International Atomic Energy Agency IWGGCR-9
International Working Group on Gas-Cooled Reactors
Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors
Dusseldorf Federal Republic of Germany
16-19 April 1984
hosted by Bundesministerium fur Forschung und Technologie
31/42 Please be aware that all of the Missing Pages in this document were originally blank pages Introduction
The Specialists' Meeting on "Heat Exchanging Components of Gas- cooled Reactors" was held at the Ministry of Economic and Trans- port of the State North Rhine Westphalia, Diisseldorf, FRG 16-19 April 1984.
The meeting was sponsored by the IAEA on the recommendation of the International Working Group on Gas-Cooled Reactors and was hosted by the Federal Ministry of Research and Technology of the Federal Rebublic of Germany.
The meeting was attended by 62 participants from Austria, France, Federal Republic of Germany, Japan, Poland, Sweden, Switzerland, United Kingdom of Great Britain and Northern Ireland and the United States of America.
The objective of the meeting was to provide a forum, both formal and informal, for the exchange and discussion of technical infor- mation relating to heat exchanging and heat conducting components for gas-cooled reactors.
The technical part of the meeting was divided into eight subject sessions:
I. Heat exchanging components for process heat application - design requirements and r/d programmes
II. Status of the design and construction of intermediate He/He heat exchangers
III. Design, construction and performance of steam generators
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V. Metallic materials and design codes
VI. Design and construction of valves and hot gas ducts
VII. Description of component test facilities and test results
VIII. Manufacturing of heat exchanging components
A total of 38 papers were presented by the participants on behalf of their organizations during the meeting, and an opportunity for open discussion of the paper topic followed each presentation. Session I
Heat exchanging components for process heat application - design requirements and R & D programmes
No 1 Status of the R&D program in the field of the heat carrying and heat transfer components of the PNP- project H. Mausbeck, W. Jansing; Interatom GbmH; FRG
No 2 Design requirements on HTR main components for process heat application K. Dumm; Interatom GmbH; FRG
No 3 Helium/helium heat exchangers and hot-gas ducts for the PNP-project according to the BBC/HRB-concept H. Schmitt, B. Jiirgens, J. Knaul; Hochtemperatur-Reak- torbau GmbH; FRG
Session II Status of the design and construction of intermediate He/He heat exchangers
No 4 Recent research and development of intermediate heat exchanger for VHTR plant A. Shimizu, N. Matsumura, H. Nishikawa, S. Yamada; Industries, LTD; Japan
No 5 Development of a helium/helium intermediate heat ex- changer (He/He-IHX) with helical coil tube bundle A. Czimczik; L. & C. Steinmuller GmbH; FRG G. Hirschle; Gebr. Sulzer AG; Switzerland
No 6 Improved spacers for high temperature gas-cooled heat exchangers L. A. Nordstrom; Swiss Federal Institute for Reactor Research; Switzerland
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No 7 Life time test of a partial model of HTGR helium-helium heat exchanger M. Kitagawa, H. Hattori, A. Ohtomo, T. Teramae, J. Hamanaka, M. Itoh, S. Urabe; Ishikawajima-Harima Heavy Industrie Co., LtD.; Japan
No 8 Development, construction and analysis of the URKO intermediate heat exchanger, R. Exner, M. Podhorskiy; Balcke-Durr AG; FRG
No 9 Development of a new type of high-temperature-insula- tion-material and its application in the PNP-project R. Burger, R. Ganz; Didier-Werke AG, FRG
No 10 Seismic analysis of a helical coil type heat exchanger I. Nishiguchi, 0. Baba, H. Yatabe; Japan Atomic Energy Research Institute; Babcock Hitachi K. K.; Japan
Session III Design, construction and performance of steam generator
No 11 Design and development of steam generators for the AGR power stations at Heysham II/Torness A. N. Charcharos, A. G. Jones; National Nuclear Corpo- rat ion Ltd.; UK
No 12 Monitoring and performance analysis of AGR boilers during commissioning and power raising M. El-Nagdy, R. M. Harrison; Nuclear Engineering Department; Babcock Power Ltd.; UK
No 13 Experience with the commissioning of helically coiled advanced gas-cooled reactor boilers D. B. Kettle; CEGB-Generation Development and Construc- tion Division; UK
- 2 - No 14 Investigations of the gas-side heat transfer and flow characteristics of the AGR steam generators J. Lis; Central Electricity Research Laboratories; UK
No 15 Effect on inlet and outlet shell side flow and heat transfer on the performance of HTGR straight tube heat exchangers D.P. Carosella; GA Technologies; USA
Session IV Design, development and fabrication of steam reformers
No 16 Status of an in-line reformer design for modular HTGR R. Gluck, W. H. Whitling, A. J. Lipps; General Electric Company; USA
No 17 Development and fabrication of a helium heated steam reformer W. Panknin, W. Nowak; L. & C. Steinmuller GmbH; FRG
No 18 Assembly and operation experience of the EVA II - steam reforming bundle H. F. NieBen, R. Harth; Kernforschungsanlage Julich GmbH; W. Kessel; Rheinische Braunkohlenwerke AG; FRG
Session V Metallic materials and design codes
No 19 Evaluation of materials for heat exchanging components in advanced helium-cooled reactors F. Schubert; Kernforschungsanlage Julich GmbH; FRG
No 20a Pressure vessel design codes: A review of their applica- bility to HTGR components at temperatures above 800°C P. T. Hughes; General Electric Company; USA K. Bieniussa; Gesellschaft fur Reaktorsicherheit GmbH, H. H. Over; Kernforschungsanlage Julich GmbH; FRG
- 3 - No 20b Status of design code work for metallic high temperature components K. Bieniussa; Gesellschaft fur Reaktorsicherheit; FRG H. J. Seehafer; Interatom? FRG H. H. Over; Kernforschungsanlage Jiilich GmbH; FRG P. Hughes; General Electric Company; USA
No 21 Oxide films on austenitic HTR heat exchanger materials as a tritium barrier H. P. Buchkremer, R. Hecker, H. Jonas, H. J. Leyers, D. Stover; Kernf orschungsanlage Jiilich GmbH; FRG
No 22 Effect of creep-fatigue damage relationships upon HTGR heat exchanging design D. P. Carosella, M. M. Kozina, J. H. King, M. Basol; GA Technologies; Combustion Engineering Inc.; USA
Session VI Design and construction of valves and hot gas ducts
No 23 The Klinger hot gas double axial valve J. Kruschik; Klinger Engineering; Austria; H. Hiltgen; Interatom GmbH; FRG
No 24 Two layers thermal insulations tests for designing of hot gas ducts T. Nakase, S. Midoriyama, K. Roko, A. Yoshizaki; Kawasaki Heavy Industries, Ltd.; Japan
No 25 Status of the development on hot gas ducts for HTRs H. Stehle, E. Klas; Interatom GmbH; FRG
No 26 Graphite and carbon-carbon components for hot gas ducts in the HTR G. Popp, U. Gruber, H. Boder, K. Janssen? Sigri Elektrographit GmbH; FRG
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No 27 Research on thermal insulation for hot gas ducts P. Brockerhoff; Kernforschungsanlage Jiilich GmbH; FRG
Session VII Description of component test facilities and test results
No 28 Facility for endurance tests of thermal insulations R. Mauersberger; Hochtemperatur-Reaktorbau GmbH; FRG
No 29 Construction and performance tests of helium engineering demonstration loop (HENDEL) for VHTR M. Hishida, T. Tanaka, H. Shimomura, K. Sanokawa; Japan Atomic Energy Research Institute; Japan
No 30 Testing of high-temperature components in the KVK W. Jansing; Interatom GmbH; FRG
No 31 WKV-operation experiences with heat exchanging compo- nents of a nuclear gasification pilot plant R. Kirchhoff, K. H. van Heek; Bergbau-Forschung GmbH; FRG
No 32 The test facility EVA II/ADAM II - Description and operational results R. Harth, H. F. Niessen, V. Vau; Kernforschungsanlage Jiilich GmbH; FRG
No 33 Modification of the AVR to a versatile nuclear test facility for high temperature components H. Barnert, N. Kirch, E. Ziermann; Kernforschungsanlage Jiilich GmbH; Arbeitsgemeinschaft Versuchsreaktor GmbH; FRG
No 34 Heat removal by natural circulation in gas-cooled rod- bundles M. Hudina; Swiss Federal Institute for Reactor Research; Switzerland
5 - Session VIII Manufacturing of heat exchanging components
No 35 Manufacture of steam generator units and components for the AGR power stations at Heysham II/Torness J. R. Glasgow, K. Parkin? Nuclear Systems Limited; UK
No 36 The use of bimetallic welds in the THTR steam genera- tors U. Blumer, H. Fricker, S, Amacker; Sulzer Brothers Ltd.; Switzerland
No 37 GMA-narrow gap welding of PNP-hot gas collectors K. Iversen, A. Palussek; Interatom GmbH; FRG
No 38 Forged hollows (alloy 617) for PNP-project F. Hofmann; Vereinigte Deutsche Metallwerke AG; FRG SUMMARY
Session I: Heat exchanging components for process heat application - design requirements and R & D programmes
Session Chairman: E. Balthesen KFA-Jiilich - PTH -, FRG
The first session of the Specialist's Meeting comprised survey papers describing the application of heat exchanging components in nuclear process heat plant concepts in the Federal Republic of Germany. These concepts are based on He-cooled high tempe- rature reactors as nuclear heat sources for steam gasification and hydrogasification of coal and for a thermochemical heat pipe system. The consideration of the strong interaction between nuclear heat source and application process illuminated the func- tion and the main design requirements for the heat ducting and exchanging components. The component designs are partially de- pending on different reactor designs. INTERATOM presented the HTR modular concept consisting of a number of independent small pebble bed reactor units with a thermal output of 170 MW each. Each modul can be combined either with a steam reformer in the primary circuit or with a He/He heat exchanger.
The system company HRB presented a modified THTR pebble bed reactor embedded in PCRV. This concept fundamentally applies an intermediate heat transfer circuit with a tandem heat ex- changer design consisting of a high temperature and a low tempe- rature unit each.
The process requirements, the gaseous media, the high temperatures and the necessary application of new materials mean a great techni- cal challenge for the development, design work as well as testing.
The result of an extensive component development up to now - al- though in a relatively early stage - indicate that manufacturing and supply will be feasible in time. No prohibitive questions
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have been discovered up to now. The present program may be considered as an effective iteration process between design work, materials qualification, code evaluation, testing and requirements specification which ultimately will also contri- bute to establish the necessary licensing criteria.
Moreover, the papers included descriptions of the capability and experience from the two large-scale test facilities, the thermochemical heat pipe system demonstration plant EVA II at KFA-Jiilich, in which successful steam reformer testing has been performed, and the multi-purpose component test facility KVK at INTERATOM Bensberg,where in particular tests of heat ex- changers in a 10 MW(th) scale are being prepared. SUMMARY:
Session II: Status of the design and construction of intermediate He/He heat exchanger
Session Chairman: K. Parkin, NEI, UK
Mr. Shimizu described the recent research and development which sought and achieved improvements both in potential performance and economy of manufacture. The use of helically rolled fin tube, automatic orbital welding and simplified assembly techniques were highlighted.Discussion included limitation of NDE on helical tubes in service and insulation on the central tube.
A comprehensive presentation by Mr. Czimczik described the fa- brication and quality control procedures of the experimental helical heat exchanger. The importance of results from experimental tests due to commence early in 1985 was stressed to allow follow up work.
Mr. Nordstrom presented work done in the selection of an improved tube spacer grid. Five different geometries were assessed and one type finally chosen. The comparison of predicted and experimental performance was considered sufficiently accurate. The work has been usefully communicated to the USA.
Life time testing of a partial model was the subject of Mr. Itoh. Three models were used, experiments were described together with results. Good agreement was observed between predicted and ex- perimental lives. Data on mechanical strength and other metall- urgical features was obtained from the experiments. The design code used was considered appropriate for this model.
Mr. Podhorsky gave a comprehensive description of the U tube compact heat exchanger.Comparison was made with previous heat exchanger and design, development work was shown to have brought
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about improvements and made savings in manufacture. The results indicated insensitivity to gas flow induced vi- bration and it was suggested that a successful conclusion will be arrived at within the timescale laid down i.e. Nov 1981 - Dec 1987.
Insulation aspects of the U tube compact heat exchanger were covered in detail by Mr. Burger. The basis of this work has provided solution for other nuclear components.
The session was concluded by a presentation from Mr. Hishida who described a finite element model of the structure of an IHX helical tube bundle, vessel and centre pipe. The model included gap elements to cover for the practical arrangements where gaps exist to accomodate thermal expansion. In addition a stopper arrangement was proposed on the centre pipe.
The seven papers presented fully identified the status of design and construction of intermediate He/He heat exchangers. Several comparisons were made with earlier designs all resulting in im- provements particularly regarding economy of manufacture. SUMMARY
Session III: Design, Construction and Performance of Steam Generators
Session Chairman: R. Gluck General Electric, USA
This session consisted of the presentation of five papers dealing with design, construction and performance testing of a large of steam generator components designed for application to High Temperature Gas Reactors.
The first four papers given by Messrs. Charcharos, Jones, El-Nagdy, Kettle and Lis dealt with the United Kingdom ex- perience with AGR power stations and showed the maturity steam generator design evolution with particular emphasis on correlation of analytical methods with performance tests.
The fifth paper given by Mr. Carosella examired the effects of 90 degree bends on straight tube heat exchanger flow and heat transfer. Draft SUMMARY:
Session IV: Design, development and fabrication of steam reformer
Session Chairman: M. Itoh,- IHI, Japan
In this session, results of recent design, fabrication and operation of steam reformer were presented. Summary of them is as follow:
1. Status of a Reformer Design for a Modular HTGR in an In-Line Configuration
The focus of the Modular Reactor System (MRS) effort is the development of generic nuclear heat source capable of supplying heat to either a steam generator/electric cycle or a high tempe- rature steam/methane reforming cacle. This paper presents the results of recent design and analytical studies conducted to evaluate the feasibility of using a steam/ methane reformer in a vertical In-Line (VIL) arrangement re- former with the generic nuclear heat source.
2. Development and Fabrication of a Helium-Heated Steam Reformer
1. For PNP a 96 MW steam reformer is to be designed. It consists of a bundle with about 300 reformer tubes which are shrouded by guide tubes. 2. In order to verify experimentally the design, to check the thermohydraulic layout and to gain know-how for the fabri- cation a small test reformer ( 5 MW, 18 tubes) is in pro- duction now. 3. The reformer, fabricated mainly of Inconel 617 will be be tested in EVA II. SUMMARY :
Session \l: Metallic materials and design codes
Session Chairman: G. Hirschle SULZER, Switzerland
Five papers were presented in this session that treated a very decisive part of the HTGR development.
In the first presentation Mr. Schubert reviewed the status of material testing for the PNP project. Data of four materials which were and still are being investigated in respect to application for heat exchanging components were presented. Emphasis was put on Inconel 617 (Nicrofer 5520 Co) for which an extrapolation of the data to about 70 000 hrs is possible today. Effects of corrosion and other failure mechanisms in different atmospheres were discussed in detail.
The present status of design codes was reviewed by Mr. Bieniussa in the second contribution. Although the existing codes have been established for materials used at temperatures below 816 °C, the authors conclude that the available design criteria and material data are sufficient to fabricate and operate compoments of a prototype plant. In addition Mr. Gluck pointed out in his presentation that ASME Code Case N-47 should be maintained as a basis for calculations, and that there is a strong need for design rules simple enough to be used during the layout period of components.
Mr. Jonas presented in the fourth paper their investigations on in situ growing oxide layers suitable as tritium barrier. Although self healing of defective layers is good, the barrier is not yet sufficient in the case of a cold start of the component.
Referring to recent material test data Mr. Carosella proposed in the last presentation to use a more conservative relationship for creep fatigue analysis than given in ASME Code Case N-47. The design of the HTGR steam generator, it is concluded, has sufficient margin to fulfil this more stringent relationship. page - 2 -
3. Assembly and Operation Experience of the EVA II Steam Reforming Bundle
1. The results of behaviour of the bundle consisting 30 tubes during the 6.000 operation hours at tempe- rature above 800 °C were presented. 2. To simulate the plug-off of tube in a nuclear heated steam reformer, some experiments were carried out with groups of tubes shut-off on the process gas side.
3. Preliminary results of the after operation inspection were presented. SUMMARY
Session VI: Design and construction of valves and hot gas ducts
Session Chairman: Mr. H. Witulski Ministry of Economic and Transport of the State North Rhine Westphalia, Duesseldorf, Federal Republic of Germany
The inner surfaces of prestressed reactor vessels and hot gas ducts of Gas-Cooled High Temperature Reactors need internal insulation to protect the pressure bearing walls from high temperatures. The design parameters of the insulation very closely depend on the reactor type. For the primary circuit a gas duct with fully ceramic insulation has been developed by INTERATOM. This insulation consists of a graphite gas liner, a ceramic fibre insulation and a metallic pressure tube (support tube). The gas liner is held in position in the support tube by means of patented ceramic spacers. Coaxial flow is used in the area of the hot gas duct, i.e. the suppport tube is cooled by the gas flowing back (42 bar, 300 C). The thermal elongations among the reactor and the heat-exchangers are compensated by means of two angular bellows.
The internal insulation of the secondary circuit consists of a metallic internal lining tube and a ceramic fibre insulation which is wrapped around the former. V-shaped spacers in the insulation prevent axial flow in the insulating material. The complete insulating system is mounted in a support tube. This unit is slid into the pressure tube as a slice-in unit. Therefore the periodic inspection is provided. This insulating system is used for all the components in the intermediate circuit such as elbows, compensators and T-pieces. - 2 -
An extensive experimental programme has to be worked through in order to qualify the hot gas ducts. In addition to the tests on prototype components in INTERATOM's component test circuit (KVK), several tests are being carried out on components and component parts, by the project-partners, namely the Juelich Nuclear Research Plant (KFA), the Juelich High Temperature Reactor Construction (HRB) and INTERATOM (Bensberg). For the primary hot gas duct thermocycling tests have been successfully carried out by INTERATOM.
For the secondary hot gas duct an original section of about 6 m has been tested in the KVK up to 2500 hours, with a max. temperature of 950°C.
The experimental work at KFA-Juelich was described by Mr. Broeckerhoff. The work was started at KFA in 1971 for HHT using three test facilities. At first metallic foil insulation and stuffed fibre insulating systems the hot gas ducting shrouds of which were made of metal have been tested. Because of the elevated helium temperature in case of PNP and the resulting lower strength of the metallic parts the interest was directed to rigid ceramic materials for the inner shrouds and spacers. This led to modified structures which were designed by INTERATOM and which were also tested at KFA.
The main object of the investigations was to study the influence of temperature and pressure on the thermal efficiency of the structures. At first the insulating systems, e.g. the INTERATOM-design with spherical spacers, their instrumentation and the experiments will be described. After that the temperature distributions within the insulation and at the pressure tube are presented. Thermal fluxes and effective thermal conductivities in axial and circumferential direction of the pressure tube are given and compared.
Coaxial double walled piping is planned to be used for a primary cooling system piping of the Very High Temperature Gas-cooled Reactor (VHTR) of - 3 -
JAERI as described by Mr. Nakase in his paper. The pipe consists of an outer pressure pipe for the reactor inlet gas flow and an inner pipe with internal insulations for the reactor outlet gas flow. The internal insulations are designed to consist of two layers; metal insulation is in the extremely inside for the higher temperature gas and fibrous insulation is between the inner pipe and the metal insulation. The thermal characteristics of inner pipe with two insulation layers are necessary for the designing of the primary cooling system piping.
Thermal characteristics tests were performed by using Kawasaki's helium test loop (KH-200), which has two hot gas ducts specimen of 267 A mm diameter and 5000 mm length with simulated two layers insulation. The one is installed in a horizontal position and the other is in a vertical position. The tests were conducted at the temperature of 500 to 1000° C, the pressure of 20 and 40 kg/cm G, and the flow rate of 100 and 200 g/s.
The distributions of temperature and heat flux at the surface of the ducts are confirmed to be within an allowable range. The test results were analyzed, and useful design data of the metal insulation and the fibrous insulation were obtained.
Valves can be a containment shut-off system or a safety device for the components of the reactor plant.
The main design features are:
Compact design due to a special valve port configuration with very low pressure loss Automatic shut-off in case of accidental pressure loss in the circuit The flexible metallic sealing system Gas-static bearings and their basic function - Cooling system The development of the gas-static bearings together with frequency analysis and manufacturing details were explained in a paper from Mr. Kruschik.
The paper also included results of stress analysis with finite-elements and its influence on the design specially of the highly stressed high-temperature components together with the analysis of the influence of uneven temperature distribution on the valve geometry. SUMMARY
Session VII: Description of component test facilities and test results
Session Chairman: M. Robin CEA, France
In this session dealing with the description of component test facilities and test results 7 papers were presented and dis- cussed .
Mr. Mauersberger, head of the Division for Experimental Engi- neering of Hochtemperatur-Reaktobau GmbH reported on their "Facility for Endurance Tests of Thermal Insulations" that includes a pressure vessel designed for a 70 bar helium pressure, a circulator with magnetic bearings capable of a 4.5 kg/sec flow- rate and the required electrical heaters allowing as 950°C He temperature and oil or water coolert in order to obtain the specified temperature gradient through the experimented insul- ation. The gas purification systems fitted for analytical mea- surement of the gas impurities. The control system provides for the thermal cycling of the insulation. A typical insulation section composed of a pressure tube lined with a graphite tube wrapped by a fibrous insulation material with a density of 130 kg/m3 has been tested.
The temperature profiles within the insulation and along the pressure tube were presented for various helium temperatures and an important heat leakage effect was apparently determined signi- ficant heat losses (up to 800 W/m2 with He at 950°C). The exper- ience gained over a peroid of more than one year shows that oper- ational requirements can be fullfilled completely.
Mr. Hishida, deputy manager of the HENDEL operating division in the Department of High Temperature Engineering (JAERI) described the "Construction and Performance Tests of the Helium Engineering Demonstration Loop (HENDEL) for VHTR". - 2 -
- two hot gas ducts were installed and tested in HENDEL. Looking at the the cross section one see a liner tube made of Hastelloy X which can be exposed to 1000°C He, a thermal insulation of the fibrous ceramic insulation divided into three sublayers by stainless steel foils, and a pressure tube made of miled steel. Temperature distributions in both axial and circumferential directions were measered with He at 10 to 40 bar, up to 950°C flowing at rates of 0.5 to 3.5 kg/sec. The maximum temperature variation range in the circumferential direction was only 35°C. In the axial direction the temperature distribution of the pressure tube was almost uniform. The effective thermal conducti- vity for both horizontal and vertical duct sections are reported with values 25 to 30 % higher in the latter case.
Thermal performances of the helium gas coolers, with U- tube or straight tube bundles are in fair agreement with the values cal- culated from the equations in the literature (Donohue). Thermal performances of the heliumheaters (graphite rods heates by Joule effect) were reported as satisfactory for the 3 .000 h operation.
Mr. Jansing who is in charge of the INTERATOM helium and sodium facilities at Bensberg described the "Testing of High Temperature Components in the KVK" facility of 10 MW(th) that is capable of a 4.3 kg/sec flowrate of helium up to 950°C under 40 bar pressure. After 5.000 h of operation including 1.200 h at temperatures in the 900-950°C range it is concluded that the helium tightness of the system is good (loss 1 kg/day/) the performances of the gasfired and of the electrical heaters are satisfactory helium circulators, steam generator and steam heated helium preheater operate reliably the materials selected for internal insulation in hot ducts are adequate no inadmissible vibration level is recorded - the helium atmosphere is easily adjusted to the required purity _ T _
a stable chromium oxide protective film developed on material surfaces in the temperature range 900-950°C the new process instrumentation and control system Teleperm-M is easy to operate and reliable - fast and effective repairs of defects in the helium heaters are possible.
In summarizing it KVK operators very satisfactorily and can be easily prepared for other series of component tests.
Mr. Kirchhoff from Bergbau-Forschung GmbH (Essen), leader of the gasification Pilot Plant reported on the "Operating Experience with Heat-Exchanging Components of the Semi-Technical Pilot Plant for Steam Gasification of Coal using heat from HTR" within the PNP project since 1976. The heat supplied by the HTR is simulated by an electrically heated helium loop operating at pressures up to 40 bar and temperatures up to 950°C. In the gasifier coal (in- cluding caking coal) has been gasified in a 1 m2 fluidized bed with a height of 4 m at a rate up to 6.4 t/day with a total of 26.600 h of operation. Besides the measurements of data on heat transfer, gasification kinetics, insulating material and heat transfer components suitable for operation at high temperature (up to 950°C) were developed and tested. Particularly interesting is the good behaviour of the helium tubes immerized in the fluidized bed of 0.3 mm coal particles. In conclusion it seems that all the experience required for the design, construction and operation of a commercial size plant is available.
Mr. Harth, leader for Project EVA 11/ADAM II at KFA Julich, des- cribed this facility and the related operational results. The helium circuit represents a complete primary loop of a HTR for process heat applications, the core being simulated by an elec- trical heater with a maximum capacity of 11 MW(th) allowing for 950°C helium temperature. _ 4 -
This high temperature He flows outside the tubes in a steam re- former, exits at 650°C and enters a steam generator which also includes an integrated gas circulator with a flowrate of 4 kg/sec. The aim of the tests was to determine the characteristics of the operation in a wide range of parameters (He pressure 15 to 40 bar, He temperature 800 to 950°C, methane flowrate 0.18 to 0.66 kg/sec. The charge of catalyst (Raschingrings) by vacuum extraction has been demonstrated, and also the replacement of a single reformer tube without removal of the bundle. Dissassembling and reinstallation of the reformer bundle has been also successfully performed. In total the helium system has been operated for 7.800 h including a 5.660 h heat transport cycle. A new steam reformer is prepared for testing whose test represents the last experimental step before a nuclear demon- stration is considered.
Dr. Barnert, who is in charge of Process Heat Application Studies at the Institute for Reactor Development (KFA-Julich) described the "Modification of the AVR for High Temperature Process Heat Systems Demonstration". He showed that the AVR structure and layout allows for this modification that is important for establishing the feasibility of coupling a nuclear reactor to a chemical plant. The major part of the discussion which followed was concerned with problems other than technical but equally important (strategy of introduction and development of nuclear process heat systems).
The last paper in the session was presented by Mr. Nordstrom, leader of the Heat Exchanger Research Group of the Swiss Federal Institute in Wiirenlingen and deals with "Heat removal by natural circulation in gas-cooled reactor bundles", a fuel bundle for a
CO2 cooled GCFR being chosen to illustrate the calculation. The conclusion is that the removal of decay heat by natural circulation of gas under pressure can be ensured up to 1.5 % of the nominal power. The temperature distributions in both axial and radial directions were presented. It is important to take into account the metal thermal conductivity in order to have a fair agreement between experiment and calculated. - 5 -
As a general conclusion, many facilities for testing heat transfer components are available and operate satisfactorily in the range of temperature and pressure of interest for the HTR development. Their operation shows that there exists no technical problem that cannot be solved before the end of this decade. SUMMARY:
Session VIII: Manufacturing of heat exchanging components
Session Chairman: J. Kruschik, KLINGER, Austria
The first paper by Mr. Glasgow and Mr. Parkin gave in a review of the design and the development: of manufacturing methods for steam generators units.
The second paper by Mr. Blumer dealt with the method for welding ferritic to austenitic tube material. The choice of the location of the weld within the heating surface of the bundle was described together with metallurgical considerations, the use of filler ma- terial, stress analysis, etc.
The third paper by Mr. Iversen gave a thorough of GMA-narrow gap welding and the inspection procedure. This new technic was development for welding collections of Inconel 617.
Paper four, presented by Mr. Hofmann explained in full detail the new process for production of forged hollows for collectors in alloy 617.
These four papers gave a good review on modern manufacturing tech- nics for nuclear components. List of Participants and Observers
AUSTRIA
H.-O. Faber Österreichisches , Forschungszentrum Seibersdorf GesmbH. Lenaugasse 10 1082 Wien
FRANCE
M.G. Robin C.E.A., CEN de Saclay DEDR/CRG B.P.No. 2 F-91190 GIf-sur-Yvette
FEDERAL REPUBLIC OF GERMANY
M. Andler INTERATOM GmbH. Postfach n-5060 Bergisch-Gladbach 1
E. Arndt Hochtemperatur-Reaktorbau GmbH Postfach 5360 D-6800 Mannheim 1
E.R. Balthesen Kernforschungsanlage Jü1 ich GmbH. Projektträger HTR Postfach 1913 D-5170 Jülich 1
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H. Barnert Kernforschungsanlage Jü1 ich GmbH Institut für Reaktorentwicklung Postfach 1913 D-5170 Jü 1 ich
F. Berdan Klinger Eng ineering Wiener Straße 17 A-2351 Wr. Neudorf
K. Bienussa Gesellschaft für Reaktorsicherheit mbH Schwerdnergasse D-5000 Köln
P. Bröckerhoff Kernforschungsanlage Jü1 ich GmbH Institut für Eaktorbaue lemente Postfach 1913 D-5170 Julien
R- Burger Didier-Werke AG Energie-Techni k Didierstraße 31 D-6200 Wiesbaden 12
A. Czimczik L. & C. Steinmüller GmbH. P.O.Box 100855 D-5270 Gummersbach
W. Dietz INTERATOM GmbH. Postfach D-5060 Bergisch-Gladbach
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K.G. Dumm INTERATOM GmbH. Postfach D-5060 Bergisch-Gladbach R. Exner Balcke-Dürr AG Hornberger Straße D-4030 Ratingen
R. Ganz Didier-Werke AG Energie-Technik Didierstraße 31 D-6200 Wiesbaden 12
U. Gruber von Rehlingen Straße 48a D-8902 Neusäss/tfestheim
R. Harth Kernforschungsanlage Jü1 ich GmbH Projekt Nukleare Fernenergie Postfach 19T3 D-5170 Jü 1 ich
F. Hofmann Vereinigte Deutsche Metallwerke AG Geschäftsoereich Nicxel-Technologie Postfach 18 20 Plettenberger Straße 2 D-5980 Werdohl
K. Iversen INTERATOM GmbH. Postfach D-5060 Bergisch-Gladbach
H. Jonas Kernforschungsanlage Jü1 ich GmbH Postfach 1913 D-5170 Jülich
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M. Jurgens Hochtemperatur-Reaktorbau GmbH Postfach 1913 D-6800 Mannheim
W.T. Jans ing INTERATOM GmbH. Postfach D-5060 Bergisch-Giadbach
J.G. Keeble • Deutsche ICI GmbH. Lyoner Strafle 36 6000 Frankfurt/Main 71
M. Schafer Kernforschungsanlage Julich GmbH Projekt Entwicklungsarbeiten fur Hochtemperaturreaktor-Anlagen Postfach 1913 D-5170 Julich
R. Kirchhoff Bergbau-Forschung GmbH P.O. Box 130140 0-4300 Essen 13
J. Knaul Hochtemperatur-Reaktorbau GmbH Postfach 5360 D-6800 Mannheim 1
J. Kruschik Kli nger Engi neeri ng Wiener StraBe 17 A-2351 Wr. Neudorf
R. Mauersberger Hochtemperatur-Reaktorbau GmbH Postfach 2080 D-5170 Julich
- 5 - - 5 -
H.J. Mausbeck INTERATOM GmbH. Postfach D-5060 Bergisch-Gladbach
Th. Monsau Der Minister für Wirtschaft, Mittelstand und Verkehr des Landes Nordrhein-Westfalen Haroldstraße 4 D-4000 Düsseldorf
H.F. Nießen Kernforschungsanlage Julien GmbH. Institut für Reaktorentwicklung Postfach 1913 D-5170 Jülich
W. Nowak L. &. C. Steinmüller GmbH. Fabrikstraße 1 P.O. Box 100855 D-5270 Gummersbach
H.-H. Over Kernforschungsanlage Jülich GmbH.- IRW P.O. Box 1913 D-5170 Jülich
W.K. Panknin L. & C. Steinmül1er GmbH, Fabrikstraße 1 P.O. Box 100855 D-5270 Gummersbach
M. Podhorsky Balcke-Dürr AG Hornberger Straße D-4030 Ratingen
- 6 - 3/ - 6 - ^ J
G. Popp SIGRI Elektrographit GmbH W.-V.-Siemens-Straße 18 D-8901 Meitingen
H. Reutler INTERATOM GmbH. Postfach D-5060 Bergisch-Gladbach
H. Schmitt Hochtemperatur-Reaktorbau GmbH Postfach 5360 D-6800 Mannheim 1
F. Schubert Kernforschungsanlage Jü1 ich GmbH Institut für Reaktorwerkstoffe Postfach 1913 D-5170 Jülich
H.J. Seehafer INTERATOM GmbH. Postfach D-5060 Bergisch-Gladbach
H. Stehle INTERATOM GmbH. Postfach D-5060 Bergisch-Gladbach
H. Stein Didier-Werke AG Energie-Technik Didierstraße 31 D-6200 Wiesbaden 12
-7 - - 7 -
H. Witulski Der Minister fur Wirtschaft, Mittelstand und Verkehr des Landes Nordrhein-Westfalen Haroldstrafte 4 D-4000 Dusseldorf
JAPAN
Mr. M. Hishida Deputy General Manager HENDEL Operation Devision, Department of High Temperature Engineering Tokai Research Establishment, Japan Atomic Energy Research Institute Tokai-mura, Ibaraki-ken 319-11
Mr. M. Itoh Section Manager, Energy Develoment Center Ishikawajima-Harima Heavy Industries Co., Ltd Marunochi 1-6-2 Chiyoda-ku, Tokyo, Japan
Mr. Akira Shimizu Nuclear Plan Engineering and Designing Section Jagasaki Shipyard and Engine Works, Mitsubishi Heavy Industries, Ltd. Nagasaki Shipyard & Engine Works 1-1 Akunouramachi Nagasaki 850-91
Nr. T. Nakase Kawasaki Heavy Industries, Ltd. 4-25 Minamisuna 2-chome, Koto-ku Tokyo
POLAND
E. Obryk Institute of Nucleare Physics Radzikowskiego 152 31-342 Krakow Poland
- 8 - - 8 -
SWEDEN
R.I. Ekholm Studsvik Energiteknik AB S-61182 Nykoeping
SWITZERLAND
U.R. Blumer Sulzer Brothers Ltd. Ch-8400 Winterthur
G. Hirschle Sulzer Brothers Ltd. Nuclear Engineering Ch.-8401 Winterthur
L.A. Nordstroem Eidg. Institut für Reaktorforschung CH-5303 Wuerenlingen
UK
A.N. enarenaros Reacotr Engineering Department National Nuclear Cooperation Ltd Booths Hall Chelford Rd. Knutsford, Cheshire WA16 8QZ
M.M. El-Nagdy Nuclear Engineering Babcock Power Ltd. 165 Great Dover Street London SEI 4YB
- 9 - - 9 -
J.R. Glasgow NEI Nuclear Systems Ltd P.O. Box 13 Saltmeadows Road Gateshead Tyne & Wear NE8 1YZ
A.G. Jones Reacotr Engineering Department Natural Neclear Cooperation Ltd Booths Hall Chelford Rd. Knutsford, Cheshire WA16 8QZ
D.B. Kettle C.E.G.B., GDCD, BARNETT WAY BARNWOOD, Gloucester, GL4 7RS
J. Lis Central Electricity Research Laboratories Kelvin Avenue Leatherhead , Surrey
K. Parkin NEI Nuclear Systems Ltd P.O. Box 13 Saltmeadows Road Gateshead Tyne & Wear NE8 1YZ
- 10 - USA
D. Carosella GA Technologies Incorporated P.O. Box 81608 San Diego, California 92138
R. Gluck Advance Reactors Systems Dept General Elektric Company P.O. Box 508 Sunnyvale, California 94086
IAEA
J. Kupitz Advanced Nuclear Power Technology Section Division of Nuclear Power WagramerstraBe 5 P.O. Box 100 A-H00 Vienna • - • No. 1
IAEA Specialists' Meeting on XA0055810 Heat Exchanging Components of Gas-Cooled Reactors Dusseldorf, 16-19 April 1984
Status of the R + D Programme in the Field of the Heat Carrying and Heat Transfer Components of the PNP Project
Hans Mausbeck, Walter Jansing INTERATOM GMBH, FRG
1. The Project: Historical Development and Motivation Since 1972 the partners Bergbau-Forschung GmbH, the Nuclear Research Center Jtilich GmbH and the Rheinische Braunkohlen- werke AG have been involved in the project "Development of processes for the conversion of solid, fossil fuels using heat from high temperature nuclear reactors".
In 1976 the partnership was extended to include the Gesell- schaft fur Hochtemperatur-Reaktortechnik mbH (GHT) and the Hochtemperatur-Reaktorbau GmbH (HRB). The new contract includes also the complete design work for the construction of a prototype plant, in order to demonstrate the feasibility of the gasification of lignite and hard coal using nuclear process heat.
At this point we should consider the motives for setting-up the PNP project and continuing intensive work on it, motives which were valid in the past and which still apply today: The balance of payments of the national economy of the Federal Republic of Germany is essentially determined by the fact that more than 50 % of the primary energy demand is covered by import oil and natural gas. In order to lower these expenses, an attempt must be made to utilize our own fossil energy reserves, namely lignite and hard coal, more effectively. -2-
The conversation of these energy resources to gaseous and liquid products enables us to do so. Of all the nuclear reactors developed today, the high temperature reactor is predestinated to play a key role, as it can supply the heat which is necessary for the conversion processes at the required high temperatures of between 800° C and 950° C.
In conjunction with special process technology, this leads to a considerable reduction of the pollutant emissions of
SO , CO2 etc. and of dust.
Above all, the COp emission; which is much lower than in autothermal gasification processes, deserves special mention,
because of the CO? influence on the temperature increase at the earth's surface.
2. The Plant Concept
In order to clarify the functions of those components which are the subject of the following presentations, I would like to consider the PNP plant with both gasification processes, namely steam gasification and hydrogasification of coal, in detail.
Let us first consider the steam gasification of coal (Fig. 1). The reaction of steam with hard coal requires high temperature heat since it is endothermic. For this process a secondary helium loop is advisable for safety reasons.
The secondary helium is heated to 900° C in the He/He inter- mediate heat exchanger and enters the gas generator at approximately this temperature. The helium is cooled to around 815° C here because of the carbon-steam reaction.
The helium leaving the gas generator is cooled in the process steam superheater and then conveyed to the steam generator. -3-
The raw gas is subjected to a number of further process steps but it would take too long to describe these in detail. They are, however, dependent on the desired end product.
In the case of the hydrogasification of coal (Fig. 2), the dried and ground lignite is fed into a gas generator into which pure hydrogen flows. The reaction taking place in the gasifier is exothermic. The residual sulphur-free coke can, for example, be further gasified in a pressurized coal gasification plant or can be used as high-grade coke. After coooling and purification, the raw gas leaving the gas genera- tor is decomposed into the fractions hydrogen, methane and carbon monoxide. The H^-gas flow is supplied to the gasifier via the preheating system, the CO-flow is added to the raw and process gas, while the CH,-flow is mixed with steam and fed into the steam reformer.
The CH./HpO mixture is cracked in the tubes of the steam reformer, which are filled with catalysts, at approximately 800° C and 50 bar in accordance with
CH4 + H2O = CO + 3
CO
This endothermic reaction is maintained by the helium flow coming from the reactor which is cooled from approximately 950° C to approximately 720° C. After further cooling in the steam generator to 3 00° C, the helium is reheated to 950° C in the HTR.
As you can see from the two flowcharts of the steam gasification and hydrogasification of coal, some components are operated at such high temperatures that it is impossible to use comparable components from either conventional technology or nuclear technology. Therefore it is not astonishing that the development of components is the major issue in the R -t D work for the nuclear heat generation system. -4-
The aim of the development of the circuit components:
Helium/helium intermediate heat exchanger Steam reformer Hot gas ducts with all sub-components Hot gas valve
is to provide functional and licensable components for a PNP plant.
3. The Large-Scale Test Facilities
Two large-scale test facilities are available to the PNP project for testing the components to be discussed here: The methane reforming plant EVA II in the Nuclear Research Center Julich and the Component Testing Facility (KVK) at INTERATOM in Bergisch Gladbach.
EVA II (Fig. 3) consists of a reformer tube bundle comprising 30 tubes which are heated with helium, and a steam generator, which is also positioned in the helium loop and supplies the steam required for reforming the methane. The helium is heated to the necessary temperature by a 10 MW electric heater. It cools down while flowing through the reformer tubes from the bottom to the top. The process gas, which is a mixture of methane and steam, heats up in counterflow to the helium in
the catalyst bed and chemically converts into C0? and H_. The product gas flows upwards through a helical tube from a collecting space in the base, whereby it cools down without any further chemical reactions. On exiting from the reformer tube bundle, the helium enters the steam generator and is then returned to the heater for reheating. One bundle has already been successfully tested at temperatures of up to 950° C in EVA II. -5-
The KVK is currently operated as a single-loop facility (Fig. 4a). The necessary heat is introduced into the loop via a natural-gas-fired and a electric helium heater. It is discharged via a steam generator, whereby part of the steam is used to preheat the helium. This regenerative circuit results in a considerable reduction in the energy consumption.
In the initial test phase, which is underway at present, it is planned to concentrate on tests, for example for the hot gas ducts and the hot gas valves, and on the component test of the "hot header" of the He/He heat exchangers.
In order to test the two 10 MW heat exchangers, the KVK will be converted to a double-loop facility (Fig. 4 b) with little expenditure, wherby the existing second blower will be connected. As in a reactor plant, the primary system is allocated the heat source, and the secondary system the steam generator as heat sink, whereby the He/He heat exchangers to be tested transfers the heat from the primary to the secondary system.
Therefore the test models of the heat carrying and heat transfer components are tested in the two large-scale test facilities EVA II and KVK, namely the two variants of the He/He intermediate heat exchanger, the hot gas ducts and the hot gas valve in the KVK and the PNP bundle for the steam reformer in EVA II. These are described and their functions explained in two subsequent papers.
4. The Components I would now like to come to the general status of the test components, and then to briefly indicate the momentary situation in a comprehensive presentation.
After initial delays in the procurement of the material 2.4663 (trade names: Inconel 617 and Nicrofer 5520) it was possible to commence fabrication of the two variants of the He/He intermediate -6-
heat exchanger in 1982. At present both heat exchangers are being manufactured without any problems.
With reference to the heat exchangers, the "hot header" is the most critical component as regards its loads. In contrast to the two complete heat exchangers, which, as modular systems, need only be subjected to a functional test, it appears necessary to subject the hot header on an original scale to a simulated life test including extreme loads due to accidents. On the basis of INTERATOM's long-term experience with large- scale plants, we have come to the conslusion that an operating time of 3000 hours is adequate for a heat exchanger test. The manufacturing work on the test object was delivered in November 198 3 and has been installed in the KVK. The tests in the KVK were started in February.
Production of the 5 MW test component of the steam reformer is on schedule, delivery is planned for the end of 1984. The planning of the instrumentation for the test component was coordinated with the Nuclear District Heating (NFE) project. Work on the major problem, the spacers for the reforming tubes, has led to the following solution: A cladding tube (cladding tube gas duct concept) is allocated to each reforming tube for flow guidance on the helium side.
On the basis of the available R + D results to date, the graphite gas liner with fibre insulation and spherical spacers has been selected for the primary hot gas duct. A second proposal for the hot gas duct is being followed up in the form of the cover plate concept with CFC coating.
The experimental examinations are concentrated on the graphite gas liner. A test tube has been tested in the high pressure channel of KfA Julich to determine the effective coefficients of thermal conductivity and was subsequently installed in the ADI loop of HRB for long-term tests at operating temperature. — 7 —
These test have been started recently.
Two constructions with metallic liner, both having the same basic concept, are available for the secondary loop. The tests on the behaviour of the hot gas duct in a horizontal position have been completed after 2900 operating hours in KVK. The results for the temperature distribution and heat losses are very positive. The test in a vertical position commenced in September 1983. At present work is centred on the constructive and analytical examination of the design of further subcomponents of the primary and secondary hot gas ducts, such as bend, compensator and T-piece.
Initially two variants were followed up for the hot gas valve, namely the ball valve and the axial valve. But since the end of 1981 we have only concentrated on the latter. As far as the constructive design is concerned, the test object has been prepared for the component test. The complete manufacturing documents have been submitted and the preliminary examination has been completed. Manufacture of the sub-component has almost been finished. It will be delivered and installed in the May of this year.
A scaled-down version of the axial valve, which has been adapted to suit the KVK, is part of the operating equipment of the KVK. Testing of this valve has already supplied valuable results.
Work performed in the past has shown that this component in particular has been the source of a number of difficulties which have caused delays in the development of the reference valve. However, they do not appear to have led to any delay in the overall component development as yet. -8-
5. Time schedule To conclude, let us consider the time schedule in the PNP project. The following facts can be deduced from the schedule (Fig. 5):
- With reference to the more closely examined components, one can say that their development will presumably be completed at the end of 1987, which is parallel to the material testing.
- As regards the development of steam and hydrogasification of coal it can be said that the results of hydrogasification development will be available according to the same time schedule. The results of the development of the steam gasi- fication process will be available only a few years later. Helium Raw Gas Coal
950 °C 40,5 bar He/He-Heat-Exchanger
Hot Gas Valve Process Steam Super Heater 815 °C 900 °C Helium 41,5 bar 41,9 bar
Cold Gas 789 °C Steam 300 °C 292 °C Valve 45,7 bar 41 bar 39,9 bar 200 °C 43,8 bar 682 °C 41 bar
Steam Generator
Primary Loop Secondary Loop
Row Sheet of the HTR with Steam Gasification of Hard Coal (1/84) Helium 950 °C ••• 50 bar
480 °C, 51 bar Cracked Gas
H2O/CH4 347 °C, 56 bar
Residue Steam Generator
300 °C Gas Separation 50,5 bar 292X949 bar
Methane
Blower Primary Loop Row Sheet of the HTR with Hydrogasif ication of Lignite (1/84) Helium V Tmax = 950 DC T °C 40 40 P =40 bar P bar 41.4 38.5 m = 3.8 kg/s m kg/s 0.619 1.234 Qel = 10 MW CH4 rel. Vol. 0.951 0.123 H2 " 0.039 0.681 CO - 0.096 C02 0.010 0.098
hpcisowasstT- nufbcroitung boiling water
EVA E Gas Heater Electric Reduction Valve
Condeneer Dtaeretor
feed Water Tank
FMd Wrter Preheater x PMd Wrtor Pump 1
-CK>- i Storage
ComprBMors Purification System
He-Filling T I •*«-
KVK Single Loop Operation With Test Sections Me-Prehe«ter Gas Heater Electric Heater
Deaerator
Feed Water lank
Feed water Preneater x Feed Water Pump
Storage Tanks Compressor* PurKteaWon System Heading y v 00
KVK Double Lxx>p Operation With He/He Heat Exchanger 90 91 92 93 Steam Gasification-Development - Operation of Small Pilot Plant - Documintatien - Bitic-Eng. of Pilot Plant - Decision for Next Step - Planning of the Pilot Plant - Erection of the Pilot Plant - Operation of the Pilot Plant ...,. Hydrogasification Development: Pilot Plant - Operation with lignite - Planning for Operation with Hard Coal - Decision for this Step - Rebuild of the Plant - Operation with Hard Coal Optimization and Evaluation of Both Gasification Procedures for Hard Coal Development of Material/Components - Material Tests - EVA: . NFE-Steim Reformer-Bundle . PNP-Steam Reformer-Bundle - KVK:. Hot Gas Ducts . He-He-Heat Exchangers . Hot Gas Valve - Research and Development Concept Review Coordination and Project Management Start of Planning for the Nuclear Process Plant with Steam Reformer and Hydrogasification with He/He-Heat Exchanger and Steam Gasification PNP Development of the Nuclear Process Heat (as per 1/84) 9* No. 2 111 _____ IAEA XA0055811 SPECIALISTS' MEETING ON HEAT EXCHANGING COMPONENTS OF GAS-COOLED REACTORS Dlisseldorf, FRG., 16. - 19. April 1984
DESIGN REQUIREMENTS ON HTR MAIN COMPONENTS FOR PROCESS HEAT APPLICATION Konrad Dumm INTERATOM GmbH., FRG.
Introduction
In the field of high temperature reactor application for direct process heat, KWU-INTERATOM developed the concept of the HTR-module. The module concept consists of a number of independent small pebble bed reactor units with a thermal output of 170 MW each. In direct process heat application each of the modules operates in combination either with a steam reformer for hydro-gasification of coal and methane reforming (fig. 1) or with an intermediate He/He- heat-exchanger for steam-gasification of coal. Reactor and heat exchanging components are connected by a coaxial gas duct, helium flow is maintained by an integrated circulator. The primary pressure inclusion is achieved by steel pressure vessels the technique of which is well proven in LWR operations.
The design requirements presented in this paper are based on the KWU-INTERATOM module concept, but to a large extend generally ty- pical for high temperature components. Besides the reactor itself the main components of interest are: a) in case of hydro-gasification and methane reforming (fig. 1)
- primary hot gas duct
- steam reformer with integrated steam generator b) in case of steam gasification (fig. 2)
- primary hot gas duct
- helium/helium intermediate heat exchanger (in 2 versions)
- secondary hot gas duct
- high temperature valve.
Some main operation and design data are given in table 1. It is assumed that a detained presentation of the components will be given by the manufacturers.
Brief description of components
The primary hot gas duct (fig. 3) serves as connection between reactor and heat exchanging component. Because of the coaxial design pressure induced loads are extremely low. But temperature loadings are comparably high. The nominal 9 50°C are superimposed by 18 K for temperature measurement errors and control deviations followed by - 20 K temperature field fluctuations from the core outlet. In addition 34 K for overtemperature induced reactor shut downs have to be taken into consideration. The poor accessibility to the primary hot gas duct however made the decision necessary to choose a design which guarantees a full reactor life time operation of 280.000 h. For this reason and because of the high temperature loadings a graphitic guide tube for the gas flow was selected. This tube is centered in a main carrier tube by means of Al«O^ adapting pieces. Thermal insulation is achieved by Al^O, multilayer mats. Differential thermal expansions and angular displacements are compensated by bellows in the main carrier tube. - 3 -
Under normal conditions the carrier tube is exposed to a pressure difference of only 1,5 bar from the outer cold gas side.
The steam reformer is shown in fig. 4. The 950°C helium enters the lower part and is distributed to the reformer tube bundle which consists of 199 individual reformer tubes of a tube in tube design. These reformer tubes are collected in the upper region by a tube plate which serves as a part of the primary;pressure inclusion. In order to allow a low temperature design of this primary barrier the lower side of the tube plate is thermally insulated. The helium leaves the reformer tube bundle at 720°C and is directed to 6 circura- ferentially arranged steam: generator modules. High temperature in- sulations around the reformer- tube bundle and the steam generator modules serve for a reliable separation of high and low temperature helium regions, by which an LWR like design of the.primary vessel becomes possible.
The process gas to be reformed, methan plus steam with a mole fraction of 1 : 4 is fed to the individual steam reformer tubes at a temperature of 34 7°C. By means of an individual recuperator in the upper part of each reformer tube the gas temperature, of the gas mixture is increased to 56 0°C. Flowing downwards through the catalizer filled reformer tube, the process gas becomes reformed. The maximum gas temperature at the lower end amounts to 810°C. Via an internal return tube the gas enters the recuperator again at 68 0°C and is cooled down to 4 80°C.
In the field of He/He intermediate heat exchangers two different designs are under development, a U-tube design and a helical coil design. Typical for the U-tube design (fig..5) is the arrangement of the secondary hot gas header in the upper region of the heat exchanger. Thereby the hot header is exposed mainly to the secondary helium temperature. The primary helium is directed to the tube bundle by an insulated internal vertical guide tube. Most of the thermal expansion of the tube bundle is compensated in the lower temperature field because of the special design of flow shroud and insulation. The tubes in the bundle region need to have only circumferential fixations but no tube weight supports. But never- theless high temperature resistant anti fretting coatings for - 4 - positions where relative movements have to be considered are necessary.
In the helix design (fig. 6) the secondary hot gas header is located in the lower region of the component. It is thermally in- sulated in order to avoid overtemperatures from the primary helium. A uniform flow distribution to the helix tube bundle is achieved by an integrated flow distributor. To a certain extend remaining temperature fluctuations from the core outlet are also reduced in this device. If a proper distribution of differential thermal expansions across the whole tube bundle can be realized under all operating conditions, relative movements between tubes and tube support structures can be minimised or eventually even avoided. But again high temperature resistant anti fretting coatings have to be taken into consideration.
The connection of both types of intermediate heat exchangers to the secondary helium system and the gasification plant is of identical design. Inside the primary confinement the secondary hot gas duct is, like the primary one, of the coaxial design. Outside the primary cell the coaxial design ends up with isolation valves, followed by normal single wall pipe work, designed to withstand the total internal pressure. While the cold leg valve will be of a normal and experienced design the hot leg valve (fig. 7) represents a completely new development. It is designed as a pneumatically controlled double acting valve. In case of a secondary system pressure relief accident, the valve closes automatically. From safety considerations the total closure time has to be < 15 s. Sealing is achieved by movement of the inner cones, which are positioned in gasstatic bearings, to the seat positions. Temperature sensitive parts in the inner bearing and control region as well as the outer pressure housing are insulated with a highly reliable high temperature insulation. The internals are cooled by helium. Design temperature for the pressure housing is 400°C. - 5 -
In the flow direction towards•the gasification plant;the high tempe- rature valve is followed by the single wall secondary hot gas duct (fig. 8). In contrast to the primary hot gas duct a metallic liner was selected. This for different reasons: The design temperature of 918°C is remarkably lower and allows a metallic system. The possibly more oxydizing secondary helium atmosphere does not allow the use of graphite or other carbon products. And of course the lower cost levels. The basic principle is well known from high temperature systems. The insulation technique is similar to that used for the primary hot gas duct. Axial gas flow in the insulation is avoided by the use of V-type spacers. The design of elbows, compensating units and T-branches is similar to the straight tube. Design pressure and temperature for the outer pressure shell are 46 bar and 400°C.
Design requirements and objectives
The general design requirements on these components can be sub- divided into 3 categories:
1. Requirements which have to be considered for quantitative analysis and design, mainly related to thermohydraulics and stress analyses.
2. Requirements for the detained engineering design taking into account results from 1 and 3.
3. A basic R+D programme initiated by 1 and 2 in order to solve open problems and to support the design.
All 3 aspects have to work in a steady and continuous interaction.
To point 1: Besides the normal operational conditions the following points have to be taken into account:
- design over pressures due to safety relief valves
- design over temperatures due to temperature measurement errors, control deviations, temperature field fluctuations, temp, increases due to over temperature induced reactor shut down
- temperature transients for normal start up and shut down and for emergency shut down (e.g. loss of heat sinks)
- pressure transients for normal start up and shut down and for emergency cases such as pressure relief accidents
- temperature induced loads resulting from hot stand by (reactor shut down followed by restart)
- thermal and mechanical effects in heat exchanging components due to failed and plugged tubes
- load effects by influences from outside like earthquakes, plane crash and gas cloud explosions.
These considerations have to be done very careful; on one side to fullfill all licencing requests on the other side to avoid any overconservatism. It is known in the high temperature stress ana- lyses that overestimations in temperatures can lead to much more severe design problems than those in pressures. This because of the very conservative results according to ASME Code Case N 47.
To point 2 In a sound engineering design a number of additional requirements are of great importance:
- due to repair concept or in case of part life time easy exchange- ability of the components
- according to regulations or customer requests easy accessibility to components and their internals for in service inspections preferably without opening the primary system
- possibilities to localize defects, esp. for repair and e.g. plugging of defect tubes
7 low stress levels, especially in high temperature regions
- load and weight supports preferably in low temperature regions - 7 -
- if in the development the construction and test of small prototyps is considered, this design has to be such that clear extrapola- tions of design and results to the larger components are possible
- suitable high temperature insulation systems which allow a strict separation of temperature and mechanical loads
- economic construction
- quality standards to meet the safety related requirements
- use of protective surface coatings to avoid fretting effects in helium
- use of materials with sufficient toughness being capable for inservice inspection techniques
- for the steam reformer: easy exchange of the catalyzer
- considerations of flow induced vibrations
- acceptable licensing procedures.
To point 3 As a whole the temperature range up to 1000°C plus the additionals to be taken into consideration including the safety requirements for a nuclear plant is a large step towards new fields. To accomplish the requirements stated before, a basic R+D programme is absolutely necessary. Reactor components of such new kind should not be designed straight on in one step on the drawing boards. A successfull development has to proceed carefully in a number of single steps:
- material development for high temp, metals and ceramics including graphit, carbon filament compositions, protective coatings, high temp, insulations etc.
- fabrication development e.g. narrow gap welding of thick-walled components such as the secondary hot gas header
- development and tests of quality assurance methods for fabrication and for inservice inspections - component parts development and tests
- development, fabrication and test of prototyps
- evaluation of prototyp test results.
These should lead to the final design of the reactor components,
Final remarks
It was the aim of this presented paper to summarize the design requirements on the main components for the HTR module concept for direct process heat application.
Papers from the different R+D organisations, manufacturers, operators of test facilities and pilot plants will give a more detailled picture to what extend the design goals have already reached. TABLE 1. OPERATIONAL DATA
ID CD IX o -52 a. LU UJ LU LU UJ o 1 LU u_ LU U_ PRESS l
PRESS l MAS S _J
PRIM. HOT GAS DUCT 50/40 1.5 950 300 50.3 2.8 * 50 720 50.3 1.4 STEAM REFORMER HE 1.5 950 STEAM REFORMER PROC.GAS. 56 6 347 480 34.8 1.4 STEAM GENERATOR HE 720 295 50.3 2.8 75 STEAM GENERATOR H2O 113 J 150 530 38 2.8
HE/HE IHX PRIM. 40 1 •ii-tt-Sf 950 295 50.3 1.4 2 HE/HE IHX SEC. 42 J 900 200 47.3 1.4 HIGH TEMP. VALVE 42 42 900 47.3 1.4 SEC. HOT GAS DUCT 42 42 900 47.3 1.4
DUE TO COAXIAL DESIGN MAX. PROC. GAS TEMP. 810°C FULL PRESSURE DIFFERENCE FOR PRESSURE RELIEF ACCIDENTS IN SECONDARY SYSTEMS HAVE TO BE CONSIDERED n
HiO * CH 34,8 kgfs 56 Dar 347°C
'51 bar
.!_i- -72CC
steam 113bar,530°C 38 kg/s
feed wafer 150°C
10 600
HTR Module 170MWfh with Steam Reformer and Integrated
Steam Generator Fig. 1 200° C 47,3 kg/s
\ 41,9 bar 900° C
HTR Module 170 MWH, with He/He JHX {U-tube design)
Fig. 2 I 2
to (U
Fig. SZX9LOL0 06000
process gas
•H20 • CH4 (4:1) 56 bar 347 °C inlet
51 bar 480 °C outlet
feed water 150 °C
50,3 kg/s, 50 bjr_ 950/300 °C
process gas connection feed water reformer tube tube plate pressure vessel process gas recuperator steam reformer shell life steam steam reformer tube bundle Helium to steam generators prim.cold gas duct steam generator tube bundle prim, hot gas distributor steam generator shell blower
170MWth Steam Reformer with Integrated Steam Generator Ficr. / , /
_200°C 47.3 kg Is
^ 41,9 bar 900°C
50,3 kg 1%, 40 bar 950 / 300 °C
0 sec. pipe connections (7) pressure vessel (2) support plate (8) heat exchanger shed (D sec. cold gas header ® flow shroud 0 sec. cold gas line (JO) prim, cold gas duct 0 sec. hot gas header @ prim, hot gas duct bundle 6t) blower
He/He intermediate heat exchanger, U-tube version of an HTR-Module
Fig. 5 50,3 kg/s, 1,0 bar 950/300 °C
sec. pipe connections pressure vessel support plate prim, cold gas duct sec. hot gas duct prim, hot gas flow distributor tube bundle sec. hot gas header heat exchanger shell blower flow shroud
Fig. 6 170MWfh Helix - He/He JHX T= 900 C P - 41,9 bar v«» 60 m/s
Double Acting High Temperature Valve 900/918"C op. /desg. 42/ 46 bar
o o o CM CM
Secondary Hot Gas Duct, oo (test prototyp) No. 3
HOCHTEMPERATUR-REAKTORBAU GMBH
XA0055812
PAPER PRESENTED AT THE SPECIALISTS' MEETING
ON HEAT EXCHANGING COMPONENTS OF GAS-COOLED
REACTORS
DLJSSELDORF, APRIL 16. - 19. 1984 HOCHTEMPERATUR-REAKTORBAU GMBH
Helium/Helium Heat Exchangers and Hot-Gas Ducts for the PNP-Project according to the BBC/HRB Concept
H. Schmitt, B. Jiirgens, J. Knaul Hochtemperatur-Reaktorbau GmbH, Mannheim
The prototype nuclear process heat plant PNP-1000 is designed for two coal gasification processes:
1. Steam gasification of hard coal and 2. Hydrogasification of lignite
In both plants the primary system is identical, i.e. that part of the plant containing the reactor and the components presented in this context. Fig. 1 shows the schematic flow diagram of the PNP-plant for steam gasification of hard coal. The primary system with the reactor, the heat exchangers and the circulator is shown on the left side.
In the heat exchangers the heat is transferred to an intermediate circuit, which is coupled to a gas-factory located outside the reactor building. At the outlet of the gas-factory SNG (substitute natural gas - CH.) is supplied for consumption. The steam required for the gasification process is extracted from the high-pressure section of the turbine.
A steam generator is installed in the intermediate circuit supplying the steam required for generating the plant electricity.
Coming back to the primary circuit, Fig. 2 shows a section of the reactor pressure vessel. It is a prestressed concrete reactor vessel in which the reactor core with the sperical fuel elements and the primary components are integrated. The arrangement of the components corresponds to the well-known THTR design. _ 2 - HQCHTEMPERAXUR-REAKTOREAU GMBH
At the periphery of the core the helium/helium heat exchangers are arranged in a tandem design. Each heat exchanger consists of a high-temperature and a low-temperature unit.
The connections between reactor core and heat exchangers are effected by the hot-gas ducts. The coupling with the intermediate circuit is indicated by the green coloured hot and cold lines of the intermediate circuit and by the connection line between low- und high-temperature unit of the heat exchanger.
The helium/helium heat exchangers
The helium/helium heat exchangers (Fig. 3) transfer the heat to the intermediate circuit, during normal operation a? well as during decay heat removal. For making optimum use of the space, the heat exchangers are designed in a tandem concept.
One objective of the current research program is to develop materials which ensure the operation of a heat exchangers for a reactor liefe time of 40 years. Based on the present state- of-the-art, this objective has not yet been fully achieved for those components which are exposed to very high temperature loads. This tandem design offers the advantage that the high-temperature unit of the heat exchanger can be removed in case that the design of the high temperature unit for the overall reactor life time is not possible. Separation between high temperature and low temperature units was selected at a temperature of 640°C so that the material NiCr 23 Co 12 Mo currently envisaged needs to be used only in the high temperature unit. The temperatures in the low temperature unit remain within the range of experience of the well known THTR steam generators. Both units are of approximately indentical dimensions.
The concept of the heat exchangers in helicoil design is presented in Fig. 3. The hot primary gas coming from the reactor core - 3 - HOCHTEMPERATUR-REAKTCRBAU GMBH
passes to the high-temperature heat exchanger, flows upward between the tube bundles and leaves the unit through windows located at the periphery of the inner shroud. The primary helium now cooled down to about 640°C flows through the annular clearance between inner and outer shroud and through a connection line well to the low-temperature unit. Flowing downwards along the tubes, the primary gas is cooled down to about 293°C is then collected in a plenum and is recirculated upwards through an annular clearance. The primary gas is then directed through a connection line and to the circulator located above.
After compressicn in the circulator the gas flows downward into the reactor cavity and is recirculated into the core through the external annular clearance of the hot-gas duct.
The "cold" secondary gas flows to the heating surface tubes passing through a cylindrical header in the lower part of the low-temperature heat exchanger. Then the gas flows upward within the tubes, heated up by the primary gas. The secondary helium is then forwarded into the high-temperature unit through a connection line. The heated gas is collected in the upper plenum; it is directed into the tube bundles through a tube plate and flows downwards. The hot secondary gas is collected in a cylindrical header and is then directed to the gas generator.
Of course it is also possible to use a U-bend design heat exchanger
The advantages of the heat exchanger in tandem design are (Fig. 4):
- Favourable utilization of space in a large single-cavity PCRV - Limitation of exchange to the high-temperature unit, if any exchange is necessary - The compensation required for thermal expansion of the hot lines is small. - 4 - HOCHTEMPERATUR-REAKTORBAU GMBH
- The loads resulting from different temperatures of the heating surface tubes and their fixing devices are low, - The dimensions cf both units are small, thus facilitating manufacture and assembly. - The number of parallel manufacturing steps is increased, thus permitting a reduction of the construction period.
The following" tasks will have to be performed in future:
- Special attention must be directed to the design of the secondary hot-gas header which is exposed to the primary hot helium. The primary hot-gas flow contains temperature streaks and velocity differences; these effects on the header have to be limited by structural measures. The influence of the header on the gas flow must be experimentally investigated. The routing of the helium lines must be designed so as to permit uncomplicated inservice inspections.
- The tube plate with a diameter of about 3 m must be designed to a temperature of about 600°C. In addition to the material strength problems, aspects of manufacturing must be taken into account.
- Seals in view of the possible exchange of the high-temperature unit.
The hot gas ducts The connection between the hot-gas plenum and the heat-exchanging components represents the hot gas duct (Fig. 5). The hot-gas duct is composed of the bearing wall and the thermal barrier applied to the inner surface. The hot gas flows through the duct at a speed of 60 m/sec.
The thermal barrier is provided to limit the heat exchange between hot arid cold gas and to protect the bearing wall of the hot- gas duct against inadmissabiy high temperatures.
The high requirements made for the hot-gas duct are mainly due to the high hot-gas temperatures. Based on the hot-gas temperature _ 5 _ HOCHTEMPERATUR-REAKTORBAU GMBH
the design temperature was specified at 10 50°C.
The hot-gas duct is designed as a coaxial line which means that the bearing wall is exposed to cold gas flowing along its outside. The material used for the thermal barrier applied to the inner wall is a ceramic fibre material containing 95 % of Al-O,,. These fibre blankets are covered by cover plates according to a design principle proven in the THTR. The thermal barrier is equipped with convection barriers in order to avoid longitudinal and transversal gas flow through the fibre blankets.
Fig, 6 shows the temperature profile in the undisturbed region of the thermal barrier.
The attachment fixture is composed of an inner stud and an outer sleeve (s. Fig. 7). Either element of the attachment fixture is able to accomodate 100 % load. They are elastically prestressed by plate springs which are cooled by the cold gas, which permits to balance different expansions. In the hypothetical event of rupture of an attachment fixture, fragments can not penetrate into the gas circuit. The fibre blankets are held down by the cover plates.
The design of a thermal barrier with cover-plates permitting free movement of each plate with the regard to expansion and shifts, is of great advantage especially when applied to complicated duct geometries.
The thermal barrier can be completely mounted in the hot-gas duct in the manufacturer's workshop. Thus the complete unit composed of hot-gas duct and thermal barrier is delivered to the construction site. This results in a considerable facilitation with regard to assembly. HOCHTEMPERATUR-REAKTORBAU GMBH
Initially the application of a metal material was analyzed for the attachment fixtures of the thermal barrier. Detailed strength analyses and verifications where performed for this material. In view of the application of a Carbon Fibre Composite - CFC - for the thermal barrier, this concept will be revised. The equivalent designs and calculations for the CFC attachment fixture are underway. Special attention will be directed to the structural and manufacturing aspects of attachment fixtures and cover-plates from CFC. The CFC is being further developed within a comprehensive program under special consideration of the components for which it is to be applied.
Based on elastic strength calculations, it has been verified for the attachment fixture of the hot-gas duct thermal barrier that the selected design is feasable and that it withstands the mechanical and thermal loads occurring during the operation of the PNP-plant.
The temperatures in the component parts are shown in Fig. 8.
The strength calculations where performed for mechanical and thermal load cases. The mechanical load cases include the event of a rapid depressurization specified as 2 bar/s. That represents the highest mechanical load.
Based on these load cases calculated within a finite-element program {temperature and depressurization) 5 load cases can be generated by superposition and linear conversion, which cover all operating and accident analyses. Fig. 9 shows the maximum stresses resulting from superposition for all 5 load cases. The calculated reference stresses are below the strength limiting value at 1000°C in all cases.
The analyses have shown that inadmissably high loads do not have to be expected. - 7 - HOCHTEMPERATUR-REAKTORBAU GMBH
We suggest the application of this concept because
1. it has a fundamentaly proven design principle
2. it is particularly well-suited for complicated geometries and
3. it can be used fur all HTR-projects currently known.
The present state of the development of the thermal barrier, the available test results, current tests and available test facilities permit the supply of this concept.
In conclusion it can be stated that draft solutions and concepts for the heat-exchanging components of a process heat plant are available. In part these solutions have been verified by calcu- lations and feasability analyses, in part also by tests. The right way towards supplying these components is being pursued. We have no doubt about their realization. It must, however, not be disregarded that comprehensive development work remains to be done both on the components and the materials. Make-up water
CH4
40 bar 3O0°C Low-temperature helium/helium heat exchanger 203°C
Intermediate Primary Circuit Cel!-typa wet Circuit Circulator cooling tower Circulator
!~?"»i»i:«wgt Decay heat removal heat -&J exchanger Decay heat Cooling vratef pump removal Circulator
Process Heat Plant for PUP HRB Steam Gasification of Coal 83.36-3
Abb. 1 Abb.P. Low temperature High temperature unit unit
Prim.-He Sec-He
PNP1000 He/He-Heat-Exchanger hRB 84^4-11
Abb.3 He/He-Heat Exchanger
Abb. 4 Header of the He/He-heat exchanger (high temperature unit)
s*—i 950°C 40 bar Primary helium inlet
Bearing wall
Thermal barrier
Secondary helium outlet
PNP1000 Hot Gas Duct hRB 84.24-10
Abb. 5 1000 -K L
1
800 - V >
fioo - \
400- \
\ ••Mi I C) 40 80 i;20 r(mm) FDerforated plate
1 > X 1X I Cov jlate Thermal barrier Bear ng tube
PNP Attachment Fixture for Hot Gas Duct Thermal Barrier PNP 1000 HRB Temperature in Thermal Barrier - Undisturbed Region 84.24-5
Abb. 6 Attachment Fixture for PNP1000 Hot Gas Duct Thermal Barrier in an PNIP Plant KRB 8424-2
Abb.7 Volume-weighted mean temperatures 1000 - Stud 5 - 629 °C - Sleeve 800 \ 600 *3""~ 400 v 40 80 120 z(mm) PNP Attachment Fixture for Hot Gas Duct Thermal Barrier PNP1000 hRB Axial Temperature of Structural-Elements 84.24-3 Abb.8 Max. Reference Stresses 2 2 379 20,9 11,0 28,5 11.8 3 46,8 25,7 38,0 74,3 56,4 / 4- 24,8 13,6 23,8 2,0 10,2 Load Case Description No. 1 p = 0,4 bar Depressurization accident 2 F=5500 N Operating prestress 3 Temperature load case 4 LC 1 + LC3 5 LC2 + LC3 PNP1000 Max. Stresses in Stud hRB 8424-9 Abb.9 No. 4 Recent Research and Development of Intermediate Heat Exchanger for VHTR Plant XA0055813 Akira Shimizu*, Noboru Matsumura* Hidetsugu Nishikawa*, Seiya Yamada* Abstract This paper describes recent tests which show progress in design of Intermediate Heat Exchanger (IHX) for high-temperature gas cooled reactor plants developed for process heat application utilizing nuclear thermal energy. As the IHX must have a large heat exchanging capacity, the most important consideration is to design a compact heat exchanger having high efficiency. For the improvement of heat transfer characteristics, tests were performed at Mitsubishi Heavy Industries such as, trial manufacturing, heat transfer characteristic tests and nondestructive inspection tests of finned tubes. Results are as follows; (1) Finned tubes produced by rolling were able to be wound helically by bending machine. (2) Laboratory tests showed this tube had good heat transfer characteristics. (3) There was not much difference between finned tubes and bare tubes for detection of defects by the eddy current tests. (4) A trial assembly using the same scale model showed that the present design is easy to assemble. (5) Automatic orbital welding can be adopted for the welding between, tube and tube sheet. As the result of the experiences gained from these successful tests, the design of the IHX has been greatly improved. 1. Introduction Design studies of the Experimental 50 MW (t) VHTR plant are being performed by the nuclear industries of Japan under the contract of the Japan Atomic Energy Research Institute (JAERI). The Intermediate Heat Exchanger (IHX) is one of the most important components, in which primary helium gas leaving the reactor at around 950°C heats the secondary helium gas for process heat application such as steel manufacture, hydrogen production, coal gasification and coal liquefaction. There are many R & D items necessary to improve the design of IHX, and these has been investigated these 10 years. In this paper tests such as trial manufacturing, heat transfer characteristics are discussed. * Nagasaki Shipyard & Engine Works Mitsubishi Heavy Industries, LTD. - 1 - 2. Description of VHTR plant The Experimental VHTR is a helium-cooled graphite-moderated reactor supplying a thermal output of 50 MW. The coolant tempera- ture at the outlet of the reactor is specified as 950°C. This type of reactor is designed so that nuclear thermal energy can be utilized not only for power generation but also for process heat application such as steel making, hydrogen manufacture coal gasification coal liquefaction. Coolant circuits consists of 2 loops (A-Loop & B-Loop) The A-loop is set up for heat removal. The system diagram of the A-loop of the VHTR plant is shown in Fig.l, and the general arrangement of the A-loop is shown in Fig.2. Primary coolant helium gas from the main gas circulator is heated in the reactor and leaves the Reactor at 950°C, and enters the Inter- mediate Heat Exchanger (IHX) at 940°C. Secondary coolant helium gas from the secondary helium gas circu- lator is heated up to the temperature of 905°C and flows to the Steam Generator where it changes water to steam. 3. The structure of the Intermediate Heat Exchanger The genaral arrangement of IHX designed by Mitsubishi Heavy Industries is shown in Fig.3, and a sectional view of the IHX is shown in Fig.4. The heat exchanger consists of a pressure vessel with concentric inlet and outlet nozzles at the bottom. This pressure vessel is called the outer shell, and it has an inner container which has an inner thermal insulation placed inside it. This inner container is called the inner shell, and it houses the helically coiled tube bundle around a center pipe with inner insula- tion. Between the inner shell and the outer shell is an annular space where low temperature cooling helium gas flows. The hot primary helium enters the pressure vessel at the bottom from the inner pipe of the concentric pipes. It flows upwards around the tube bundle and heats the secondary helium gas which flows inside the tube. The primary helium gas leaves the IHX near the top of the vessel and flows to a main helium gas circulator and there it is pressu- rized and flows into the annular space of the concentric duct near the top of the heat exchanger. It flows downwards through the annular space between inner shell and outer shell and leaves the exchanger through the annular space of the concentric pipes. Then it flows to the reactor vessel. Cold secondary helium gas enters the pressure vessel from the four inlet nozzles near the top of the vessel. It flows into the heat transfer tube through the low temperature manifold, and then flows downwards through the helically wound tubes. The secondary helium gas receives the heat from the primary helium gas and flows into the center pipe through the high temperature manifold. It flows upwards through the center pipe and leaves the IHX for the steam generator. - 2 - Items of Research and Development for IHX Fig.5 shows the design items with the related R & D for each part of the IHX. There are many items necessary to design the IHX. These have been investigated in Japanese laboratories and industries. In this paper the manufacturing of heat transfer tubes, trial assembly of the tube bundle and the high temperature manifold, heat transfer tests and the results of the eddy current test are discussed. Test results 5.1 Trial assembly 5.1.1 Purpose of trial assembly Objectives of manufacture are shown in Fig.6. This model was selected to test different manufacturing procedures affecting the tube bundle in which heat transfer takes place. This tube bundle in which heat transfer takes place is the most important part from the view point of heat transfer characteristics and manufacturing cost. In addition to the tube bundle a high temperature manifold was also selected. Test Items are as follows; 1 Development of finned tube 2 Bending of tube into a helix 3 Assembly of tubes and support plates 4 Welding between connecting tube and high temperature manifold 5.1.2 Results of trial assembly (1) Development of finned tube It is well known that the heat transfer characteristics of finned tube is more efficient than that of plain tube. But finned tubes have not been helically wound for the tube bundle of a heat exchanger for temperatures higher than 900°C. To improve the characteristics father, the angle at which the fin is attached to the tube has been increased. In other words the lead angle is usually small but that of the improved finned tube is almost 45 degrees. As gas flow is almost perpendicular to the axis of the tube, the fin causes turbu- lance in the gas flow and improves heat transfer. In order to make such a finned tube, cutting by lathe or rolling were considered as a manufacturing method. The cost of manufacture by lathe was expensive, therefore rolling was chosen. Though the material was Hasteiloy-XR which is harder than steel and a large lead angle was difficult to produce, satisfactory fins were made by rolling. Fig.7 shows this choice of this manufacturing method for the finned tube. Fig.17 shows the finned tube manufactured by lathe. The shape of the fin was very sharp but it was expensive to make this finned tube. Fig.18 shows the finned tube manufactured by rolling. The cost of manufacture is reasonable and high heat transfer characteristics were demonstrated as will be seen later. - 3 - (2) Bending tube into helical coil Finned tube made of Hastelloy-XR, and plain tubes made of Hastelloy-XR or of carbon steel were bent. Dimensions such as diameter satisfied the accuracy of design. (3) Assembly of tube bundle Trial assembly was performed for the new type of tube bundle. New type of support structure was confirmed to be easy to assemble and not injure the accuracy of coil diameter. Fig.8 shows the procedure of assembly of the IHX. At first finned tubes are produced from plain tubes by rolling. After that, the tubes are wound into a helical coil by a bender. Diameter and pitch are amended by hand bender to necessary tolerances. Cover rings are attached on the tubes for the protection from wear from the friction between tubes and support plates. Tubes are sandwiched by the 3 support plates at the cover rings. The inside support plates and outside support plates are welded to each other. Support plates are attached at 8 spots around the circumference. Next, the inner tube bundle is inserted into the outer tube bundle. After this operation, each tube layer is hung by tie- rods at the top. The procedure of the model is similar to actual IHX. Table 1 shows one of the advantages of this type of tube bundle is that the each tube can expand freely if its tempera- ture changes because it is not attached to the center pipe. Compared to the type which has support plates attached to a center pipe, thermal stress of this type is small. Another advantage of this type of tube bundle is that the tubes and cover rings are not scrached when being attached to the support plate. Table 2 shows the specifications of the assembly model and the actual IHX. And Fig.19 shows a total view of the model. (4) Welding between tube and tube sheet To attach the tube to the high temperature manifold, welding was adopted. Butt welding is used between the connecting tube and the manifold. The outer surface of manifold rises at each hole like a crater and it is butt welded at the top. As a result using Hastelloy-XR the weldability between two tubes and the tube sheet was very good. An X-ray test showed that there was no incomplete fusion and there was no defect. The model for the weldability test is shown in Fig.20. As many tubes are connected to the manifold in a small area, welding in a narrow space was tested using 5 tubes and a tube sheet. They were made of carbon steel. The pitch between the tubes was enough to adopt automatic welding called orbital welding. - 4 - 5.2 Heat transfer characteristic test 5.2.1 Test apparatus The heat transfer test for the rolled finned tubes was performed using the test apparatus shown in Fig.9. The apparatus used for this experiment consisted of a square duct, part of which housed a tube band, connected to a blower by a series of reducers and a round duct. Air was sucked by an induced draft fan from an inlet nozzle and was regulated by an intake damper. In heat transfer test section 150 (25 x 6) tubes were arranged in 40 mm longitudinal and transverse pitch, pitot tubes are installed at the inlet and outlet of test section to measure the flow rate and pressure loss of air. A turbine meter was installed at the inlet pipe for the measure- ment of the hot water flow rate. Thermocouples were installed to measure the inlet and outlet temperature of air and water. The velocity of air was 3 m/s ^ 50 m/s and the Reynolds number was between 5,000 and 100,000. 5.2.2 Test results The pressure loss and heat transfer coefficients at the tube inside were the same as for round plain tube because the roughness inside of the tube was very small. On the other hand the heat transfer coefficient outside of the tube was remarkably improved. Though the pressure drop outside of these finned tubes was similar to that of plain tubes, the heat transfer coefficient increased by 30% ^ 60% compared to that of plain tubes. Fig.10 shows the heat transfer coefficient and Fig.11 shows the pressure drop of finned tube outer surface. 5.3 Eddy current testing (ECT) For non-destructive testing of the heat transfer tubes of the IHX, eddy current testing (ECT) is most useful. ECT is widely used overseas and in Japan, too. A flaw detection test was performed. Some artificial flaws were prepared on plain tube and finned tubes. The result of the ECT shows that there was not much difference between finned tubes and plain tubes for the detection of defects. 6. Design progress As the result of the experiences gained from these successful tests the design of the IHX has been greatly improved as shown in Fig.12. An increase of the heat transfer characteristics by 30% resulted in decrease of total length by 34%. The specification of IHX is shown in Table 3. Further design work of the IHX is the stress analysis of high temperature structures such as the high temperature manifold, the reducer and lower connecting tubes. Fig.13 shows the abstract of these components. Example of a temperature and stress analysis is as follows. Fig.14 shows the manifold. The outside is covered by thermal insulation to decrease the thermal stress at normal operation (100% output). - 5 - A part of the manifold which has many holes is called the ligament. High thermal stress occures when the reactor increases or decreases its output because the change of helium gas temperature cause the ligament to change its temperature but other parts are slower to change. Fig.15 and Fig.16 show an example of this phenomenon. The temperature of the ligament is the same as that of the helium gas, but at spots away from the ligament the temperature is not so quickly changed by the temperature change of the helium gas. So a temperature gradient in the direction of the axis of the manifold causes thermal stress. The result of inellastic stress analysis showed that the manifold satisfies its life time. The analyses for other high temperature structures showed the structures are sound. 7. Conclusion As the result of the experiences gained from these successful tests, many data of characteristic, assembly and inspection were obtained. The main results are as follows; (1) Finned tubes produced by rolling were able to be wound helically by bending machine. (2) Laboratory tests showed this tube had good heat trasfer characteristics. (3) There was not much difference between finned tubes and bare tubes for detection of defects by the eddy current tests. (4) A trial assembly using the same scale model showed that the present design is easy to assemble. (5) Automatic orbital welding can be adopted for the welding between tube and tube sheet. By adding the data of corrosion, weldabil ity, creep, creep- fatigue and tribology has been obtained from other tests, the design of the IHX has been greatly improved. The Research and Development of IHX is in a final stage. REFERENCE - (1) JAERI, "Present Status of Research and Development for Multi- purpose VHTR" Annual report, 1980 % 1983. (2) JAERI, "Multi-purpose VHTR plant Detail Design (I)" Design Book. - 6 - Table 1 Comparison of structure Table 3 Specification of the IHX of support plates for VHTR plant 1 2 3 Type Helical coil, onethrough type ri No. of unit 1 1" Fluid Secondary helium I Flow rate 27.600kg/h A Tube side 1 Temp, ent./leav. 283/905 C 1 }\ —• Press. 41.3kg/cm'g —* , r g i Fluid Primary helium Flow rate 30.700kg/h Shell side Temp, ent./leav. 940/378' C Press. 40kg/cm*g Support plates Support plates Support plates Heat transfer 2.15x10'kcal/h[25MW| are attached to are not attached are divided into the center pipe to the centerpipe, individual layer Material HASTELLOY-XR they are hung and are hung Size 31.8mmO.D.(Fin outside) X 4.0mm1 by tie rods by tie rods Tube Explanatio n Number 296 Pitch 40mmfl trans. )x40mm|long.) No. of layer 13 Helical coil Dia. max./min. 1.760/800mm stres s Therma l X o o Material 2 1/4Cr. -IMo Steel Shell Outside dia. 2.380mm Height 14.300mm assembl y X X o tub e bundl Suitabilit y o f This specification shows the IHX for A loop Table 2 Specifications of model for trial assembly Specifications Items Trial assembly model Actual IHX |25MWt| Height(mm) About 3.000 About 7.500 No. of layers 3 13 Diameter min. max. min. max. of layers(mm) 800 1.000 800 1.760 Lead of coil 600 750 560 1.240 (mm) No. of tubes 26 296 --..-111 Lead angle of About 13 About 13 helical coil Helica l coi tub e bundl Trans, pitch 50 (mm) 40 Longi. pitch (mm) 50 40 Outer diameter (mm) 31.8 31.8 Thickness 4 4 (mm) Tub e Hastellov XR. Material Hastel oy XR Carbon Steel Fig.l Schematic flow diagram Length of About 14.000 About 33.400 one tube(mm) - 7 - container vesse Inner concreat Secondary He gas circulator Fig. 2 Outline of general arrangement of the heat removal loop o ...p. C«..M H.l H-. 1 J P... >, ! i^ 1.1.1 1.-0 i 111 *c Coll., T—,, )H 'C i ..1 •c Secondary helium gas 16 ) T/h' ! 11. i outlet D,_..,^. ...„ (to Steam generator) @ Primary hellum gas Tu.. S,,. .,, . . outlet (to Crrculatorl I © Primary helium ga (from Circulator) hol.um gas outlet]! (T)Pnmnty helium 93s inlet ho HencloO Itrom Raaclof) Fig. 3 General arrangement of IHX Fig. 4 Sectional view of IHX Process of design Assembly of h*Uc*J coil tub* bundt* Tub* (finrMd tub*, pUin tub*) Support pUt* W«kiiog b*tw«wi manifokl •nd conn*cting tub* A Automatic wtldmg t**t B W*4ding t*st of HuttNoy XR Mat«rial : Hutattoy XR Fig. 5 Research and Development items Fig. 6 Components for IHX for trial assembly Eiiinm el b«ndmg tub«« into • h#*C*f H**1 trMMfM CCHI ch»r»ctMiatics cott a Fig. 7 Choosing design for finned tube ••-1. Fig. 8 Assembly of the tube bundle - 9 - V 10 G-Weight v^octty N : No. of tub* row Dtnsity f : Friction factor Round tub« bundb Pr«ssur« loss t*st Hwt transfer twt 5x10* 10* 5X104 10* BxlO1 Reynolds numb«r (R«) Fig. 9 Heat transfer test apparatus Fig, 11 Friction factor of finned tube outer surface onventions typ* N«w typ* Pl*in tu6« typ« Fk>n*d tub* typ« FINNED TUBE BUNDLE Nu = 0.139R«°'"Pr" PLAIN TUBE BUNDLE r = 0.7 =0.95 10 10 10- 10' Re=.Gdo Fig. 10 Heat transfer coefficient at tube outside —~~~—___Conv*ntion»t N*w tVM Cut town Tot«J Uo«tNml 21.Cm 14.3m lt*H diam^wlml 2.ftm 1% ^•l«ht of iwb« bundMm] 12.8m 7.Un 42% Fig. 12 Comparison between Conventional and new design of IHX - 10 - Fig. 15 Temperature change in shut down Fig. 13 Abstract of high temperature structure T*mp*ra1ur* Location for stress evaluation Tim«© 69m*n. Fig. 16 Temperature Thermal insulation and stress distribution after manual trip Fig. 14 Analysis model for high temperature manifold - "11 - Fig. 17 Finned tube manufactured by lathe (SUS) Fig. 18 Rolled finned tube (Hastelloy-XR) Fig. 20 Partial model of high temperature manifold (Hastelloy-XR) Fig. 19 View of assembled model No. 5 XA0055814 IAEA Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors Development of a Helium/Helium Intermediate Heat Exchanger with Helical Coil Tube Bundle by A. Czimczik (L. & C. Steinmiiller GmbH, FR Germany) G. Hirschle (Gebr. Sulzer AG, Switzerland) Dusseldorf, 16. April 1984 — 1 — Development of a Helium/Helium Intermediate Heat Exchanger (He/He-IHX) with Helical Coil Tube Bundle 1. Preface Besides the research and development work leading to the design basis for the reference heat exchanger, the stages of development up to the realisation of a nuclear He/He IHX include, selected experiments in manufacturing techniques and with individual test components, as well as the design, manufacture and trial of a representative prototype compo- nent. The general features of the helical construction will first of all be explained, taking into account the manifold spe- cific requirements placed upon a nuclear component. The primary and secondary flow path will be discribed. Special design measures included to fulfil specific requirements imposed by quality assurance and the need to facilitate re- peated examination or ease of repair are also explained. A comparison between reference and test component shows that the capacity of 10 MW represents a sensible compromise between a unit capable of providing answers to the techni- cal problems involved and the cost of this step forward. The selected size provides a useful test unit for material processing, manufacture of components and quality control. The materials used and the new design problems meant that in constructing the 10 MW test component certain manufac- turing processes had to be modified or to be developped to- tally new. The successful use of these methods in the manu- facture of the test component marks an important step in the direction of producing such full-scale components. The process of manufacture and the experience gained so far are then described. - 2 - 2. Technical requirements The more important requirements demanded of the heat ex- changer, which forms part of the primary enclosure, can be considered under three aspects. These are: - Safety requirements - Limiting process conditions - Mechanical design aspects Among other things, considerations of safety motivated the following decisions: - The heat exchanger should be located within the pressure vessel. This dictates a compact construction and arrange- ment of all assemblies. - All surfaces enclosing primary circuits, particularly weld joints, must be accessible for repeated examination without having to open the primary circuit. - In the event of component failure there must be no risk of primary gas infiltrating into the secondary circuit. Only one of the limiting conditions imposed by the process will be mentioned: The primary helium hinders the formation of oxide layers. It is therefore possible, that in the higher temperature range parts which are in contact might weld together by diffusion. The contact areas of any parts which move rela- tive by to one another must therefore be protected against undesired welding or wear by suitable coatings. For reasons of mechanical design the loads on components, especially those exposed to high temperatures, must be minimized. In the given temperature range of 950 °C, even with the use of high temperature resistant alloys, the available design reserves are minimal, particularly looking to the life of 140,000 hrs specified. - 3 - 3. Main development steps Although process equipment has been successfully operated at similar temperatures in many industrial applications, the design, manufacturing and operational experience avail- able from such conventional components is not by itself an adequate foundation for attaining the development objective. This applies particularly to the problem of approval of components for use in the nuclear field. For one thing, work on the relevant codes to be applied in the high temperature field is not yet complete; for another, the long-term characteristics of the reference material Inco- nell 617 or Nicrofer 5520 Co (2.4663) have not so far been ascertained. It is therefore necessary to find new ways of proving the safety of this type of equipment. The construction and testing of a component which shall be as representative as possible is regarded as the most im- portant step in this direction. This component, with a heat transfer capacity of 10 MW, is at present under construc- tion. Before this, and in association with it, manufacturing tests were carried out on the machining and welding of the material 2.4663. Special experimental components were tried out to settle particular questions regarding thermohydrau- lics and the mechanical behaviour of important parts (hot headers, coatings, tube sleeves, test bundles for thermal transfer and flow measurements). Additionally, elastic and inelastic analysis was undertaken for the assessment of all critical components. The soundest basis for ensuring that a product will pass all acceptance requirements is however a well thought-out design, which from the very beginning reduces the load on critical areas to a minimum. Here the helical variant of the intermediate heat exchanger offers excellent possibili- ties, since because of its specific constructional features it provides good compensation for thermal expansion and can rapidly disperse hot gas stratifications in the tube bundle. - 4 - 4. Description of the component 4.1 Structure and flow paths The essential component parts of the He/He IHX and the gas flow paths can be seen from the general drawing of the 10 MW apparatus (Fig. 1). From the covering support closure which supports the whole heat exchanger, the central duct for the returning hot se- condary helium extends downwards. This hot gas return duct (HGRD) is welded to the support closure via a thermal sleeve. On the HGRD ten cylinders forming the tube bundle support are attached one after the other. Each cylinder bears perforated support plates arranged radially around the circumference in order to locate the tubes. The seg- mental arrangement of the supporting elements allows for the relative axial expansion as a result of different oper- ating temperatures in an optimal manner. The helices surround the hot gas return duct in cylindric assemblies. At both ends connecting tubes link up the tube bundle with the support closure and the hot header respec- tively. The secondary helium is distributed evenly to the helices by an annular header. It flows in counterflow with the primary helium and is heated thereby from 220 °C to 900 °C. Subsequently it flows back through the internally insulated return duct. The primary helium enters the heat exchanger at a tempera- ture of 950 °Cr at the bottom via a mixing and baffle cham- ber and flows upwards through the tube bundle. The inside boundary for the gas flow is the supporting cylinder. The outside boundary is given by a flow shroud welded to the support closure. The helium cools down to about 300 °C, turns below the support closure and flows back through the outer annulus between pressure vessel and external flow guide shell to the outlet. The external flow guide shell also serves for the attachment of the insulation. _ c _ 4.2 Technical data and their transferablility In designing this test component, strong emphasis was laid on the possibility of applying the experience gained in manufacture and test running as far as possible to the re- ference heat exchanger. Pressures, temperatures, tube dimensions, tube pitching, height of the bundle and helix pitch are therefore practically identical for both units. The only major difference is the number of tubes (117 as against 1482) , which are arranged in 3 and 19 cylindric tube assemblies respectively. The test bundle is therefore very largely typical in respect of manufacture, assembly, selection of material and design details. The flow and heat transfer characteristics however are strongly influenced by the number of tube assemblies i.e. where only a few cylin- ders are involved the boundary bypass gets a disproportio- nately large influence. For this reason supplementary component tests and basic thermohydraulic research are being carried out to prove the design of this heat exchanger. The tube bundle built for this purpose is shown in Fig. 2. The cross-wise arrangement of the tubes can be clearly seen, as well as the support plates for the tubes. This bundle is at present undergoing trials at the KFA/Jiilich. f&O - 6 - 4.3 Design Features The central part of the test unit the hot gas return duct, approximately 20 m long, is constructed in 2.4663 (plus sections in the cold area of 10 CrMo 910) with a diameter of approx. 762 mm (Fig. 3). The wall thickness depends on the local operational temperature and varies between 35 - 60 mm. The individual duct build up from sections which are welded together by narrow-gap welding. The hot gas return duct is internally insulated. The insu- lating system consists of a continuous outer tube in which shaped sections of compressed Al^O- fibre are inserted (Fig. 4). Internally, the insulation is covered by indivi- dual overlapping liner tubes. The internal rings between these two concentric tubes support the weight of the insu- lation and also hinder convection through the insulation. The selected mode of construction permits the complete in- sulation to be withdrawn and allows periodic inspections of the hot gas return duct from inside. The insulation of the outer shell is built up in a similar manner. It was however not necessary to arrange for complete cladding of the insulation, since it is only surrounded by a stagnant helium layer. The supporting structure of the helices (Fig. 5) consists, as already has been said, of separate elements which are slid onto the hot gas return duct one after another and each welded into position at one end. The points at which the cylinders overlap and move relatively to one another are suitably coated. The helices (Fig. 6) have an inclination angular of approx. 25 ° and are located in the supporting plates with the aid of sleeves. These tubular sleeves must fulfil three re- quirements, viz: - 7 - - they must compensate for installation tolerances - they must take the weight of the tubes - they must function as a protection against wear. Pairs of sleeves wedged onto the tubes are envisaged for the temperature range below 700 °C. They however cannot be used for temperatures above 700 °C, as frictional-type con- nections become ineffective as a result of the relaxation of material and subsequent loosening. Various proposals have been developed for this temperature range, of which two are illustrated here by way of example (Fig. 7). To accept the downward forces imposed by the weight of the tubes, the tube itself must exhibit a thickening such as a welded-on ring or a built-up weld, as well as a suitable thrust piece on the supporting plate. As a protection against high-temperature fretting wear and to prevent spontaneous welding in the helium atmosphere, ceramic coatings are applied. The current view is that layers of zirconium oxide stabilized with yttrium are best suited to use under PNP conditions. However proof of the reliability of such protective layers will only be possible as a result of the projected component tests and especially test running with the 10 MW unit. For this reason the intention is to try out various alternative types of tube holder and coating in the test unit. The heating surface tubes lead into the hot header (Fig. 8). Because of its exposed position in the hottest part of the gas temperature range, this is regarded as being one of the most heavily-loaded components. A number of design measures were therefore taken to limit the severity of the load on this component as much as possible. First of all, the header and the incoming heat exchanger tubes are protected against direct impingement by the primary helium, which might well contain hot gas stratifi- cations. As the secondary helium has a temperature of 900 °C and is therefore some 50 °C cooler than the primary helium, an additional internal flow guide tube is fitted to improve the internal heat transfer in order to keep the temperature as close as possible to that of the colder helium. A third possibility of lessening or preventing impermis- sible thermal stresses in thickwalled structures as a re- sult of steep temperature gradients is by careful insula- tion of endangered areas. - 9 - 5. Fabrication and quality control 5.1 Welding 5.1.1 Gas-shielded metal arc narrow gap welding For welding the pipe sections of the hot gas return duct (HGRD) together, a process known as "gas-shielded metal arc narrow gap welding" was utilised. This process has the following special features: 1. A. specially developed wire-weaving mechanism produces a weaving arc in a narrow gap. 2. The weld puddle is controlled by the shield gas flow to produce an optimum contact angle between the weld sur- face and the flanks of the weld. 3. The use of a pulsed arc power source produces a stable arc and a weld edge with minimum spatter. The process will be more fully dealt with in another paper. Welding equipment from Interatom was used in the fabri- cation. - 10 - 5.1.2 Other welding work Tube bundle supporting system Together with the "Internal Flow Shroud" this assembly constitutes the major welded-up components which determine the geometry of the active part of the heat exchanger. Exact conformity with the prescribed geometry of the flow duct when locating the helical tube bundle is therefore important and for this reason the welding and machining of this parts is a complicated business. The permissible radial tolerance between the support plate edges and the internal flow shroud on installation requires special jigs for welding up the segments of the supporting structure, consisting of support cylinder, weld-on ring and support plates. The weld preparation for the connection between the support plate and support cylinder is shown in Fig. 9a and that for the connection between the support cylinder and HGRD in Fig. 9b. For these welds the hand-manipulated electrode Gritherm 85 R proved reliable in service. Testing according to the specifications has been demonstrated for all weld connec- tions whose use is indicated for the full-scale PNP compo- nent and has been proven by component tests. Tube weld connections The weld seams in the heat exchanger tubes (similar-metal welds in 2.4663 and dissimilar-metal welds 10 CrMo 910/ 2.4663) are carried out using the TIG process with station- ary torch in the downhand position and rotating workpiece. This process was easy to apply and produced good results. All welds were radiographed and subjected to surface crack detection procedures to check their soundness. - 11 - Homogenous welding of tubes to the support closure The 117 tubes of 10 CrMo 910 material were welded from in- side, using a hand welding gun, to the support closure which is also of 10 CrMo 910. As preparation for welding a fine centring ring was machined into the support closure at the downside Wherein the tube was located. (Fig. 9c). Special arrangements had to be made to ensure optimal gas shielding. The welds are made with the tube axes in the vertical plane, to produce the best possible bead confor- mation. 5.2 Machining The more important machining operations on the helical type intermediate heat exchanger are: - Deep-hole drilling of the holes and countersinking of the ring grooves for connection of piping to the support closure. This work in 10 CrMo 910 material was undertaken by a specialist firm and presented no problems. Drilling of the holes and countersinking of the ring grooves on the hot header (Pig. 8). The toughness of the 2.4663 material blunted the tools, which imposed the necessity of frequent tool changes. In comparison to "normal" steels the cutting speed also had to be kept very low. Because the area to be machined was relatively small, it was possible to carry the work out with normal machine tools. For full-size components it must be attempted to develop special tools of a suitable type to facilitate these operations. - 12 - - Machining of supporting cylinder with support plate. The prescribed narrow tolerances meant that special jigs and fixtures had to be used in machining the supporting structure segments as well. All surfaces of the segments which immediately contact other components were speci- ally machined. The drilling of the support plates was effected using drilling templates in a specially fitted holding and chucking device mounted on a horizontal dril- ling machine with numerically-controlled coordinate input (for a finished segment after machining, see Fig. 5). - Final machining of the support closure/HGRD assembly. The great length of this component {about 21 m) and the required positional tolerances in respect of linearity and angularity meant that machining on an oversize lathe was necessary (Fig. 3). Machining presented no problems and the tolerances could easily be complied with. The actual tolerance measured amounted only to some 10 % of that permitted. 5.3 Protective coating For the experimental helical type heat exchanger one HGRD section and one set of connecting tubes, both of 2.4663, were coated with a triple layer of ZrO2 13 % Y-O- NiCr A1Y, using the APS-process. The choice of this type of coating represents the best available knowledge and was made as a result of the research whicl- is still continuing on the problem. - 13 - 5.4 Bending of the helix tubes Up to now the 3-roll bending process had only been used for shallow-pitch tube helices. For the He/He Heat Exchanger the process has now been expanded in experimental proce- dures to cope also with steeply pitched helices. Besides the modifications of the machines, this meant that the new bending parameters had to be found out. The permissible manufacturing tolerances were worked out by screwing tests on individual helices. The tubes of the KFA bundle for thermohydraulic tests were the first production application of this process. For the 10 MW He/He IHX the tubes were welded together from 4 standard lengths to a total length of around 45 m and then bent to form helices of about 18 m length and three diffe- rent diameters of 918, 990 and 1062 mm. The intermediate diameter helices were given a left-handed helical coil, the others a right-handed coil. A particular feature in the production of steep-pitch helices is the great amount of room required for bending and intermediate storage (Fig.6). 6. Conclusion The development of a large He/He-IHX (125 MW) for the pro- totype plant for nuclear process heat (PNP) requires not only basic engineering but also various experiments on special structural parts of the component in order to test manufacture procedures and important details of the design. Additionally to this, a representative experimental IHX was designed, which is now in fabrication and which will be tested early in 1985. Until now already many requirements connected with IHX are fulfilled. Therefore we are very confident that after the test runs a large nuclear heat exchanger can be realized. - 14 - 7. List of Figures Fig. 1 10 MW test unit Fig. 2 Helical test bundle Fig. 3 Hot gas return duct Fig. 4 Insulating system Fig. 5 One element of the supporting structure Fig. 6 Helical tubes Fig. 7 Tube locating elements (sleeves) Fig. 8 Hot gas header Fig. 9 Weld preparations a) Weld preparation for the connection between support plate/support cylinder b) Weld preparation for the connection between support cylinder/HGRD c) Weld preparation for welding heat exchanger tubes to support closure 900TV _ Secondary helium -t- 220 °C Ring distributor Support closure Hot gas return duct Insulation Operational Data Primary Secondary Flow rate 2,95 kg/s 2,85 kg/s Temperature 950/293 °C 900/220 °C Pressure 39,9 bar 41,9 bar Diff.Pressure 0,55 bar 1,65 bar Power 10 MW Dimensions and Material Number of tubes 117 Dim. of tubes 22 X 2,0 Tube Material 2.4663 (Nicrofer 5520) Vessel Material 1.6368 (WB 36) Structure Material 1.7380 (10 CrMo 910) 1.4876 (Incoloy 800 H) 2.4663 (Inconel 617) 950 °C Gas mixing device ITEIMMUUER Fig. 1: 10 MW He/He Heat Exchanger Fig. 2: Helical test bundle Fig. 3: Hot gas return duct Fig. 4: Insulating system o o oi o o o o O( o o o< o v-~ •* V Fig. 5: One element of the supporting structure Fig. 6: Helical tubes Fig. 8: Hot gas; header 300-700 t 700-900 °C 700-900 C Fig. 7: Tube locating elements (sleeves) support cylinder support plate support cylinder HSRD gas • tube Fig. 9: Weld preparations No. 6 XA0055815 IMPROVED SPACERS FOR HIGH TEMPERAUTRE GAS-COOLED HEAT EXCHANGERS L,A. Nordstrom Swiss Federal Institute for Reactor Research 5303 Wurenlingen / Switzerland Paper presented at Specialists's Meeting on Heat Exchanging Components of Gas-Cooled Reactors Diisseldorf, BRD April 16-19, 1984 -. 1 - IMPROVED SPACERS FOR HIGH TEMPERATURE GAS-COOLED HEAT EXCHANGERS by L.A, Nordstrom / Swiss Federal Institute for Reactor Research 1. INTRODUCTION Experimental and analytical investigations in the field of heat exchanger thermohydraulics have been performed at EIR for many years. Basic studies have been carried out on heat transfer and pressure loss for tube bundles of different geometries and tube surfaces, As a part of this overall R+D programme for heat exchangers, investigations have been carried out on spacer pressure loss in bundles with longitudinal flow. An analytical spacer pressure loss model was developed which could handle different types of subchannel within the bundle. The model has been evaluated against experiments, using about 25 spacers of widely differing geometries. In a gas-cooled reactor it is important to keep the pressure loss over the primary circuit heat exchangers to a minimum. In exchangers with grid spacers these contribute a significant proportion of the overall bundle losses. For example, in the HHT Recuperator, with a shell-side pressure loss of 3.5 % of the inlet pressure, the spacers cause about one half of this loss. Reducing the loss to, say, 2.5 % results in an overall increase in plant efficiency by more than 1 % - a significant improvement. Preliminary analysis identified 5 geometries in particular which were chosen for experimental evaluation as part of a joint project with the SULZER Com- pany, to develop a low pressure-loss spacer for HHT heat exchangers (longi- tudinal counter-flow He/He and He/H20 designs). The aim of the tests was to verify the low pressure-loss characteristics of these spacer grid types, as well as the quality of the results calculated by the computer code analytical model. The experimental and analytical results are compared in this report. 2. EXPERIMENTAL APPARATUS An open air loop at EIR was used for the tests. The required air flow was obtained from a rotary compressor. It gave a maximum air flow of 1 kg/s at a pressure of 0,5 bar (above atmospheric). A remotely - controlled motorised valve in a bypass pipeline was used to regulate the air flow, and the excess air released to atmosphere. A cooler was used to keep the air temperature at a constant level. A silencer was added to the compressor in order to suppress undesired resonant vibrations. Pressure tapping (18 mm rod pitch) or blocked tapping (24 mm rod pitch) Bundle tubes 0 12 x 1 mm Spacer grid Blocked tapping (18 mm rod pitch) or pressure tapping N> (24 mm rod pitch) Wall half-rods Attachment for screwed to the the grid test-section wall Fig- 1: Cross-section of the test module (18 mm rod pitch) Fig- 2: Test-section in the laboratory with connected pressure tappings — 3 — The test-section allowed two bundle arrangements to be installed, with a tube-pitch of 18 or 24 mm (i.e. a pitch-to-^diameter ratio of 1,50 and 2,00, respectively, for a tube of 12 mm O.D,). Half-rods were fixed to the test- section walls to reduce the influence of the wall friction as much as possi- ble. The corner rods were an integral part of the test-section wall. The two bundles consisted of 37 full- and 18 half-tubes (Fig, 1), and 19 full- and 12 half-tubes, respectively. The test-section itself was a precision fabricated piece of 2 meters length (Pig. 2) in two separable halves. The tubes for the bundle and the wall half- rods were also of high precision construction with O.D. of 12,01 t 0.01 mm. The test-section wall and the tubes can be considered as smooth. The bundle tubes had solid ends which were rounded at the inlet end. The wall rods were tapered towards the inlet end (Fig. 4). The longitudinal attachment of the tube bundle to the test-section shell was carried out at a special grid (Level 6, Fig, 3), holding the tubes at their ends in such a way as to create no extra pressure loss in excess of that of a normal spacer grid. As ct "flow equaliser" (Level 1) a spacer grid of the same type as the one to be investigated was normally installed. The spacer grids could be mounted at any level between 2 and 5, allowing one alone or more together to be tested in the same run. 50 pressure taps, 19x 100 mm 50 j I I I I 1 I I I I I I I I I I Inlet X Outlet level of spacer grids: 1. flow equaliser spacer used as longi- possible positions tudinal fixing point for the test grids for the tube-bundle Figure 3: Axial arrangement of spacer grids and pressure tappings in the test-section. - 4 - 3. MEASUREMENT AND DATA HANDLING To determine the spacer grid pressure-loss coefficient, Z, as a function of the Reynolds Number, Re, the air mass flow was changed and other parameters kept constant (e>g, inlet air temperature). The laminar bundle flow cases are not of interest here and therefore not considered. The range of Re co- vered here is from 10^ up to the maximum achievable for each geometry, i.e. about 105 (p/d = 1.5) and 2 • 1O5 (p/d = 2.0). In order to have good mea- suring conditions, the Mach-number was kept below a value of 0.3, At each of the 19 measuring levels (Fig. 3), all three pressure tappings were connected together to obtain an average value of pressure. From experience with other experiments, this arrangement results in a significantly more accurate value than the use of only one pressure tapping at each measuring level. Obtaining a mean value from many single-point measurements by the use of a rapid-collection data acquisition system, has reduced the scatter signi- ficantly, to about 20 % of the scatter of the single-point values. For the data analysis and calculation of the spacer pressure-loss coefficients, the EIR computer code "TESERA 2" was used. It plots the pressure data and the resulting Z versus Re. The gradient of the pressure loss between two grids was calculated by the method of least squares, without fixing any measured point as a reference. Pressures measured in the disturbed region just downstream of a grid were neglected when extrapolating inter-spacer pressure profile. This eliminated error caused by high turbulence in the near-spacer region, and the loss of a few data points did not significantly affect the accuracy of the fitting of the line to the data. 4, SPACER GRID TYPES In cooperation with Sulzer Bros., 11 new spacer grid types were constructed as possible low pressure-loss grids suitable for the HHT heat exchangers. Following calculation with the computer model (see Section 6), 5 types were chosen for the experimental investigation reported here, see Fig. 5 to 9. Noted for each type is the ratio of pitch-to-tube O.D. (p/d), the total height (h), the solidity (O = spacer blocked area in the bundle channel / free flow area in the channel) and the resulting typical spacer pressure-drop coefficients (Z, see Section 5). Defining the blocked area of a spacer grid is not always straight forward, and depends upon the geometry. Of the grids presented here, the calculation is trivial for Types B and D, with constant geometry throughout. For Type C, with axially overlapping plates, the blockage was defined at the level with the smallest free flow area. In Type A the blocked area does not vary along the spacer, the tube-locating dimple just causing transverse flow inside the spa- cer. However this effect does produce a pressure drop and, after investigation, it was found that approximately 3/4 of the projected area of the Type A dimples should be added to the spacer plate blocked area. Supporting this, similar spa- cer geometries have been treated in a corresponding way, and good agreement with experiment has been obtained. Fig. 4; View of the test-section Fig. 5: Type A, p/d=1.50, h=38 mm inlet. Bundle-rods with rounded a - 0.291, Z = 0.76+0.1 ends. Jv*» ^.4" ^ 7 4 Fig. 6: Type B, p/d=1.50, h=50 mm Fig. 7: Type C, p/d=1.50, h=38 mm 0 = 0.285, Z = 0.5+0,05 a = 0.191, Z = 0,6+0.05 Fig. 8; Type B, p/d=2.00, h=38 mm Fig, 9: Type D, p/d=2.00, h=38 mm a = 0.171, Z = 0.21+0.02 a = 0,180, Z = 0.26+0.03 - 6 T 5. EXPERIMENTAL RESULTS From the measured pressure characteristic on each side of the test spacer, the axial pressure drop gradient for the undisturbed bundle was evaluated. Axial extrapolation to the edge of the spacer gives the total pressure drop over the length of the grid, Different spacer grid pressure-loss coefficients are compared by considering just the extra pressure loss which a grid adds to a free bundle (i.e. without the influence of tube wall friction). This spacer grid loss coefficient can be defined as: 2 2 Z = Ap • 2p • A /m - A h/d. h where Ap = pressure drop over the spacer grid length (N/n/ p = density (kg/m ) A = total free-flow area in the bundle (m2) in = mass flow rate (kg/s) A = tube wall friction factor (-) h = axial length of spacer grid (m) d = hydraulic diameter (m) The experimental results of all the grid types investigated are given in Fig. 10 in the form of Z versus Re, After consideration of all inaccuracies (in test-section geometry, measuring equipment, mathematical fit etc.), the accuracy of the evaluated pressure-drop coefficients is within the range ± 10 %. This was only achieved by having very small tolerances on all geome- trical dimensions. o 5-10 o Q. 2 10 w -i a> 210 5'10' 10a CL Re Spacer grid A Type A, P/d = 1.5 A Type D, P/d =2.0 • Type B, P/d = 1.5 O Type B, p/d = 2.0 • Type C, p/d = 1.5 Figure 10: Experimetal spacer grid pressure-loss coefficient (Z) versus bundle Reynolds number (Re). _ 1 — 6, ANALYTICAL MODEL AND CALCULATED RESULTS The EIR analytical model for the prediction of spacer pressure-^loss coeffi- cients has been evaluated and improved over a number of years. To date it has been tested against experimental results from more than 25 spacer grids of widely differing geometries. The model can handle different types of subchan- nel within the bundle, such as triangular and square arrays, and different wall channels, with smooth as well as rough bundles. It separates the pressure- loss coefficient into three components; plate friction (the effect of friction on the spacer plate), irrecoverable profile loss (effect of the flow area restriction) and channel wall friction (the effect of friction on the sur- rounding bundle rods or sheath). The model is fully described in /I/. One of its essential features is that the plate friction coefficient is based on the assumption of the free flow principle. It is necessary to define a Reynolds- number, Re , which governs the transition from a laminar to a turbulent boun- c dary layer on the spacer grid plate, based on the distance from the leading edge of the plate. In /I/ a typical value for Re of 8 x 10 is suggested, ^ c but this can increase to 5 x 10 for very low freestream turbulence and very smooth surfaces. Calculation with the analytical model computer code "SPATES" gave the results shown in Fig. 11. Three different values of Re were considered: O (direct c transition to turbulent flow and no laminar boundary from the leading edge of the spacer plate), 8 x 10 (transition inside the spacer grid) and 3 x 10^ (flow with mainly laminar boundary layer over the spacer plate). The predicted spacer pressure—drop coefficients are, as can be seen from Fig. 11, always lower than the experimental results. This confirms the trend from other expe- riments. The best fit is obtained with Re =0. The difference between expe- riment and prediction decreases with increasing bundle Reynolds number and, for Re > 3 x 104, is less than 0.1 (p/d = 1.5) or 0.05 (p/d = 2.0) . The rela- tive difference is thus less than - 20 % for the whole range of Re investi- gated. 7. DISCUSSION A spacer grid represents a local blockage of the bundle channel, characterised by the solidity (G). The spacer pressure-loss coefficient, Z, is significantly influenced by 0, and is very sensitive to changes in a especially at higher a (e.g. a > 0.4). At the design stage of the new spacer types an attempt was made to: 1) keep the a-value as low as possible and, 2) distribute the blockage so that the material of the spacer was as close as possible to the tube walls, and not in the middle of the subchannel where the gas-velocity is highest. Secondary effects, such as the influence on the heat transfer, were not considered. It is worth noting here that smaller o- and therefore Z-values, can be obtai- ned by increasing the tube pitch, see for example Type B in Fig. 10. The a- values of the spacer grids reported here are low - giving a small dependence upon Re. It is therefore possible to define a characteristic mean Z for each type (see Fig. 5 to 9), which is good enough for lay-out calculations of heat exchangers. Type A, % = 1.5 Rec = Rec = 800CX) Rec = 300000 experimental results 2-10 Rec=0 Rec=80000 Rec = 300 000 experimental results 10 2-10 2-10 Figure 11: Comparison of predicted and experimental spacer pressure-loss coefficient (Z) versus bundle Reynolds number (Re). 9 - The influence of the transition Reynolds number as a parameter is obvious and considerable (Fig. 11), Therefore Re must be known as accurately as possible in order to obtain good analytical prediction of spacer pressure loss. The value of Re can be found experimentally and this is probably best c done by means of accurate velocity measurements in scaled-Hip geometries of a few typical bundle channels, at a Reynolds number corresponding to this transition region. A value of Re =0 gives the best fit between experimen- tal and predicted results for the spacer geometries reported here. This indi- cates that the grid is such a flow disturbance that no laminar boundary layer is possible over it, and the flow is turbulent throughout. The particular shape of the predicted curve for Re > 0 corresponds to the transition from laminar to turbulent flow within the spacer. Obviously the handling of this phenomenon by SPATES is not satisfactory. The transition occurs progressively within the spacer, instead of abruptly as programmed in the code /I/. The Type B grid was chosen as the best spacer for the recuperators for the HHT-Demonstration Plant. As the spacer constitutes an essential and delicate part of high temperature heat exchangers with shell-side, longitudinal flow this type was extensively investigated. Mechanical stability and fretting behaviour, as well as pressure drop, were tested experimentally and the in- fluence of spacer design and axial separation of grids on the heat exchanger dimensions were investigated analytically /2/. The R+D work in this field, which EIR is continuing to pursue, is the syste- matic investigation of the influence of different geometrical parameters, such as plate profile and thickness, plate axial length and distribution of solidity axially, on the pressure-loss coefficient. The prediction model is also being critically reviewed and an analytical method is being drawn up for the calculation of the components for profile pressure drop, instead of the semi-empirical equations used in SPATES. Another logical, but as yet un- planned, step would be to investigate the influence of different grid geo- metries used in a heat exchanger on the heat transfer. 10 8. CONCLUSION • The improvement of low pressure-drop spacer grids for high temperature gas heat exchangers with longitudinal tube-bundles has been successful. From these results the spacer grids which have been tested have pressure- loss coefficients which are significantly lower than those currently used. • The accuracy of experimental results of within ±. 10 % is very satisfactory and was only possible by using very small tolerances in the geometrical dimensions. This accuracy should be adequate for practical use in the layout of heat exchangers, • The analytical model slightly underpredicts the spacer pressure loss, in particular at low bundle Reynolds Number. However, results are sufficient- ly accurate to justify the use of this code for calculation of the pres- sure-loss coefficient of widely differing spacer grid geometries. It is fairly independent of the bundle geometry and can also treat artificially roughened rods. The possible spacer geometries it can handle are typical- ly those of honeycombs and rings, but this does not exclude other shapes. Generally speaking, a wide range of grid types are allowed provided their webs can be approximated to semiinfinite plates in a free flow, or rings around the rods /3/. • The prediction in the lower Reynolds Number range is not as accurate as at higher Re, and additional work is also required at much higher Rey- nolds Numbers (e.g. 10 ) to prove the validity of the free flow principle under these conditions. Attemps are therefore still being made to improve the model. REFERENCES /I/ Barroyer, P.; "Analytical Model for the Prediction of Spacer Pressure- Drop Coefficients", EIR-Bericht No. 350, September 1978. /2/ Nordstrom, L.A., Fischli, H., Sulzer Brothers Ltd, Winterthur, CH-8400, Naegelin, R., Formerly Sulzer Brothers Ltd, now at Swiss Nuclear Safety Dep. (HSK), Wiirenlingen, CH-5303: "Recuperators for the HHT-Demonstration Plant", Transactions of the ASME, Engineering for Power, Vol. 102, No. 104, October, 1980. /3/ Barroyer, P.: "Verification of an Analytical Model for the Prediction of Spacer Pressure Drop Coefficients", EIR-Report SL-R/30, May 1979, Presen- ted at the Specialist Meeting on Heat Transfer, Session VI, Paper 1, Wurenlingen 14 to 16 May, 1979, OECD/NEA GCFR Programme. XA0055816 IAEA Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors in Duesseldorf, FRG 16-19 April 1984 Life Time Test of A Partial Model of HTGR Helium-Helium Heat Exchanger Masaki Kitagawa, Hiroshi Hattori, Akira Ohtomo, Tetsuo Teramae, Junichi Hamanaka, Mitsuyoshi Itoh and Shigemi Urabe Ishikawajima-Harima Heavy Industries Co., Ltd Japan, 100 AB STRACT Authors had proposed a design guide for the HTGR components and applied it to the design and construction of the 1.5 Mwt helium heat exchanger test loop for the nuclear steel making under the financial support of the Japanese Ministry of International Trade and Industry In order to assure that the design method covers all the conceivable failure mode and has enough safety margin, a series of life time tests of partial model may be needed. For this project, three types of model tests were performed. A life time test of an partial model of the center manifold pipe and eight heat exchanger tubes were described in this report. A damage criterion with a set of material constants and a simplified method for stress-strain analysis for stub tube under three dimensional load were newly developed and used to predict the lives of each tube. The predicted lives were compared with the experimental lives and good agreement was found between the two. The life time test model was evaluated according to the proposed design guide and it was found that the guide has a safety factor of approximately 200 in life for this particular model. - 1 - i. INTRODUCTION To develop the structural strength design procedures for the HTGR components, authors had already proposed an structural design guide which utilizing the Inconel 617 and Hastelloy X as the high temperature material, and applied it to the design of a 1.5 Mwt helium heat exchangers test loop. The guide was developed by extend- ing the ASME B & PV Code Case N-47 to the HTGR temperature and HTGR environments. The main items to have been modified (Refs 1,2) were listed in Table 1. The guide was intended to cover the conceivable fracture modes of the HTGR high temperature components and to include certain amount of safety margin to assure the total safety of the designed components. In order to increase the reliability of the guide, the life time test of the partial model of the HTGR components is needed. Three type of models were tested in relation to the above mentioned design guide. CD 1.5 Mwt Helium Test Loop (experimental test loop for nuclear steel making plant) which had been operated at 1000°C for 2000 hrs (design time was 5000 hrs). (2) Partial model of the center portion of above 1.5 Mwt Helium Heat Exchanger (center pipe + tubes) which had been tested to life time under accelerated loading condition. (3) Small components such as heat exchanging tubes which were usually run to failure under the variety of loading conditions. In this report, the some details of life time test of the partial heat-exchanger-center-pipe-and-tubes model is to be described. The model of center pipe and tubes were selected for life time test, because the preliminary strength analysis had revealed that the highest temperature portion of center pipe and heat exchanger tube assembly is going to be subjected to the severest damage during the life. 2. EXPERIMENTS Figure 1 shows a schematic diagram of the partial model tested. The model consists of eight 600 mm diameter helical coils (25.40 OD, 4t tubes) and 2000 x 20t center pipe made of Inconel 617. The eight tubes were welded to the stubs machined on the lower part of the center pipe and the other ends were fixed to the tube supports connected to the vessel wall. The center pipe is guided and is capable to be electrically driven along the vessel axis (max. stroke 150 mm; 0 ^ 25 mm/hour). - 2 - 2.1 Applied Load and Temperature The center pipe was cyclically moved up and down along the vessel axis to simulate the thermal expansion stress. The tangential angle between the end of helical coil and the first support of the each tube was different, and consequently, the stress on each tube caused by cyclic motion of center pipe differs each other. To apply the hoop stress to the tubes, the helium gas of 99.995% purity was charged to the predetermined pressure. The pressures for tubes varied to produce the variety of damage conditions. The load conditions of each tube were summarized in Fig. 1. All of the employed load conditions were set to be more deleterious than load conditions of actual components in order to accelerate the failure of the model. Load history is shown in Fig. 2. The helium gas temperature was controlled by electric heater so that the metal temperature of lowest portion of the helical coil was kept at 9 50°C. 3. EXPERIMENTAL RESULTS The fractured model is shown in Fig. 3. The fracture occured at the stub portion of the center pipe (not on the welded joints between the stub and tube). Tubes no. 1^ no.5 failed or leaked during the test at the time shown in Fig.2. The time shown in the fiqure is the accumulated time at max. operating temperature 9 50°C. As for tube no.3 and no.6, the cracks were found from the inside surface of the tube. Tubes no.7 and no.8 had not failed even after the prolonged exposure to 14.7 MPa (150 kgf/cm2). An example of fractured tube may be seen in Fig. 3. The tubes were generally expanded radially due to the internal pressure and cyclic bending and the cracks were initiated from inside surface of the tubes. Damage of other part of the tested model was generally small. Slight amount of decarburization were observed on the outside and inside surface of the tubes. The internally oxidized layer was observed near the inside surface of the tubes. All of the welded joint between cubes and stubs were found sound after the test. - 3 - 3° 4. LIFE PREDICTION AND SAFETY MARGIN OF THE DESIGN GUIDE In order to evaluate the accuracy of the life predic- tion method, the lives of the five fractured tubes were calculated using the following life prediction criteria and structural analysis method. 4.1 Damage Criteria for Life Prediction The loading condition of the life time test was such that the tubes were expanded outwards under the internal pressure and cyclic bending. The amount of the damage due to diametral expansion of the tube is so large that the linear creep fatigue damage summation rule of the code case N-47 can not be used. Therefore, a following failure criteria which includes the damage due to the strain accumulation was employed. ^f + ^c + *D = 1 (1) where re - 4 - where t , t are cyclic and static creep rupture time under a given stress. For Inconel 617, the cyclic rupture time is approximately 10 times as large as the static creep rupture data, the damage criteria may be written as /|t_ + o.9 /§£ = 1 (3) Src r Fig.4 shows the results of increasing mean strain fatigue test compared with the above failure criteria. It is seen that the above criteria is adequate for this type of loading. 4.2 Material Properties of Inconel 617 For accurate life time prediction, it is desirable to prepare the material properties of the particular tubes employed in the model at the test temperature. However, that is impractical time- and cost-wise. Therefore, in this work, material properties of the model tubes at the tested temperature in the test atmosphere were obtained from the data which had been existing at the time of life time test. Constitutive Equations of Creep Constitutive equation of the code was generated on the creep data of the plate material, and is shown as follows. e = kant + 6aqtr + a/E (4) e : total strain t : time (hr) a : stress (kg/mm ) k, B, n, q, r, E : material constants However, the tube and forged material has the different grain size and the different product form. Therefore, their creep equations should be different each other. Considering these differences, the creep equation of pipe and forged material was assumed to be same and 1.13 times as large as the code equation value. Strength reduction factor due to effect of decarburization was also considered. — 5 — Cyclic creep rupture strength Cyclic creep rupture strength of the tube and forged materials were estimated from the static creep rupture data of the tube materials and comparison of static and cyclic creep strength at 1000°C. The effect of the grain size on the creep rupture strength was estimated. Rupture elongation Although rupture elongation usually depends on the rupture time, the rupture elongation was assumed to be constant of 0.40 because the variation of the rupture time in the life test was small. 4.3 Creep Analysis of the Model and Life time Prediction For accurate prediction of the life time of the model, the finite element method with 3 dimensional solid element may be employed to analyze details of the structural behaviors including the analysis of the crack propagation. But this kind of approach is usually unpractical because it takes too much computing time. Here, a simplified analysis method was proposed and utilized under justified assumptions. The details of the method will be published elsewhere (3). The predicted life time obtained using the above symplified analysis method under the fracture criteria and the materials constants described in the preceding paragraphs were summarized in the Table 2. The results clearly show that in the tubes no.1 and 2 where thermal expansion stress is high, the creep damage The predicted life is seen to be in good agreement with the experimental rupture time shown in Table 2. Although the calculation shows the outer surface fails earlier than the inside surface, the actual failure occured from inside surface. As described earlier, this may have been caused by the internal oxidation on the inside surface. 4.4 Safety Margin of the Proposed Code Because this life time test was performed under the enhanced loading condition, it is difficult to meet the design requirements given in the code. However, the safety margin of the code may be estimated by calculating the creep life or creep-fatigue life on the basis of the code method and comparing them with the experimental life. - 6 - The creep damage limit for primary stress was evaluated for the tube no.4 and no.5 because these tubes were fractured mainly under primary stress, the creep fatigue limit was evaluated for tubes nos.l, 2, 4 and 5. The results are summarized in Table 3. Besides the safety margin of these items, the proposed other rules had been checked. From the results of calcu- lation, it was found that the most likely fracture mode was the creep-fatigue fracture, and that the minimum safety margin of the proposed code is more than 200 in life. 5. SUMMARY Following conclusions were obtained from the life time test of the partial model of the HTGR heat exchanger. (1) The good agreement between the predicted lives and experimental lives was observed. (2) A variety of mechanical and metallurgical informations were obtained. (3) An appropriateness of the proposed design code was proved for this model. ACKNOWLEDGEMENTS Authors sincerely acknowledge the appropriate suggestions given by Professor Yasuo Mori, Prof. Teruyoshi Udoguchi, Prof. Yoshiaki Yamada,Prof. Masateru Ohnami and other members of the special committee on structural strength of high temperature heat exchanger. REFERENCES 1. M. Kitagawa, J. Hamanaka, T. Umeda, T. Goto, Y. Saiga, M. Ohnami, T. Udoguchi, "A New Design Code For 1.5 Mwt Helium Heat Exchanger", 5th Int. Conf. on Structural Mechanics in Reactor Technology, Berlin, 1979. 2. T. Nakanishi, T. Nakata, M. Ito, and M. Kitagawa, "Development of High Temperature Helium Technology for Nuclear Heat Utilization — Review", J. of Society of Mechanical Engineers Japan, Vol. 84 no.757, 1981, pp. 1296 ^1303. 3. T. Teramae and H. Ohya, "Analysis and Design of Structures and Machinery Employed at High Temperature 1st Report ", IHI Engineering Review, Vol.14, No.2, 1981, pp. 69 - 77. - 7 - Table 1 Proposed Design Code for 1.5 MWt Heat Exchanger Scope Max. Temperature 1273"K (1OOCTC) High Temperature Materials inconel 617 (Hastelloy X) Environment Helium Items which differs from C.C.N-47 Items Modified or Added Items 1. Allowable Stress (Environmental Added Effect, Cold (Strength Reduction Factors, Work, Thermal for SQ, St, Sm etc.) Aging, Stress Aging, Welding} 2. ' Evaluation of Creep-Fatigue Modification of Evaluation Interaction Effect Method 3. Detailed Examination on Reflected the Temperature j Evaluation Method of Thermal Dependency of 4. | Evaluation of Creep Buckling Early Start of 3rd Stage Creep 5. ! Piping and Earthquake Design Revised and Added for B, i Method 6. i Evaluation Method of Elastic Adoption of Quantitative I Follow-up Evaluation Method i 7. Design Guide for Faulted Modified on Ductility Loss Condition Table 2 The results of simulated calculations Simulated Calculation TP. No. Outer surface Inner surface Experimental Rupture Failure Failure Time Time Time (tf) exp: hr 0c 0D (tf)o:hr (tf)j:hr TP. 1 and 0.620 TP. 2 0.380 793 0.652 0.348 1,033 695 and 666 TP.3 0.359 0.641 787 0.404 0.596 1,117 >700 TP.4 0.209 0.791 421 0.245 0.755 557 400 TP.5 0.153|0.847 341 0.184 0.816 445 372 Table 3 Safety Margin of the proposed code Allowable ITime to Safety Margin Limit tube no. Time (hr) I failure (hr) no. 4 40 330 8.3 for primary no. 5 2.4 215 90 stress no. 1 1.0 584 584 no. 2 1.0 555 555 0f+0c^1 no. 4 1.0 330 330 no. 5 1.0 215 215 - 9 - £Motor Driven Microstructure Examination = 890°C •TP.1-9 25.40D, 4t :=930°C 600 ^ Hencal No.2 Support Coil Dia. \ TP.1-8 TP.1-7 TP.1-6W NQ.1 Support TP.1-5 2000D TP.1-4W 20t TP.1-3 = 952°C TP.1-2 •TP.1-1 Temperature TP.1(as an example) Distribution Fracture Loading Tested Tube No. Differential PressureMPa(kgf/cm2) Support Location 8 (deg) 1 0.98(10) 117.25° 2 0.98(10) 117.25° 3 2.94(30) 163.25° 4 5.88(60) 140.25° 5 8.83(90) 140.25° 6 7.85(80) 207.25° 7 0.98^10.78(10-+110) 297.25° 8 0.98—10.78( 10—110) 320.25° Pig.l Loading Condition of A Mock Up Model - 10 - pio . CD 100 i£=25mm/h E 11cycles 75 8= 100mm ila c CJ- C =75mm E CO 1 tf= 25mm/h ^5 !B 50 47cycles •a ,, pip e b 4cycles 25 CD I CD o O a 0 (48) (91) (237) (613) (701) 15.69 u/ 1 1 1 TP 8 TP.7&8 (1 123) "I 160 (-• cd - 1 CL 13.73 -TP.5 ( 79) TP.5 Ruptured 140 "> o (117) '"(372) TP.6 11.77 120 CD TP.6 TP.4 Ruptured CO 9.80 100 CD CO TP.4 * • jns s CD (440) Q. 7.85 "TP.3 TP.3 80 TP 1 finishe d CD CD t—i IP.3 ,647 TP.3&6 - 'Ruptured Q. 5.88 (189) TP.1.-2.7&8 (666) 7 60 "o • • t (980) Tes t CD CO ,TP.1. 2,7 &8 3.92 (695 40 Z3 (125) (237) TP.2 Ruptured (. (866) O 721) to 1.96 i 20 III! . . i 100 200 300 400 500 600 700 800 900 1000 1100 1200 Time (h) Note : Numbers in brackets denote elapsed time Fig.£ Loading histories of fractured heat exchanger tubes (b) Center pipe manifold (a) Over-all view (c) Tube No. 1 (Upper side) Fig. 3 General appearance of the fractured model Jnconel 617 1000"C in Air e=1%. e=103 £1.0 . Eq(3) / £ ^.=0.5 .£0.5- • Experiment -L Prediction T 0 10 20 €f ; Strain at Failure, % Fig 4 Comparison of experimental and predicted lives under cyclic loading with increasing mean strain - 12 - /2f No. 8 XA0055817 Development, construction and analysis of the URKO intermediate heat exchanger R. Exner, M. Podhorsky, Balcke-Durr AG, Ratingen 1. Objective The objective of our development programme is to develop a functionally efficient He/He intermediate heat exchanger, capable for a heat transfer capacity of 125 or 170 MW and to prove its suitability by carrying out representative tests. As the vessel represents a barrier between the primary and secondary circuit the requirements to be met with regard to stability, reliability and availability of the components are particularly high. Special requirements arise as a result of - the high application temperature up to 950 °C - the necessary service life of 140.000 hours - the high temperature transients and differential pressures to be expected during operation - the large wall thicknesses needed as a result of this in the high temperature section where the temperature gradients are significant. In addition to the requirements made by the plant itself both the economic efficiency and the ability of the vessel to gain approval have to be taken into consideration when carrying out the development work. Furthermore it must be possible to carry out reexamination and repairs on the components. Fig. 1 shows the complete project divided up into its main components. Two years ago when we started work on this the availability of both high- alloyed steels (e.g. NICROFER 5520) and also the ceramic materials was very limited. Today, not least as a result of the influence provided by the heat exchanger development all the necessary semi-products are available. Furthermore with the He/He intermediate heat exchanger advances have been made into temperature ranges for which neither standards nor regulations with regard to dimensioning for stress exist. The guidelines of the ASME- Code (Case 1592) only apply up to temperature limits of max. 815 °C. Parallel to our detailed work on the U-tube concept a series of tests is being carried out to prove the safety of the concept and to optimize the components: - Basic tests - Component tets - Tests on partial components (up to a capacity of 10 MW) - Integral tests (for capacities above 10 A 10 MW test module is being manufactured and will be tested to prove the safety of the component development. The engineering work for the reference component commenced at the same time as the work on the He/He intermediate heat exchanger. This planning work together with the test results and the experience gained in the manufacture and testing of the test module will enable the reference solution to be worked out and an offer for an economic nuclear component to be submitted. 2. Function of the He/He intermediate heat exchanger The primary hot gas (950 °C) flows upwards centrally in a duct within the heat exchanger, is deflected in front of the hot secondary gas header and is drawn off around the heat exchanger tubes in a concentric annulus. After a further deflection in the U-bend section the primary helium cools to a cold gas temperature of 293 °C in a further annulus. The cold secondary helium (220 °C) is directed into a circular header and then into the heat exchanger tubes via helical compensation tubes. It flows into the hot gas header through the U-tubes in counterflow to the primary helium. The heat exchanger tubes are arranged on involutes in order to achieve the same packing density in the cold and hot annuli. As a result of the compact type of construction the diameter of the concrete cavern is small. The most important components of the He/He intermediate reheater are: - Hot gas header and supporting plate. The URKO heat exchanger differs from the conventional U-tube heat exchanger in that the separation of the media is guaranteed by two different components instead of by one single tubesheet. If} A coaxial structure with medium at a high temperature in the internal section of the vessel which guarantees both the thermal symmetry and the symmetry specific to the-Wow is achieved with this construction. The form of the hot gas header and the adjacent tube supporting plate was optimized taking into account all loads (internal pressure, external forces, temperature gradients) (Fig. 3) The construction of this element enables it to be placed in a space where the flow is non-existant. The header and the tube supporting plate are made of NICROFER 5520 forgings. Tube bundle and tube supports . Another part of the exchanger which can be counted as belonging to the critical sections of the heat exchanger is the tube bundle with the tube supports in the high temperature section. As mentioned at the beginning the heat exchanger tubes are arranged concentrically around the primary hot gas central duct in the case of the compact U-tube heat exchanger. The hot gas header with the secondary gas return tube is a determining factor as far as the minimum diameter of the tube bundle is concerned. For this reason not more than one tube cylinder or several separate tube packs can be selected for the above output. From the point of view of the greatest possible representativeness with regard to the reference component a study proved the optimum to be to divide the unit into 6 tube packs. (Fig. 4) The free non-tubed sections were filled with insulation material. With a diameter of approx.~-2 m the test module therefore corresponds geometrically to a 30 MW-heat exchanger. Thus the test heat exchanger in this arrangement is capable of providing all the required information and in addition representative conclusions can be drawn with regard to a 30 MW compact U-tube heat exchanger to a certain extent. The material to be used for the heat exchanger tubes and tube supports is NICROFER 5520. Support shells. The support shells connect the hot and cold tubesheets and form a significant part of the supporting system. They also convey the primary gas and very quickly take on its temperature. The tube supports are supported against the walls of the support shells and in the case of an earthquake they absorb the horizontal loads, which can result from the acceleration of the tubebundle. Insulation. Thermal insulation is used in all critical sections of the URKO-He/He intermediate heat exchanger. This insulation minimalizes the temperature losses in the central duct and the hot gas header, guarantees the separation between the hot and cold branch of the tubebundle and makes sure that the wall temperature of some components is reduced. Pressed fiber mats made of Al203/Al2Si02 will be used as insulation material for the 10 MW test module. tty Load supporting and reduction system. The efficiency of the load reduction system is of particular significance for the construction of the vessel. The high temperature sections can be relieved in such a way that a minimum primary load arises as a result of dead weight and thus the problem zone of the hot section including the hot bundle section is eased. The construction described briefly here is distinguished by the following features: - the hot header is not charged with primary mass flow and therefore it is not subjected to any direct primary temperature stress. - the thermal expansion balance can be determined by the design of the lengths of the arms (cold/hot branch). - the secondary gas return tube is not involved in load reduction during operation. The hot sections are relieved during operation because the load is reduced via the supporting shell, cold circular header and spring packs. - the simple tube shape is the precondition for economic manufacture and easy reexamination. 3. Analysis of the He/He intermediate heat exchangers The characteristic feature-of the He-He intermediate heat exchanger of the U-tube compact construction is the single U-tube bundle which was proved itself in practice and which expands independently of the load transfer system. In addition to the pressure and temperature load, however, the hot header is subjected to a dead load which is dependent on the temperature. The construction can, however, be designed in such an advantageous way that this primary load decreases as the temperature increases and that overloading of the components in the high temperature section is prevented by automatic regulation. The aim of the analyses which were carried out was: - to determine the constructive parameters which guarantee an advantageous load transfer during the individual load cases, - to calculate the temperature and stress distributions during the individual load cases, - to ensure the stress and strain values, - to make a statement about material fatigue, - to prove the insensitivity to vibration. The analysis has been carried out with the FE-Programme ANSYS on the CYBER 175 computer. 4. Thermohydraulic calculation model of the He-He intermediate heat exchanger As already mentioned it is essential to know the temperature distribution in the vessel during operation and during the individual load cases. Fig. 5 is a diagrammatic drawing of the load transfer system. The dis- placement of the cold tubesheet depends on the support shell temperature, the header temperature and the relevant temperature expansion coefficients. The heat exchanger was "divided up" into 20 sections over its total height. One element was provided for each material layer in a radial direction. The primary gas flow was represented by three flow tubes connected in parallel. The convection from the primary flow to the outer wall of the secondary gas tubes was simulated by the first of the three tubes connected in parallel. The second tube served to provide the convection connection between the primary flow and the support shell and the third flow tube provides the connection between he primary flow and the flow shell. The tube wall of the secondary gas flows was replaced by a concentric cylinder shell of the same thickness and the equivalent surface area. The ANSYS-gas flow elements with convection coupling are each constructed of a transport element and two convection elements. The element is designed in such a way that the convection heat flows only into the knots but not over the edge of the element between the knots. For this reason it is permissible to connect several elements in parallel with a mass flow which is distributed accordingly. By connecting the elements a set of equations is obtained which show pressure and temperature as independent values. The temperature solution approach which is of interest to us proves not to be linear due to the temperature dependence of the material values. In addition to the actual material values the so-called real constants must be defined for gas flow elements. Geometrical coefficients, the value of which cannot be- determined from the knot geometry alone, such as the hydraulic diameter, the area through which the flow passes, the loss length or the proportional convection area, are determined using these values. The results of the intermittent calculation for the load case "cooling down" can be seen in Figs. 6 to 10. In this load case the primary and secondary helium cools down with a gradient of -1 °C/min. Fig. 6 shows the change in the primary gas temperature along the path in the hot and cold channel for a period up to 4 hours. Figs. 7 and 8 indicate the same temperature patterns as fig. 6 for the secondary gas temperature and for the tube wall temperature. As is to be expected it can be seen that the greatest temperature changes occur in the hot channel. Figs. 9 and 10 show the changes in the radial temperature pattern as a function of time. Fig. 9 shows the temperature patterns for the primary and secondary helium temperature, the heat exchanger tubes and the central tube average temperature as a function at a distance of 3,6 m above the flow return. A similar curve is shown in Fig. 10 for the cold channel at a distance of 7,2 m. Based on the knowledge obtained with regard to the support shell temperature distribution which is dependent on time it is easy to calculate the relative displacement of the cold tubesheet as a function of time. (Fig. 11) This displacement pattern is then required in order to dimension the sets of springs. It determines the weight distribution and the distribution of the dead weights during this load case. The solution described here for load case "cooling down" should then, of course, be repeated for all relevant load cases. 10 5. The structural model of the hot header The hot header is of interest from the point of view of material fatigue because of the relatively high wall thickness of 90 mm, the complicated geometry with some discontinuities and the high temperatures and therefore a great deal of attention was paid to it. The calculation of non steady state temperature distribution were also carried out using the FE-programme ANSYS. The temperature changes based on the reuslts of the thermohydraulic system calculation were used as input data. For example, Fig. 12 shows the curves for the same temperature in the load case already mentioned, "cooling down". The greatest temperature difference over the wall occurs in section A-A (Fig. 13) and determines the stresses calculated elastically here. These can be seen in Fig. 14. As is to be expected the branch-off point is the point with the highest reference stresses. Similar calculation models were also set up for all the other interesting components of the He-He intermediate heat exchanger. The analyses carried out have confirmed that the thick-walled "hot header" is the component which can be expected to suffer the greatest fatigue damage during its service life. It was possible to safeguard against the primary stress in all specified load cases without any problem. This safeguard against stresses is carried out by comparison with the admissible stresses formed from those material values which are a function of time as well as those which are not. The ratcheting analyses according to T-1320 of the ASME Code, Case N47 could also be carried out successfully in all load cases. The calculation of the intermittent temperature field has shown that the header is only subject to a greatly reduced temperature change. The reason for this is the shielding of the "hot header" and not least the stored system heat which has a \/&ry good damping effect especially in the case of low mass flows. 11 6. Vibration analysis of the heat exchanger tubes The operability of the He-He intermediate heat exchanger is only guaranteed if both the static and the dynamic stress are kept within limits. It is particularly important to examine this type of stress for the heat exchanger tubes around which helium flows. It is important to ensure that the heat exchanger tubes cannot be damaged during the course of time by a possible vibration. In order to fulfill this requirement a vibration test was carried out on a model heat exchanger. A segment of the tube field of which .the geometry was retained, was built into the wind channel in such a way that the flow conditions which would prevail in the heat exchanger were correctly simulated (see Fig. 15). The measured values were recorded on a 32 channel magnetic tape device. The vibration behaviour of the tubes which had been determined by calculation was checked in a preliminary test. For this the tubes were stimulated with harmonic and random vibration using an oscillator. The vibration behaviour of the tubes tested proved to be similar and corresponded well with the calculations. In the case of harmonic vibration the four lower frequencies and a frequency range around 900 Hz was measured at a constant power level. The random vibration was simulated by means of a generator which was located in the Fourier analyzer used. The test equipment was fitted with 90 tubes for the flow test. It was run under a total of ten operating conditions at speeds of between 15,7 m/s and 50 m/s in the smallest inlet cross-section. IP 12 An existing shell programme was used to determine the vibration conditions of the tubes and the required free values were calculated. Fig. 16 illustrates the -strain distribution over the whole rolled out tube surfaced. A realistic outline of the overall vibration state of the tubes was obtained with the simulation method. The phase angles were generated with the help of a random number generator. These angles were inserted in the individual vibration conditions and superimposed. The result of the simulation process is the frequency distribution of the strains (Fig. 17). /si 13 7. Summary The overall objective of- this development order is indicated in this lecture and the construction of the new type of heat exchanger desribed in brief including the problems involved with carrying out the design in the time available. Observations up to now have shown that the compact U-tube heat exchanger is in a position to fulfill the specified requirements. The results of the tests carried out and the advanced manufacturing status of the test module point to a successful conclusion of our development project. To conclude we would like to show you some pictures of the interesting phases in the production of the test module. - 14 - Legend to the illustrations Fig. 1 - The components of PNP Fig. 2 - Components of the He/He intermediate reheater (URKO) Fig. 3 - Hot gas header Fig. 4 - The tubed and non-tubed sections of the reheater Fig. 5 - Load transfer system He-He-intermediate heat exchanger of U-tube compact construction. Fig. 6 - Primary gas temperature in load case "cooling down". Fig. 7 - Secondary gas temperature in load case "cooling down". Fig. 8 - Tube wall temperature in load case "cooling down". Fig. 9 - Temperature pattern in hot channel H = 3,6 m. Fig. 10 - Temperature pattern in cold channel H = 7,2 m. Fig. 11 - Relative displacement in load case "cooling down" Fig. 12 - Isotherms in the structure of the hot header in the load case "cool ing down". Fig. 13 - The largest temperature difference in section A-A of the hot header in load case "cooling down". Fig. 14 - Reference stresses in the hot header in load case "cooling down". - 15 - Fig. 15 - Arrangement of the measuring points for the vibration test. Fig. 16 - Type of vibration with the maximum load. Fig. 17 - Frequency distribution of the strains in the tube. IL_JF=IF=R NOV 1981 AUFTRAG - GmndUgenvcrsuche \z> - BautmhtBrtucbe VERSUCHSDURCH- APRIL 1984 - TeilksmpwwntsnvertuctM FtiHRUNG - lnt«gr*tveriuche - Komtr MaAnahnwn - Nachwws tUr *e Werfcttsff- •rtchApfung 125-rl70MW HERSTEU.UNG emw nia4lunB der VP H»4WZWT (URKO) 10 MW-TESTMODULS f - Boachmns, EKwurt, Koratr. MARZ1985 - Erftilluni - R*pariturfconz«p( ERPftOBUNG dM MARZ 1986 TESTMODULS \Z> \7? ANGEBOT Oil 1987 HfrHa/ZWr |URKO) 125+ 170 mt PNP-HE-He/ZWT(URKO) Fig. 1 Sek. Sek. Austritt Eintritt Tragplatte Zentralrohr Kaltgassammler Äußerer Ringraum Druckmantel Heißer Sammler Betriebsdaten Primär Sekundär Durchsatz 3,0 kg/s 2,9 kg/s Temperaturen 950/293 °C 900/2 20 °C Druck 39,9 bar ¿»3,5 bar Druckdifferenz 0,5 bar 1,0 bar CO er- Leistung 10,1 MW Abmessungen und Werkstoffe Rohrzahl 180 Rohrabmessung '20" 2,0 Rohrwerksfoff 2.4663 (Nicrofer5520Co) Behälterwerkstoff 1.6368 (WB36) Strukturwerkstoffe 1.7380 (10CrMo910) 1.6311 (20MnMoNi55) 1.5415 (15Mo3) Isolierung Heißgas Steigrohr Prim. Eintritt Vermischungseinrichtung 10MW He/He-U-Rohr- Warmetauscher Fig. 2 Werkstoff 10 CrMo 910 Nicrofer 5520 PNP He-He/ZWT (URKO) Heifer Sammler Fig. 3 : K: ^ — Innerer Stromungsmantel Sekundargas Isolierung Aufterer Stromungs- mantel Tragmantel KalterAst Druckbehalter Primargas HeiBerAst Primargas Zentrales Steigrohr Aufbau des PNP He-He/ZWT (URKO) Fig. 4 F=l I CZZ V Tragplatte - heifJer Sammler heiRePlatte WT-Rohre heifter Kanal Hattemantel Stromungsmantel isolierung Lastabtragesystem URKO Fig. 5 PrimargastemperalM Auskuhlen 240 0.0 16.6 33.3 50.0 66.6 83.3 100 Fig. 6 Lauflange (%l Sekundargastempacatur Auskiihlen i 320 160 0.0 166 33.3 50.0 66.6 83.3 100 Fig. 7 Loufldnge (% j ch ll 8 _ •5 1 fandfem f a Io s.c £E :3 i Kalte r K o o ID //// Kana l ID tlh ] a> U3 c II 7/ a) I 7 d o o oCO to o CO in fNJ [Go] en LL rsi Q 01 i I [0o] il * m 0 III 1 I O O o o 00 to o ID If) Relativverschiebung beziiglich der heiflen Rohrplatte [mm] in o o o o in 1 o o MIf Q Is 3 * / Auskuhle n 31 / M I Fig.12 15.0 Fig.13 9- ZEITeO.h ZEIT. 2.25h 31.3 %ri? Fig.U ••*•• 4. + OJ •»• • +_• • * •_* Bazeichnung des + ••*••• o*+ • •» +-T5 •* CL+ + + •• ••»••* M»flrohre$ • •*•••••••++•• o MeOrohr* 3850 XJ a a I in 3950 3750 1550 3 750 C 4 3 I *^^9} ^Wr ^W^^ 1 U30 1670 2170 2490 2650 3690 DM5 Bezeichnunq der Meflebcnen Fig. 15 ID D 2 1 8 10 I L I It !| 8 I z I 8 en LL 0 u 1 1 No. 9 A TWT¥1T¥?W »DUFTET! WTK? ÉJMfflMíE FEUERFEST-ERZEUGNISSE 1 INGENIEURLEISTUNGEN 1 ABTEILUNG ENERGIE-TECHNIK UIUIEII ANLAGENTECHNIK 1 1 1 XA0055818 1 DEVELOPMENT OF A NEW TYPE OF HIGH-TEMPERATURE INSULATION 1 MATERIAL AND ITS APPLICATION IN THE PNP PROJECT 1 •—,— 1 ~l— by DipL-Mineralogist R. Burger Dr.-Engineer R. Ganz DIDIER-WERKE AG Dept. Energie-Technik Didierstraße 31 D-6200 Wiesbaden Federal Republic of Germany Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors 1 Düsseldorf, Federal Republic of Germany, 16-19 April 1984 ZU zu zu ZI] Aufsichtsrat: Dr. Horst Burgard, Vorsitzender Hauptsitz: Wiesbaden, Ussingstraße 1¿-18 Abteilung Energie-Technik: Wiesbaden 12 Vorstand: Dipl.-Kfm. Dr. ¡ur. Martin Bieneck, Vorsitzender Registergericht: Wiesbaden HRB 2376 (Biebrich), DidierstraSe 31 Dietrich von Knoop, stellv. Vorsitzender Telefon: (0 01 21) 359-1 Telefon: (0 61 21) 605-0 zu Dr.-Ing. Gerhard Reinhardt, Dr.-Ing. Hans Stollenwerk Telex: 4 186 681 diw d Telex: 4 186 461 difo d ! DEVELOPMENT OF A NEW TYPE OF HIGH-TEMPERATURE INSULATION MATERIAL AND ITS APPLICATION IN THE PNP PROJECT 1. Basis of the Research and Development Program The following significant operating conditions apply to the prototype plant for nuclear process heat : . Temperatures of up to 950 °C . Pressures of up to 42 bar Maximum gas velocities of 65 m/sec Reducing or oxidizing atmosphere, depending on the loop. All heat-carrying and heat-exchanging components of the PNP have to be insulated to protect the metallic structures against temperature, and to maintain the thermal efficiency of individual components and of the entire plant. Therefore, thermal insulation is of major significance in the PNP project. Metallic liners in the primary loop fail because of the required service life under high-temperature conditions. Therefore, in this case, other materials such as graphite are used. Insulation is achieved by means of stuffed ceramic fibers. So this involves an insulation system where the insulating material is protected against erosion by a non-metallic liner. The ideal solution is the internal insulation loaded by the gas flow and which does not require an additional liner. In conventional technology, also erosion-resistant types of insulation are well-known which can safely withstand temperatures of up to 1700 °C and under inert conditions up to far beyond 2000 °C. The question arises why these materials and insulation systems cannot be employed. - 2 The reason is the particular requirements specific to the operation of a nuclear plant. Hence, the methods applied in conventional technology cannot be transferred directly to nuclear technology. Therefore, before the application of insulating materials and systems in the PNP, a development program became necessary to satisfy the specific requirements connected with nuclear components. The work for developing a PNP-compatible insulating material and system for the intermediate heat exchanger (URKO and HELIX design) are performed at DIDlER's in close cooperation and coordination with . INTERATOM/GHT and the component suppliers . BALCKE-DORR AG and . consortium STEINMULLER GmbH/SULZER AG. The main fields of the working program are two areas of the Helium/Helium Intermediate Heat Exchanger (He/He-IHX) : The primary hot gas riser duct (central line) with internal insulation loaded by the gas flow (without liner) and the intermediate heat exchanger external shell with internal insulation in the spaces not subjected to gas flow. The results obtained from this work can be applied directly to the steam reformer and can be used for hot gas valves and hot gas ducts. - 3 - 2. Development of the Materials 2.1 Material requirements The material requirements are governed by the operating data of the PNP . the experience gained from conventional insulation techniques, and the experience from test facilities (KVK, EVA I/I I AVR, etc.) The major requirements for the mateiial are . thermodynamic endurance for an operating period of not less than 140,000 hours of operation . bulk density < 1.0 g/cms . good insulating properties in helium, 40 bar . resistance in the event of accidents involving sudden pressure loss adequate mechanical strength . good thermal shock resistance . little abrasion during gas flows (v < 65 m/sec) 2.2 Feasible conventional materials The next figure shows conventional insulating materials that were discussed as basis at the beginning of the development work. It can be seen that none of these materials combines all the properties required. As experience shows, insulation fire bricks are characterized by- just satisfactory, but mainly poor resistance to thermal shock. parameter vacuum-formed insulation firebrick l_ 3^0 fiber-blankets fiber-based shapes (bubble alumina) /dimension A -mats content of A12O3 /weight-% bulk density < 18o ++ < 220 ++ /kg-nf3 porosity > g5 ++ 64 - /volume-% cold crushing strength /N-rnirf2 none none 20 ++ thermal conductivity (in air) , 0,12 ++ 0,13 ++ 1,3 - /W-(m-K)"1 thermal shock resistance /cycles ace. to DIN > 30 + < 10 - o workability o + machinabllity ++ excellent resistance against + good pressure-loss 0 o satisfactory thermal conductivity (in helium) + bad Conventional types of high-alumina insulation material DIDIER-WERKE S / - 5 - Densified fiber mats and vacuum shaped fiber parts exhibit excellent resistance to thermal shock, low thermal conductivity, and favorable behavior in the event of sudden loss accidents. A disadvantage is their lack of mechanical stability and the low resistance to abrasion and erosion. Thus, there is a fair chance that materials on the basis of ceramic fibers can satisfy the PNP-specific requirements if it becomes possible to increase the strength without reducing the positive thermal properties. Hence, the development program was defined : Adaptation of the conventional fiber materials to the severe marginal conditions of the PNP by producing Fiber Ceramics of adequate mechanical strength. 2.3 Fiber Ceramics 2.3.1 Basis The ceramic fibers used for the development work are obtained from the binary system SiO2-Al2O3. Two types of ceramic fibers can be distinguished : . The glassy fibers manufactured from the molten state by spinning or blast drawing, the Al2 O3 to SiO2 ratio being 1:1 or 2:1. . The fiber obtained from a salt solution by the spinning technique with subsequent annealing, with 95 % by weight Al2 O3 and 5 % by weight SiO2 . As far as we know, the glassy fibers are suitable in the PNP project only for temperatures < 650 °C because of the low thermodynamic stability. - 6 - °F ^000 - 3500 - 3000 - 2500 80 A12O3 cristallin fiber glassy fiber Binary system SiO2 -AI2 O3 2.3.2 Manufacturing process The ceramic fibers are prepared together with system-compatible inorganic binders and, in most cases, with organic preparing and shaping aids. The suspension or mix obtained is then processed into blanks or semifinished parts. Subsequently, the materiai is dried and fired. By the firing process, the organic additives are removed without residues. - 7 - Ceraalc d lnder systei Preparation Forming Drying Firing Mac n l n i ng General production flow By selection of the appropriate firing temperature, shrinkage of the material under service conditions is avoided. The semifinished product can be formed into shaped parts of accurate dimensions. - 8 - 2.3.3 Properties of Fiber Ceramics and Qualification Tests With the manufacturing techniques developed by DIDIER, Fiber Ceramics with bulk densities of 200 kg/m5 to 800 kg/m3 can be produced. In the course of preliminary tests a special binder system was developed by means of which the strength of the Fiber Ceramics is increased by the factor 10 to 20 compared with conventional vacuum shaped parts of the same bulk density. By varying the density of the Fiber Ceramics, the strength of the material can be adapted to the respective requirements. bulk density / kg «m 600 Al2 O, / SiO2 6,7 thermal shock resistance ace. DIN < 7 / cycles _2 cold crushing strength / N«mm 3,2 | 1,5 _2 cold flexural strength / M'mm 3,7 j 1,6 _2 4,3 -103 | 0,1 '103 E-Modulus / N *mm thermal expansion 20 °C - 100 °C 7,5 | 10,0 fiber J_ fiber Properties of a DIDIER Fiber Ceramic •— = Fiber Ceramic bulk density 600 kg/m3 3 1.0 - ,o_—o— = Fiber Ceramic bulk density 250 kg/m 0,9 - 0,8 - Z,0bar He 0,7 - 40 bar He > 0,6 - 1 bar He "u 1 bar He TO 0,5 - c o — — o 0,4 - — " —- *~ ' " Ibar N2 0.3 - --1bar N2 0,2 - 0,1 - 0,0 0 100 200 300 £00 500 600 700 800 900 1000 temperature / °C Thermal conductivity of Fiber Ceramics in air and helium. DIDIER-WERKE S - 10 - It has been proven that at temperatures > 500 °C materials with bulk densities ^400 kg/m3 have a lower thermal conductivity than conven- tional materials. Before being applied to a nuclear plant component, the Fiber Ceramics must pass several qualification tests. These have been initiated and are being performed at present. Abrasion test The abrasion test series served for determining the abrasion behavior of the parts of Fiber Ceramics with plane contact under normal atmospheric conditions. The following combinations were selected : Fiber Ceramics/Fiber Ceramics Fiber Ceramics/steel Fiber Ceramics/graphite . Graphite/graphite Reactor graphite was used as reference material in these tests. The cyclic frictional motions and the loads per unit area were selected to be similar to the He/He intermediate heat exchanger. As was to be expected, it was revealed that the Fiber Ceramics exhibit higher abrasion relative to one another than with respect to steel and graphite, or as regards the graphite/graphite pair. At the beginning, abrasion is high, but is reduced after approximately 200 cycles and then increases noticeably less. . Erosion test In dust-loaden gas flows the produced Fiber Ceramics exhibit markedly greater abrasion than graphite, but under the test conditions that could be realized also the behavior of graphite which is qualified for gas cooled reactors was not acceptable. - 11 - In air with a very low dust content and at gas flow velocities of 200 m/sec, Fiber Ceramics show no signs of abrasion according to our knowledge and experience. . Other tests Existing computer programs are reviewed at present for their suitability to describe the behavior of Fiber Ceramics under PNP conditions. In addition, the first series of tests were initiated to determine the long-term behavior of Fiber Ceramics by means of fatigue tests. - 12 - in Microstructure of DIDIER - Fiber - Ceramic Examination by Scanning Electron Microscope DIDIER-WERKE 13 - Microstructure of DIDIER - Fiber - Ceramic Examination by Scanning Electron Microscope DIDIER-WERKE - 14 - 3. Testing of Structural Parts, Engineering 3.1 Requirements of the insulation system An insulation system suitable for PNP must satisfy the following requirements of a general and . a specific nature related to the structural parts. The general requirements are . high insulating capability insensitivity to sudden pressure loss accidents ease of installation . prevention or limitation of bypass effects. Examples of requirements specific to the structural parts are e.g. the withdrawability of the insulation for repeat tests special geometric configurations 3.2 Basic design features The insulation concept for riser ducts proposed by DIDIER at present is described below : An external steel shell will be insulated with segmenttype shaped fiber parts. These shaped parts can be machined to very close tolerances. Axial joints will be sealed with cylindrical parts, preferably made of ceramic materials. The segments can be provided with holes so that an interlinked insulation can be built up with staggered joints. 15 - sealing bolt fiber ceramic densified fiber mat steel shell Basic design features of the DIDIER insulation concept. A densified fiber mat located between steel shell and shaped parts provides the necessary initial stress and at the same time serves for accommodating differential thermal expansion. With the concept presented, damaging thermal stresses are avoided which may occur in the shaped insulating parts due to stationary or non-stationary differential thermal expansion. The concept is characterized by . division of the annulus into a practical number of individual segments . reduction of the heat storage capacity of the individual segments by reduction of the mass (axial holes) reduction or prevention of bypass effects - by cylindrical sealing bolts on the cylindrical axial joints - by a bracket structure with associated ceramic seals . clamping of the shaped parts with a flexible fiber mat within the metal shell results in frictional connection on the axial joints. - 16 - The concept developed can be adapted, e.g. by adding holders, depending on the insulating materials, metallic structural, and marginal conditions. The concept exhibits the following characteristics : . low weight in relation to the mechanical strength pre-assembly is possible outside of the pressure tube (modular construction) . the insulation build-up is flexible and can be adapted to a great many marginal conditions setting free of fiber particles is reduced to a minimum the system is capable of compensating manufacturing tolerances of the metallic components when installed in place. 3.3 Preliminary test of the insulation concept The insulation system described above was subjected to a preliminary cyclic test. Marginal conditions : air, 1 bar . temperatures from 20 °C to 1100 °C heating and cooling down velocity < 10 K. The shaped fiber parts installed withstood the transients. Calculated and real operational behavior were compatible as regards the thermal behavior; the operating conditions were reproducible. In spite of extensive data acquisition, there was no evidence of damage to the insulating material or of opening joints. - 17 - Shaped Fiber Ceramic for preliminary test of the DIDIER insulation concept. - 18 - Testing equipment with data acquisition 4. The New Insulating Material in the PNP Project The material and insulation system were not yet applied in the PNP project because the R & D program is not completed and the final qualification is still pending. However, the manufacturing process has been developed far enough so that the insulation of the PNP half axial valve prototype could be performed like any normal supply and assembly contract. This lining will be subjected to helium, 40 bar, 918 °C and flow velocities of up to 70 m/sec. The following figures show various shaped insulation parts and the valve body lined with them. /fo - 19 - Insulation parts shaped of DIDIER Fiber Ceramic for the PNP half axial valve prototype. - 20 - /I) Outer shell of the PNP half axial valve prototype after insulation assembling. On the basis of the materials and insulation system developed by DIDIER, solutions also for other nuclear plant components can be worked out. No. 10 XA0055819 Seismic Analysis of a Helical Coil Type Heat Exchanger Isoharu NISHIGUCHI, Osamu BABA and * Hiroshi YATABE Japan Atomic Energy Research Institute * BABCOCK HITACHI K. K. ABSTRACT The intermediate heat exchanger (IHX) which forms the reactor coolant pressure boundary is one of the most important components of the Multi-purpose Experimental Very High Temperature Gas-cooled Reactor (ex. VHTR) under development at Japan Atomic Energy Research Institute. This paper presents the results of the finite element modeling, eigenvalue analysis and dynamic response analysis of the IHX. For the model ing, the structure of the IHX was separated into a he Iical tube bundle, inner and outer vessels, and a centerpipe. The eigenvalue analysis was made for each structure with a detailed three-dimensional finite element mode!. Then the simplified model of the whole structure of the IHX was constructed using the result of the eigenvalue analysis. A dynamic response analysis was made for the simplified model with and without stoppers of the helical tube bundle supports and the centerpipe. The effect of stoppers on the behavior of the centerpipe, the helical tube, and the connecting tube is discussed. 1. INTRODUCTION The finite element model (FEM) is a powerful method for structual analysis, but it is expensive to model a complicated component such as the IHX which is made up of a centerpipe, inner and outer vessels, tube bundle supports and so on. It is especially expensive when a dynamic response analysis is performed because of a large amount of CPU-time. It is therefore necessary to simplify the model to some extent. Simplified models using beam, spring and mass elements are often used to carry out dynamic response analyses. When the modeling is not adequate, however, it is difficult to predict the local vibration behavior of the parts of a component under consideration. We therefore decided to carry out the seismic analysis of the IHX as follows: i) Regard inner structure of the IHX as an assemblage of several structure elements and make a detailed local model (DL-model) for each structure element. ii) Make the detailed inner structure model (Dl-model) by combining DL-models. iii) Make the simplified inner structure model (Si-model) using beam, spring and mass elements. In this step, static FEM analyses are carried out to estimate the stiffness of the joint between structure elements in case of need. iv) Evaluate the propriety of the Si-model by comparing the mode shapes and frequencies of the Si-model with those of the Dl-model. Modify the stiffness of the Si-model when necessary. v) Make the simplified whole structure model (SW-model) by combining the Si-model with the vessel model which is made up of beam, spring and mass elements. vi) Perform the dynamic response analysis of the SW-model using an adequate seismic wave. vii) Perform dynamic response analyses or response spectrum analyses of structure elements as lower connecting tubes where local vibration behavior should be predicted exactly. At this stage, time history responses of joints between structure elements obtained in vi) are used as the boundary conditions. To allow for the thermal expansion, the heat tube bundle supports (HTBS) and the centerpipe are suspended from the upper part of the IHX, and gaps are provided between the centerpipe and the HTBS and between the HTBS and the inner vessel. In dynamic response analysis, therefore, the effect of gaps should be taken into consideration. 2. DESCRIPTION OF THE IHX The ex. VHTR is made up of 2 loops (A-loop and B-loop) and the IHX is located in each loop. The structure of the A-loop IHX in 1981 is shown in Fig.l and Table 1 shows the main specifications of the IHX. Both IHXs have similar structural arrangements. In each IHX, secondary helium gas flows inside heat transfer tubes, while primary helium gas flows outside in the opposite direction. The low temperature secondary helium gas flows downward in helical tubes from the upper tube sheet to the high temperature tube sheet at the bottom of the centerpipe. In the centerpipe the gas flows upward from the high temperature tube sheet to the upper exit nozzle. On the other hand, the hot primary helium gas enters the IHX at the 'Ty bottom and flows upward outside the helical tubes and exits from the upper nozzle on the side of the pressure vessel. After being pressurized by the gas circulator, primary helium gas enters the IHX once more at the nozzle located at the upper part of the pressure vessel, and flows down the annular space between the inner vessel and the outer vessel for cooling. Helical heat transfer tubes are supported by the HTBS as shown in Fig.2 and the HTBS, the centerpipe, and the high temperature tube sheet are suspended from the upper part of the IHX and are allowed thermal expansion downward. The centerpipe and the manifold type high temperature tube sheet are in a body and the relative displacement between the tube sheet and heat tube bottom ends is absorbed by flexibility of the connecting tubes. The IHX is located on the floor of the IHX room within the reactor containment vessel as shown in Fig.3. 3. MODELING OF THE IHX Modeling of the A-loop IHX was carried out according to the flow mentioned above. * DL-model * To begin with the DL-fliodels were made for the structure elements; -upper and lower connecting tubes, -helical tube bundles and the HTBS, -the centerpipe and the tube sheet. Fig.4 shows the DL-models of the helical tube bundle and the HTBS and Fig.5 shows the mode shapes of these models obtained by the eigenvalue analysis. It can be observed that vibration behavior of this structure element is governed by the stiffness of upper and lower connecting tubes and supporting rods. * Dl-model * Fig.6 shows the detailed inner structure model which consists of the helical tube bundle, the HTBS, the centerpipe and so on. Because the vertical stiffness of helical tubes is low enough compared with that of the HTBS, tubes were treated as an additional mass to HTBS. Horizontally, however, because the movement of the HTBS is restrained by helical tubes, beam elements which model the HTBS are jointed by rigid circles. In the A-loop IHX, 270 tubes are connected to the lower and upper tube sheet and these are modeled by 8 pipes as shown in Fig.6. tff * Si-model and SW-ciodel * Based on the Dl-model, a two dimensional Si-model which consists of beam, pipe and mass elements is made as shown in Fig.7. Fig.8 and Fig.9 show the result of eigenvalue analyses of the SI-model and the Dl-model using SAP-V^). Mode shapes and frequencies of these models agree well. Then the SW-model is made by combining the Si-model with the vessel model which is made of beam, spring and mass elements. 4. DYNAMIC RESPONSE ANALYSIS OF THE IHX 4.1 Modeling and Analytical Conditions Time history analysis was performed with the SW-model shown in Fig.10 using the floor response of a design basis earthquake Sg for ex. VHTR. Gap elements are connected between the centerpipe and the HTBS and between the HTBS and the inner vessel. By combining 2 gap elements the condition below is posed between 2 nodes at both ends of gap. |uj_-uj| > Lgap : closed (works as a spring) I UJ_-Uj | < Lgap : open where UJ_,UJ '• horizontal displacements of 2 nodes at both ends of gap. Lgap : gap length. By employing the constraint above, behavior of the centerpipe, the HTBS and the inner vessel which contact each other at the right and the left side of the central axis of the IHX can be simulated. A general purpose FEM program ANSYS^) was used for the response analysis and Rayleigh damping was assumed with 3 % critical damping. The Houbolt integration scheme with time step 0.0005 sec. was employed in direct integration. 4.2 Results The results of the response analysis are shown in Fig.11 to Fig.15. Gap stiffness k=10' kg/cm is assumed in the calculations. Fig.11 and Fig.12 show horizontal displacement response at the bottom of the high temperature tube sheet calculated with Lgap=0.5 cm. In the calculation of Fig.11 the gap element at the bottom of the tube sheet is not used and a pair of gap elements between node 1 and node 41 which simulates the stopper was employed in the calculation of Fig.12. It can be observed that horizontal displacement reduced by a factor of four by the stopper at the bottom. Fig.13 to Fig.15 show horizontal relative displacement between both ends of the lower connecting tube. Analytical conditions of Fig.13 and Fig.14 are the same as those of Fig.11 and Fig.12. Amplitude decreases and frequency increases by locating the stopper. Fig.15 shows the result of changing Lgap to 0.1 cm. Other relevant conditions are the same as Fig.14. Amplitude decreases and frequency increases more compared with the result of Fig.14. 5. CONCLUSIONS Modeling, eigenvalue analyses and dynamic response analyses in an earthquake using FEM was performed for the IHX. In the stage of modeling, good agreement between Dl-model and Si-model was obtained. In the dynamic response analyses using SW-model, gaps between the centerpipe and the HTBS and gaps between the HTBS and the inner vessel are modeled by gap elements and the effect of the stopper on the behavior of the centerpipe and connecting tubes is examined. The results show that gap elements used can simulate the behavior of the structure elements in earthquakes and that the vibration of the centerpipe and the connecting tube can be restrained effectively by the stopper at the bottom of the centerpipe. ACNOVLEDGMENT The authors would like to thank S. Fujita for assistance in performing FEM analyses. Thank is also due to C. Winsel for useful suggestions. 1 ) SAP-V; A Structural Analysis Program for Static and Dynamic Response of Linear Systems, USC version July 1976. 2) DeSalvo, G. J. and Swanson, J. A., 'ANSYS User's Manual , SASI, Rev. 4, 1 February 1982. Table 7 Main specifications of IHX A loop B loop type helical coil helical coil heating arcs (m*) i j eo ] 4B5 heat tube tf 3 1.6 x 41 tf> 2 5.4 x 4t ;size (mm) ; number 27 0 474 number of layers 1 3 ] 7 of coils material Hastelloy X"R Hastelloy XR 20870 12750 ?ry Helium Oolltl Helium Fig -1 Intermediate Heat Exchanger (A-loop) 4 Supiiotl Lvg Fig. 2 Tube support structure of the A-loop IHX _ CL +52.0 ftaaetor Canutmwrit Vessel kafnellag Machine Kcaecor Auxiliary (Auxiliary Kaulpacnt Ant) Kaactor Ancillary kttlMlag (Control Systaa Ar»») A-loop IEX B-leop IKX Reactor CL -i3.S * CL -20.C Fig.3 Reactor building of the experimental VHTR (A-A section) tbe colunm the aedion colmm the outenoost column Fig.4 Modeling of helical tubes and support plates tff j V\ 1st imi Jrt) moit Fig.5 Mode shapes of the helical tubes 21570.0 7D.0 Fig.6 Detailed inner structure model Fig.7 Simplified inner structure model II 1.9BIIZ 2.38Hz 7.76Ht 7.OHz 3rd ante 4th »ode O.42HE 0.41Hz lit sodc ode Fig.8 Comparison of the mode shape of the detailed model with that of the simplified model F<55) 38 3.0- 16 0 @ ~$mP% 13 !U££. l.o- 6 *-rf _21 JUT*5 dlr ^ -9fr (^ SO 1.D © © ID ?c (_) : beam element -w -B6- 7S A : pipe element S) D : spring element -&fr -[Hi-: gap element 5 IWILJJ 7 27 47. £3 Fig.9 Comparison of frequencies between 1) 1© 1® Dl-model end Sl-model t 2^,, «ifilhi_S A «i 60'^ V 25j •A SE £> f ' *^ I I TT 56 T 2.5 iFig.10 Simplified whole structure model 2.0 1.0 -1.0 -2.0 without a pair of gap elements between node 1 and node 41 -2.5 0 1 2 3 4 sec5 Fig.11 Horizontal displacement at node 1 with a pair of gap elements between node 1 and node 41 0.5 0.25 B 0 -0.25 -0.5 0 12 3 4 sec Fig.12 Horizontal displacement at node 1 2.4 without a pair of gap elements between node 1 and node 41 2.0 1.2 0.8 in CM -0.4 -0.8 -1.2 -1.6 0 1 2 3 4 sec Fig.13 Horizontal relative displacement between both ends of the lower connecting tube 1.2 Lgap*0-5 cm 1.0 with a pair of gap elements between node 1 and node 41 0.8 0.6 0.4 'm 0.2 CM 0 -0.2 -0.4 -0.6 -0.8 sec Fig.14 Horizontal relative displacement between both ends of the lower connecting tube 0.20 with a pair of gap elements between node 1 and node 41 u 0.12 in c\j 0.04 -0.04 -0.12 2 3 45 sec Fig.15 Horizontal relative displacement between both ends of the lower connecting tube No. 11 PAPER 58 DESIGN AND DEVELOPMENT OF STEAM GENERATORS XA0055820 FOR THE AGR POWER STATIONS AT HEYSHAM II/TORNESS A N Charcharos, A G Jones, National Nuclear Corporation Ltd. SYNOPSIS The current AGR steam generator design is a development of the successful once-through units supplied for the Oldbury Magnox and Hinkley/Hunterston AGR power stations. These units have demonstrated proven control and reliability in service. In this paper the factors which have dictated the design and layout of the latest AGR steam generators are described and reference made to the latest high temperature design techniques that have been employed. Details of development work to support the design and establish the performance characteristics over the range of plant operating conditions are also given. To comply with current UK safety standards, the AGR steam generators and associated plant are designed to accommodate seismic loadings. In addition, provision is made for an independent heat removal system for post reactor trip operations. 1 DESCRIPTION 1.3 Tube elements for the HP and reheater units 1.1 General The HP elements are continuous (2 flow/platen) from the feed inlet penetration to the superheater outlet There are four boilers in each reactor and each at which point the elements are bifurcated into boiler comprises three once—through HP units and tailpipes which are routed through the steam three single stage reheater units which in normal penetrations. The material of the HP tube elements operation generate steam to drive the main turbine. is graded from austenitic stainless steel type 316 H A further bank of tubing is provided beneath each HP at the top through 9% Cr 1Z Mo to 1% Cr \% Mo at the unit and operates in conjunction with the HP unit to bottom. The tubes for the primary economiser are remove decay heat when the reactor is shut down. 25.4 m o.d. made of \X Cr i% Mo with carbon steel This bank also assists in the maintenance of the fins on a staggered arrangement; the 9% Cr banks reactor gas inlet temperature within the required are made of 28 mm o.d. plain tube on an in-line range during reactor start-up and shutdown pitch; and the stainless steel tube bank consists operations. The general layout of the steam of 36 mm o.d. plain tube on a staggered pitch generators within the reactor is shown in Fig 1 and arrangement. Inconel 600 and 5% Cr transition a cross-section of an HP unit with an associated joints are used at the upper and lower material reheater and decay boiler is shown in Fig 2. change locations respectively. The boiler units, rectangular in section, are The elements are supported by means of a welded located in the annulus between the reactor gas spacer system of the same material as the tubing. baffle and vessel wall. The tube arrangement for The elements are connected at various intervals down the heating surfaces, consisting of plain and finned the unit by links to transverse support beams which tubing, is formed from horizontal parallel straight transfer tube bank loads to the main casing tubes and associated return bends. Tubes axes are structure. Support beams and associated links are arranged in the circumferential direction to take of austenitic stainless steel. advantage of the maximum length of straight tube. The reheater elements comprise 38 mm o.d. plain 1.2 Plant layout and steam'and feed tube, each element having four flow paths. The connections tubes are bifurcated to reduce the number of branches on the large bore headers. The elements The gas is constrained to flow through the reheater are supported by means of 316 H stainless steel and HP boiler units by means of permanent casings, spacers and hangers In a similar manner to that of gas seals and annular plates. The boiler units are the HP boiler. The tailpipes between the tube banks supported from below by two carbon steel beams which and the headers are attached in the shops and the are suspended from supports on the gas baffle and assembly is transported to site for unit erection. from the vessel wall. The reheater banks are suspended from the vessel roof and connected to the 1.4 Decay heat boiler HP unit with a flexible seal. The tube banks for the decay heat boiler are located Each quadrant has its own feed water system directly beneath the main boiler economiser sections incorporating the usual complement of main and (Fig 2). One inlet feed penetration and one outlet start-up feed regulating valves, see Fig 3. steam/water penetration are provided for each boiler Pipework distributes the feed water to six quadrant. External valve Isolations allow the penetrations through which tailpipes pass to the operation of the decay heat system when fewer than three boiler units. A control valve in the pipework four quadrants are required. to each penetration is provided to ensure distribution stability and steam temperature Each decay heat bank consists of twelve rows of control. High pressure steam is conveyed from each finned tubing of the same geometry and material as boiler through nine penetrations in the pressure the main boiler primary economiser. Ferritic vessel wall and is collected into steam headers. material is used throughout the platens so that The reheat steam is fed into and out of the pressure operation with lower quality water than that vessel through a total of six combined headers and specified for the main boilers can be tolerated. penetrations per quadrant. External reheater The boilers are brought into service for reactor pipework incorporates a bypass to limit the outlet start-up duty and are also initiated automatically steam temperature by means of steam attemperators. after reactor trip as part of the reactor shutdown sequence. Facilities for control of feed flow and pressure are provided. R737-PAPER58( 1)HB R7 37-PAPER58C 2)HB 1.5 Tailpipes approximately 1% of the boiler gas flow. The coolant flow, in association with baffles around the 1% Cr i% Mo tailpipes pass from the tube plate in unit at various levels serves to prevent hot CO2 the feed inlet penetration to the bottom of the bypassing the boiler units and flowing down to the boiler units with sufficient flexibility to lower part of the annulus. Division plates are accommodate Che relative thermal movements. Super- provided between each quadrant below the main gas heater internal tailpipes are routed from the top-of seals to isolate adjacent boiler/circulator systems the boiler units through a tube plate in each outlet in the event of one quadrant being shut down. The penetration to external headers. Tailpipes connect division plates are made of mild steel plates the platens of the reheaters to large bore internal suitably stiffened, the structure being adequate to headers. withstand the full differential pressure across the boilers. Differential loads on the division plate are reacted via attachments to the liner and gas baffle. A door is provided in the division plate at The boiler casings are designed to support the tube floor level for inter-quadrant access. banks in a horizontal position during assembly at works, transportation and up-ending prior to lifting 1.9 Penetrations into the reactor vessel. In addition to supporting the tube banks, the casing is designed to withstand Two economiser penetrations are provided for each gas pressure and temperature differentials during boiler unit, see Fig 6. The penetration has a tube normal operation of the plant. Provision is made at plate at the outer end to which the economiser tubes the bottom of the HP unit casings for support are welded. Flow stabilising orifices are provided pedestals to the main support beams. Reheater at each tube location. The tube plate extension is casings and main unit casings are fabricated from welded to the penetration liner tube on site. austenitic stainless steel. Vertical baffles are Attached to the tube plate is a large diameter provided over the full height and breadth of the flanged header comprising feed inlet branch and ceheater and HP units to improve the acoustic drain branch. The flange can be removed for behaviour of the casing cavity. These baffles form plugging tubes in the event of a boiler element an integral part of the casing structure. Flights leak. are included to minimise gas bypassing between tube banks and casing walls. The three superheater outlet penetrations associated with each boiler unit are fabricated into units 1.7 Boiler supporting structure comprising steam tubes and insulated sheath tube and complete with a forged end plate at the inboard end, An annular ring common to all four boilers is see Fig 7. attached to the lower end of the HP units to control radial expansion of the units and provide convenient The annulus formed by the sheath and the insulated attachment points for the main gas seal. Each main penetration liner tube is provided so that the unit is supported at its base by two fabricated penetration sheath tube acting as a cantilever can radial beams through pin joint connection between take up the vertical expansion of the boiler units. the beams and the base of the unit casing'. The radial beams are supported by flexible slings Cold pull is applied to limit the stresses in the connected to tubular supports attached to the gas sheath tube during normal working conditions. The baffle and to dummy penetration in the liner. The individual steam tubes pass through the forged end slings have hinged connections to permit transverse plate and connect into 3team headers mounted on the movement and are preset so that in normal operation outside of the reactor vessel. Access is provided the flexural stresses are low, see Fig 4. Connect- to the steam tubes for blanking in the event of a ions are provided at the top of the HP unit between boiler element leak. superheat penetrations and casing to control thermal movements in a radial direction, and to accommodate Plate baffles and insulation assemblies are provided vertical and horizontal expansion of the units. on both feed and steam penetrations to control gas circulation and heat transfer. Each reheater is supported by four slings suspended from the concrete pressure vessel roof, see Fig 1 The reheater inlet and outlet steam penetrations and 2. The reheaters are also connected to the HP comprise large bore stainless steel tubes externally boiler unit3 by means of a rolling seal which allows insulated and contained in a sheath tube. The outer vertical and radial differential expansion between end of the steam tubes is flanged to provide access the reheater and the HP units. A sliding bracket to the header for inspection and remote tube maintains radial and circumferential alignment plugging in the event of a reheater element leak. A thereby preventing undue loads on the seal. Seismic branch is provided for connection to the reheater loads on the boiler structure are reacted through steam pipework. restraints connected between the casings and dummy supports on the vessel liner wall and differential Shielding at the outer end of the feed and steam loadings on the reheater slings and the steam penetrations limits external radiation to acceptable penetrations. levels. All boiler feed and steam penetrations are provided with external secondary retention features 1.8 Ga3 seals and division plates to limit gas side discharge to acceptable levels in the unlikely event of a penetration weld failure. The main gas seals are continuous around the Inner and outer walls of the annulus at the bottom of the 1.10 Boiler annexe components main unit casing. Each seal comprises two tubes spaced apart and sandwiched between layers of shims Each boiler has an independent feed supply with which are backed and protected by plates. One tube associated control and trip valves which divides of the seal assembly is connected via a membrane into six branches which connect to the outer end of plate to the boiler units, the other tube via a the feed penetrations. Separate branches are membrane plate to ledger sections on the liner and provided for connection of emergency feed systems, gas baffle skirt plates. Joints In the system are see Fig 3. covered by lap sealing plates. The seals allow relative thermal movement in both the horizontal and The HP steam pipework is routed from the outlet vertical directions, see Fig 5. Low temperature branch of the external superheate headers; each coolant tapped off the gas baffle is piped around header serves three boiler units. The pipework is the bottom of the units to maintain the surrounding anchored back to the concrete vessel. In addition voids cool and to offset leakage through the seal. to the steam mains to the turbine, connections from The gas leakage through the seals will be the HP steam pipework permit discharge of steam or R737-PAPER58C 3)HB R737-PAPER58( water to the start-up vessels, and to the LP vent Occasional wetting of austenitic tube material Is system, via appropriate control valve stations. acceptable in the short term provided water quality Provision is also made for pipework to bypass the HP Is controlled within defined limits. Potential long boiler for circuit clean-up operations. term problems are avoided by using high purity feed water during power operation and by controlling Reheater bypass pipework is. provided between the steam temperature at entry to the austenitic section reheat inlet pipework and the branches on the of the secondary superheater to a nominal minimum of reheater outlet penetrations, complete with an 70°C superheat at which droplets are absent. All attemperator device to obtain the required final weldments are heat treated to condition the reheat steam temperature. materials. 2 FACTORS DICTATING LAY-OUT AMD DESIGN Erosion corrosion 2.1 Materials In regions of high water velocities such as down- stream of control orifices and at tube bends, Apart from the requirement to use materials with erosion corrosion can occur resulting In severe adequate strength at the specified temperature and metal wastage. The process is a function of a pressure, there are four factors governing the number of factors including the material composition choice of materials in the various sections of the of the attacked surface. Tests have shown that steam generator. These are: alloy steels are resistant to attack and so these steels are used for the water sections of the boiler (i) Gas side oxidation units. (ii) General water/steam side corrosion (iii) Stress corrosion (iv) Erosion corrosion. Taking account of all the above factors 1% Cr i% Mo Gas side oxidation steel Is used for the feed Inlet, primary ecotiotnlser and decay heat tube banks. The secondary Temperature limits are necessary to avoid excessive economiser, evaporator and primary superheater are gas side oxidation of components for an economic made from 9% Cr 1% Mo steel while the upper tube boiler design. It is also necessary to use alloy banks are made of 316 H austenitic stainless steel. steels that are compatible with other operational factors. The following metal temperature limits are The transition from one tube material to another is employed with associated metal loss and oxide growth provided by welded joints at interbank locations. A for 30 years operation exposed to AGR gas. 5% Cr \% Mo tube Insert is used between the 1% Cr and 9% Cr tubes and an Inconel 600 insert is used Material Temperature Metal loss Oxide growth between the 9% Cr tubes and the 316 H tubes. All tube welds are subject to heat treatment. Carbon steel 350"C 0 .17 mm 0.29 mm (0.1% Si min) TABLE 1 Carbon steel 370°C 0.29 mm 0.38 mm BASIC DIMENSIONS AND MATERIALS (0.2% Si min) Component Dimensions Material 1% Cr i% Mo 370°C 0.29 mm 0.38 mm (0.2% Si min) Reheater tubes 38 o.d. x 316 H SS 4 thick 9% Cr 1% Mo 550"C 0.51 mm 0.81 mm Reheater tailpipes 51 o.d. x 316 H SS (0.62! Si min) 4 thick Reheater platens 36/unit 316 SS 700°C 0.33 mm 0.28 mm (staggered) Reheater slings 4/unit Nimonic 80A For structural components interface joints at Reheater casing Stiffened 316 H SS temperatures above 250°C are fully seal welded where plate 1.5 m possible. Where gas tightness cannot be guaranteed, long prescribed gaps are adopted beween faces of weld Reheater headers 368 o.d. x 316 H SS connections to accommodate oxide growth. Bolted 30 arrangements are designed to cater for oxide jacking Secondary super- 36 o.d. x 316 H SS strains and metal loss at thread forms to prevent heater tubes 4 thick disengagement. Secondary super- 38 o.d. x 316 H SS heater tailpipe Water/steam side corrosion Secondary super- 44/unit 316 H SS heater platens (staggered) Allowances for general water/steam side corrosion Transition tubes 36 o.d. x 4 Inconel 600 are included in the tube wall thickness assessment Primary super- 28 o.d. x 9% Cr 1% Mo N4T for the respective materials of construction. In heater tubes 3.5 thick addition allowances to cater for periodic chemical Primary super- 44/unit cleaning of the boiler units due to accumulated heater platens (In-line) oxides result in the following totals: Evaporator tubes 28 o.d. x 9% Cr 1% Mo N4T 3.5 Carbon and 1% Cr i% Mo steels 0.61 mm Evaporator platens 44/unit (in-line) 9% Cr 1% Mo steel 0.89 mm Secondary economiser 28 o.d. x 9% Cr 1% Mo N4T tubes 0.35 316 H SS 0.18 mm Secondary economiser 44/unit 9% Cr 1% Mo N4T platens (In-line) Stress corrosion Transition tubes 28 o.d. x 5% Cr 1% Mo T c J. J Because of metal temperature limits due to gas side Primary economiser 25.4 o.d. x 1% Cr i% Mo oxidation, it Is necessary to use aus-tenitlc stain- tubes 4 (finned) less steel tube for the second stage superheater and Primary economiser 44/unit (MS fins) reheater tube banks. This material is subject to platens (staggered) stress corrosion in aggressive chemical environ- Decay heat boiler 25.4 o.d. x 1% Cr iZ Mo ments . tubes 4 (finned) R737-PAPER58( 5)HB R737-PAPER58( 6)HB TABLE I (cont'd) Should a boiler unit tube leak occur (detected by water ingress to reactor gas) the associated Component Dimensions Material quadrant would be shutdown, isolated and dried out. When operationally convenient the reactor would be Decay heat boiler 44/unlt (MS fins) shut down and depressurised and the flanged end platens plate connections at Eeed headers removed to gain Boiler unit casing Stiffened 316 H SS access to the orifice holders. The affected tube plate 12 m circuit would be identified using CO2 analysers long and then isolated by fitting blanks in place of the Economiser tailpipes 25.4 o.d. x 1% Cr \% Mo control orifices. At the superheater outlet 4 x 44 off penetration the affected tube circuit would be Feedtube restrictors 22 o.d. x 3 1% Cr \% Mo- isolated by cutting the associated external x 44 off superheater tailpipe and welding on domed ends. Decay heat inlet 38 o.d. x 4 1% Cr \% Mo Access to blank off tubes In the reheater is made tails x 12 off possible by end flange plates on each penetration. Decay heat outlet 51 o.d. x U Cr \% Mo A purpose made, remote-ly operated tube plugging 4.5 x 12 off machine can be inserted into the penetration to the Decay heat 76 o.d. x 8 1% Cr \% Mo tailpipes headers to plug weld the nozzle serving manifolds x 4 off the affected platen. Boiler support 2/unit Carbon steel beams 2.3 Fabrication Boiler support 4 sets/beam 2i% Cr U Mo slings The tube platens are supported by means of a welded spacer system of the same material as the tubing. 2.2 Inspection/repairs In preference to manual welding an automated welding process is employed for the majority of the boiler To assist inspection/repair of the steam generators and reheater spacer welds using robot welding heads. under man access conditions permanent ladders and As a consequence the spacers are arranged longitud- platforms are provided in the reactor annulus and to inally with single spacers at tube centreline where the boiler and reheater casings. Doors are fitted access for the welding heads is permitted by the in the main boiler unit casing at two levels to tube pitch and tangential spacers at close pitch permit access to superheater tailpipes and to the 9% configurations. Cr/316 H stainless steel transition joints and associated thermocouples. Viewing panels are 2.4 Vibration available also for inspection of the 9% Cr/1% Cr transition joints and other components. Access Tube plates are supported at two locations by hanger doors are fitted in the angular ring cheese plates, bars pinned to welded spacer attchment. To prevent platforms and quadrant division plates for passage excessive gas flow induced vibration, tube spacers from one quadrant area to another. are arranged so that natural trequencies are in excess of generated frequencies. For remote inspection additional standpipes are The fundamental frequency of acoustic resonance located at pilecap level to enable TV cameras to be within the boiler casings is determined by the unit introduced into the reactor vessel annulus. dimensions. In order to Increase the acoustic Identifiers are provided on the boiler units to frequency of the cavity and provide separation of establish locations and for record purposes. TV the acoustic frequencies and the flow induced camera access to the annulus below the main gas seal frequencies and hence avoid coupling, a full length level is arranged via standpipe penetrations in the baffle is incorporated in the tube banks at .mid annulus floor. Access route provisions incorpora- position, thus splitting the boiler into two half ting guide tubes and tundlshes allows inspection of units. the following items: 2.5 Special requirements (i) reheater headers, tailpipes, casings and sling supports 2.5.1 Poat-trip cooling (ii) main unit casing (ill) 9% Cr/316 transition joints The boiler system performs a major role in the (iv) seismic restraints safety of the reactor by virtue of its heat removal (v) main gas seal functions. This role is enhanced by the provision (vi) internal cooling pipework of a separate and independent decay heat removal (vii) main unit support system boiler that is accommodated by an additional tube (viii) economiser and decay heat tailpipes bank of finned tube platens located below the (ix) quadrant division plates. primary economiser and contained within the main boiler unit casing. In each quadrant the three Camera routes are also used to permit access to decay heat tube banks are Integrally connected at corrosion specimen containers attached to the main inlet and outlet by tailpipes that are formed into boiler casings. These containers are installed as tube bundles and pass through two penetrations to part of the component oxidation monitoring system. connect to external pipework. .The associated feed system is independent of the main supply and draws It is a requirement that isolation of a single main feed water from reserve tanks. The discharge from boiler tube element Is possible by external tube the decay heat tube banks passes through a pressure blanking. This is made possible by single feed tail control valve (35 bar setting) to a flash vessel and tube connections to each flow circuit of which there hence to the dump condenser. are two per element, and by a single steam tube connection at exit to the element. The feed tail The prime safety function of the system is to tubes are welded to tube plates located in external provide decay heat removal for pressurised reactor water headers that are fitted with bolted flanged faults and is capable of continued long term plates giving hand access to tube ends when removed. operation. It is designed to be aseismic and to The steam tailpipes however, are welded to internal cool a pressurised shutdown reactor to approximately nozzles of a tube plate located at the inboard end 100°C gas temperature. As an added function the of a penetration sheath tube. The 3team circuits system is used to maintain reactor gas inlet are continued through the tubeplate to external temperature within limits during reactor start-up or tail-pipes which pass through the sheath tube to controlled shutdown, or following a pressurised externally mounted superheater headers. fault. It is also employed during commissioning tests to control reactor gas temperatures. R737-PAPER58( 7)HB R7 37-PAPER58( 8)HB 2.5.2 Seismic reactor and boiler plant a series of Plant Operating Conditions has been identified, together with The stations are required to withstand the effects associated frequencies over reactor life, and of an earthquake having a prescribed maximum free subsequently assessed for the design substantiation field ground motion considered applicable to UK of boiler components. Table 2 gives a simplified sites, with a peak acceleration of 0.25 g. Dynamic summary of POCs and frequencies. analysis of the complete nuclear island has enabled aseismic design conditions to be derived for The POCs are classified as Normal, Frequent, Infrequent or individual boiler components. These conditions are Limiting events; test POCs are also identified. in the form of floor response spectra and static coefficients. Normal POCs occur during the course of planned operation of the reactor and include start-up, power Initial design has been based on loadings defined by operation over the load range, boiler quadrant shut- static coefficients derived from a design base down and start-up, refuelling activities and earthquake having a peak ground acceleration of controlled reactor shutdown operations. 0.125 g. These seismic loadings are combined with operational loads and the resulting stresses are Frequent POCs include reactor trips with variants of required to be within the elastic limits of the post-trip cooling, turbine trips, quadrant trips and materials of construction. This approach is other plant faults expected to occur several times strictly applicable only to components having a during reactor life. lowest natural frequency of 33 Hz or greater. As major components in the steam generators have lower Infrequent POCs are expected to occur once or less natural frequencies a dynamic response of a complete during reactor life and component design must assume representation of the twelve boiler and reheater a once per lifetime basis. Included are major units together with annular ring, supports and boiler leak, steam and feed pipework failure, minor restraining structures has been performed with safe reactor depressurisation faults and quadrant trip shutdown earthquake inputs. The resultant loads and protection failures. displacements are applied to the individual components in combination with operational Limiting POCs are events which are not expected to conditions. occur but are included in the design otherwise the consequences could include release of significant Restraining structures include seismic restraint radioactivity. It is accepted that these POCs may brackets located at two levels of the boiler casing. require extensive remedial action or even write-off A /5th scale model of a boiler unit has been of plant. Included in these POCs are major reactor dynamically tested to substantiate the theoretical depressurisation, hypothetical failure of a boiler modelling. For external pipework and associated half unit feed tube plate, reactor trip from full equipment conventional dynamic analysis methods have power with minimum post-trip cooling by one main been used to establish positions of seismically boiler supplied with emergency boiler feed or by two qualified restraints and snubbers necessary to decay heat boilers in service. accommodate the safe shutdown earthquake. Test POCs include boiler testing during the combined Shaker tests have been carried out on valves and engineering tests and power phase commissioning of actuator assemblies in operating mode to demonstrate the station. seismic capability. Control and instrument cubicles have been similarly tested. TABLE 2 2.5.3 Safety SUMMARY OF NORMAL AND FREQUENT POCs It is a requirement for reactor safety that in the Plant Operating Conditions Frequency unlikely event of a failure of a gas pressure con- tainment component, the rate of depressurlsation of Reactor temperature raising to standby 35 the reactor is limited to an acceptable value. Consequently, all boiler steam and feed penetrations Reactor start-ups with three and four 300 are fitted with gas flow restrictors designed to boilers limit the free flow area of Q.006 m2 for any Cold turbine start-ups to 100% load 100 single weld failure. The restraints for the major Hot turbine start-ups to 100% load 200 penetrations take the form of long seamless pipes Power cycling 100%-80%-100% load 5400 extending internally from the outboard end of the Power cycling 100%-60%-100% load 2400 penetrations to low stress regions of the liner Power cycling 100%-40%-100% load 650 shutter tubes, (see Fig 7). Refuelling cycles at loads from 30% 200 to 100% load Single boiler planned shutdowns 150 For the superheater tailpipe bundles and the minor Single boiler trips 150 penetrations, restrictor plates are used as Controlled shutdown from three or four 150 indicated. Externally, secondary retention flanges quadrant operation to hot reactor standby are provided at the end forgings and are tied back Reactor trip from operating power level 60 to the reactor vessel to an anchor plate. Turbine trip - reactor trip 60 Loss of feed trip - turbine trip - 30 In order to limit external radiation to an reactor trip acceptable level (a maximum of 20 ntrem/hour during refuelling), radiological shielding is provided at 3.2 Operating procedures the outer end of the reheater and superheater penetrations and on the inner end of the reheater Operation of the reactor at power requires the use headers. The feed inlet and decay heat boilers of three or four boiler quadrants in service penetrations are similarly shielded. Internal producing steam conditions scheduled as appropriate shields are fitted to the instrument penetrations. for the turbine load. The station auto control The layout of the external shielding is arranged in system ensures that the boilers are operated over sections to permit examination of penetration gas the control range within the defined constraints of pressure retaining welds during station life. the boiler plant, in particular tube metal temperatures and steam superheat at the 9 Cr 1 Mo/ 3 PLANT OPERATING CONDITIONS TP 316 transition jonts. 3.1 General For the start-up of the boilers a 'dry start' technique is employed with feed admitted to From a study of the operational envelope for the initially empty boilers which are uniformly at R737-PAPER58( 9)HB R737-PAPER58( 10)HB approximately 300°C. Low load steaming of the histogram is established to envelope the duty cycles boilers allows the generation of steam at 360°C/ on the plant including loadings from: 90 bar a, which is controlled by means of steam pressure reducing/desuperheating valves at entry to (I) pressure differentials between reactor gas and the boiler start-vip vessels, see Fig 3. These boiler fluid steam conditions are suitable for starting a cold (ii) structural loadings arising from constraint of turbine. For a hot shutdown turbine the boiler differential thermal expansion of tubing steam conditions are raised to 460°C/140 bar a within the platen before admission to the turbine takes place. (iii) local thermal stresses in the welded spacer to tube connections Over the scheduled load range the high pressure (iv) steady and fluctuating aerodynamic loadings steam conditions from the boiler rise from the hot (v) self weight and seismic loadings start conditions to 538°C/160 bar a. Power cycling (vi) system loadings from connected tailpipes or from the 100% MCR load to reduced outputs takes structural members accommodating gross thermal place under control of the station auto control expansions of the boiler system and supports system to match reactor/boiler operation to within the reactor vessel. variations in turbine load demand. In undertaking the component design analyses due The station auto control system is based on a gas attention is given to the load combinations and forcing/steam pressure governing strategy to match appropriate conditions of design. The load turbine load demand. As part of the overall control combinations comprise the parameters associated with system the boiler half unit feed valves are the plant operating condition either alone, or, regulated to adjust feed flow and control the degree where relevant, in combination with an Internal or of steam superheat at the 9 Cr 1 Mo/316 transition external hazard. joints, or to control the decay heat boiler outlet gas temperature at loads below 50%. In addition the The boiler system performs a major role in the average differential pressure across the half unit safety of the reactor by virtue of its heat removal feed valves is maintained at a constant value over functions and therefore safety class 1 designation the load range by regulation of the boiler feed pump is appropriate to certain components. These speed. components are the main boiler reheater and decay heat boiler pressure parts and their respective In the event of a reactor trip or turbine trip the support provisions, penetration assemblies and high pressure boiler steam is diverted from the secondary containment devices. The loading turbine, to a low pressure vent system which allows combinations on these items are analysed to provide discharge of steam to atmosphere via a set of steam design substantiation against the following modes of pressure reducing valves and flash vessels provided failure where applicable: with silencers. These control valves also serve to reduce boiler pressure at a predetermined race down (i) ductile or creep rupture to 80 bar a. This approach allows the boilers to (ii) instantaneous or creep Instability continue to remove heat from the reactor gas circuit (ill) excessive deformation after the trip, whilst ensuring that the boiler (iv) incremental collapse or ratcheting platen stresses remain acceptable and also permit- (v) creep and fatigue Including creep/fatigue ting any subsequent use of low pressure emergency interaction. feed pumps if this is necessary. In parallel with post-trip operation of the main boilers the decay No single design code is available which meets all heat boilers are brought into service automatically requirements for substantiation against the above after a reactor trip generating steam at 35 bar a modes of failure. It has therefore been necessary and providing the longer term heat removal system. to develop a substantiation route based on the principles of code BS 5500 and ASME Code Case N47- The above start-up and shutdown procedures described 21, together with concepts of limit load and in outline are all based on the successful exper- reference stress techniques as applicable using ISO ience gained with the boilers at Hinkley Point B and material data and corresponding UK data for 9 Cr Hunterston B Power Stations. 1 Mo and 1 Cr 0.5 Mo steels. The above approach allows for primary stress assessment, shakedown 3.3 Analysis of POCs assessment and creep-fatigue damage assessment. Fatigue strength reduction factors for specific The POCs have been studied extensively utilising geometrical configurations e.g. spacer tube welds, computer models for the complete station as well as have been derived experimentally in order to relate specific boiler models to establish thermal maximum stresses from analysis to peak stresses for performance data during steady-state and transient fatigue assessment. operations. Particular development of the boiler models has allowed the most critical tube temper- Cumulative damage evaluation of components ature profiles to be identified during rapid recognises the contribution of low cycle fatigue due transients such as following reactor trip. to plant cycling from the POCs, high cycle fatigue due to aerodynamically induced vibration where Validation of the theoretical boiler models has been appropriate, and creep damage to components demonstrated by a programme of laboratory testing on operating at the higher temperature levels. The representative platen configurations covering both result of this substantiation work demonstrates that steady-state and transient thermal performance. the components can accommodate the duties imposed by the plant operating conditions over the reactor Results from the POC calculations provide the life. definitive thermal input data for the structural analysis models used in the design assessment. 4 PERFORMANCE ASPECTS Steady-state temperature profiles throughout the 4.1 General boiler are shown In Fig 8, whilst typical temper- ature variations with time following a reactor trip Optimisation of the operating parameters has been are presented In Fig 9. based on total plant model studies covering reactor, boiler and turbine performance. Layout of the 3.4 Design assessment reactor circuit and boiler arrangements within the pre—stressed concrete pressure vessel provides for a From a study of the POCs and associated frequencies, downwards gas flow through the boiler units, with loading histograms have been compiled for the carbon dioxide gas conditions at reheater inlet of various components in the boiler systems. The approximately 615°C/41 bar a. These provisions R737-PAPER58( 11)HB R737-PAPER58( 12)HB allow the design of once through main boilers for Development work in manufacturing and service per- generation of high pressure steam at 538°C/ formance aspects of restrictor assemblies has been 160 bar a, together with reheat steam at 538°C/ carried out and also erosion corrosion studies. In 39 bar a, to suit 660 MW(e) turbo-alternator the case of the decay heat boilers, stabilising machines as developed in the UK for conventional requirements are satisfied by the inclusion of fossil fuelled power stations. Operating parameters orifice holes in the distribution manifolds at the are summarised in Table 3. bottom of the boiler units. TABLE 3 4.4 Tube bank vibration OPERATING PARAMETERS AT 100% CMR LOAD Boiler equipment in the reactor vessel is designed to resist damage from gas flow or noise induced Gas flow/reactor 4203 kg/s vibrations. Component design is suitable for Gas pressure 41 bar a operation at gas flows up to 375 kg/s per unit which Gas temperatures: assumes utilisation of circulator output margin Inlet to reheater 615°C beyond the best estimate operating point. For the Inlet to HP boiler 573°C short term transient associated with transfer from Outlet from HP boiler 290°C four quadrant to three quadrant operation gas flows HP steam flow/reactor 500 kg/s up to 425 kg/s per unit have been considered. HP steam outlet 540°C/166 bar a HP feed temperature 156°C Tube platens are supported at two locations by means Hot RH steam after 538°C/41 bar a of a system of welded hanger/spacer attachments. attemperator Spacers are also provided at the mid-span of the Cold RH steam 342°C/43 bar a platen which, together with the supports, provides sufficient restraint to prevent excessive vibration. 4.2 Tube bank heat transfer and gas mixing In addition, welded spacers are provided on a characteristics staggered arrangement between the spans formed by the support and centre restraint system, and also 4.2.1 From the definition of overall thermal duty local to the bends at the outer ends of the platen and interface data to suit turbine and to give further restraint against vibration. reactor operation, the gas 3ide and water steam side boundary parameters are estab- The fundamental frequency of acoustic resonance lished for the steam generators. Sizing within the boiler casings is determined from the calculations for the installed heat transfer unit dimensions. A full height baffle is incorpor- areas are based on heat transfer correlations ated in the tube bands at the mid-position, in order established from laboratory tests on the to increase the acoustic frequency of the cavity and selected tube pitching configurations and so prevent coupling between acoustic and flow- published data. In addition development work induced frequencies. has been undertaken to confirm gas side heat transfer correlations for tube part rows and The welded spacer support system for the tube banks tailpipe geometries. ensures natural frequency values in excess of the frequencies which could be generated by flow-induced Performance tests at Hinkley Point B and phenomena. Furthermore, the design of the support Hunterston B Power Stations have also been system includes dynamic effects due to fluctuating analysed and shown to give overall confirm- aerodynamic loadings which are treated as ation of the correlations used in the current distributed loads on the platens both in-plane and steam generator design. out-of-plane. 4.2.2 Laboratory tests on the tube bank geometries The general arrangement of boiler tubing and support have been carried out to establish the gas systems is similar to the Hinkley Point B/Hunterston mixing characteristics. Particularly good B design, which has been demonstrated to have satis- mixing has been demonstrated for the 9 Cr factory vibration characteristics from laboratory 1 Mo in-line tube pitch arrangement. The and site tests together with service operation. mixing data is important for the assessment Hence there is confidence that no fundamentally new of effects of blanked flow paths. Appro- vibration problems will occur. priate boiler models are used to determine any required changes in tube orifice sizes, In order to demonstrate that platen support and tube whilst ensuring that platen temperature spacer design is adequate to prevent damage, vibra- constraint values are still satisfied. tion tests at gas flows in excess of lOOX CMR are being carried out in the laboratory on represent- 4.3 Boiler water flow stability ative platen and tailpipe configurations. Dynamic models have been developed to predict the On completion of construction and during the commis- threshold parameter values associated with the onset sioning phase, vibration testing will be carried out of instability, and to define the necessary on the first reactors at Heysham II and Torness as stabilising orifice pressure drops at inlet to each follows: feed tube. This analysis work has also been valid- ated by laboratory tests on full scale electrically (i) determination of natural frequency, mode and heated rigs. amplitudes of vibration of components in the reactor Stability beween the parallel flow paths is assured (ii) determination of response of the gas within by the provision of orifice assemblies at the tube the boiler geometries plate of each economiser penetration, which are also (iii) combined system tests for which strain gauges, accessible in service when the plant is shut down. accelerometers, and pressure transducers will In the event of a tube leak, access is therefore be monitored during the tests and signals possible to permit blanking at the tube plate and to recorded and analysed. The tests will be re—orifice neighbouring flow paths as required. designed to simulate as far as possible, or allow assessment of the relevant operating For stability between half boiler units, in addition conditions applicable to the boiler to the restrictor tube and orifices, the half unit components, including maximum flow feed control valves also serve to ensure stable conditions. distribution. To minimise post-trip instability between half units the post-trip sequence equipment High temperature vibration monitoring equipment is will initiate closure of the half unit control provided to monitor vibration responses during valves to their minimum opening. operation at high power. R737-PAPER58( 13)HB R737-PAPER58( 14)HB 241 5 DEVELOPMENT Extensive development work has been undertaken in support of the design of the steam generators and a schedule of topics covered is given below. Some of these will be seen to have application not only to gas cooled nuclear steam generators but also to boiler designs in general. On-going work requiring long term data includes mechanical testing and oxidation/corrosion studies. 5.1 Fabrication development Spacer-tube geometries. Transition joints. Tube plates. Casings. Penetrations. Non-destructive examination techniques. Heat treatment gas composition. 5.2 Performance Temperature measurement at transition joints. Heat transfer at spacer-tube connections. Gas flow distributions and mixing. Seal leakage. Tube inlet flow measurement. 5.3 Tribology Fretting and wear tests on selected components. 5.4 Mechanical testing Tensile tests, creep tests, ambient bursting tests on tube butt welded and spacer attachment specimens. Transition joint creep tests. 5.5 Noise vibration, fatigue High cycle fatigue testing spacer tube assemblies. Natural frequencies and damping studies. Acoustic tests in pressurised facility. Flow induced vibration tests in CO2 pressurised rig on platens and tailpipe configurations. 5.6 Oxidation and corrosion Gas side oxidation. Erosion-corrosion studies. Stress corrosion, corrosion fatigue of transition joints. R737-PAPER58( 15)HB GENERAL LAYOUT OF STEAM GENERATORS FIG 1 WITHIN REACTOR 113 SECTION THROUGH ANWLU5 SECTION THROUGH REHEATER^ MAIN UNIT CROSS-SECTION OF STEAM GENERATOR UNIT FIG 2 ATTEMPERATOR : REHEAT OTHER BYPASS REHEATERS REHEATER STEAMo ,-&-& L.P. VENT FLASH VESSEL —tS}—1*|—d TURBINE WATER T STOP VALVE -iXj 1*»—•- «>TEAM SPRAY WATER OTHER BOILERS HP L.R WATER I TURBINES PAIRED START UP VESSEL DEAERATOR BOILER CONDENSER •a f o POST TRIP FEED J -©- START UP/STANDBY MAIN BOILER HALF - EMER6tNCY START UP OTHER J^ PUMPS UNIT — FEED VALVES — FEED BOILERS MAIN FEED PUMP MAIN FEED DECAY HEAT DECAY HEAT FLASH VESSEL RESERVE BOILER FEED WATER •A { TANK CONDENSE1 R •&• DECAY HEAT FEED PUMPS AIR OO COOLER CIRCULATORS OO CHECK PLATES OlACRID SUPPORT DUMMY PENETRATION SOILER MAIN SUPPORT BEAM ARRANGEMENT OF MAIN BOILER SUPPORT BEAM FIG 4 AND SLINGS GAS SEAL TU9ES MAIN BOILER UNITS ANNULAR RING AND GAS SEAL FIG 5 RETENTION FI_*.MSL. - END CAP RESTRICTOH. —:THERMOCOUPLE. PENETRATION t OUTCD PQCSTOEMIN6 TIE BOLT ANCMOD FLANGE. FEED THERMOCOUPLE AND INSTRUMENT PENETRATIONS FIG 6 REHEATER PENETRATION SUPERHEATER PENETRATION STEAM PENETRATIONS FIG 7 XI? STEADY-STATE TEMPERATURE PROFILES FIG 8 © R.H. INLET &flSTEMPERfffURE © H.P. INLET (&flS TEMPERATURE (3) H. P. OUTLET STEW-1 TEriP. © M.P. OUTLET © STEflM "© WfflTER BOU^DfiRy REACTOR TRIP TEMPERATURES FIG 9 No. 12 Babcock MONITORING AND PERFORMANCE ANALYSIS OF AGR BOILERS XA0055821 DURING COMMISSIONING AND POWER RAISING M.EL-NAGDY AND R.M. HARRISON NUCLEAR ENGINEERING DEPARTMENT, BABCOCK POWER LTD., U.K. SUMMARY The installed boiler plant, for two 1300 MW AGR stations, is comprehensively instrumented for boiler control, performance assessment and component monitoring to ensure the integrity and safe operation of the plant during normal and faulty operating conditions. Plant instrumentation and computer systems installed at site for vibration analysis during the engineering runs and data acquisition during the power raising stage have been described. The results, from early rig investigations and the vibration testing during the unfuelled engineering runs, indicate that the behaviour of the plant within the practical range of operating conditions is free from vibration problems. Also the analysis of the steady state thermal and hydrodynamic behaviour of the boiler plant during the power raising phase confirms the methods and computer models used for the boiler design. 1. INTRODUCTION The boiler plant for two recently commissioned U.K. AGR stations (Hartlepool and Heysham) consists of 4 boilers per reactor. Each of these boilers incorporates 2 boiler units (pods) which are separated on the primary (gas) side and connected on the feed side. The boiler section is of the high pressure once-through type with the reheater section above it. These heating surfaces are formed from helically coiled tubes composed of three materials; mild steel, 9% Cr & 1% Mo and 316 stainless steel in the HP section and 316 stainless steel in the reheating surface. They are arranged (Fig. 1) with all the boiler penetrations at the top (passing through a concrete head), thus enabling the withdrawal of the boiler units for inspection and maintenance. The boiler heating surfaces are composed of tubes with equal length and connected to common headers at the secondary inlet and outlet. For uniform primary temperature distribution at the boiler inlet and equal primary and secondary flow rates, the axial temperature distribution and thermal duty are the same for all the tubes. The tube C0£ gas side and steam side corrosion, together with the tube material strength properties have imposed upper limits on the temperatures of the tube walls for the three materials used and a lower limit on the temperatures of the steam in the austenitic section of the boiler. To achieve optimum power without violating these temperature constraints, careful scheduling of the boiler's terminal parameters is essential. Moreover, due to manufacturing and assembly tolerances, the boiler behaviour Babcock cannot be adequately represented using a single tube measurement and the 2D characteristics of the boiler must be investigated to fully assess its performance. The boiler control requirements together with the need to understand the boiler behaviour have led to a comprehensive measurement and instrumentation scheme (Fig. 2) designed to monitor the boiler constraints and measure the terminal parameters during the various operating conditions of the plant. The instruments provide primary data of the terminal parameters for all the boilers and additional metal, gas and steam temperatures at various horizontal levels of a specially instrumented boiler unit which is treated as a thermal analogue of the other units. The measurements from the instrumented pod are used to interpret and extrapolate to the conditions of those uninstrumented units. Also the other pod of the boiler containing the specially instrumented unit, is fitted with strain gauges and accelerometers for the measurement of the vibration response of the boilers particularly during the Unfuelled Engineering Runs (UFER). Although various components had been tested individually in rigs, the Unf uelled Engineering Runs were the first and only opportunity to see how the components would behave assembled and in conditions which were dynamically very similar to those which would be experienced during power generation. The UFER programme of testing was carried out on the Hartlepool Unit MA7 and the Heysham I Unit MA23 during 1980/81 and on the Heysham I unit MA31 during early 1983. This programme was instigated in order to confirm the integrity of the boiler components in the reactor gas circuit with respect to vibration due to circulator noise, gas flow or mechanical transmission from other affected components. It would also serve to establish operating margins for power generation over the station lifetime. The boiler testing programme during power raising which started in 1983 is principally aimed at ensuring that operating and design limits are not transgressed, providing supplementary data necessary to understand boiler behaviour and gaining operating experience of the plant. To achieve these objectives two types of testing are carried out; steady state and dynamic tests. The latter type of tests involving the whole of the plant rather than simply the boilers themselves, is for monitoring performance and closely associated control parameters likely to approach operating limits. The steady state testing which includes symmetric and asymmetric operations of the plant is designed to check and if necessary, amend the operating schedules and to assess the boilers operating margins by investigating the sensitivity of the constraints to the boiler terminal conditions and their spread due to systematic and random factors. For symmetrical operations, the terminal conditions of all the units are matched up with the instrumented unit and for asymmetric tests, the instrumented pod is selected to be the most affected by the prevailing conditions. Babcock Measurements, during the testing and power raising programmes, have been continuously collected by the station data processing system, the commissioning data logger and temporary recorders and data loggers. The thermal and hydrodynamic measurements after suitable corrections, are subsequently used to assist in understanding the boiler behaviour and assess the operating margins. Collected data are also employed to validate the computer model predictions of the plant performance. 2. VIBRATION MONITORING The boiler components can be divided into two categories for the purposes of vibration testing and monitoring, according to the dominant excitation mechanism. Plate-like structures, for instance baffles, are excited acoustically whilst the vibration of tubes is largely due to gas flow effects such as vortex - shedding and turbulent buffeting. Theoretical analysis of the vibrational behaviour of the HP and reheater tube banks indicated a large number of modes with some of them close in frequency to predicted excitation frequencies. Thus the possibility of large-amplitude vibration occurring due to positive-feedback effects was considered. It was also postulated that, in an extreme case, the coils themselves, rather than individual tube spans could pulsate (whole boiler modes) with large amplitudes, perhaps assisted by an acoustic standing wave. 2.1 Preliminary Testing Expected stress levels during UFER were established before commissioning by a combination of testing and analysis. For acoustically dominated components, the instrumented pod (unit MA5) was excited by a loudspeaker. The stresses obtained were extrapolated to the expected SPL during commissioning. Also, a test bank of 4 helical multistart coils each 40 rows deep was constructed to investigate the behaviour of the HP bank. This half scale bank, with the tube span length of the outer coil as that of the outer coil of the boiler, was extensively instrumented with accelerometers on both tubes and strap supports and installed inside a pressurised CO2 loop. Contract conditions of gas density and velocity could be achieved or even exceeded in the rig. The rig provided data on bank response as functions of gas flow, density and temperature. It was also operated continuously over two 1000 hour periods in order to provide information on fretting wear at the tube/strap interface. The level of response, for the conditions prevailing in the AGR boilers was low (accelerometer reading less than lg with calculated amplitudes less than 0.1mm). Some accelerometer response curves (plotted against gas velocity) showed peaks due to narrow-band excitation, but none was high. The peak to baseline response ratio (Q) was only around 2 to 5, indicating high damping. The accelerometer results were used to calculate the expected stresses in the boiler and reheater banks. Babcock 2.2 Unfuelled Engineering Runs Testing 2.2.1 Instrumentation The Hartlepool boiler unit MA7 was instrumented with 300 strain gauges fitted to around 20 boiler components. Also, a limited number of microphones at the top and bottom of the boiler were used for the detection of acoustic standing waves and to confirm whether a component response was due to acoustic excitation. The strain gauge signals were fed via a multiplexer unit to a PDP11 computer connected to a Genrad signal analyser. Fast-Fourier-Transformation was performed by the analyser on the signals to provide r.m.s. and Power Spectral Density plots. The analyser could also provide the cross-correlation and phase between two selected signals. All transformed signals were stored on hard disk and a 14-track tape recorder was available for recording untransformed signals for more detailed analysis at a later date. Ancillary equipment included a Hewlett-Packard 3582 signal analyser for spot checks and an oscilloscope with storage facility for visual checks on gauge signals. Heysham I unit MA31 was similarly instrumented with 300 accelerometers of which 50 were available to Babcock during UFER. The signals were analysed with a Hewlett- Packard HP85 computer and recorded on 14 channel tape. 2.2.2 UFER Testing Contract conditions of gas velocity, density and acoustic velocity at three planes (boiler gas outlet, mid-HP-surface and superheater gas inlet) were achieved by means of varying the Inlet Guide Vanes (IGV's) position, gas temperature and CO2 mixture (with nitrogen). Testing was carried out at three temperatures 150,246 and 288°C. At each temperature, the gas density was increased in stages and the IGV angle varied in increments of 10°to 2°from fully open (-10°) to almost closed (82°) positions. Testing began with the IGV at an intermediate setting (70°) since the maximum acoustic excitation was expected at the nearly closed position. On the other hand, maximum flow- induced excitation was expected at the fully open position. Initially, the gauges were monitored up to 2000 Hz frequency, but subsequently up to 250 Hz owing to the low level of response above this frequency. Also, to save time and improve resolution, only a selection of 60 'priority' gauges with consistently high readings, was monitored at every IGV setting. The acquisition of data for the Genrad was automatically controlled by programming the PDP11. Babcock At the end of each test, the r.m.s. stresses would be compared with the component expected maximum stresses and fatigue/fretting stress limits. For on-line interpretation and assessment, PSD plots of some gauge readings were printed. All the gauges, however, would be monitored on the overnight conditions and whenever a particular contract condition was reached. Faulty or noisy gauges were often identified by unusually high stress readings. Examination of the PSD plots would confirm the fault/noise. D/C switching or 50 Hz mains pickup due to poor insulation resistance would be observed on these plots. The Heysham unit MA31 with accelerometers on coil 19 was tested in the same way, but with fewer readings (50) to monitor at a time. Also, the settings of the IGV were refined to 0.5 degree intervals. The main aim of these tests was to investigate boiler resonance over very narrow frequency/velocity ranges. 2.2.3 Test Results and Conclusions Stresses measured in all the boiler components were lower than the previously calculated maximum expected values (see Table 1). The stresses of some of the acoustically excited components, such as the top gas baffle and debris screen were close to the expected values owing to the small degree of pessimism involved in calculating these values. Careful analyses of the test results of the coiled tube banks indicate the following: a) There was no sudden and large take-off in response that would indicate the presence of fluidelastic instability. b) The excitation was chiefly due to turbulent buffeting since the dominant frequencies did not change with gas velocity, and stresses were approximately proportional to velocity head, (" V . c) There was evidence of Strouhal type (proportional to velocity) excitation of the tubes at around 30Hz but the contribution to r.m.s. stress was minimal. A value of Strouhal number = 0.12 was derived which agrees well with earlier test work. d) Cross correlation with pressure transducers revealed no significant acoustic excitation. e) There was no coherence between gauges at the top and bottom of the boiler that would indicate a whole boiler mode of vibration. £23 Babcock 3. THERMAL AND HYDRODYNAMIC DATA MONITORING 3.1 Data Acquisition and Processing The various data items, necessary for Babcock to effectively monitor the boiler plant during the power raising of the Hartlepool and Heysham stations are collected and recorded on disks by the CEGB Station Commissioning and Optimization (SCOOP) computer system [1]. This involves the recording of some 200 signals for each station representing the terminal parameters of all the boiler units, the metal and gas temperatures of the heating surfaces of the specially instrumented unit and the helical coils outlet steam temperatures in the superheater headers. The recorded disks are despatched regularly from both sites to the Babcock1s London Offices for processing. Twenty four hours worth of thermal and hydrodynamic data items, on each disk, are read and processed by the LSI 11/2 computer, in London, which is connected to a line printer and a graphics screen attached to a dot printer (fig. 4). The data processing is normally carried out at three stages, the first of which involves the conversion of the recorded signals to engineering units (if not already converted at site), detection of transmission, standard types and out of range errors and the production of selected (initial) plots representing the variation, with time, of the main plant parameters. These initial plots (usually for 24 hours periods) are chosen to enable the following of the main operating conditions of the plant and the specification of the data processing to be carried out at the following stages. A flow chart, of the main processing operations in the first stage, is shown in fig. 3. An error logger as well as the initial plots are produced at the end of this stage. The instrument signals from the stations are scanned at two different rates, a slow scan rate (1 or 10 minutes) and a fast scan rate (3 or 8 seconds). The signals are usually recorded at the low scan rate of 1 or 10 minutes and the fast scan rate is only carried out following the initiation of and during selected plant transients. The slow scan data items are processed further in the second stage which includes the production of a diary containing the boilers terminal parameters and a description of the plant states identified from the interrogation of selected plant parameters such as feed flow, IGV position and gas circulator speed. Each of the main plant operating states is associated with a unique set of values or range of values of these parameters. Additional graphical or tabular outputs may also be obtained at this stage to allow more detailed investigations of the plant conditions to be carried out. The processing of the slow scan signals is shown diagrammatically in Fig. 4. Babcock Following the completion of the initial processing described earlier, the recorded data items are reduced by performing statistical analysis on the signals during the steady state periods of the plant operation. The reduced information, which includes the parameter values at the beginning and end of the steady state, means and standard deviations, is stored permanently on disks for future analyses of the plant performance. Measurements associated with non-steady state operating conditions are also reduced and stored on the permanent disks. The third stage of processing is mainly concerned with the preparation of plant data for use as input data to the computer models developed specially for investigating the thermal performance of the Hartlepool/Heysham boilers. Typical plots, produced in the 1st stage of data processing, are given in Fig. 5 which shows the variation, with time, of the outlet steam temperatures of boilers C and D and the feed flow of boiler C. Fig. 6, showing the reheater inlet and outlet steam temperatures and the bypass valve position during the connection of the reheater, is typical of the plots which can be produced on request at the 2nd stage of processing. 3.2 Components Integrity Monitoring To safeguard the integrity of the boiler components, that are predicted to operate close to their design temperature limits, a close watch is kept on the temperatures of these components during the various power raising stages particularly when the plant is operated for the first time at a new set of conditions. These temperatures are continuously checked and their values compared with the predictions and the limits established prior to the power raising programme. When these temperatures exceed the recommended limits, the plant conditions, whenever possible, are adjusted to satisfy the limits. Also the boiler plant terminal conditions are monitored regularly to compare their values with previous theoretical predictions and to assess the thermal and hydrodynamic performance of the boilers. 4. BOILERS PERFORMANCE ANALYSIS The design of the boilers involves the calculation of the heating surfaces and the optimization of the dimensions/arrangement of the tubes forming these surfaces which are required to perform a specific thermal duty within limited space envelope and specified design margins. The design calculations are frequently performed using computer models representing the geometrical details and thermal and hydrodynamic characteristics of the boilers and associated plant. These characteristics are based on the results of experimental investigations carried out in test rigs, which are usually idealization of the boiler geometry, to derive general correlations economically and within limited time scale. Under the actual plant operating conditions, the characteristics may ISA Babcock vary from those derived experimentally and to gain sufficient confidence in the computer models using the experimentally determined characteristics, it is thus necessary to verify the model predictions using plant data. This data can be obtained from the specially designed tests planned during the power raising programme. Following the completion of the thermal design, investigations of the performance of the boilers under the anticipated operating conditions are carried out to determine the thermal operating cycles for the various boiler components. These cycles are subsequently used for the stress analysis and the mechanical design of the components. Both simple and sophisticated computer programs, including some of those used in the design stage of the boiler, are employed to determine the operating cycles and to establish the plant thermal duty at the various operating conditions. Two main computer models were developed and extensively used by Babcock for the design and performance analyses of the two AGR boiler plants under consideration. The first of these models is a multibank unidimensional representation of the plant while the other model is a two-dimensional representation of the boiler main heating surfaces. The application of these models and comparisons of their predictions with early commissioning plant data are briefly discussed below. 4.1 Multibank Boiler Model Babcock developed a boiler performance computer program [2] capable of carrying out mono-tube performance calculations on several tube banks within a multibank network including once- through steam generators. The primary (heating) fluids which can be handled by the program include carbon dioxide and helium and the secondary fluids are water and steam. The program contains a number of standard heat transfer and pressure loss correlations and can accept user written FORTRAN coding for special analysis of boiler peripherals. It is also structured in such a way that a large amount of information, such as the geometrical details and boiler configuration etc. which are the same for a number of runs, is stored in data banks. These banks together with the loading conditions form the input data to the program. Although the data banks can be edited by the user, he is not permitted to modify the data banks and only authorized users are allowed to add new banks which may be accessed by the program. Initial estimate, for the flows and the pressures and enthalpies at the terminal points of the heating surfaces, is included in the data banks to enable fast convergence to the final solution. Extensive user testing of the program was carried out before its release for general use, and it is hoped to be able to verify the program using plant data from the power raising programme testing. The program has also been successfully used for the assessment of the boiler performance during start-up [3 and 4] of the Hartlepool and Heysham stations. Typical results of this assessment are plotted in Figs. 7 and 8 for Hartlepool and Heysham respectively. The plots show the axial profiles, in the HP and R/H sections of the boiler, of the tube, metal and the primary and secondary fluid temperatures. Babcock 4.2 Two-dimensional Boiler Model A two-dimensional analysis capability has been added to the Babcock multibank computer program [2] to enable the investigation of the asymmetric behaviour of the boiler and allow mixing between the primary streams while the secondary fluid flow inside the parallel tube paths can vary to maintain a common header to header pressure drop. Two mixing calculation options are available in the program, in the first of which the primary flow streams are isolated and equal pressure drop across the bank can be maintained. In the other option (equi-pressure) equalisation of the pressure at every section of the tube bank is carried out, thus allowing a transverse flow between the primary streams to occur. In the longitudinal direction, the tube bank can be subdivided up to 18 sections, and in the transverse direction, the primary fluid is divided into 20 streams. Heat transfer to or from the boundaries of the 2-D banks can be allowed for in the calculation. Also for the graphical presentation of the results, the program accesses a plotting file on which the results from a series of runs for the same configuration can be stored. Early site data during the start up of the Heysham boilers (phase 3) have been analysed using the part load program [5], The temperature profiles calculated for this phase at three selected axial locations in the boiler are shown in Figs. 9-12 for the isolated streams option. The inlet gas temperature profile to the 9% CR section of the boiler and the superheater outlet steam temperature profiles are plotted in Figs. 9 and 10. The metal temperature profiles at the stainless steel/9%Cr and the 9% CR/Mild Steel tube transitions are shown in Figs. 11 and 12 for the same start-up phase. 5. ACKNOWLEDGEMENT The authors would like to thank Babcock Power Ltd. for their permission to publish this work. REFERENCES 1. Main, J. - Specification of Plant Data available on the SCOOP (PDP 11) System at Hartlepool Power Station CEGB, Issue 3, July 1981. 2. Woodison, P.M. and Sodha, M.N. - PNB01001 Nuclear Boiler Performance Program User Manual Babcock Power Ltd., Issue 1, November 1982. 3. Teasdale, E.N. - Hartlepool Commissioning - An Initial Assessment of Steady State Boiler Thermal Performance during start-up using One Dimensional Models, Babcock Power Ltd., NEDR/07/0288, Issue 1, November 1983. Babcock 4. Teasdale, E.N. - Heysham Commissioning - An Initial Assessment of Steady State Boiler Thermal Performance during start-up using One Dimensional Models, Babcock Power Ltd., NEDR/07/0287, Issue 1, November 1983. 5. Roy, R. - Heysham NPS 2-D Analysis of Phase III using the Program PNB010. Babcock Power Ltd. NEDR/07/0296, Issue 1, March 1984. 6. Graham, J.P. and Hill, R.B. - Hartlepool Nuclear Power Station, Boiler Unit Vibration Tests carried out during the Unfuelled Engineering Runs, Preliminary Babcock Report No. (05)/81/14, March 1981. Babcock E Maximum Stresse =s at:- Boiler 150°C 246°C 282°C Component H'pool Heysham H'pool Heysham H'pool Heysham Debris Screen 483 860 570 566 640 676 Gas Seal 24 13 25 12 23 11 Noise Baffle Plate 101 30 65 25 59 22 Feed Inlet Tail Strap 33 77 27 72 23 78 Feed Inlet Tail 105 27 84 102 125 78 Radial Arm 28 55 35 19 16 11 Boiler Casing 15 37 25 14 17 12 S/H Outlet Tails 82 60 107 59 104 66 R/H Inlet Tails 82 26 60 20 75 23 R/H Surface 64 28 71 25 82 26 Top Gas Baffle 219 170 312 140 204 114 R/H Outlet Tails 136 159 107 224 104 96 Spine 33 136 39 13 21 11 GFR Support Strap 90 84 65 52 103 40 Acoustic Angle 77 54 80 23 66 18 C.S. Tubes 100 80 103 62 - 80 S.S. Tubes 61 110 81 84 - 120 9% Cr Finned Tubes 151 194 91 95 - 97 TABLE 1 COMPARISON OF HARTLEPOOL & HEYSHAM UFER VIBRATION TESTS (6) Babcock MAIN FLOW REVERSE FLOW - TOP GAS BAFFLE BOILER GAS -GAS CIRCULATOR NOLSE BAFFLE —— Fig la Gas Flow Diagram Fig lb Pod Boiler Unit V THERMOCOUPLE R/H • • ORIFICE PLATE I x/ — 1 J \7 \/ I SUFV I X -tx}- FIG. 2 PLANT MAIN INSTRUMENTATION 'CEGB / SITE DISK / [SITE TAPEj 1 (Slow Scan 1 V(D4ta) \ Data) \ f PLANT STATES (Diary) \ STATISTICS VlANT DATA / & DATA CHECKING \ PREDICT-/ \ION / \r DATA EDITING \HEADINGS I CONVERSION & TO REDUCTION \ DATA / ENGINEERING UNITS CHECK SIGNALS (Within Range) / PERMANENT DATA DISK \ SELECT SIGNALS FOR PLOTTING Fig. 4 Data Reduction/Storage (LSI 11/2) Fig. 3 1st Stage Processing (Error Logger and Initial Plots) Fig. 6 Additional Plots of the Reheater Outlet Steam Temperatures before Fig. 5 Initial Plots of the Superheater Outlet Steam TemperatureB and Boiler and after the Bypass and the Bypass Valve Position Feed Flow BfiBCOCK POWER LTD BfiBCOCK POWER LTD a a -.•.•> SW (3 33O 3an±aa3cU3i to 33Q) 3aniaa3dw3i O3S/3X) (101J Babcock D PR (H ?Y TEhFt R£ HF H9 F H ft1 R3HT R2 9 0 FIN T p TEKIPE m.RE HP H fi!H R1 Rt •fT m A TUBE )UT 5ICE T IMPERf URE HPIR 3 FtH R] » HT RE + BEccK WR!f 1D TU HPRl hi F;?H Rt •fT R£ OEO.C 450. -r • i s i - 400. 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BS. 90 TUBE LENGTH METRES ---- FJ _„_.__ . -- t (D _ — . 7 ~" a: 2 __ J .3 X _ _ en m r- s £ ::;:" ~\~:~ :;:: ID ss ' in £ £ S § J gg :::: ::;-::::: b( __ - - _ __. _,;:>. i: = _, __. j( j. _ __ _ = ^^ „ o o S L^ a ^ o i/> c L/D CJ) Ui D oo t- r- ID •r n 'ff ^ GAS TEMPERATURE AT 9CR PLAIN/9CR FINNED TRANSITION STEAM TEMPERATURE AT S/H OUTLET FIGURE {0 CM 0) 0] r- ID Ul -t •« \ \ \ * m \ 1 i — en NUT - ^ 10 . \ X 1 m •> ? r CD \ J a a a s 4 <;(_- r- / si & U- (0 aa 38 K t- s a ui - T J m 5 5 5 5 § § CvJ _ i; El 0 u • C S S S S ii S S S S3 S 3aniJ3dW31 ' TUBE OUTSIDE TEMPERATURE 9CR TUBE OUTSIDE TEMPERATURE AT SS/9CR PLAIN TRANSITION FIGURE 11 v\ FINNED/MS TRANSITIO FIGURE 12. No. 13 XA0055822 INTERNATIONAL ATOMIC ENERGY AGENCY INTERNATIONAL WORKING GROUP ON GAS-COOLED REACTORS EXPERIENCE WITH THE COMMISSIONING OF HELICALLY COILED ADVANCED GAS COOLED REACTOR BOILERS Paper Presented to the Specialist's Meeting on Heat Exchanging Components of Gas Cooled Reactors Dusseldorf, Federal Republic of Germany 16 - 19 April 1984 PRESENTED BY: D.B. KETTLE CEGB — GENERATION DEVELOPMENT AND CONSTRUCTION DIVISION ABSTRACT The paper describes aspects of the experience gained during commissioning of the helically coiled pod boilers for an advanced gas-cooled reactor. The boiler geometry is shown to be a factor contributing to gas-side and water-side convection phenomena encountered during commissioning. Detailed information on thermal performance and vibrational response was obtained from commissioning tests on specially instrumented boiler units. Experience with the Commissioning of Helically Coiled Advanced Gas Cooled Reactor Boilers Contents 1. Introduction 2. Water-side convection 3. Post-trip syphon 4. Hot gas convection 5. Thermal performance 6. Vibration response Figures 1. Location of boilers in pressure vessel 2. Boiler steam flow 3. Boiler feed and steam pipework 4. Boiler gas flow BOILER CLOSURE " (CONCRETE FILLED) "• BOILER LINER - GAS INLET DUCT FIG. 1 LOCATION OF BOILERS IN PRESSURE VESSEL Introduction Hartlepool and Heysham are both twin reactor stations with nominally replicated designs of advanced gas cooled reactors. The design of the concrete pressure vessel differs markedly from the earlier designs for Dungeness and Hinkley Point in that the boilers are located in vertical pods within the wall of the vessel (figure 1). The most efficient utilisation of the space available within the boiler pods is achieved with a helical boiler geometry. Each boiler unit incorporates a once-through high pressure boiler and a separate reheater. The boiler unit is suspended within its pod from the closure head, which seats on a flange in the pod liner. This joint is sealed by 3 concentric O-rings, 2 metallic and one rubber. The boiler tubing is arranged in 19 concentric, contra-rotating helices. The innermost helix has 6 tubes on a helix radius of 30cm, and the outermost helix has 24 tubes on a helix radius of 120 cm. Inner row tubes are supported at 4 radial positions, while outer row tubes are supported at 8 radial positions. Three boiler tube materials are used. The reheater and secondary superheater tubing is made from Type 316 austenitic stainless steel. This material is resistant to corrosion in carbon dioxide at elevated temperatures, but, owing to concern about the possibility of stress corrosion in wet steam, a minimum superheat margin is specified. The primary superheater, evaporator and secondary economiser are made from a ferritic 9% chromium, 1% molybdenum steel. This material is resistant to waterside corrosion and can be used at temperatures up to about 520°C. The primary economiser is made from carbon steel, which is suitable for operation at temperatures up to about 350°C. Enhanced heat transfer is obtained by rolled finning of the tubing. The tubing is subsequently heat treated to restore the required material properties. There are 8 boiler pod units per reactor and for control purposes these are connected in pairs on the feed and steam sides. Each boiler pod unit has one feed header feeding 285 tubes via the feed tube plate. Tubes are bifurcated at the superheater outlet, and the tailpipes are then routed to the 4 superheater headers. There are, in addition, 2 reheater inlet and 2 reheater outlet headers. A schematic arrangement is shown in figure 2. Waterside Convection Before power raising commences on the completed reactor, engineering tests are carried out in which the plant is run at conditions closely matching the proposed operational state. During this period, reactor temperatures are regulated by using the gas circulators to input heat, and with boilers operated in a flooded mode to take out heat. 2 1} FEED WATER INLET II SUPERHEATER OUTLET REHEATER INLET :. REHEATER OUTLET HOT GAS INLET FROM REACTOR GAS OUTLET TO REACTOR FIG. 2 BOILER STEAM FLOW Now, it is a feature of the podded boiler design that both the feedwater inlet header and the superheater outlet headers are located at the top of the boiler unit. During the engineering test period, feedwater temperature was typically 100°C and reactor gas temperature 280cC. The density of water decreases by almost 30% as it heats up over this temperature range. The static pressure increase down the downcomer is therefore greater than the decrease up the riser, and for low flow rates the frictional losses can be insufficient to maintain a higher pressure in the inlet header than the outlet header. Where circuits are connected in parallel a potential then exists for flow reversal in some circuits. During the engineering tests, thermocouples attached to the feed water inlet header were observed to increase in temperature on one pod unit per pair, when the feed flow rate was dropped below a certain level. It was inferred that some slight assymetry in the operating conditions of the two pod units had led to the feed water flow going preferentially to one unit, with reverse flow being established in the adjacent unit. Figure 3 illustrates the pipework geometry: the secondary steam header is the point at which the flow from one pod unit can return in reverse flow into the adjacent pod unit. There was some evidence to suggest that recirculation could also occur within a boiler pod unit. This can occur via the primary steam header or between individual tubes connected by a bifurcation piece. The evidence for this behaviour came from an observation that the measured temperature could rise on both of the feed headers associated with a pair of pod units. A series of tests were carried out to establish the minimum flow rate required to avoid recirculation. It was noted during these tests that a hysteresis effect occurred, with the flowrate required to suppress recirculation being greater than the flowrate at which it was initiated. Post-Trip Syphon On 10 April 1983, power raising had progressed on Reactor 1 at Heysham Power Station to an output of 270 MW(Th). By this time the boilers had passed through the boil-back phase of start-up, and were passing steam to the dump system. A reactor trip then occurred, and all shut-down systems operated satisfactorily. However, some 20 minutes after the trip, although reactor temperatures were dropping in a controlled manner, it was noticed that the boiler outlet gas temperatures from pairs of boiler pod units were diverging significantly. Differential temperatures as high as 100°C were measured between adjacent boiler units. It was also observed that the feed header temperature was rising on one pod unit in each quadrant. A *122-5i(PILE CAP LEVEL) '20 \ B1C/SR/8 FIG. 3 BOILER FEED AND STEAM PIPEWORK The explanation for these effects is found in the geometry of the steam pipework (figure 3). The method of boiler operation post trip is that a feed flow rate of 6-8% MCR is maintained from the emergency boiler feed pumps, with the boilers slowly flooding through as reactor circuit temperatures drop. Once the water level reaches the boiler steam outlet header, it falls through the steam pipework to the primary header, and thence to the secondary header, where it combines with flow from the adjacent boiler pod unit. The secondary header is approximately 16 metres below the level of the superheater outlet headers. Once one boiler unit has flooded over, the syphon effect from the column of hot water in the outlet pipework pulls feedwater preferentialy into the unit, starving the adjacent unit. In fact, reverse flow can be induced in the adjacent unit, with water being sucked backwards through the feed header. Analysis of the post-trip transient showed that increasing the pressure loss across the pod feed trim valves could prevent the syphon occurring. Post-trip operation was therefore modified to provide closure of these valves to a pre-set mechanical stop. Subsequent experience has shown this modification to be effective. It was also apparent that the static head of the syphon could be substantially reduced by providing a pipework cross-connection between the boiler primary headers. Consideration is being given to this as an alternative solution to the syphon problem. Hot Gas Convection The advanced gas-cooled reactors are designed for continued operation with one shut-down quadrant. When operating in this condition, relatively high temperatures have been measured on the feed headers of the shut-down boiler units. A shut-down boiler unit can be either wet stored or dry stored. Feed header temperatures were higher for a wet stored boiler unit. The pressure differential across the reactor core ensures that there is a flow of gas at reactor inlet temperature up through a shut-down boiler unit. However, commissioning experience has shown that there is still some ingress of hot gas through the boiler gas inlet duct, which circulates into the dead space below the boiler closure and across the top gas seal (figure 4). This gas heats the boiler support spine, which in turn heats the boiler inlet feed tubing by a mixture of radiation and natural convection. The fluid within the tubing then circulates into the feed header, producing the high temperatures measured on the plant. The temperature cycle to which the feed header is subjected when the boiler is reconnected is therefore more severe than originally anticipated, but can be accommodated by margins in the design. MAIN FLOW REVERSE FLOW TOP GAS BAFFLE BOILER GAS INLET ANNULUS GAS BAFFLE EIGHT GAS BLEED HOLES CENTRAL SUPPORT SPINE ANNULUS REVERSE FLOW BOILER GAS SEAL GAS TO W REACTOR IQV GAS CIRCULATOR GAS CIRCULATOR CO2 FEED FROM OUTLET DUCT SEAL AUXILIARY GAS CIRCUIT FIG. 4 BOILER GAS FLOW J1Y3 Thermal Performance Prior to Hartlepool and Heysham, advanced gas cooled reactor boilers had been of a rectangular section design, built up from a number of identical tube platens. The helical boiler design, however, comprises tubes formed to coiled radii varying between 30 cm and 120cm. A number of radially asymmetric features were identified which could affect the thermal performance of a coiled boiler unit. The most important of these arises because the reheater tube bank has a smaller diameter than the high pressure boiler bank. Gas flowing through the reheater and its surrounding annulus, emerges with a radial temperature and velocity profile which influences the performance of the boiler unit situated below the reheater. Other less important factors taken into consideration are the heat losses to the boiler support spine and casing, and the effect of coil radius on water-side frictional pressure losses. Variations in the heat duty for different coils are accommodated by adjusting the feedwater flow to individual tubes by the selection of different inlet orifices, fitted at the feed inlet tubeplate. To confirm that the sizing of the orifices achieves an acceptable boiler thermal performance, two boiler units were extensively instrumented with thermocouples measuring gas and metal temperatures. To date operating data has only been obtained at relatively low power levels. Tube metal temperatures at the transition between 9% chromium ferritic tubing and type 316 austenitic tubing were lower for small diameter coils adjacent to the central support spine. This is believed to be attributable to greater heat losses to the spine resulting from lower feedwater temperatures at low loads, and there is evidence to show that more uniform temperature distributions are achieved as power is increased. Extrapolation to higher power levels will be carried out using the two-dimensional performance programme described by Mr. Lis. Vibration Response Experimental work on a straight tube cross-inclined tube bank representing the Hartlepool and Heysham boiler geometry showed a susceptibility to flow induced vibration. This result prompted more detailed experimental work on a fully representative coiled tube bank, and led to the fitting of instrumentation to monitor vibration levels on an operating boiler unit. Accelerometers capable of operating at full reactor temperature in a high pressure carbon dioxide atmosphere were fitted extensively on boiler tubes and support straps. In addition, a number of microphones and displacement transducers were fitted at selected locations. 8 Before power raising commenced, tests were carried out at operating gas densities and velocities. Measured vibration levels were low and no restriction on power raising was identified. Monitoring of the vibration instrumentation has continued as power has been increased, and the low level of response observed during the engineering runs has been confirmed. The monitoring exercise will continue until full output is achieved in order to establish whether temperature levels affect the structural response of the boilers. Meanwhile a comprehensive exercise is being undertaken to fully analyse the results being obtained. No. 14 XA0055823 Technology Planning and Research Division Central Electricity Research Laboratories INVESTIGATIONS OF THE GAS-SIDE HEAT TRANSFER AND FLOW CHARACTERISTICS OF STEAM GENERATORS IN AGR STATIONS by J. Lis April 1984 Sfi INVESTIGATIONS OF THE GAS-SIDE HEAT TRANSFER AND FLOW CHARACTERISTICS OF STEAM GENERATORS IN AGR STATIONS J. Lis Central Electricity Research Laboratories, Leatherhead, Surrey, England ABSTRACT This paper describes the experimental and analytical investigations of the gas-side heat transfer and flow characteristics of steam generators in the AGR stations carried out by CERL. The majority of the experimental work on heat transfer and flow characteristics of close-packed tube arrangements in cross-flow of gases is carried out in a pressurised heat exchanger rig. The rig is operated on-line by a dedicated PDP 11/40 computer over the range of Reynolds number lO4 to 3 x 10 . Atmospheric wind tunnels employing either small or large scale models of the specific sections of steam generators are used for a variety of supplementary and development studies. Various measurements techniques and, in particular, LDA and hot wire anemometry employed in these studies are described. The more important aspects of various investigations are illustrated by typical results. In order to ensure the efficient operation and integrity of steam generators under asymmetric boundary conditions a MIX suite of 2- dimensional codes has been developed. The codes calculate the gas and water/steam flow and temperature distributions in each channel of the steam generator taking into account thermal mixing in the gas as it passes through the generator. Application of the MIX codes to the solution of various operational problems is illustrated by typical examples and the continuing exercise of validating the codes against plant operational data is discussed. 1. INTRODUCTION The thermal performance of steam generators in gas cooled reactor stations depends primarily on the gas-side heat transfer and flow characteristics of the generator tubing. There was very little information available in the literature on these characteristics at the start of the nuclear power programme in the UK. In particular, there were no data on the heat transfer and pressure losses in close-pitched arrangements of finned tubes in cross-flow of gases and almost complete lack of understanding of the physical processes in such arrangements. Recognising the importance of accurate information in this field to the efficient and safe operation of the Magnox stations, the CEGB initiated a programme of investigations and set up appropriate test facilities at CERL in 1958. The broad objective of these investigations was to develop analytical methods and provide experimental data on heat transfer and flow characteristics of tube arrangements in cross-flow of gases as required for the assessment of the design, operational performance and control of steam generators in the gas cooled reactor stations. The commissioning and operation of once-through steam generators in AGR stations has brought to light a variety of new thermal problems specific to generator tube geometry in each station. In order to provide rapid and effective solutions to these problems, additional test facilities have been set up and appropriate experimental and analytical studies carried out. The paper describes the major test facilities employed at CERL for the investigation of gas-side heat transfer and flow characteristics of steam generators in the AGR stations. Examples of some more important experimental studies are given and illustrated by typical results. The development of a MIX suite of multitube steam generator models is briefly described and the application of these models to the solution of the design and operational problems discussed. 2. TEST FACILITIES 2.1 Pressurised Gas Heat Exchanger Rig The first heat exchanger rig at CERL (Davies and Lis, 1960) was designed to cover the range of gas-side Reynolds numbers in steam generators of Magnox stations. This range was not wide enough to cover the envisaged operational requirements of the AGR stations and a new, larger and far more flexible test facility was commissioned in 1965. With some modifications and improvements incorporating the recent advances in control and measuring techniques, this basic test facility has been used almost continuously over the past nineteen years for a variety of investigations. A general view and a schematic diagram of the pressurised gas heat exchanger rig are shown in Fig. 1 and 2, respectively. Essentially, the rig consists of a closed loop circular duct incorporating a centrifugal fan, electrical heater bank, test section and the necessary control and measuring instrumentation. The gas is circulated by the fan designed to deliver 2.35 m3/s of C02 at 7 bar and 80°C against a static head of 250 mbar. The fan is driven by a 165 kW dc motor with a thyristor controlled stepless regulation of speed from 30 to 1500 rpm. Gas mass flow rate is measured by a Dall tube with a 1:1.755 throat ratio situated in a 17.25 in bore duct about 25 diameters downstream of the blower. After passing through a supplementary cooler and a 180° elbow incorporating an expansion bellow, the gas flows through a 1:4.5 area ratio diffuser into a 550 kW electric heater where its temperature is raised up to 150°C. The gas temperature at the heater outlet is controlled by a thyristor unit in conjunction with a set of sensing thermocouples placed at the inlet to the test section. The transition from the heater bank to the test section is through a 1:1.6 area ratio diffuser and a 7:1 area ratio contraction. Each diffuser contains a number of gauzes of appropriate mesh. The pressure vessel containing the test section and its upstream and downstream instrumented ducts consists of three cylinders. Access to the test section is obtained by sliding the large diameter cylinder in the centre (see Fig. 1) over the smaller diameter cylinder on the right hand side. 2.1.1 Test section The test section (see Fig. la) was designed to allow a quick and easy assembly of test baskets from a standard length of tubing. It consists essentially of a frame 460 mm wide and 530 mm high * 1300 mm long, incorporating two detachable plates with holes drilled to the required tube pitching configuration. The tubes are usually arranged vertically between these plates and are connected on the outside by 180° bends to form two or more passes for the cooling water. The same method of construction is used for assembling cross-inclined test baskets in which straight tubes are employed to simulate the configurations of counter-rotating helical coils. The two side walls of the frame incorporate thermocouple and hot wire anemometer traversing assemblies. The positioning of these walls and hence the width of the test basket depends on the transverse pitch of the tubing under investigation and is suitably adjusted to preserve the symmetry of tube arrangement. In the case of staggered arrangements, this symmetry is also ensured by fitting dummy half tubes to the side walls. The general arrangement of the rig is sufficiently flexible to incorporate, if required, other types of test section. Thus, for example, by fitting a suitable transition piece to the inlet and outlet inner ducts it was possible to carry out a series of tests on various arrangements of helical coil heat exchangers with the maximum diameter of 1 m. 2.1.2 Cooling Water Circuit A diagramatic arrangement of the cooling water circuit is shown in Fig. 3. The circuit comprises: (i) a closed loop, filled with de-ionised water, which cools the test section, and (ii) a secondary, raw water loop which provides the heat sink and temperature control for the primary water loop. The two loops are interconnected through two shell and tube heat exchangers arranged in series. - 3 - The water in the primary loop is circulated by a centrifugal pump delivering up to 4 kg/s at 27.5 bar and driven by a thyristor controlled motor. The flow rate is adjusted to the required value by a control valve situated immediately downstream of the pump and the water then flows through the shell and tube heat exchangers, the metering section incorporating an orifice, the test bank, an automatic weigh tank or the by-pass sleeve and then through a de-aerator and back to the pump. The water flow orifice is regularly calibrated using the automatic weigh tank and the orifice constant is updated if it has altered by ±1% from the value originally determined during commissioning. The weigh tank load cell is separately calibrated by temporarily applying a pivoted beam with a known effective mass to the weigh tank suspension linkage. 2.1.3 Control and Instrumentation The rig is operated on line by a dedicated PDP 11/40 computer. The rig is connected to the computer via a high level (±10 mV) analogue input/output multiplexer and by a multichannel digital input output system. All measurements on the rig such as temperatures, absolute and differential pressures, flow rates, etc. are converted into 'mV level signals by means of suitable transducers. The outputs from these transducers are converted into the ±10 mV range for input to the computer by amplifiers which are calibrated on-line for 'zero' and 'gain'. The system also incorporates suitable facilities for on-line calibration of all transducers. Control signals for switching on pumps or providing pulses for controlling stepper motors are generated using the digital output system and position information of the on/off variety from the digital input system such as a valve fully open or traverser at limit of travel. Proportional control signals required for various thyristors are obtained using the analogue output system. The continuous data processing, available from the computer, allows comprehensive instrumentation checks and measurement verification to be carried out prior to and during an experimental run. During the test sequence, instrument calibrations are carried out at each gas pressure change. Any discrepancy promotes a second calibration check and if this is still unsuccessful the rig is automatically closed down. For heat transfer and pressure loss tests five complete sets of measurements are taken at each programmed value of gas-side Reynolds number. If conditions are found to be unsteady during the measuring period or the energy balance is outside the specified range of 0.97 < Iw^g ^ 1.03 the analysis of data is suspended and the calibration routine re- introduced. The five sets of measurements are then repeated and, unless they are found to be acceptable, the rig is shut down. All rig circuits are arranged on a 'fail safe1 principle with the loss of any supply closing the rig down. In addition, a continuous train of pulses is generated as a part of the computer program and applied as a series of signals to an output channel. The suspension of these signals for more than 1 s, indicating a computer failure, immediately closes down the rig. - 4 - The entire software package for the control of rig operations and analysis of test results is written in Coral 66. All temperature measurements on the rig are made with insulated junction, sheathed, mineral insulated nickel-chromium/constantan thermocouples. To ensure consistency, the thermocouples were made from cable produced from the same batch of starting material. Calibrations of the representative sample of thermocouples showed deviations less than ±0.1K at the maximum operational temperature of 150°C. All the thermocouples are taken to a thermally insulated ambient temperature reference enclosure, the temperature of which is separately determined with thermocouples referred to a melting ice cold junction. Strain gauged, diaphragm type pressure transducers are fitted to measure gas static pressure in the rig and the differential pressures across the Dall tube, the test section, the gas circulator and the water flow orifices. 2.2 Atmospheric Wind Tunnels A general arrangement of two atmospheric wind tunnels employed in the investigations of gas-side heat transfer and flow characteristics of steam generators in the AGR stations is shown in Fig. 4. Both tunnels can accommodate test baskets from the pressurised heat exchanger for a complementary or exploratory study of some specific feature of a given tube arrangement. The main application of these tunnels is, however, to the investigations of gas flow distribution in the scaled up or down models of various sections of steam generators using LDA and hot wire anemometry measuring techniques. 2.3 High Temperature Rig The rig was designed to assist the investigations of specific gas-side problems in steam generators of High Temperature Reactors and to extend the operational range of test facilities for general purpose studies. The maximum values of the more important operating parameters of the rig are as follows: (i) gas pressure, temperature and flow rate - 15 bar, 500°C, 1.4 kg/s of He (ii) cooling water pressure, temperature and flow rate - 105 bar, 280°C, 1.5 kg/s Similarly to the already described pressurised gas heat exchanger test facility, this rig consists of a closed loop circular duct incorporating a Rateau gas blower (ex Dragon project), a 500 kW electrical heater, a 30 mm wide x 30 mm high x 60 mm long test section, a helical coil gas to water heat exchanger and the necessary control and measuring instrumentation. The plain or finned tubes of the test section are arranged in cross-flow to gas and are cooled by the high pressure water circulating inside them. The rig is operated on-line by a dedicated computer on the same general principles of control and data acquisition as those used for the pressurised heat exchanger rig. - 5 - 2.4 Miscellaneous Rigs From time to time a specific problem arises which cannot be resolved satisfactorily by investigations in the available test facilities or there is a need to supplement the already obtained information by additional data. In such cases, special purpose test rigs ranging in size from bench top facilities to large loops are quickly constructed and after providing the required data, dismantled. Over the past twenty five years many test facilities of this kind have been used at CERL and only few examples can be given here to illustrate the variety and range of applications. One of the first of such rigs was a water loop designed to measure heat transfer to water flowing inside short sections of tube connected by 180° bends, (Lis and Thelwell, 1963). Accurate data on heat transfer in such arrangements were not available at that time and the information was urgently required for the analysis of test results from the pressurised heat exchanger rig. The correlation derived from this investigation is still employed by CERL and other organisations working in this field in the UK. One of the requirements of the early design studies of the AGR stations was to assess the effects of gas-side oxidation on the performance of steam generators. A survey of the literature indicated that published data on the thermal conductivity of iron oxides differed by almost two orders of magnitude and there were no data on thin oxide layers in the 'as-formed' conditions. A special apparatus was, therefore, developed and used to measure the thermal conductivity of iron oxide layers 0.025 mm to 0.48 mm thick which formed on low alloy steels in CO-CO2 gas containing up to 1200 vpm of water at pressures up to 20 bar and temperatures up to 600°C (Lis and Kellard, 1968). It was found that the thermal conductivity decreased with oxide porosity and the measured values were effectively correlated as a function of porosity (see Fig. 5). Prior to the construction of the high temperature rig a preliminary investigation was carried out to assess the effect of varying gas physical properties on convective heat transfer to a bank of tubes in cross-flow (Preece, Lis and Hitchcock, 1975). The general arrangement of the gas circuit in a rig specially adapted for this purpose is shown in Fig. 6. Both the required mean bulk temperature of the air and the flow rate at the test section were achieved by mixing cool air supplied from a blower with combustion products from a gas-oil-fired combustion chamber. The air bulk to tube wall temperature differences could be varied between 100°C and 400°C with the tube wall remaining constant at about 40°C or 120°C. It was found that the variation in physical properties could be accommodated by evaluating all properties in the dimensionless groups of the conventional type of correlation at the air bulk temperature. More recently, a test facility has been constructed and is still being employed for determining the steady state and transient characteristics of various thermocouple assemblies installed in the AGR steam generators. Precise knowledge of these characteristics is especially important in the case of CS/9Cr 1 Mo and 9Cr lMo/316 SS transition thermocouples which form a part of the plant protection system - 6 - and are also used for control of steam generator operations. A schematic diagram of the thermocouple response rig is shown in Fig. 7. The thermocouple assembly under test is mounted in a duct fitted with a bellmouth intake through which atmospheric air is drawn. Water is initially circulated round a by-pass loop containing a heater and when the required constant temperature is obtained the flow is diverted by a three-way valve to the test section to provide a step change in temperature. All test section and rig instrumentation is continuously monitored until a new steady state has been achieved. The water is then diverted back to the by-pass loop and air and water flow rates adjusted for the next test. The rig is controlled on-line to a PDP 11/40 computer and rig operation is entirely automatic. A new test facility is now being designed to measure the heat transfer and flow characteristics of close-pitched arrangements of tubing in cross-flow under the combined forced/natural convection gas flow conditions. With the increased emphasis on all aspects of safety in the AGR plant operations, accurate information on these characteristics is required for the assessment of the decay heat removal duties and temperature distribution in steam generators under various post-trip conditions. The rig will consist essentially of a vertical closed loop duct about 10 m high, one side of which, incorporating a 125 kW electrical heater and a gas flow venturi, will constitute the hot leg and the other side, housing the standard CERL test basket section with the automatic traversing devices in side walls, the cold leg for the natural convection. A centrifugal fan driven by a variable speed motor will be placed in the bottom return leg of the loop to provide the required forced circulation. Full simulation of the steam generator dynamic conditions will be obtained by using SFg gas at 1.4 bar pressure. 2.5 Temperature Calibration Facilities Comprehensive facilities for calibrating temperature measuring devices against the accepted references standards form an essential part of support services in the CERL programme of work on steam generators. The most important components of these facilities are: water bath for temperatures up to 85°C, silicone fluid bath, 80°C to 225°C and nitrate bath, 180°C to 625°C. The silicon fluid and nitrate baths are based on the NPL models (Grace and Hall, 1943) which have been modified by reducing the thermal capacity, fitting directly immersed heaters and using proportional temperature controllers. The temperature stability and uniformity obtained is ±0.02K for the silicon bath and ±0.03K for the nitrate bath. 3. EXPERIMENTAL INVESTIGATIONS 3.1 Heat Transfer and Pressure Loss Characteristics of Close- pitched Tube Arrangements in Cross-flow The main objective of these investigations is to provide accurate data required for the assessment of design, control and operational performance of steam generators in the Magnox and AGR stations. The arrangements of tubing in each steam generator are first tested at the tender design stage to assess the validity of the design submissions and, subsequently, the tests are repeated, as required, to determine the implications of any design modifications and the effects of manufacturing and assembly tolerances. - 7 - The measurements of the gas-side heat transfer and pressure loss in a given arrangement of tubing are analyzed using generally accepted procedures. The overall heat transfer coefficient, u^, is evaluated from the installed heating surface area and the measured values of heat gained by water, qw, and the logarithmic mean of terminal temperature differences between gas and water, 9m. The mean gas-side heat transfer coefficient, hg, is then derived from the usual relationship for thermal resistances of gas, tube wall and water in series. In calculation of the water-side thermal resistance, an allowance is made for the effects of bends, thermal boundary layer development and radial temperature distribution by a general correlation proposed by Lis and Thelwell (1963). In the case of finned tubes, the measured coefficient, ho, is converted into the gas-film heat transfer coefficient, hf, using the fin efficiency proposed by Gardner (1945) with the fin height modified by the addition of a half fin thickness. The heat transfer results are correlated in the form, j = Nu Pr~1/3 = a Rem and log j = A (log Re) (log Re) + B log Re + C The measurements of pressure loss across the test bank are carried out under conditions of zero heat flux and correlated in the form, f = b Ren and log f = D (log Re) (log Re) + E log Re + F where f is the pressure loss coefficient defined as the number of velocity heads lost per one row of tubing. For both the heat transfer and pressure loss data, all physical properties in the dimensionless parameters are evaluated at the mean gas bulk temperature. For each set of test results, the regression line and the 95% confidence limits are evaluated by statistical analysis. Typical experimental data obtained in the pressurised heat exchanger rig are shown in Fig. 8. It will be seen that the plot of j against Re in the log-log coordinates has a definite curvature and the best correlation of data is obtained by using a second degree polynomial curve. There is also a characteristic change in the slope of f against Re plot at Re - 1.2 x 105 indicating a transition to a flow regime where the pressure loss coefficient becomes independent of the Reynolds number. For the majority of the tube configurations tested so far at CERL, the curvature of the heat transfer data plots is usually negligible over the range of Reynolds numbers covered and the traditional straight line correlation is perfectly adequate. The characteristic change in the slope of the pressure loss characteristics has been observed to occur for all arrangements of finned tubes at about 8 x 10^ < Re < 1.2 x 105, except for the cross-inclined configurations where the transition in flow regime takes place at appreciably lower Reynolds numbers. The reproducibility of results obtained in this type of tests on the pressurised heat exchanger rig is illustrated in Fig. 9. The data were derived over a period of about nine months using the same configuration of plain tubes in three separate series of tests employing air, carbon dioxide and helium. 3.2 Gas Flow Distribution One of the problems encountered occasionally in the large steam generator units employing curved tube arrangements is the maldistribution of gas flow caused by the asymmetries of tube pitching at the casing walls. With the casing/edge tube gaps larger than those required to preserve the symmetry of tube arrangement there could be a considerable amount of gas by-passing the steam generator, whereas smaller gaps will cause a starvation of gas flow at the edge tubes. In order to provide the information required for the operational control of steam generators with such asymmetries (see Section 4) a comprehensive investigation of this problem is being carried out at CERL. The measurements of gas flow distribution are most conveniently carried out on scale-down models of various sections of steam generator. This allows the correct representation of the steam generator aspect ratio and depth within the frame of the standard CERL test section. Banks of full scale finned, tubes representing a relevant part of the steam generator section, are also used to provide supplementary information. The measurements of velocity distribution are carried out using laser Doppler and hot wire anemometers. For the preliminary measurements (Fallows and Massey, 1982a) a 5 mW He-Ne laser provided the source for the forward scatter operation of one component counter based DISA system. More recently, the LDA facility has been extended by the addition of a 2W Ar laser and a back scatter module. For measurements of velocities between the narrowly spaced tubes and within the casing/edge tube gaps, the recently developed DISA triple split fibre probes were found to be most useful. The experience gained so far have shown these probes to be more rebust and less sensitive to contamination than the classical wire probes. Trials are now being carried out on the suitability of these probes for measurements in the pressurised heat exchanger rig. The more important characteristics of the by-pass flows were found to be generally as expected. At the inlet, the velocity profile is distorted by the movement of gas towards the gap. The velocity within the gap increases gradually by the inflow of gas from the body of the bank until an asymptotic value is attained at a bank depth which is function of the gap width. Downstream of the bank there is usually an extensive recirculation region of the gas emerging from the gap. It is evident that in order to describe the distribution of gas flow in this complex situation it is necessary to define the magnitude of the transverse flow component and the development of flow within the gap. The evaluation of the transverse flow is especially important, since, apart from affecting the overall flow distribution, it provides also a convective component of the overall gas thermal mixing. The magnitude of the transverse flow will depend on the transverse pressure gradients generated by the differences in gas velocities and densities across the bank and on the transverse flow resistance factor of a given tube - 9 - arrangement. Realising that the direct measurements of the transverse flow resistance would be very difficult because of the complicated nature of flow, a simplifying concept of a pressure equalisation factor, X, was used in a mathematical model developed at CERL (Fallows and Massey, 1982b) to predict the flow distribution and thermal diffusion in models of steam generators with gas by-passing. X = 1 represents a complete equalisation of pressure at a given level of the tube bank and X = 0 defines zero pressure equalisation, i.e. the gas flow profile remains unchanged at that level of the bank. A comparison of this model predictions with a typical set of velocity measurements carried out in a 1/4 scale model of a section of the Dungeness TBr steam generator is shown in Fig. 10. 3.3 Gas Thermal Mixing The maldistribution of fluid and metal temperatures in a steam generator operating under asymmetric conditions will depend essentially on the gas thermal mixing. The overall mixing comprises the already mentioned convective component and in most situations a predominant diffusive component. Heat transfer by diffusion is assumed to take place normal to the main flow direction and can be conveniently expressed as a turbulent Peclet number, PE, based on the total thermal conductivity, Xtot = \nol + ^turbulent' At Present there is no complete theory of diffusive mixing in banks of tubes in cross-flow and it is necessary to determine experimentally the turbulent Peclet number for each arrangement of tubing. Preliminary investigations of diffusive mixing in staggered and in-line arrangements of both plain and finned tubes in cross-flow were carried out in an atmospheric wind tunnel at zero heat flux conditions (Jones, Lis and Massey, 1978). A thermal spike was generated upstream of the test bank and the decay of this spike was measured by thermocouples traversing the gas flow at various positions within and downstream of the test bank. For the great majority of tube arrangements covered in this investigation there was a uniform decay of the temperature spike as shown in Fig. 11. It is assumed that the mixing process may be described by the simple diffusion equation where T is the gas temperature, De is the equivalent diameter and x and y are distances in the direction and normal to the overall gas flow, respectively. This equation was solved numerically for PE by the Crank- Nicolson method (1947) using three points at the corresponding positions near and at the peak on each of two experimentaly determined temperature profile curves. It was found that in the range of Reynolds numbers covered, 1.1 x lO*4 < Re < 6.6 x 10^, the turbulent Peclet number for different tube configurations varied between 12 and 30. There was no strong dependence of PE on Reynolds number. One striking exception to these general characteristics was a close-pitched in-line arrangement of plain tubes. It will be seen from Fig. 12 that there is a distinct sideways drift of the peak on each successive temperature profile curve. The direction of the drift was found to be random with respect to - 10 - Reynolds number and (x,y) plane. The occurrence of drift has been confirmed by smoke and tuft tests and Batham (1973) has deduced a similar effect from measurements of pressure distribution on tubes in an in-line square arrangement with P^/D = 1.25. The effects were absent, however, when the pitch to tube diameter ratio was increased to 2.0. The physical processes involved in drift are not fully understood at present and the most often advanced explanation is some form of Coanda jet switching mechanism. It is readily apparent that the simple diffusion equation cannot be used to describe the complex mixing processes involving drift. In fact this method of analysis fails also to describe symmetrical diffusion adequately because it defines only a part of rather than the whole process. In both cases it would be preferable to use the entire temperature profile curve in the analysis. At present, this cannot be done directly and a method of fitting temperature profiles predicted by the mathematical model (Fallows and Massey, 1978b) for a range of PE with the measured profiles has been employed for all subsequent investigations. The great majority of these investigations were carried out in the pressurised heat exchanger rig over the operationally relevant range of Reynolds numbers and heat fluxes. Two techniques were employed: (i) A stream of hot gas with uniform temperature was passed over a bank of water cooled tubes in which one or more longitudinal tube rows were blanked off. (ii) A temperature profile was generated at the bank inlet by injecting isokinetically a stream of hotter gas into the hot mainstream gas and the decay of this temperature spike in a bank of water cooled tubes was measured. Both these techniques provide close simulation of the more important causes of asymmetrical conditions in the AGR steam generators and as such are a considerable improvement on the originally employed test procedure. Also, the close control of the pressurised heat exchanger rig operations by the dedicated computer and the use the automatic thermocouple traversers (see Fig. la) has appreciably improved the accuracy of experimental data and increased the scope of investigations. A typical set of experimental data obtained using technique (i) is shown in Fig. 13 together with the predictions of the mathematical model. It will be seen that the predicted development of the temperature spike for PE = 12 is in good agreement with the measurements. It is of some interest to note that a very similar shape of the gas outlet temperature profile was generated by a large casing/edge tubes by-pass gap with all tubes cooled by water. In this case, however, the best fit PE was about 30 indicating appreciably lower mixing. Repeat tests on the in-line arrangement of plain tubes using technique (i) have confirmed the presence of drift in this configuration of tubing and, what is even more Important, showed that with PE ranging from 0.3 to 0.7, thermal mixing in this tubing is more than an order of magnitude higher than in all other arrangements of tubing tested at CERL. This confirmatory finding is of particular importance to the operation and control of steam generators in the Hinkley Point 'Br, Hunterston 'B1, Heysham II and Torness AGR station where this particular In-line arrangement of plain tubes is used in the 9%Cr IMo sections of the generators. - 11 - The improved experimental and analytical techniques have revealed many new interesting features and, in some cases, very considerable complexity of gas thermal mixing processes in banks of tubes in cross-flow. In an attempt to throw some light on these processes, comprehensive investigations combining detailed measurements of gas flow and temperature fields are being carried out. 3.4 Other Experimental Investigations The scope of this paper does not allow even brief description of all the major investigations carried out by CERL on the gas-side heat transfer and flow characteristics of steam generators in the AGR and Magnox stations. Some typical examples of these investigations are given in Section 2.4 in conjunction with the description of miscellaneous test facilities and few more will be described briefly here to illustrate the wide range of our studies. A series of comprehensive studies have been undertaken at an early stage of our programme of work to obtain a better understanding of the heat transfer and flow processes within banks of finned tubes in cross-flow. In one series of such investigations (Neal and Hitchcock, 1966), a four times scale model bank of annular finned tubes in staggered arrangement was used. The test section was situated inside a Perspex duct of an atmospheric wind tunnel and comprised nine transverse rows of finned tubes. One of these tubes, which could be positioned in any row, had an internal radiant heater and was comprehensively instrumented with thermocouples and miniature disc heat-flux meters. A constant temperature 0.005 mm dia. hot wire anemometer was used for measurements of velocity, turbulence and boundary layer development. These measurements were supplemented by flow visualisation studies using smoke and fine silk tufts. The investigations yielded most valuable information on the overall and local heat transfer processes and revealed several adverse characteristics of this type of extended surface. The next logical step in this study was to apply the obtained information to the development of improved heat transfer surfaces for tubes in cross-flow (Neal and Hitchcock, 1970). Preliminary investigations indicated that a rib roughened surface offered the most promising solution and a series of tests on the large scale model were carried out to determine the optimum rib size and pitch. The pitch/height ratio finally selected was 7.54:1 and tests in the pressurised rig on a set of 38 mm dia. tubes with integrally machined ribs showed a 41% increase in heat transfer at Re = 1.4 x 105 over the corresponding arrangement of plain tubes. The main reason for this increase is the markedly improved heat transfer over the front half of the tube (see Fig. 14). Further investigations in this field have led to the development of tubes with two or more start helically wound ribs. It is worth noting that this roughened surface should be generally less susceptible to the flow induced vibration problems because it cannot sustain any regular shedding of vortices. 4. MIX SUITE OF 2-DIMENSI0NAL STEAM GENERATOR CODES The MIX suite of codes has been developed to predict the steady state performance of Magnox and AGR steam generators operating under asymmetric conditions. The codes calculate the distribution of gas, - 12 - water/steam and tube wall temperatures along each tube of the steam generator for a number of different boundary conditions. The diversity of the various designs of gas heated steam generators in nuclear stations made it impractical to perform multi tube calculations for all stations with a single code. Consequently, codes have been developed for individual stations or groups of similar stations employing a number of common subroutines, steam tables, gas physical properties and the input and output routines. The codes comprising the MIX suite are listed in Fig. 15. All codes of the MIX suite are based on the mathematical model, of a multi-pass cross-flow heat exchanger with primary fluid mixing described in a paper by Fallows, Gane, Jones, Lis and Massey (1979). Taking into account the computational capacity of the available main frame computer and the need to provide a practical and flexible tool for the steam generator designers and operators, it was assumed that the variations of the temperature and flow fields in the direction of tube axes can be ignored and that 2-dimensional representation would be perfectly adequate for the great majority of applications. Fig. 16 shows a 2-dimensional representation of typical in-line and staggered tube banks. The shaded region in each of these diagrams indicates the symmetry units of the banks which are used as the basis of the numerical solution. On the gas side the model solves the equations of conservation of heat and momentum to determine the gas temperature and flow conditions respectively. The heat balance equation relates the convective and diffusive components of heat transfer in the gas to the loss of heat to water/steam. The turbulent Peclet number, PE, and the gas pressure equalisation factor, X, are input to the model to account for the gas thermal mixing. The distribution of the mainstream gas flow is evaluated using appropriate friction factors for each gas stream and any gas by- pass lane present in the tube arrangement. The water/steam flow distribution is determined by consideration of the pressure losses due to friction, gravity and acceleration. For a given gas side flow and temperature distribution the water side temperature/enthalpy distribution is determined from the heat transfer rate equation. The overall heat transfer between the gas and water/steam is evaluated using appropriate heat transfer correlations for gas, metal, water, 2-phase water/steam and superheated steam. Additional heat transfer resistances on the gas and water/steam side may be included as required. The boundary conditions specify the gas and water/steam temperature profiles at opposite ends of the counterflow steam generator and hence the gas and water-side equations can only be solved iteratively. Thus the calculation procedure consists of an inner iteration, which evaluates the gas side heat transfer and flow for a given water flow and temperature distribution, and an outer iteration which adjusts the water flow distribution to meet the new boundary conditions. The process is repeated until the change in gas temperatures between the successive iterations is less than the specified value. Each code can be used either to evaluate steam generator conditions with the installed feed water flow orifices or to define the orifices which would give the required water/steam enthalpy at a given position in the steam generator. Additional iterations are included in each code to meet the specific operational and control requirements of a given type of steam generator. Thus, for example, there are 53 calculation options available in DUNMIX. - 13 - 5. APPLICATION AND VALIDATION OF MIX CODES The validation of any new complex mathematical model of the once-through steam generator is bound to be a prolonged, or, perhaps, even a continuing exercise. The initial steps of such an exercise will involve usually comparisons with the predictions of other established, but generally simpler, models and with data obtained from the laboratory rigs simulating a specific aspect of plant operations. Each completed step along this route will increase the level of confidence in the model validity. The complete validity can, however, be demonstrated only by a good agreement between the model predictions and plant data obtained over a wide range of operating conditions. Some comparisons of the predictions of the mathematical model incorporated into the MIX suite of codes with the rig data have been already discussed in Sections 3.2 and 3.3 of this paper. Several examples of the initial application of this model to the design and operational problems in once-through steam generators of the AGR and Magnox stations are given in a paper by Balfour et al. (1979). One of the first cases of the model application concerned a severe maldistribution of feed flows in the Wylfa steam generator operating with a number of water passes blanked off. The problem was compounded by the dynamic instability of water/steam flows in some platens. WYLMIX code was used to assess the existing situation and then to evaluate the feed ferrule distribution required to restore the operating conditions to the design level. Measurements of the individual tube flow rates made at the start of the exercise and after re-ferruling were found to be in good agreement with the model predictions. At an early stage of power raising at Hinkley Point fB' a number of feed flow orifices became defective causing a severe violation of the 90K minimum steam superheat constraint at the entry to the austenitic steel sections of the steam generator operating at 82% MCR. As a temporary remedial measure it was decided to operate the generators in a 'pinch' mode at a reduced power level. This mode of operation required a careful choice of boundary conditions and, in particular, an accurate definition of feed pressure to ensure the preservation of the superheat constraint and the dynamic stability of water/steam flows. An early version of HEYMIX was used in conjunction with the dynamic stability code SPOTS to evaluate the required conditions and the steam generators operated successfully at about 60% MCR until the modified feed orifices became available. A comparison of the PODMIX prediction with the plant data obtained at the early commissioning phases of Heysham I is shown in Fig. 17. The main point of interest in the steam generator behaviour was the rather unexpected reduction in the tube wall temperatures of the innermost helices during Phase III which all but disappeared during Phase IV operations. The power level for both phases was about the same, but Phase IV operated with lower gas flow rate at higher inlet temprature. These differences in the operational boundary conditions had a large effect on the steam generator heat losses to the spine and it will be seen that PODMIX accounted for these differences very accurately. - 14 - Finally, the application of the MIX codes to design of steam generators is illustrated by considering the effects of inlet feed throttling on the sensitivity of a once-through steam generator to perturbations in boundary conditions such as gas inlet temperature profile, heat transfer coefficients, water/steam side friction factors, heat losses through the containing wall, blanked off tubes, etc. Preliminary investigations indicated that the uncertainties in the individual parameters can be conveniently combined into three types of perturbations, namely water-side, gas-side and blocked tubes. For each level of inlet impedance the distribution of ferrule sizes required to give 50K of steam superheat at the 9%Cr lMo/316 SS transition in each tube for the best estimate of all parameter values was calculated. Using the ferrules defined in this manner, the steam temperatures at the transition were then evaluated for the perturbed boundary conditions and the standard deviation, a, of the perturbed temperature from the corresponding reference temperature determined. The results are shown in Fig. 18 where it will be seen that the rate of change of a becomes negligible above 12 bar inlet impedance for 100% load and 2.5 bar for 40.5% load. Taking into account the contribution of the restrictor tubes and hydrostatic head, a ferrule giving effective impedance of 12 bar at 100% will provide only 0.6 bar impedance at 40.5% load. Such a ferrule will obviously have little effect on the steam generator sensitivity, which may also become statically unstable. In order to reduce the steam generator sensitivity and ensure stability at low loads the impedance of 2.5 bar is clearly required. However, the corresponding effective impedance at 100% load will be 24 bar, i.e. twice the value required by the stability and sensitivity considerations. The problem of excessive parasitic pressure losses at higher loads could, of course, be solved by the installation of a variable throat ferrule, if such devices were available. 5. ACKNOWLEDGEMENT The work reported here was carried out at the Central Electricity Research Laboratories and this paper is published by permission of the Central Electricity Generating Board. 6. REFERENCES Balfour, J.D., Fallows, T., Gane, C.R., Gill, G.M., Jones, R.C., Lis, J. and Preece, R.J., 1979. The application of two-dimensional model to design and operational problems in AGR and Magnox stations. BNES 2nd Int. Conf. Bournemouth, Paper 11 Batham, J.P., 1973, Pressure distribution on in-line tube arrays in cross-flow. Int. Symposium on Vibration Problems in Industry. Keswick, England, Paper 411 Crank, J. and Nicolson, P., 1947, Proc. Cambridge Phil. Soc. 43, p. 50- 67 Davies, F.V. and Lis, J., 1960, Heat transfer and pressure drop characteristics of concentric arrangements of helical coils. Instn. Mech. Engrs Symposium on the Use of Secondary Surfaces for Heat Transfer with Clean Gases, London, Paper 5 - 15 - Fallows, T., Gane, C.R., Jones, R.C., Lis, J. and Massey, T.H., 1979, CROSSMIX, A mathematical model of multi-pass crossflow heat exchanger with primary fluid mixing, BNES 2nd Int. Conf, Bournemouth, Paper 38 Fallows, T. and Massey, T.H., 1982a, Studies of nuclear boiler models with LDA, Int. Symposium on the Application of LDA to Fluid Mechanics, Lisbon Fallows, T. and Massey, T.H., 1982b, Flow and gas mixing in heat exchangers with gas by-passing. 7th Int. Heat Transfer Conf., Mlinchen, Paper HX6 Gardner, K.A., 1945, Trans ASME, 6_7_ pp. 621-632 Grace, A. and Hall, J.A., 1943, J. Sci. Inst. 20, 60-63 Jones, R.C., Lis, J. and Massey, T.H., An experimental investigation of thermal mixing in crossflow tube banks 6th Int. Heat Transfer Conf., Toronto, Paper HX12 Lis, J. and Thelwell, M.J., 1963, Experimental investigation of turbulent heat transfer in a pipe preceded by a 180° bend. Proc. Instn. Mech. Engrs, 1963-64, 178 (Pt 31(IV)), 17 Lis, J. and Kellard, P.O., 1968, Measurements of the thermal conductivity of thin films of magnetite, Brit. J. Appl. Phys. (J. Phys. D) Ser. 2, Vol. 1, pp 1117-1123 Neal, S.B.H.C. and Hitchcock, 1966, J.A., A study of heat transfer processes in banks of finned tubes in cross flow, using a large scale model technique, 3rd Int. Heat Transfer Conference, Chicago, Paper 110 Neal, S.B.H.C. and Hitchcock, J.A., 1970. The development of improved heat transfer surfaces for tubes in cross flow using a large scale model technique, 4th Int. Heat Transfer Conference, Paris - Versailles, Paper FC7.8 Preece, R.J., Lis, J. and Hitchcock, J.A., 1975. Effect of gas-side physical property variations on the heat transfer to a bank of tubes in cross flow. Proc. Instn. Mech. Engrs, Vol. 189 17/75, pp 69-75 Tachometer Gas/Water Gas Temperature Gas Dump Heat Exchanger Thermocouples Valve BAS FLOW rff To Static To aP Hotorised Pressure Transducers TransducerlH Regulator To Static Pressure U/S Gas Thermocouple Transducer Lirid (16 off) ••" * Heater Control Thermocouples 550 KW Heater Fig.2 Schematic Diagram of Heat Exchanger Rig. Fig 1 General View of Heat Exchanger Rig TUBE SUPPORT PLATE DRILLED THERMAL INSULATION TO GIVE REQUIRED TUBE PITCHING STEPPER MOTOR FOR THERMOCOUPLES -MICROSWITCHES METALLIC SEAL THERMOCOUPLE Fig la Test Section with Traversing Assemblies Fig. 3 Rig Water Circuit. ~n *> ^ Pressure Loss Parameter, f Fi n Thicknes s Tl I D 100262 6 o 3 -a X a o oO o Ul 9090 0 DIMENSIONLESS PRESSURE DROP PARAMETER, f TO M fD t-i H- fD 3 l-h O H i-j fD Co OP fD CO ?3 fD - 18 - )ooooooooooo°d X INLET (RO«S 0 1) OH0ISIM! 7 OUTLET (RO*S 16 -) Fimw 7//4 n~77[ \3+mr\ By pa: Ton* Ffxtor ! I ft t 0-0? 2 0-7F 3 0-80 4 0-005 f n i / A f-S mm By pot* to) X : 1-C DISTANCE ACROSS DUCT, m Axial Position Fig 11 Thermal Mixing in a stag- gered Tube Arrangement o o o oo. J6 Longitudinal ftcwi oo Hotel Seal* 1 4-29 oo BLANKED COOLED O O O ooo TUBUBE TUBES ooo I 2 3 4 3 Fig.TO: Comparison of Predicted Bypass Velocity Profiles With Measured Values. STAGGERED TUBE BANK, 14 ROWS OF TUBES TRANSVERSE PITCH 19.22 mm LONGITUDINAL PITCH 32.76mm O 96 FINNED TUBES 0.0.18.04mm, FIN HEIGHT3.0mm OOOOOOOOOOOOOOOOOOOO £ 9* • MEASURED BANK OUTLET GAS TEMPERATURE 5 92 oooooooooooooooooooo PREDICTED TEMPERATURE LU FOR P« -12 a- 90 - K 88 - in S 86 - 84 - «2 — 100 200 300 400 DISTANCE ACROSS DUCT, mm Fig 13 Predicted and Measured DO 300 DISTANCE ACROSS DUCT, m Outlet-Temperatures for a Fig 12 Thermal Mixing in a Closely- Bank with an Edge Tube Pitched In-line Tube Blanked Arrangement STREAMS i-l j j + l oooo LEVELS 80 OCJOO 40 NODE '-r ' i +l fSYMMETRY 111 lUNIT -L. • I + -L * i j_ o LEVELS o 0 25rt 05n 0 75n n i-l ANGULAR POSITION ON TUBE RADIANS Fig 14 Circumferential Q ®6. • Variation of Heat i + l Transfer over a Ribbed and Plain : Q ; O Fig 16 Node Distribution for Typical Tube Surface In-line and staggered Tube feanks CROSSMIX WYLMIX DUNMIX PODMIX HEYMIX WATER-SIDE PERTURBATIONS HINKLEY POINT A' WYLFA DUNGENESS B' HEYSHAM A' HINKLEY POINT B' 30 SIZEWELL A1 HARTLEPOOL HEYSHAM 'B' 40.25% LOAD 25 MAGNOX MAGNOX AGR AGR AGR DRUM BOILERS ONCE-THROUGH 20 SERPENTINE TUBES IN SERPENTINE SERPENTINE TUBES HELICAL TUBES SERPENTINE TUBES CYLINDRICAL CASING. TUBES IN IN ANNULUS IN CYLINDRICAL IN RECTANGULAR BOILERS OUTSIDE ANNULUS BETWEEN REACTOR PODS. MULTI-START BOX. COMBINATION CONCRETE PRESSURE BETWEEN REACTOR AND CONCRETE CONTRA-ROTATING OF PLAIN AND VESSEL AND CONCRETE PRESSURE VESSEL. HELICES OF FINNED TUBES IN PRESSURE VESSEL. INTERLEAVED EQUAL LENGTH STAGGERED AND BOILERS INSIDE REHEATER ACOUSTIC INLINE ARRANGEMENTS VESSEL BAFFLES GAS-SIDE PERTURBATIONS Fig 15 MIX-Suite of Codes 9%Cr1Mo/3l6SS TRANSITION o o o o «• S S 8 % Ftias* 3 Phise U Measured o Predicted • + BLOCKED TUBES 350 i S 1 CS/9%Cr1Mo * * > TRTRANSITIOA N 10 _L 5 10 15 20 8 10 12 HELIX NUMBER INLET IMPEDANCE-BAR Fig.17 Comparison of PODMIX Predictions with Measured Temperatures etHeysham I during Phase 3 &4 Operations. Fig 18 Sensitivity of Boiler to Inlet Impedance No. 15 XA0055824 THE EFFECT OF INLET AND OUTLET SHELL - SIDE FLOW AND HEAT TRANSFER ON THE PERFORMANCE OF HTGR STRAIGHT TUBE HEAT EXCHANGERS D. P. CAROSELLA Staff Engineer GA Technologies Inc. San Diego, California, USA ratio is small (less than 1.3), the 90-deg turns can ABSTRACT have a critical effect on the performance. However, Since the mid-1970s, various high temperature use of design concepts such as flow baffles or small gas-cooled reactor (HTGR) steam generator, aux- modular bundle designs can greatly reduce or elim- iliary heat exchanger (AHE), recuperator, and in- inate these problems. termediate heat exchanger (IHX) designs have 3. The inlet turns are more critical to the design been proposed that use straight tube configurations. than the outlet turns. Each of these designs requires 90-deg turns in the 4. The non-unform temperature profiles result- helium gas flow at the inlet and/or outlet of the ing from poor inlet design can result in high thermal tube bundle. stresses in the tubes. The design of the steam generator for the HTGR 5. Testing to correlate the shell-side heat trans- steam cycle/cogeneration lead plant includes a fer in the region of a 90-deg bend across a tube straight tube superheater (STSH) which incorpo- bundle is required to more fully understand the rates both a 90-deg inlet and outlet turn across the characteristics of this problem. tube bundle. The AHE includes a 90-deg outlet turn 6. Although the 90-deg turns are not desirable, across the tube bundle. Previous GA Technologies they are frequently necessary in realistic plant and Inc. (GA) recuperator and IHX designs for gas tur- component designs. However, with the proper anal- bine and process heat HTGRs have also considered ysis and testing, the designs can be developed so as straight tube designs with 90-deg bends at the inlet to minimize their adverse effects on tube bundle and outlet. performance and thermal stress. To evaluate the effect of these turns on the tube bundle performance, two model air flow tests have NOMENCLATURE been performed, and a third is being planned. Fluid flow and heat transfer computer models have also The following symbols are used in the text and on been used to try to determine the effect of these the figures in this paper: 90-deg turns on the tube bundle performance. XT tube pitch-to-diameter ratio As a result of these studies, the following con- RQ bundle outside radius clusions can be made: Rj bundle inside radius 1. For all of the designs investigated, the 90- Rl local bundle radius deg turns reduced the bundle performance. V^ local velocity 2. In designs where the tube pitch-to-diameter Vm mean velocity INTRODUCTION DESCRIPTION The primary coolant heat exchangers used in an Inlet Turns HTGR are located in cavities in a prestressed con- crete reactor vessel (PCRV). The axes of these heat In the inlet region, flow enters the bundle per- exchangers are parallel to the vertical axis of the pendicular to the tube bundle vertical axis (Fig. 2). PCRV (Fig. 1). The primary coolant is transported The flow sees both a vertical and a horizontal flow to and from the heat exchangers via ducts. The path. The relative resistance of these paths will es- geometrical arrangement of the flow paths requires tablish the flow distribution. In bundles where the the primary coolant to make one or more 90-deg tube pitch-to-diameter ratio (XT) is small (less than turns in the heat exchanger cavity. 1.3), the radial flow resistance will cause the flow For straight tube configurations, the 90-deg turn to travel upward before it reaches the center of the must occur while crossing the tube bundle. The flow bundle, resulting in excessive flow at the outside of and heat transfer characteristics of three such con- the bundle and flow starvation toward the center figurations have been investigated by a combination of the bundle. In bundles with large (greater than of analysis and flow model tests. 2.0) XT values, the flow distribution will be nearly References 1 and 2 reported the results of two uniform. such investigations. By using these results and an- The inlet flow maldistribution can reduce the alytical results for a third configuration, an attempt bundle performance by two mechanisms. In the will be made to characterize both the flow and heat flow-starved region, the shell-side film coefficient transfer characteristics of straight tube heat ex- will be less than nominal. This reduction of film changers with 90-deg inlet and/or outlet turns. coefficient is partially offset by the higher film coef- ficients in the outer regions of the bundle; generally, CONTROL ROD DRIVE AND REFUELING PENETRATIONS AUXILIARY CIRCULATOR CIRCULATOR FEEDWATER PRESTRESSED ACCESS SHAFT CONCRETE REACTOR VESSEL STEAM CORE GENERATOR LINEAR CIRCUMFERENTIAL PRESTRESSING PRESTRESSING SYSTEM SYSTEM PCRV SUPPORT STRUCTURE CORE AUXILIARY HEAT EXCHANGER N-028 A17351 Fig. I. HTGR steam supply system TUBE (1 FLOW / U-J Outlet Turns The outlet turn compounds the problems result- OUTLET ing from the inlet turn. As the shell-side fluid travels SHELL SCREEN radially outward, a radial temperature profile de- velops in the outlet region, further reducing the FLOW mean effective temperature difference between the primary and secondary coolants. Both the inlet and outlet turns contribute to re- duced heat exchanger performance. However, since the inlet effect is propagated throughout the length of the bundle, it is by far the more critical effect. Bundles with only an outlet turn are not greatly affected by the flow maldistribution in the outlet region. INVESTIGATIONS Fluid Flow The fluid flow characteristics of three straight tube configurations have been investigated by a combination of air flow tests and/or analysis. Each of these heat exchangers has a different geometry. BAFFLES IN LET They represent a spectrum of XT values ranging SCREEN from 1.25 to 2.13. One (the AHE) does not have an inlet turn in the tube bundle. Investigation of these three heat exchangers establishes a base for TUBE-SIDE FLOW looking at other straight tube designs that might be A17351 encountered in gas-cooled reactor configurations. Straight Tube Recuperator/IHX Model. A model Fig. 2. Straight tube configuration of the inlet region of a heat exchanger characteristic of a typical recuperator or IHX used in various HTGR gas cycle designs was tested to evaluate the however, the maldistribution of film coefficients flow characteristics of such a design. This model results in lower performance in the turning region. had an XT value equal to 1.25. The second mechanism can be much more crit- Reference 1 detailed the results of this investi- ical than the first. As the flow crosses the bundle gation. They are summarized here. in the radial direction, heat transfer between the The model represents a sector of the inlet region primary and secondary coolants causes a radial tem- of a recuperator or IHX. Figures 3 and 4 illustrate perature profile to develop in the shell-side fluid. this test model and a typical flow path. In bundles with flow starvation near the center, the The measured results from this test model were slope of this profile can be very steep. This profile compared against calculated results using a net- becomes the inlet temperature distribution into the work flow computer model (SNIFFS, Ref. 3). Fig- axial flow region of the bundle. This temperature ure 5 shows typical test and analytical results for maldistribution reduces the mean effective tem- this test model. The solid line represents the ana- perature difference between the primary and sec- lytical results, while the test data points are rep- ondary coolants, reducing the performance of the resented by circles and squares. Table 1 compares axial portion of the bundle. The combination of measured and calculated loss coefficients for three these two mechanisms can have a major effect on different inlet conditions, as tested by this model. the performance oi the heat exchanger. The loss coefficient is defined as the ratio of the 1.8 1° \D 1.4 - O MEASURED DATA: \ LEFTSIDE QF MODEL O MEASURED DATA: 1.2 RIGHT SIDE OF MODEL 0.8 : - \ D O \ o \ a > 0.6 o o \ RADIAL 0.4 - \ FLOW \ DIRECTION <- \ 0.2 - ANALYTICAL RESULTS i i 1 20 40 60 80 OISTANCE FROM INLET FACE (CM) A17351 Photograph of recuperator inlet air flow Fig. 5. Recuperator inlet air flow test: velocity test model profile, configuration 8 LOCATION OF PITOT-STATIC PROBES TOPOFTUBEBUNDLE 104 CM VARIABLE BLOCKED SUPPORT SCREEN BAFFLES GRID AIR FLOW A17351 Fig. 4. Schematic of recuperator inlet air flow test model TABLE 1 COMPARISON OF TEST AND ANALYTICAL LOSS COEFFICIENTS FOR RECUPERATOR/IHX TEST MODEL PRIMARY INLET COOLANT' DUCT HEXAGONAL Test Loss Coefficients SHROUD Configuration Test Analytical 6 125.6 158.7 8 32.1 33.6 ^BAYONET TUBE 9 100.2 104.4 BUNDLE AXIAL FLOW pressure drop to the average velocity head in the REGION HONEYCOMBED tube bundle. GRID B The following conclusions were made from the results of this test: OUTLET i C REGION X ! 1. The SNIFFS network flow model could be used to predict the flow distribution and pressure loss characteristics of this type of heat exchanger. PERFORATED 2. The small XT value results in flow starvation A17351 GRID at the center of the bundle. 3. A SNIFFS analysis of a complete heat ex- Fig. 6. AHE primary coolant flow path changer of this type showed that the flow maldis- tribution is very rapidly dissipated after leaving the 25 inlet region. However, the maldistribution rede- CURVE RAOIUS TEST velops in the outlet turn region. NO. (Ml DATA 1 0.022 0 Auxiliary Heat Exchanger. References 2 and 4 2 0.064 V 20 report an air flow test on a half-size AHE. This heat 3 0.108 A exchanger is typical of a straight tube configuration 4 0.149 a (Fig. 6), with a 90-deg outlet bend and a moderate 5 0.192 0 pitch-to-diameter ratio (1.43). For this test, the COMMIX measured data were compared against both a SNIFFS SNIFFS model and a turbulent flow model (COM- MIX, Ref. 5). Figure 7 and Table 2 show typical results of this comparison. 10 The following conclusions can be drawn from I these results: 1. Both the SNIFFS model and the COMMIX model gave adequate predictions of the flow dis- tribution and the pressure drop. 2. As compared to the recuperator/IHX re- sults, the larger XT value resulted in less flow star- vation at the center of the heat exchanger bundle. 0.3 0.2 0.1 0.1 0.2 0.3 DISTANCE ABOVE TOP DISTANCE BELOW TOP BOTTOM OF Straight Tube Superheater. The geometry of the OF EXIT SCREEN (M) OF EXIT SCREEN (M) BUNDLE STSH is similar to that of the recuperator/IHX. AI7351 However, the STSH design has a 2.13 XT value. This XT value is 1.7 times greater than the recu- Fig. 7. Radial velocity comparison test: SNIFFS, perator/IHX XT value. COMMIX TABLE 2 3. The inlet and outlet regions are treated as TYPICAL AHE PRESSURE LOSS cross-flow heat exchangers, with the shell-side fluid COMPARISON mixed and the tube-side fluid unmixed. The effec- tiveness of each node is evaluated separately. The calculated effectiveness is used to calculate the temperature change across the node. The output Loss Coefficient from one node provides input to its adjacent nodes. By using a stepwise process, the radial and axial Test data 4.85 temperature profiles are developed. 4. The axial region is treated as a counterflow SNIFFS model 4.15 heat exchanger. The counterflow effectiveness COMMIX model 4.77 equation is used to calculate the temperature change across each node. After each node, a gas mixing correlation is used to account for heat trans- fer between adjacent radial columns. A stepwise Flow testing of a one-sixth scale model of the process is used to develop both the radial and axial STSH is scheduled for 1985. Preliminary flow anal- temperature profiles. ysis of the STSH has shown a relatively flat velocity 5. The shell-side film coefficients in all three profile with very little flow starvation at the center. regions are based on input velocity data, which are The near-uniform flow distribution is the result of taken from either test data or analytical results the large XT value. from codes such as SNIFFS or COMMIX. In the The 1985 air flow test is anticipated to again axial region, the local velocity is used in the Dittus- confirm the use of SNIFFS and/or COMMIX for Boelter equation (Ref. 6) to establish local film flow analysis in straight tube heat exchangers. coefficients. The evaluation of film coefficients in the turning region presents a problem. No correlation for this Heat Transfer Performance type of flow is available. In the codes, the film coef- As previously described, the 90-deg inlet and ficients are estimated by using the two velocity outlet turns can have a major effect on the heat components for each node. The axial component is transfer performance of the bundle. used to calculate an axial film coefficient using the No known published data are available on heat Dittus-Boelter equation. A crossflow coefficient is transfer in a 90-deg turn across a tube bundle. With calculated using the radial velocity component and a lack of such test data and with no plans for such the Grimison correlation (Ref. 7) for flow across a a test, the only way to investigate the performance tube bundle. The weighted average of these two of such a configuration was concluded to be via coefficients is used as an estimate of the coefficient analytical means. in the turning region. In order to do this, a generic heat exchanger 6. In operating reactors, the primary coolant sizing and performance computer code was devel- temperature will not be uniform. These codes allow oped. Different versions of this code were developed the input of hot streaks. for the various types of heat exchangers (recuper- 7. The codes perform a tube-side flow balance. ator/IHX, STSH, and AHE). Although these var- The shell-side flow maldistribution, caused by the ious codes differ in some details, they all have the 90-deg turns, results in non-uniform temperature following features: distribution in the bundle; thus, the tube-side cool- ant has a radial temperature distribution, which 1. The heat exchanger is modeled as an axi- leads to more flow in cold tubes and less in hot symmetric (R-Z) model, which is divided into a tubes. The computer codes readjust tube-side flow number of concentric radial regions. Each radial so that all tubes have the same plenum-to-plenum region is divided into a series of axial nodes. pressure drop. 2. The axial length of the heat exchanger is The computer codes were used to evaluate the divided into three regions (two for the AHE). These performance of several straight tube configurations. regions include an inlet region, an axial flow region, Three different recuperator/IHX designs were and an outlet region. evaluated. Each of these designs had an XT value of 1.25. The recuperator/IHX configurations in- 300 cluded the following: 1. A large monolithic module without inlet 280 flow baffles. 2. A large monlithic module with inlet flow baffles. 260 MONOLITHIC MODULE 3. A cluster of seven small modules, which had NO BAFFLES. a total frontal area and a total number of tubes equal to the frontal area and a number of tubes for the monolithic module. 240 The AHE and STSH were also evaluated. All of the configurations were compared to the appro- 220 priate ideal counterflow designs which do not have MONOLITHIC MODULE 90-deg inlet and/or outlet turns. WITH BAFFLES 200 7-MODULE RESULTS CLUSTER From this study of three basic configurations, knowledge of the critical parameters which affect 180 the design of straight tube heat exchangers can be 0.2 0.4 0.6 0.8 1.0 developed. RELATIVE RADIUS, (Ro - R 4 )/(Ra - R,) Figures 8 through 13 illustrate the shell-side flow A17351 and temperature characteristics of the various con- Fig. 9. Calculated inlet temperature profiles for figurations. Table 3 compares the various recuper- three recuperator(IHX flow configurations ator/IHX designs to the ideal counterflow designs for the same heat duty requirements. Likewise, Table 4 compares the AHE and STSH to their re- spective ideal counterflow designs. 2.0 2.0 MONOLITHIC MODULE , NO BAFFLES 1.6 1.6 MONOLITHIC MODULE /TOP OF OUTLET WINDOW /"\ WITH BAFFLES 2 1.2 1.2 , BUNDLE MIDPLANE 7-MODULE < CLUSTER 0.8 0.8 > < 7.6 CM FROM WINDOW 0.4 - RADIAL FLOW DIRECTION J_ 1 0.2 0.4 0.6 0.8 1.0 0.2 0.4 0.6 0.8 1.0 RELATIVE RADIUS, IR0 - R £ )/(R0 - R|) RELATIVE RADIUS Fig. 8. Calculated inlet velocity profiles for three Fig. 10. AHE velocity distribution recuperator/IHX flow configurations 700 INLET RADIAL FLOW TOP OF 680 OUTLET WINDOW' INLET TEMP 30 CM FROM WINDOW EDGE 380 660 OUTLET TEMP 28 CM — 360 — FROMWINDOWEDGE 340 BOTTOM OF BUNDLE 320 600 OUTLET RADIAL FLOW 300 RADIAL FLOW 580 I I DIRECTION 0.2 0.4 0.6 0.8 1.0 RELATIVE RADIUS 280 A17351 0.2 0.4 0.6 0.8 1.0 RELATIVE RADIUS (R0-R,)/(Ra-R,) A17351 Fig. 13. STSH shell-side temperature distribution Fig. 11. AHE shell-side temperature distribution From the figures and tables, the following ob- servations can be made: —*- INLET RADIAL FLOW 1. The configurations with the small XT values are most easily affected by the 90-deg turns. How- 1.6 BUNDLE MIDPLANE I ever, the effect of the 9O-deg turn can be greatly INLET WINDOW reduced by considering certain design changes. 30 CM FROM EDGE These design changes include the use of flow baffles or the use of a small-diameter modular design. Flow baffles force the flow to the center of the bundle, thus reducing flow maldistribution. In small-di- —^ ~~ v \ ameter modular designs, the radial flow resistance is not large enough to allow maldistributions to de- OUTLET WINDOW 28 CM FROM EDGE velop. Both of these changes improve the flow and < 0.4 temperature distribution (Figs. 8 and 9), but at a -*— OUTLET RADIAL FLOW cost of pressure drop and/or envelope space. For the HTGR, both envelope space and pressure drop are critical to the design. 0.2 0.4 0.6 0.8 1.0 2. Configurations, such as the AHE configu- RELATIVE RADIUS ration, that have only an outlet turn in the tube (R0-R1)/(R0-R1) A17351 bundle are least likely to be affected by the turn. Even though flow and temperature maldistribution develops in the outlet region, the effect does not Fig. 12. STSH velocity distribution travel the length of the bundle; thus, it is minimized. TABLE 3 SIZING AND PERFORMANCE PARAMETERS FOR RECUPERATOR/IHX DESIGNS Monolithic Ref. Without Monolithic Cluster of Design Baffles With Baffles 7 Modules Required length, m 16.16 31.24 15.95 16.67 Pressure drop, kPa Shell side 13.69 44.43 73.09 15.07 Tube side 23.72 50.12 25.30 26.68 TABLE 4 SIZING AND PERFORMANCE PARAMETERS FOR AHE AND STSH AHE STSH Ref. Required Ref. Required Design Design Design Design Required length, m 3.92 4.06 15.25 16.34 Pressure drop, kPa Shell side 2.55 2.57 5.55 5.95 Tube side 522.6 541.2 130.8 136.8 3. Bundles with a large tube pitch-to-diameter ratio have the smallest effect on flow and temper- ature distribution (Figs. 12 and 13). However, if the effect occurs at the inlet (as in the STSH), this TABLE 5 small effect is propagated along the length of the TUBE TEMPERATURE VARIATION bundle, thus causing some loss of performance. In addition to the thermal performance degra- Temperature dation that occurs as a result of the 90-deg bends, Variation the resulting temperature maldistributions will re- sult in corresponding tube temperature variations. Table 5 shows that tube temperature variation cal- Recuperator/ IHX culated for the various configurations that have Monolithic without baffles 50 been investigated as part of this study. In designs where the tubes are attached at each end to fixed Monolithic with baffles 39 headers, tube temperature distribution could result Cluster design 13 in excessive tube stress. The possibility of excessive AHE 9 tube thermal stress must be part of the evaluation STSH 23 of straight tube designs with fixed headers. CONCLUSIONS REFERENCES Based on the results of the two air flow tests, the 1. Carosella, D. P., "Experimental and Analytical two analytical codes, SNIFFS and COMMIX, have Investigation of Shell-Side Flow Distribution Ef- been shown to adequately predict flow distribution fects on Axial Flow Heat Exchangers," Regenera- in the 90-deg turns. This conclusion will be further tive and Recuperative Heat Exchangers, ASME verified by the steam generator one-sixth scale air Symposium, Volume HTD-21, November 1981. flow tests. 2. Carosella, D. P., and P. N. Pavlics, "Comparison There are no test data to establish a heat transfer of Calculated Results from Two Analytical Models correlation for flow turning 90 deg across a tube with Measured Data from a Heat Exchanger Flow bundle. Testing to establish such a correlation is Test," ASME Paper 83-JPGC-NE-15. required before a general correlation can be devel- 3. Baccaglini, G. M., "SNIFFS: Single Phase oped to predict the effect of the inlet and outlet Nonisothermal Fluid Flow Simulator," GA Report turns on a straight tube heat exchanger. GA-A14501, June 1977. Lack of the necessary test data led to the use of 4. Kaufman, J. S., and M. M. Bressler, "Auxiliary analytical models to predict the heat exchanger Heat Exchanger Flow Distribution Test," ASME performance. Al though such models cannot exactly Paper 83-JPGC-NE-16. predict performances, they can provide data to 5. Sha, W. T., et ai, "COMMIX-1: A Three-Di- make possible the investigation of various straight mensional Transient Single-Phase Component Com- tube designs and the elimination of undesirable de- puter Program for Thermal-Hydraulics Analysis," sign features such as small pitch-to-diameter ratios Argonne National Laboratory Report ANL-77-96, and large tube thermal stress in Fixed header September 1978. designs. 6. Dittus, F. W., and L. M. K. Boelter, Engineer- ing, v. 2, University of California Publications, ACKNOWLEDGMENT 1930, p. 443. 7. Grimison, E. D., "Correlation and Utilization This work was funded by the San Francisco Op- of New Data on Flow Resistance and Heat Transfer erations Office of the Department of Energy under for Cross Flow of Gases Over Tube Banks," Trans. Contract DE-AT03-84SF11962. ASME 59, 583-594 (1937). 10 No. 16 GEFR-SP-314 DATE APRIL, 1984 TITLE: STATUS OF A REFORMER DESIGN FOR A MODULAR HTGR IN AN IN-LINE CONFIGURATION AUTHORS: R. GLUCK, W. H. WHITLING, AND A. J. LIPPS XA0055825 Prepared for presentation at IAEA SPECIALISTS MEETING ON HEAT EXCHANGING COMPONENTS OF GAS COOLED REACTORS Conference Held in DUSSELDORF, FRG City, State On 16-19 APRIL, 1984 Date This paper contains material resulting from work performed for U.S. DEPARTMENT OF ENERGY Under Contract No. DE-AC03-80ET 34034 This paper has been authored by a contractor of the U.S. Government under Contract No. DE-AC03-80ET 34034 Accordingly, the U.S. Government retains a nonexclusive, royalty-free license to publish or reproduce the published form of this contribution, or allow others to do so, for U.S. Government purposes. ADVANCED NUCLEAR TECHNOLOGY OPERATION • GENERAL ELECTRIC COMPANY SUNNYVALE, CALIFORNIA 94088 GENERAL W ELECTRIC 84-02-11 STATUS OF A REFORMER DESIGN FOR A MODULAR HTGR IN AN IN-LINE CONFIGURATION by R.GLUCK, W.H.WMTLING and A.J.LIPPS GENERAL ELECTRIC COMPANY SUNNYVALE, CALIFORNIA INTRODUCTION For the past several years the General Electric Company has had the technical lead on advanced concept studies for the Modular High Temperature Gas Cooled Reactor (HTGR) programs sponsored by the United States Depart- ment of Energy. The focus of the Modular Reactor System (MRS) effort is the development of a generic nuclear heat source capable of supplying heat to either a steam generator/electric cycle or a high temperature steam /methane reforming cycle. Some early ground rules for this study were that the reactor be designed for 950°C direct cycle reforming and that the core be of the prismatic type. Since the prismatic core required control rods near the center of the core, the vertical in-line concept was selected to promote natural circulation cooling of the core for all potential transients except the depressurized core heatup transient. Although the requirement for a prismatic core has been eliminated for recent cost reduction studies, the vertical in-line configuration has been retained for its potential as the lowest cost configuration. This paper presents the results of recent design and analytical studies conducted to evaluate the feasibility of using a steam/methane reformer in a Vertical In-Line (VIL) arrangement with the generic nuclear heat source. DESIGN REQUIREMENTS The requirements for the study of the VIL reformer are presented below in two broad categories. The first establishes those requirements arising from the overall thrust of the MRS program, while the second establishes the component functional requirements for the steam reforming process. MRS PROGRAM REQUIREMENTS The MRS program has as its primary goal the development of a generic nuclear heat source capable of being utilized either in a steam/electric cycle operating'mode or a steam/methane reforming mode. In arriving at this peal the project has put primary emphasis on achieving passive safety and reducing construction costs by: a - utilization of a standardized modular design approach to minimize site construction activity, b - utilization of an in ground silo for housing and supporting the pressure vessel, c - providing passive decay heat removal capability. Preliminary evaluations showed that a vertical arrangement in which a steam generator or steam reformer mounted directly above the reactor vessel would have many desirable features towards achieving the objectives. The follow- ing programmatic and functional requirements were established for the VIL reformer: 1 - Streamlined reactor vessel configuration with the reformer mounted in the pressure vessel vertically above the pebble bed core. 2 - Decay heat removal to be by natural or forced circulation of primary coolant gas to the main cooling systems under normal and upset conditions and by natural circulation of primary coolant and natural circulation of external cavity coolant under emer- gency conditions. [Normal, upset and emergency conditions are defined in terms of frequency and probability for ASME Code design purposes.] VIL REFORMER REQUIREMENTS The detailed performance and design parameters for the VIL reformer are based on overall plant system analyses, i.e. heat balances, fuel performance, cere heat-up analyses, etc.. These requirements are summarized in Table 1. REFERENCE DESIGN A typical arrangement of a VIL-MRS is shown in Figure 1. The reactor has a pebble bed core surrounded by graphite reflectors in the bottom half of the reactor vessel. The reformer is mounted above the core and directly supported by the upper half of the lower pressure vessel. Fuel loading, circulator mounting and control rod penetrations are all accomplished through the reactor pressure vessel structure. The reformer is a baffled shell and tube heat exchanger using bayonet tubes attached to tubesheets as shown in Figure 2. The vessel is constructed as a cylindrical shell 4.76 meter (187.5 inch) in diameter, flanged at each end. The lower section has a conical transition to the flange to enable mating with the reactor pressure vessel. The upper flange mates with an elliptical bolted head that provides closure of the pressure boundary. All access to the vessel internals for installation, inspection and maintenance of the reformer and steam generator is accomplished through this closure. Penetrations are provided through the cylindrical portion of the vessel for piping connections of the process gas inlet and outlet, buffer helium, feedwater inlet and steam outlet. The vessel shell also has a ring ledge for supporting the reformer internal structure and brackets for supporting the helical coil steam generator. The reformer internal structure consists of inlet and outlet process gas plenums, each connected by means of bellows expansion joints to the outer vessel nozzles. The inlet plenum is connected to both the upper and lower tubesheet while the outlet plenum is connected to the upper tubesheet only. The reformer tubes consist of an outer tube 64.5 mm (2.54 inch) in diameter, 10.2 m (33 ft. - 4 in.) long with a monolithic type of reformer catalyst and an inner return tube. The outer tube is attached to and supported by the lower tubesheet. The inner return tube supports the mono- lithic catalyst and is attached to and supported by the upper tubesheet. The plenums, in addition to providing flow separation, must provide the structural flexibility to accommodate the differential thermal growths resulting from the differences in inlet and outlet process gas tempera- tures. This is accomplished by providing sufficient length to the plenum cylinders. Directing the process gas flow from the pressure vessel nozzles to the inlet and outlet plenums is accomplished by utilizing flanged, bolted in sections having bellows expansion joints. The use of the bellows joints provides a means for accommodating the differential thermal growths of the reformer structure relative to the pressure vessel. The flexibility of the bellows joint also allows the tolerances for manufacture, assembly and replacement to be generous. The use of the bolted flanges provides e means for inspection and replacement of internal structural elements. The reformer tube bundle is shrouded by a cylinder 3.13 meter (123.1 inch) in diameter, connected to the lower tubesheet. This cylinder inter- faces with internal reactor flow distribution baffles to direct circulating core helium to the reformer inlet and to direct the return path of the helium. Baffle plates are provided on the shroud to direct the helium flow in a cross flow pattern and to provide lateral support to the reformer tubes against flow induced vibration and seismic accelerations. At the top of the shroud, the helium under cross flow conditions, exits at one side only. An eccentric distribution baffle is provided to uniformly distribute the flow into the steam generator section. The entire reformer structure is supported by a conical support skirt between the lower tube sheet and the outer vessel. All internal surfaces of the outer vessel above the top of the steam generator, the reformer support cone, and the underside of the lower tubesheet are insulated to minimize temperature gradients and/or long term exposure to elevated temperature conditions. The lower end of the insula- tion will form a juncture with the steam generator shrouding. FLOW ARRANGEMENT AND LOADINGS Under normal operating conditions, helium flows upward through the core and top reflector and exits at 950 C (1742 F). Flow is then directed by internal baffling to the shroud encompassing the reformer tubes. The helium after passing through the reformer tube bundle exits the tube shroud at 625 C (1158 F) and flows down the annul us between the reformer and pressure vessel to a helical coil steam generator. Helium exits the steam generator at 350 C (660 F) and is returned to the core inlet. The steam generator tubes are enclosed in their own shroud, which forms an annulus with the pressure vessel wall. This annulus is closed at the top to prevent flow from bypassing the steam generator. This creates a "stagnant" helium buffer between the steam generator and the pressure vessel wall which limits the heat transfer to the vessel wall to natural circulation convection in the annulus and radiation from the steam generator shroud. The vessel wall under normal conditions is calculated to be 393 C (740 F) on the inside surface with a mean wall temperature of 371 C (700 F). The design temperature of the vessel is established at 400 C (750 F). The corresponding process gas flow during normal conditions is ducted to the inlet plenum where it flows down through the nickel catalyst in the reformer tubes and returns through the inner tube to the outlet plenum. The inlet process gas enters the reforming tubes at 482 C (900 F) and is heated by both the reactor helium on the outside of the tube and by the process gas in the return line to a maximum of 820 C (1508 F). The process gas exits the return line to the outlet plenum at 596 C (1105 F). During normal operating conditions the reactor helium and process gas are maintained at pressure such that the maximum differential pressure load on the reformer components is 0.42M Pa (60 psi). The differential design pressures for the lifetime loading (300,000 hour) of the reformer compo- nents are shown on Table 2. During upset and emergency transient conditions the reformer structure is generally exposed to both increased temperature and increased differential pressure loads for short time loadings (< 100 hour). The actual service temperature, pressure and duration are functions of the transients imposed. Maximum pressure loads occur as a result of depressur- ization of coolant on one side of the boundary while pressure is maintained on the other. Typically these are events such as core depressurization or process line rupture accidents. An envelope of the differential pressures that can exist on the reformer structures for these type events is also shown on Table 2. These short term loading events generally govern the size of the structural elements of the reformer. THERMOCHEMICAL PERFORMANCE Analysis for the steady state reforming behavior of the steam reformer bundles was performed using the DSRDSRN computer code (Ref. 1). This code uses an empirical chemical kinetics model for the steam methane reaction and assumes that the water shift reaction is at equilibrium. An iterative solution is employed in order to calculate heat transfer rates and chemical compositions in an axial model of an average reformer tube in a tube bundle. The accuracy of this code was verified for single tubes in tests conducted at KFA in the EVA-I test rig. DSRDSGN is a modification of the DSR1 code (Ref. 2) that incorporates the improved crossflow heat exchange resulting from the inclusion of the baffles. Each baffled flow region is assumed to be characterized by the average helium temperature for that region. The pressure drop across each baffled region was calculated in three steps; the frictional pressure loss across the tubes, a 180 degree turn from one region to the next, and an orifice pressure drop as the helium passes between the end of a baffle and the bundle shroud. Heliun leakage flow through the reformer tube penetrations in the baffle plates causes a reduction in the heat transfer coefficient from true cross flow. This effect was included in the serpentine flow heat transfer correlation as a 10% reduction in heat transfer coefficient (Ref. 3). This helium leakage also results in a 20% reduction in helium pressure drop a- cross the reformer tubes and baffles. The performance requirements for the reformer tube bundle were shown on Table 1. The tube bundle design geometry selected to meet those Jf} requirements consists of 1039 64.5 mm (2.54 inch) outside diameter tubes with an active reforming length of 10 m (32 feet 9.7 inch). The tube wall thickness required to satisfy structural requirements is 15.2 mm (0.6 inch). Significant results of the analyses with the DSRDSGN code are shown on Table 3. Typical conditions as they exist along the active length of the reformer tube are shown on Figure 3. DECAY HEAT REMOVAL Decay heat from the reactor core is, under normal shutdown conditions, removed by continuation of coolant flow to both the reformer and steam generator tubes with forced circulation of reactor helium. The Vertical In-Line Arrangement, however, also provides a is that the vertical height separation between the core and the heat exchangers giving a thermal driving head capable of sustaining core cooling under natural circulation conditions. This feature enables removal of decay heat without reliance on the circulator or heat exchangers during abnormal events. This is accom- plished using either the heat exchangers or the vessel cavity cooling. The various modes of decay heat removal are summarized in Table 4. Analyses have shown that a reactor trip coupled with decay heat removal by natural circulation using the steam generator as a heat sink does not result in severe transients for the reformer steam generator steam generator or reactor structures. Except for the reformer tube temperature, the maximum value for all key parameters occurs at the design point. Isolation of reformer cooling after reactor trip results in a severe transient for the upper portion of the reformer tubes. Even more important is the extreme temperature gradient across the lower reformer tubesheet during these transients. When no reformer cooling is available, the steam generator is also subjected to a severe transient. To reduce the severity of this transient, several reformer cooling options were investigated. The use of an auxiliary boiler to provide steam cooling to the reformer was evaluated. Decay heat removal transients with reformer cooling present showed that the reformer, steam generator and pressure vessel transient problems were mitigated. The vessel design pressure and temperature are maintained within design condition values. Decay heat removal without reformer cooling results in heat up of the reformer tubes and tube sheets to elevated temperature extremes, 950 C (1742 F) for limited time durations. The severity of the temperature, duration and number of cycles varies with the concurrent condition of the steam generator availability for heat removal; the most severe condition occurring without steam generator cooling. This condition also results in peak vessel wall temperatures. Decay heat is removed by the dry steam generator structure rising in temperature to reject heat to the vessel wall by radiation and natural circulation in the annulus, and rejection of the vessel wall heat to the cavity cooling system. Below the steam generator the vessel wall also increases in temperature to reject heat to the cavity cooling system. Structural analysis of the reformer have concentrated on this condition and results are presented below. For cool down following a helium depressurization accident and loss of circulator, heat is removed from the core radially by conduction and radiation through the reflectors, core barrel, vessel wall and finally to the cavity cooling system, with no reliance on the heat exchangers. STRUCTURAL EVALUATION The structural evaluations conducted have been limited to establishing structural feasibility rather than a fully comprehensive analysis as required for ASME code certification. This was done using stress equations and techniques as outlined in Appendices 2000 through 8000 of Section III of the ASME Code, that utilize pressure and temperature differentials in the metal to evaluate primary and secondary stress responses. All analyses were based on elastic methods. No attempt was made to use the more sophis- ticated finite element technique that would be used for a final analysis. Table 2 gives the pressure differentials analyzed and Table 5 is a summary of the major transient conditions considered. The thermal response of the metal to these transients was analyzed using a finite difference computer code, SINDA (Ref. 4) and the response to these transients at a typical location, i.e. center of the lower tubesheet is shown on Figure 4. This is provided to give an indication of the thermal severity of the transients at an arbitrarily selected location. The actual stress state is a function of the variation in temperature throughout the region including mismatch with other regions at different temperatures. The various regions analyzed together with the primary (PI ,Pb) stress and secondary (Q,F) stress results are shown in Table 6. The complete stress analysis is given in Appendix A of Reference 5. The associated material property curves used to deduce the material allowables are shown on Figures 5 and 6. From these it is seen that the material data at high temperature are uncertain and extensive extrapolations are used for both Alloy 800 and Inconel 617. Pending confirmation of this data the material selections for the reformer tubes are not yet firm. Table 6 shows the various failure modes that have been considered for the selected locations. These include ductility limits for primary stresses, stability limits, and strain range control associated with low cycle fatigue. This is not the fully comprehensive set of conditions as required by the ASME for qualification of a final design, but it is consi- dered sufficient in the preliminary phase to show structural feasibility. On these bases, the structural margins contained in Table 6 are shown to be positive, and it is expected that, following confirmation of a material data basis, and completion of a full Code analysis this conclusion should not chanoe. OPERATION AND MAINTAINANCE The design of the reformer shown in Figure 2. is, because of balanced pressure between the helium and the process gas, able to be designed using bolted joints while retaining confidence against leakage. The use of bellows expansion points also minimizes loadings that would tend to unload sealing surfaces, and as noted earlier, allows for generous manufacturing and assembly tolerances. The free use of bolted joints enables access to many structural elements for in-service inspection and replacement of sub-assemblies if necessary. FUTURE DEVELOPMENT The design and analysis of the VIL reformer has progressed sufficiently to be considered a feasible option in future applications of the modular HTGR program. Near term plans are to continue developing the design with supporting stress analyses in the creep-fatigue regime. For the efforts to be meaningful, parallel efforts concentrating on obtaining elevated temperature properties of candidate reformer tube and tube sheet materials along with development of failure models are required. REFERENCES 1. Meyer, D. J., "Description of the DSRDSGN Code: Steam Reformer Design Code," General Electric Company Report GEFR-OO535, August 1980. 2. Meyer, D. J., Parker, K. M., "DSR1 Analytical Computer Program," DSR 2841-6.l-(2), January 1980. 3. Fraas, A. P., Ojisik, M. N., Heat Exchange Design, John Wiley & Sons, 1965. 4. SINDA Users Manual, J. F. Smith, NASA Contract 9-10435, TRW Report No. 14690-H001-R0-00, April, 1971. 5. "MRS Recomrnendations and Evaluation of a Vertical In Line Reformer," Gluck, R., Whifling, W. H., Hill, R., Unpublished, Feb. 1984. TABLE 1 EL. 24.4m - DESIGN REQUIREMENTS FOR A 250 MW(t) (BOh-din.) VIL-MRS REFORMER HEAT BALANCE TOTAL REACTOR HEAT AT 100% POWER - MW{0 250 PROCESS GAS OUTLET-- -- PROCESS GAS INLET REFORMER HEAT TRANSFERRED - MWIl) 135 STEAM GENERATOR HEAT TRANSFERRED - MWIl] 115.8 SYSTEM LOSSES • MWW 3 EXTERNAL VESSEL CAVITY CIRCULATOR INPUT MWft) 3.8 MAIN STEAM OUTLET-- — FEEDWATER INLET REFORMER REQUIREMENTS PRIMARY HELIUM REFORMER INLET TEMPERATURE C (F) 950(1742) REFORMER OUTLET TEMPERATURE • C (F) 625(1158) REFORMER INLET PRESSURE • MPa (psia) 4.137 1600) PRESSURE DROP • MPa Ipsil FLOW RATE - kg/s (Ib/M PROCESS GAS STEAM/METHANE RATIO 4:1 METHANE CONVERSION RATIO 60% INLET TEMPERATURE C IF) 482 1900] OUTLET TEMPERATURE -C(F) 596 11105) INLET PRESSURE MPa Ipsia] 4.447 (645) PRESSURE DROP • MPa (psi) 0.31 (45) FLOW RATE - kg/s (Ib/hrl 79.0 (627400) REACTOR VESSEL CRDMREMOVAL CELL EL.-25.9m - <-85fl-0in.) Figure 1. VERTICAL IN-LINE MODULAR REACTOR SYSTEM RUPTURE DISKS (2)- INSULATION OUTLINE STEAM OUTLET TABLE 2 BAYONET REFORMER TUBES DIFFERENTIAL PRESSURE LOADS ON [1039) INTERNAL REFORMER STRUCTURES STEAM GENERATOR NORMAL CONDITION ACCIDENT CONDITION TUBE BUNDLE INTERNAL DIFFERENTIAL DIFFERENTIAL COMPONENT MPa (psi) MPa (psi) INLET PLENUM 0.03 ( 5) 4.38 (635) OUTLET PLENUM 0.35 (50) 4.3B 1635) UPPER TUBE SHEET 0.31 (45) 0.35 1 50)* LOWER TUBE SHEET 0.3 155) 5.52 1800) INTERNAL SUPPORT CONE 0.42 160) 4.38 (635) REFORMER TUBE 0.38 (55) 5.52 (800) TUBE BUNDLE SHROUD 0.10(15) 0.10 1 15) -3.13 m diam—*- \ ;: ,' (123.1 in) -t— REACTOR PRESSURE 'Limited bypass rupture disks :;, ..• I \ VESSEL * PRIMARY •'• HELIUM Figure 2. VERTICAL IN-LINE REFORMER TABLE 3 ENERGY BALANCE TOTAL ENERGY TRANSFERRED FROM HELIUM 135 MW I I I I I I I I I ENERGY STORED IN CHEMICAL REACTION 109.3 MW SENSIBLE HEAT ENERGY 25.7 MW 1000 - 80 PROCESS GAS IN PRESSURE DROPS INNER RETURN —v PRIMARY HELIUM, MPa (psi) .104 (15) 900 - TUBE \ - PROCESS GAS-REFORMER TUBE, MPa (psi) .170 (25) RETURN TUBE, MPa Ipsi) .118(17) 60 | CONVERSION ACHIEVED >t PROCESS GAS IN ^__3i— ' 60% JJ UJ REFORMER^ < ^== AVERAGE HEAT TRANSFER COEFFICIENT §700 - TUBE \^""^ ^^- 50 z REFORMER TUBE 1036W/m2-K(182BTU/Hr-ft2-FI - O 2 2 RETURN TUBE 785W/m -K(138BTU/Hr-fl -F) J 600 - > / u *- H 500 - 30 ! 400 - - 10 / I I I I I | I I I TABLE 4 01 23456789 10 VIL-MRS MODES OF DECAY HEAT REMOVAL ACTIVE REFORMER TUBE LENGTH - METERS STEAM HEAT REMOVALS MODE CIRCULATOR REFORMER GENERATOR SYSTEM TRANSIENT DESIGNATION Figure 3. DSRDSGN CODE ANALYSES RESULTS # * * A - REACTOR SYSTEM PRESS RIZEO FORCED CIRCULATION WITH MAIN 1 X X X LOOP COOLING THROUGH STEAM GENERATORS. REFORMER FORCED CIRCULATION WITH STEAM 2 X X X COOLING THRU REFORMER NATURAL CIRCULATION WITH MAIN 3 X X X LOOP COOLING THRU STEAM GENER- ATOR & REFORMER NATURAL CIRCULATION WITH HEAT 4 X X X REJECTION THROUGH STEAM GENER- ATOR (B17& C7I NATURAL CIRCULATION IN REACTOR VESSEL. PASSIVE HEAT REJECTION 5 X X X BY RADIATION AND NATURAL CIRCU- LATION FROM PVWALL TO VESSEL CAVITY B - REACTOR SYSTEM EPRESSURIZE o 1 X X X SAME AS A-1 RADIAL CONDUCTION FROM CORE TO 2 X X X LOWER VESSEL WALLS. PASSIVE HEAT REJECTION AS A-4. = OPERATIONAL = NON -OPERATIONAL TABLE 5 UMBRELLA TRANSIENTS FOR STRUCTURAL EVALUATION CLASSIFICATION DESCRIPTION AND NUMERICAL CONDITIONS COMMENT DESIGNATION ISOLATION AND SLOWDOWN iiP - 700 psi (EXTERNAL) OF STEAM GENERATOR. UPSET Tm3x SG * 1150°F lor REFORMER AND CIRCULATOR B11 160 HRS. NO. OF CONTINUE TO OPERATE. OCCURRENCES N = 20 DEACTIVATION OF THE APPLIES FOH 10% OF REFORMER CATALYST. UPSET REFORMER LIFE. 1 STEAM GENERATOR, REFORMER, B16 AP =G5p!iTTUBEWALL = CIRCULATOR ALL OPERATING. 173S°F for 14700 HRS. ISOLATION OF REFORMER LOWER T/S ^P-330p.iTTUB£SHEET. TEMPERATURE IS PROCESS GAS. UPSET SUPPRESSED BELOW STEAM GENERATOR FUNCTIONS. B17 1600°F* 100 HRS ABOVE 1600° F BY SUPPRESSING CIRCULATOR IS TRIPPED. 1200°F N » 20 NC IN TUBESHEET ISOLATION ANDBLOWDOWN 4P-330piiT = OF STEAMGENERATOR WITH EMERGENCY TS 176d°F FOR 520 HRS. AS ABOVE REFORMER ISOLATION C2a N-7 CIRCULATOR TRIPPED. PROCESS GAS LINE RUPTURE. AP-710piiTTS- PRECLUDES ISOLATION OF EMERGENCY 1350°F FOR 32 HRS. AS ABOVE REFORMER, FOLLOWED BY C7 TWALL " "S^F FOR REACTOR TRIP. 2HRSN " 7 * THESE CONDITIONS WERE FOUND TO 8E UNACCEPTABLE STRUCTURALLY IN THE LOWER TUBESHEET AND THE CERAMIC INSERTS WERE EXTENDED FROM THE REFORMER TUBES INTO THE TUBESHEET REGION THEREBY REDUCING THE NATURAL CIRCULATION HEAT INPUT TO THE TUBESHEET. °c 1000 TEMPERATURE - °C 427 538 649 760 871 982 1 1 1 1 1 1 1 1 W ' 1 /— 300,000 HRS 20 — 100,000 HRS //~ 1600 — 10,000 HRS 1,000 HRS 16 r- Sm '— 100 HRS 10 HRS 12 800 8 - . 1400 — 4 ~ ALLOY BOOH B17 (WITH NATURAL CIRCULATION) E 700 IN TUBE SHEET 0 1 1 1 1 1 1 1 ?"""~ r- — ^n*— UJ ECL UJ l- 1200 — III)) I I J s I 1 U3J a: s y—300.000 HRS ~ t- 20 — m ~—_ /,— 100.000 HRS ~^~ ///— 10,000 HRS 16 \VV//— 1,000 HBS B17 IVWO NATURAL CIRCULATION) Vrvv \X/\//. 100 HRS 1000 — IN TUBE SHEET 12 v\ VAA/XV— 10 HRS 3 A INCONEL617 400 0 I I I I I I 1 1 ^ T~ 10 15 20 1200 1400 1600 TIME-HOURS TEMPERATURE - °F Figure 4. LOWER TUBESHEET AVERAGE <£. TEMPERATURE Figure 5. SMT FOR ALLOY 800H AND INCONEL 617 I ! /— 538C (1000FI / — 760C (1400F) //— 871C (1600F) / ,— 982C (1800F) • • 1 ALLOY 800H ~~ —• I I i I IU.U I I 1 1 /— 538C (1000F) /.— 740C (1300F) 1.0 ///—815C (1500F) i ^—. 0.1 ^^ INCONEL 617 0.01 1 1 I 1 102 103 104 105 106 Nf-CYCLESTO FAILURE Figure 6. DESIGN FATIGUE CURVES FOR INCONEL 617 AND ALLOY 800H 10 84-176-02 TABLE 6 STRESS ANALYSIS RESULTS LOADING STRUCTURAL VALUE ALLOWABLE MARGIN PARAMETER EVENT NORMAL IA-3) 239 psi 570 psi 1.38 REFORMER P UPSET (8-16) 305 psi 720 psi 1.36 m 04 TUBES INSTABILITY EMERG. (C-7) 710 psi 72987/2.5 = LARGE 800H 29195 psi PL + PB + Q+F ALL TRANSIENTS 0.18 1.0 4.56 DAMAGE FACTOR) PRIMARY MEMBRANE STRESS AT 1735°F 1.51 617 MARGIN > 800H MARGIN REFORMER (A-3) 239 psi 600 psi TUBES (B-16) 305 psi 800 psi 1.62 (C-7) 3332 psi 4830 psi 617 INSTABILITY LARGE (DAMAGE FACTOR) LARGE S P m EMERG. (C-7) 10956 psi Smt @950°F and 0.33 TUBESHEET LOWER T/S 300Hrs = 14,600 psi 0.84 RIM/HOLE INTERACTION SUPPORT PL + PB + Q (C-7) 23782 psi CONE 800 H 43,800 psi P|_ + Pg + Q+ F LARGE p + P L B PRESSURE ONLY 2403 psi Smt PL + PB + Q+ F 37,500 psi CONSERVATIVE BASE ON USE OF "HOLD TIME" Pm NORMAL 6490 psi 13,400 psi 1.06 FATIGUE CURVES UPPER T/S P EMERG. (C-7) 11800 psi 16,560 psi 0.40 SUPPORT m CYLINDER INSTABILITY (C-7) 200 psi 1692/2.5 = 2.38 677 psi Pm NORMAL 1740 psi 13,400 psi 6.70 Pm EMERG. (C 7) 15950 psi AP REDUCED BY USE OF RUPTURE DISKS CYLINDER INSTABILITY (C-7) 550 psi 4794/2.5 = 2.49 1917 psi Pm EMERG. (C-7) 14575 psi 17,160 psi 0.18 HEAD INSTABILITY (C-7) LARGE 84-176-01 No. 17 - 1 - XA0055826 Development and Fabrication of a Helium-Heated Steam Reformer W. Panknin, W. Nowak L. & C. Steinmiiller GmbH Gummersbach, F.R. Germany 1. Introduction The development of a helium-heated steam reformer for the Prototype Plant Nuclear Process Heat (PNP) or for a High- Temperature Reactor Module Plant must satisfy the specific requirements connected with nuclear components. Some of the main aspects will be mentionend: Because the steam reformer is part of the barrier between primary helium and secondary process gas, all parts of the primary closure must be accessible without opening the primary circuit, and in-service inspection should be possible. Leakage of the primary gas into the secondary circuit, due to the failure of component parts, must be impossible. This is accomplished by maintaining a slight pressure difference of 2 bar between the two circuits. Gas streaks in the primary helium with a temperature difference of about + 20 K in comparison to the average gas temperature are expected to occur. The resultant effects must be considered and kept under control. The layout of the steam reformer must be designed against the impacts of external events such as seismic loads. - 2 - The loss of the secondary heat sink ought not result in any major damage. The whole tube bundle must be replacable. Replacement of the catalyst must be simple and fast. The design lifetime of the component is 140.000 hrs. All these requirements are more stringent than those for conventional steam reformers or other comparable heat ex- changers. Therefore, a new concept had to be worked out first, before beginning with the layout and design of the component. Here our main objective was, to separate those assemblies which are exposed to high in-service strains or high temperatures into structures of a simple geometry and to assign only one function to each element. This way we found a concept and design with such a low stress level in all its relevant parts that the licensing procedure of the stress analysis is expected to be fea- sible. The steam reformer for PNP will have a heat transfer ca- pacity of about 96 MW. The bundle will consist of roughly 300 tubes. In order to test the new concept and design ex- perimentally, a steam reformer of smaller capacity (5 MW) with only 18 tubes was designed and is being fabricated at the moment. This bundle will be tested in the EVA II pilot plant at KFA Julich. Because this bundle is comparable in almost all important details to the large steam reformer, the flow path and the design, as well as some interesting manufacturing steps will be explained with respect to this test steam reformer. - 3 - 2. Description of the Design 2.1 Primary and secondary ciruits A longitudinal section through the test steam reformer illustrates the primary and secondary circuits (see fig. 1). Primary helium enters at 950 °C and 40 bar from the reac- tor core through the hot gas duct. It flows upwards on the shell side of the reformer tubes and is cooled to 700 °C. Below the insulated tube sheet the primary helium is taken to the steam generator. Having cooled to 300 °C, the helium is returned through the concentric annulus back to the reactor. In this way, the walls of the reactor pres- sure vessel are kept at a low temperature. In order to obtain the heat transfer coefficient on the helium side resulting from the layout, the reformer tubes are shrouded by guide tubes. Therefore, a concentric annulus of 9 mm width results around each reformer tube. To separate the entrance and exit regions of the helium, a partition plate is located in the upper part of the bundle. The process gas circuit has its entrance and exit at the top of the steam reformer. The process gas, a steam/meth- ane mixture at about 330 °C, enters the process gas cham- ber and is distributed among the reformer tubes. It cooles the support plate before being preheated in a small re- cuperator which is integrated in each reformer tube (see fig. 2). Hereafter it enters the catalyst region, is heated up to about 810 °C by the hot helium stream and is simultaneously reformed to a synthesis gas with a high content of hydrogen. - 4 - This synthesis gas is cycled back through a return tube, cooled down to 460 °C in the recuperator and leaves the steam reformer through the product gas chamber. 2.2 Reformer Tube Bundle As already mentioned, the tube bundle of the 96 MW com- ponent consists of about 300 tubes, 120 0 x 10 x 17.000 mm. The test bundle has 18 tubes and one inspection tube. They are arranged in a triangular pitch of 200 mm. 13 tubes are made out of Inconel 617 (2.4663). For comparison reasons 5 tubes are made out of Incoloy 800 H (1.4876 H). Because such long tubes cannot be manufactured in one step, shorter tubes, about 2.500 mm long are welded togehter (see fig. 4). Then they are bored to the final internal diameter and the outer diameter is machined to the correct size with the aid of a special lathe. The tubes are welded to the underside of the support plate, which is manufactured of material 10 CrMo 910. This material has to be annealed after welding. Because of the tube length this is not practicable for the whole bundle. Therefore the plate is deposition welded first with the corresponding tube material. Then the plate is heat treated and finally the tubes are welded to it. Hereafter no further heat treatment of the plate is necessary. At the level of the partition plate the guide tubes are welded to the reformer tubes so that they can freely ex- pand downwards in the axial direction (see figs. 3 and 5). The guide tubes with an outer diameter of 144 mm have a wallthickness of only 3 mm. They are manufactured by a special flow forming process which produces tubes to very close tolerances (see fig. 10). In addition this is a material-saving process. — 5 — To maintain the annulus between reformer and guide tube, spacers are placed every 2 m (see fig. 3). The spacers and the contacting surfaces of the reformer tubes are coated with a protective layer against fretting in the helium atmosphere. The tubes are freely suspended from the support plate and held in position by only one more plate, called tube guide plate (see figs. 1 and 2). Through this guide plate at the lower end of the reformer they can expand during opera- tion. The sealing between the inlet and outlet helium flow is achieved by a partition plate as mentioned above. Ra- dial and axial displacement of the tubes is possible, be- cause of the use of expansion joints connecting the par- tition plate and the guide tubes (see figs. 3 and 7). There are some advantages of the new concept in comparison to other wellknown systems. One important characteristic of the design is that the reformer tubes are kept in position by only two elements which are as far apart as possible. Therefore a certain tube distortion is accept- able and the reaction forces between plates and tubes, as well as the bending stresses are minimized. No unallowable stress or even deformation of the tube guide plate or the tube itself is expected to occur. Another advantage of this design is the fact that guide tubes are used for the helium flow. Firstly, this design does not have an undefinable boundary flow as can usually be found in similar components. Secondly, the blocking of individual tubes with a resulting local temperature change will cause no unallowed interference to the adjoining tu- bes. Therefore steam reformers of different capacity can be built by simply chosing the correct number of tubes. In addition no overall heat transfer measurements are needed because of the same regular and simple flow in each of the annuli. - 6 - 2.3 Integrated Recuperator and Catalyst The recuperators which are integrated in each reformer tube are helix coil type (see figs. 8 and 9). They consist of three coils, 0 18 x 1.5 mm arranged around a central displacement tube. The mean diameter of the coils is only 78 mm. This design combines the qualities of good compen- sation of thermal expansion with a low pressure drop and relatively short length. The recuperator can be used in two modifications depending on the type of catalyst to be tested. Two different catalysts, the Raschig ring bed and a newly developed disk catalyst, are proposed. When Raschig rings are to be used the three coils are extended into three return tubes which reach down to the lower end of the reformer tube. The Raschig rings can be filled in after the central dis- placement tube is opened. For the replacement it is planned to use a new hydraulic process, developed by KFA Jiilich, which promises a replacement of the whole bed within only a few minutes. The second type of catalyst consists of two types of disks. The disks are alternately piled up on a straight return tube, which is connected to the recuperator. This arrangement also allows easy and quick replacement by simply exchanging the complete unit. This could already be verified in the EVA I plant. Fig. 11 shows the catalyst column before being inserted in the tube. — 7 — 3. Outlook As already said, the test bundle is in production at the moment. It is planned to start the experiments at the end of this year. The whole test period will last between 1 and 2 years, thereafter the test bundle will be removed and subjected to several extensive checks and post-experiment investiga- tions . With the experience already gained by the fabrication and testing of various reformer tube configurations, as well as with the enlarged knowledge from the design studies, the development standard should allow the licensing and subsequent construction of a steam reformer for a nuclear power plant. product gas chamber product gas process process gas gas chamber support plate insulation recuperator primary helium partition plate reformer tube tube guide plate primary helium ~ Fig. 1: Steam Reformer (Test-bundle for EVA II) Fig. 2: Reformer tube 1 Separation plate 2 Tube sheet, (support plate) 3 Reformer tube 4 Recuperator 5 Expansion joint 6 Partition plate 7 Guide tube 8 tube guide plate 9 Spacer 10 Internal return tube 8 reformer tube Spaltrohr Kompensafor expansion joint Zwischenplatfe partition plate Festpunkt fix point 5Dist-anzhalter 5 spacers guide tube Hullrohr Fig. 3: Reformer tube (helium side) Fig. 4: Welding of the reformer tubes Fig. 5: Reformer tube - guide tube connection element Fig. 6: Deposition welding on the tube sheet HP iili Fig. 7: Expansion joints (between partition plate and guide-tube) Fig. 8 and 9: Recuperator •Fig. 10: Flow forming of guide tubes Fig. 11: Disk catalyst for EVA I experiments No. 18 Assembly and Operation Experience XA0055827 of EVA II Steam Reforming Bundle H.F. Nieften, R. Harth Kernforschungsanlage Jiilich GmbH, Jiilich FRG W. Kesel Rheinische Braunkohlenwerke AG, Koln, FRG The main test component of the experimental facility EVA 11/ ADAM II is a helium heated steam reformer bundle with 30 tubes. The dimensions of the tubes are 120 mm OD/100 mm ID and a. heated length of 11 m. The tubes are arranged in the tube sheet in a triangular pitch, and the diameter of the bundle is 1.2 m. To achieve a better heat transfer on the helium side, there are baffles (disk and doughnut) arranged every 250 mm, resulting an intensive cross counterflow. The helium entering the lower end of the bundle with a temperature of 950 CC flows upwards, and is cooled down to a temperature of 600 - 650 °C. After- wards the helium leaving the bundle is guided to the steam generator. The bundle is surrounded by an insulation of car- bon stone bricks, which are inside metallic liners on both sides. This component is arranged inside a metallic vessel with a diameter of 2.3 m and a height of 18 m. The "cold" he- lium flowing back from the circulator to the helium heater is guided in the gap between the metallic vessel and the car- bon stone insulation. Fig. 1 shows a scheme of the steam re- former and fig. 2 a photograph of the bundle itself. The steam reformer tubes are filled with a catalyst of raschig ring type. Every 3 tubes have a common feed- and product gas line. Outside the vessel the 10 feed and the 10 product lines are connected in central headers. The feed gas streaming down- wards is heated up and reacts at the same time according to the steam reforming reaction. The product gas is guided at the lower end of the steam reformer tubes to two small coiled returning tubes and flows then upwards. Additional to the well known advantages of the inner gas return tube (heat back to the process, lower temperature level for the penetration etc.) the coiled version has the following advantages: - 2 - D 101 RSO L 18.126 mm 0 2.300 mm 6 54 mm Wst 13 Cr Mo 44 M 171 t T 500 P 45 bar i ill in. Wlhi Fig. 1: Scheme of the bundle arrangement in the vessel Fig. 2: Steam reformer bundle 3 - 3 - 1) Good compensation of differential length in the start-up and shut-down phase. 2) No possibility of sinking of the catalyst because each coil carries a part of the catalyst. After 8 000 hrs operation time, 6 000 hrs with a helium tempera- ture above 800 °C and a lot of start-ups and shut-downs no sinking of the catalyst was registrated. The main tests of the component were related with the power dependence. The several loads are adjusted by massflow control but also by the helium enterance temperature. Fig. 3 shows the temperature of the product gas at the end of the catalyst bed versus the relation of the mass flows of helium to process gas at the helium inlet temperatures of 950 °C and 900 °C. At the design point (helium temperature 950 °C, pressure 40 bar, helium mass flow 3.2 kg/s) a product gas temperature of 825 °C results. The approach to the chemical equilibrium is 9 K that means the content of unvonverted methane in the product gas is equivalent to an equilibrium temperature of 816 °C (fig. 4). The dependence of the productgas temperature versus the mass flow relation is represented by a straight line with an accep- table accuracy. This type of control has the disadvantage that the helium exit temperature of the bundle also decreases and this results in an additional reduced steam production and extremely there is not enough process steam produced. Therefore it seems to be necessary at smaller part loads (70 % or less) to have not only a mass flow control but also a combined mass- flow-/temperature control. While the maximum process gas tem- perature has only a very small pressure dependence, there is a significant dependence for the unconverted methane content in the product gas, that means at lower pressure that part of helium heat used for conversion of methane increases. This results for the Nuclear Chemical Heat Pipe System in a larger production of chemical heat and for the nuclear process heat applications in larger production of hydrogen + carbon monoxide. This could be verified by the experiments. The results are shown in fig. 5, where the relation of the chemical heat to the energy - 4 - 900 'PG.MAX D Ausiegung 850 o RSO-Eintr.-Temp. Helium o ca. 950 °C 800 • ca.900 °C 1.5 m'PG Fig. 3: Maximum process gas temperature versus relative mass- flow RSO-Eintr.-Temp. Helium / n Ausiegung Systemdruck 0,10 ca.900°C/30bar /ca. 900°C/40bar RSO.A 0,05 .ca. 950°C /40bar 1,5 PG Fig. 4: Unconverted methane content versus relative massflow _ cr _ supplied by the helium heater versus the system pressure is shown. A similar effect like the pressure reduction has a higher helium inlet temeperature to the steam reformer bundle. The temperature dependence is given in fig. 6. Starting with these results, one can see that it is possible to get a higher effi- ciency with a lower pressure, or the same efficiency at a lower helium- and process gas temperature level which results lower wall temperatures too. For nuclear heated steam reformers it is planned that in case of tube cracks the damaged tube has to be plug off. This pro- cedure causes consequences for the operation of the steam re- former. A series of experiments, done in the way shown after- wards, served the investigation of these effects. After a stable operation point was reached with 3 0 tubes, 3, 6 and upto 9 tubes were plugged off on the feed side. That means the steam re- former was operated under constant conditions like mass flow (helium and process gas), inlet temperatures, pressure etc. but with 30, 27, 24 and in some cases 21 tubes. The result of these experiments in the direction of the maximum process gas temperature are shown in fig. 7. For each plugged-off group of 3 tubes the process gas temperature decreases 7 K and the conversion too. The non-fed tubes nearly have helium tempera- ture in these experiments. Because of the mixing effect of the cross counter flow on the helium side there are no parts with significant higher helium temperatures. But it has to be mentioned, that in case of 6 or more plugged-off tubes, all loacated in the same quarter, the fed tubes surrounded by plugged-off tubes have a systematically higher temperature than those only surrounded by fed tubes. Always the wall tem- perature of the original situation (30 tubes) will not be reached. A last series of experiment dealed with the influence of the steam/methane ratio in the direction of carbon deposit for- mation. Because of economic reasons a steam/methane ratio as low as possible inpreferable. Therefore the ratio was system- atically decreased and at ratios below 2.5 mol/mol in the - 6 - .5 30 bar Fig. 5: Efficiency versus process pressure 850 30O 350 Fig. 6: Efficiency versus helium inlet temperature - 7 - process condensate carbon deposits were found. All observations (no activity loss of catalyst, pressure drop etc.) show that the carbon was not deposited in the catalyst itself but in the cooling devices like inner gas return tubes process gas coolers. Therefore the carbon seems to be formed by the Bondouard reaction (2 CO —*- C + C0~ ) . With the end of the year 1983 the succesful experiments with the bundle were ended. At this time an after-operation inspec- tion program is on the way. 900 Systemdruck ca.40bar 'PG.MAX [ °C ] RSOEintr.-Temp Helium 850 ca.950°C 800 ca. 900 °C berikksichtigte Meflstellerv. 4,7,13,14,20,21 1,5 "He m'PC Fig. 7: Maximum process gas temperature with tubes plugged-off No. 1 9 IAEA - INTERNATIONAL WORKING GROUP ON GAS-COOLED REACTORS •III XA0055828 Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors Duesseldorf Federal Republic of Germany 16-19 April 1984 'EVALUATION OF MATERIALS FOR HEAT EXCHANGING COMPONENTS IN ADVANCED HELIUM-COOLED REACTORS by F . Schubert Kernforschungsanlage Juelich GmbH Institute for Reactor Materials Evaluation of Materials for Heat Exchanging Components in Advanced Helium-Cooled Reactors by F. Schubert Kernforschungsanlage Juelich GmbH Institute for Reactor Materials Vortrag anlaSlich des IAEA-Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors vom 16. - 19. April 1984 in Diisseldorf - Summary The qualification of metallic materials for advanced HTR applications is based on creep behaviour, fatigue properties, structural stability and corrosion resistance. A brief state of the art is provided for the materials for heat exchanging components. The experimental results are treated with respect to the importance for the design, the characteristic of time-depend materials behaviour are evaluated. Of specific interest are the possible effects of helium on the mechanical properties. Helium, which serves as primary coolant, contains traces of reactive impurities such as hydrogen, methane, carbon monoxide and watervapor. The evaluation of the HTR materials program serves as basis for structural design rules of components with operation temperatures above 800 °C. The materials mechanical topics are discussed. Alloy improvement and the progress in development of new alloys are reviewed. 3// 1 . Introduction The fule elements and the structural graphite for the core of helium cooled HTR's are developed and qualified. It is generally accepted that they can be used without problems up to temperatures of about 1000 °C. A comprehensive survey of the state of the art is given in the 1977 - special issue of Nuclear Technology /I/. In the steam cycle HTR's (Fort St. Vrain and THTR) the highest metals temperatures are around 750 °C. The iron-base alloy X 10 NiCr 32 20 AITi (INCONEL 800) can be used. For the heat exchanging components of advanced HTR for process heat creep resistance metallic alloys are needed for long operation times at temperatures above 760 °C. The availability of creep resistance material for these temperature ranges is one of the most important aspects for the feasibility of advanced HTR-systems. Therefore, in the frame of German HTR research and development work, the evaluation and qualification of creep resistant high temperature alloys is one of the main tasks. There exist also current material programs for advanced HTR in Japan /2/ and USA /3/, partly in cooperation with Germany. A selection of typical results should demonstrate the state art of the German activities in this contribution. A general and more detailed survey on the German programs has been provided in 1982 /4/, the states of international effort will be summarized in a special issue of Nuclear Technology, which is now being published /5/. The conventional technology, e.g. gas turbine and petro-chemical plant construction, offers solutions for components in the temperature range of 800 to 1000 °C. Alloys, developed for these non-nuclear applications are commercially available. For nuclear systems, however, a specific completion of the materials qualification is recommended because of the necessesibility to predict the properties of materials for very long operation times of more than 100 000 hrs. -2- A further aspect of the qualification of alloys for HTR applic- ations is the effect of the working environments on the mechanical behaviour. The first investigations of the effects of impure helium simulation the primary coolant gas of an HTR on the properties of alloys were started in the early 1970s by the OECD Dragon Project /6/. The stress rupture and corrosion behaviour of various alloys in simulated HTR helium were investigated. The results of these tests showed a significant influence of the gas composition and temperature on corrosion effects (e.g., at 850 - 1000 °C, internal oxidation and severe carburization were found). However, an effect of corrosion on the stress rupture properties was not clearly observed. 2. Expected operational conditions The present paper is essentially oriented towards the use of a pebble bed HTR for nuclear process heat. In these cases the highest temperatures of metallic components occur in the hot gas ducts, the methane reformer tubes and the helium/helium heat exchanger. These components are the topics of this specialists' meeting. In the case of nuclear process heat, the thermal power of the core is transfered by helium either to the reformer tubes (maximum metals temperature 900 °C) in which hydrogen is generated for the hydrogasification of coal by methane reforming or to the tubes of the intermediate heat exchanger (metals temperature of about 930 °C), which is required in the steam gasification of hard coal (Fig. 1). In order to achieve long service lives the stress under normal operational conditions is 2 kept below 5 N/mm . The component must withstand additional temperature increases and/or pressure transients under up set and emergency conditions. For thick-walled components with wall thickness translations, particular attention must be paid to fatigue phenomena due to starting up and shut down operations. The coolant is the inert gas helium, but because it contains slight impurities in mbar range /8/ of CO, CH., H~, H^O, CO- and N_ reactions can be expected with certain alloy constituents which may cause material property changes. In Fig. 2 some of the simulated HTR-heliums are presented. 3/3 -3- Under these operational boundary conditions the design and the analysis of component behaviour must be outlined. Today, no accepted structural design rules or codes are avialable so far for nucler application. The ASME-CC N 47 /I/ contains rules for the design of nuclear components based on the time dependent properties for application temperatures up to around 800 °C and may provide guidelines for high temperature design methods. The work at present in progress has, however, shown that the rules and tests given in the code are of only limited use at the higher temperatures, for two main reasons. Firstly, the candidate materials have not yet been adopted for nuclear applications and secondly the material response to the predicted HTR service conditions proves to be rather different to that at the lower temperatures for which the code is valid. The stringent safety and reliability requirements which are applied to nuclear components demand that, for all postulated service and upset conditions the design shall ensure against failure. This can only be achieved if all modes and causes of failure are known, characterized and controlled by adequate methods. For design with time dependent properties at high temperatures, the following failure modes need to be considered: failure due to short-term loadings, failure due to long-term loadings, failure due to creep-fatigue loadings, distortion due to incremental collapse and creep ratcheting. In addition to the design against failure, a design against excessive deformation is required to exclude loss of function and to maintain as far as possible the original design status of the component. The calculation of component deformation behaviour requires practicable constitutive equations, which must be incorporated into the elastic, simplified inelastic or full inelastic analysis. The equations must reasonably describe the material response to the loading conditions. Material failure and deformation limits are derived from the material properties data. -4- 3. Materials evaluations The selection of candidate alloys for the advanced HTR projects was based on long term creep rupture properties, structural stability and fabricability. For the highest operating temperatures, solid solution hardened nickel-base alloys were therefore chosen. To qualify these alloys for nuclear applications, extensive test programs have been initiated /4, 5, 8, 9/, the principle tasks of which are: creep-rupture testing of the basic alloys and weldments determination of the changes in mechanical and physical properties after long-term exposure under simulated service conditions, particularly ductility (end-of-life properties) high and low cycla fatigue testing gas-metal-reaction kinetics influence of high temperature corrosion in simulated service environments on mechanical and physical properties creep-fatigue interactions fracture mechanics studies with special emphasis on the effects of service environments on crack initiation and propagation fracture toughness characteristics constitutive equations damage accumulation and estimation of service life transferability of mechanical properties data to multiaxial loadings and complex geometries fretting, friction and wear development of coatings influence of irradiation on materials behaviour (for metal- lic control and absorber rod cladding only) In addition to the extensive research and test programs for the qualification of commercial alloys for HTR process heat applic- ations, efforts have been expended to develop new alloys designed specifically for high temperature HTR components. Generally, the primary goals of these developments are improved high-temperature creep and fatigue strength, better corrosion resistance in the HTR primary coolant. -5- 4. Selected commercial alloys From the groups of high temperature alloys, some NiCr-alloys have been selected. The nominal compositions are given in Fig. 3. The most important criteria for the selection of alloys are the creep behaviour and the availability of the needed semi-finished products. In Fig. 4 average 1 ?o creep strain limit and creep rupture strength are shown. These data originate from the extrapolation of a scatterband evaluation /7/. The magnitude of creep strength indicated the capacity for practicle use. Among those materials which can still be processed into plate and tube geometries by hot and cold working INCONEL 617 exhibits the highest level of creep strength. INCOLOY 800 H is a well characterized and readily available material which will be of significance in HTR's for medium temperatures. In the INCOLOY 800 version it is also used for tubes in LWR steam generators. HASTELLOY X has creep strengths between those of INCONEL 617 and INCOLOY 800 H. 5. Status of materials qualification 5.1 Creep behaviour With regard to design against creep failure, an alloy is only sufficiently qualified for technical application if the long-term data and their scatterbands are fully known and the influence of opperational atmospheres on the creep properties is established, so that the design data can be derived. -6- From the comparison of the test results INCONEL 617 reveals at the high temperatures the best creep strength characteristics /10/. There is no influence of the simulated HTR-helium on the l?o creep strain limit and on the creep rupture strength for this material up to about 20 000 h (Fig. 5). This observation is valid for all alloys investigated. The creep rupture deformation of specimens in simulated HTR-helium atmosphere is in the scatterband of values obtained in air. Fig. 6 shows the results obtained on specimens on various INCOLOY 800 H heats tested in simulated primary circuit helium (mainly carburizing under the conditions used in the creep rigs) in methane reforming gas and in air. Up to now, no systematic effect of the test atmosphere can be seen and a common scatterband including all the data points can be drawn. The rupture elongations of INCONEL 617 (Fig. 7) shows a trend towards lower ductility in HTR helium than in air for specimens of small diameter. It is well demonstrated, however, that the behaviour of pre-car- burized specimen differs from that of specimens subject to continous carburization during creep testing. For long time design values, the extrapolation methods must be carefully proved. The usual parameter methods, e. g. according to Larson-Miller, as well as graphical methods are compared in respect to their applicability. For the design of heat exchanger and methane reformer long term design data can be extrapolated on the basis of experimental results between 20 000 h and 25 000 h. An extrapolation in time with a factor 3 is usual, above that, the extrapolated design values become more and more uncertain. For the inelastic analysis creep laws are required. They are derived from creep curves. Some typical creep curves are shown in Fig. 8. The curves for INCONEL 617 at 850 °C show the more or less classical form with recognizable primary, secondary and tertiary creep regimes. At 950 °C the creep rate increase from the beginning of the test with no recognizable secondary creep 2/7- -7- regime. However, using a more detailed analysis, the parameter for Norton's creep law can be derived. 5.2 Behaviour during short-term load For evaluation short-time loads and inelastic analysis, stress strain curves are to be used. Several stress-strain diagrams for HASTLLOY X are shown in schematic form in Fig. 9. At room temperature and elevated temperatures tensile tests show the expected stress-strain curves with a straight rise in the elastic region, following by plastic strengthening above the 0.2?o yield. But, at very high test temperatures (in tests with constant strain rate) the specimen reach a yield point and deform than in a way, which can be approximately described as creep. The beginning of this "creep" depends on the strain rate . The deformation at ambient and elevated temperature for the candidate material can be described by an equation for the strain in the form of £ = £ , + £ , + £ *- el pi cr where C - elastic strain; At high temperatures, however, the plastic fraction may be neglected and the strain results as s = S. . + E <~ el cr The validity of this simplified equation is temperature and deformation rate dependend. Results from strain-controlled low cycle fatigue (LCF) tests are required to assess the influence of thermal cycling during start-up and shut-down operations and during power changes. The allowable start-up and shut-down cycles as well as the cycle temperature changes are determined by using fatigue curves. These design curves for different temperatures are derived from experimental LCF-results using adequate safety margins. Test results for the material INCONEL 617 (determined in air and helium) are given in Fig. 10 /ll/. It may be seen that the specimens frequently exhibit higher fatigue life endurance with equal strain range in simulated HTR helium. However, detailed investigations concerning the influence of HTR helium on crack formation and crack propagation are still required in order to understand this effect. At increasing temperatures the fatigue behaviour is increasingly dominated by creep mechanisms, this becomes obvious when the deformation rate in LCF tests are reduced. The maximum stress in the 200th cycle is plotted against the strain rate as a characteristic value in Fig. 11. The stress decreases remarkably with decreasing strain rates. This fact must be adjusted for, if an inelastic analysis of the component behaviour has to be performed. The experimental results, obtained so far, reasonably allow the development of design curves, the effect of holding time should be further evaluated. 5.4 Properties after long-term annealing After long-term application at high temperatures, all high temperature alloy tend to alter their microstructure. At elevated temperatures even the creep resistant materials are subject to structural changes which may influence their short time properties. In the solution annealed state only a restricted amount of carbide precipitation is within the grains of the grain boundaries /12/. After an exposure, the precipitates, especially those formed at the grain boundaries, lead to a change in the deformability of the alloys. Exposed at 900 °C the RT-impact strength of INCONEL 617 and INCOLOY 800 H is reduced due to the i/f -9- described changes in carbide precipitation (Fig. 12). But tested at ageing temperature, these alloys change their temperature impact strength only slidely. The RT-impact strength of these resistance alloys, which have to be used for the high temperature exposed HTR component, are of minor importance for the design. For the nuclear acceptance precedures the guide lines cannot be used, which are prescribed in LWR-codes for steels used at lower application temperatures. 5.5 Corrosion caused by the operating gases The mechanical properties can also be changed by gas/metal-in- teraction, e. g. oxidation and internal oxidation, carburization and decarburization, which are caused by impurities of the primary coolant gas, helium (compare the nominal impurity levels in Fig. 2 ) . The different operation gases lead to different surface scales dependent on exposure temperature and time and the given alloy composition. Some examples of corrosion scales on three alloys after exposure of approx. 10 000 h in HTR-helium are demonstrated schematically (Fig. 12) /13/. In the spinel layer of INCOLOY 800 H, formed at 850 °C both carbides and oxides are visible. Up to 40 \im beneath the scale internal oxidation of aluminium is observed. NIMONIC 86, exposed at 900 °C, displays a Cr?0^ oxide layer with inclusions of carbides and small particles of residual metallic matrix. At 950 °C INCONEL 617 in water depleted simulated HTR-helium forms an Al~0, layer, which completly covers the surface thus preventing carburization and internal oxidation. Experimental evidence illuminated also the importance of the operational gas atmospheres - especially above 900 °C - for the corrosion in HTR-helium. Small changes in the content of certain impurities can cause different corrosion phenomena. Extensive research is under way to understand the importance of the impurities for the corrosion. -10- A recently developed modell serves as a basis for the lay out of the tests. It was found that the stability of the corrosion system is determined by the interaction of carbon in the alloy with the surface. It is controlled by the stability of carbides in the alloy matrix and by oxidation and carburization potential of the atmosphere (Fig. 14). Assumed, only chromium acts as a carbide and oxide former in the alloy /1A/, the different corrosion effect which can occur may be deduced from a stability diagram if the kinetic conditions of high gas velocities are being considered. If the modell predicts the occurance of the corrosion products properbly - which has to be proved by the ongoing tests - areas can be identified, where no detrimental corrosion should be expected. Whenever those boundaries are exceeded remarkable corrosion effects such as carburization can occur, which is frequently observed after tests in simulating HTR-helium atmospheres. In simulated PNP-helium up to operation temperatures of about 900 °C the gas/metal-reaction are of minor significance for the mechanical properties. In dependence of the local impurities levels or at higher temperatures, however, the possibility of significant decarburization or carburization must be taken into account. For the design, corrosion effects can be handled by a wall thickness margin when surface corrosion or internal oxidation are dominant. If there are significant decarburization and carburization effects one has to analyse their influence on the mechanical properties and their impact on the behaviour of the component. 6 . Possibilities of improving materials for heat exchanging components Operational limitation due to the creep strength and corrosion effects of the commercial alloys asks for improved alloys. Since alloys for tubes have to be workable and weldable the main strengthening mechanism such as y- -hardening, which is suitable for cast turbine blade materials cannot be used. The most promising opportunity left is solid solution strengthening of Ni-base alloys. The optimizing studies aim at both the in- creasing of the creep strength and the improving of the resist- -11- ance to carburization in HTR-helium /15/. With an alloy of the nominal composition: 0.05 C, 1 Si, 1 Mn, 32 Ni, 25 Cr, 20 Fe, 12.5 W (in mass. %) , Thyssen Edelstahlwerke have developed a material which is highly resistant to carburization and decar- burization in simulated PNP-helium while featuring the same creep strength as INCONEL 617 (Fig. 15). In recent work /16/ it could be demonstrated that in the temper- ature range, where chromium oxide is no longer stable, a protective scale due to Titanium and Aluminium content is formed. Alloy development programs aimed at high temperature creep resistant alloys for nuclear heat exchanger have already been under way in Japan /17/. The creep strength data display certain improvements compared to those of INCONEL 617. This essential improvement is achieved by tungsten contents of more than 15 %. Above 15 ?o NiCrW-alloys are additionally strengthened by the formation of \A -W precipitates within the grain. The micrograph (Fig. 16) illustrates their shape and their homogeneous destribution within the grain. This favourable strengthening mechanism provides a new alloy development line for HTR application. The accessibility high temperature creep strengthening mechanisms provide a potential for the improvement of creep strength. In addition the enhanced resistance against carburization or decarburization can be achieved for a certain helium impurity composition. 7. Remarks concerning analysis of component behaviour There is a lack of nuclear structural design rules for the dimensioning and analysis of the operating behaviour in the very high temperature range. The art of understanding will be presen- ted by the paper of Bieniussa et al at this meeting /18/, which discussed, collected and proposed basics for a design code for advanced HTR components at temperatures above 800 °C. -12- The results of the ongoing materials program is one of the basis for these investigations. Recommandations concerning the analytical procedure are presently worked out. Although the data derived from the materials program are sufficient for determining preliminary design characteristics, further support is required from tests with specimen under multiaxial loading and with component relevant geometries. The following experiments with multiaxial loads are being carried out in order to verify model computations of the deformation and failure behaviour /19/: creep under external pressure (ovality influence) creep under internal pressure creep under internal pressure with superimposed tensile stressing strain cycling torsional stressing fracture mechanics crack propagation studies. The first results are compared in the Fig. 17. In addition to the evaluation of safety problems these experiments, serve for verifying model concepts, simplified analyses and finite-element methods. A further point of increasing effort must be experiments for life-time prediction rules. The creep fatigue interaction is frequently determined by means of LCF tests with hold times. When these results (Fig. 20) are plotted in the coordinates for exhaustion of fatigue and of creep, these experimentally obtained values deviate considerably from the straight line expected. A functional description of the results can be achieved, however, using a modified model in which creep is taken more into consideration. -13- 8. Concluding remarks The evaluation of commercial alloys for high temperature applic- ation are discussed and their significance for design is shortly reviewed. Concerning creep behaviour, the so far available results allow the definition of long term design data for heat exchanger and methane reformer by extrapolation up to about 70 000 h. The development or preliminary fatigue design curves is possible. It is postulated that room temperature properties after overageing are of minor importance for the design. Corrosion effects can be handled by a wall thickness margin when surface corrosion or internal oxidation are dominant. In the case of significant decarburization or carburization effects their influence on component behaviour has to be analysed. There is a potential to increase the high temperature creep strength and improve the corrosion resistance of commercial alloys by alloy development. The consideration of structural design rules ask for continuing research efforts on the following topics: measurement of long-time data for the base material and weldments establishing applicable rules for life time prediction examining the transferabi1ity of uniaxial data to component geometries with multiaxial loading. Some problems are not yet completely resolved, the results obtained so far in the materials testing program indicate that it is possible to utilize nuclear high temperature heat for processes such as coal gasification or methane reforming. -14- LITERATURE /I/ "Coated Particle Fuels", Special Issue, Nuclear Technology, Vol. 35, pp. 205 - 573 (1977), ed. T. D. Gulden and H. Nickel /2/ R. Tanaka, T. Kondo, "Research and development on heat resisting alloys for nuclear process heating in Japan", Nuclear Technology, in print. /3/ R. E. Ellus, I. H. Caturilo, 0. F. Kimball, "Effects of simulated HTGR- Primary Coolant on the Structure and Properties of Structural Alloys", Nuclear Technology, in print. /4/ Statusseminar "Metallische Werkstoffe", Dusseldorf, Januar 1982, Bd. 14, "Energiepolitik in Nordrhein-Westfalen /5/ Status of Metallic Materials Development for Application in Advanced High Temperature Gas Cooled Reactors", special issue of "Nuclear Technology", in print, eds. H. Nickel, T. Kondo, P. L. Rittenhouse /6/ R. A. K. Huddle, "Metals and alloys for very high temperature reactors", Reprint BNES Intern. Conf. "High Temperature Reactors and Process Application", London, 1974, ISBN 7277 0049 /!/ ASME-Code Case N 47-15 "Class 1 Components in Elevated Temperature Service, Division 1", The American Society of Mechanical Engineers, New York (1974) /8/ H. Nickel, P.L. Ennis, F. Schubert, H. Schuster: "Qualifi- cation of metallic materials for application in advanced high temperature gas cooled reactors", Nuclear Technology, Vol. 58 (1982), p. 90 - 106 -15- /9/ H. Nickel, F. Schubert, H. Schuster: "Structural Materials in Helium-Cooled Reactors Today", Vol. 2, British Nuclear Energy Soc., London, 1982 /107 F. Schubert, U. Bruch, R. Cook, H. Diehl, M. TeHeesen, G. Ullrich, H. Weber: "Zeitstandverhalten" in Statusseminar "Metallische Werkstoffe", Dusseldorf, Jan. 1982, Bd. 14, "Energiepolitik in NRW" /ll/ H. P. Meurer, G. GnirG, W. Mergler, G. Raule, H. Schuster: "Untersuchungen zum Ermudungsverhalten von HTR-Werkstoffen bei Temperaturen bis zu 1000 °C" in Statusseminar "Metallische Werkstoffe", Dusseldorf, Jan. 1982, Bd. 14, "Energypolitik in NRW". /12/ H. Kirchhofer: "Beitrag zum isothermen Ausscheidungsver- halten von hochwarmfesten Nickellegierungen", Dissertation RWTH Aachen (1983) /13/ H. Schuster, R. Bauer, L. Graham, G. Menken, W. Thiele: "Corrosion of High Temperature Alloys in the Primary Circuit Gas of Helium Cooled High Temperature Reactors", Proceeding 8th Int. Conf. Metallic Corrosion, Dechema (1981), p. 1601 /14/ W. Quadakkers, H. Schuster: "Corrosion mechanism of nickel base alloys in the primary coolant gas of high temperature reactors", anl. 9. ICMC, Toronto, Canada, 1984, Paper 005F- 000299 /15/ B. Huchtemann, L. W. Graham, W. Schendler, P. Schuler, H. Weber: " Legierungsentwicklung fur einen He/He-Warmetau- scher" in Statusseminar "Metallische Werkstoffe", Dusseldorf, Jan. 1982, Bd. 14, "Energiepolitik in NRW" /16/ P. J. Ennis, A. W. Dean: "Alloy Development for HTR-Helium", to be published in Nuclear Technology. reformer tube! 900°C tH4+H2O-»|CQ+3H2 (to gasworks); sjprocess gas+catalyst)j ^ 950°C Helium+ hO-15l- mm impurities immersion IHX heater CO.CH, CO2,N2) ^doped helium) I jrocess gas+coal) 875°C, -2,2 mm ~6 mm Fig. 1 : Tube wall and types of gas in circuits of the proto- type plant for nuclear process heat Test facility Gas designation Nominal impuri ty content, ,ubar H2 H20 CH4 CO co2 KFA, BBC, HRB HHT-helium * 50 5 5 50 5 KFA, HRB, IA PNP-helium + 500 1,5 20 15 — HTMP (Oslo)*) Oslo Phase 4 + 500 1.5 50 40 — — GE NPH helium + 400 2,5 20 40 0,2 6 ORNl HTGR helium + 500 0,2 50 40 — 1 JAERI Helium "B1 + 200 1 5 100 2 — *' former Dragon Project ** direct cycle helium turbine HTR + nuclear process heat HTR Fiq. 2: Nominial compositions of simulated HTR helium test atmospheres nominal composition {wt °/o) alloy C Fe Ni Cr Co Ti Al Mo others INCOLOY 800 H — — — (X 10 NiCrAITi 32 20) 0.08 bai. 32.0 21.0 0.4 0.4 HASTELLOY X — — (NiCr22 Fe 18 Mo) 0.07 18.0 bal. 22.0 1.5 9.0 0.6 W INCONEL 617 0.07 - bai. 22.0 12.5 0.4 1,0 9.0 — NIMONIC 80 A 0.08 - bal. 19.5 — 2.2 1,4 — Fig. 3: Nominal compositions of selected alloys for heat- exchanging components 1% ereep :>traiiT lirnit stress rupture strength 10 £ 6 U 01 •••-.... X)1 101 8 8 6 6- INCONEL 617 - N HASTELLOY X NIMONIC 86 2- HASTELLOY S" 4 3 io5 102 3 103 3 10A 3 105 time /h time /h Fig. 4: Creep rupture strength of materials und^r consider- ation for advanced HTR's 200 00 O. *O * "*** n*^ ~ '^*f o *iSiO^ " fl \ T/-C 600 850 900 950 1000 HTGR-He — tests continuing air o V o I I 10 102 10J 10* 10 102 10s time to 1%strain/h time to rupture /h Fig. 5: Creep properties of INCONEL 617 10z s8 m - 6 ^ 4 *^ 800°C K • a 4 •a w m T 1 10 V 8 °\ n\ 6 HTR He T - 9508C process gas • s 4 air • > 10' 10J 105 time to rupture / h Fig. 6: Comparison of results for INCOLOY 800 H specimens tested in different atmospheres §70 900 °C diameter (mm) air helium a en -4,5 a • M 60- 4,6-5,0 A • at 5,1 - 6,0 V T 50 6,1-7,0 0 • 7.1- 0 40 30 • 20 10- cm 0 103 io4 rupture time/h Fig. 7: Rupture ductility of INCONEL 617 single specimen test — — — - mulH specimen tesHaverage of two tests} 5x103 10* 1,5x10* 2,0x104 2,5*10* time/h Fig. 8: Creep-curves of INCONEL 617 RT 0 0,5 1 strain E (%) «- Fig. 9: Stress-strain curves for HASTELLOY X (schematic) 750°C 850°C 950°C range for design curves 10' 10 103 cycles to failure Fig. 10: LCF results for INCONEL 617 T=850°C I T I T I 6 10"5 2 4 6 r4 2 4 6 10'3 2 4 6 10"2 at c c Fig. 11: Max. tensile stress ( 3 72no^ Y le nr. 200 (saturation stress) vs. strain rate in LCF tests « test at RT i o test at 900 »C j .initial values •/(solutioX n treated) INCOLOY 800 H ageing time [ h ] Fig. 12: Impact energy for HTR-alloys after isothermal exposure at 900 °C Carbide Spinet I A12O3 Carbide ^SiO 2 Cr - V 2°3 j 1 / Carbide /I A12O3 INCOLOY 800H NIMONIC86 INCONEL 617 20>jm 1 900°C8600h i 950°C 10000h 1 ODU L IU UUUn Fig. 13 : Different surface layers of high temperature alloys formed in HTR helium carbid Cr C n m ^oxid \ O sV IVb v\ 1 II p°r Cr- metal log p0. Fig. 14: Different corrosion effects with chromium in chromium containing nickel basse alloys, a : carbon aactivit< y c partial pressure oxygen lowest CO-pressure for preserva- CO tion of acceptable carbide structure stable coexistence of chromium and corrosion products 333 Prüftemperatur: 950 °C Streubandgrenzen NiCrCoMo 22 12 9 o NiCrCoMo 22 12 9 E# Fe-Ni-Cr-Modellegierungen TypI ~& Ni-Cr-Modellegierung Typ H LOJ 100 1000 10000 100000 Beanspruchungsdauer bis zum Bruch in h Fig. 15: Creep results of the new Thyssen alloys a- tungsten ' V::l •'"•; . y Co \ i o 'St ' V- i , 50jim- Fig. 16 : Precipitation of flC-tungsten in alloy 55 113 MA (Ni-23.5 Cr-18.5 W-0.2 C) •V Ml1 •! I * t 3 c. ' ' L Fig. 17: Failure of INCOLOY 800 H tubes after multiaxial creep tests 1 1 1 1 1 1 • 1 1 1 1 i i i 1 ®30 * - CO - 1.0 INCOLOY 800 H _ K 850°C, air • A \^o 3^3 0,3 % 0,6% B 1,0 % 0,5- \ ,X 0=1 1,5 % y°Ai ^\~~ 30 h B 3 \. V^o°« ^ - f 1 1 1 f 1 I 1 1 1 0,5 1,0 1,5 JL NiB Fig. 18: Linear damage accumulation for creep-fatigue inter- action No. 20 a GEFR-SP 315 DATE April, 1984 TITLE: Pressure Vessel Design Codes: A Review of Their Applicability to HTGR Components at Temperatures above 800 C. AUTHORS: P.T. Hughes, H.H. Over & K. Bieniussa XA0055829 Prepared for presentation at IAEA Specialists Meeting on Heat Exchanging Components of Gas Cooled Reactors Conference Held in DuesseldorfCity, Stat, eFRG On 16-19 April 1984 Date This paper contains material resulting from work performed for U.S. Department of Energy Under Contract No. DE-AC03-80ET34034 This paper has been authored by a contractor of the U.S. Government under Contract No. DE-AC03-80ET34034 Accordingly, the U.S. Government retains a nonexclusive, royalty-free license to publish or reproduce the published form of this contribution, or allow others to do so, for U.S. Government purposes. ADVANCED NUCLEAR TECHNOLOGY OPERATION ^GENERAL ELECTRIC COMPANY SUNNYVALE, CALIFORNIA 94088 GENERAL W ELECTRIC 84-02-11 INTRODUCTION The governments of the United States and the Federal Republic of Germany have approved of cooperation between the two countries in an endeavor to establish a structural design code for gas reactor components intended to operate at temperatures exceeding 800°C. The basis of existing codes and their applicability to gas reactor component design are briefly reviewed in this paper. This review has raised a number of important questions as to the direct applicability of present codes. The status of the US and FRG cooperative efforts to obtain answers to these questions are presented in a companion paper at this conference. PRESSURE VESSEL CODE POLICIES All pressure vessel codes are either based on a policy of "design by rules" or a policy of "design by analysis". Compliance with the rules is sufficient for safe design with codes of the "design by rules" type. A formal structural evaluation of critical vessel regions and loadinys, for comparison with structural criteria, is the basis of safety in "design by analysis" codes. The regulations and requirements to be observed in design by rules concern geometrical arrangements, dimensions, loading conditions materials and fabrication methods. Estimation of the necessary wall thickness is obtained from elementary calculations which relate a permitted stress at temperature, the principal vessel dimensions and the pressure loading. The regulations and requirements represent a synthesis of satisfactory vessel design and operation. For this reason, they may not be applied to vessels intended for operation at temperatures and in geometrical forms under load conditions which differ from those from which the rules were made. Codes based on a policy of "design by analysis" enable vessel design to proceed for complex loading conditions or geometric arrangements or both. In such codes, specific criteria which define the permissible stress, strain or displacement are stated for both elastic and inelastic behavior. Specification of a criterion for either stress, strain or displacement uniquely specifies the remaining two quantities for the case of elastic vessel response. This cannot be so for inelastic behavior, and stress, strain and sometimes displacement, each require separate criteria. The structural response to load and temperature for comparison with criterion may be obtained from experiment but is usually the result of calculation. -2- LIMITS OF EXISTING CODE APPLICABILITY The principal pressure vessel codes in use in the U.S. and the FRG are the following: FRG- AD - Instructions (3) FRG- TRD - Regulations (4) FRG- KTA - Rules, primary circuit components of light water reactors, 3201 (5) USA- ASME - Code, Section III (6) USA- ASME - Code, Section VIII (9) USA- ASME - Code, Case N47-17 (1592-17) Class 1 components in elevated temperature service, Section III, Division (7) The AD - Instructions, ASME-Code, Section VIII and TRD - Regulations are based on the policy of "design by rules". They are intended to govern the design and construction of conventional boilers and pressure vessels. The design restrictions imposed by such codes in conjunction with the need for a more specific evaluation of safety margin have resulted in the evolution of "design by analysis" codes for nuclear quality pressure vessels and reactor structures. In the FRG, the KTA - Rules based substantially on ASME - Code, Section III are in force. The design of components to 800°C is governed by ASME- Code Case N-47 in the US, since the applicability of ASME - Code, Section III is limited to temperatures to about 400°C as are the KTA - Rules. The limits of applicability of these codes may be summarized as follows: AD - Instructions, TRD - Regulations and ASME - Code, Section VIII provide means by which to calculate vessel wall thicknesses for defined pressure loading and temperature, for the required service life. The allowable stresses are specified based on both time independent and time dependent materials behavior. Time dependent allowable stresses are related to both a creep rate and total service life creep strain to offer protection against failure by creep rupture, but other elevated temperature structural phenomenon such as creep-fatigue interaction, creep ratcheting and creep caused enhancement of geometric imperfections are not explicitly covered. KTA - Rules and the ASME - Code, Section III permit only an evaluation of structural response to ensure freedom from time independent failure modes because of the specified and limited temperature range at which they apply. They require first an evaluation of the structural response to primary loads which could cause bursting or unacceptable deformation, followed by an evaluation of the magnitude and frequency of primary and secondary (self limiting) stresses for comparison with fatigue criteria. -3- Whilst stress or strain due to external forces, the self-weight, and pressure sustained by a component, are covered similarly to the procedures of (3) and (4), elastically calculated stresses, due to all types of loading, for the fatigue evaluation, are permitted to vary over a range of magnitude equal to twice the yield stress (3 Sm-criterion). Criterion must be met to ensure that plastic deformation sustained from strain cycling in the plastic range does not impair integrity or function ("shakedown"). Loads from other than conditions of normal operation are classified by frequency of occurrence and successively greater primary stress limits are permitted as the loads become less probable. ASME-Code, Case N-47 includes criteria for "design by analysis" in the time dependent failure temperature range. These criteria were developed through a number of preceding code cases from the fundamental elevated temperature considerations of ASME-Code, Section VIII. The limit of elastic (linear) approximation to non-linear creep behavior is defined in ASME-Code, Case N-47 by restricting to about the 0.1% yield stress, any cycle composed of elastfcally calculated load and deformation (thermal) stresses. If this restriction cannot be observed, inelastic analysis is required to ensure that stress, and cyclic and cumulative strains, meet separate inelastic criteria. A definition of secondary stress in the sense of the ASME-Code(s) is not applicable in such circumstances and stresses are instead defined as load or deformation controlled. The upper temperature limit to the use of a material is governed by creep rupture behavior for which a satisfactory definition requires data which permits well described scatter bands. STRUCTURAL DESIGN CONSIDERATIONS FOR HTGR COMPONENTS The design of components for the high temperature regions of an HTGR plant requires consideration of both time independent and time dependent material properties, the latter characterized by irrecoverable strain and deformation which increase with time. By a process of selection through a formal description of material behavior, and the imposition of geometric and fabrication limts in an elevated temperature code, conditions may be arranged such that the list of possible failure modes is restricted. Criteria can then be devised to ensure a remote probability of failure in any mode. For pressure vessel structures the failure mode list should include - 77 r -4- -ductile rupture -creep rupture due to long term loading -fatigue failure -creep-fatigue failure -fast fracture (failure due to unstable crack growth) -excessive strain due to incremental deformation (cyclic strain accumulation doe to creep or plasticity or both) -loss of function due to excessive deformation -loss of stability due to short term loading (elastic or elastic-plastic buckling) -loss of stability due to long term loading (creep enhancement of initial geometric imperfection) -environmentally caused material failure (carburization, corrosion, irradiation, etc.) Quantitative criteria, observance of which will avoid failure in eight of these ten modes are stated in ASME-Code, Case N-47. Failures due to either ductile fracture or creep rupture are avoided by the definition of load controlled stress limits. Loss of stability due to both short and long term loading is avoided by stipulations such that working loads are substantially less than divergence loads. Calculation of the cumulative strain for comparison with permitted values is the policy used to avoid failure due to creep-fatigue interaction, excessive strain accumulation from incremental deformation and loss of function because of large deformation. These criteria may indeed be suitable for gas reactor components operating at ^er^/ high temperatures but a number of aspects associated with these criteria require critical investigation. The effects of service environments on the behavior of materials and the failure modes of components are not satisfactorily covered in any of the available design codes. ASPECTS OF AN HTGR PRESSURE VESSEL DESIGN CODE An HTGR pressure vessel design code may very well be based on the general policies of the ASME-Code, Case N-47. A general level of security should be obtained by placing limits on the magnitude of load controlled stress and strain and a specific security by further stress and strain limits for the combination of load and displacement controlled strains. The security necessary for prolonged gas reactor operation may also require the development of fracture mechanics based criteria, environmental effects criteria and until sufficient experience accrues, periodic testing of material test coupons and components sections which have been exposed to reactor operation. "0™ For the drafting of a design code: material data sufficient to permit a satisfactory calculation of the force distribution in a component must be available material data to permit evaluation of force distribution in terms of failure modes is required design margins, which offer protection against approximations and uncertainties must be determined for criteria of stress, strain, load and time : the nature and extent of periodic component inspection and possible testing to ensure structural reliability must be considered. The data required for the materials of HTGR components to permit calculation of force distribution and force effects in the application temperature range include: ; ,: : Physcial properties data, such.as density - thermal conductivity .,.<•. , . : specific heat ..'.._, thermal expansion coefficient modulus of elasticity Poisson's ratio Data for constitutive equations (strain-time laws) to permit • inelastic analysis such as: form of the equations constants and coefficients v environmental effects ' ' The data required to assess the significance of force distribution and force effects to stress, strain and displacement for the various failure modes in the application temperature range must include: : Impact Strength ; Yield Strength Maximum Tensile Strength Fracture Elongation and Reduction of Area Creep Strain Rates . Creep Rupture Strength ' Rupture Elongation , Fatigue Parameters ...,' ..-.;.' ; -6- The structural engineering problems of the sodium-cooled reactor and the high temperature behavior of low alloy steel and austenitic stainless steel have been major Influences in the drafting of ASME-Code, Case N-47. The use of nickel based alloys for components in a gas reactor environment at temperatures above 800cC however, presents somewhat different engineering problems and materials behavior. Thus, a number of questions arise for which resolution must be found within the general ASME-Code, Case N-47 policy of "design by analysis" and the description of stress as load controlled or deformation controlled. A failure mode entitled "environmentally caused material failure" has been introduced. The design measures to avoid this failure mode include: -selection of materials -thickness allowance -surface protection -consideration of the changes in material properties Use of any one or combination of these measures required clarification of the following questions: -Which environmental parameters should be used to describe the process of material property changes lead to environmentally caused failure or accelerated failure in other modes? -How should material property alterations and environmentally caused structural effects be specified and presented? -What should be the basis for placing limits on environmental effects? -In what form should limits to permissible environmental effects be embodied in design criteria? A decision to adopt the ASME-Code, Case N-47 policy for the limitation of primary stresses in gas reactor components operating at temperatures above 800°C requires resolution of numerous questions which include the f ol 1 owi ng: -With what strain rate are the time independent strength properties to be determined? -Which cross-section factor K is sufficient for restriction of the primary bending stress in time dependent materials behavior? -Is the stated equation for accumulation of creep damage appropriate for the various operating conditions of a gas reactor? -Is it necessary to specify specific stress limits for welds which differ from those applicable in all regions away from welds? -7- Similarly, a number of questions arise when considering the ASME-Code, Case N-47 philosophy with respect to the accumulation of component material damage due to creep-fatigue: -Which strain rate and hold time should be used in determining cyclic strength properties? -Is the criterion for evaluation of creep-fatigue damage also suitable for HTGR components? -What significance have the strain limits criteria and are the allowable values acceptable? -Does it remain an acceptable procedure to dispense with an evaluation of creep-fatigue damage against region specific criteria for weldments? To provide a sound and thorough basis for the evaluation of creep-fatigue damage using inelastic calculations, resolution of the following types of questions must be achieved: -What manner of tests are practicable and necessary for determining the constants and coefficients of constitutive equations? -In what form should the qualification and verification of computation methods for enelastic analysis be cast? The definition of criteria to avoid the failure mode of unstable crack growth at temperatures below 800°C because of material changes encountered at temperatures above 800°C will be achieved only by answers to the following questions: -What test methods and analysis criteria should be used to characterize toughness alternations and crack growth properties? -In view of inspection difficulties, especially of weldments, should fast fracture criteria be formed in terms of fracture mechanics techniques or a reduction in load and deformation controlled stress and strain criteria, in combination with certain geometric criteria? There have been technical arguments to the effect that ASME-Code, Case N-47 incorporates criteria for creep enhancement of geometric imperfections which are too conservative for load to compensate for lack of margin on time. This situation requires answers to the following type of question. -Should there be criteria and margins for time in addition to, or supplementing, direct criteria and margins for load? -8- The intention of a design code is to offer criteria such that for the factors under design control, the probability of failure is remote. Uncertainties in defining criteria and uncertainties and approximations in the evaluation of a design to meet these same criteria, require the definition of design margins. The factors and effects covered by ASME-Code, Case N-47 design margins are not well defined. In addition, there is little useful background of operating vessel experience above 800°C by which to justify these design margins. Thus, the following questions arise: -How should a design specification and an HTGR structural code be formally linked from the viewpoint of acceptable failure rates? -To define specific and quantitative meaning to criteria which embody design margins what is the best approach and what theoretical and physical data base developments are required? -Which components must be tested to validate criteria and design margins? -What possibilities for testing are practicable? -How are the results of physical testing to be used in comparison with design calculations to verify that satisfactory criteria and design margins have been provided? -What possibilities for testing are practicable? -How are the results of physical testing to be used in comparison with design calculations to verify that satisfactory criteria and design margins have been provided? -Which components when in service require periodic inspection to provide final code varification and information for improvements? -Which in service inspection possibilities are practicable and how is the inspection information to be used to verify or modify criteria and margins? (5) REFERENCES 1. Dokumentation Fachkreis "Regelwerk", Nr. 4.01, Anwendung des ASME-Code, Case N-47 auf HTR-Komponenten, erforderliche Erganzungenbzw. Überprüfungen, July 1980. 2. Dokumentation Fachkreis "Regelwerk" Nr. 10, Vergleich der Aussagen über Werkstoffe, Werkstoffverhalten und Werkstoffversagen verschiedener deutscher und amerikanischer Regelwerke hinsichtlich deren Anwendungsmoglichkeiten fur den Werkstoffeinsatz bei Temperaturen oberhalo 800°C, üuni 1980. 3. Ad-Merkblatter 4. Dampfkessel Bestimmungen (TRD) 5. KTA-Regeln, Komponenten des Primarkreises von Leichtwasserreaktoren, 3201. 6. ASME: ASME Boiler and Pressure Vessel Code, Section III, Rules for Construction of Nuclear Power Plant Components, Division 1, ASME, New York, 1977. 7. ASME: Cases of ASME Boiler and Pressure Vessel Code, ASME-Code Case N-47 -17 (1592-17), Class 1 Components in Elevated Temperature Service, Section III, Division 1, ASME, New York, 1978. 8. ASME: Criteria for Design of Elevated Temperature Class I Components in Section III, Division 1 of the ASME Boiler and Pressure Vessel Code, ASME, New York, 1976. 9. ASME: ASME Boiler and Pressure Vessel Code, Section VIII - Division 1, Pressure Vessels, ASME, New York, 1980. No. 20b .===. Gesellschaft fur Reaktorsicherheit (GRS) mbH XA0055830 Specialists1 Meeting on Heat Exchanging Components of Gas-Cooled Reactors (Duesseldorf, FRG, 16.-19. April 1984) Status of Design Code Work for Metallic High Temperature Components K. Bieniussa Gesellschaft fur Reaktorsicherheit 5000 Koln 1, Schwertnergasse H.-J. Seehafer INTERATOM 5060 Bergisch Gladbach 1, Postfach H.H. Over Kernforschungsanlage Jiilich, IRW 5170 Julich, Postfach P. Hughes General Electric Company Sunnyvale, CA USA Schwertnergasse 1 • 5000 Koln 1 -Telefon (02 21) 20 68-0-Telex 8881 807grsd Gesellschaftfur Reaktorsicherheit (GRS) mbH - i - 1. INTRODUCTION The mechanical components of high temperature gas-cooled reactors, HTGR, (fig. 1), are exposed to - temperatures up to about 1000°C and this - in a more or less corrosive gas environment. Under these conditions metallic structural materials show a time-dependent structural behavior. Furthermore changes in the structure of the material and loss of material in the surface can result. The structural material of the components will be stressed originating from - load-controlled quantities, for example pressure or dead weight, and/or - deformation-controlled quantities, for example thermal expansion or temperature distribution, and thus it can suffer - growing permanent strains and deformations and - an exhaustion of the material (damage) both followed by failure. To avoid a failure of the components the design requires the consideration of the following structural failure modes (fig. 2) - ductile rupture due to short-term loadings - creep rupture due to long-term loadings - creep-fatigue failure due to cyclic loadings - excessive strains due to incremental deformation or creep ratcheting Schwertnergasse 1 • 5000 Koln 1 -Telefon (02 21) 20 68-0 -Telex 8 881 807 grs d 29* Gesellschaft fur Reaktorsicherheit (GRS) mbH - 2 - - loss of function due to excessive deformations - loss of stability due to short-term loadings - loss of stability due to long-terra loadings - environmentally caused material failure (excessive corrosion) - fast fracture due to instable crack growth. With exception of the last mentioned two failure modes, there exists a design code (ASME Code Case N-47 /I/) for the different failure modes in the creep regime. The use of this code is restricted to a small number of materials; the highest tempe- rature of use, T = 816°C, lying substantially below the intended temperature of T = 1000°C. Environmentally caused material failure will be avoided by an appropriate material selection and/or by taking into account this effect in the design procedure. Up to now, a concept for treating fast fracture due to instable crack growth is not yet available. Nevertheless first test results indicate, that at higher temperatures a loss of ducti- lity need not be expected. Based on the design philosophy of this ASME Code Case /2/ and further considering the German and American design codes (fig. 3), the German BMI sponsored Working Group "HTR-Regelwerk" (HTR-Design Code) has developed first bases for a design code for HTR- components /3/. Some of the intentions of this group have been discussed with the American participants of the Subprogram Hate- rials PWS-Rl "Establishment of the Basis for Structural Design of Elevated Temperature HTGR-Components". The good agreement in the object of view has encouraged us to this joint presentation. Schwertnergasse 1 • 5000 Koln 1 • Telefon (02 21) 20 68-0 • Telex 8 881 807 grs d Gesellschaftfur Reaktorsicherheit (GRS) mbH — 3 — As just have been shown, a great number of structural failure modes have to be discussed. In the course of this presentation the current status of design code work for metallic high tempe- rature components exemplarily shall be demonstrated at the issue: - exclusion of creep-fatigue failure due to cyclic loadings. According to our investigations concerning the different compo- nents and the presupposed loadings, this type of failure mode is significant for many parts of the HTGR components. Simultaneously this includes the consideration of the first five failure modes of (fig. 2). Schwertnergasse 1 • 5000 Koln 1 • Telefon (02 21) 20 68-0 • Telex 8 881 807 grs d 3?f Gesellschaftfur Reaktorsicherheit (GRS) mbH - 4 - 2. DESIGN PHILOSOPHY For the exclusion of creep-fatigue failure due to cyclic loadings the design philosophy of the ASME Code, Case N-47, will be kept. In detail this means (fig. 4): - The load-controlled stresses alone have to be limited to the maximum carrying load, taking into account the possibilities of ductile and creep rupture, - the load- and deformation-controlled stresses together have to be discussed with respect to material exhaustion due to hold-times (creep) and alternating loads (fatigue), - the accumulated strains calculated.from the load histogram shall be compared with ratcheting limits; the deformations caused by the strains shall be checked against functional requirements. During the design phase for economic reasons the use of elastic analysis should be strived for. Up to now, this cannot be achieved in a sufficient way (fig. 5). The satisfaction of strain limits using elastic analysis (test 1 to test 4 of the Case N-47) is no longer possible for the tempe- rature regime above 800°C, because - either they are too conservative and do not allow a positive stress report for a reasonable loading time (test 1, test 4), - or there are restrictions for a stress extreme below the creep regime, which cannot be guaranteed (test 2, test 3). Furthermore the creep-fatigue evaluation using elastic analysis cannot be adopted, because of the great conservatisms - in the fatigue curve with hold-time effects and - in the procedure for the creep damage evaluation. Schwertnergasse 1 • 5000 Koln 1 • Telefon (02 21) 20 68-0 • Telex 8 881 807 grs d Gesellschaft fur Reaktorsicherheit (GRS) mbH First comparative calculations, considering the realistic mate- rial behavior and estimating the inelastic behavior by a simple relaxation model, demonstrate, that this simplified method results in - a too fast decrease of the stresses and - a too small rise of the strains in the course of time. In connection with the relatively low material strength at these temperatures, this procedure cannot be tolerated. For this reason at this moment the most promising way in the computing strategy is - elastic analysis for load-controlled stresses alone, - detailed inelastic analysis for load- and deformation-con- trolled stresses combined and - further development of simplified methods based on the results of representative detailed inelastic analyses. There is no operating experience so far with HTGR-component (fig. 6). Further more - the design codes are unproven, - component test results are rare and - loading conditions are presupposed. That is why - inservice inspection of the components, - recurrent tests of the components and - periodic testing of material test coupons are very important in the design philosophy and require com- ponents with both simple geometries and a good accessibility for testing. As mentioned in other presentations, several component tests are running. They can help to remove safety additions not needed and existing uncertainties. Schwertnergasse 1 • 5000 Koln 1 • Telefon (02 21) 20 68-0 • Telex 8 881 807 grs d Gesellschaft fur Reaktorsicherheit (GRS) mbH - 6 - 3 . DESIGN CRITERIA The load-controlled stresses are classified as primary stresses (fig. 7) and have to be limited as usual by the time-independent stress value Si n and the time-dependent stress value S,t. Further- more different loading conditions have to meet the use-fraction sum criterion, for avoiding a creep rupture. At high temperatures the time independent strength is a function of the strain-rate, and the S -value has to be fixed in accor- ' m dance with the loading-rate and perhaps also with modified cri- teria. The criteria for the definition of the S.-value will be kept, excepted the criterion of onset of tertiary creep. For the provided materials the available stress-to-rupture data allow 70,000 operating hours. This statement is based on the assumption, that extrapolation of data from about 25,000 h to 70,000 h is permissible. The combined stress state is evaluated with regard to creep- fatigue failure (fig. 8), using the linear damage interaction hypothesis. Therefore fatigue and creep damage (exhaustion) are computed in accordance with the Case N-47 using the Miner-rule and the Robinson-rule respectively. Both damage portions added have to be less than the total creep-fatigue damage factor. The design curve for computing the fatigue damage portion has been constructed from the cyclic failure curve as usual incor- porating a design factor of 2 on strain range or 20 on cycles, whichever is less. The design curve for computing the creep damage portion has been constructed from the stress-to-rupture curve incorporating a design factor of 1/0.9 on stress and additionally 2 on time, whichever is less. For the different materials the final shape of the allovable creep-fatigue damage factor curve will be fixed at a later date. The strain limits, given in the Case N-47 were kept for this moment, also the restrictions for weld regions. Schwertnergasse 1 • 5000 Koln 1 -Telefon (02 21) 20 68-0 -Telex 8 881 807 grsd Gesellschaftfur Reaktorsicherheit (GRS) mbH - 7 - For these design criteria it is presupposed as working hypothesis that the structural failures for base material and for weld regions are similar. To clarify this point the following tests are running: - creep rupture tests for welds, - cyclic failure tests for welds and later on - creep-fatigue failure tests for welds. Having these results, the very important question can be answered: Does it remain an acceptable procedure to dispense with an evaluation of creep-fatigue damage againt region specific criteria-for weldments? As a conclusion (fig. 9) it can be stated, that the available design criteria and material data are sufficient to fabricate and operate the metallic components of a prototype HTGR. With respect to economy and availability further investigations for the following aspects are desirable: - long-term material data, - environmental influence, - constitutive equations, - simplified inelastic methods. Schwertnergasse 1 • 5000 Koln 1 -Telefon (02 21) 20 68-0 -Telex 8 881 807grsd Gesellschaftfur Reaktorsicherheit (GRS) mbH 4. REFERENCES /I/ The American Society of Mechanical Engineers: Cases of ASME-Boiler and Pressure Vessel Code, Case N-47-18, Class 1 Components in Elevated Temperature Service, Section III, Division 1 ASME, New York, 198 0 /2/ The American Society of Mechanical Engineers: Criteria for Design of Elevated Temperature Class 1 Components in Section III, Division 1, of the ASME Boiler and Pressure Vessel Code ASME, New York, 1976 /3/ Fachkreis "Regelwerk" AbschluBbericht fur Sonderforschungsvorhaben SR 191 des BMI: Erarbeitung von Grundlagen zu einem Regelwerk iiber die Auslegung von HTR-Komponenten fur Anwendungstemperaturen oberhalb 800°C KFA Julich, 1984 Schwertnergasse 1 • 5000 Koln 1 • Telefon (02 21) 20 68-0 • Telex 8 881 807 grs d 1 Operating condition -temperature <1000°C - corrosive gas environment Loading condition - load-controlled (pressure) - deformation-controlled (thermal expansion) Material behavior - growing strains design restrictions - material exhaustion (damage) Component response \7 - failure of component safe operation - loss of function * HTGR METALLIC COMPONENTS ™ (DESIGN ASPECTS) FIG, 1 ductile rupture due to short-term loadings creep rupture due to long-term loadings creep-fatigue failure due to cyclic loadings excessive strains due to incremental deformation on creep ratcheting loss of function due to excessive deformations loss of stability due to short-term loadings loss of stability due to long-term loadings environmentally caused material failure (excessive corrosion) fast fracture due to instable crack growth £ HTGR METALLIC COMPONENTS w (FAILURE MODES) FIG. 2 FRG: KTA safety standard 3201 T < 400 °C design by analysis FRG: AD - instructional sheets T ^ 550 °C design by rules FRG: TRD - regulations T ^ 550 °C design by rules USA: ASME - code, section III T < 427 °C design by analysis USA: ASME - code, section VIII T < 816 °C design by rules USA: ASME - code, case N-47 T ^ 816 °C design by analysis g DESIGN CODES FOR w METALLIC COMPONENTS FIG. 3 relaxation elastic /— time /. inelastic load-controlled stresses (SL) - maximum carrying load for ductile rupture - maximum carrying load for creep rupture load- and deformation controlled stresses (SD) - material exhaustion due to creep - material exhaustion due to fatigue accumulated strains and deformations - ratcheting limits - functional requirements EXCLUSION OF SI CREEP-FATIGUE FAILURE FIG. using elastic analysis -for economic reason strived for - satisfaction of strain limits, no way - creep fatigue evaluation, no way comparative calculations - realistic material behavior - estimation with relaxation model - not conservative in results intended computing strategy - elastic for load-controlled stresses - inelastic for combined stresses - development of simplified methods * HTGR COMPONENT COMPUTING STRATEGY CO C\J FIG, 5 component behavior evaluated with - unproved design codes - lack of component test experience - presupposed loading conditions important for a prototype - inservice inspection of components - recurrent tests of components - periodic testing of material test coupons planned or running component tests - HE/HE - heat exchanger - hot gas ducts - hot gas valves * HTGR DESIGN QUALIFICATION 8 BY TESTING FIG. 6 QUALIFACATION Q. 3 CD t to time t time-independent? time-dependent if) £ = 5%/min -<^SR (stress-to- to to ^^i rupture) Si So B = 0,5%/min (stress 1 %) t strain 7-104 time 'm =s S/? O( — Opj/1,3 ^ S2/1,5 ^ S^/1,0 stress limit: Pm < Sm use-fraction sum: 2 t/tim ^ B = 1,0 g LIMITATION OF 8 LOAD-CONTROLLED STRESSES FIG. 7 time rupture design cycles time Q D = ZN/N CD'I Dc = J DT/TD F D D §^XDF+DC E "D y CD \1 "c3 creep damage Dc LIMITATION OF LOAD-AND CO DEFORMATION-CONTROLLED STRESSES FIG. 8 design criteria and material data are sufficient for fabrication and operation of a prototype HTGR for economy and availability reasons the following investigations are desirable - long-term material data - environmental influence - constitutive equations -simplified inelastic methods I HTGR DESIGN CODE (CONCLUSION) FIG. 9 XA0055831 OXIDE FILMS ON AUSTENITIC HTR HEAT EXCHANGER MATERIALS AS A TRITIUM PERMEATION BARRIER H.P. Buchkremer, R. Hecker, H. Jonas, H.J. Leyers, D. Stover, Kernforschungsanlage Jtilich GmbH Paper for the IAEA Specialist Meeting on Heat Exchanging Components of Gas-Cooled Reactors Dusseldorf April 16-19, 1984 In a high-temperature process heat reactor (1), tritium may inadvertently permeate from the primary circuit into the secondary circuit through the metallic walls of the heat exchanging components due to the high gas outlet temperatures around 1223 K. The German radiation protection regulations only permit a concentration of 10 pCi tritium per gramme of the product gas. On the other hand, hydrogen capable of permeating into the reactor primary circuit is produced on the secondary side of the heat exchangers during the foreseen processes. In order not to reach this limit of 10 pCi/gr it is necessary to produce very effective tritium permeation barriers, which also respond to the hydrogen, in order to keep the existing gas purification system economically efficient. It turned out during our investigation here that such a permeation barrier is optimally formed by oxide coatings (corrosion layers) grown in oxidizing atmospheres where the thickness of the coatings can reach up to a few |im /3/. Figure 1 shows the composition of some steels examined by us. These are centrifugally cast and wrought alloys. The high-alloy steels on a nickel or iron base envisaged for the HTR usually have a chromium content of approx. 20 - 25 %. For this reason, the formation of oxidic corrosion layers must be expected even for our process gas compositions (presence of hydrogen possibly in excess of steam, hence relatively low oxygen partial pressures). - 2 - These oxide films grown in situ, as well as oxide coatings produced in a different way, are known to have an impeding effect on hydrogen permeation (2), (3). The attempt to significantly influence even the permeation behaviour of the bare metal by adding certain alloy components cannot be successful. Figure 2 shows that various candidate alloys in the bare state only differ insignificantly with regard to hydrogen permeation, whereas an oxide coating may effect changes by orders of magnitude. It can therefore be stated that the metals predefined for our applications exert an influence on hydrogen permeation primarily through their capability to form coherent dense oxide films. In the case of bare metal, the /p-dependence claimed by Sieverts is observed over wide pressure ranges. An activation energy for permeation of approx. 63 kJ/mol is observed for austenitic high-temperature alloys and of 33 kJ/mol for ferritic materials. Permeation through the latter is always higher up to 1223 K as compared to permeation through austenitic materials (assuming otherwise identical conditions). The investigations on "coated material" led to the following results: impeding factors (related to the bare metal, otherwise identical conditions) of approx. 200 to over 1000 were measured "in situ", i.e. under process gas atmospheres, with lower values being obtained at low coating temperatures and higher values at high temperatures. It was found that a well impeding layer is obtained if parabolic growth can be observed. It was equally found that the activation energy for this layer growth is apparently determined in essence by the migration of chromium cations with an activation energy of approx. 250-420 kJ/mol which, however, varies significantly depending on the growth conditions of the layer. This is in agreement with observations made by other authors (e.g. Hagel (4), Zink (6) ) . Oxide layers on the material Incoloy 800H, for example, which - 3 - were produced at a certain oxidation potential, are not stable if they are subsequently subjected to a markedly lower potential They obviously transform, reaching distinctly elevated impeding factors after some time - proved by permeation measurements - in conjunction with activation energies for the layer growth rising towards 420 kJ/mol. These layers transmuted in such a way, however, are then stable over wide ranges of the oxygen potential. (We generally work at potentials <10 -1 5 bar.) It may be tentatively assumed that these layers still exhibit pronounced fractions of spinel structures in iron-containing materials. A "more impurity-free" chromium oxide layer is probably produced by reducing the oxidation potential and simultaneously supplying additional chromium from the substrate (noticeable from 1123 K upwards to 1173 K). This assumption is supported by post-examinations which have shown in particular cases that spinel-richer structures are associated with poorer impeding factors. The time-dependent course of the formation of in-situ oxide scales at different oxidation temperatures is shown in Fig. 3. After the usual pretreatment by H annealing, rapid oxidation sets in when process gas is applied, with a clear gradation being observed as the oxidation temperature rises. The increase in oxidation temperature from 923 to 1223 K first leads to a rearrangement of the layer, following which the permeation flux drops to values similar to those for specimens run iso- thermally at T=1223 K. Our oxide films produced in situ show a peculiar behaviour with regard to the pressure dependence of hydrogen permeation. In contrast to observations made by other authors and contrary to the Vp-dependence towards lower pressures as predicted in general theory, we observe a transition to p-dependency for good layers (high impeding factor) at relatively high hydrogen pressures. Below this level, we observe a /p-dependence virtually over 9 decades! Our observations are illustrated in Fig. 4. - 4 - The deviations from Sieverts ' law observed in the high-pressure range can be tentatively interpreted by the introduction of surface reactions. The diffusion process is assumed to govern the time and all surface reactions are to proceed rapidly as compared to diffusion. Of the four possible reaction cases for bare metal in our pressure ranges (cf . e.g. Ash and Barrer ( 5 ) ) R» = k ' p-(1-u) - k ' \J~ physisorption 2 a ~ 2 R2 = kp''l~i^ v - k © dissociation R = k (1-c)6> - k (1-9)-c solution path 1 (A) R, = k, p ( 1 - ©)( 1-c ) - k ,C'9 one atom goes into solution, one atom stays on the dissociation site ; solution path 2 with k as constants 17 area distribution of physisorbed molecules c hydrogen concentration, normalized to saturation H = (1 + D../D. 6/d C.. ./C . ) (1) M Ox M,equi nOx,equi - 5 - in which C_ ., CM . stand for the respective hydrogen UX ^ SQUl ri , e q U1 equilibrium concentrations in the oxide or metal. These can be calculated from the equations (A). Furthermore,'O D.x, and . D,M, denote the diffusion constants of the metal or oxide, )J and d the respective thicknesses of the oxide film and membrane. If the solubilities or diffusion constants known from the literature for metal oxides (8) are tentatively substituted in equation (1), very high values for H are obtained, exceeding any value ever observed by us. We therefore arrive at the conclusion that we are basically concerned with hydrogen conduction via "point defects", which is also indicated by the above observation to the effect that layers defined as "spinel-containing" in surface investi- gations show poorer inhibition than "purer" Cr?0 layers. In addition, an activation energy for hydrogen permeation of approx. 146-162 kJ/mol is observed for good layers as compared to that of 63 kJ/mol of the metallic substrate. Pure chromium specimens /9/, which were oxidized under our conditions, show a qualitatively equal behaviour which also leads us to conclude that the formation of "pure" chromium oxide layers is decisive for reaching high impeding factors. A similar interpretation is suggested by the results shown in the following Figure 5. They indicate that the activation energies for H permeation slowly develop towards 146-167 kJ/mol in the case of coatings under hydrogen/water vapour atmosphere with rising coating temperature. It has not yet been possible to clarify in our investigations how the permeation of point defects (e.g. deviations from the stoichiometry) is influenced in detail. It is interesting to note that, concerning the influence of a "pure" chromium oxide layer, our permeation studies coincide with corrosion studies on these high-temperature alloys where, in fact, the formation of such a Cr 0. layer has also been recognized as decisive for protection. - 6 - Further findings will have to be derived in order to substantiate this working hypothesis. This also appears important because further layer growth might presumably take place through such "defects" which, of course, could also influence the corrosion protection properties of the layer. Moreover - as shown by our experiments - such "defects" are apparently also points of departure for mechanical defects in connection with thermal cycles (10). It can also be seen from Fig. 3 that there is initially no cover on the first start-up of heat-exchanging units, that there is no sufficient permeation barrier for the first few days. For this reason, we have adopted a precoating programme in cooperation with our industrial partners MAN, Munich, and Nukem GmbH for the reproducible production of oxide films by preoxidation using CVD methods (T = 1223-1273 K) with initial impeding factors of approx. 500. When exposed in process gas, partial cooling damage can be detected with subsequent good recovery in the oxidation phases. These layers already show impeding factors H > 1000 after a few days. Post-examinations reveal coherent layers mainly consisting of Cr 0 in this case, too. Much technical effort is presently being devoted to these preoxidation procedures including the study of alternative coating methods with a view to short recovery times. Effective Al 0 layers are being generated, for instance, by pre-alitizing the base material. The quality of a permeation protection layer has so far been characterized by the impeding factor achieved. The tritium problem can be basically regarded as solved with impeding 3 factors of the order of magnitude of H=10 for the steady-state operation of plants producing process gas. But what will happen in reality in the event of load fluctuations with decreases in temperature or even in the case of accidental shutdowns? In this connection, it is important to examine the temperature cycling resistance of the oxide layers. Selective thermal cycling studies with hydrogen as the model gas have so far been conducted on permeation specimens of the materials IN-519, Incoloy 800H and Hastelloy X. The material IN-617 - 7 - is being cycled at present. Two load cases are distinguished in connection with cycling operations: 1. hot start conditions: starting from the oxidation temper- ature T=1223 K, cooling to T=573 K is effected at a temperature change rate of dT/dt = 1 K/min, followed by heating to T=1223 K; 2. cold start conditions: cooling proceeds to room temperature, conditions otherwise equal to hot start cycling. The cooling process involves more or less pronounced damage of the oxide layers. Our detection systems enable us to separate different damage mechanisms, for example dominant microcracking in the temperature range 923 K>T>RT with a peak at T = 473 K, macrocracking delayed in time at room temperature as well as a reversible effect (during both cooling and heating) in the range T 1073 K which we may regard as phase transformation of spinels . Macrocracking and phase transformation occur in permeation layers of IN-519 and Incoloy 800H (both 50 % Fe), increasingly for coatings with a high oxidation potential, but not in Hastelloy X (18 % Fe). The detection system available includes gas chromatography for H-permeation measurements as well as a measuring channel for acoustic emission analysis. Figure £> shows the temperature course of a cycling operation. Starting from T=1223 K, the sample is cooled to room temperature and heated again to T=1223 K after a preset holding time. The values measured by gas chromatography are given in the lower part of the figure. Starting from a small permeation flux § at T=1233 K, the signal drops with decreasing temperature. After reheating to the original oxidation temperature there is a damage-induced rise in the permeation flux by the amount A$. With the simultaneous operation of a measuring channel for acoustic emission analysis, intense swelling acoustic emission occurs in the range T<675 K, which we assume to indicate damage events due to microcracking even after room temperature has been reached. When the temperature rises "rest" results. The first conclusion is that severe damage occurs in the cooling phases. The effects of such damage become noticeable by increased tritium permeation during the restart of a plant producing process gas. Quality assessment of permeation impeding scales has been carried out to date by means of the so-called impeding factor Y\ pt vi o (-\ v" H =0 /$ , i.e. by the ratio of the permeation fluxes of the bare sample to those of the oxidized one. This character- ization can be supplemented by specifying the temperature cycling resistance (TCR) for non-steady state operating con- ditions. To this end a parameter U = A$/$OX is defined, where A$ is the damage-induced change in the permeation flux. U assumes the value U=0 at 0 % damage and the value U = H-1 at 100 % damage, i.e. the value of the impeding factor originally present before cycling. When plotting the acoustic counts during the damage phase of a thermal cycle against the parameter U in log-log scale then a correlation is obtained (Fig. 7) from the large number of measurements on different materials and coating methods. This correlation shows a clear dependence of the TCR on the base material of the sample on which the coating has been grown in situ. Measurements of good TCR can be seen at the bottom left for the materials Hastelloy X and Incoloy 800H at small acoustic counts and U 1. Measurements of poor TCR of the material IN-519 at high acoustic counts and U 10 are top right. Two measuring series on one of the Hastelloy X samples, which was however coated at an extremely unfavourable oxidation potential, are located in the transition region. A cycling operation with a sample of centrifugally cast IN-519 in the tritium permeation test stand will serve as an example of 100 % damage caused exclusively by microcracking (Fig. 8). Starting from impeding factors H>1000 for both hydrogen per- _ 9 _ meation as well as tritium permeation in the opposite direction, the sample was cooled at a low rate of temperature change. Intense acoustic emission already began at T=875 K. At the same time the tritium activity rose in the permeation sample and at T=775 K already displayed a complete loss of impeding effect. The cycling operation was continued after a holding time. Acoustic emission showed an intensity peak in the range of T=475 K. Complete damage to the oxide layer was also established in the case of hydrogen permeation measurements during heating and when reaching the initial temperature again. A study with numerical computations was compiled for a 50 MW plant covering the effect of graduated damage of 100, 90 and 50 % on tritium contamination. Figure 9 shows the time-dependent course of product gas con- tamination at different damage rates of the permeation-impeding oxide scales. The steady-state conditions of the plant (0 % damage) result in contaminations below the statutory limit of 10 pCi/g. Under so-called "hot start" conditions, i.e. after cooling the plant to T=573 K, damage rates of 50 % occur. Subsequently starting heating to T=1223 K does not result in the 10 pCi/g level being exceeded here either. Only a "cold start", i.e. after cooling to RT with 90 % damage rates, exceeds this level for 2-3 days. 100 % damage was only established in the thermocyclical model investigations for the material IN-519. The good regenerability of damaged oxide scales is striking and can be seen from Fig. 9. The current status of our investigations can be described as follows. Suitable methods for highly effective coating have been developed, by means of which reproducibly high - 10 - impeding factors are achieved. The action of the layer has been described in a phenomenologically consistent manner by the introduction of surface reactions and is thus largely clarified /11/, although further supporting data are still required . The aim of further work is the industrially required and economically acceptable precoating of large components including the use of alternative methods as well as the establishment of sufficient temperature cycling resistance for these layers. This also includes improved characterization in connection with methods of metallographic and surface investigations. Suitable methods of industrial fabrication are currently being developed by our partners. Similar problems and thus in part more stringent requirements arise in connection with the tritium problem of the fusion reactor, so that our work described here is also of great interest there. An additional effect of oxide poisons such as sulphuric compounds may have to be considered for the applications envisaged by us (e.g. coal gasification). Our investigations in this field, which may be of significance for the petrochemical industry, have only just been commenced. Literature /1/ R. Schulten, H. Barnert: Entwicklung der nuklearen Prozeß- wärme. Erdöl, Erdgas 3/79, Volume 95, March 1979 /2/ P.S. Flint: KAPL 659 Atomic Power Lab. Dec. 51 /3/ H.D. Röhrig, R. Hecker, J. Blumensaat, J. Schaefer: Nucl. Eng. Des. 34 (1975) 157-167 /4/ W.C. Hagel: Cation Diffusion in Cr 0 . J. Electrochem. Soc. 108 No. 12 ( 1961 ) /5/ R. Ash, R.M. Barrer: Phil. Mag. 4 (1959) 9 /6/ ü. Zink: Jül-Rep. 1880 (1983) /!/ J. Ali-Khan, K.J. Dietz, F.G. Waelbroeck, P. Wienhold: J. of Nucl. Materials 76 and 77 (1978) 337-343 /8/ T.S. Ellemann, L.R. Zumwaldt: Proc. of the Third Topical Meeting on Controlled Nucl. Fusion Vol. 2 p. 763; Conf. 780508 Nat. Techn. Inf. Service U.S. Dept. of Commerce Springfield V.A. 1978 /9/ D. Stöver, H.P. Buchkremer: Paper on "Gase in Metallen" Deutsche Gesellschaft für Metallkunde e.V., Darmstadt, March 28 - 30, 1984 /10/ H. Jonas, R. Heckert, D. Stöver: Symposium on Acoustic Emission Monitoring and Analysis in Manufacturing, New Orleans 1984 711/ H.J. Leyers : KFA Report in preparation Figure Captions Fig. 1: Chemical composition of the high-temperature alloys examined Fig. 2: Permeation values for uncoated and coated high-temper- ature alloys Fig. 3: Histogram of in-situ oxidations Fig. 4: Dependence of hydrogen permeation through oxide films produced in situ on hydrogen pressure Fig. 5: Temperature dependence of hydrogen permeation through Incoloy 802 and oxide layers applied at different coating temperatures Fig. 6: Thermal cycling of an Incoloy 800H sample, response of hydrogen permeation and acoustic emission Fig, 7: Correlation between acoustic emission and the temperature stability parameter U Fig. 8: Thermal cycling of an In-519 sample and effect on tritium permeation Fig. 9: Calculated tritium contamination of process gas in a 50 MW HTR plant in different regeneration phases of the oxide scales Nominalzusammensetzung (Gew. %) Legierung C Hn Si Cr Ni Co Ti Al Nb Fe Mo Andere Knetlegierunq INCOLOY 800 0,05 0,75 0,35 20,5 32 - - - - Rest - 0,25 Cu INCOLOY 800H 0,08 0,75 0,5 21,0 32,5 - 0,4 0,4 - Rest - INCOLOY 802 0,35 0,75 0,1 21,0 32,5 - 0,75 0,6 - Rest - INCOLOY 807 0,08 0,75 0,1 20,5 40,0 8,0 0,45 0,35 - Rest - 5,0 M IN 586 0,05 - - ?5,0 Rest - - - 10,0 0,03 Ce INCONEL 617 0,07 - - 22,0 Rest 12,5 - 1,0 - - 9.0 HASTELLOY X 0,10 0,5 0,5 22,0 Rest 1,5 - - - 18,0 9,0 0,6 W SchleuderguB HK 40 0,45 0,65 1,75 24,5 19,5 - _ - - Rest - MANAURITE 36X 0,4 1,5 1,5 25,0 33,0 - - - 1,0 Rest - IN 519 0,3 0,75 1,0 24,0 24,0 - - - 1,5 Rest - IN 638 0,5 0,5 0,5 26 Rest 15 - - 1 16,5 - 5,0 W IN 643 0,5 0,3 25 Rest 12 0,1 2 3 0,5 0,1 Zr, 9,0 W Fig. 1 Hj-Konientration Ivpm ) HASTELLOY X o D INCOLOY 800H l Incolcy 600 H 6 Tubes (H) - 2 IN 519 1 Tube (K*T) 3 HasteUoy X 10 Tubes (H*Ti i. IN 566 5 Tubes IH) 5 Hastellcy X 6 Disks ID) 6 ^conel 617 2 Tubes (H) Fig. 2 -Fig. 3 .it; 10-' io-' 10'' Wasserstoff - 10'° Tritium 10'"' 10-' os 10" io3 io-' 10-' 10° • PvJ Fig. 4 Fig. 5 _ ... » 1 ° _ x IN-519 o HASTELLOY X • JNCOLOY 800 H ] / TIMEIh) 10' u 10? Fig. 6 Fig. 7 /IpC./g] \ '00% DAMAGE RATE i <° .0' z \ IUMCONT ; a 10' 15 I Ml 20 10 10 » IB 140 160 tlhl Fig. 8 Fig. 9 XA0055832 • GA-A17351 THE EFFECT OF INLET AND OUTLET SHELL-SIDE FLOW AND HEAT TRANSFER ON THE PERFORMANCE OF HTGR STRAIGHT TUBE HEAT EXCHANGERS by D. P. CAROSELLA MARCH 1984 THE EFFECT OF CREEP-FATIGUE DAMAGE RELATIONSHIPS UPON HTGR HEAT EXCHANGER DESIGN M. M. KOZINA, J. H. KING Staff Engineers GA Technologies Inc. San Diego, California, USA and M.BASOL Staff Engineer Combustion Engineering Inc. Chattanooga, Tennessee, USA ABSTRACT Materials for heat exchangers in the high tem- relationship consists of a design locus drawn from perature gas-cooled reactor (HTGR) are subjected DF = 1.0, Dc = 0 to DF = 0.1, Dc = 0.1 to DF to cyclic loading, extending the necessity to design = 0, Dc = 1.0. DF is the fatigue damage and Dc against fatigue failure into the temperature region is the creep damage. A more conservative damage where creep processes become significant. There- relationship for 2-1/4 Cr-1 Mo material consisted fore, the fatigue life must be considered in terms of including factors that degrade the fatigue curves. of creep-fatigue interaction. In addition, since These revised relationships were used in a structural HTGR heat exchangers are subjected to holds at evaluation of the HTGR steam cycle/cogeneration constant strain levels or constant stress levels in (SC/C) steam generator design. high-temperature environments, the cyclic life is substantially reduced. The HTGR-SC/C steam generator, a once- through type, is comprised of an economizer-evap- Of major concern in the design and analysis of orator-superheater (ESS) helical bundle of 2-1/4 HTGR heat exchangers is the accounting for the Cr-1 Mo tubes followed by a superheater of straight interaction of creep and fatigue. The accounting is tubes of Alloy 800H in the central zone of the steam done in conformance to the American Society of generator. The high-temperature components af- Mechanical Engineers Boiler and Pressure Vessel fected by creep-fatigue interaction are the tubing Code, Code Case N-47, which allows the use of the and the superheated steam tubesheet of Alloy linear damage criterion for interaction of creep and 800H. fatigue. This method separates the damage incurred in the material into two parts: one due to fatigue and one due to creep. The summation of the creep- The effects of the revised creep-fatigue damage fatigue damage must be less than 1.0. relationships were evaluated by: (1) calculating the temperature-dependent allowable strain range with Recent material test data have indicated that the the assumed bilinear damage relationship for Alloy assumption that the summation of creep and fatigue 800H; (2) calculating the temperature-dependent damage equals unity at failure may not always be allowable strain range with reduced fatigue allow- valid for materials like Alloy 800H, which is used ables for 2-1/4 Cr-1 Mo; and (3) predicting the in the higher temperature sections of HTGR steam strain range in the critical parts by extrapolation generators. Therefore, a more conservative creep- of finite element calculations for the second or last fatigue damage relationship was postulated for Al- cycle analyzed for each transient to the expected loy 800H. This more conservative bilinear damage number of cycles and hold times. The preliminary analyses indicate that the Alloy The HTGR steam generator is a once-through 800H tubing and tubesheets have predicted strains type. (See Figs. 1 and 2.) It is comprised of a mul- substantially under the allowables based upon the titube helical coil configuration in the economizer- bilinear damage relationship but that the 2-1/4 evaporator-superheater (EES) region, followed by Cr-1 Mo tubing at the hot end inner radius portion straight tube finishing superheater (STSH) in the of the EES tube bundle presents a slightly over- central core of the module. stressed situation. It is believed that there is suf- ficient design latitude to resolve this problem in the Feedwater is supplied to the bottom of the EES continuing preliminary design. It is concluded that by a single side penetration duct and enters the the HTGR-SC/C steam generator design has suf- economizer through 2-1/4 Cr-1 Mo tubes extend- ficient margin to accommodate the more conserv- ing from the feedwater tubesheet located in the ative creep-fatigue damage relationships. discharge helium flow. The feedwater is then di- rected upward through the tubes in the EES bundle INTRODUCTION where it is converted to steam with a small amount of superheat. In high-temperature power plant components, materials are frequently subjected to cyclic loading The tubes exit from the top of the helical bundle extending the necessity to design against fatigue and are routed to the perimeter of the bundle where failure into the temperature region where creep they are anchored. From here they pass through a processes become significant. Creep-fatigue inter- stagnant helium region where expansion legs and action therefore becomes a key phenomenon lim- the bimetallic welds, which join the 2-1/4 Cr-1 Mo iting design life. tubes to Alloy 800H tubes, are located. The Alloy 800H tubes, which are somewhat larger than the EES tubes, are then routed into the active helium In the high temperature gas-cooled reactor flow region in the center of the module and straight (HTGR) steam generators, some of the high-tem- down to the helium inlet plenum, thereby forming perature sections are subjected to creep-fatigue in- the STSH. In this section, the steam temperature teraction. Load changes and shutdowns produce is raised to its discharge condition. Below the he- many temperature and stress cycles throughout the lium inlet plenum, the tubes are led to the Alloy 40-year life of the plant. The two principal materials 800H superheater tubesheet. of construction, Alloy 800H and 2-1/4 Cr-1 Mo, are both operated in their respective creep ranges in certain sections of the component. In addition, The hot helium gas emerging from the core flows due to the load changes, they are both subjected to the steam generator through a lower cross-duct to cumulative fatigue damage. Since the materials and enters the STSH radially through a flow dis- are used to form part of the primary pressure bound- tribution screen. As the gas enters the tube bundle, ary of the system, their use is governed by Section it takes a 90-deg turn and flows upward through III, Code Case N-47 of the American Society of the STSH section parallel to the tubes. The gas Mechanical Engineers (ASME) Boiler and Pressure exits at the top of the STSH and takes a 180-deg Vessel (B&PV) Code (Ref. 1). turn, passes through another flow distribution screen and proceeds downward over the helically coiled tubes in the EES section counterflow to the This paper will discuss the effects of current and steam and water. postulated creep-fatigue damage relationships on the HTGR steam generator design. The gas flow leaves the EES at its lower end, turns and flows radially outward through discharge HTGR STEAM GENERATOR DESIGN ports in the outer shroud and enters the annular region between the heat exchanger module and the In the United States, the HTGR steam cycle/ cavity thermal barrier. The helium then exits at the cogeneration (SC/C) reactor is designed for gen- top end of the annulus and flows to the main eration of electricity and steam. The steam gener- circulator. ation is for chemical process applications. The HTGR-SC/C reactor is similar in concept to the Alloy 800H is employed for the higher temper- 33O-MW(e) Fort St. Vrain and 40-MW(e) Peach ature superheater and tubesheet. The lower tem- Bottom prototype reactors. The Federal Republic perature helical superheater, together with the of Germany's demonstration thorium high-temper- evaporator and economizer sections are designed to ature reactor is based upon the same principles and be manufactured from 2-1/4 Cr-1 Mo material in many similar concepts and technology. The HTGR- the annealed condition. All must be designed and SC/C steam generator design is being developed analyzed to Section III, Div. 1 and the high-tem- by Combustion Engineering Inc. under subcontract perature Code Case N-47 of the ASME B&PV to GA Technologies Inc. Code. BIMETALLIC TO CIRCULATOR WELD OUTER SHROUD HELIUM FLOW I ••••-• •> BAFFLE--- ECONOMISER EVAPORATOR MAIN ACCESS > SUPERHEATER PENETRATION , HELICAL BUNDLE /SUPPORT PLATE STRAIGHT-TUBE SUPERHEATER SUPPORT GRID STEAM GENERATOR / TO LINER SUPPORT " FEEDWATER FLANGE WELD PENETRATION STRAIGHT-TUBE FEEDWATER SUPERHEATER Fig. 1. MK IVA steam generator STRAIGHT-TUBE SUPERHEATER HELIUM INLET DUCT GAS INLET SCREEN SUPERHEATER TUBESHEET SEALWELD SUPERHEAT PENETRATION FLOW RESTRICTOR/ PIPE RESTRAINT Fig. 2. MK IVA steam generator penetration CREEP-FATIGUE DAMAGE time. TD is the allowable time duration from exist- RELATIONSHIPS ing stress to rupture curves for a given stress and the maximum temperature at the point of interest. To account for creep damage during the shakedown Because of the HTGR 40-year cyclical life and period and for the incorporation of damage due to high-temperature requirements, the effects of creep unexpected transients, the allowable duration, TD, and fatigue and their interaction are concerns in is reduced by increasing the value for applied stress steam generator design. Present rules for analyzing from the stress-to-rupture curves by 10%. the creep and fatigue interaction in the existing ASME Code Case N-47 include a "linear" damage Recent experimental data (Ref. 2) has indicated criterion wherein the sum of the creep damage and that the observed sum of creep and fatigue damage the fatigue damage must be less than 1.0 (see Fig. fractions at failure for Alloy 800H metallurgical 3). This method separates the damage incurred in property laboratory test specimens will often be sub- the material into two parts: one due to fatigue and stantially less than unity. This is particularly so for one due to creep. The fatigue damage is given by: typical HTGR heat exchanger design conditions DF = Sn/N, where n = number of creep fatigue that incorporate relatively long hold times. Because cycles accumulated and N is the number of fatigue of this apparent discrepancy between the "linear" cycles to failure at the same strain range but with damage criterion and the observed experimental a zero hold time at the maximum and minimum results, a new relationship was postulated for the strains. The creep damage is given by Dc = 2At/ design of heat exchangers when low strain ranges TD where At is the time at load during the hold and long hold times are experienced. Zt/T,, 0.2 Fig. 3. Creep-fatigue linear damage envelope Analysis of test data has suggested that a "bilin- = hours/Dc). Stress to rupture values are then ob- ear" relationship between creep and fatigue dam- tained from Ref. 1 to find the stress to cause rupture age limits be adopted instead of the 1.0 to 1.0 linear (trSR) at the chosen temperature and at the calcu- summation. In this case, the bilinear loci for Alloy lated allowable hours—(TD). The allowable cyclic 800H would be drawn from DF = 1.0 and Dc = strain range may then be calculated for the monthly 1.0 and intersecting of DF and Dc = 0.1 (see Fig. or weekly cycles as follows. 4). Monthly shutdown cycle: For 2-1/4 Cr-1 Mo, the fatigue curves would be degraded by time-dependent factors which are pres- 0.9 (TCR CTpr ently being developed. These are chosen so that eR = ET "D" values would come out at 1.0. However, more -RT test data are needed, for both Alloy 800H and Weekly cycle: 2-1/4 Cr-1 Mo, under low strain range and long hold time conditions to ensure that even this new _ 2 (0.9) qSR creep-fatigue damage criterion is not unconservative. ET ALLOWABLE STRAIN RANGES FOR where eR = allowable cyclic strain range, HTGR-SC/C STEAM GENERATOR ERT = Young's modulus at temperature, ERT = Young's modulus at room A method to determine the allowable cyclic temperature, strain ranges is described in Ref. 3. Basically, the <7PL = room temperature proportional limit time at temperature is determined and a creep dam- stress, age (Dc) value is chosen between 0.1 and 1.0. The 0.9 = factor required by Ref. 1, allowable number of hours is then calculated (TD POSTULATED DAMAGE CD') RELATIONSHIP FOR ALLOY 800H FOR ALL TEMPERATURES 0.1 0.2 0.3 0.4 0.5 C.6 0.7 0.8 0.9 2n/l\ld Fig. 4. Creep-fatigue bilinear damage envelope 6 APPLICATION TO HTGR STEAM GENERATOR DESIGN In high-temperature components, thermal gra- dients that fluctuate with time lead to three types of strain-controlled loading cycles. These three ge- neric creep-fatigue loading cycles are shown in Fig. 1A) TENSILE HOLD SHUTDOWN CYCLE 5. Curves (A) and (B) are representative of the strain cycles that are produced during startup/shut- down cycles with holds at power. Curve (C) is rep- resentative of the loading cycle during load follow- ing, where fluctuations of the thermal gradient occur while the material remains at high temper- ature. It is worth noting that, although the strain histories, (A), (B), and (C) comprise strain in one direction only, creep relaxation will cause the cor- responding stress cycles to shake down to conditions IB) COMPRESSIVE HOLD SHUTDOWN CYCLE in which both tensile and compressive stresses are produced in the cycle. The important parameter, however, is the strain range. In the present HTGR-SC/C steam generator structural analysis, the thermal gradients are lumped into two types of critical cyclic loadings: a weekly load following cycle and a monthly shutdown cycle. The monthly cycle is based on the projected design TIME total number of shutdowns, which averages ap- (Cl POWER FLUCTUATION CYCLE proximately one a month. Consequently, allowable cyclic strain ranges for Fig. 5. Generic creep-fatigue strain cycles both 2-1/4 Cr-1 Mo and Alloy 800H material were formulated based upon the lumped loading cycles. Two conservative assumptions used are (1) that the steam generator tube maximum metal temperatures do not vary with power changes and (2) that the CONCLUSIONS steam generator goes to room temperature and zero pressure differential at each of the 400 shutdowns The preliminary structural analysis indicates (one per month over the life of the plant). that the more critical components of the steam gen- erator satisfy the ASME Code Case N-47 linear Using the calculated strain range (£R), the allowable creep-fatigue interaction damage criterion. number of cycles (N) can be determined from fa- tigue stress versus number of cycles curves in Ref. Preliminary design analyses of the steam gen- 1. Knowing the desired number of cycles the fatigue erator are continuing based upon the postulated damage factor DF can be calculated by: bilinear creep-fatigue interaction relationship. In- dications are that the Alloy 800H tubing and tube- DF = number of cycles/N sheets have predicted strains substantially less than the allowables based upon the postulated bilinear The creep damage and fatigue damage factors (Dc damage relationship. However, the preliminary and Dp) are then compared to the total allowable stress analysis for the 2-1/4 Cr-1 Mo tubing in the damage factor (1.0 for 2-1/4 Cr-1 Mo and Fig. 4 hot end inner radius portion of the EES tube bundle for Alloy 800H). The allowable strain range can indicates a slightly overstressed situation. It is be- then be determined by plotting strain range (eR) lieved that there is sufficient design latitude to re- against damage as in Figs. 6 and 7 and using 90% solve this problem in the continuing preliminary of the allowed damage. Figure 8 shows the allowed design. Therefore, it is concluded that the HTGR- strain range for Alloy 800H plotted as a function SC/C steam generator design has sufficient margin of maximum material temperature for both weekly to accommodate the postulated bilinear creep-fa- and monthly cycles. tigue damage relationship. 1.1 ALLOY 800H AT 650°C(1200°F) FATIGUE DAMAGED 0.001 1.0 DAMAGE 0.9 LIMIT -WEEKLY LOAD FOLLOWING 0.8 MONTHLY BASE LOADED 0.7 0.6 MONTHLY LOAD FOLLOWING 0.5 0.4 I 0.6 0.7 0.8 0.9 1.0 1.1 CYCLIC STRAIN RANGE x 10~3 (IN./IN.) Fig, 6. Allowable cyclic strain range at 650°C (1200°F) 1.1 MONTHLY LOAD 2-1/4 Cr-1 MoAT538°C (100i°F) FOLLOWING. 1 1.0 - -WEEKLY L@AD FOLLOWING I DAMA 3E 0.9 - I LIMIT < r MONTHLY [ 0.8 BASE L0ADED-~_J 0.7 - 1 0.6 I 1 1 0.6 0.7 0.8 0.9 1.C CYCLIC STRAIN RANGE x 10"3 (IN./IN.) Fig. 7. Allowable cyclic strain range at 538"C (WOO'F) 2.0 ALLOWABLE CYCLIC STRAIN RANGES 1.6 cr> 1.4 I o MONTHLY LOAD X FOLLOWING LU o I 1.2 cr. 1.0 0.8 - WEEKLY LOAD FOLLOWING 0.6 0.4 800 900 1000 1100 1200 1300 MAX METAL TEMPERATURE (°F) Fig. 8. Alloy 8O0H allowable cyclic strain range as a function of temperature 10 ACKNOWLEDGMENT This work was funded by the San Francisco Op- erations Office of the Department of Energy under Contract No. DE-AT03-84SF11963. REFERENCES 1. "Class 1 Components in Elevated Temperature Service," ASME B&PV Code, Section III, Division 1. Code Case N-47-21, December 1981. 2. Johnson, W. R., and D. I. Roberts "Materials Development for HTGR Heat Exchangers," Jour- nal of Engineering Power 105, October 1983. 3. Lewis, A. C, "Allowable Cyclic Strain Ranges for High Temperature Preliminary Design," GA Technologies unpublished data, August 1983. 11 srf No. 23 XA0055833 THE KLINGER HOT GAS DOUBLE AXIAL VALVE 1) by J. Kruschik 2) and H. Hiltgen 3) Introduction The Klinger hot gas valve is a medium controlled double axial valve with advanced design features and safety function. It was first proposed by Klinger early in 1976 for the PNP-Project as a containment shut-off for hot helium (918°C and 42 bar), because a market research has shown that such a valve is not state of present technics. In the first stage of development a feasibility study had to be made by detailed design, calculation and by basic experiments for key components in close collaboration with Interatom/GHT. This was 1981 the basis for further design, calculation, construction and experimental work for such a valve prototype within the new development contract under the support of the MWMV. The stage of knowledge to that time revealed the following key priority development areas: - Finite element stress analysis for the highly stressed high temperature main components. - Development of an insulation layout based on knowledge of Messrs. Didier a firm also working under the support of MWMV. - Detailed experimental tests of functionally important structural compo- nents or units of the valve, partly at Klingers (gasstatic bearings, flexible metallic sealing element, aerodynamic and thermohydraulic tests), partly at Interatom (actuator unit and also gasstatic bearings), partly at HRB in Julich (flexible metallic sealing system, aerodynamic and thermo- hydraulic tests). - Design of a test valve for experimental work in the KVK (test circuit at Interatom) for evaluation of temperature distribution and reliability of operation. - Design of a prototype and extensive testing in the KVK. 1 Presented at the Specialists Meeting on Heat Exchanging Components of Gas-Cooled Reactors of the International Atomic Energy Agency Dusseldorf, 16. - 19. April 1984 2 Klinger Engineering, A-2351 Wiener Neudorf, WienerstraBe 17 3 Interatom, D-5O6O Bergisch Gladbach, Friedr.-Ebert StraBe - 2 - Valve Specification Working/Design pressure 41,9/51 bar Working temperature 900 + 18° C Helium mass flow 35,6 kg/sec Temperature gradient i 2 K/min Max closing time 15 sec Leakage rate 10"1 mbar L/sec Closing cycles per life time 1000 Closing cycles at full temperature 100 Life time over all 40 a Diameter of gas duct 700 mm Valve Description The Klinger double axial valve, Fig. 1, is loaded between the seats with clean pressurized helium in the closed position to prevent leakage from inlet to outlet. Its dimensions are very compact. It fits practically into the hot gas duct. The weight of this valve is roughly 4,2 to compared with 3 to of the equal length of hot gas duct. In case of a sudden accidental pressure loss the valve closes automatically by the flow forces. The special arrangement of the flow path is responsible for the small dimensions and also for the low pressure loss: Accelerated flow at the inlet, followed by a ring-diffusor, leading to a ring channel with the three struts which connect the internal cylinder with the two actuators to the outer shell, followed by a ring- infusor and an outlet diffusor after the second seat. This channel is symmetric and there is no preference to flow direction. This flow path design results in a relatively small seat diameter together with very low pressure loss, in that case 0,025 bar or^r = 1. These values are theoretically calculated, but they are sound as many tests with a model in a flow channel at the Klinger facilities have shown, Fig. 13, 14. The measured value had been^*"= 0,8. But to include adverse influences from fabrication we remain with the calculated value of __^= 1. These very detailed tests will be repeated and completed at the HRB facilities in Julich in the next future for absolute safety. The corrugated flexible metallic sealing element is a further specific feature of this axial valve. This sealing system together with the basic valve design is subject of Klinger patents. The piston rod of the actuator slides in gasstatic bearings, having no metal contact (prevention of welding-together in a clean helium atmosphere) and very low friction. A further specific feature of that valve. A shock absorber decelerates the moving mass at the end of the closing stroke. These basic principles were already included in the very first design early 1976, Fig. 2. The intense development work led to the prototype design, Fig. 1. The valve actuator is shown in detail in Fig. 3. A prototype actuator has been built in the last years by Kraftwerk Union Berlin and extensively tested at Interatom, Fig. 15, 16. The results of those tests are already included in the present design, Fig. 3, and this actuator is part of the test valve for the KVK and has also been manufactured at Kraftwerk Union Berlin. This modified actuator of the test valve was already tested succesfully at Interatom short time ago. Normal closing time is 2 sec. Accidental closing time will be within about 0,3 sec. - 3 - The bends in the gas flow path create a rotation of flow and this could induce a rotation of the piston rod with the shut-off cone. Therefore the piston is guided and can't turn. The sealing elements on it are special spring assisted PTFE-rings. The end positions of the piston rod are indicated pneumatically. The system consists of two orifices in series with different bore sizes, whereby the one with the bigger bore is closed in the end position. The resulting pressure rise is the position signal, see Fig. 6. A redundant-diverse capacitor system operates in the piston rod center line. Both actuators as both valve halves are fully independent in respect of actuation and cooling system. The gas static bearings operate on the fact, that gas flowing through porous bushes under a certain pressure difference creates a gas cushion, which keeps the piston rod in suspension. The load bearing capacity is proportional to the pressure difference. The heavy mass of the piston rod and shut-off cone needs quite a load bearing capacity and to get also a reasonable clearance in the bearings we took high porosity sintered metal bushes. This system worked quite well from the start but we got air hammer in certain cases. The problem could be solved in collaboration with Interatom by extensive experimental and mathematical work. Today we have manufacturing knowledge to produce such sintered bearings with high load capacity within a flow accurancy of 10%, creating no air hammer at all. . Part of the extensive test programme was the evaluation of frequency response of this bearing system. We found out by tests and by mathematical analysis that the resonant frequency is at 20 and 40 Hz. The Klinger frequency test bay is shown in Fig. 17. These low resonant frequencies prevent excitation of vibrations from outside, mainly from the gas-stream. The helium for the gas static bearings is used at the same time as coolant for the internal valve components, Fig. 4 and 5. The high pressure cooling helium is fed into the valve through the three struts, cools first the inner cylinder and flows from the plenum chamber round the actuating cylinder to the gas static bearings, Fig. 4. Than it cools the hollow piston rod on the way from the cylinder to the shut-off cone. In the cone nose it is heated first to prevent too high temperature gradients in the shut-off cone and than it flows along the inner face of that cone, Fig. 5, to the head of it, cooling there the flexible metallic sealing element and than it leaves the valve through the struts. In the closed position of the valve this coolant flow of pressurized clean helium cools the valve down and acts as a barrier to prevent leakage of contaminated helium from inlet to the outlet. The struts are cooled by a separate closed circuit helium loop under low pressure, The flexible metallic sealing system is shown in Fig. 5. It is a corrugated sheet metal part of Incoloy 800 H, electron-beam welded to the shut-off cone head of Inconel 617. A stop on that head prevents excessive load on the sealing element in the closed position. A new element has deeper corrugations and they are at the first shut-off deformed plastically and adapted to cone and seat. When the valve is opened again an elastic return spring action takes place and this is the remaining elastic deflection of the element. It took intensive experimental work to find the best way of fabrication of such large elements, but at the end we get now quite reasonable results. - 4 - The testing of such sealing elements is done at HRB Julich. Fig. 18 shows the test apparatus and Fig. 19 profile recordings of a sealing element undeformed and after the first plastic deformation. To prevent welding together of seat and sealing element, we proposed a detonation gun applied chromium carbide coating on the seat and a silver coating on the element. During the tests we found, that at 750° C the silver coating starts to stick on the chromium carbide coated seat. Coating tests on the other side have shown that an uncoated 800 H should not weld together with a coated seat. The next series of tests will be made now with uncoated sealing elements and a seat with ceramic coating of yttria-stabilized zirconia (ZrO2), because this type of coating is more suitable for higher temperatures. We have also in mind other coating types for the element, if the uncoated one would not work properly. Besides that a general research program for testing coating combinations for the PNP-project is on the way. Results of stress analysis with finite elements All the stress analysis work has been done and is still going on at Kraftwerk Union Berlin (KWU), department TB-TDR. The very high helium temperature of 918° C needs a relatively high loading of the important components. The permissible primary loads -loads from internal pressure and closing force- are dictated by the permissible creep deformations. The secondary loads -loads from hindered thermal expansion- are dictated mainly by the permissible alternating stress for the material. Plausible temperature distributions by hand-calculation had been sufficient for the preliminary design for steady state conditions. But then these data had to be corrected by simplified linear heat flow calculations and thermohydraulic tests in the flow channel, Fig. 13 and 14. At the end we had to proceed to linear elastic and partly elasto-plastic calculations with three-dimensional finite elements for the critical components. The instationary temperature distribution in the warming up and cooling down periods after valve shut-off led to rather high stress concentrations with appropriate strain. The basis for these calculations is the ASME-CODE. Unfortunately there is no chart for the chosen material Inconel 617. First of all such a diagram had to be developed. The material-data are from the manufacturer, from KFA-Julich and from Interatom, Fig. 7. Highest temperature with high primary stress remains after a shut-off only for a very short time. 100 such shut-offs are specified. The boundary line for permissible stress at high temperature is the 1 % creep-strain for 10 hours. Below 800° C it is 2/3 of the stress-rupture curve, below 600° C it is 90 % of the 0,2 % proof stress and below 400° C 2/3 of that. For secondary loads the 3 Sm criterium for the low temperature range was extended up to temperatures above 900° C. Partly plastic alternating stress- strain calculations with relaxation periods inbetween have demonstrated, that the assumption -primary plus secondery stresses equal 2 times the.0,2 % proof stress-is sufficient conservative. This allowes a linear elastic method for the more extensive finite element structures and therefore a comparison with fictive stresses. Note please that this is only valid for the examined components under the specified load cycles. Fig. 8 shows the result. A fictive stress of maximum 2 times the 0,2 % proof stress from primary plus secondary loads leads to a plastic deformation of a few hundreds of one percent in the highest loaded zone after the first cycle (running up and closing down). This value raises due to further stress dis- tribution and relaxation asymptotically to a just slightly higher value. The plastic alternating stress strains go up to a permissible maximum of about 0,06 %. There is no ratcheting after shakedown. - 5 - Finite element stress analysis was made so far for the essential structural components, as central body with struts and inner cylinder, body extension with seat-ring and shut-off cone with flexible sealing element. Instationary heat flow calculations have been carried out for the body extension with seat-ring and for the shut-off cone head. A specified running up and closing down period under extreme conditions had been simulated for these calculations. For the rest of the structural components the stationary temperatures were calculated sufficiently accurate by linear heat flow analysis. Uneven tempe- rature distribution by temperature differences in the hot gas stream are thereby hardly recordable. But extreme assumptions for resulting temperature distributions in the structure, especially in the struts, helped to get a conservative picture of that effect. Due to this very elaborate analysis, the design of those structural compo- nents had to be altered many times. For the seat-ring for instance 65 calcu- lations and 8 design variations were necessary to finish up with the present uncooled shape. The temperature distribution and the resultant linear-elastic calculated fictive secondary stresses are shown in Fig. 9 and 10. The maximum possible fictive primary stresses due to differential pressure and closing force and the resulting deflection are also shown. The influence of an uneven temperature distribution on the struts is demon- strated in Fig. 11. The stresses hereby are very low. These components have now their final shape with the necessary factor of safety. For the shut-off cone also many variants had to be calculated together with the accompaning design alterations, mainly on the head. But making the wall thickness larger due to the closing forces, the secondary stresses came up, needing an alteration of the cooling ducts and so on. Fig.12 shows the present state. Temperature gradients come up to 15 K/mm. This leads with permissible primary stresses to unrealistic secondary stresses of about 10 times the 0,2 % proof stress. The about 1300 drillings for the proper coolant flow are still not taken into account. In the chapter before we have mentioned the difficulties with the silver coating on the flexible metallic sealing element. The temperature limit of this silver coating is 800° C and therefore we had to cool extensively. If we accept an uncoated element -or other types of coatings which stand higher temperatures- we can go up with the temperature and we can use a just slightly cooled shut-off cone, which leads to much lower temperature gradients and therefore to lower secondary stresses. This will be the next set of calculations. But for verification of all the assumptions and theoretical calculations we do need urgently measured parameters. Further test programme To get those parameters we have designed a test valve for experiments in the KVK, Fig. 4 and 6, which is practically a half prototype valve. This test valve has also been built by Kraftwerk Union Berlin and is now ready for insertion in the KVK. The design was going on at the same time as the mentioned structural calculations. We noticed soon, that all necessary design alterations could not be introduced into that test valve, otherwise our test schedule would be in danger. So we uncoupled it. And that was right, because at the present state of calculation it is evident, that due to the complex interconnections only ex- perimentally measured parameters can lead to a final layout for the highly loaded shut-off cone. But we are shure to have in hand a suitable solution in a relatively short time. The test schedule is shown in Fig. 20. The test valve will be first in the state shown in Fig. 6. This narrow channel gives with the mass flow in the KVK the same C<.-values for heat transfer as the full mass flow in the prototype. We should therefore get the real temperatures of all the highly stressed parts. - 6 - Further it is possible to install in this test valve a new designed slightly cooled shut-off cone and measure the temperature distribution there. This will give us the knowledge we need about the necessary parameters. The valve operation can then be tested with the internals to Fig. 4. This picture shows a cooled seat in form of a disc. This was also a preliminary design, but it gave too high stress concentrations. But for the test valve we use it, to run full stroke aad to examine the function of the shock absorber as well as the overall tightness after shut-off. Fig. 21 shows the center part of that valve and Fig. 22 the shut-off cone. Besides these extensive tests in the KVK at Interatom, further experimental work will be done at HRB Julich on the flexible metallic sealing system and on a valve model of final shape in the flow channel for investigation of the pressure loss, friction factor and heat transfer value on important sections. A further development programmeis started for research on the coating side specifically tailored for the PNP-conditions. The Insulation concept The insulation concept has been laid out in collaboration with Messrs. Didier. Thi: concept is used in the test valve and will be used also in the prototype, if it shows to be sound. At the inlet and outlet side stuffed A12O3-fibres are used within liners for the gas duct. Between the seats and the center part with the struts vacuum-formed A12O3-parts within liners will be used. The center part itself bears a linerless type of insulation. This insulation is built up of moulded shapes on the inner and outer ring channel surface. Behind these shapes are vacuum-formed rings and between them and the steel walls is a cushion of stuffed Al203 fibres. Summary The preceding chapters have shown the intense development work, carried out by Klinger, by Interatom, by HRB-Julich and by Kraftwerk-Union, Berlin. Thanks to the excellent coordinating work by Interatom, all development lines started and finished at the right time and with the right results. As demonstrated, there are only a few points, which wait for a final solution, mainly the shut-off cone, but we are already on the way with it. The same can be said about the final concept for coatings. We have in hand some solutions but we have to find out the optimum by the coming experimental work. As far as we can foresee the time schedule presented can be kept working, so that 10 years after the first design sketch a prototype valve for these extreme conditions will be on test. KLINGER Pig.1.Prototype of the KLINGER double axial valve KLINGER Fig.2.First rough draft of the KLINGER double axial valve KLINGER Cooling helium Pressure helium for actuator Fig.3.Actuator KLINGER Strut cooling in Strut cooling out Cooling helium Connections for cooling cf tail I Pressure helium for_gclugtor Pressure helium for Gctuator Fig.4.Test-Valve with internals for function and thightness tests KLINGER Cooling helium Fig.5.Shut-off cone with flexible metallic sealing element KLINGER Front position indication Fig.6.Test-Valve with internals for temperature measurements 3ft Kraftwork Union Kraftwork Union I Temperature °C 300^ Fig.9. Instctionary temperatures Fig.10. Stresses and deflections. Kraftwark Union Kraffwork Union loaded Tempercture difference unloaded agcinst the outer shed Stress N/mm2 Fig.11. Deflections due to uneven temperature Fig.12. Stresses and deflections. distribution. Kraftv/ark Union Kraftwark Union Plastic alternating stress s Fig. 7 Permissible primary stress for Fig* 8 Design limits Fig. 13 Flow channel for aerodynamic and Inconel 617 according to ASME' thermohydraulic tests on a valve-model \ Fig. 14 Valve model Fig. 15 Actuator unit (Interatom) Fig. 16 Test chamber for the actuator unit at Interatom ERB-JULICH Fig. 17 Test bay for measuring frequency Fig. 13.Test-bay for hot helium test with the Fig. 19.Profile of the corrugated flexible response sealing element at. HRB Julich metallic sealing element tf&Q 1931 1982 1983 1984 1985 1986 1987 1988 Engineering attendance by ~ GHT Preliminary testa Structural component test3 Testa with the Tast-valve Design and preparation for fabrication rabrication Exparl.T«ntal tasts Prototype Valve Design i Preparation for fabrication T" IT" ' ami '-. '-rsj1^1 — _ —-» — —» — —• — ^ Fabrication Experimental teats Evaluation and documentation "1 Series Valve Final design and quotation (bv aanufacturer) Actual state (Doc.33) Planned stata (1981) Fig. 20. Time schedule for development and test of a PNF-Hot-Gas-Vaive DN 7C0 Fig. 21.Center part with struts and Fig. 22.Shut-off cone inner cylinder No. 24 XA0055834 TWO LAYERS THERMAL INSULATIONS TESTS FOR DESIGNING OF HOT GAS DUCTS by T. Nakase, S. Midoriyama, K. Roko, A. Yoshizaki KAWASAKI HEAVY INDUSTRIES, LTD. ABSTRACT Coaxial double walled piping is planned to be used for a primary cooling system piping of the Very High Temperature Gas-cooled Reactor (VHTR) of JAERI. The piping consists of an outer pressure pipe for the reactor inlet gas flow and an inner pipe with internal insulations for the reactor outlet gas flow. The internal insulations are designed to consist of two layers; metal insulation is in the extremely inside for the higher tempera- ture gas and fibrous insulation is between the inner pipe and the metal insulation. The thermal characteristics of inner pipe with two insulation layers are necessary for the designing of the primary cooling system piping. Authors performed thermal characteristics tests by using Kawasaki's helium test loop (KH-200), which has two hot gas ducts test sections of 267.4 mm diameter and 5,000 mm length with simulated two layers insulation. The one is installed in a horizontal position and the other is in a vertical position. The tests were conducted at the temperature of 500 to l,000°C, the pressure of 20 and 40 kg/cm2G, and the flow rate of 100 and 200 g/s. The distributions of temperatures and heat flux on the surface of the ducts are confirmed to be within an allowable range. The test results were analyzed, and useful design data of the metal insulation and the fibrous insulation were obtained. 1. Introduction The Very High Temperature Gas-cooled Reactor (VHTR) of JAERI is planned that the reactor outlet temperature is 950°C, the reactor inlet temperature is 395°C and the pressure is 40 kg/cm2G. For this operating condition, the primary cooling system piping has been designed as the coaxial double walled type shown in Fig. 1, which consists of an outer pressure pipe for the reactor inlet gas flow and an inner pipe with internal insulation for the reactor outlet gas flow. In order to keep the structual integrity of the inner piping, it is protected from very high temperature by the internal insulation and the continuous cooling with the reactor inlet gas flow. The internal insulation is divided into two layers by a partition pipe to suppress a natural convection flow across the insulation. To confirm the thermal performance of the two layers insulation, thermal characteristics tests were performed on the condition of helium gas flow by using Kawasaki's helium test loop (KH-200), including two hot gas duct test sections shown in Fig. 2. — 1 — 2. Description of the test sections and parameters (1) Hot gas duct test sections Two hot gas duct test sections were fabricated simulating the inner pipe of the coaxial double walled piping of VHTR. These two test sections have the same dimension except fibrous insulation's packing density. The diameter of the test section was 267.4 mm and the length was approximately 5,000 mm including a bend pipe. The internal insulation between the pressure pipe and the liner con- sisted of two insulation layers with a partition pipe. The metal insulation made of inconel 600 was used in the extremely inside insulation layer for the higher temperature gas, and Kaowool made of A&2O3 and SiO2 was used in the outside insulation layer. The core section of metal insulation was made of 32 foils, each 0.05 mm thickness. The insulation consists of alternate layers of plane and dimpled foils. Three sides of the core section was covered with 0.8 mm thick cover plates. The cover plates were welded in each other. One side of the metal insulation was opened to make helium gas passage freely. The total thickness of the metal insulation was 25.3 mm. A gap between the inner cover plate and the liner was 2 mm, and a gap between the outer cover plate and the partition pipe is 2.4 mm. Z shaped transition pieces were welded on the outside of the liner and on the inside of the partition pipe. The end cover plate of the metal insulation is welded on the inside of the partition pipe. Thus there is no bypass flow through the gap between the metal insulation and surrounding pipe. In the case of outside insulation layer, Kaowool was packed between the pressure pipe and the partition pipe. The packing density of the test section "A" is 300 kg/m3, and the test section "B" is 250 kg/m3. The total thickness of the fibrous insulation was 44.5 mm. Conical shaped transition pieces were welded on the inside of the pressure pipe to suppress a bypass flow and to support the internal insulation. Two thermocouples were inserted into the center of the liner pipe to measure the flowing helium gas temperature. In order to measure the tempera- ture distribution of the insulation, 33 thermocouples are installed on the outside surface of the partition pipe, 12 thermocouples on the outside surface of the transition pieces and 27 thermocouples on the outside surface of the pressure pipe. (2) Test conditions and measurement Thermal characteristics tests were performed by using the test loop. At first, the hot gas duct test section "A" was installed in a horizontal position and test section "B" was installed in a vertical position. The tests were conducted ranges between 500 and l,000°C of the heater outlet helium temperature. The test pressure was 20 and 40 kg/cm2G with the flow rate of 100 and 200 g/s. Then the position of the test section "A" and "B" were exchanged in each other, and tests were continued. - 2 - The temperature distribution in the insulation layer and at the surface of the pressure pipe were measured with thermocouples and a radiative thermo- meter. The heat flux in the vicinity of the thermocouples on the surface of the pressure pipe was measured with a heat flux meter. 3. Test results (1) Temperature distribution in the insulation layer and on the surface of the pressure pipe Figs. 3 and 4 show the temperature distribution of the test section "A" and "B" respectively at the gas temperature of l,000°C, the pressure of 40 kg/cm2G and the flow rate of 200 g/s. In the case of the test section "A" in the horizontal position, the average temperature of the partition pipe was approximately 66t)°C, and the circumferential and longitudinal temperature distribution was within a range of ±25°C of the average temperature. The outside location of metal insulation joint section was approximately 30°C higher than the other section. This was caused by the thermal conduction of the metal of the insulation connecting section. The average temperature on the surface of the pressure pipe was approximately 195°C, and the circumferential and longitudinal temperature distribution was within a range of ±15°C of the average temperature. No hot spot was observed. In the case of the test section "B" in the vertical position, the tem- perature distribution was almost same as that of the test section "A". Fig. 5 shows the temperature distribution at the transition piece of the test section "A" at the gas temperature of approximately ls000°C, the pressure of 40 kg/cm^G and the flow rate of 200 g/s. The temperature dis- tribution at the transition piece was close to linear. The circumferencial temperature difference was within 15°C. As the temperature difference of the pressure pipe in the vicinity of the transition piece was approximately 5°C, the effect of the transition piece on the temperature distribution of the pressure pipe was low. Fig. 6 shows the heat flux distribution of the test section "A" in the horizontal position. Each heat flux was measured at the pressure pipe near the thermocouples installed on the pressure pipe. The average heat flux was approximately 2,060 kcal/m^h and its distribution range was approximately within ±300 kcal/m^h of the average heat flux. There was no abnormal heat flux. Fig. 7 shows the temperature distribution on the pressure pipe surface of the test section "A" measured by a radiative thermometer. Test conditions were the same as shown in Fig. 3. Both upper side pictures (picture 1A and 2A) show the location of the measurement. Center line of each graph in the lower side pictures (picture IB and 2B) shows a temperature of 190°C, and one scale shows 10°C. The temperature distribution measured by the radiative thermometer was within 180 to 210°C. A peak temperature of 240°C in the picture IB is an influence of an emission of a paint (emissivity = 0.99) of — 3 — + mark shown in the picture 1A. The lower flat temperature distribution in the picture 2B is an influence of a clamp for a pipe hanger. Other sharp negative peaks shown in the picture IB and 2B are due to an influence of compensating lead of thermocouples. Therefore, no hot spot was detected, and this result agrees with the result shown in Fig. 3. In case of the test section "B", almost the same results were also obtained. Further, to confirm the effect of the vibration, the authors performed a seismic test by using the test section "B" and compared the thermal dis- tributions before and after seismic loading test. Both the temperature distributions agreed well. (2) Effective thermal conductivity of the insulation layers Effective thermal conductivity at each test positions were obtained by using the heat flux, partition pipe surface temperature, pressure pipe surface temperature and room temperature. The calculation method of the effective thermal conductivity is described in Appendix. The estimated effective thermal conductivities of metal insulation and fibrous insulation were plotted in the Fig. 8 and Fig. 9. The estimated effective thermal conductivities were arranged by using simple regression. 4. Discussion (1) The regression lines of the effective thermal conductivity Fig. 10 shows the regression lines of the effective thermal conductivity of the metal insulation of the test section "A" in the horizontal position. The regression line at the helium pressure of 40 kg/cm^G and the flow rate of 200 g/s is higher than the others. The conductivity at the pressure of 20 kg/cm2G shows about 10% lower value than that of the test pressure of 40 kg/cm^G, and this is thought to be the gas density difference. As the conductivity at the flow rate of 100 g/s and pressure of 40 kg/cm^G is a little lower than that of the flow rate of 200 g/s and pressure of 40 kg/cm2G, the difference of flow rate is considered to be small effect. Fig. 11 shows the effective thermal conductivity of the metal insulation of the test section "A" and "B" at the helium pressure of 40 kg/cm2G and the flow rate of 200 g/s. The test results in the horizontal and vertical posi- tion showed a good agreement. The test results of the test section "B" shows relatively higher thermal conductivity than those of the test section "A". Fig. 12 shows the effective thermal conductivity of the fibrous insula- tion at the helium pressure of 40 kg/cm2G and flow rate of 200 g/s. The results were slightly difference each other, but it is thought to be no remarkable difference due to test positions and insulation packing density. - 4 - (2) Comparison of the thermal conductivity Fig. 13 shows a comparison of the thermal conductivity between afore- mentioned flowing helium test and the stagnant helium test of KHI. The representative lines used in the former test is shown in Fig. 11 and Fig. 12. The stagnant helium test was conducted by using the full sized diameter test section of the inner piping of VHTR. The internal insulation consisted of two insulation layers; metal insulation was used for the higher tempera- ture insulation layer, and Kaowool was used for the low temperature. The test section length was 4,500 mm. Helium gas was charged in the test section and heated by an electric heater. The test was conducted at the helium temperature of l,000°C and the helium pressure of 40 ^ The measured data of KH-200 almost agreed with those of the full sized stagnant test data in both cases of the metal insulation and the fibrous insulation. 5. Conclusions The following conclusions are obtained- (1) The temperature distribution with two layers thermal insulations is almost uniform and no hot spot is observed. (2) Useful design data of effective thermal conductivity of the metal insulation and fibrous insulation are obtained. Reference (1) T. Nakano, T. lino, T. Hagiwara, S. Takano, High Temperature Piping Insulation Testing. FAPIG No. 94, 1980-3 (in Japanese). - 5 - Appendix. Calculating method of the effective thermal conductivity _ The heat balance model of the two layers thermal insulation system is shown xn the figure at the lower part. The heat balance of the system is dxfxned as following equations. Q = q • TT • Dp0 = aHe. TT. Du(THe - Ti) Partition - T ) 2 pipe Pressure Pipe Liner (T2 - T3) a air AD —-i A V 2TTX TT(T3 " T4) 2TTAF o a c 2TTAT " T6) = »air * TT • Dpo(T6 - Tair) where Q : heat flux per unit length aHe : heat transfer coefficient at the inside surface of the liner aair: heat transfer coefficient at the surface of the pressure pipe AM : effective thermal conductivity of metal insulation AF : effective thermal conductivity of fibrous insulation ^£ : thermal conductivity of liner As : thermal conductivity of pertition pipe Ap : thermal conductivity of pressure pipe Effective thermal conductivity is obtained by substituting measured q (kcal/m^h), THe, T4, T6 and Tair into above equations. - 6 - OUTER PIPE -FIBROUS INSULATION INNER PIPE PARTITION PIPE INSULATION LINER "^W. •: •'. •:•:•:•:•*:• ']•: 77 HIGH TEMP. He GAS 950°C 03 LOW. TEMP. He GAS FIG. 1 CONCEPTIONAL DRAWING OF PRIMARY COOLING SYSTEM PIPING FOR VHTR 4510 to jcsl j-e- -METAL LINER INSULATION -FIBROUS •TRANSITION PIECE PARTITION PIPE INSULATION PRESSURE PIPE -METAL INSULATION \-FIBROUS INSULATION METAL INSULATION (TEST SECTION "A" AND "B") INNER COVER PLATE : INCONEL 600 DETAIL "a' OUTER COVER PLATE : INCONEL 600 CORE SECTION : INCONEL 600, 0.05mm THICKNESS x 32 LAYERS T0P MEASURING POINTS FIBROUS INSULATION (KAOWOOL) He PACKING DENSITY, TEST SECTION "A"; 300 TEST SECTION "B"; 250 kg/m3 BOTTOM FIG. 2 SECTIONAL DRAWING OF HOT GAS DUCT TEST SECTION "A", "B" - 7 - 1000 x(He GAS) 900 800 700 o UJ 600 (PARTITION PIPE SURFACE) Z) < 500 LU 400 LU 300 200 Be B PARTITION (PRESSURE PIPE SURFACE) -PIPE 100 .PRESSURE \PIPE n in ijn BOTTOM LINER rt m to FIG. 3 TEMPERATURE DISTRIBUTION OF HORIZONTAL TEST SECTION "A" (He GAS TEMPERATURE 1000°C, PRESSURE 40 kg/cm2G, FLOW RATE 200 g/s} 1000 x (He GAS) 900 800 _ 700 S o So © 600 (PARTITION PIPE SURFACE) < 500 UJ 400 300 200 PARTITION (PRESSURE PIPE SURFACE) PRESSUREl o TOP in ii n LEFT"* 'BOTTOM -LINER FIG. 4 TEMPERATURE DISTRIBUTION OF VERTICAL TEST SECTION "B" (He GAS TEMPERATURE 1000°C, PRESSURE 40 kg/cm2G, FLOW RATE 200 g/s) 700 600 500 L> cc 400 cc 300 LLJ (- 200 TEST POSITION HORIZONTAL PARTITION He GAS TEMP. 1000°C vPIPE 100 TEST PRESSURE 40 kg/cm2G \ PRESSURE \PIPE FLOW RATE 200 g/s \TOP x V LEFT O 93 :50!il7l 293 KAOWOOL D 'BOTTOM METAL 'LINER INSULATION •=* He GAS FLOW FIG. 5 TEMPERATURE DISTRIBUTION AT THE TRANSITION PIECE AND VICINITY; TEST SECTION "A" 3000 E o c e 2000 © c e X ©5 © _i o U- < 1X1 1000 POSITION HORIZONTAL He TEMP. 1000°C PARTITION TEST PRESSURE 40kg/cm2G -PIPE FLOW RATE 200 g/s PRESSURE 0 \TOP in I n LEFT© BOTTOM LINER FIG. 6 HEAT FLUX DISTRIBUTION OF THE TEST SECTION "A' - 9 - (PICTURE 1A) (PICTURE 2A) o I (PICTURE 1B) (PICTURE 2B) FIG. 7 TEMPERATURE DISTRIBUTION ON THE PRESSURE PIPE SURFACE MEASURED BY A RADIATIVE THERMOMETER MEASURING POINTS • TOP u G LEFT A BOTTOM > O D 2 O u Ior LU X LU TEST SECTION > "A" REGRESSION LINE OF Xeff (- POSITION HORIZONTAL LU He TEMP. 1000°C TEST PRESSURE 40kg/cm2G FLOW RATE 200g/s 0.00 100.00 200.00 300.00 400.00 SOO.OO 600.00 700.00 BOO.00 900.00 1000.00 MEAN INSULATION TEMPERATURE (°C) FIG. 8 CALCULATION RESULT OF EFFECTIVE THERMAL CONDUCTIVITY (METAL INSULATION) MEASURING POINTS , to • TOP ©LEFT JE - ^BOTTOM 0.7 2 ' (Kc a ~IVIT \ p REGRESSION LINE OF Xeff 1— (D CJ in- Q ° 2 O \ u n < o \ ° Ai £ X S i a*^?,———^ JJ H S- TEST SECTION "A" — POSITION HORIZONTAL o He TEMP. 1000°C ULUL » 2 J. —- TESTi PRESSURE 40kg/cm G 1X1 o FLOW RATE 200g/s 1.0 0 0.00 100.00 ZOO.00 300.00 400.00 500.00 600.00 700.00 BOO.00 SOO.OO 1000-00 MEAN INSULATION TEMPERATURE (°C) FIG. 9 CALCULATION RESULT OF EFFECTIVE THERMAL CONDUCTIVITY (FIBROUS INSULATION) - 11 - 0.7 0.7 TEST PRESSURE 40kg/cm2G u FLOW RATE 200 g/s u 0.6 0.6 o JZ E o o — 0.5 0.5 0.4 3 0.4 O Z O o O o 0.3 0.3 EC cr UJ X I 0.2 0.2 TEST SECTION "A", HORIZONTAL 40kg/cm2G, 200 g/s H O LU TEST SECTION "A", VERTICAL 40kg/cm2G, 100 g/s LU n 1 £ 0. u_ U.I LJJ TEST SECTION "B", HORIZONTAL 20kg/cm2G, 100 g/s TEST SECTION "B'\ VERTICAL 0 200 400 600 800 1000 0 200 400 600 800 1000 MEAN INSULATION TEMPERATURE (°C) MEAN INSULATION TEMPERATURE (°C) FIG. 10 EFFECTIVE THERMAL FIG. 11 EFFECTIVE THERMAL CONDUCTIVITY OF THE CONDUCTIVITY OF THE METAL INSULATION OF METAL INSULATION TEST SECTION "A" (HORIZONTAL) 0.7 0.7 He PRESSURE 40kg/cm2G FLOW RATE 200 g/s o 0.6 0.6 JZ E CO u 5 - 0.5 >- > °' 0.4 a Q 2 2 O O u O -J 0.3 0.3 < UJ UJ METAL INSULATION 0.2 KH-200 TEST UJ TEST SECTION "A", HORIZONTAL STAGNANT He TEST OF KHI o TEST SECTION "A". VERTICAL u S o.i FIBROUS INSULATION TEST SECTION "B", HORIZONTAL E o.i KH-200 TEST TEST SECTION "B", VERTICAL STAGNANT He TEST OF KHI I i I i I i I i J | , I I I T 0 200 400 600 800 1000 0 200 400 600 800 1000 MEAN INSULATION TEMPERATURE (°C) MEAN INSULATION TEMPERATURE FIG. 12 EFFECTIVE THERMAL FIG. 13 COMPARISON OF CONDUCTIVITY OF THE THERMAL CONDUCTIVITY FIBROUS INSULATION - 12 - No. 25 XA0055835 IAEA Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors Diisseldorf, 16-19 April 1984 Status of the development of hot gas ducts for HTRs H. Stehle, E. Klas INTERATOM GMBH, FRG 1. Introduction In the PNP nuclear process heat system the heat generated in the helium cooled core is transfered to the steam re- former and to the successive steam generator or to the intermediate heat exchanger by the primary helium via suitable hot gas ducts. The heat is carried over to the steam gasifier by the intermediate heat exchanger and a secondary helium loop. In both the primary and the secondary loop, the hot gas ducts are internally insulated by a ceramic fibre insu- lation to protect the support tube and the pressure housing from the high helium temperatures. A graphite hot gas liner will be used for the coaxial primary duct with an annular gap between support tube and pressure shell for the cold gas counterflow. A metallic hot gas liner will be installed in the secondary duct. 2. Operational data and design criteria The most important operational data for dimensioning and design are as follows: Primary loop Secondary loop He-temperature 950° C 900° C He-pressure 4 0 bar 4 2 bar Mass flow 37 kg/s 37 kg/s Flow velocity 60 m/s 60 m/s Heat loss 50 kW/m 10 kW/m Outer cooling coaxial flow with free convection helium at 300° C and radiation to ambient air Table 1: Operational data for hot gas ducts The most important design criteria concern both, primary and secondary ducts: - Compensation of thermal expansion - Integrity against pressure and temperature transients during normal and upset operating conditions - Homogeneous insulation to avoid thermal loads on the metallic supporting structures - 30 years lifetime - Feasibility of inservice inspection of the supporting tube and the pressure housing - Dismantling and replacing possibilities in the case of failure - Demonstration of the integrity for the licensing proce- dure (analytical and experimental) In addition in the case of the primary duct the leakages between hot and cold flow have to be minimized. -3- 3. Design and construction of the hot gas ducts 3.1 Primary hot gas duct The reference primary duct developed by INTERATOM is shown in Fig. 1. Fibre mats consisting of aluminium oxide are wrapped round the graphite hot gas liner. Graphite foils inserted between the single mats are used as radial convection barriers. A metal supporting tube separates the hot gas from the cold gas.. The liner is positioned radially and axially in the support tube by high density aluminium oxide supports. These are arranged in two cross-sections for each liner assembly unit. Four ceramic ball supports are located in each cross-setion for the radial alignment. The lower two supports are rigidly set, the upper two are positioned with a spring and limit stop. One cone-shaped support in each cross-section takes care of the axial positioning. All support elements are inserted from the outside of the supporting tube through corresponding windows. After installation, the inserts are welded to be gas-tight. The single liner units are connected by overlapping grooves with a calculated axial clearance at ambient temperature. This clearance will be closed at design temperature. In principle the same design is used for two compensating units in the hot gas duct between reactor and heat exchanger The total unit, hot gas duct and compensating unit are inser ted into the pressure retaining shell with an annular gap for the cold helium backflow. This pressure shell is insula- ted at the outside. 3.2 Secondary hot gas duct The secondary duct consists of a inner section with a metallic liner and fibre insulation for flow guidance and heat insulation, and a outer shell for pressure containment. This concept has already been applied for different high temperature loops. The significant advantage of this design to -4- is based on two facts: - The simple procedure for dismantling and for exchange of internals - Accessibility of the pressure housing for inservice inspection Fig. 2 shows the reference secondary hot gas duct, developped by INTERATOM. The different parts viewed from the inside to the outside are: - The metallic hot gas liner - The perforated metallic tube for the inner limitation of the insulation - The ceramic friction bearing spacers between the two tubes to exclude the danger of friction welding. The annular gap allows a pressure balance between the hot gas channel and the insulation - The wrapped fibre mats of 95 % aluminium oxide - The intermediate tube to reduce the radial convection space - The wrapped fibre mats of 55 % aluminium oxide - The support tube - The V-shaped thermosleeves to limit the axial convection and to centre the flow guidance tubes - The support tube for all internals - The pressure tube into which the assembled support tube is inserted as a slide-in unit. Each internal assembly is sealed and fixed at the flange of the pressure tube. This system, including the slide-in technique, is also used for the secondary loop elbows and compensators (Fig. 3). -5- 4. R + D work The R + D work for the development of the hot gas ducts can be divided into four main tasks: a) Material development and testing b) Component part tests c) Scaled down component tests d) 1 : 1 scale component tests In the meantime, the materials, metals, graphite and ceramics, have been investigated with good results. The tests listed under b) to d) are carried out under all expected demands such as steady state thermal loads, transient temperature and pressure conditions and, in some cases, superimposed additional mechanical loads. The support elements of the primary duct, the ceramic balls and cone shaped parts, were tested with temperature gradients and transients and, in addition, with alternating mechanical loads. These tests demonstrated the integrity of these support elements. A similar test programme will be carried out for the pri- mary graphite liner. A suitable test rig is under construction. Test operation will start in summer 84. The assembly method for wrapping the fibre mats was developed during preliminary tests and the specific assembly data for the applied insulation qualities were optimized. The fibre mat insulation can be compressed to such an extent that it is almost possible to exclude the formation of cavities during the lifetime. The extreme requirements in the case of accidents are gene- rally simulated during the component part tests. The main objective of the large component tests is to demonstrate the integrity of the overall assembly under original operating conditions. -6- Important details for improving the design can be obtained during the assembly work of the component test sections. The efficiency of the insulation was proved during thermal examinations in the high pressure helium channel of the Nuclear Research Centre Jiilich. First results indicated that the open insulation systems were severely impaired by free convection, resulting in thermal loads on the metallic support structures. There- fore the insulation had to be modified with adequate con- vection barriers. With reference to the secondary hot duct insulation a lot of experience and test results had been obtained from the preceeding tests for the hot gas piping of the High Temperature Helium Test Loop KVK at INTERATOM and from other test facilities. As a result, it was possible to take these test data into account for the planning and design of an originally-scaled test section. It was there- fore possible to install such a large test component of 6 m in lenght in the KVK before the first start-up period of the test loop. In the meantime this component has endured about 5000 hours of operation under various conditions without any indications of failures. The test component was equipped with thermocouples in different axial cross-sections. In each measuring plane, chromel-alumel thermocouples were installed over two per- pendicular diameters in different radial positions. The corresponding outer wall temperature can be measured using resistance thermocouples. The measurements were carried out in the KVK under different operating conditions between helium temperatures of 400° C to 950° C and helium pressures of 14 bar to 4 0 bar. Figs. 4 and 5 present the radial temperature profiles in two different measuring planes for a helium temperature of 936° C and a pressure of 20 bar. The two measuring planes are situated upstream (B) and downstream (B1) of the axial extention gap between two liner units. The radial -7- temperature profiles are pointed out in all four angular directions. There are small variations for the different directions due to geometric disalignments of the metallic structures, possible inhomogeneities and convection in the insulation. A remarkable temperature decrease can be noted across the annular gap between the support tube and the pressure shell. This decrease is different in the two measuring planes. A definite explanation of this difference can be given after improvement of the alignment during dismantling. Fig. 6 shows the outer wall temperature of the pressure tube as a function of the helium temperature. The mean values of the four measured temperatures in each measuring plane are marked together with the maximum deviations. All these deviations are smaller than 10 % of the mean value, an indication that hot spots do not exist. The azimuthal temperature distribution around the pressure tube surface is shown in Fig. 7 for different helium tem- peratures. The curves represent the relative deviation in percent from the mean temperature. There is a strong asymmetry in the vertical direction for low helium tempera- tures decreasing with increasing helium temperature. This behaviour clearly shows the influence of free convection in the insulation, which is characterized by a modified 4 Raleigh number for porous materials proportional to p2/T . The specific heat loss to the environment as a function of the helium temperature in two measuring planes is shown in Fig. 8 in W/m2 of the pressure tube surface and in W/m tube length. The upper curves represent the total heat loss, the lower ones the convection ratio. Summarizing all experimental and operational results, the integrity of the component has been demonstrated and the design calculations have been confirmed as being conserva- tive. The measured heat loss at helium design temperature was less than 70 % of the calculated value. 6000 I-,—, , I; 7N X X X ! *\ X x X VK X>C"TN x XX A , xxx xxX-^^-yvM O CO S CD O o O oo "Si \/. '/. i-r-'X'j-K XXX X X X > r'T A VX rh X x x -' X x x XViX \->X AXXX X X X X X fh X ^ Pressure vessel Support structure — Insulation Axial support element Radial support element Liner Displacement body Detail B Detail C Cross-section A-A Test Section Primary Hot Gas Duct Fig-1 5970 Insulation Displacement Space Cross Section A-B Support Tube Pressure Vessel 15Mo3 Liner Incoloy 800H Secondary Hot Gas Duct with Metallic Liner and Fibre Insulation Fig.2 \ s Secondary Compensator Fig. 3 -2- The conversation of these energy resources to gaseous and liquid products enables us to do so. Of all the nuclear reactors developed today, the high temperature reactor is predestinated to play a key role, as it can supply the heat which is necessary for the conversion processes at the required high temperatures of between 800° C and 950° C. In conjunction with special process technology, this leads to a considerable reduction of the pollutant emissions of SO , CO« etc. and of dust. Above all, the CO- emission, which is much lower than in autothermal gasification processes, deserves special mention, because of the C0_ influence on the temperature increase at the earth's surface. 2. The Plant Concept In order to clarify the functions of those components which are the subject of the following presentations, I would like to consider the PNP plant with both gasification processes, namely steam gasification and hydrogasification of coal, in more detail. Let us first consider the steam gasification of coal (Fig. 1). The reaction of steam with hard coal requires high temperature heat since it is endothermic. For this process a secondary helium loop is advisable for safety reasons. The secondary helium is heated to 900° C in the He/He inter- mediate heat exchanger and enters the gas generator at approximately this temperature. The helium is cooled to around 815° C here because of the carbon-steam reaction. The helium leaving the gas generator is cooled in the process steam superheater and then conveyed to the steam generator. tit O n—• 1-0 1 i US OH R • RUCK R U O CO _1 <\J 0 \ a. hD I O cr /// Vtil a ** !3 $^ CD / C\J > "S. 7/ LO CL O UJ \ /// a I V O CO a: z UJ m o cr UJ >• 2 u. i- • *~ _i z cr ZWISCH E !• j tj i C~i i . j *T~ O »- Z Q UJ a »- cc m r _j a x Q: or D: »- ec UJ o u v) uj D > »z— LkJo —O U. ZD 3 UJ O *- =3 »- 2T or x r** x n UJ (J i_> O •* i* UJ z z a: a: J — LU t- A / RflDIU S — — O • JZ / * ** • • tx. 1 T Ill I 3. 6 4. 0 4 • < c %'•'//////////// pi* C\J a L» u —i DO - - a 11 m a: z z JQ. « « 1 2 3 > • CD CM n r iij x: UJ LJ t— Z CD UJ Z —' o tt: Z3 _i CM — ~ — ~ CVJ to O •«• co a) co •) — X X X X a • a a D 0021 0001 008 009 OOfr 002 o) yniuy3dW3i becondary duct test component Radial LJ _ t smperature distributions Measuring plane B' g- K CZZ) I ID —a a. O I r • RUCK R a CD OJ _J OJ J CO ( '* * f \ i 10 I i TRAGROH F CM d in LU o 7i CO NROH R j c u 21 f O i- Z. O iu j a i it K '** •* ai *— cc UJ CD w in u 3 > i ft/ >— UJ O U. z o — IU S Q i? UJ ^ j f °C£ —O . —Z UoJ »- r> ( j 10 NJ i UJ UJ t / a: cc < ii. X X ( CD e: ^ u a o ro z 10 o "\ U_j »— CD — a | 1 a> m CM LJ U t- • Q.-2 i — — a: : z _J a. - o: -or a ^ 3 3 > o CO C\J CO 1— Z CO z: Lu ic LLJ C_> h— Z CO UJ Z — o a: => _> •«? CNJ ~ CD O -fl- CO CD r~ r- t— X It » « B o < + X 002 I 000 I 009 009 OOfr 002 f^l \JPlll-JVI13_JII13l L J J QIIXUCJJJCJWJJ.L S econdary duct test component Radial n?mperature distributions Measuring plane B' F c1-5 1 _>u c_) cD { ^ 120- •x. // | 110- -U // inn - Ai IUU r nn i/xi 80- A •7 fi _ V / u / (. A _ bvj 400 600 800 1000 ^e(ium Temperature °C 0 Measuring Plane B • Measuring Plane B' Mean values of the oufs de pressure tu be temperature as a funktion of helium Fig.6 temperature and maximum deviations 12 % 400 °C • 600°C * 802°C * 890°C + 936°C Azimuthal ternperatur distributies on the pressure tube outside for different helium Fig. 7 temperatures 1200 4600 1000 3832 i/i 800 3066 a CD 600 2300 400 1533 200 760 Convection 400 500 600 700 800 900 Helium Temperature °C o Measuring Plane B • Measuring Plane B' Specific heat loss in W/m2 and W/m duct length as a function of the Fig. 8 helium temperature U59> 16-S0/4 No. 26 INTERNATIONAL ATOMIC ENERGY AGENCY XA0055836 Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors Diisseldorf Federal Republic of Germany 16.-19. April 1984 Graphite and Carbon/Carbon Components for Hot Gas Ducts by G. Popp, U. Gruber, H. Boder, K. Janssen SIGRI ELEKTROGRAPHIT GMBH GRAPHITE AND CARBON/CARBON COMPONENTS FOR HOT GAS DUCTS G. Popp, U. Gruber, H. Boder, K. Janssen Introduction The large coal reserves in the Federal Republic of Germany and the uncertainty of the future energy situation on the world market make it appear sound policy to devote some thought to the gasification of coal. For certain chemical processes, moreover, it would be advantageous to have a reasonably priced source of process heat available. In the Federal Republic of Germany this process heat shall be produced in a high-temperature nuclear reactor (HTR), the primary heating temperatures being in the range between 950 °C and 1050 °C. One serious problem in utilisation of high temperature heat is the temperature resistance of the construction materials. Even special alloys are scarcely suited for a lifetime of 40 years. Ceramic materials with high tempera- ture resistance therefore come into consideration. The ma- terial include graphite and also CC carbon fibre reinforced carbon. As a result of the projects promoted by MWMV Diisseldorf it has now been demonstrated that both CC and graphite ma- nufactured from SIGRI GmbH are well suited for use in high temperature reactor. *) Ministerium fur Wirtschaft, Mittelstand und Verkehr ministry for economy, middle class and traffic - 2 - 2. Properties required by CC and graphite In HTRs The properties required by graphite and CC as materials for the primary hot gas line will now be considered. Table 1 sets out the specification nowadays required of a material for the primary hot gas line in a hightempera- ture reactor. The normal operating temperature will range from 950 °C to 1050 °C but may rise as high as 1150 "C in the event of faults. Not even these temperature peaks, constitutes a*ny problem for CC and graphite. Both these materials can be used at up to 2500 °C. Another feature deserving particular mention is the high thermal flux of 4000 W/m2: here, a material like CC, with thermal conduc- tivity substantially lower than that of super alloys, has advantages. A further factor to note is the high gas pressure of 40 bar although this is taken up by a steel tube surrounding the inner liner. Approximately, the only forces acting on the CC/graphite inner tube are the restoring forces of the fibre insulating mats and any pulsation of the helium. The primary gas line does not need to be gas-impermeable. The integral radiation dose amounts to about 10 neutrons/cm2. In order to prevent corrosion of the secondary hot gas line, being made of stainless steel, the graphite and CC-material must be virtually free of chloride and sulphur. This high degree of purity can be provided by the choice of materials and appropriate process control during manufacture. - 3 - 3. Properties of CC and graphite Table 2 compares the properties of CC and graphite in a summary. CC and graphite differ substantially in anisotropy, mechanical properties a summary and fracture mechanics. Further measurements are neccessary concerning the corrosion resistance and erosion rate of cc. Should these values pro- ve adequate, CC will be superior to graphite. 4. Manufacture, Properties 4.1. Manufacture of Graphite The raw materials used for the manufacture of graphites are petroleum and pitch cokes and pitch or, less commonly, synthetic resins as binder. The components are mixed and the mixture is shaped into "green" artifacts. In a next step the green artifact is baked at temperatures of 800 to 1000 °C (pyrolysis of the binder). The porous baked artifact may be impregnated with pitch or resins and subsequently heated for carbonisation of the impregnation. The final stage of the manufacturing process is graphiti- zation at temperatures above 2500 °C. If 33 - 4 - 4.2. Average graphite properties At present there are two grade of reactor graphite with a good level of properties: a) Graphite ASR-1RG (SIGRI GmbH) b) Graphite V 483 H (RINGSDORFF-Werke GmbH) The most important properties are summarized in Table 3. The higher strength of the grade V 483 H is due to the smaller grain size of this grade (0,1 mm) and the isos^atic compaction procedure (see also Fig. 1). The ASR-1RG graphite is compacted by vibration moulding (see Fig. 2) and con- tains larger coke grains ( -\, 1 mm) . 4.3. Dimensions of components with could be made from reactor graphite Table 4 provides a survey of the component sizes which can be manufactured at present. Components of considerably lar- ger sizes can be manufactured from ASR-1RG graphit rather than V 4 83 H grade. The surface roughness of ASR-1RG graphite, however, is not quite as small as that of grade V 483 H, whose excellent strength is particularly noteworthy. 4.4. CC manufacture and machining CC-artifacts are manufactured from carbon fibres and a resin which is transfered into carbon by thermal pyrolysis, The actual loading of an artifact can be compensated by the fibre alignment. CC can be machined by any of the usual cutting machining operations, such as turning, milling and drilling. If at all possible, machining should be carried out under wet conditions. 4.5. Level of the CC properties The mechanical properties of two directional CC composite materials are given in Table 5. One of these materials is a (0/90°) composite and another is (0AM5/900) composite which is quasi-isotropic. This is illustrated in the polare chart (see Fig. 3), for flexural strength and Young's modulus in flexure. Subjected to a corresponding high temperature treatment, the strength values will decline slightly and the Young's moduli will increase substantially. The 0°-direction al- ways refers to the warp direction of the composite material, and the 90°-direction to the direction of the weft. The advantages of CC are its comparatively high flexural, tensile strengths, and Young's modulus. The weak point of CC is the interlaminar shear strength, which is about 10 N/mm2. Fortunately the effect can be compensated for by a suitable fibre arrangement within the artifact. - 6 - The low density of the material, is attractive in compari- son with super alloys, which are approximately six times as heavy. One feature of its mechanical behaviour in the event of damage is controlled fracture, governed by the nature of the applied load (bending/shear stress) or nearly brittle fracture in the case of pure tensile stress (see Fig. 4). The curves apply to room and high temperature. An idea of the extreme energy absorption of a CC pipe section under pressure is proveded by Fig. 5. The pipe can be pressed inwards by approximately 5 cm without any disastrous crack propagation. The sequence of damage pro- ceeds fiber fractures to delamination and debonding, fol- lowed by controlled failure. The high creep resistance of the material is shown in Fig. 6. Sample B enters a stationary state without passing a pronounced transition range, this state being characte- rized by an extremely low creep rate which is only just capable of measurement ( ^ 1 . 10 /h). In contrast sample A has a more easily measureable creep rate, but even this is extremely small. Fig. 7 depicts the thermal conductivity as a function of the temperature. This is valid for both types of CC composite. Perpendicular to the plane of the laminates the thermal conductivity appears to be largely unaffected by a rise in temperature, whereas int he laminate plane a slight increase is observed from room temperature to 1100 °C. So far as the function of CC as a material for the hot gas line is concerned (insulating effect), this low thermal conductivity is a desired property. 5. Components made from CC 5.1. Pipes and similar components The requirements for pipes and the current feasabiXity of manufacturing such component are set out for ready refe- rence in Table 6. The pipe primarily requires straight lengths up to 700 to 900 mm. These are already technically feasible. Fig. 8 illustrates, for instance, a pipe with external ribbed reinforcement, a diameter of 1200 mm, a height of approximately 1500 mm and wall thickness of 12 mm. Fig. 9 shows a blow-moulding tool for the superplastic for- ming of titanium at about 950 "C. This CC tool is approxi- mately 700 x 1400 mm in size. The design is most complica- ted due to flanges, retaining loops, bolts with nuts around the circumference and also the conical geometry. Segmented securing elements for a large-diameter pipemade for reactor applications are illustrated in Fig. 10. The boreholes for securing the segments are clearly visible. The segments are about 80 mm in height. 5.2. Sheet-like CC components Among the first components to be manufactured by SIGRI GmbH were those with sheet-like geometry. One particularly interesting component is the CC swage block head for a plant engineering vessel measuring 700 mm diameter x 50 mm is shown in picture 11. The boreholes and turned-out hollows were produced afterwards. 5.3. Connecting elements The use of CC in the high-temperature range as for primary gas lines, securing structures or heat exchangers demands thermally resistant connecting elements. Projects promoted by MWMV-Dusseldorf, however, now enable SIGRI GmbH to de- monstrate the notable level it has achieved in the produc- tion of bolts, nuts, loops etc., which are suitable for HT use and nearly as versatile as those from metal. A few typical connecting components made from CC are shown. Fig. 12 and 13 illustrates bolts and nuts of various sizes with metric threads, which are made from different CC-types developed with SIGRI GmbH. Prospects It has been shown in this paper that CC composite fibre mate- rial and also graphite are genuine alternatives to metals in hot gas lines. CC and graphite know-how has now reached a level which is arousing great expectations that still more applications will be develop, eg in the field of high-temperature heat exchangers, In particular, both materials are suited for use at far higher temperatures which are totally excluded to metals. We are convinced that both materials are still at the beginning of their exploitation in the high-temperature sector. 93? - 9 - Especially in the case of CC, however, knowledge of component dimensioning needs to be augmented so that the fibre composite material can be used under conditions appropriate to its pro- perties and hence also economically. Service Conditions: hot gas temperature 950 - 1050 °C ^ disruption case max. 1150 °C 1ife time 40 years thermal flow "^ 4000 W/m2 temperature interval 1 2 k/min disruption case max. 200 K/min THESE CONDITIONS HAVE TO hot gas pressure 40 bar ? BE FULFILLED FROM CC- disruption case ^ 0 bar AND GRAPHITE COMPONENTS max pressure-inner tube 1 bar hot gas velocity 60 m/s irradiation dosis «^1o''" n/cm2 hot gas medium HTR-He chlorides ^ 10 ppm sulfure ^ 15 ppm J CO cr SERVICE CONDITIONS OF THE INNER PART OF THE CD (SIGRI) PRIMARY HOT GAS TUBE OF HTR CC-CHARACTERISTICS Graph i te-CHARACTERISTICS excellent high temperature resistance,useable excellent high temperature resistance, useable to2500°C to 2500 °C anisotropy of mechanical/physical properties low anisotropy of mechanical/physical low inter laminar shear strength, to consider with constr. high bending and outstanding tensile strength medium bending strength, low tensile strength small fatigue, creep nearly zero small fatigue and very small creep controlled load-strain failure no controlled load-strain failure excellent thermoschock resistance excellent thermoschock resistance high self-damping good self-damping very low density low density oxidation sensible oxidation sensible low thermal expansion, low heat conductivity low thermal expansion, low heat conductivity properties of erosion? properties of erosion? irradion properties? good irradion properties good heat isolation average heat isolation Cl, S-content attainable ^ 10 ppm Cl, S-content attainable / 10 ppm CO (SIGRO TYPICAL PROPERTIES OF CC AND GRARilTE IN RELATION TO SERVICE CONDITIONS OF n> THE HTR table 3 ASR-1RG V 483 H Variant N 1) Variant S2) 3 density g/cm 1.79 1.75 1.79 (1.81)3) dyn, Young's 8.7 8.9 9.6 103 modulus 1 7,7 8.6 9.2 bending 19.4 26.5 (30.5)3j 29.3 (39.7)3) strength 1 17.1 27.9 (30.8)3) 30.2 (39.3)3) tensile 12.0 19.2 21.4 strength 1 10.9 16.6 19.4 compression 47.0 67.7 77.8 strength 1 47.2 70.5 81.3 linear thermal ex- 3.5 3.3 3.3 pansion coeff. 10~5/K 5.9 3.5 3.5 (20-500 °C) thermal 157 116 117 conductivity W/mK 1 136 116 114 ash ppm 130 29 32 permeability cm2/sec 3.2 0.33 0.13 1) Variant N: no post densification 2) Variant S: one post densification with pitch impregnation 3) first, new properties of a new, not fully tested variant of V 483 H Characteristical properties of the CSIGRl) Graphite types ASR-1RG and V 483 H Fig. 1 1 = Frame 2 = Pressure Vessel 3 = High Pressure Liquid 4 = Flexible Form 5 = Form Body (S1GRI) ISOSTATIC PRESSING 93234 993 Fig. 2 \ \ c 5 2 D> 1 = Vibrator Table 2 = Form 3 = Load 4 = Load Fixing 5 = Vacuum Bonnet (SlGRl) VIBRATOR 93334 ASR-1-R6 V-483-H Poss i b 1 e Sizes cylindrical blocks cylindrica 1 blocks max. size 1250 0 mm x 1000 mm max. size 630 0 x 1600 mm prismatic blocks prismatic blocks 500 x WO x 1600 mm max. size 600 x 700 x 2700 mm tubes 1250 0 x 1000 mm advantages : low cost, easy advantages : high strength in production very fine grained large blocks possible high tooling tolerance disadvantages: lower strength than V 483 H disadvantages: sophisticated coarse grains, molding procedure limited tooling tolerance smaller blocks than with ASR-1RG CO cr CD CSIGRI) SIZES/DIMENSIONS AND EVALUATION OF REACTOR GRAPHITE table 5 CC 1501 0/90° CC 1501 0M5/900 Bending strength (N/mm2. 23 °C) 237 156 (10~5 mbar. 1200 °C) 296 Youngs modulus (bending) 58 43 (kN/mm2, 23 °C) (10~5 mbar. 1200 °C) 54 Tensile strength 332 107 (N/mm2, 23 °C) Youngs modulus (tensile) 66 43 (kN/mm2. 23 °C) Compression strength 160 91 (N/mm2. 23 °C) Interlaminar shear strength 12 10 (N/mm2. 23 °C) Poisson number 0.03 0.21 Density (g/cm3) 1.45 1.36 CSIGR!) MECHANICAL PROPERTIES OF CCJ501 0/90° and 0/+45/90o 0° - DIRECTION Fig. 3 (fjNImm) bb E (KNImm) 0 0 CSIGRQ Polar diagram within laminat-plane 92434 Bending strength and Young's modulus o o tension (in 45°) tension (in 0°) ELONGATION Principal Load-Elongation Curves in Tension and Bending Mode IQ CSIGR1) 9313/; CC 1501r [0/90°J , 1200 °C, 10~5 mbar Fig. 5 CC-pipe section under pressure in a universal testing machine 2.5 2,0 Mat. A (low temp, mat.) oo Mat.B (high temp, mat.) to CD 0.5 TIME (h) (Yo.) 10 15 20 25 30 (Q Creep Experiments - 3-Point-Bending CC 1501- O/9OC (SIGRI) 92834 cn 1200 °c, 10"5 mbar, i = 100 MPa = k -15 I -10 900 [°C] 1100 IQ Thermal Conductivity as a Function of Temperature CSIGRO CC 1501-[0/90°J und 9293/i CC-Component Requirements Actual CC-Component Possibilities straight CC-components, curved, CC-prototype components produced with large dia- conical tube components with great meter including flanges dimensions and max flange 2500 mm 0) Plane, flat and curved planes, plates CC-prototyp parts with and without bracing ribs (max 500 x 500 mm) retainers, attachments and joints like screws, nuts and stretching loops produced of screws, nuts, tensile and compression of high loading capacity bushings high accuracy to size without tooling - accuracy to size of very great components in the range of millimeters good tooling good tooling and machinability; high potential for integral constructions ° light weight constructions about 6 times lighter than super alloys cr CSIGRO FEASIBILITY OF CC-COMPONENTS FOR THE INNER PART OF THE PRIMARY CD HOT CAS SYSTEM OF THE HTR en Fig. 8 cc-pipe with external ribbed reinforcement, diameter 1200 mm, height of 1500 mm Fig. 9 cc-blow-moulding tool for superplastic forming of titanium i—'• cc-segmented securing element for a special reactor application with boreholes Fig. 11 cc-swage block head for a plant engineering vessel, 700 mm diameter Fig. 12 cc-bolts and screws in various sizes Fig. 13 cc-bolts in metric threads No. 27 XA0055837 Research on thermal insulation for hot gas ducts P. Brockerhoff Kernfors chungsanlage Jiilich GmbH Institut fiir Reaktorbauelemente Abstract The inner surfaces of prestressed reactor vessels and hot gas ducts of Gas Cooled High Temperature Reactors need internal thermal insulation to protect the pressure bearing walls from high temperatures. The design para- meters of the insulation depend on the reactor type. In a PNP-plant tempe- rature and pressure of the cooling medium helium are proposed to be 950 C and 40 bars, respectively. The experimental work was started at KFA in 1971 for the HHT-project using three test facilities. At first metallic foil insulation and stuffed fibre insulating systems, the hot gas ducting shrouds of which were made of metal, have been tested. Because of the elevated helium temperature in case of PNP and the resulting lower strength of the metallic parts the interest was directed to rigid ceramic materials for the spacers and the inner shrouds. This led to modified structures designed by the INTERATOM company. Tests were performed at KFA. The main object of the investigations was to study the influence of temper- ature, pressure and axial pressure gradients on the thermal efficiency of the structures. Moreover, the temperatures within the insulation, at the pressure tube, and at the elements which bear the inner shrouds were measu- red. Thermal fluxes and effective thermal conductivities in axial and cir- cumferential direction of the pressure tube are given, mainly for the INTER- ATOM-design with sperical spacers. 1. Introduction In a PNP-plant (Prototype Nuclear Process heat project) the heat generated in the core is transferred either to the steam reformer for producing hydro- gen and methanol (HVK) or to the intermediate heat exchanger and steam gas- ifier for producing methanol (WKV)' through primary and secondary hot gas ducts, respectively. Because of the helium temperature of 950 °C and the pressure of 40 bars these ducts need an inner thermal protection system. Since the insulating systems are filled with pressurized helium and are in contact with the hot coolant, they have to meet severe requirements. Mate- rials used must endure all operating conditions over their lifetime. The components must be able to withstand high pressure transients, when sudden depressurization will take place. This causes a high pressure difference between the insulation and the inner cross section. A high pressure within the insulation and axial pressure gradients may produce natural or even forced convection. Convection, howe- ver, must be kept at a low level otherwise the effectiveness of the insu- lation will be decreased. In the following at first the test facilities and the insulating systems tested will be described. After that investigations on a design of INTERATOM and experimental results are discussed. Finally, some results of fibre in- sulations tested previously will be presented for comparison. 2. Test facilities and insulations tested Experimental work was started at KFA in 1971 using the ARGAS-loop described by Bruners et al. /I/ and the high pressure wind tunnel (HD-channel), see Grosse and Scholz /2/. An advantage of the HD-channel is the high volume flow of about 7 m /s thus being four times higher than the one of the ARGAS- loop. The maximum temperature of the HD-channel is 400 °C compared to 1000 °C of the ARGAS-test facility. The corresponding pressures are 40 bars and 10 bars, respectively. In the high temperature helium test rig (HHV) which was erected for testing HHT (direct cycle High Temperature Reactor) components, the maximum temperature is 850 °C at a pressure of 51 bars. The mass flow is approximately 200 kg/s. Further details are given by Noack and Weiskopf /3/. The test objects were mounted horizontally in the test facilities described above. Experiments in vertical position of the insulations were also car- ried out with stagnant gas. By means of electrical heaters hot face temper- atures of 760 C were reached. The insulating systems which were tested up to now are listed in table 1 1) Metallic foil insulations TABLE 1 a) Bobbin design (ARGAS-loop) b) Element design (HD-channel and vert.) Insulating systems c) Element design (HHV-loop) 2) Stuffed fibre insulations a) One interm. tube (ARGAS-loop, p = 280 kg/m3) b) One interm. tube (HD-channel, p=400 kg/m3) c) Two interm. tubes (HD-channel, p=280kg/m3) d) One interm. tube (HHV-loop, p = 290kg/m3) e) One interm. tube (bend) (HD-channel, p = 217 kg/m3) 3) Ceramic insulation a) Carbon rings (HD-channel and vert.) U) Fibre blanket insulations a) Cover plate design (HD-channel, p=178kg/m3) b) KWU/IA design (CFC spacers) (HD-channel, p=130 kg/m3) c) GHT/IA design (Ceramic spacers) (HD-channel, | One can roughly discern between metallic, fibrous and rigid ceramic struc- tures, /4/. The table contains also the test facilities used and the den- sities of the fibre insulations. It is also mentioned, when experiments were carried out with the specimen in vertical arrangement. The metallic insulation of bobbin type was delivered by Darchem/UK for experiments in the ARGAS-loop. A special foil element insulation was also designed by Darchem for HHT applications for high axial pressure gradients and depressurization rates. The foil elements were held together by means of studs and cover plates. End sections bear the inner gas ducting shrouds. First measurements showed an excessive influence of gas pressure on the distribution of thermal fluxes and temperatures around the pressure tube. This was caused by natural convection within circumferential gaps. By clos- ing these gaps the thermal efficiency was improved. As a consequence a similar foil insulation was manufactured for a section of the HHV tubing system. Stuffed fibrous insulating systems designed by BBC/Switzerland for straight tubes were tested as well in the ARGAS-loop as in the HD-channel. The main components are metallic ducting shrouds, v-shaped end pieces and perforated yj-f and intermediate tubes. The annuli were filled with Kaowool the densities 3 3 of which are between 280 kg/m and 400 kg/m . The insulation for the HHV- loop was of the same design. The average fibre density of the test section 3 3 3 3 was 290 kg/m . Densities of only 230 kg/m , 214 kg/m and 208 kg/m were reached for the segments of an insulated bend, which was also delivered by BBC. The design is similar to the insulation for straight tubes. To study the effects of gaps and fabrication tolerances on the thermal per- formance an insulation made of five carbon rings was tested in the HD- channel. It was supplied by the Sigri company in Meitingen. First experi- ments showed that the existing gaps were too large. As a measure of impro- vement the rings were sealed to suppress bypass-flow due to axial pressure gradient. A fibre blanket insulation with metallic cover plates was manufactured by BBC. The density of 178 kg/m was achieved by compressing the Kaowool blankets by means of the cover plates and studs. The additional blanket insulations were constructed by INTERATOM. Contrary to the BBC design they did not contain metallic parts. The inner shrouds made of graphite were held by massive carbon fibre composite (CFC) and ceramic spacers, respec- tively. The blankets were wound around the inner graphite tubes. 3. Description of insulation with spherical spacers Fig. 1 shows the scheme of a section of the INTERATOM insulation. The insu- lation consisted of inner shrouds made of graphite on which the fibre blan- kets were wrapped. At the hot side Saffil fibres of approximately 95% Al-O^, at the outer side Cerablanket fibres of 55% Al^ and 45% SiO2 were used. The outer and inner diameters of the inner shrouds were 780 mm and 630 mm, respectively. The Saffil blankets were compressed by means of wire mesh and \\ water cooling system \Cerablanket FIG. 1. \ pressure tube Saffil inner shroud Scheme of INTERATOM insulation /spacer elements the blankets at the outside by means of a sheet of metal. The mean fibre 3 . density was approximately 160 kg/m . Ceramic balls consisting of high den- sity Al 0, or Si^N, were located at one end of the shrouds for supporting and earthquake damping of the tubes. Additionally, Al-0« or Si-N, elements are used for fixing the liner in axial direction. The whole assembly and further details are given in Fig. 2. The insulation FIG. 2 Pressure tube, insulation and instrumention consisted of 5 sections, the first and fifth of which were used as entrance and exit passages. Sections 2, 3 and 4 were test pieces. The gaps between the particular shrouds were for thermal expansion. The length of the pres- sure tube is 4710 mm. The inner diameter of 930 mm was equivalent to that of tubes already used earlier. Half tubes are welded to the outside for heat removal. The eight measuring systems enable to measure thermal fluxes of the sectors top, bottom, right, and left for the middle part of the insula- tion. In the flange region there are auxiliary systems for cooling the non insulated section of the pressure tube. Thermocouples were installed into the cooling systems and along the outer wall. The positions of the thermo- couples within the insulation, at spacers and shrouds can be seen from the six cross sections. In particular, the cross sections A-A and D-D in the middle of the parts 3 and 4 were instrumentated. The experiments were carried out in the HD-channel of KFA with air and helium as cooling media at pressures up to 40 bars. Because of the graphite corrosion the maximum temperature of the air experiments were restricted to 300 C. Using helium the highest temperature was 400 °C. The axial pres- sure gradient was also varied. In total four test runs were conducted. During the first run the maximum velocity was approximately 20 m/s. In order to get higher velocities a displacement body was mounted which enabled a maximum velocity of about 39 m/s. Since the thermal fluxes of these test runs depended on pressure and pressure gradient another run was conducted with a closed metallic inner shroud which should eliminate the influence of axial pressure gradients. The results of this run, however, were not satis- factory, too. Therefore the insulation was dismantled and rebuilt after having coiled additional Sigraflex foils together with the fibre blankets as convection barriers made of graphite. Because of the lack of time only one test run with the displacement body could be carried out. During the last series the outer cooling was switched off. Fig. 3 shows the temperatures of the pressure tube of parts 2, 3 and 4 for FIG. 3 Temperatures of the after modification without water cooling — pressure tube pG [barl 37.8 376 39.4 TG I°C) 402 395 395 wG [m/s] 39.4 375 35.7 279 277 282 ^^^^^^^^Si^^^^^^^^^^^^^^S^li/icM^^^^^^^^^i the top position and three test runs. The data for pressure, temperature, mean velocity and pressure difference of the coolant are given. The lower curves which stand for the experiments with outer water cooling can be compared. The slope of the temperatures in the region of the second and fourth section is caused by the increase of temperature of cooling water. The temperatures of the third section are approximately constant except the downstream region for the experiments before the modification, where the average temperature is 90 °C. The influence of the higher thermal conducti- vity of the spacer elements is demonstrated. The additional foils increase the pressure drop in radial direction, thus decreasing wall temperatures to 65 C. Only in the region of the axial spacer the temperature is 70 °C. When the water cooling is switched off the temperature of the outer wall increases to 120 C and 150 C, respectively. The heat passing the insulation is trans- ferred to the outside by natural convection and thermal radiation. Since the thickness of the fibre blankets was only 75 mm the increase of tempera- ture by a factor of 2 seems to be high. In Fig. 4 the total thermal fluxes in dependence upon pressure are given for helium, part I, which includes the whole section 2 and half of section 3, and for the first test run. The dependence upon pressure is almost linear. qj^l helium TG[°C] FIG. 4. 2500 wG = 20.8 m/s Thermal fluxes versus helium 2000 397- pressure 1500 301 1000 -O 201 500 -O 104 I • PG I bar] 10 20 30 The maximum flux of 2000 W/m was measured at the gas temperature of 397 °C and the pressure of 38 bars. It is lower by 55% than in case of the insula- tion with CFC spacer elements previously tested, /5/. The improvement is caused by the reduction of the thermal conductivity of the spacer material. In Fig. 5 the total thermal fluxes of the experiments before and after the modification are compared. The results are given in dependence upon pressure 3400 3000 - FIG. 5. Thermal fluxes versus helium pressure 2000 1000 40 3 for helium, part I and the experiments with displacement body. In comparison with the results previously discussed it is striking that with growing velocity from 20.8 m/s to 38.8 m/s thermal fluxes increase by 50% at the highest pressure. This is valid for all temperature steps. In the lower pressure region, however, thermal fluxes are nearly identical in spite of different velocities. After having modified the insulation the dependence of the thermal fluxes upon pressure is only weak, as demonstrated by the dashed lines. Thermal fluxes are remerkably reduced compared to the results which stand for the unmodified insulation. At the highest pressure and 2 temperature thermal flux was only 1500 W/m and thus two times lower. This reduction is also valid for the other temperature levels. Due to the reduction of the permeability in radial direction consequently the thermal fluxes were also reduced. The efficiency of the convection barriers with regard to the distribution around the circumference is demonstrated in Fig. 6, where Nusselt numbers Nu before after modification helium FIG. 6. Nu number versus helium pressure 10 20 30 40 for helium versus pressure are given. The Nusselt number Nu is defined as the ratio of the effective thermal conductivity of each sector to the con- ductivity within the insulation, e.g. when gas movement is suppressed. The results hold for part I, the maximum gas temperature and the average velo- city of 37.4 m/s. Before the modification of the structure the differences between top and bottom sectors increase strongly with growing pressure. Whereas the Nusselt number of the bottom only increases weakly from 1.1 to 1.45, in the top position it increases from 1.7 to 4.4. As shown by the dashed line the additional foils cause in particular an obvious uniformity in addition to a reduction of Nusselt number of all sectors. The highest and lowest values at the highest pressure are 1.38 and 1.15, respectively. This improvement may be compared with the investigations on the metallic foil insulation, /6/. With growing gas temperature the differences between top and bottom sector are lowered. Since the influence of axial pressure gradients on the Nusselt number is only weak, the thermal behaviour of the insulation can be described to be good. Similar results were gained for the stuffed fibre insulation. The experi- ments were carried out in the HD-channel, the ARGAS-loop and HHV-test facility /4/. The design is described in chapter 2. It is the basis for the concept developed by INTERATOM for the KVK-test facility. The Nusselt numbers in dependence upon the average temperature are given in Fig. 7 for various fibre densities and pressures. As described by Bruners et al. /I/ the influence of gas pressure is only weak. The dashed curve fits the experimental results rather well. Contrary to the metallic foil insulation HD- channel helium fHHV-loop FIG. 7 = 38.2 bars. "? p=51 bars^^^ 3 = iOO Kg/m3 [p = 290 Kg/m v Nu number versus helium pressure Nu A (fibre insulations) a Bruners et al. p = 6 bars horizontal vertical p=280 Kg/m3 O top A right average a bottom V left 100 200 300 400 500 tested in HD-channel the insulation whose density was 400 kg/m exhibits only Nusselt numbers of 1.95 and 1.8 for top and bottom sectors, also in horizontal position. Because of the small influence of gas pressure the mean value of 2.07 of the HHV-experiments agrees with the other results, though the differences between the four sectors are higher than expected. This may be caused by irregularities of fibre densities, mainly in the region of the v-shaped end pieces. The reason for the higher values compared with those of Fig. 6 is that the effective thermal conductivity of the stuffed insulating systems is related to the total thickness. Basing on the thick- nesses of the graphite tubes and blankets the Nusselt numbers of the fibre blanket insulation would be increased by a factor of 2.2. In case of the stuffed insulation the Nusselt number reaches a minimum of 1.85 at the temperature of 320 C. Then it increases continuously. In the low tempera- ture region the influence of free convection is stronger than that of ther- mal radiation, which is dominating at elevated temperatures. 10 4. Conclusions A thermal insulation for the primary circuit of a nuclear power station was tested at KFA. The insulation has been developed by INTERATOM. At first the design, the instrumentation and the various test runs were described. After that the experimental results were discussed for helium as coolant. In particular the constructive improvements were mentioned. Natural and forced convection were almost suppressed by means of Sigraflex foils. Consequently, the amount of thermal fluxes and effective thermal conductivities of the four sectors was almost equal. Finally, the results of the stuffed fibre insulation previously tested were discussed, the thermal behaviour of which was quite similar. 5. References /I/ Bruners, R., Lang, H., Noack, G., JuL-1227 (1975) 111 Grosse, H., Scholz, F., Kerntechnik 7 (4) (1965) 150-158 131 Noack, G., Weiskopf, H. , JiiL-1403 (1973) I hi Brockerhoff, P., BNES, Vol. 1, London (1982) 145-150 151 Brockerhoff, P., Stausebach, D., JiiL-1840 (1983) 161 Brockerhoff, P., Scholz, F., IAEA-SM-2OO/3O (1976) 353-362 9G6 No. 28 Facility for Endurance Tests of Thermal Insulations R. Mauersberger XA0055838 Hochtemperatur-Reaktorbau GmbH Federal Republik of Germany In the following report the design and construction of an experimental facility for endurance tests of thermal insulations is presented. It's name in abreviation is "ADI" standing for the German words "Anlage zum Dauer- test von _Isolierungen" . This test facility was build by HRB in order to investi- gate the performance of thermal insulation systems of hot gas ducts for the process heat-reactor-project. The tests are intended to simulate the conditions of reactor operation, They include short-time experiments for selection of insu- lation-concepts and in a second step long-time experiments as performance tests. During these tests are measured the effective heat conductivity the local heat losses the temperatur profiles of the insulation, of the fixing elements and along the wall of the duct The design-data required to perform all these tasks are shown in the first picture: H9 - 2 - The gas-atmosphere must be Helium in tests like in reactor with regared to the special thermal and hydraulic properties of Helium and to the influence of Helium on mechanicle friction and wear. The hot gas temperature in the PNP-reactor will be 950° C and should be equal in the experiments. The temperature on the cold side of the insulation has to be adjustable from 50° C up to 300° C. The Helium pressure in the hot gas ducts of a HTR- plant is about 4 2 bar. The ADI was layed out for 70 bar to cover the hole range of interest. A Helium mass flow has to stream through the insu- lated test duct in order to realize equal tempera- tures on the hot side of the insulation. A flow rate of 4,5 kg/s is sufficient for this requirement. The axial pressure gradient along the insulation must be the same as in the reactor, because this has an essential influence on the heat losses. This pressure gradient is about 40 Pa/m. - 3 - An important part of the test programm is the realization of temperature cycles. The temperature transient should be approximately 3° C/min. This value is - however - depending on the quality of the test insulation. By temperature cycling the start-up and shut-down procedures and conditions of a PNP-reactor are simulated. Last not least the test of full scaled insulations must be possible. A longitudinal section of the ADI-test-vessel is shown in fig. 2. It has an inner diameter of 2 m and a total length of 10 m. The test insulation is mounted in a tube, which is fastened to two special designed rings. On the out- side of the tube is welded a cooling-system; it is divided into longitudinal sections and into circum- ferential segments in order to measure local heat losses. The cooling fluid is a special heat-transfer- oil which can be used up to 350° C. An electrical heater which is built in a tube, is inserted into the hot gas duct. Between this tube and the hot side of the test insulation is an annular space for the Helium flow. - 4 - The hot gas Helium flow along the test insulation is circulated by an integrated blower. The blower has an electric motor drive with speed control. The shaft of this drive has magnetic bearings wor- king without any friction. This special kind of bearings allow the endurance tests under Helium and at high temperature without interruption for changing the bearings. The axial wheel of the blower is working at full test temperature up to 950° C. The blower exhausts the heated Helium from the electrical heater and presses it into the annular space along the test insulation mentioned above. A ceramic guide for the Helium flow is placed behind the blower wheel. It protects the housing of the blower drive against high temperature and guides the Helium flow into the wanted direction. A similar ceramic guide is placed on the other side of test duct leading the Helium flow back to the heater. A Helium/water-heat-exchanger is integrated in the other cap of the pressure vessel. It is used for cooling the Helium purge flow to the gas-purification and for the Helium which has to be circulated if the test bench is performing temperature cycles. The empty volume in the pressure vessel is fulfilled with an auxiliary insulation of fibrous material. - 5 - A survey of the various systems for supply and control outside of the pressure vessel is given in fig. 3: The control system for the blower drive and the magnetic bearings. The oil-cooling system for the test duct. The measurement devices for the test-insulation. The water-cooling system for the whole pressure vessel. The energy supply and control for the electrical heater. The gas purification system including the analytical measurement of gas impurities. All measured data from the test insulation and the operational components are collected by a data acquisition system. Then follows data processing and documentation. The ADI can be driven automatically in normal operation and in cycling operation. In both cases all security precautions are taken into account. This self-controlled operation is a great advantage cause it decreases costs. - 6 - In the mean time the experiments with the first test insulation have been finished and some results are presented in the fig. 5 to 8. The set-up construction of the test insulation is very similar to that one, which was explained in the report of Mr. Brockerhoff, just before (fig. 5). Within a pressure tube is a inner liner tube made of graphite. This tube is supported by special spacer rings. Blan- ketts of fibrous insulation material are wrapped around the graphite tube. The density of the fibrous material is about 130 kg/m3. A temperature profile within the insulation is shown in fig. 6. It is remarkable that the temperatures near the duct decrease so steeply. The temperature distribution along the wall of the pressure tube is given in fig. 7. There are remarkable temperature peaks in the neighbourhood of the spacers and the temperature differences between upper and lower side are sometimes considerable. The heat losses as a function of gas temperature are shown in fig. 8. The heat losses of section 2 are higher because this section includes two spacers. At present experiments are running with another test insulation with spherical spacers of Interatom Company. Mr. Brockerhoff has reported before on the experimental results of this insulation in the temperature range up to 400° C. In ADI the temperature range of the measure- ments will be extended up to 950° C. The experiences with ADI gained over a period of more than one year have shown that the specifications and operational requirements have been fulfilled completely and very satisfactorily. The erection of the test facility ADI and the performance of the tests were sponsored by "Minister fur Wirtschaft, Mittelstand und Verkehr des Landes Nordrhein-Westfalen". Medium: Helium / Max. \ ,1 \ / Dimensions \ >' Hot Gas \ I of Test Ducts: Temperature: j V D = 2000 mm 950°C J \L = 4700 mm ^^ ^^^ / \^ I Pressure Tube \ / Temperature - \ transient: K > C )( Temperature: 1 \ ± 3°C/min J ADI } V 50-330°C J [ Pressure \ X / Helium \ / Gradient: \ ^-*-^ / \/"^ \/ Pressure: 1 1 40 Pa/m r N / A 70 bar I \(950°C/40 bar)/f Helium \ Massflow: i 4,5 kg/s Jf^—^^ ADI PNP fig. 1 KRB Design Data 84.27- 2 5 6 6 2 1 Pressure Vessel 4 Heater 2 Heat Exchanger 5 Hot Gas Blower 3 Test Insulation 6 Guide for Helium Flow Endurance Test Facility PNP fig. 2 ADI of Thermal Insulation 84.27- 1 Oil-cooling Water Cooling Heater Analytical Blower System System 200 kW Measurements JL Digital and Analog Data Data Acquisition —^ Data Processing Display and Output ADI PNP fig. 3 Principle Design of Test Facility 84.27- 4 ADI PNP fig. 4 Test Facility in Operation 84.27-16 f~\ /^\ /^\ /^N /^k / ////////////// } )••:'•''"•':'•'. ''•yi'l'l'V, •' Cerablanket ;.'•;'; • r ; r •K -V;': •/.•:•::••: Saffll •::"••.'•:•.••. 4 796 4 0 93 K^i Inner Liner Tube S 1" 1 Spacer Q X/\ Pressure Tube i :;•;:•: Y ''///////////// Length of the Test Duct: 4700 mm ADI PNP fig. 5 MB Test Insulation with CFC-Spacer 84.27- 3 « Cerablanket 800 — Upper Side Lower Side Temperature tTube [°C] tTube (average) Medium : Helium Ap' : 15 Pa/m tGas = 946°C (average) Peas = 40 bar trube = 147°C (average) fig. 7 ADI PNP Temper Distrib. along Pressure Tube 84.27-10 P = 40 bar x = Section 1 2 q [W/m tTube = 150°C (average) a = Section 2 1000 800 600 / y 400 Y 200 200 400 600 800 1000 tGas[°C] ADI PNP fig. 8 Heat Loss of Test Insulation 84.27-12 No. 29 Construction and Performance Tests of Helium Engineering Demonstration Loop (HENDEL ) for VHTR 111 III 111 III Uli III III XA0055839 Makoto HISHIDA, Toshiyuki TANAKA, Hiroaki SHIMOMURA and Knonomo SANOKAWA Department of High Temperature Engineering Japan Atomic Energy Research Institute, Japan 1. Introduction A helium engineering demonstration loop (HENDEL) has been constructed and operated in the Japan Atomic Energy Research Institute (JAERI) in order to develop the high-temperature key components of an experimental very high- temperature gas cooled reactor, like fuel stack, in-core reactor structure, hot gas duct, intermediate heat exchanger and so on " . Performance tests as well as demonstration of integrity are carried out with large-size or actual - size models of key components. The key components to be tested in HENDEL are: (1) fuel stack and control rod (2) core-supporting structure, or bottom structure of reactor core exposed to direct impingement of high-temperature core outlet flow (3) reactor internal components and structure (4) high-temperature components in heat removal system: primary and secondary cooling systems. HENDEL consists of mother section, adapter section and test section " '. The mother and adapter section (M+A section) supplies test sections with He gas flow of required temperature, pressure and flow rate. The first stage of HENDEL project was to construct the M+A section. The design work of this section was started in 1978 and construction was completed in March 1982. Up to now, eight cycles of test operations have been carried out with the M+A section, including the preliminary test operation in March 1982. During the operation, performance tests of hot gas ducts and major heat exchanging components of HENDEL were carriec out. At present HENDEL project is in the second stage, that is, tests with the first test section (T-, ) have been started and the successive test sections are being constructed or planned. This report describes outline of HENDEL test facility and performance test results of the hot gas ducts and the heat exchanging components. - 1 - 2. Outline of HENDEL Figure 1 shows the flow sheet of HENDEL. The M+A section of HENDEL1~5^ consists of No.l mother loop (M, loop), No.2 mother loop (M? loop), adapter section (A section), purification system, make-up system, water cooling system, instrumentation and control system, and electrical system,HENDEL has four test sections of T, - T,. A fuel stack test section, which is the first test sectior (T,), was constructed in March 1983 and tests are now being performed. The design work of the second test section (T?) is now under way and it will be put into operation in March 1986. The third and fourth test section (T3 and T.) are now being planned. The M, loop supplies T, test section with He gas flow of the maximum tempera ture 400°C, the maximum pressure 4.0 MPa and the maximum flow rate 0.4 kg/sec. The temperature of 400°C and pressure of 4.0 MPa are equal to those of inlet He gas flow of VHTR reactor core. And the flow rate of 0.4 kg/sec is equal to the one of two colums fuel stack. The M, loop consists of a gas circulator (B-,), a He gas heater (H,), a He gas cooler (C,), a mixing thank, a filter, orifice flow meters and so on. The M~ loop supplies the test sections T2 - T« with He gas flow of the maximum temperature 400°C, the maximum pressure 4.0 MPa and the maximum flow rate 4.0 kg/sec. The flow rate of 4.0 kg/sec is equal to half the flow rate in one of the two primary cooling loops for VHTR. And the M2+A loop supplies the test sections with He gas flow of the maximum temperature 950 - 1000°C, the maximum pressure 4.0 MPa and the maximum flow rate of 4.0 kg/sec. The maximum temperature 950 - 1000°C is equal to the mixed mean temperature at the outlet of VHTR reactor core. The mother and adapter loops consist of He B He gas neaters H H H a hot gas ducts gas circulator (B-,, B22, B^o, 24^' ( 2' 31 ' 32^ (hot gas ducts A,Bj, He gas coolers (Cp, C.,,, C^), mixing tanks, filters, orifice flow meters and so on. The purification system consists of CuO beds, molecular sieve traps, active carbon traps, compressors, heat exchangers and so on. Impurity gases of H2 and CO become H^O and C02 by oxidation reaction in the CuO bed. Produced Wfl and C0? together with H?0 and C0? contained in inlet He gas flow are absorbed in the molecular sieve bed. Impurity gases of 0o, N^ CH- are absorbed in active carbon trap of -196°C. The purification system is so designed that contents of H20, 02, N2, C02 and CO can be reduced to bellow 0.1 p.p.m. and those of H2 and CH. bellow 0.2 p.p.m. at the outlet of the system. Major heat exchanging components in the M, loop are He gas heater H, and He gas cooler C-,. And those in the M~ loops and the adapter section are He - 2 - gas heaters H~, H.,, and H_?, and He gas coolers C^, C-, and C^. Up to now, the M, loop has been operated 7 cycles and total operation hours are about 3000 hours. And the Mp loop and NL+A loop have been operated 8 cycles, operation hours being about 3200 hours. During the operation, performance tests of the hot gas duct A and B installed in the adapter section and major heat exchanging components in the M+A section were carried out. 3. Thermal performance test of hot gas ducts As shown in Fig.l, two hot gas ducts are installed in HENDEL. Hot gas duct A is connected with the main heater H-? and main cooler C~,, and the hot gas duct B is connected with the main heaters H,, and W~?. Table 2 shows the main items of the two hot gas ducts. Figure 2 shows a conceptional view of hot gas duct A. Hastelloy-X is chosen for the liner tube, because it is supposed to be exposed to He gas flow of 1000°C. The pressure tube is made of mild steel. In order to maintain the temperature of the pressure tube lower, the internal thermal insulation is installed outside the liner tube, and is divided into three sub-layers by stainless steel foils. Fibrous ceramic insulation (Kaowool) whose main composition is SiO,, and AlpO- is packed in these layers. In the two inner sub-layers, Kaowool blanket of 0.2 g/cm in density is packed, and in the outmost 3 sub-layer Kaowool bulk of 0.25 g/cm in density. V-shape end plates of Hastelloy- are welded to the pressure and liner tubes at intervals of 0.7 - 1.4 m so that they might prevent bypass or permeation flow of He gas in the insulation layer, which gives rise to deterioration of insulating characteristics. The basic configuration of hot gas duct B is almost the same as that of hot gas duct A. In the present test, temperature distribution of pressure tube and internal thermal insulation, and heat flux on pressure tube were measured. Additionally, air flow around the hot gas ducts was visualized with mist of dry ice. Temperature distribution of whole surface of pressure tubes was measured by C.A. thermocouples fixed on the pressure tube and a radiation pyrometer. Temperature distribution of internal thermal insulation was measured by C.A. thermocouples with metal sheath. Heat flux was measured with heat flux meters. The present test was performed in the range of He gas temperature of 400 - 950°C, pressure of 1.0 - 4.0 MPa and flow rate of 0.5 - 3.5 kg/sec. Figure 3 shows circumferential surface temperature distributions of the horizontal, vertical and bent tubes of hot gas duct A, measured by fixed thermo- couples. The maximum temperature variation range in the circumferential direction was only 15°C. The surface temperature of pressure tube was maximum at the locations where studs and V-shape end plates were welded to the pressure - 3 - tube, and was minimum at flanges and the locations where supporting hangers were fixed. The maximum and minimum temperatures, however, were ±30°C higher or lower than average temperature. That is, temperature distribution of pressure tube was almost uniform. Radial temperature distribution in the internal thermal insulation was in good agreement with the one calculated by a conduction trudel taking into consideriation of temperature dependence of effective thermal conductivity of insulation material. Figure 4 shows circumferential heat flux distribution of the horizontal, vertical and bent tubes of hot gas duct A. Heat flux distribution of tubes of the hot gas duct B was also measured; the heat flux distribution of hot gas duct A and B was almost uniform. Effective thermal conductivity of internal thermal insulation layer was evaluated from the measured heat flux and temperature. Figure 5 shows the effective thermal conductivity of the present hot gas ducts together with those of cither hot gas ducts " ' ' or fibrous insulation material itself ' '. Effective thermal conductivity of the present horizontal tubes of hot gas ducts A and B is correlated by the following equation: 4 Xeff(W/m.K) = 0.01963 + 4.702 x 10~ T(K) (1) and the one of the vertical tubes is correlated by the following equation: 4 X ff(W/m«K) = 0.02014 + 6.039 x 10" T(K) (2) Effective thermal conductivity of the vertical tubes was 25 - 30% higher than that of the horizontal tubes, which might be caused by natural convection established in internal thermal insulation layers. The presented effective thermal conductivities shown in Fig.5 are 1 - 2 times larger than the thermal conductivity of He gas. The difference between the data ' " ' might be attributed to the difference of composition of fibrous insulation material, packing method of insulation material, design of internal insulation structure and so on. Most data lie within the range of -10 - 45% of the present equation (1). 4. Thermal performance test of He gas coolers In the M+A section of HENDEL, four He gas coolers are installed, which are C,, Cp, C.,, and C-p. Main items of the coolers are listed in Table 3. Basic structure of He gas coolers C-,, C«, C^2 is almost the same. U-tubes are used for heat exchanging tubes which absorb thermal expansion at free end. Cooling water of atmospheric pressure flows inside the tubes and He gas in the shell with segmental baffle plates. Arrangement of tubes is staggered equilateral - 4 - triangle. Material of tubes is mild steel because of rather lower temperature. Detailed structure and drawing are presented in references(l-4). Figure 6 shows a schematic view of He gas cooler C.,,. The structure of He gas cooler C,-,1-4 )' is different from that of other coolers. Straight tubes are connected to the top and bottom water ring headers, and both ends of tubes are bent in order to absorb thermal expansion. Inside of the tubes flow pressurized cooling water for preventing boiling. Cross flow of He gas is produced by step baffle plates. The pressure vessels of cooler C_, and C,? have internal thermal insulation to keep temperature lower, while those of other coolers C-, and C? have no internal thermal insulation because of rather lower He gas temperature. Measured cooling capacities of the coolers C,, C2, C~, and C-- were respectively 20%, 40 - 50%, 50 - 70% and 40 - 50% higher than the design values. Figure 7 shows the relatioship between Nusselt number of He gas flow in the shell and Reynolds number. Nusselt number and Reynolds number are defined as follows: a *d Nu --V2- (3) Re = ^~- (4) g where, a was calculated from the equation: a = where, a. was obtained from the Dittus-Boelter's equation: a- = 0.023 x^x (^i)0-8 x Pr0'4 d. v i w The straight lines in the figure are the correlations for tube bank with segmental baffle plates, which were presented by Donohue ' and Kern '. The data measured with the cooler C, agree fairly well with Donohe's correlation, while the data measured with the coolers C? and C~? are 10 - 20% higher than the correlation. The chain lines are the correlations for tube bank in cross-flow presented - 5 - by Zukauskas ' and Fishenden-Saunders . Measured data with the He gas cooler C_, are 10 - 30% higher than the correlations. After about 3000 hours' operation, no change in overall heat transfer coefficient was detected for all the coolers. 5. Thermal performance test of He gas heaters In the M+A section of HENDEL, four electric heaters are installed, which are H-, , Ho, Hol and H_o. Main items of the heaters are listed in Table 4. Basic structural concept is almost the same for all the heaters , schematic view of the heater W~~ being presented in Fig.8. They have 15-72 heated tubes in which He gas flows upward and is heated up. The tubes are heated by alternating 3 phase electrical current, heating capacity being 160 - 4700 kW. Material of the heated tubes for the heaters H,, H? and H_, is Incolloy-800H, because of relatively lower temperature. On the ohter hand, the heated tubes of heater H_2 are made of graphite, because tube temperature was estimated to be about 1300°C. The heated tubes are fixed to the top or bottom tube sheet at one end and kept free at the opposite end. Three tube sheets are provided in each heater, one for supporting tubes and the rests for preventing vibration. The heated tubes are electrically insluated from the earth by A1203 or Boron-Nitride in- sulators. In order to prevent heat loss from the heaters, pressure vessels of the heaters H. and H? have thermal insulation wrapped on the outer surface. And the pressure vessel of the heaters H31 and H-p have internal thermal insulation. Figure 9 shows thermal efficiency of the heaters, which is the ratio of enthalpy rize of He gas to electric input. Thermal efficiency of all the heaters is higher than 70%. Namely thermal insulation performance of pressure vessels of heaters was satisfactory. Temperatures of heated tubes were also measured and they were in good agreement with the values calculated by Dittus- Boelter's equation (6). It seemed to indeicate that flow distribution among heated tubes was almost uniform and no significant bypass flow, which might deteriorate heat transfer performance, occurred. 6. Conclusion Thermal performance tests were carried out with the hot gas ducts and the major heat exchanging components of the M+A section of HENDEL. The following results were obtained: (1) The major components were operated for more than 3000 hours without trouble. (2) Temperature and heat flux distribution on the pressure tube of hot gas - 6 - ducts were almost uniform in both circumferential and axial directions. No hot spot was found. (3) The correlation for the effective thermal conductivity of the internal thermal insulation was obtained as follows: 4 Xeff(W/m.K) = 0.01963 + 4.702 x 10~ -T(K) (4) The cooling capacity of He gas coolers was 40 - 60% higher than the design values. (5) Heat transfer coefficient of shell side of the cooler C, was in good agreement with Donohne's correlation, while those of the coolers C~ and C_p were 1° - 20% higher than the correlation and that of the cooler C_, was 10 - 30% higher than Zukauskas1 and Fishenden -Saunders' correlations. was (6) Thermal efficiency of He gas heaters H,, H?, H-, and H-p as high as 70 - 95%. Flow distribution among heated tubes seemed to be uniform and no bypass flow likely to occur. NOMENCLATURE d : diameter of tube K overall heat transfer coefficient Nu : Nusselt number Pr Prandtle number Re : Reynolds number T Temperature u : average velocity of fluid u average velocity of fluid in m through tube bank a tube A : effective thermal con- X thermal conductivity eff ductivity v : kinematic viscosity SUBSCRIPT i : inside of tube 0 outside of tube 9 : gas s solid w : water REFERENCES 1) Japan Atomic Energy Research Institute Status of R & D on VHTR in JAERI, (1980 - 1983). 2) Ishikawa, H. et al . : ASME 81-WA/NE-9 (1981). 3) Okamoto, Y. et al. : JAERI-M 82-133 (1982). 4) Izawa, N. et al. : JAERI-M 82-122 (1982). 5) Tone, H. et al., : JAERI-M 8309 (1979). 6) Umeda, T. et al. :• IHI Tech. Bull., 20(6), 416 (1980). 7) Nakano, T. et al. : FAPIG, 94(3) 9 (1980). - 7 - 8) Ogawa, M. et al. : Private communication. 9) Katagiri, M. et al. : Trans. JSCM, 7(1), 11 (1981). TO) Tokita, U. et al. : Mitsubishi Tech. Bull., 27, 14 (1982). 11) Jones, G. et al. : ASME 74-WA/HT-l (1974). 12) Nakanishi, T. et al. : J. At. Energy Soc. Japan, 21(2), 13 (1979). 13) Brtickerhoff, P. : J. Non-equilib. Thermo., 3(4), 231 (1978). 14) Donohue, D. A. : Ind. and Eng. Chem, 41, 2499 (1949). 15) Kern, D. L. : Process Heat Transfer, McGraw-Hill (1950). 16) Zukauskas, A. : Advans in Heat Transfer, Academic Press. (1972). 17) Fishenden, M. et al. : Introduction to Heat Transfer, 132, Oxford (1950) TABLE 1. MAIN ITEMS OF HENDEL M+A SECTION M1 LOOP M2 LOOP M2+A LOOP TEST SECTION T l 2' 3' 4 T2'T4 TEMPERATURE 400°C 400°C 880°C, 1000°C PRESSURE 4.0 MPa 4.0 MPa 4.0 MPa FLOW RATE 0.4 kg/s 4.0 kg/s 2.8 kg/s,4.0 kg/s HEATER / u \ (H2) (H31+H32)or(H2+H31+H32) HEATER POWER HT 160 KW 2000 kW 2000 kW H31 4700 kW 4360 kW H32 (D \ (B +B ) (B +B )or(B - +B ) B1OWER ID- I 21 22 23 24 2 ! 22 HEAD 0.2 MPa 0.1+0.1 MPa 0.1+0.1 MPa REVOLUTION 12000 r.p.m 12000 r.p.m. 12000 r.p.m. POWER 150 kW 250kW+250kW 250kW+250kW PIPING DIAMETER 100 mm* 250mm ITEMS HOT GAS DUCT A HOT GAS DUCT B Ci C2 C31 HE GAS TEMPERATURE (MAX.) 1000°C 700°C TYPE EGHENTAL-BAFFLED EGMENTAL-BAFFLED TEP-UP-BAFFLED EGMENTAL-BAFFLED -TUBE -TUBE TRAIGHT-TUBE -TUBE HE GAS PRESSURE (MAX.) 4.0 MPa 4.0 MPa FLUID He/WATER He/WATER e/PRESSURIZED- He/WATER DESIGN TEMPERATURE OF ATER 350°C 350°C FLOW RATE (kg/s) O.4/2.I 0.4^4.0/28 ,0.4-\>4.0/26 0.4VI.O/4O THE PRESSURE TUBE INLET TEMPERATURE (°C) ^400/32 -^400/32 "M 000/50 550/32 DIMENSION (DESIGN TEMPERATURE) PRESSURE 660.4°-D-x22t OUTLET TEMPERATURE (°C) 313/42 288/52 550/110 382/42 D D LINER TUBE (DESIGN TEMPERATURE) 355.6°- -x 6* 355.6°- -x 6* PRESSURE (MPa) *l. 0/0.4 4.0/0.4 4.0/3.6 4.0/0.4 LINGTH 14 m 23.6 m COOLING CAPACITY (kW) 170 2300 6700 3500 MATERIAL REYNOLDS NUMBER ^26000/24000 ^-21000/44000 M2000/72000 •V47O0O/51OOO PRESSURE TUBE SB42 SB42 OVERALL HEAT TRANSFER 430 ^640 "M32O LINER TUBE HASTELLOY X INCOLLOY 800H COEFFICIENT (W/m*K) V-SHAPE END PLATE HASTELLOY X INCOLLOY 800H TOTAL HEAT TRANSFER 1.22 21.8 16.6 18.7 SUS 304 AREA (m2) SUS 316 NUMBER OF BAFFLE PLATES 6 3 3 4 STUD HASTELLOY X INCOLLOY 800H DISTANCE BETWEEN 150 530 900 450 BAFFLE PLATES (mm) INCOLLOY 800H THERMAL INSULATION OUTER INSULATION OUTER INSULATION INTERNAL INTERNAL INSULATION KAOWOOL 1260S KAOWOOL 1260S INSULATION INSULATION A12O3 47.3% A12O3 47.3% TUBE SiO2 52.3% SiO2 52.3% NUMBER 14 132 54 64 Fe2O 0.05% Fe2O3 0.05% ARRANGEMENT STAGGERED (EQUI- STAGGERED (EQUI- IN-LINE STAGGERED (EQUI- .ATERAL TRIANGLE) LATERAL TRIANGLE) LATERAL TRIANGLE) PITCH (mm) 40 40 100 70 OUTER DIAMETER (mm) 27.2 25.4 27.2 48.6 INNER DIAMETER (mm) 20.8 18.4 22.0 41.6 THICKNESS (mm) 3.2 3.5 2.6 3-5 LENGTH (mm) 2040 2070 3600 1910 MATERIAL STB35S STBA22 STBA22S STB35S VESSEL DIMENSION (mm) 230<{ix2300hx20t 922 H.i H H2 32 TYPE ELECTRIC HEATER FLOW RATE (kg/s) 0.04ML4 0.4M.0 O.4M.0 O.4"M.O REYNOLDS NUMBER 54000 68000 83000 83000 INLET TEMPERATURE (°C) 320 300 400 700 (DESIGN TEMPERATURE; OUTLET TEMPERATURE (°C) 400 400 720 1000 (DESIGN TEMPERATURE) PRESSURE (MPa) 4.0 4.0 4.0 4.0 HEATER TUBES MATERIAL INCOLOY 800H INCOLOY 800H INCOLOY 800H GRAPHITE NUMBER 42 36 t DIMENSION 27.2°-D-x3.5t 38J°-D-x2.0t 54.0°-D-x6.5t 7O.0°-D-x15 X28391 X30001 X50401 x3850t VESSEL MATERIAL SCMV2 SCMV2 SB49 SB46 DIMENSION 1280°-D-x40t leoo0-0^* 237O°-D>x6Ot 1684°-D-x42t X56001 X77OO1 xllllO1 X90741 CAPACITY (kW) 10"M 60 50^2000 3OOM7OO 250M360 THERMAL INSULATION OUTER OUTER INTERNAL INTERNAL INSULATION INSULATION INSULATION INSULATION VOLTAGE 3* 6.3KV 50Hz 3* 6.3KV 50Hz 3* 6.3KV 50Hz 3 HEATERIH,) T,-sl FUEL STACK TEST T,-J SECTION T : IN-CORE STRUCTURE T,.. 2 COOLERICI b I-M TEST SECTION T3 : IN-CORE FLOW HOT GAS DUCT(B) TEST SECTION L, = HEAT REMOVAL HENDEL M^T, HEATER{H3() TEST SECTION BLOWER * COOLERS COOLER (C32) COOLER! Cj, WATER PUMP HOT GAS DUCT (A) FIG.l FLOW SHEET OF HENDEL - 10 - 300 STUD 200 PRESSURE TUBE SEPARATING PLATE INSULATION SIGN. LEGEND • • HORIZONTAL TUBE LINER TUBE 100- o o VERTICAL TU8E a s BENT TUBE He GAS TEMPERATURE = 900 °C He GAS PRESSURE = 4.0 MPa AIR TEMPERATURE = 30 °C 0 0 [80 360 CIRCUMFERENTIAL ANGLE (DEGREE) FIG.2 CONCEPTIONAL VIEW OF FIG.4 CIRCUMFERENTIAL HEAT FLUX HOT GAS DUCT A DISTRIBUTION ON PRESSURE TUBE SURFACE (HOT GAS DUCT A) NQia6«.IHEEPRE53MPd *PENS.(g/cro3 150 LU CC SIGN. LEGEND a: • • HORIZONTAL TUBE oL_U o o VERTICAL TUBE ©—-—« BENT TUBE LU He GAS TEMPERATURE = 900 °C He GAS(4.0MPa) 100 - He GAS PRESSURE =4.0MPa AIR TEMPERATURE =30°C • *H:HORIZONTAL TUBE *M:INSUUTI0N MATERIAL: *VIVERTICAL TUBE **: PRESENT WORK 0 0 180 360 0 200 400 600 800 CIRCUMFERENTIAL ANGLE (DEGREE) AVERAGE TEMPERATURE OF INSULATION (°C) FIG.3 CIRCUMFERENTIAL TEMPERATURE FIG.5 SURVEY OF EFFECTIVE THERMAL DISTRIBUTION AT PRESSURE CONDUCTIVITY OF HOT GAS TUBE SURFACE (HOT GAS DUCT A) DUCTS WITH INTERNAL THERMAL INSULATION AND FIBER INSULATION MATERIALS - 11 - HEAT EXCHANGING TU8E OUTER SHELL INNER SHELL STEP BAFFLE PLATE STEP BAFFLE [I6OO INTERNAL PLATE INSULATION OUTLET (WATER) NT 3O He OUTLET =g (1000°C) UPPER TUBE B=Q Rsn HEATER ELEMENT SHEET A (GRAPHITE TUBE) INNER RING HEADER WATER JIE ROD SHELL JACKET (WATER) MIDDLE TUBE SHEET OUTER INTERNAL SHELL STEP BAFFLE PRESSURE VESSEL PLATE INSULATION HEAT t EXCHANGING LOWER TUBE SHEET 8 TUBE INTERNAL POWER SUPPLY He INLET INSULATION TERMINAL (700 °C) RING HEADER (WATER) PRESSURE VESSEL INLET WATER JACKET FIG.8 SCHEMATIC DRAWING OF (WATER) He GAS HEATER W^ FIG.6 SCHEMATIC DRAWING OF He GAS COOLER (C31) i 1 • HEATER Hi ) -I-] I I I I—I ( O HEATER (H2 ) • COOLER (Ci ) A HEATER HJI) o COOLER (Cz) ( H32) * COOLER (C ,) a HEATER 1 3 — a COOLER { C32> 100 - Nu"0.33ReMPr0-3;(FlSHENDEN-SAUNDER5) cc 2 ~ O C LU O CQ • o o A 2 UJ i !0 80 A. 8 b 6 ••* EFFIC I fiu=0.23-R8fl'Pr^(-fc) (DONOHUE) 60 - - (KERN) 10 i L I , ,T 6 8 10 2 4 6 810 2 4 6 810 6 8(0 2 4 6 8|0 2 4 6 8|0 REYNOLDS NUMBER ELECTRICAL INPUT (KW) FIG.7 HEAT TRANSFER CHARACTERISTICS OF FIG.9 EFFICIENCY OF HEATERS TUBE BUNDLES OF He GAS COOLERS - 12 - No. 30 ••III XA0055840 Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors Diisseldorf, 16-19 April 1984 Testing of High Temperature Components in the Component Testing Facility (KVK) W. Jansing INTERATOM GMBH, FRG 1. Introduction The Component Testing Facility (KVK) is used for the experimental testing of high temperature components for nuclear coal gasification. It went into operation in August 82 ofter a planning and construction period of two and a half years. 2. Design and mode of operation The main operating data of the KVK are shown in Fig. 1 - The thermal power is 10 MW (maximum 12,8 MW) - The temperature in the primary system amounts to 950° C (maximum 1000° C) - The system pressure is 40 bar (maximum 46 bar) - The nominal flow is 3 kg/s (maximum 4,3 kg/s) - The helium velocity in the hot-gas duct is 60 m/s - The maximum achievable temperature transients are ± 200 K/min - The maximum achievable pressure transient is 5 bar/s. iffO -2- Fig. 2 presents single-loop operation of the KVK. The initial construction stage includes experiments on hot-gas ducts and hot-gas valves and on the hot header of the helium heat exchangers. The helium is circulated by a radial blower. The necessary- heat is introduced into the circuit via a heating system fired by natural gas and electricity and is discharged via a steam generator, whereby part of the steam is used to preheat the helium in a helium preheater. This regenerative circuit results in a 4 0 % energy saving. The construction of the test facility alone led to important technological progress. Thus seamless tubes made of Nicrofer 5520, which were fabricated in the Federal Republic of Germany for the first time, were used in the natural-gas-fired helium heater with very positive results. There were no significant diffi- culties either in the bending or the welding of the tubes. The 145 m long operational hot-gas duct with metallic liner and fibre insulation was manufactured without any problems. The operational hot-gas duct is essentially in conformity with the planned secondary hot-gas duct of the PNP. The facility is converted with low expenditure for the testing of the helium heat exchangers. Fig. 3 shows double- loop operation of the KVK. The primary system contains the heat source - natural-gas-fired and/or electrically powered helium heater - the secondary system contains the steam generator as heat sink. The helium heat exchanger to be tested transfers the heat from the primary to the secondary system. In addition, a test section of the primary hot-gas duct and the hot header of the helium heat exchangers are included in the primary system. Test sections for the secondary hot-gas duct and the hot-gas valves are installed in the secondary system. -3- Fig. 4 presents the working and demonstration model of the facility. The building is 50 m long, 20 m wide and 25 m high. Some components such as the helium storage tanks, the coolers and the natural-gas-fired helium heater are installed outside the building. The left-hand side of the building accomodates the control room, parts of the water steam system, the helium auxiliary facilities and the blowers. The other loop components and test objects, for example the 10 MW helium heat exchanger, are located on the right-hand side of the building. Fig. 5 is a photograph of the outside of the KVK. In the foreground you can see the helium storage tanks with the control room and on the left in the background the natural- gas-fired helium heater. 3. Test projects Let me now briefly describe the test objects. Fig. 6 shows the 10 MW He/He heat exchanger with helical tube design manufactured by the firms Steinmiiller and Sulzer. The primary helium enters the shell space from the bottom at a temperature of 950° C and is cooled down to 290° C by cross flow of the helical heat exchanger tubes. It then flows back to the primary outlet through the outer annular space, whereby the pressure shell is simultaneously cooled. The secondary helium enters the helical tubes with a temperature of 220° C via the cold gas header and exits from the heat exchanger with a temperature of 900° C via the hot header and the central tube. In comparison to the helical tube heat exchanger, the U-tube heat exchanger of the firm Balcke-Durr is differently designed (see-Fig. 7). The hot header is positioned in the upper part of the heat exchanger and subsequently the central -4- tube is shorter. The cold gas header is separated from the support plate and is suspended by springs. The heat exchanger tubes are bent in a U-shape. The primary helium flows along them on the outside. Fig. 8 presents parts of the test vessel for the two He/He heat exchangers in the manufacturer's factory. The photograph gives an impression of the size of the vessel, which has a diamter of 2,4 m and an overall height of 25 m. The pressure test will be performed in May of this year and construction will be completed in August Fig. 9 is a cross section of the hot header of the two helium heat exchangers with measuring insert for the creep buckling test. The test is performed at a differential pressure of 44 bar and a maximum initial temperature of 992° C, which is steadily decreased to approximately 750° C in the course of 10 hours. The next picture (Fig. 10) is a photograph of the hot header with tube studs and thermocouples for recording the radial and axial temperature distribution- The following photograph (Fig. 11) presents the measuring insert of the hot header for recording the creepage and buckling distances during assembly. You can see the cooling coils, which are covered with aluminum foil, and the ceramic coupling rods. Measuring insert and coupling rods are protected by an additional insulation. The first part of this experimental project was performed on February, 23 For this test, the hot header was heated to a temperature of 930° C at a pressure of 43,5 bar. This resulted in a pressure balance between the air inside the header and the helium on the outside. -5- The inside was then relieved to atmospheric pressure by blowing off the air. Subsequently the initial temperature of 930° C was reduced to 680° C at a rate of 25 K/h in the course of 10 hours. The overall deformation was determined using 24 dis- placement transducers. The results are currently being evaluated. The next creep buckling test at a temperature of 992° C is intended to be performed on April 24 e.g. on Tuesday of next week. In order to perform creep fatigue tests, the caps of all 8 64 tube studs are removed and the test object is exposed to hermocycles between 950° C and 700° C at 40 K/min as shown in Fig. 12. This is followed by another creep buckling test. The next picture (Fig. 13) shows the test section secondary hot-gas duct with metallic liner and fibre insulation. The radial dimensions, outer diameter 1220 mm, liner diameter 700 mm are in compliance with those of the PNP. The next picture (Fig. 14) presents the primary hot-gas duct with fibre insulation with 95 % AL2O3. The gas liner is made of graphite. The adjustable supporting elements made of isostatically compressed AL?O_ position the gas liner radially and axially. The following picture (Fig. 15) depicts the sub-component of the axial hot-gas valve. The valve is cooled and actuated by helium. -6- Fig. 16 gives an overview of the tests which are scheduled for the KVK in the coming years. These will be described in greater depth in subsequent lectures. - The experiments on a test section of the secondary hot-gas duct, which commenced in 82, will be completed this year, with the exception of the long-term test. A corresponding tube bend will be tested in 85/86. - The primary hot-gas duct and primary compensator will be tested with the first helium heat exchanger in an integral test in 85/86. - The sub-component of the axial hot-gas valve will be subjected to a number of different tests relating to thermohydraulics, tightness etc. before a prototype of the reactor valve is included in a long-term test in 86/87. - The tests on the hot header, creep buckling test 1, fatigue test, creep buckling test 2 will take place in 8 4/8 5. They commenced last February. - The first helium heat exchanger will be delivered at the beginning of May 1985. The tests will start at the beginning of October 85 and will continue for a period of 7 months corresponding to an operating time of approximately 3 000 h. The second helium heat exchanger will then be installed. 4. Operating experience and test results To date the KVK has been in operation for 5000 h,1200 h of which have been at reactor temperatures between 900° C and 950° C. Apart from two leakages in the natural-gas-fired helium heater and a ground fault in the electric heater, the operating experience has been very positive. After approximately 2500 h, a leakage was detected in the natural-gas-fired helium heater when shutting down the facility. There was a crack in a bend made of Incoloy 800 H, which must have formed during the fabrication of the tubes. -7- With reference to the electric heater, part of the perforated liner of the inner insulation had loosened from a holding plate and caused a ground fault over one of the segments of the power supply lines. In compliance with the electric circuitry, 2 of 12 modules were switched off and the electric heater continued operation with 10 mo- dules, as planned for such a case. The defective point has been repaired in the meantime. To conclude, I would like to list the most important facts which have been amassed to date. They are summarized in headlines in Fig. 17. - The helium tightness of the overall facility is good. It amounts to < 1 kg/d. - The combined application of a natural-gas-fired and an electrically-powered helium heater ensures the high economy and availability of the facility. - He-blowers, steam generator and steam-heated helium preheater operate reliably. - The selected materials for internal insulations proved themselves right away. - The components do not show any inadmissible vibrations. - The required helium atmosphere can be easily adjusted using the available helium purification and dosing system. - The surfaces of materials in the high-temperature zone between 900° C and 950° C exhibit a stable chromium oxide protective coating. This is also valid for the included metal samples, which can be exchanged during operation. -8- To date, it has been possible to convert and extend the KVK without any problems and within a short time. The new process instrumentation and control system Teleperm-M is easy to operate and reliable. New operating conditions can be easily realized by software alterations. Fast repair of the defects and disturbances in the heaters was possible. Component Test Facility Operating data: Thermal power 10 MW (max. 12,8 MW) Temperature 950 °C (max. 1000 °C) Pressure 40 bar (max. 46 bar) Row rate 3 kg/s (max. 4,3 kg/s) Helium velocity 60m/s Temperature transient ± 200 K/min Pressure transient 5 bar/s 900 T 950 C 1AK 1T2 5X1 220 °C 1A3 1A11 1A12 240°C 1X3 1X 2 PrelTeater 1T2 Valves Test Section 1H 1 Gas Heater 5X1 Waste Heat Boiler 1H 2 Elec:tnc Heater 1X3 Cooler 1T 1 Hot Gas Duct 1k Test 1P2 Blower Hot Gas Header f Section SL Shock Line KVK Komponenfenyersuchskreislauf SINGLE CYCLE ARRANGEMENT 220°C 1X2 Preheater 1T1 Hot Gas Duct Test Section 1H1 Gas Heater 1T2 Valves Test Section 1H2 Electric Heater 5X1 Waste Heat Boiler 1X1 He/He Heat Exchanger 1X4 Cooler 1X3 Cooler 1 P2 Secondary Blower 1 P1 Primary Blower SL Shock Line KVK KomponentGnyGrsuchskreislauf Dual-Cycle Arrangement with HG/HG HX Working and Demonstration Model of the Facility The KVK is sponsored by the Ministry of Economics, Small Business and Traffic Abb. 4 View of the KVK Facilit- Abb. 5 fbl. Secondary Outtet r Cold Gas Header Secondary Inlet Support Plate Central Hot Gas Duct Outer Annulus Pressure Vessel insulation Operational Data ; i Primary Secondary I Flow rate 2,95 kg/s ; 2,85 kq/s j Temperature 950/293°C , 900/220°C : I Pressure , 39,9 bar 41,9 bar i Diff. Pressure • 0,55 bar j 1,65 bar Power I 10 MW Dimensions and r•tatenal Number of tubes 117 DtTn. of tubes '22 "2,0 Tube Material 2.4663 (Nicrofer 55201 Vessel Material 16368 (WB36) Structure Material 1.7380 (10CrMo910) 1.4876 (Incoloy 800H1 Hot Header Mixing Device Primary Inlet 10MW He/He Heat Exchanger in Helical Tube Construction L_ Secondary Secondary Outlet Inlet Support Plate Central Duct Cold Gas Header Outer Annutus Pressure Vessel Hot Header Operational Data Primary Secondary Flow Rate 3,0 kg/s 2,9 kg/s Temperature 950/293 'C 900>220°C Pressure 39,9 bar 43.5 bar Diff. Pressure 0,5 bar 1,0 bar Power 10. MW Number of tubes 180 Tube Dimension "20 * 2,0 Tube Material 24663 (Nicrofer5520Co) Vessel Material 1.6368 (WB36) Structure Material 17380 (10CrMo910) 16311 (20MnMoNi55) 1.5415 (15 Mo 3) Insulation Hot Gas Central Duct Primary Intet Mixing Device 10MW He/He U-Tube Heat Exchanger Pressure Vessel of the 10 MW He/He Heat Exchanger Abb. 8 Instrument Penetration Water Cooling Strain Gauge Insertion Test Vessel Insulation Liner Mot Header 864 Tube Nippel Helium - Inlet Operational Oata Temperature 950 °C • 1000°C Diff Pressure 44 bar Dimensions Material Tube Nippet "22x2 2 4663 (Nicroter 5S20 Co) Hot Headsr * 1020 x 100 Wd 2.4663 (Nicroftr 5520Co) Ttst Visstl •2300x4490 1.5415 (15Mo3l Lintr 1 4876 (Incoly 80QH) Hot Header Creep Buckling Test s Hot Header -. Assembly of Thermocouples Abb. 10 Strain Gauge Insertion of the Hot Header Abb. 11 Helium-Inlet Instrument Penetration Test Vessel Insulation Operational Data Flow 2 kg/s by 9S0°C Pressure 3 kg/s by 700 °t Temperature 700 °C * 9S0°C Ramp Dimensions Material Tube Nippel "22x2 2 1.663 (Nicrofer SS20 Col Hot Header '1020 « 100 Wd 2 i.663 (NlcroferSS20 Co) Test Vessel - 2300 « 1 5/.15 (15Mo3l Liner 1 (.876 llncoloy 800H1 Hot Header Fatique Test 5970 Insulation ,t _LL -U Displacement Space - -• •*•- — N \ \ \ • 7" Cross Section A-B Support Tube Pressure Vessel 15Mo3 Liner Incoloy 800 H OESIbN DATA TtST DAIA P 30 bar GQi ' 1.0 bar PGQS = TTub« • 1S5°C TTube = 123°C Qlm 6<.S0 W/m Q/m 5023 W/m (Heut loss) *F 0.6 W/mK XF 01.8 W/mK (Heat conductivity) Secondary Hot Gas Duct with Metallic Liner and Fibre Insulation 6000 L < / (_, L. A x \ X XXXX V x x x x x x x s CO o csj o o 00 •a l x xxx XX XX JXrh x xy,t x x x^x •>t1x >f- x>;x xx xxxx x ' ' ^ ^ ^ C £ Displacement body Detail B Detail C Cross-section A-A Test Section Primary Hot Gas Duct Staling ferrule with He Outlet He Inlet Control Cat bellow ••aling element Cylinder Seating 0a> static bearing Design Data Operational Data Medium Htlium Temperature 900 °C Temperature 9S0*C Pressure 1,3 bar Pressure ti bar Flow rate 3.0 kg/s Diameter 13*0 MI Length Hot Gas Valve (System;Axial Valve) Versuche VERSUCHSTERMINPLAN KVK STAND DEZEMBER 1983 Montage 1982 1983 1984 1985 1986 1987 8 10 2468 10 2468 10 2468 10 2468 10 2468 10 _i—i—i i_ i i i i i i i i_ _i i i i i I i i i i ' Heißgasleitungen Sekundärheißgasleitung mit metallischem Liner Rohrbogen mit metallischem Liner Primärheißgasleitung Primärkompensator Axial-Heißgasarmatur Teilkomponente • Thermohydraulik • Dichtigkeitsuntersuchung _ .Inspektion • Stoßdämpfertest • Funktionsspiele V Betriebsheißgasarmatur (axial) Prototyp-Armatur Heißer Sammler • Kriechbeulversuch Nr. 1 • Ermüdungsversuch • Kriechbeulversuch Nr. 2 Helium-Wärmetauscher 1. He-Wärmetauscher 2. He-Wärmetauscher KVK-Experience in Statements Good helium leaktightness (<1 kg/d) High availability and economy through combined helium heating system Reliable performance of helium blower, steam generator and helium preheater The inner insulation systems attest to their design suitability No undue vibrations of the components Good attainment of the required helium atmosphere Stable chromiumoxide layer in the high temperature region, neither decarburisation nor carburisation Impediment free carrying out of construction modifications and extensions Reliable and congenial performance of the applied process control system Damages and disturbances of the heaters could be expeditiously removed No. 31 XA0055841 OPERATING EXPERIENCES WITH HEAT-EXCHANGING COMPONENTS OF A SEMI-TECHNICAL PILOT PLANT FOR STEAM GASIFICATION OF COAL USING HEAT FROM HTR Dr.-Ing. R. Kirchhoff and Priv.-Doz. Dr. rer. nat. K.H. van Heek Bergbau-Forschung GmbH, Essen, FRG 1. Concept and Operation of the Semi-Technical Pilot Plant Within the framework of the PNP-Project Bergbau-Forschung GmbH of Essen has been operating a semi-technical plant for the development of a process of gasifying coal by means of nuclear heat /I,2,3/. Here gasification is for the first time implemented in a fluidized bed using the heat of an electrically heated helium cycle at pressures up to 40 bar and temperatures normal with a HTR /4,5,6/. The plant - a general view of which is given on Fig. 1 - serves in a first Fig. 1: General view of the semi-technical coal gasification plant on the test premises of Bergbau-Forschung in Essen - 2 - line for testing and developing various components as immer- sion heater, insulations, dosing devices, etc. and, secondly, for gathering sound data for further planning. The second objective can, of course, be met only if any and all plant components function in a way so as to guarantee trouble- free stationary experimental operation. Fig. 2 is a flow- sheet of the experimental plant. COAL COAL INJECTION Fig. 2: Flowsheet of the semi-technical pilot plant The non-caking coal is dose-fed via the pressure lock system to the top of the gasifier. Caking coal is introduced via an injection feeder in the fluidized bed on the gasifier bottom. The injector concept is to prevent agglomeration of the feed coal. Fresh coal and partly gasified coke shall be mixed fast and intimately enough so as to avoid any con- tact and, consequently, sticking together of the fresh coal particles. The gasification residues will be discharged through the gasifier bottom by means of a chamber-wheel serving a system of two parallel locks. Jib - 3 - The helium of the heat carrier cycle is heated in an electric helium heater up to temperatures between 900 and 1000 °C; from there it flows through the heat exchanger in the fluidized bed and provides the necessary process heat. The steam is raised in a gas-fueled steam generator as well as by heat recovery from raw gas in a pressurized gas cooler. The product gas leaves the gasifier through the top and is then fed to a cyclone where entrained fine dust is removed at relatively high temperatures and discharged through a system of locks. Subsequently the gas will be cooled. A scrubbing system serves for cleaning the raw gas and removing the remaining, untransformed process steam by condensation. The product gas is then measured and analysed and subsequently burnt in a flare burner. Planning of the experimental plant started from 1974. Con- struction work was commenced in spring of 1975, and the plant was commissioned in July 1976. Fig. 3 gives the yearly operational hours, coal throughputs and transformation rates. The overall operation including commissioning, closing down, and test operation, amounted to 24 571 hours until Decem- ber 31st, 1983; 17 848 hours thereof account for hot operation and 12 611 hours for gasification. The years 1976 and 1977 were characterized by commissioning and functional tests as well as by gasification tests for the purpose of overcoming toothing troubles with process technology. It was only after this phase that systematical gasification trials could be run on non-caking feed coals. The design figure of 4.8 t/d of carbon gasified was slightly exceeded in 1978 (4.9 t/d). The objective from 1980 onward was to establish sound data for the gasification of high volatile bituminous coal with and without addition of a catalyst. J7? - 4 - Hours of operation : h/a Total hoif^ 4259 of operation 7H»» •:•;:• 24571h 2841 • Hours of hot operation 16B1 1271 111 I Illll 17B£.8h 1500 EH Hours of rod Illl 2404 qasificotiofi Illll Hfiilh Caking cool Coal throughput: t/a Non ^^^ Total coal operating frjljltl throughput 21311 Coking cool 13 Nonpretreolec 211: mI 1 | | Pretreated . cokuioled carbon Carbon gasified t/d Decreasing gasification rate 'j up TO 5,6 ;;•: "up'to 4,9 •: of 'inonces '.'r •.'•."•'• :.'v ;.'•'•'•'• "-•-•'':' ;' up to 2,4'£: ! 1978 1979 1980 1981 1982 19B3 1976-1983 Fig. 3: Figures on test operation at the semi-technical plant In the course of these trials and based on practical experience we were able to permanently improve the technical performance of the injector with appertaining feed hoppers, the feed bottom, the heat exchanger of the gasifier, and the ash discharge system. It was ultimately the successful tech- nical achievement of the above components which allows to entirely control the mechanism of pyrolysis. During 1983 the carbon transformation was increased to 6.4 t/d; coal throughput was 572 t, 407 t thereof were high volatile bi- tuminous coal /8,9/. 2. Development of several components In the following operating experiences concerning gasifier, helium circuit and the big gas heaters installed in the plant will be reported. - 5 - 2.1 Gasifier_with_the_Immersed_Heat_Exchanger The gasifier (Fig. 4) is a detail taken over from the commer- cial gasifier in so far as fluidized bed height and arrange- ment of the heat exchanger tubes are identical with the large-scale concept. The fluidized bed has a cross-section 2 of 1 m and may be up to 4 m high. Helium Helium Coal feeding i"ube Injection feeding steam Superheated steam Fig. 4: Gasifier of the semi-technical coal gasification plant Jli - 6 - It turned up already during the initial operating phases during 1976 and 1977 that the insulating properties of the gasifier brickwork lining were not sufficient. Heat transfer into the brickwork was not by conduction but rather by con- vection. Under gasification conditions the steam condensed in the lining, at saturated steam temperature. The heat had, consequently, to be taken down within the remaining insulating brickwork. The chemical and physical structure of the insulation became that much damaged thereby that steam by-passes through the brickwork were created. As this presented a hazard to the pressure shell and also caused substantially increasing heat losses the insulation had to be replaced by a concept, in which steam condensation was prevented. The heat exchanger - flown through by helium - in the gasi- fier underwent several modifications as practical experience was broadened and for complying with the different test objectives. The first apparatus was a single-flow heat ex- 2 changer of roughly 42 m effective surface; the helium flow was directed around semi-circular tube arches through the rising descending tubes. Fig. 5 shows the tube bundle prior to installation. From October 1977 until June 1979 the device had operated for 4900 hours, 4350 hours thereof under gasification con- ditions and with a coal/coke throughput of 470 t. The tem- perature range was between 800 and 900 °C. After this period the tubes were in perfect condition. In the summer of 1979 the heat exchanger was replaced by a shorter version. This was done because subsequently to the trials on non-caking feed coal it was planned to dose-feed high volatile bituminous coal through the injector. The tubes had to be shorter to provide the necessary free space in the coal feed section. During 1980 the heat exchanger was re-modified since, due - 7 - Fig. 5; Single-flow heat exchanger including semi-circular tube arches to the good heat transfer conditions, the relative tempera- tures of helium and fluidized bed had become so close to each other after just about 50 % of the heat exchange sur- face that no further heat reduction was brought about on the remaining surface. So, to make better use of the heat, the entering helium flow was distributed over the two halves of the surface. The helium circulation volume was increased at the same time. Fig. 6 is a view of the apparatus instal- led early in 1982 and which differs from the previous con- cept mainly by hair pin tube curves instead of semi-cir- cular tube arches, leaving no space between the rising and descending immersion-type tubes. This avoids reduction of the process space in the hair pin area. Such concept en- larged at the same time the absolute distance between tube pairs from 62 mm to 142 mm which provides more vertical freedom of movement to the fluidized bed. Fig. 6: Heat exchanger in gasifier, hair pin tube arrangement 2.2 Helium_circuit Another innovation is the helium cycle whose components operate at 40 bar and up to 1000 °C. Fig. 7 is a drawing of the different units operating within the high temperature range with appertaining pipelines conveying hot helium. The heart of the system is the helium heater consisting of one vertical pressure vessel of about 9 m height and an outer diameter of 1.5 m. The helium is heated with elec- tricity. The resistance heater is a pitch coke fixed bed. Tube 3 Tubel Fig. 7: Arrangement of the pipelines for hot helium within the semi-technical coal gasification plant Early in 1976 were carried out initial functional tests (pressure test, test runs), followed by a two months' trial run. During this period the circuit was operated both as - 10 - open cycle on air and as closed cycle at 3 bar operational pressure and nitrogen as heat carrier. The hot gas lines with interior insulating brickwork were pre-dried simul- taneously. During this period the plant was operated on nitrogen under a pressure of 15 bar. The unexpectedly high humidity of the inner bricklining led, during the subse- quent trial phases, to heat carrier losses and some damages at the inner lining of the helium heater, this due to gas formation by the reaction of coke with steam. As soon as the lining brickwork was dry enough helium was used as heat carrier, and operational pressure was increased to 40 bar. Mid-1978 the helium heater was scaled up to 0.9 MW. Given the optimized conditions for use of high volatile bituminous coal, a second up-scaling of the electric capacity of the helium heater to 1.7 MW was implemented during the third quarter of 1980. At the same time a new helium circulating compressor of 21 000 m /h (i.N.) capacity was installed (the capacity of the previous compressor had been 10 360 m /h (i.N.)). The above modifications proved their usefulness in the course of the further test operation. During all of the trial phase thermal investigations were carried out concerning the helium/helium heat exchanger since practical experience on this apparatus is significant for the planning of bigger plants under the PNP project. Table no. 1 compares design and operational data. Table no. 1: Design and operational data of the helium/helium heat exchanger Design Operational data data Surface m2 103 103 Helium volume flow m3/h (i.N.) 22 000 20 116 Operating pressure bar 40 39.3 Heat transfer W/m2K 234 217.1 6ln 57.9 74.7 Heat withdrawn kW 1 396 1 671 % 91.5 91.6 Efficiency - 11 - 2.3 Gas_heaters Aside from the electric helium heater there are two other gas heaters installed at the semi-technical gasification plant. They are one gas-fueled and one electric steam heater. The operational data of these apparatuses can be seen, further to the data of the electric helium heater, on Fig. 8. The gas-fueled superheater brings the steam raised in the steam boiler to 900 °C. During this the tubes of the heat exchanger - Fig. 9 - have to resist to the full pressure gradient of 40 bar between fueling section and steam sec- tion. This heater went defective in February 1978. Electrically heated Gas. fired by wire 810 kW 600 kW Power: 2 t/h 21000 Nm3/h; 1700 kW 2 t/h Fig. 8: Big gas heaters at the semi-technical coal gasi- fication plant The cause of this were three cracks in the supporting tubes, The complete heat exchanger system had to be dismantled and checked. It was found out that, aside from the cracks - 12 - in the supporting tubes made of 15 Mo 3, tube bundles no. 3 and 4, of 13 Cr Mo 44, showed considerable scaling and were badly distorted. Tube bundle no. 2, of Incoloy 800, was free of scaling and hardly distorted. Bundle no. 1, of HK 40, was in perfect condition. Upon this the supporting tubes as well as bundles 3 and 4 were manufactured out of Incoloy 800 The new tube bundle - Fig. 9 - has resisted to test operation without any damages so far. Fig. 9: Heat exchanger tube bundle of the gas-fueled steam superheater - 13 - The electric steam superheater consists of a vertical insu- lated pressure vessel into which process steam is entered and passed evenly distributed through 36 annular tube gaps arranged in parallel whereby it is heated up to maximum 900 °C discharge temperature. In the center of the double tube construction is the electric heat conductor, mounted on ceramic material, including tmeperature controls at two different levels in order to protect the heat conductor from overheating by a cut-out system. In Fig. 10 is repre- sented the top section of the electric steam superheater including the heat conductor arrangement. Certain problems were encountered with the electric steam superheater during the past. During the spring of 1979 - like in previous years - the heating wires of the conductor repeatedly went defective due to inadequate welding connections and excessive material deposits. Fig. 10: Heating arrangement in the electric steam superheater - 14 - Damages of the above kind can be seen on Fig. 11. One of the causes was found out to be the excessive nickel content so that in April 1979 the heating cartridges were replaced by items of different quality. Further improvements were possible by using heat conductors of Incoloy 800 whose varying cross-sections were obtained by hammering. Fig. 11: Damaged heat conductor The new heating arrangement was equipped additionally with an inner shell preventing major heat differences thanks to its aluminium oxide fibre lining. Major heat differences entraining varying elongations had, indeed, contributed to the damages and, along with unfavourable mounting arrange- ments, led to distortions of the heat cartridges and thus, to ruptures of the heating wires. - 15 - Based on practical experience the installation has been optimized so that the basical problems are overcome by now. And there is a sufficient amount of sound data available to back up the erection of a plant scaled-up in size. 3. Bibliography /I/ Juntgen, H. and van Heek, K.H.: Coal gasification, - fundamentals and technical application (Karl Thiemig, Munich, 1982) /2/ Schilling, H.-D.; Bonn, B., and Kraufi, U.: Coal gasi- fication, 3rd edition (Gluckauf GmbH, Essen, 1981) /3/ Neef, J.H., and Weisbrodt, I.: Coal gasification with heat from high temperature reactors. Objectives and status of the project "Prototype plant for nuclear process heat (PNP)", Nucl. Engr. Des. 54 (1979), 157-174 /4/ Juntgen, H. and van Heek, K.H.: Gasification of coal with steam using heat from HTRs, Nucl. Engrg. Des. 34 (1975), 59-63 /5/ Feistel, P.P.; Diirrfeld, R. ; van Heek, K.H. and Juntgen, H. Layout of an internally heated gas generator for the steam gasification of coal, Nucl. Engrg. Des. 34 (1975), 147-155 /6/ van Heek, K.H.; Juntgen, H., and Peters, W.: Steam gasification of coal using process heat from HT reactors. Atomenergie/Kerntechnik 40 (1982), 225-246 /I/ Kirchhoff, R. ; van Heek, K.H. and Juntgen, H.: Operation of a semi-technical pilot plant for allothermal pressure gasification of coal using steam. Compendium 80/81, Erdol und Kohle (supplement), pp. 177-179 /8/ van Heek, K.H. and Kirchhoff, R.: Present state of coal gasification with nuclear heat. Publication of papers read at a conference held at Haus der Technik, Essen. Issue 453, 1982, pp. 59-67. /9/ Kirchhoff, R.; van Heek, K.H.; Juntgen, H. and Peters, W.: Operation of a Semi-Technical Pilot Plant for Nuclear aided Steam Gasification of Coal. Nucl. Eng. Des. 78 (1984), pp. 233-239 No.. 32 XA0055842 IAEA Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors Diisseldorf 16.-19. April 1984 THE TEST FACILITY EVA II/ADAM II Description and Operational Results R. Harth, H.F. Niessen, V. Vau Kernforschungsanlage Julich GmbH R. Merken Rheinbraun-Koln THE TEST FACILITY EVA II/ADAM II - Description and Operational Results - R. Harth, H.F. Niessen, V. Vau, KFA-Julich R. Merken, Rheinbraun-Koln The Nuclear Research Center Juelich (KFA) and the Rheinische Braunkohlenwerke AG, Cologne, signed the contract for the R&D project "Nukleare Fernenergie (NFE)" in December 1975. Among others one task of this project has been the construction and operation of the test facility EVA II/ADAM II. General aim of the project work is to elaborate the data necessary for the design and construction of a heat transport system using the thermo-chemical cycle steam reforming/methanation. Fig. 1 shows a view of the complete facility sited at KFA Juelich It consists mainly of a helium system as heating unit, a steam reforming system and a methanation unit. The large building in the background contains the helium loop EVA II. The methanator ADAM II is located in the foreground. There the synthesis gas produced in the steam reformer is converted to methane again and the heat transported by the thermochemical cycle is re- leased at 650 °C. The complete facility EVA II/ADAM II has been constructed by Lurgi Kohle und Mineraloltechnik GmbH, Frankfurt. In the following the facility EVA II is discribed. The flow scheme is shown in fig. 2 The helium circuit represents a complete primary loop of a HTGR for process heat application. The core is simulated by an elec- trical heater with a maximum power input of 11 MWe. The helium is heated in the electrical heater up to 950 C. After passing the hot gas duct the heat is transferred succes- sively primarily to the steam reformer tubes where the helium JJ) - 2 - is cooled down to 650 C and then to the steam generator for process steam production where the helium is cooled down to 350 C. An integrated circulator transports the helium back to the electrical heater by following a coaxial flow principle. As process gas a mixture of methane and steam enters the steam reformer. By means of the endothermic chemical conversion to synthesis gas it absorbs the heat transferred from the helium. Fig. 3 shows the steam reformer bundle tested in EVA II. It consists of 30 tubes. The tube dimensions corresponds to those in the conventional technique: . length 11 m . diameter (internal) 100 mm. Further characteristics of the steam reformer bundle: . 4 different alloys have been used for the reformer tubes . Raschig-ring-catalyst as conventionally approved . reformer tubes with internal return pipes . baffles (disc-and-doughnuts) to intensify the helium heat transfer. The steam generator is shown in Fig. 4. This is a induced-single- circulation boiler, characterized by . helix design . 2 layers contrary directed coils . helium temperature max. 700 C . steam temperature max. 700 C . steam pressure max. 55 bar. Electrical heater, steam reformer and steam generator are linked together by coaxial hot gas ducts. The thermal insulation sepa- rating the hot and cold helium flow is made of carbon bricks. On both sides the insulation is covered by a metallic liner. Just the colder outside liner is gastight. - 3 - The tasks for the plant EVA II are as follows: . tests of the heat transport by a thermo-chemical cycle . tests of a complete helium loop equivalent to that of a nuclear process heat plant . tests of steam reformer bundles in a representative size and design . functional tests of a steam generator under HTGR like conditions . investigation of operational behaviour on normal operation, partial load and break-down conditions . description of the operating characteristics by mathematical models. The main tasks of the investigations in EVA II aim not prior at the life time of the components but at operation characte- ristics and behaviour in a wide range of different process parameters. The main parameters have been changed between the following limits: electrical power input 3,8 •+• 11 MW helium mass flow rate 1 -*- 4 kg/s helium pressure 15 -*- 40 bar helium temperature 800 + 950 °C methane mass flow rate o, 18 •*• 0 ,66 kg/s. All test runs of the facility EVA II in linkage with ADAM II could be performed successfully. In the course of the test program the following tasks have been worked off: . change of catalyst by vacuum extraction . replacement of a single reforming tube without removal of the whole bundle . disassembling and reinstallation of the steam reforming bundle. In total the helium system has been operated for 7,800 hours. Thereby the EVA/ADAM heat transport cycle reached 5,660 hours. - 4 - The test of the first steam reformer has been finished 1983. The bundle has been removed. It will be disassembled and the steam reformer tubes will be checked in detail. The facili- ty EVA II now is under preparation for installation of a new bundle with a different design. The test of the component steam reformer in the facility EVA II represents the last experimental step before entering into nuclear demonstration. Acknowledgements The work described in this paper was performed within the framework of the project "Nukleare Fernenergie" between the Kernforschungsanlage Jiilich GmbH, Jvilich and the Rheini- sche Braunkohlenwerke AG, Cologne, sponsored by the Federal Minister of Research and Technology, FRG. Fig. 1: View of the test facility EVA II/ADAM II T °C 40 40 P bar 41.4 3a5 m kglB 0.619 1.234 CH, rel.Vol. 0.9S1 0.123 Hj „ 0.039 0.6B1 CO — QO96 CO, 0.010 0098 EVA 31 Fig. 2: Flow scheme of the facility EVA II Fig. 3: Steam reformer tube bundle Fig. 4: The helium heated generator of process steam No. 33 XA0055843 Modification of the AVR for High Temperature Process Heat Systems-Demonstration * * ** Barnert, H. , Kirch, N. , Ziermann, E. * Kernf orschungsanlage Jiilich GmbH ( ) ** Arbeitsgemeinschaft Versuchsreaktor GmbH ( ) IAEA Specialists' Meeting on Heat Exchanging Components of Gas-Cooled-Reactors Diisseldorf, Federal Republic of Germany 16-19 April 1984 Sponsored by Ministerium fur Wirtschaft, Mittelstand und Verkehr des Landes Nordrhein-Westfalen Hodification of tiu. A Systems-Demonstration for High Temperature Process Heat Applications 2. Systems- Demonstration for Refinement of Fossils, eg. Ctol into Energy Alcohol 3. Contribution to Specific trtas : ProcCa ct ^peci fi cat (on s Licensing Procedures A-pproado of 2ero E 4. Investigations for the Horizontal Integration Systems-Demonstration with AVR Gasification Steam Generator Heavy Oil Methane Splitting Reformer Steam Coal Gasification \^ y Synthesis Energy-Alcohol Intermediate Heat Exchanger Hydrogen Coal Gasification •/ 1. Reactor AVR 1.1 Thermal power of AVR Core 46 MWt 1.2 Mean outlet temperature of coolant He 950 °C 1.3 Pressure of coolant He 10 bar 1.4 Mass flow of coolant He 13 kg/s 1.5 Power existing steam turbine plant: HTP-Loop 50 50 2 Existing Steam Turbine Plant 2.1 Thermal power of the steam generator 23 MWt 2.2 Electrical power of the turbine generator 9 MWe 2.2 Comparable to (to days) part load 50 % 3 High Temperature Process Heat Loop 3.1 Thermal power of the loop 2 3 MWt 3.2 Temperature of the helium at the components 950 °C 3.3 Design temperature of the hot gas duct 1000 °C 3.4 Length of the hot gas duct 60 m 3.5 Inner diameter of the hot gas duct 480 mm 4 Reformer 4.1 Thermal power of the reformer 8,5 MWt 4.2 Helium temperature, inlet/outlet 950/700 °C 4.3 Feed methane, volume flow, ca. 2500 mVh 4.4 Molar ratio H2O:C, ca. 4 : 1 4.5 Process gas maximum temperature 825 °C 4.6 Number of reformer tubes 65 5 Steam generator 5.1 Thermal power of the steam generator 14,5 MW 5.2 Temperature of the steam 530 °C Loop circulator 6.1 Pressure increase of the circulator 1,0 bar 6.2 Power of the circulator 500 kW 7 Measurements of the buildings 7.1 Height of the AVR building 49 m 7.2 Height of the loop building with hall 48 m 7.3 Height of the loop building without hall 28 m 7.4 Length x width of the loop building 23 x 14 m Frischd&mpf Produkt, z.B. Synthesegas Einsatz, z.B. Methan AVR High Temperature Process Heat Loop 1 Core 10 Hot gas tube 2 Steam generator 11 Hot gas duct 3 2 Circulators 12 Cold gas duct 4 He T = 950 ° C 13 Process heat components 14 Circulator for the loop Fig. 2: Principle of the Modification [ oo CM cn fN od < o col o x: O o o frj o a • o t?> £ > o 95 0 co •MM ————— • A •••r—i ICQf X tn \ LU Q A a 6, 5 ^o ^. o O P k in J CO cn QO 1f o 1 i n F E (J i E o o O CO tri e • CaO) • i—i /i o c_ P > in en i" < c lin © CM i LU > < KFA- I a," =^ r AVR High Temperature Process Heat Loop 1 Core 10 Hot gas tube 2 Steam generator HHot gas duct 3 2 Circulators 12 Cold gas duct 4 Reactor pressure vessel 13 Reformer 5 Stages 14 Steam generator 6 Containment 15 Circulator of the loop 16 Loop building 17 Conncection 18 Hall with crane 19 Hot cell 20 Stack Fig. 3: AVR-Process Heat Plant, Cross Section 1 AVR High Temperature Process Heat Loop 1 Reactor 10 Loop building 2 Turbine house with High Temperature Process Heat Loop 3 Hot shop 19 Hot cell 4 Coolers 20 Stack 5 Pump house 6 Main gate Fig. 4: AVR Process Heat Plant, Buildings 'IT // • •////•;'>/ , ,;^\ br s Kamera Befeuchtungskorper -••22,3m Rohrbefestigung +18,2 m Ron/ 075x43 innerer Reaktorbehalter auBerer + 10,1m Vorschubeinrichtung • 5m AVR - Deckenref fektor - Inspektion Zentrales BE- Forderrohr AVR - Deckenref lektor- Inspektion Core - Kamera 640 hydr. Kamera- CO Bewegungseinricht Core-Kamera TrennsteJIe / }\\\\\\\\\\\\\\\\} J Kupplung 155 —4 r=200 Tht Work Principles Conditions of Safety Life Time Limits Special tasks attars 85 Decision 85 -86 PL a em (19 a I ndus% -90 Operati'on. Of -Wie. AV^ Temp. P+i Loop A Novel Horizontaly Integrated Energy System, NHIES Zero Emission Decomposition Stochio metric Synthesis and Complementation with and Final Energy Cleaning Internalization of Costs Conversion f(NOx) fcO2 CH,OH Motor Fuels / EL Electricity t (N0x) t CO; NPH NucI.E. -EL Heat Uranium HTR IFP NPH No. 34 •III XA0055844 HEAT REMOVAL BY NATURAL CIRCULATION IN GAS-COOLED ROD-BUNDLES M. Hudina Swiss Federal Institute for Reactor Research 5303 Wurenlingen / Switzerland Paper presented at Specialists' Meeting on Heat Exchanging Components of Gas-Cooled Reactors Dusseldorf, BRD April 16-19, 1984 HEAT REMOVAL BY NATURAL CIRCULATION IN GAS-COOLED ROD-BUNDLES by M. Hudina / Swiss Federal Institute for Reactor Research 1. INTRODUCTION Experimental and analytical investigations in the field of gas-cooled reac- tor core thermohydraulics have been performed at EIR since 1964, Temperature and pressure distributions in rod bundles of different geometries and rod surface finish have been measured over a wide range of flow conditions, star- ting at high turbulent Reynolds numbers (up to 10^), over the transition re- gion and down to the laminar flow regime. The aim of these rod bundle heat transfer and fluid flow experiments was to prove the quality of the results calculated by comprehensive subchannel analysis thermal-hydraulic computer codes (code verification procedure) under nominal as well as transitional (part load and low flow) reactor conditions. The measurements were used to assess the reliability of the analytical models and the accuracy of their predictions for design purposes. More detailed information about the experiments, together with measured re- sults, have been reported earlier /I/, /2/, /3/. Parallel to the experiments, the fundamental analytical models and correla- tions for bundle thermohydraulics have been studied and incorporated in sub- channel-analysis computer codes /4/, /5/. The experimental results were then compared with analytical predictions, partially within the framework of inter- national benchmark calculations /6/. After completion of the standard measurement programme, additional test were made in order to: - measure the axial and radial temperature distribution in the bundle together with the pressure drop over the bundle length under natural circulation conditions - calculate these distributions with the aid of subchannel analysis co- des and compare the results with measured data - prove the capability of analytical predictions to be used for gas—coo- led rod bundle design purposes. - 2 - 2. EXPERIMENTAL EQUIPMENT For the tests a high pressure, high temperature loop with CO2 as cooling gas with the following characteristics was used: Coolant pressure 1-60 bar Coolant temperature 30-500 °C Maximum mass flow 4,5 kg/s Heating power 0-1000 kW The loop was designed to carry out steady state fluid flow and heat transfer single pin or rod bundle tests over a wide range of fluid flow and heat flux (uniform and power tilt) conditions. The flow scheme of the loop is presented in Fig. 1. For the investigation of natural circulation the correct positio- ning of the heat exchanger (about 5,7 m higher than mean heat source level) is very important. Further information about the loop can by found in /2/ and /3/. The longitudinal section of the pressure vessel with the 34-rod hexagonal test section and the cross-section of the bundle with spacer geometry is shown in Fig. 2. The rod pitch-to-diameter ratio of this bundle (Bundle 3 of the AGATHE HEX Experiment) was 1.5. HEAT EXCHANGER = temperature sensor Figure 1: Experimental Loop 'AGATHE1 3 - POWER TEMPERATURE SENSORS SUPPLY ON THE SHROUD WALL AXIAL LOCATIONS OP THE PRESSURE TAPPINGS - HEATED LENGTH INSULATION SHROUD WALL SPACER GRID PRESSURE VESSEL POWER SUPPLY Figure 2: AGATHE HEX bundle test section and cross-section of Bundle 3 (including spacer geometry), used in natural circulation tests. Jlf — 4 — 3. MEASUREMENTS AND DATA HANDLING To determine the heat removal capacity of the Loop by natural circulation, runes had been carried out under different system pressure levels, starting from vacuum and up to a maximum of 2 MPa. At constant system pressure, the mass flow was varied to simulate various situations between blower failure and full isolation of the bundle, using different combinations of valve clo- sure and metering orifices. As a limiting case (isolated bundle) the main inlet valve was closed. The maximum rod temperature was limited to approxi- mately 1000 K. In cases with gas circulation in the whole loop (open main inlet valve) the inlet and outlet gas temperatures (max. 500°C), as well as maximum cladding temperature, were kept approximately constant. The power distribution in the bundle was uniform. For all the tests detailed axial and radial temperature distributions in the bundle were measured and recorded. These data were then used as input for the DRABEX evaluation code which pro- vides a printout of all important information including punched cards as in- put to analytical calculation code. The EIR subchannel analysis computer code SCRIMP/4/ was used in this investigation. 4. RESULTS In Fig. 3 the results of the natural circulation heat removal capacity of the AGATHE loop are summarised. The results are strongly influenced by the system pressure but the pressure drop over the section of the loop with built-in ori- fices is also important. 18 % of the nominal heating power can be removed by natural circulation at nominal system pressure. For the isolated bundle core and at full system pressure this capacity drops to under 1,5 %, According to these results it can be assumed that with a suitable arrangement of the pri- mary reactor circuit the residual heat can be removed by natural circulation as long as fast depressurisation of the system and/or extreme blockages of the loop free flow area are prevented. The axial distribution of temperatures for the isolated bundle case is given in Fig. 4. In contrast to the symmetrical temperature distribution obtained for tests under vacuum (heat removal by conduction and radiation only) the influence of natural convection in the bundle was clearly observed. In Fig. 5 the measured and calculated axial temperature distributions for one particular run with opened inlet valve is presented. The radial temperature distribution (over the bundle cross-section) of this run is given in Fig. 6 and 7. The analysis of these results shows that both conduction in the rods and radiant heat transfer are important heat transfer mechanisms and are to be included in the analytical models of the codes. The radial temperature distributions for one further test run (with maximum mass flow) can be seen in Fig. 8. - 5 - In all analysed cases the analytical results slightly overpredict the wall temperatures. Considering that this is on the safe side the predictions can be judged to be acceptable to fulfil the gas-cooled bundle design accuracy requirements. 20 NOMINAL CONDITIONS : PRESSURE 2MPa (C02) MAX. ROD SURFACE TEMPERATURE 750°C Ixl o 15 0. 5 o O 10 o Q- occ Q_ PRESSURE DROP OVER VALVE/ t ORIFICE MAXIMUM COMBINATION ISOLATED BUNDLE 1 2 —• PRESSURE (MPa) Figure 3: Cooling capacity of the "AGATHE" test loop with natural circulation. J7? - 6 - T 700 •• INNER ROD (x :) SUPPORT ROD {•) 600-- 500-- 400-• 300 200-• INSULATION (•) 100- 500 1000 1500 2000 (mm) ENTRANCE 860 HEATED 950 250 R0UGHENED1200 iimiiniiinniiinl Spacers 1 2 3 4 5 6 7 8 Figure 4: Axial temperature distribution (bundle 3 of "AGATHE HEX" experiment) under natural convection conditions (isolated bundle). CENTRAL RODS (A) 700 •• WALL CHANNEL RODS(ak :) CORNER RODS(v)- 600 500- SUPPORT ROD 400- HROUD (+) GAS (•) 300- 200-- 100 •• 0 500 1000 1500 2000 X(mm) ENTRANCE 860 HEATED 950 250 1200 ROUGHENED HIM Illlll III HUH III III I I HIM III III M Mil III lit II Illllll 111 II IIIMtlll III Mill Illllltll ••••111 I Spoors: 8 Figure 5: Axial calculated temperature distribution in the bundle under natural circulation conditions - calculated (curves) and measured (symbols). jrf - 7 - a) without consideration of conduction in the rods and of radiant heat transfer b) without consideration of radiant heat transfer Figure 6: Comparison between measured and calculated temperatures (°c) at the end of heated section (Test run No. 36). Analytical predictions are written inside the rods and shroud. Figure 7: Comparison between measured and calculated temperatures (°c) at the end of heated section (Analytical predictions for test run No. 36 are written inside the rods and shroud). - 9 - Figure 8: Comparison between measured and calculated temperatures ( C) at the end of heated section (Analytical predictions for test run No. 35 are written inside the rods and shroud). - 10 - 5. CONCLUSIONS The results of this investigation lead to the following conclusions: - The heat removal capacity of a given gas loop depends considerably upon the system pressure, and also on local reduction of the free flow area (e.g. by blockages). Assuming a suitable arrangement of primary circuit components, the residual heat of a gas-cooled reac- tor can be removed by natural circulation. - For the steady-state situation after blower failure, the temperature distribution in the rod bundle can be closely predicted by the sub- channel analysis code SCRIMP. The inclusion of conduction in the rods and radiant heat transfer in the code was found to be important. - The analytical results slightly overpredict the local cladding tem- peratures (which is on the safe side). However, results are suffi- ciently accurate to justify the use of this code for fuel element design calculations under steady-state natural circulation conditions. REFERENCES /I/ Huggenberger, M. and Markdczy, G,; "Verification of Subchannel Analy- sis Computer Codes by Full-Scale Experiments", EIR-Bericht Nr. 283, June (1975) . /2/ Hudina, M. and MarkSczy, G.: "The Hexagonal Bundle Heat Transfer and Fluid Flow Experiment AGATHE HEX", Nuclear Engineering and Design, Vol. 40, No. 1, Special Issue GCFR, pp. 121-131, (1977). /3/ Hudina, M. and Mark6czy, G.: "The Hexagonal Bundle Heat Transfer and Fluid Flow Experiment AGATHE HEX", Proceedings of the ANS/ASME/NRC International Topical Meeting on Nuclear Reactor Thermal-Hydraulics, NUREG/CP-0014, Vol. 3, pp. 2278-2295, Oct. 5-8, Saratoga, (1980). /4/ Huggenberger, M.: "SCRIMP, A Thermal-Hydraulic Subchannel Analysis Computer Code", EIR-Bericht No. 322, June (1977). /5/ Barroyer, P.: "CLUHET: A Computer Code for Steady State and Single- Phase Thermohydraulic Analysis of Rod Bundles", Proceedings of the ANS/ ASME/NRC International Topical Meeting on Nuclear Reactor Thermal-Hy- draulics, NUREG/CP-0014, Vol. 3, pp. 2204-2221, Oct.5-8, Saratoga (1980) /6/ Barroyer, P., Hudina, M., Huggenberger, M.: "Benchmark Thermal-Hydrau- lic Analysis with the AGATHE HEX 37-Rod Bundle"' EIR-Bericht No. 439, September (1981) . JZb No. 35 XA0055845 MANUFACTURE OF STEAM GENERATOR UNITS AND COMPONENTS FOR THE AGF POWER STATIONS AT HEYSHAM II AND TORNESS J.R. Glasgow, K. Parkin, N.E.I. Nuclear Systems Limited synopsis The current AGR Steam. Generator is a development of the successful once-through units supplied for the Oldbury Magnox and Hinkley B/Hunterston B AGR power stations. In this paper a brief outline of the evolution of the steam generator design from the earlier gas cooled reactor stations is presented. A description of the main items of fabrication development is given. The production facilities for the manufacture of the units are described. Reference is also made to some of the work on associated components. The early experience on the construction site of installation of the steam generators is briefly outlined. 1. INTRODUCTION higher coolant gas pressures which have been progressively exploited to the Heysham II and A description of the Steam Generators Torness Design. Simultaneously, integral together with the layout, materials, environ- arrangement of boilers within the reactor concrete ment and constraints are given in the NNC Paper pressure vessel and once-through steam generators, No. 58. The major manufacturing change from instead of re-circulation steam generators, were Hinkley/Hunterston AGR is application of adopted. The main reason for the use of once- automated spacer welding process rather than through steam generators was to reduce the the manual process used earlier. The automated penetrations of the concrete pressure vessel. spacer welding process produces higher quality more repeatable welds. This has dictated the The second major change was the development adoption of longitudinal spacer configurations of uranium dioxide fuel with stainless cladding throughout the Element Supports. in the Windscale AGR leading to its application in Hinkley & Hunterston 'B' AGR Power Stations. 2. EVOLUTION OF DESIGN This allowed the advances in fuel, coolant temperatures and steam conditions which give the The early Magnox series of stations were parameters for the current stations with the limited in coolant gas temperature by the resulting improved overall cycle efficiencies. Magnox fuel cans. For these reactors, the gas outlet temperature varied from about 3^5°C The principal features of the current AGR in the first stations to 400°C in the latest. Steam Generators are:- Steam conditions were not conducive to high cycle efficiencies. A dual pressure steam (a) Large shop manufactured units of approx. cycle was adopted for all the magnox series of 90 tons. stations and a tight thermal design with close (b) A single pressure steam cycle with reheat gas to steam/water temperature approaches to made possible by the higher gas temperature. achieve cycle efficiencies as high as J,k%. 3. FABRICATION DEVELOPMENT The low heat fluxes, together with the relatively low heat transfer associated with The principal areas of fabrication develop- the low pressure carbon dioxide and steam ment were:- pressures applicable to early stations, 3.1 Tube Manipulation The special purpose twin dictated the use of extended surface tubing headed bending machine was defined and tooling for most zones of the heat exchangers to developed. Test bends produced and sectioned to restrict the steam generator pressure vessels confirm general slope and acceptable ovality and to reasonable proportions. thinning. Proof testing of typical bends was carried out with special reference to ID radius Two step changes in the evolution of the bends. design brought us to the current status of 3.2 Tube Spacer Welding Steam Generator design. (a) Development of weld process to arrive at the optimum welding procedure in terms of First was the adoption of concrete quality and strength. pressure vessels to house the reactors for (b) Consistency trials on jigging and investigations into welding shrinkage and restraint effects. Oldbury Magnox (1962). This allowed the use of (c) Confirmation of quality and consistency by the production and destructive examination of full scale production platens. - 1 - U. SHOP MANUFACTURE elements, 2112 decay heat elements and 1728 THMLEY 'B' & HUNTERSTON 'B1 reheater elements. Some 116,000 butt welds join the tubes together; the platen tubes are A decision was taken some eighteen years supported and separated by small spacers across ago to assemble the steam generators for the tubes, requiring approx. 8.48 million spacer/ Hinkley 'B1 and Hunterston 'B' AGR Power tube welds. (Figs.la, 1b, 1c). Stations in the works to standards of clean- liness that would comply in full with Taking into account the advances made in requirements for plant in nuclear installations. manufacturing technology, N.E.I. Nuclear Systems In addition, the Units were much larger than set out to plan and install a small batch/flow had previously been attempted and thus line manufacturing capability using some of the production facilities were critically reviewed techniques established in earlier A.G.R^ resulting in conversion of the existing programmes. However, significant changes manufacturing areas to provide the capability involved the introduction of:- to produce the A.G.R. boiler units to these standards. (i) improved fixed and orbital head welding machines. The main production areas were, tube (ii) C.N.C. bending, bending, tube butt welding, welding of tube (iii) robotic spacer/tube welding, supports, heat treatment, testing and final (iv) computer controlled inert gas atmosphere assembly. stress relieving furnace. The above processes take place prior to entry Ferritic materials were kept separate but into the Clean Condition final assembly area but fabricated in parallel with austenitic materials the manufacturing environment was raised to a level and the principal innovation was the of control designed to avoid any acid cleaning introduction of mechanised argon arc tube butt and thus ensure maintenance of cleanliness up to welding. transfer to clean conditions. Tubes were first welded into long lengths 6. A BOILER ASSEMBLY LINE (up to 30 metres), using fixed head argon arc welding machines, prior to bending into Raw tubes are fed onto a tube end preparation serpentine sub flow path form. Many of the facility - incorporating cutting and finishing tube bends were made to a mean radius dimension machines, (separate machines are used for ferritic equal to one tube diameter. and austenitic tubes). They are them moved to a mechanical handling system (Fig. 2) for butt Sub platens were produced by welding the welding into total lengths of 42 metres using sub flow paths together using orbital argon arc Fixed Head TIG welding machines (Fig. 3) which welding machines and all butt welds were subject were designed and manufactured by N.E.I. Nuclear to radiographic examination. Systems. The mechanical handling system then transports the tube across the shop through an The sub platens were then made into platens N.D.T. station onto the bending machine conveyor. by the attachment of many small spacers across The C.N.C. bending machine can bend up to 90 mm the tubes using manual C02 MIG Dip Transfer. diameter left and right-hand and the completed sub flow path can have as many as 16 bends, The platens were then heat treated in an 1 (Fig. t). The radius of bend can be as small as electric inert gas 'Top Hat furnace followed one tube diameter with close control on ovality by hydraulic test and a whole surface leak test. and thinning. Where necessary the ferritic steel platens were cleaned in a special acid cleaning plant prior to transfer into the clean conditions shop where The sub flow paths from the bending machine element assembly and loading was carried out. are jig assembled for orbital welding, employing a 'U' head orbital TIG welding machine (Fig. 5) developed and manufactured in house. The The units were all delivered to site by equipment incorporates a weld programmable April/May 1971, and then there was a long pause sequencer and arc voltage control, allowing a until early 1978 when A.G.R. stations at fully automatic weld of high quality and Heysham II and Torness were announced. consistancy. 5. SHOP MANUFACTURE HEYSHAM II & TORNESS To weld the tube spacers, a fully automatic method has been developed incorporating robots. General views of the installation are shown in The production facilities were re- Fig. 6 & 7. established along the lines already described but a large production investment was made to It is fundamental to the design of the introduce advanced manufacturing techniques and boilers that consistant high quality of welds 'A boiler assembly line1 was born. is produced. In this case the quality is dependent on:- It had earlier been recognised that nuclear boiler platen manufacture had many features of (a) consistancy of dimensions of tubes and spacers, repeatability. For two power stations some (b) consistancy/repeatability of the equipment 12,288 platens are required, subsequently to be employed. further joined together to make 2112 main boiler - 2 - The system is divided into three distinct called up whenever it is required. Different sections:- spacer geometries have their own individual sub- routines. (a) The welding jig. (b) The weld torch positioning system. Before commencement of platen welding a (c) The welding equipment. test piece is welded and the weld examined for shape and penetration. If accepted the gantry The function of the jig is to position the is moved to the first position, the appropriate tube and spacer consistantly in relation to the welding parameter sub routine is called up and welding torch positioning system. welding is commenced. The two robots then weld all of the spacers within their operating, range. This latter system employs a gantry running On completion the gantry indexes to a new along a track. Suspended from each gantry are position and the process is repeated. two inverted robots, each carrying a welding torch. The sequence of indexing and welding is continued until the first side is fully welded. The welding equipment is based on the MIG/ MAG process backed up by improvements in the "he second side master base will be already form of electronic power sources, better wire loaded with a platen previously welded on the feed units and redesigned welding torches. first side. The gantry is moved, to this base and the whole process recommenced. In this The workpiece is loaded into a Tube Clamp- way the gantry can be kept in continuous ing Frame, and the tubes clamped into position. production. This ensures correct tube pitching. The frame is then lowered onto a fixed master base and Four interacting computers control the located accurately by holes in each corner of operations:- the frame, which sits onto dowels attached to the base. The jig master base is fixed in a (a) Robot microprocessors. permanent position on the gantry bed. Besides (b) Gantry mini computer. the location dowels the base has numerous (c) Weld programmable controller. finger devices to receive and hold the spacers (d) Memory extension mini computer. in accurately known positions. The spacers are loaded off line, into a magazine which is The installation is believed to be one of basically a rotating barrel with 'pockets' for the most advanced applications of robotic weld- holding individual spacers. The magazine which ing in the world and todate almost seven is on wheels is located on the master base. million welds have been produced. Individual rows of spacers are then loaded by lowering the magazine towards the workpiece. 7. HEAT TREATMENT The master base fingers are then actuated and the magazine raised leaving the spacers in On completion, the boiler platens are position for welding. The magazine is then subject to post fabrication heat treatment to moved across the base and each row of spacers relieve stress and restore properties affected loaded. The magazine is removed from the base by bending and welding operations. But good and the platen is ready for welding. surface condition of the tubes has to be preserved to meet the stringent engineering The range of weld torch positions is requirements. The heat treatment is carried limited by robot reach and is optimised by a out in an electric 'Top Hat' furnace which combination of gantry and robot movements has facility for air evacuation prior to filling allowing full traverse of the workpiece. with nitrogen to provide an inert atmosphere with less than 7 parts per million oxygen and The gantries are programmed to stop at any with a dew point of the order of minus 25°C. required position along the track. Hence each The furnace is computer controlled can be robot is programmed to carry out the desired operated up to 1060°C and the load can be welding sequence associated with that track rapidly cooled by circulating the nitrogen stop position. through integral heat exchangers. In more detail the gantry is griven to any- There are two loading bays to allow the one of a number of pre-determined positions on completed load to be removed from the hearth the gantry runway. Associated with each of and a fresh one returned to the furnace without these gantry positions is the appropriate disruption. The loading bays are shown in programme for moving the robots (and hence the Fig. 8. The furance is in the background. welding torch), through the required welding path at the required speed. All positioning 8. ASSEMBLY OF ELEMENTS and movement of the gantry and robots has to be individually programmed. The four platens designated, economiser, lower chrome, upper chrome and superheater are The welding parameters, e.g., current, arc joined, together via six tube butt welds to form voltage, wire feed speed, shroud gas flow and an element. The economiser 5% Cr. transition the timing of any changes in these parameters piece is joined to the 9% Cr. lower platen, the required to weld a specific spacer configuration lower to upper plates are 9% Cr. throughout and are programmed into a sub routine which can be the upper chrome platen is joined to the 'Inconel' 3 ~ transition piece on the Superheater. This beams and ties. At the same time a decay heat latter transition piece already incorporates platen is loaded into position alongside each main a 9% Cr. stub allowing like materials to be unit element. This assembly process is repeated welded to complete the element. until 22 elements are in position at which point the acoustic baffle is lowered into position and These six welds are ultrasonically secured. Figs.11a. 4 11b show units at different examined followed by heat treatment using stages of assembly. localised heating elements. A further 22 elements are then assembled to The completed element is 'balled complete the sequence. The decay heat elements are through' then hydraulically tested. This test welded to manifolds and the top casing is then facility (Fig 9) takes one element at a time lowered into position and connected to the internal supported on a manipulator to assist filling structural beam and tie members. The two side and emptying of the deionised water. It also casings are now positioned, clearances checked then allows positioning at any elevation for bolted and seal welded. Temporary steel covers are inspection. fitted to each end permitting visual inspection of the tube surface condition through an integral Drying is carried out by blowing dry observation window which houses a calibrated nitrogen through the tube bores after which hygrometer. The unit and tubes are purged and a whole surface leak test is performed. The filled with dry nitrogen, lifted (Fig. 12) and elements are enclosed in a large box (Fig 10) transferred through the air lock (Fig.13) onto which is provided with helium gas under slight transportation for delivery to site. pressure. Gas samples are taken from each tube flow path under vacuum conditions and A nitrogen purge gas unit is despatched with analysed for traces of helium using a mass each boiler unit to make good for any nitrogen spectrometer. The element is given a final leakage and thus adjust the internal relative clean followed by inspection and then humidity should this be necessary. Fig. 11 provides transferred via a bogie to the clean a view of a unit leaving the factory. conditions area. The reheater section is assembled by building 9. ASSEMBLY OF BOILER UNITS the containment casing and loading it with 36 stainless steel elements separated at mid point by The objective of keeping site work to a an acoustic baffle plate. The elements are minimum requires assembly of the boiler units positioned, tied and locked as described earlier in permanent workshops large enough to handle for the main unit. Two stainless steel headers completed units and under appropriate clean are set up relative to the reheater elements and conditions. also lined up radially to match the shield wall penetration to which they are eventually joined. The facility at Gateshead is 20 metres wide by 18 metres high by 130 metres long and Assembly continues by welding 1tH tailpipes is equipped with two 50 ton overhead non drip between the elements and stubs on the headers. cranes. The cranes can be synchronised to Each weld which is manually made by the T.I.C. lift up to 100 tons with control from one process is ultrasonically examined and leak tested crane only. using helium mass spectrometry. A general view of the assembly area is shown in Fig. 15. The shop has a sealed floor and is internally lined to maintain clean conditions. The completed unit is upended for transportation Air conditioning plant provides a slightly provided with temporary covers and purged in a pressurised heated atmosphere within the similar manner to the Main Unit for despatch to site area to prevent ingress of dirt and moisture (Figs. 16 & IT). and to limit relative humidity to less than 50%. In addition to the element transfer 10. ASSOCIATED COMPONENTS roller shutter door, a large steel panelled air lock is provided with removable roof Various components which connect or support the section to allow entry of large components heat exchangers are being produced within the N.E.I. e.g. casing plates and the despatch of group and at other companies. completed Boiler Units. Man access is controlled through a manned security gate Penetrations associated with the Economiser Feed, and is provided through changing rooms and Decay Heat, Superheater, Reheater and Instrumentation mess rooms which adjoin the shop. are being manufactured to standards and controls already described above. Automated welding techniques Assembly begins by setting out one are utilised wherever possible and stringent casing side in the horizontal position together inspection criteria are being complied with. with the main tube bank support members to form a jig frame. The elements are then The manufacture of the headers and pipework for transferred using a special handling frame reheater and superheater sections however currently over the jig and lowered onto the casing utilise a high degree of manual processes which structure and located in position by means require great attention to the training, qual- of permanent tube element support attach- ification and performance of individual craftsmen. ments and spacers to the casing structural Development work is, however progressing, aimed at using robot techniques for header manufacture pivot and beam supports and tied back at the and it is hoped to commence pilot production on upper end of the vessel wall. The reheater is other projects in the very near future. The attached to its slings and suspended from the possibility of producing headers within a flex- roof. ible manufacturing system (FMS) is also being investigated continuing an emphasis on inherently The sequence is repeated until all 12 units reliable mechanised production methods. are in position. After datum setting and align- ment of units,connections to the various penetr- 11. QUALITY ASSURANCE ations are made to complete the installation of the pressure part flow paths within the reactor The manufacture of the steam generators for vessel. the Heysham II/Torness A.G.R. programme demands The installation of the steam generators tne highest standards of Quality Assurance based within the primary circuit is completed by on B.S.5750 augmented by National Nuclear attachment of the annular ring, gas deflectors Corporation requirements G7622. and seals, cooling gas pipework, the quadrant division plates, guide restraints and removal A Quality Programme is defined which refers of debris screen protection covers. to the NEI Nuclear Systems Quality Manual together with associated documentation in the In parallel with these activities and form of quality procedures, manufacturing and external to the reactor pressure vessel,penetr- process specifications, welding specifications ation connections are completed and integral and non destructive test procedures. header, pipework, valves, mountings and instru- mentation are erected and connected to complete The programme embodies the familiar eighteen the boiler circuits. principles found also in the top quality stand- ards applied throughout the world to nuclear 13- CURRENT SITUATION power plant. This spans from bought out materials and sub-contracted components through in-house Production of the Steam Generators at manufacture to documentation control. Gateshead commenced in earnest at the beginning of 1983. To date, 36 of the 18 units (includ- Quality Audits are now commonplace to demon- ing reheater) have or are ready to be despatched strate not only that the quality programme conf- to site. Heysham II will very shortly have all orms to the required standards but also that it units for both reactors available. Work to is being faithfully implemented throughout all of complete delivery to Torness for the second the manufacturers involved. reactor is well underway and the Last Unit is on schedule for delivery late September/early A.G.R. steam generators are thus produced October 1981. to quality standards as rigorous as those any - where in the world. Installation of all units for the first reactors is complete at both Heysham and Torness 12. SITE CONSTRUCTION and work is proceeding to complete the circuits. The boiler units are off-loaded upon receipt 1H. CONCLUSION at site by a mobile transportation vehicle (shown in Fig.18) which removes the unit to covered The intent of this paper has been to storage. The dry nitrogen atmosphere is maint- provide a broad view of the evolution of gas ained during this period and appropriate atmos- cooled reactor steam generators. pheric control is maintained in the storage area. Temporary debris screen protection covers are The production facilities have been fitted at the superheater end. described and the robotic welding installation has been emphasised. When required the units are removed from the store and transported to the reactor building. The contracts are continuing to programme Covers are removed and lifting tackle secured to both in terms of progress and cost. the casing structure. Each main unit is upended from the horizontal position and lifted over the It is clear that N.E.I. Nuclear Systems top of the concrete reactor pressure vessel to a have made substantial investment in new plant, slot above the annulus around the reactor core. in software, in utilising the skills of people Fig.19 shows a unit which has just been lifted and in the management of change to advanced over the reactor vessel. manufacturing processes. The unit is then lowered (Fig.2.0) through the That equipment is obviously available for slot in the roof into the annulus and onto a a continuing programme of A.G.R.'s if required, hand propelled trolley. The associated reheater or also for any successions to later Magnox unit is lifted, and placed on top of the main steam generators where components are similar unit and temporarily bolted to it. The entire to A.G.R. assembly is traversed around the annulus to its final position. However, the Engineering skills and Manufacturing practices illustrated in this The main boiler unit is then jacked up and paper are relevant to other systems and transferred from the trolley to its permanent - 5 - ffl indeed to other industrial projects to which they could be adapted in line with customer requirements. ACKNOWLEDGEMENTS The authors wish to thank the Management of N.E.I. Nuclear Systems for permission to produce this paper and to those colleagues who provided assistance in the preparation of the paper. - 6 - NEl A.G.R. STEAM GENERATORS aftmn co»tnuCT». KVMAM • * TOTMM NEl UOt •nnOM MU -1 M A.G.R. TUBE BUTT WELDING UKATOt MAMLUL MlDt t*O* MAM tOUM UWT CaWIM* • M CLCMCKTt * «4 MC&T t^ MHUTM / 43,94* 18,0* » TMJ» FON Tt* two ram IT*TWN« Wf H*M TO KUWMKN • t.Ml UAM HU • 1.111 MCAT MA CVAMMMTOA SO.65S 12.672 tetmmitx or • n» tui e# T fig.1a AGR Production Statistics Fig.1b AGR Production Statistics MEl A.G.FL SPACER WELDING ••HMTn / tUHIIHUTn 2,108,018 EVAPORATOR •COHOMtU / HCU MAT 1,030,858 Fig.ic AGR Production Statistics t'ig.2 Tube Handling System Fig.3 Fixed Head TIG Welding. Fig.4 CNC Bending Machine _ 7 - Fig.6 General View of Robot Installation Fig. 5 . Grbita.l TIG Welding Fig.7 Robot Gantry Welding 9%Cr platen Fig.8 Furnace Loading Hearths Fig.9 Hydraulic Test Facility Fig.10 Helium Leak Testing Box fig. 11a Main Unit assembly in c.l?nn aren. Fig.11b Main Unit assembly in clean area. Fig.12 Main Unit being lifted. Fig.13 Main Unit being lowered into Airlock. Fig.1t [-rain Unit leaving Factory - 9 - No. 36 SULZER II III III III II11II XA0055846 IAEA Specialist's Meeting on Heat Exchanging Components of Gas- Cooled Reactors Dvisseldorf, 16 - 19 April 1984 THE USE OF BIMETALLIC WELDS IN THE THTR STEAM GENERATORS U. Blumer, H. Fricker, S. Amacker, SULZER BROTHERS Ltd. Winterthur 1. Introduction Heat exchanger tubes that operate in the elevated temjterature re- gion often have to be designed in two qualities of material. The part which has relatively low tube wall temperatures can be de- signed with the use of a ferritic material for economical rea- sons. At a certain temperature level of the wall however, the creep strength and the stability of this material are no longer sufficient, and an austenitic tube material has to be provided for the higher temperature section of the heat exchanger bundle. This paper deals with welds between the two tubing sections, with emphasis on their application in the Thorium High Temperature Re- actor (THTR) steam generators (Fig. 1). While the tubing of heat exchanger equipment in general needs ca- reful design to withstand a number of different loading types, the use of bimetallic welds (BMW) requires special attention to prevent it from becoming a weak spot in the design. 2. Description of weld Fig. 2 shows such a weld. The tube dimensions are 25 mm OD and 3.2 mm wall thickness on both sides. The ferritic side is made of the material 10 Cr Mo 910, a steel containing 21/4 % chromium SULZER - - - and 1 % molybdenuin. The austenitic side consists of the alloy 800, a steel with 32 % nickel, 20% chromium and additions of Ti and Al. The welding is accomplished by TIG using a weld fil- ler metal in the form of an insert ring. This intermediate ma- terial is a nickelbase alloy with 20 % chromium,with molybdenum and Niobium. The high nickel content of this intermediate layer acts as a barrier for carbon migration and therefore prevents the decarburization of the ferritic tube. The welds were heat-treated at 730 °C in order to reduce the hardness of the ferrite and the residual welding stresses. 3. Locating the weld For the location of the weld within the heat exchanging surface, there are reasons to go to higher temperatures and others for lower temperature. The following considerations show the deter- mination of an optimum temperature level for the selection of the proper position of the weld. Fig. 2 shows the distribution of the wall temperatures along the tube length for minimum (40 %) and maximum (100 %) power level. Reasons for selecting higher temperature are: - saving material costs (non technical) - the BMW and the following austenitic tubes should not experi- ence wetting.by evaporating water, as this might cause inter- cristalline stress corrosion cracking in the alloy 800. There- fore the location should be well above the vaporization tempe- rature. SULZER - 3 - However, reasons for upper limits of wall temperatures are: - the creep strength of the ferritic material. - diffusion of carbon from ferritic to austenitic material, which could cause a weakened decarbonized zone in the ferrite near the weld. - the strain range due to the different coefficients of expansion increase with the temperature range that the weld experiences. This might affect fatigue life in addition to the generally re- duced fatigue strength at higher temperatures. The distance from the evaporation points is governed by the full load condition, as can be seen in fig. 3. It should be 29 °C above evaporation temperature. At the chosen location a maximum wall temperature of 526 °C can be reached at 40 % load. This inculdes 31 °C for deviation from the ideal, calculated condition. This temperature is well below the 550 C which we consider as an upper limit. This comes from experiences that have been gained from failures in the heat-affec- ted zone of the ferrite above this temperature, as well as from carbon migration measurements. 4. Tests with THTR welds Corrosion tests have been performed in order to check the sus- ceptibility against stress corrosion cracking. Under the applied stress of 2/3 yield stress at 1 % strain, we did not find any ruptures. In addition, fatigue tests have been performed that combine ther- mal shock with temperature load cycling. The thermal shock stress is caused by a radial temperature gradient in the tube wall, that SULZER - 4 - arises from rapidly cooling the tube inner surface. The tempera- ture level cycling from ambient temperature to 550 °C and back causes stresses that have their origin in the different thermal expansion of the two joining materials. The following two types of thermal fatigue tests have been per- formed {fig. 4): The first test is a long duration cycling with one hour hold time at 550 C. The thermal gradient across the wall, which is produ- ced by air cooling after the holdtime, causes a calculated gra- dient stress of only 54 MPa. Together with the strain range due to the temperature level cycling, this loading did not have a fai- lure in 23'000 cycles totalling 35'000 hours of test duration. The second fatigue tests are short term tests where the specimens are heated to only 500 C and water quenched. Figure 5 shows the induction coil configuration for the rapid homogeneous heating of the weld region. 36*000 cycles were performed without producing any cracks on the inner or outer tube surface. Fig. 6 shows the quenching temperatures as a function of time and the maximum stress that occurs after about 0.5 sec. The comparison of the cal- culated superimposed stress with fatigue curves showed that the allowable stresses were exceeded by a factor of 2 and the allow- able cycles by a factor of 15, respectively. As these factors are about equal to the inherent safety factors in the fatigue curves, we can draw the conclusions that the weld behaves as well as the parent material and can be evaluated for fatigue damage by the same curves. 5. Stress analysis for the licensing procedure In the real surrounding of the THTR steam generators the region of the tube bimetallic weld is subject to a number of loadings that have to be considered in the design and the analysis. Fig. 7 SULZER S shows the different kinds of loadings that affect the behaviour of the weld, in the sequence of their importance on a basis of calculated stress. In the following, the evaluation of the stresses for the diffe- rent loadings is presented in more detail. Only the conditions at full load are adressed here for simplicity, although for a complete analysis all relevant operational states (including transient) have to be considered. 5.1 Thermal gradients The thermal gradients that are caused by the heat transfer through the tube walls, have the following values which include a nonuni- form gas side heat transfer around the circumference: 36 on the ferrite side and 56 on the austenitic side. The maximum occuring stresses on the inside of the tube are 73 and 161 MPa respectively. The higher expansion coefficient and the lower conductivity of the alloy 800 lead to that much higher stresses. 5.2 Bimetallic thermal stress The calculation of the stresses due to the different thermal ex- pansion, assuming constant temperature, is a considerable task, that has been attempted in different ways in the past. Among the main problems is the lack of knowledge about the distribution of the material properties in the weld region. There are three different basic materials, from the two joining tubes and the weld filler metal. Through welding, intermediate alloys are formed, the properties of which are poorly defined for detailed analysis. SULZER ~ 6 - Finite element analyses often have the drawback that they show singularities at material interfaces and the results are there- fore dependent on mesh refinement. In view of these difficulties we consider a simple hand calcu- lation approach as a sufficient and as the most efficient way to arrive at the proper order of magnitude of the stresses. We used thin shell analysis assuming a sudden transition bet- ween the parent metals. This is a conservative assumption as the filler metal has properties that are approximately in bet- ween. The resulting maximum stresses amount to 115 MPa. 5.3 Thermal tube bending In ref. (1), the calculation of thermal bending moments in tubes of helical bundles has been explained and shall not be repeated here. Due to different temperatures of bundle tubes and support plates as well as due to nonlinear thermal expansion of the tubes as a function of bundle height, bending moments are exerted on the tubes. Fig. 8 shows an additional detailed analysis model that was used for the assessment of the maximum bending moment. The calculated stress due to thermal bending at full load was calculated to be 69 MPa. However, special consideration was given to the fact, that tubes can show a phenomenon called "elastic-follow-up" which results in strain accumulation at the support points. An inelastic ana- lysis of a pipe segment has shown, that the inelastic strain accu- mulation can be assessed by multiplying the elastic strain by a factor of 2.6 for this geometry, material and temperature level, and in the absence of other loadings. : •.. •• J99- SULZER 7 5.4 Internal pressure, weight and earthquake loadings The stress calculations for these loadings is straight forward and needs no further comments. The maximum pressure stress is 56 MPa on the inside and 39 MPa on the outside. The loading of weight is counteracted to a large amount by the flow forces of the helium flowing upwards. In the THTR steam generators the earthquake loading is negligi- ble for the bundle tubes. 5.5 Combination of stresses and assessment All the stresses from the mentioned loadings have to be super- imposed to form stress range intensities for the important ope- rational cycles. For the cycle from full load to cold shutdown addressed here the maximum stress intensities amount to 264 MPa on the ferritic and 251 MPa on the austenitic side. Both stresses occur on the tube inner surface and can be classified as localized peak values that have to be assessed with creep-fatigue damage criteria. In this case, the linear damage rule of ref. (2) has been applied, which states that damage can be separated in two different damage por- tions. The first portion, the so called fatigue damage, is found to be negligibly small, that is less than 1 % on both sides of the weld. The second portion, the creep damage is assessed with integration of damage during relaxation of the thermal stresses, as it is also shown in Ref. (1). The creep damage only is important on the fer- ritic side and was calculated to be 14 % of the allowable damage. SULZER 8 Calculations have also been done to demonstrate the absence of ratchetting, which means the accumulation of strain due to repea- ted cyclic loadings. The elastic-follow-up effects mentioned ear- lier have been taken into account to show the adequacy of the de- sign. 6. Conclusions We have tried to assess the behaviour of the weld by the follo- wing different approaches: - design considerations towards proper operational temperature location - tests to show the adequacy in corrosion and fatigue behaviour - detailed analysis of the combined effect of all loadings and operational cycles. The results of this work have shown an adequate behaviour of these bi- metallic welds in the THTR steam generators. References "Analysis of heat exchanger bundles operated at eleva- ted temperatures", U. Blumer, Specialist's Meeting on Process Heat Applications Technology, IAEA, 1979. ASME Boiler and Pressure Vessel Code, Code Case N-47-12 SULZER Schnitt durch eine Dampferzeugereinheit zum 300-MWe-THTR-Prototyp-Kernkraftwerk Schmehausen, BRD 1 Äußerer Abschluß 2 Innerer Abschluß 3 Zwischenüberhitzer-Dampfaustritt 535°C, 49bar 4 Dehnzone 5 Hochdruckdampf-Austritt 550 °Cr 186 bar 6 Spannbetonbehälter-Durchdringung 7 Zwischenüberhitzer-Dampfeintritt 365 °C 8 Speisewassereintritt 180 °C 9 Kernrohr 10 Heliumaustritt 250 °C 11 HD-I-Bündel 12 Dampferzeugerhemd 13 HD-Il-Bündel 14 Zwischenüberhitzer-Bündel 15 Heliumeintritt 750 °C Sectional view of a steam generator unit for the 300-MWeTHTR prototype nuclear power station Schmehausen, FRG 11 1 Outer closure 2 Inner closure fe 3 Reheater steam outlet 535 °C, 49 bar 4 Expansion modulus 5 HP steam outlet 550°C, 186 bar 6 Prestressed concrete vessel penetration -12 7 Reheater steam inlet 365 °C 8 Feed-water inlet 180 °C 9 Central column 10 Helium outlet 250°C 11 HP-l-bundle 1m 12 Outer shroud 13 HP-ll-bundle 14 Reheater bundle 15 Helium inlet 750°C -13 Section d'un générateur de vapeur pour la centrale nucléaire prototype de 300 MWe THTR de Schmehausen, RFA 1 Fond extérieur 2 Fond intérieur 3 Sortie de la vapeur du resurchauffeur 535 °C, 49 bars 4 Zone de souplesse 5 Sortie de la vapeur H P 550 °C, 186 bars 6 Pénétration du caisson en béton précontraint 7 Entrée de la vapeur du resurchauffeur 365 °C 8 Entrée de l'eau d'alimentation 180 °C 9 Tube central 10 Sortie d'hélium 250°C 11 Faisceau HPI 12 Chemise extérieure 13 Faisceau HP II i ' 14 Faisceau resurchauffeur 15 Entrée d'hélium 750 °C JV> SULZER Makroschliff Macrographie »1 \ /• • KZ] Welding material Mikroschliff x100 Alloy 800 Micrographie 21/4 Cr 1 Mo Welding material Figure 2 : Section of weld SULZER 5LM Tube Length Fig 3 Nominal Wall Temperature Profiles Test I Test II - upper temperature 550 °C - upper temperature 500 °C - heating by furnace - heating by induction coil - air cooling - water cooling - holdtimes totalling - no hold times 23'000 hours - 23'000 cycles - 36'000 cycles - no cracks detected - no cracks detected Fig. 4: Fatigue tests Figure 3 Water Cooled Heater Coil Adapted to the different materials in order to provide an equal heat up period. SY3 5ULZER SLM T [XJ [N/m1] Stress Tube mid wall temp •new cycle Tube inside temp. 0 2 3 4 5 6 7 10 tfsj Time Fig. 6 Temperatures and Stresses during thermal shock test - Thermal gradient across wall - Bimetallic thermal stress - Thermal tube bending - Internal pressure - Tube bending by weight and gas flow - Tube bending by earthquake Fig. 7: Loadings to be considered SULZER fixed sipport plate floating plate FJQ.8 Tube model for analysis of bimetallic weld bending load No. 37 1 Gas Metal Arc Narrow-Gap Welding of Pressure Vessels Made from the Nickel Alloy 2.4663 Ill XA0055847 by K. Iversen and A. Palussek 1. Introduction The highly heat-resistant nickel alloy 2.4663 is used for the construction of test components of the nuclear heat reactor PUP as structural material. This material is given preference for strength reasons in particular in the range of peak temperatures of 950 °C at pressures up to 40 bar. Since no construction and operation experience is yet available with primary components for the process heat reactor, test components shall be developed, manu- factured, and tested. These works are sponsered by the German Minister of Virtschaft, Mittelstand und Verkehr of the state of Nordrhein-Westfalen. With the helium intermediate heat exchanger, two 10 MW types come under consideration, these being the helical tube and the straight tube versions. The hot gas collector component part places the highest demands on the wel- ding and testing technology. Workpieces of 1000 mm diameter and wall thicknesses of 42 to 100 mm are to be forged from material 2.4663, to be joined together, to be nondestructively tested and to be tested in a largescale test plant under operating conditions. 2. Design and materials of a hot gas collector test model Before the construction of the two 10 MW heat ex- changers, a hot gas collector model shortened in the longitudinal direction with original wall thicknesses and diameters was manufactured amongst other things for the development of the manufacturing technology, in order to gather in good time experience for the construction of the components. Figure 1 shows the design of the hot gas collector model. Three welds are to be made in the wall thickness range between 35 - 70 mm. The nickel alloy 2.4663 was selected as base material and filler wire of the same composition 1.2 mm diameter was chosen. Table 1 shows the chemical analyses of base metal, filler wire and welded joint. All figures had been within the standards. The burn-out of the elements aluminum and titanium amounted to 0,19 %• Therefore limitations of titanium and aluminum to £1,5 % from the point of embrittlement of the deposited metal is not necessary according to the latest experiences. Maximum values of Ti + Al ~ 1#9 % (Al max 1.3 % and Ti max 0#6 %) were permitted. 3. Welding and testing the hot gas collector model 3.1 Selection of the suitable welding process for the manufacture_of_the_circular_welds The manufacture of pressure vessels for the reactor con- struction necessitates the use of proven reliable welding procedures. However, with the selection of the welding procedure neither tested electrodes nor powder for sub- merged arc have been available for the material 2.4663* Experience had been gained only for the manual arc wel- ding processes TIG and GMA each with argon as shielding gas. Other welding procedures with high deposition rates like electro-slag and submerged-arc process could not be used by means of their hot cracking sensivity caused by the high heat input. On the other hand efficieny must be an extremely important point of view if new components are to be developed. Therefore the decision was made to chose narrow gap welding with an inert gas. With expensive base metals and filler wire as well as high labour costs, the gap width and the side wall angle effect decisively the quantity of the weld metal to be filled in, the melting rate and the manufacturing costs. This applies in particular with larger wall thicknesses, Fig. 2. According to experience made so far with the nickel alloy 2.4663, however, the intermediate layer temperature should be limited to 150 °C, the weld pool size should be small and overheating of the weld puddle should be avoided. This led to the selection of a GKA process with a melting rate of maximum 3-5 kg/h. Therefore orLJ-y single wire processes came into consideration. For reasons including experience available with pressure vessels in Japan, the Babcock-Hitachi oscillation flap process was chosen /~1_7. Fig. 3 and 4 shows the method of process operation C^.J\ f^bj* 3.2 Qualification 2£_i^£_£§§_S®5ai_§£c_Ba£r!2HzS§2 E£°£eSS Unfortunately no experience was available for the pro- duction of circular welds on forged shells of the nickel alloy 2.4663 with the above-mentioned narrow-gap process. The same applied for stainless Cr-Ni steels too. There- fore suitability for welding had to be proven and the first parameters had to be found with plate samples of s » 60 mm by fundamental research at the ISF in Aachen/~4^7- Here it was demonstrated that neither the commercially used gas mixture of 82 % Ar/18 % COp nor pure argon came into consideration because of the high arc length and the too small side wall penetration, Fig. 5« Only the use of pure helium (99*995 °/°) in combination with the pulse technique led to a short arc and to a low risk of burn- back as well as to reduced spattering in the 9 - 11 mm wide gap. The side wall penetration could be increased decisively from approx. 0.5 nua to>1 mm and produced sound welds, Fig. 6. The low density of the helium had a disadvantageous effect in the narrow gap because of amongst other things the formation of very adhesive titanium, aluminum and chromium oxides on the layer surfaces which could only be removed by grinding. At the welding equipment itself, the wire feed system above all had to be adapted to the significantly stiffer nickel wire (2.-4-663). The tests in the ISF in Aachen were concluded successfully after the above-mentioned improve- ments C^l' The plate samples had to be provided with an angle of 18° out of the flat for the start of welding by means of shrinkage. The continuation of the tests at Interatom took place initially on circular seams 1000 mm o.d. x 125 mm wall thickness of 2.4-663. For this the ISF welding data were taken over and the welding system was completed appro- priately, Fig. 7. However, because of the impeded shrinking, heat cracking occurred in the center of the weld. Only by reducing the layer height, the heat absorption, the weld pool overheating and the weld pool size welds could be done successfully without cracks. Before the optimum data were achieved, single pores and lack of fusion were found by means of radiographic and ultrasonic testing. The first circular weld was followed by the german official TuY process inspection with the same dimensions. The root path had been welded manually from the inside while fillet and the final runs had been welded from the outside automatically by the narrow gap G-MA process. For the manual weld Nicrofer S 5520 of 2,4- mm 0 was used while the automatic GMA process consumed Nicrofer S 5520 wire of 1,2 mm 0 diameter. All welds had been done in the horizontal position while the workpiece was turned and the torch was stable. The opening of the weld preparation was app. 2° to avoid the squeezing of weld head which was just 7 mm in width by means of transverse shrinkage, Fig. 8. The opening of the root srf gap on the surface was app. 17,8 mm before welding and lowered down according picture 8 to app. 14 mm before welding the last path, Fig. 9. The narrow gap welding with the Babcock-Hitachi-Process is capable of one path per layer for a gap of 9 - 15 mm and two pathes per layer till 18 mm. During this inspection, influences on the welding results such as not grinding the bead surfaces, welding over of start and stop points, working out and filling in local repair places in seam depth and seam width were also tried. It was shown that welding over ungrinded layers led to cracks and lack of fusion. All filling layers and the repair simulations met the required values for BS according to DIN 8563> Part J. For tbis, however, all parameters had to be matched exactly to one another and kept within a close tolerance band. The computer-assisted welding data monitoring system was here of outstanding benefit. After completion of the test weld (81 layers) it was x-rayed by a linear accelerator. An evaluation of the films did not lead to a conclusion about weld quality. It must be stated that good results can be obtained til 80 mm wall thickness only. The evaluation of the ultrasonic testing, manually as well as automatically, Fig. 10, did not show any defect outside tolerance band. The dye check did not show any defects too. According to the fixed plan of specimen the forged ring had to be split up. Six cross sections had to be examined by TtJV while another 39 had been studied by INTERATOM. Macro sections did not show any defects without some minor pores, Fig. 11. Micro sections (v » 50 or 100:1) did show layers according to speci- fication but with the exception of some lower area, where the layers had not been grinded. In that ungrinded and unsound area lack of fusion as well as porosity had to "be stated. Most of these defects had been detected already during x-ray and ultrasonic testing. The figures of the other testing procedures are given in table 2. - £ound_tensile S2ecimen_according_to_DIN_50125 The requirements for the base metal values of Em 2" 700 N/mm2 and Ep 0,2 - 300 N/mm2 were met extremely close because the lowest value was 302 N/mm . On the other hand the values of the deposited metal (4-23 N/mm ) as well as the welded Joint (486 N/mm ) exceeded the specification by far. The requirements for the yield point had been exceeded with Ep 0,2 = 328 N/mm but the figures for the ulti- mate tensile strength did not meet the requirements. The rupture had been secured in the base metal beyond the joint. Therefore the weld had met the requirements. former - 0 : 3 x wall thickness The following specimen had been taken: " out of a repaired section bend test specimen had been taken two for root bend testing and another two for final layer bend test * four bend test specimen (two of each for root bend test and final layer bend test) had been taken from all around the weld. All specimen did show 180 bend angle without any cracks. 7 notched tar_impact_test_according to_DIN 50115 The specimen had been taken from a place of repair and from other areas all around the joint. The requirements of - 40 J was met by all specimen. The lowest value was shown with 71 J at the weld edge while the highest was shown with 151 J in the deposited metal. It can be emphasized that narrow-gap-welds do show much better results than there is the demand according to HP 2/1. With these results the process qualification for the narrow .gap welding with the single wire snake-wave process (VP 70) was successfully finished. According to Fig. 1 three circular seams of different wall thicknesses (35? 42 and 70 mm) had to be produced. Narrow-gap welding took place with the welding data proven in the above-mentioned process inspection. About JO layers were required for the 70 mm wall thickness, this meant an average layer height of 2.*? mm. The heat input was 14 kJ/cm, the welding speed 30 cm/min and the melting rate app.3»8 kg/h. Radiographic inspections were carried out after about 4-0 mm of weld metal and after laying the final layer (70 mm). The results of the radiation testing and the results of the manual as well as computerassisted mechanized ultrasonic testing, Fig. 9 did show good harmony. With the exception of some single pores neither lack of fusion nor cracks had been detected. After finishing of the welds (T) and (2) the workpiece had to be oval shaped followed by welding (seam (^) ) to the bottom, figure 12. Fig. 13 shows the figures of transverse and longitudinal shrinkage depending on wall thickness. 8 4. Summary The GMA narrow-gap welding process with helium as shielding gas showed circular welds free of defects up to s = 126 mm with the nickel alloy 2.4663. The commer- cially available pulse power source had to be modified at the shielding gas supply and the wire feed. The selection of the welding parameters had particular significance. It was shown that with shrinking-impeded welding of the circular welds, an overheated weld pool must absolutely be avoided and that a small weld pool size must be main- tained. The nickel alloy 2.4663 which contains aluminum and titanium necessitates grinding of the single layer surfaces. However, it remains for the future to show with further applications if a technical breakthrough of the narrow- gap welding process with the manufacture of thick-walled pressure vessels of stainless Cr-Ni steels and nickel alloys can been made. 5. References /""1J7 S.Swada, K.Hori, M.Kawahara, M.Takao, I.Asano; Application of Narrow-Gap Welding Process; AVS 60th Annual Meeting, 5 April 1979 /"~3 7 K. Iversen; Review of the process of narrow-gap welding; Colloquium narrow-gap welding SLV Duisburg, June 1982 /~4_7 P. Eichhorn, P.Groger; Gas metal arc narrow-gap welding tests on austenitic chrome-nickel steels and nickel based alloys; DVS report Vol. 75, 1982 £3 p 13 Z3 3 Z3 p —t- —f- —s D O cr P- S. I ro T3 fD c 3 a> CL CL U3 O TD rx> fD 3 CL UD O rV fD 40 CL CL P. "8 o P 3 3 3 D P. X ^< o o o o O O o Q o o o o o» oo •r> ao *^ balance balance balance balance ro ro IO ro __k k eu o o UJ O* >O o o oo J3O oo k 00 en UJ ° o O -* _-* o o UJ UJ •F- OS OS '— o o o o o c> —J ON >o en o, "ro o IA IA o o o 3: o o Vj 13 en o tA IA o O o en ro Vj x k UJ 2$ r§ oo en o o o 0,00 4 0,00 3 0,00 2 IA 1A O o 1 en o o IA IA nn to en en ft) o IA IA O o o o en1 en en o IA ~g •8 GD en "r8o 0,0 1 0,0 4 tA IA O o c » IA IA OO en "en en RpO,2 Rp 1.0 Rm A5 Z Av [|] Av [|] Bending 2 2 < [N/mrf] [N/mm ] [N/mm ] [%] [%] S PrU/P0-U [°] Requirement- ^300 — ^700 ^35 — >40 >40 180 — Base Mefal 302-323 — 710-742 35-52 — 105-108/1150°C — B.M. N125mm HeafNo.22541 316 — 723 42 — 107 — B.M. Depositedfiller 443-464 — 762-775 42-44 51-53 134-143 — — B.M. HeafNo.97097 — 01,2 453 768 43 52 139 — — Weld Tensill-est 486-553 510-581 753-792 43-47 46-51 — — — Weld DIN50125 523 549 776 45 48 — — — Weld — — Tensiltesf- 328-336 356-376 630-667 — — — Weld DIN 50120 332 364 649 — — — — — Weld joi n Impact-Tesf — — — — — 118-151 71-100 — DIN50115 — — — — — 137 82 — Bending- all 12 Specimen 3xf Welde d Tesf Bending Mandre DIN 50121 180° Properties of Base Metal, Filler Metal and Welded Joints (Narrow Gap ) 1350 1020 weld (1 Csl [7/7 // / \ \ \ \ _.\i \ \ \ \ \ 1 <./// // / & O CO \ \ \ \r\ \ \ \ \ zzzzzzzi> ^ Fig. Hot gas collector T3 6 Y=[b-S+2sZ^H)x length 20 40 50 30 100 120 140 160 *80 200 220 thickness [mm] Fig: 2 Seam volumes as a function of wall thickness and aoerature anale We ! Fig.3 Process with oscillating flap - a) Flap plate b) Feed rollers c) Contact nozzle d) Upper gas protection e) Welding gap f) Wire electrode g) Water cooling h) Front gas protection nozzle i) Rear gas protection nozzle input vD wire feed speed 8 wire deflection fp oscillation frequency output X wave length Y oscillation Wire. Def6rsn/r?3 •Jff Fig. 5 Penetration depending upon type of shielding gas Base metal: X 5 CrNi 18 9 Filler Metal: X 5 CrNi 18 9 The first 6 layers were welded with argon, the following layers with helium 1 Helium • I * t jjj I > > Fig. 6 Influence of the shielding gas on the arc length and side wall penetration Fig. "$- G'tA narrow gap welding equipment 81 pathes MIG -Narrow Gap mTIG (2pathes) Wsld preparation for MIG-Narrow-Gap Groove width on vesselsurface [mm] 50 60 70 80 125 Welded thickness [mm] Shrinkage of MIG-(Narrow- Gap)-Welding Fig 1o Fully mechanized ultra-sonic testing of the hot gas collector with computer evaluation Fig. 44 Narrow gap weld (wall thickness 12^ nun) Fig 72, Hot <3as collector with 3 circular welds Transverse shrinkage Longitudinal shrinkage 125 [mm] wall thickness a Longitudinal and transverse shrinkage depending on wall thickness A XA0055848 FORGED HOLLOWS (ALLOY 617) for PNP-Hot gas collectors by F. Hofmann VEREINIGTE DEUTSCHE METALLWERKE AG Geschaftsbereich Nickel-Technologie D-5980 Werdohl / West Germany - 1 - A VDMi Introduction and Purpose When the partners of the PNP-project decided to manufacture PNP-components, such as hot gas collectors, from material of the type "alloy 617" (DIN material No. 2.4663) the problem was encountered that the required semi-fabricated products, especially forged hollows weighing several tons each, were not available. Also at that time it was not known, whether products in this high alloyed high-temperature material could even be produced in the required dimensions. As VDM had already gathered experience in the production of other semi-fabricated products of this alloy, attempts were made based on this know- ledge to develop manufacturing methods for forged hollows. The aim was to produce hollows as long as possible, to keep costly welding to a minimum. Welded seams are always critical, during fabrication, as well as on later inspection under actual operating conditions. On the other hand, of course, the economics of the production method had to be kept in mind in reaching the goal of this development. A three stage plan illustrates the development aims, whereby stage 3 is currently being worked at (figure I). The first two stages involved the production of forged hollows for hot gas collectors for 10 MW heatexchangers designed by Steinmliller/ Sulzer and Balcke-Dlirr. Stage 3 encompasses the development of necessary forgings for a hot gas collector for a future 125 MW heatexchanger. As, according to current thinking, this entails approaching the limits of what is technically feasible and as such involves a high economical risk. This project is subsi- dized as part of the overall PNP-project. A Procedure A prerequisite for this project were the previously started investigations studying the influence of the melting method on the most important properties of the alloy, as for example the creep rupture properties at high temperatures. These in- vestigations revealed, that the expensive method of melting and casting under vacuum (VIM), which is frequently stipulated for such material, is in fact not necessary. This alloy can be melted in an open electric arc-furnace, similar to highly alloyed stainless steels in large heats of for example 30 tons, and cast into ingots after a VOD treatment. The advantage of this method is that it facilitates the use of economical raw materials through corresponding metallur- gical treatments, as well as the production of large ingots. Large ingots, however, have to be remelted to minimize segre- gations in the ingot. These are caused by elements such as chromium and molybdenum which have a great tendency to segre- gate. For remelting, the electroslag remelt process (ESR) was chosen. Compared to other remelt methods, it is economical and offers certain advantages during subsequent fabrication. Figure 2 shows the relationship of creep rupture properties of alloy NICROFER 5520 Co (Alloy 617) to the respectively discussed melting methods. This diagram clearly reveals that the melting method selected for this project does not adversely affect the creep rupture properties of NICROFER 5520 Co (Alloy 617) parti- cularly at high temperatures. A heat which was melted and cast under vacuum was used as reference. Of the two casts investigated, it was found that the creep rupture properties were actually better above 850 °C for the VOD/ESR processed heat as compared to the material melted and cast under vacuum. A For stage 1 of the project, remelt ingots of approximately 750 mm in diameter with a weight between 6 to 7 t were produced according to the melting method described. After dressing, the ingots were heated up to the forging temperature, upset forged (figure 3), pierced with a mandrel (figure 4), and forged to final dimension over a mandrel in several forging steps (figure 5). Prior to the final forging operation, it is important to homogenize the forgings in order to reduce unavoidable microsegregations. The temperature during the last forging operation is of great siginificance as it determines the properties of the forged part lateron. Thus the forgings must be sufficiently and uniformly deformed to obtain a defined grain structure during the final thermal treatment (fig.6). This thermal treatment is carried out within the temperature range 1150° - 1200 °C. With this thermal treatment an average grain size of ASTM 0 could be attained. This avoids any problems during ultra- sonic testing of the forged hollows for interior defects. The in- spected forgings (figure 7) were then machined to final dimensions (figure 8). This production step also necessitated detailed investi- gations, as alloy NICROFER 5520 Co (Alloy 617) is very difficult to machine. For stage 2 of the developmental work, i.e., for the production of a forged hollow, approx. 2000 mm in length, a larger ingot was needed to obtain the desired final dimensions. In this case an ingot 850 mm in diameter and weighing approx.8,5 t, was chosen. Fabrication parameters, similar to those used during stage 1, could be applied. The increased occurance of microsegregations caused by the larger ingot diameter had to be compensated by suitable homogenizing treat- ment. On this occasion, it was noted, that the operating limits of a 7000 t forging press was reached at individual deformation stages. During final inspection the hollows were shown to meet the specified properties and were consequently further processed to hot gas collectors by the fabricators, (figure 9) A Currently work is persued for stage 3 of the developmental project which entails production of a forged hollow for a hot gas collector for a future 125 MW heatexchanger. At least in the area of critical temperature the hot gas collector should consist of one piece, i.e., it should be approx. 4000 mm in length. After reviewing the forging capacities of suitable equipment available in West Germany, it was determined that an ingot, 1000 mm in diameter and 2700 mm in length, weighing approx. 17 t was required. For this highly alloyed material, which is very prone to segregations as already mentioned, this is certainly a most unusual size, which probably has never been pro- duced in the past. During processing the remelted ingot unsuspected problems were en- countered during electroslag remelting. These, however, were overcome by adapting process parameters, involving for example the development of a suitable slag. Figure 10 shows the cast ingot with a diameter of 1000 mm x 2700 mm and a weight of 17 t. After dressing the surface, the ingot was heated up to forging temperature. The ingot had to be upset forged directly without a supporting fine grained shell. Upset forging was carried out with a forging power of 9000 t. In two stages the ingot could be reduced approx. 50 % in height by upset forging to 1300 mm. The ingot was forged back to its original shape and once more upset forged to 1300 mm. This rather costly procedure was deemed necessary to obtain an allround homogeneous deformation of the ingot structure, as well as a breaking up of the segregation zones for the subsequent homogenizing heat treatment. The ingot was then pierced with a 300 mm diameter mandrel. Before forging could be continued, the rough forging had to be dressed, to remove upset folds and cracks inside the hollow. During further processing of the forging, problems were en- countered with widening and stretching, as the available tools, i.e., the forging mandrels were breaking due to insufficient strength. This problem was overcome by forging in the upper forging temperature range, A - 5 - WDM, which necessitated frequent reheating. The final dimensions attained were o.d. 1050 mm and i.d. 780 mm with a length of 4400 mm (figure 11). In order to determine the thermal treatment procedure a sample ring approx. 300 mm in length was cut off. This sample ring was given a trial heat treatment in the annealing furnace foreseen to be used for the actual forged hollow. Results obtained so far, indicate a structure essentially free of segregations with an average grain size of ASTM 0. Utilizing these experimentally determined heat treatment parameters, final annealing of the actual forged hollow is currently beeing con- ducted. Future Outlook Following completion of the thermal treatment of the forged hollow, it is planned to take samples in various places, in order to examine the properties of the alloy in the existing product form. This should particularly show, whether production of forgings of that dimensions are technically and economically feasible in sophisticated nickel alloys such as NICROFER 5520 Co (Alloy 617) in future. This is of importance, not only for the PNP-project, as other applications also require at least the capability of producing large ingot sizes in similar alloys. Up to now this could not be considered to be the latest state of the art, due to the high production risks involved. As an extension to this development, it is planned to utilize the forged hollows, to study economic joining methods, such as narrow gap submerged arc welding. This additional work focuses particularly the question of economics. Initial experiments have been conducted and can be said to be promising. Abb.1 /\ ] Stufenplan zur Entwicklung geschmiedeter NOW /VDMV Hohlkorper aus dem Werkstoff Nicrofer 5520Co( alloy 617) Hohlkorperabmessung Blockgewicht Stufe 1: ca. 1090'740 0 x 1000 (mm) ca. 6,7 to Roh-Gew. ~ 5to Stufe 2 ca. 880/550 0 x2100(mm) ca. 8 to Roh Gew. ~ 6,5to Stufe3: ca. 1050/780 0 x 4400(mm) ca. 17 to Roh-Gew. ~ 16,6 to Fig. 1 Plan of developmental work 100 10- o o CO E en 700 800 900 1000 Temperatur in °C Fig. 2 Creep rupture strength of NICROFER 5520 Co (alloy 617) related on melting procedure ( 0 VIM; XAVOD/ESR) Fig . 3 Upsetting of the ingot 4 Piercing of the upset inqot Fig. 5 Widening of the pierced ingot Fig. 6 Heat treatment of the forged hollow 9- Fig. 7 Forged hollow before machining Fig. '8 Machined hollow ready for delivery A HeiBgassammler fiir U-Rohr Warmetauscher aus NQ W VDM Nicrofer 5520 Co(alloy 617) Konstruktion;Balcke-Diirr 820/630 0x1850 820/674 0x1650 Diese Abbildung setzt sich aus Teilstucke der Abmessungen: 820/620© x 265 zusammen Fig. ^ Hot gas collector for He/He heat exchanger Fabricator: Baicke-Dlirr C/f Fig. 40 17 t ESR ingot in NICROFER 5520 Co A geschmiedeter Hohlkorper aus Nicrofer 5520 Co Schrniedeabmessung: 1050/7800 x 4400mm NOW Fig. AA Forged hollow in NICROFER 5520 Co Dimension: o.d. 1050 mm i.d. 780 mm, length 4400 mml118°Fand PL + PB PRESSURE ONLY 6352 psi mt 1.33 LOWER 300Hrs = TUBESHEET 14,100 psi PL + PB + Q 25991 psi 0.63 3Sm,@1067°Fand 800H 300Hrs = PL + PB + O+ F 42,300 psi ALL TRANSIENTS 0.81 1.0 0.23 - NATURAL CIRCULATION SUPPRESSED IN