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Proceedings of the Workshop on Structural Design Criteria for HTR

Jiilich, 31. January - 1. February 1989

Editors: G. Breitbach F. Schubert H. Nickel

Jiil-Conf-71 April 1989 ISSN 0344-5798 Als Manuskript gedruckt

Berichte der Kernforschungsanlage Jülich - Jül-Conf-71

Zu beziehen durch: ZENTRALBIBLIOTHEK der Kernforschungsanlage Jülich GmbH Postfach 1913 • D-5170 Jülich (Bundesrepublik Deutschland) Telefon: 02461/610 • Telex: 833556-0 kf d Proceedings of the Workshop on Structural Design Criteria for HTR

Jiilich, 31. January - 1. February 1989

Editors: G. Breitbach F. Schubert H. Nickel Workshop on Structural Design Criteria for HTR

Introductural remarks

Most of the presentations given in this workshop are based on the German research and development project "HTR Design Criteria" carried out under the sponsorship of the Federal Ministry of Research and Technology. The main emphasis of this work was to acquire the fundamental principles and basic data for the establishment of German KTA-rules (KTA: Nuclear Safety Standards Commission) for the design of HTR-structural components.

The project began in 1984 and the research work divided among several working groups and task forces, with participation from several institutions and companies. The role of coordination has been carried out by the Institute for Reactor Materials, Nuclear Research Centre Julien, headed by Prof. Dr. H. Nickel.

The work has been organized into four working groups: a) Technical safety boundary conditions; b) Metallic structural components; c) Prestressed concrete pressure vessel; d) Graphitic structural components.

The required work in each group was divided between a number of task forces. The membership of each group and task force is given in the appendix. Foreign participants in the workshop had the opportunity to present the status of the HTR-related structural design code work being carried out in their own countries. Ill

Table of contents

Objective of the workshop

Objective of the Workshop on Structural Design Criteria for HTR; Survey of the Research Activities in the Federal Republic of Germany H. Nickel KFA-Jülich, Institut für Reaktorwerkstoffe

Section I: HTR Projects and status of licensing principles

Chairman; H. Schuster, KFA-Jüüch

Present Status of MHTGR program in USA 15 P.L Rittenhouse Oak Ridge National Laboratory, Oak Ridge

Present status of HTTR project in Japan 33 T. Tanaka, S. Saito Japan Atomic Energy Research Institute

HTR situation in China 48 D. Wang, S. Xu Institute of Nuclear Energy, Tsinghua University, Peking

Present status of HTR in FRG 59 H. Nickel KFA Jülich, Institut für Reaktorwerkstoffe

Section II: Technical safety boundary conditions

Chairman: R. Trumpfheller, Essen

HTR Safety Features and the Integrity Concept • ,. * 83 J. Wolters - - KFA Jülich, Institut für Nukleare Sicherheitsforschung . .-~ • IV

Classification of systems and components into safety classes and quality standard classes 97 M. Dette Rheinisch-Westfälischer Technischer Überwachungs-Verein e. V., Essen

Section III: Metallic high temperature components

Section 111.1

Chairman: H. Clausmeyer, MAN-GHH, Oberhausen

Metallurgical and physical fundamentals for the design of high temperature components 113 F. Schubert KFA Jülich, Institut für Reaktorwerkstoffe

Load levels, stresses, failure modes and design criteria 133 K. Bieniussa Gesellschaft für Reaktorsicherheit (GRS) mbH, Köln

Basic requirements relating to quality assurance of safety related HTR materials and components 159 J. Just Rheinisch Westfälischer Technischer Überwachungs-Verein e. V.,Essen

Non-destructive detection of flaws during manufacture and operation of components 171 F.Walte Fraunhofer-Institut für zerstörungsfreie Prüfverfahren, Saarbrücken

Section III.2

Chairman: W. Dahl, RWTH-Aachen

Material Data and Constitutive Equations 185 HJ. Penkalla KFA Jülich, Institut für Reaktorwerkstoffe

Methods for very high temperature design 206 J. J. Blass, J. M. Corum and S. J. Chang Oak Ridge National Laboratory, Oak Ridge Life fraction rules 228 K. Maile Staatliche Materialprüfungsanstalt, Stuttgart

The present status of research and development works for the preparation of the high temperature design code 243 Y.-Muto Japan Atomic Energy Research Institute (JAERI), Japan

Creep rupture characteristics in the HTGR simulated helium gas environment and their relevance to structural design 275 Y. Kurata, Y. Ogawa, H. Nakajima, T. Kondo Japan Atomic Research Institute (JAERI), Japan •

Section III.3

Chairman: H. P. Alder, PSI-WCirenlingen

Assessment of primary and secondary stresses for component design , 293 E. Bodmann Hochtemperatur-Reaktorbau GmbH, Mannheim

Elastic and inelastic analysis for component behaviour . . 309 H.-J. Seehafer Interatom GmbH, Bensberg

Significance of Fracture Mechanics 32g K. .Schneider - i ASEA Brown Boveri AG, Mannheim

Section IV: Reactor pressure vessels

Chairman: J. Altes, KFA-Jülich

Design criteria for prestressed concrete pressure vessels 34g K. Schimmelpfennig. Stangenberg, Schnellenbach & Partner, Bochum

Design criteria for liners of concrete vessels 370 R. Oberpichler ... ' •• ' Stangenberg, Schnellenbach & Partner, Bochum VI

Special features of the design of pressure vessel closures and heat inulations 385 J. Pschowski Hochtemperatur-Reaktorbau GmbH, Mannheim

The HTR-module pressure vessel unit; design criteria and safety philosophy 404 G. Neumann, Siemens AG, Unternehmensbereich KWU, Erlangen K. Dumm, Interatom GmbH, Bensberg

Design principles for MHTGR pressure vessels 442 C. Hoffmann Combustion Engineering, Windsor, Connecticut

Section V: Structural graphite components

Chairman: G. Wintermann, RWTÜV-Essen

Materials behaviour and design values 467 G. Haag KFA JCilich, Institut für Reaktorwerkstoffe

Design Methods and Criteria for Graphite Components 480 A. Schmidt Hochtemperatur-Reaktorbau GmbH, Mannheim

Analysis of the graphite side reflector block of the HTR-Module 493 P. Rathjen Interatom GmbH, Bensberg

Design criteria for graphite components of HTTR 506 T. lyoku, S. Shiozawa Japan Atomic Energy Research Institute (JAERI), Japan

Section VI: General Summary

Chairman: R. Schulten, KFA-Jülich

Statements on current HTR structure design criteria 527 R. Trumpfheller vii

Comments concerning the "Workshop on Structural Design Criteria for HTR" 535 W. von Lensa

Summary of the final discussion 539

Appendix 541

Members of working groups and task forces of the German Design Criteria project Objective of the workshop Objective of the Workshop on Structural Design Criteria for HTR Survey of the Research Activities in the Federal Republic of Germany

H. Nickel

Kernforschungsanlage Jülich GmbH, Institut für Reaktorwerkstoffe, Fed. Rep. of Germany

Abstract Technical guidelines and design codes for nuclear plants in the Federal Republic of Germany are tailored to the light water reactor (LWR) systems, so that a design code for the helium-cooled high temperature reactor (HTR) has to be formulated. In an extensive research project, the underlying principles for such a design code have been worked out. The aim of the workshop is to present the current status of the work.

1. Introduction

The purpose of this meeting is the presentation of the results of a nine-year effort to set out the fundamental principles and basic data for a nuclear design code covering high temperature reactor components. At the beginning, work was concentrated on metallic heat exchanger components, but, on completion of the preliminary work about five years ago, it was decided to increase the scope to include all structural components of an HTR plant. This was not least from the plant constructors desired, who wished to move away from the often time-consuming methods applied for the THTR licensing and control procedures. Instead they wanted to build on the good results of and positive experience with the methods of the specialists' working party 'HTR design criteria'.. The advantage of cooperation between materials scientists, plant constructors, stress analysts and safety engineers of the different organisations should be further utilized and transferred to the basic work for the establishment of guidelines and design rules for the HTR. There are two aspects which are of particular significance in the drafting of guidelines and rules:

- adherence to the established status of science and technology; - formulation of generally accepted rules and regulations in such a form that their development for technical systems remains transparent and comprehensible for all concerned.

With regard to a complicated technology such as nuclear reactors, which involves the participation of different institutions and organizations, it. is the second aspect that plays a crucial role. Clearly defined procedures and regulations simplify and ease the carrying out of appropriate steps in which many participants are involved. In the end, the regulations ensure that everyone speaks the 'same language1.

Appropriate guidelines and procedures have been laid down for the light water reactors, an established and commercially marketed system. For the advanced reactor types that have not yet been commercially established such guidelines are not available. The HTR suffers a deficit in this respect.

2. Legal basis for the use of nuclear reactors in the FRG

In this section, the standard framework within which the use of nuclear technology, including of course the HTR, is controlled in the Federal Republic of Germany will be described.

2.1 Highest level of control hierarchy

The planning, construction, commissioning and operation of nuclear plants have to fulfil the appropriate nuclear design codes. The Atomic Energy Act IM and the regulations based on it, for instance the Radiation Protection Ordinance /2/, have to be strictly applied to the HTR (see Figure 1). 2.2 Intermediate level of control hierarchy .

Basic Technical safety regulations for the plant and surrounding area are laid down in the Safety Criteria for Nuclear Power Plants issued by the Federal Minister of the Interior (BMI) /3/. Although they cover all reactor types, they are especially directed towards and applied very strictly to light water reactors (LWR). The analogous transfer of the BMI Criteria from LWR to HTR is related to in-plant inspection, the shut-down systems, the residual heat removal systems, the pressure- retaining containment of the cooling medium and the safety containment. In order to avoid the problems which the analogous transfer of licensing procedure from LWR to HTR caused for the THTR, a draft for the safety criteria for the HTR /4/ and a new draft of the BMI criteria /5/ which should include the HTR were formulated.

The guidelines of the Reactor Safety Commission for pressurized water reactors /6/ or the up-set condition guidelines HI can in part be applied to the HTR. For example, one may consider using them for the design of the reactor pressure vessel of the HTR Module or the HTR 100.

2.3 Lower level of control hierarchy

The regulations of the Nuclear Safety Standards Committee (Technology Committee) (KTA)1 are used for the detailed design of nuclear power stations. The HTR is, however, hardly considered in these rules. It is here on the lowest level of the control pyramid that the most extensive work is needed. The most advanced KTA code contains nuclear design procedures that only apply to the LWR, those which are applicable for all nuclear reactor types, and a few procedures specially drawn up for the HTR. These HTR rules cover the thermal and thermohydraulic design of the reactor. Aspects of the HTR which are not covered by the KTA rules for the licensing procedure are based on specifications that are orientated towards non-nuclear design rules, for example, the German Technique Standards (DIN)2 or the Pressure Vessel Association {AD3,TRD4}

1) KTA = Kerntechnischer Ausschuß 2) DIN = Deutsches Institut für Normen 3) AD = Arbeitsgemeinschaft Druckbehälter 4) TRD = Technische Regeln für Dampfkessel codes of practice. Foreign codes are in part used as appropriate. The whole procedure outlined above can take a long time because of difficulties in obtaining a concensus of agreement.

3. Endeavours for the formulation of design code principles

The described lack of nuclear design codes and guidelines led to delays in the licensing of the THTR. For the various HTR components, there is therefore a need for KTA regulations. Endeavours concerning the establishment of the fundamental principles of a nuclear design code for HTR components goes back to 1979. In a special research programme of the Federal Minister of the Interior /8/, the experience gained from the projects 'Nuclear Process Heat1 (PNP) and 'HTR with helium turbine1 (HHT) was reviewed with emphasis on metallic, high temperature components. Up-set condition topology was at the same time being investigated in a probabilistic study in which the technical safety characteristics of the HTR were clarified /9/.

Based on this preliminary work, an extensive research programme 'Design Criteria for High Temperature Metallic and Ceramic Components and the Prestressed Concrete Pressure Vessel of Future HTR Plants' was begun in 1984, with the support of the Federal Minister of Research and Technology /10/.

The programme was organized as a joint venture. The participants and sponsors are shown in Fig. 2. The activities are divided into four sections (see Fig. 3):

A: technical safety boundary conditions B: metallic components C: pre-stressed concrete pressure vessel D: graphitic structural components

In section A, the HTR integrity concept was developed, based on the special properties of the HTR. The properties of the fuel elements which are stable to very high temperatures play a key role in the concept. Proposals for technical safety classification of components were developed, taking into account the radiological consequences of failure.

In section B, metallic components operating at temperatures above 400°C were examined, based on the experience which has been described in /&/. This work broke new ground, for the application spectrum of the HTR includes components which reach operating temperatures of up to 950 °C. ] The time dependent mechanical behaviour of the materials requires verification methodes which can no longer be covered by purely elastic analysis. Inelastic analysis methods have to be applied for the evaluation of various loading conditions"]] This means a procedure according to 'design by analysis'. The decisive phenomena were recognized and methods for their evaluation shown./There are foreign design codes in existence which go beyond the LWR regulations, specifying procedures for dealing with components operating above 400°C. The relevant design codes are the ASME Code Case N-47 (USAj/and-t-he- -RCC=M—(fiance-)-,- which however are specia 1 ly directed towards the requirements of the fast breeder reactor. They can therefore only cover partial aspects of HTR components."]

HTR plants of limited capacity, up to about 200 MW^n can be constructed using a steel pressure vessel, which may be designed essentially using the principles of the LWR pressure vessels. Plants of larger power output are more advantageously constructed with a prestressed concrete pressure vessel (section C of the programme). For the design of the vessel, the extensive construction guidelines laid down for the building industry can be applied, and additionally the experience obtained in connection with the British Advanced Gas-cooled Reactors (AGR's) is available.

The current status of knowledge for graphitic and ceramic core '•• components (section D of the programme) allows the formulation of a design code, because investigations concerning stress measurement, the influence of neutron irradiation and the corrosion effects during normal service and upset conditions were carried out and comprehensively evaluated in the nineteen sixties. The whole programme with a total of 24 sub-groups comprised 200 individual contributions, and was carried out between 1984 and 1988 /10/. Based on these data the KTA board ask a specialist group to define a first draft of KTA-rules in the HTR-material area. The HTR regulation themes evolving from this programme are:

- Metallic HTR components - Technical safety requirements for the design of HTR prestressed concrete pressure vessels - Ceramic components in the HTR pressure vessel.

4. Acknowledgements

During this workshop, we would like to give an account of the investigations and results of the research and development activities which have taken place over the last nine years. The aim was to establish the principles and data base for structural design rules covering HTR components.

As director of the Institute for Reactor Materials of the Nuclear Research Centre Julien, the organisation responsible for the coordination of the activities, it gives me great pleasure to be able to demonstrate together with our partners the progress we have made in this field. Without the excellent and friendly cooperation of all the partners involved in the different working groups and task forces, we would not have achieved the target set {Fig. 2).

The realization of the project was made possible only by the financial support of the Federal Minister for Science and Technology (BMFT) who continued the sponsorship, which for the first four years came from the Federal Minister for Interior. To these two ministers we record our special gratitude.

Furthermore we would like to thank the Ministry of Economics, Small Business and Technology of the State of North Rhine Westphalia who released the results of the PNP project for use in the design code working groups; the Federal Minister of the Environment, Nature Conservation and Reactor Safety who gave support to the safety-related experimental tasks; the board members of the participating institutions and companies for their favourable support of all the various activities.

References IM Gesetz über die friedliche Verwendung der Kernenergie und den Schutz gegen ihre Gefahren {Atomgesetz) - Act on the peaceful use of nuclear energy and the protection against its hazards (Nuclear Energy Act) - Publication of the revised text of the Atomic Energy Act of 15.07.1985, Bundesgesetzblatt I, Nr. 8, 21.02.1986 HI Verordnung über den Schutz vor Schäden durch ionisierende Strahlen (Strahlenschutzverordnung (StrlSchV)), Regulation on protection against damage caused by ionising rays (Radiation Protection ordinance) of 13.10.1976, last amended on 20.05.1981, Bundes- gesetzblatt I, Nr. 19, 20.05.1981 13/ Der Bundesminister des Inneren, Sicherheitskriterien für Kern- kraftwerke, Federal Minister of the Interior, Safety criteria for nuclear power plants, Bundesanzeiger, Nr. 206, 03.11.1977 /4/ TÜV-Arbeitsgemeinschaft Kerntechnik West, Sicherheitskriterien für Anlagen zur Energieerzeugung mit gasgekühlten Hochtemperaturreaktoren Safety criteria for energy producing plants equipped with gas-cooled high temperature reactors, Draft September 1980, Essen IS/ Der Bundesminister des Inneren, Sicherheitskriterien für Kern- kraftwerke, The Federal Minister of the Interior, Safety criteria for nuclear power plants, Draft, 21.05.1984 /6/ RSK-Leitlinien für Druckwasserreaktoren, RSK guidelines for pressurised water reactors, 3rd edition of 14.10.1981 as amended on 21.03.1984, Bundesanzeiger, Nr. 104, 05.06.1984 /7/1 Der Bundesminister des Inneren, Leitlinien zur Beurteilung der Auslegung von Kernkraftwerken mit Druckwasserreaktoren gegen Störfälle im Sinne des §28 Abs.3 StrlSchV - Störfall-Leitlinien Federal Minister of the Interior, Guidelines on the assessment and design of nuclear power plants with pressurised water reactors with regard to accidents in the meaning of Article 28 para. 3 StrlSchV - Accident Guidelines, Bundesanzeiger, Nr. 245, 31.12.1983 /8/ Nickel H. et al., Erarbeitung von Grundlagen zu einem Regelwerk über die Auslegung von HTR-Komponenten für Anwendungstemperaturen oberhalb 800 °C, Jül-Spez-248, March 1984 to

/9/ Kernforschungsanlage Jülich, Institut für nukleare Sicherheits- forschung/Gesellschaft für Reaktorsicherheit mbH, Sicherheitsstudie für HTR-Konzepte unter deutschen Standortbedingungen (Safety Study for HTR Concepts under German Site Conditions), Jül-Spez-136, vol. 1, 1981 /1G7 Endbericht zum Verbund-Forschungsvorhaben des BMFT: Auslegungs- kriterien für hochtemperaturbelastete metallische und keramische Komponenten sowie des Spannbeton-Reaktordruckbehälters zukünftiger HTR-Anlagen, Jül-Spez (1989) (in Preparation)

Law e.g. Radiation Protection Ordinances Ordinance Safety Criteria for Safety Codes and Nuclear Power Plants Administrative RSK-Guidelinesfor Regulations PWR / BWR

e.g. TRD / AD / DIN / KTA-Rules Technical Rules

Specifications for Specifications Components and Systems for Licensing

Fig. 1: Law, Codes and Guides for Nuclear Power Plants in the FRG

Section I:

HTR Projects and status of licensing principles 15

PRESENT STATUS OF MHTGR PROGRAM IN USA*

Compiled from Contributions from the MHTGR Program Team

for Presentation by P. L. Rittenhouse at Workshop on Structural Design Criteria for HTR

January 31 - February 1, 1989 at Kernforschungsanlage, Juelich, FRG

•"The submitted manuscript haa ba«n authored by a contractor of th« U.S. Government under contract No. DC- AC05-64OR21400. Accordingly, the U.S. Government retains a noneiduvwe. rovsrry-fres beans« to pubksn or raproduc« the pubkshed torm of this contribution, or alow others to do so. for U.S. Government purposes."

*Research sponsored by the Division of HTGRs, Office of Advanced Reactor Programs, U.S. Department of Energy, under Contract No. DE-AC05-84OR21400 with Martin Marietta Energy Systems, Inc. 16

PRESENT STATUS OF MHTGR PROGRAM IN USA

Compiled from Contributions from the MHTGR Program Team

ABSTRACT

The U.S. Department of Energy (DOE) Modular High-Temperature Gas-Cooled Reactor (MHTGR) program has produced a conceptual design which has been reviewed by the U.S. Nuclear Regulatory Commission (NRC). The results of the review were generally favorable, and the program team has now moved into the preliminary design phase. The program team consists of a nuclear island engineering (NIE) team, an energy conversion area (ECA) team, a design integration organization, and a technology development team. Utility user requirements are provided by a utility organization which also participates in design and programmatic reviews/evaluations. This paper will review the direction and accomplishments of each participating organization.

INTRODUCTION

High-Temperature Gas-Cooled Reactor (HTGR) technology in the United States evolved through a "normal" progression from the early 1960s through the mid-1980s. The first plant was the 40 MW(e) Peach Bottom Reactor (pilot scale), which was followed by the 330 Mtt(e) Fort St. Vrain Reactor (demonstration scale). In the early 1970s, 2000 and 3000 MW(t) designs for the Delmarva and Fulton nuclear power stations were completed to the stage of initial fabrication of major components. An NRC review and safety evaluation report (SER) were completed, and NRC was about to issue a construction permit when the projects were cancelled. Subsequently, in response to utility and congressional surveys, conceptual designs for an HTGR gas turbine and a 2240 MW(t) cogeneration plant were produced. Later, in 1982/1983, in response to congressional input and the vendor's perception of the needs of the U.S. nuclear market, a Modular HTGR (MHTGR) was developed. The current reference MHTGR concept comprising 4 x 350 MW(t) modules coupled with two steam turbines was selected in September 1985.

After assessing the factors which led to the rapid decline in interest in nuclear energy during the mid-1970s among utilities, government, and the public, HTGR program management concluded that fundamental changes are needed for the next generation (Millunzi 19881). According to this assessment, the next generation of reactors will be smaller, safer, and simpler than those in operation today. Licensing risk will be reduced and improvements in safety will be achieved through standardization and simplicity of design and use of passive safety features. Investment risk will be reduced through application of a risk-adverse design approach. The MHTGR design process is using an integrated approach that begins with top-level requirements and proceeds downward through the design. Design 17

information from the Large .HTGR (LHTGR) program was used where appropriate. Where design data were missing, assumptions were made, and a technology program formulated to validate (or modify) the assumptions.

A Preliminary Safety Information Document (PSID) was completed in 1986. This document, which was based on the conceptual design of the MHTGR, was forwarded to the NRC in September 1986 to initiate an early dialogue with the NRC on this unique design. Substantial discussions with the NRC took place for the next two years, leading to a sound understanding by the NRC staff of the attributes of the MHTGR, and a well-documented regulatory position to guide DOE in continued design and development of the MHTGR concept. A conceptual design for the MHTGR was completed in early 1988. The preliminary design phase began during the second half of FY 1988.

The remainder of this paper will focus on the major facets of reactor design, licensing, and technology development. Program participants and their roles will be identified and discussed.

DESIGN ACTIVITIES

NUCLEAR ISLAND (NI) •

The nuclear island includes the four reactor modules, each producing 350 MW(t) (see Fig. 1 [Neylan 19882]). The modules are headered in pairs to feed two turbine generators of 300 MW(e) each. The turbines operate in parallel. All systems containing radionuclides and all systems essential to nuclear safety are located within the NI. Therefore nuclear standards and practices are implemented for the NI. The NI includes the reactor core, the reflector, three steel vessels (the reactor vessel, the steam generator vessel, and the crossduct vessel), the shutdown cooling system (heat exchanger and shutdown circulator), the control rod drives, the reserve shutdown system, the main circulator and other components of the heat transfer system, and the reactor cavity cooling system (RCCS).

ENERGY CONVERSION AREA (ECA) ;

The ECA includes all those components not in the NI. The use of high levels of conventional standards and practices will be utilized to meet requirements of the ECA. Included in the ECA are the turbine generator building, operations center, and cooling towers (see Fig: 2 [Neylan 19882]). .

DESIGN INTEGRATION

This important activity includes-: '

t . • Writing and implementation of program policies and procedures Maintenance of the Overall Plant Design Specifications ' 18

Support to program development, planning and control Maintenance of the Summary Level Program Plan (SLPP) Program technical management support and design evaluation Functional analysis implementation Specification of data storage and retrieval requirements Integration of the utility/user requirements Maintenance of licensing plan criteria, methods, and standards Planning and coordination of licensing submittals to the NRC and supporting NRC briefings Interface control Scheduling Quality assurance

LICENSING ACTIVITIES

The licensing objective of the MHTGR program is to eventually obtain a certification rulemaking for the MHTGR from the NRC. The current plan to reach this objective includes three major steps:

(1) Obtain a favorable SER on the MHTGR conceptual design,

(2) Obtain a Preliminary Design Approval, and

(3) Obtain a Final Design Approval for a standard MHTGR

These activities, in combination with experience gained from construction and operation of the first MHTGR, will form the basis for design certification.

It is the policy of the NRC to interact as early as possible with proponents of advanced reactors (applicants, vendors, and government agencies) [see Commission's advanced reactor policy statement 51 FR 24643]. The NRC review of the MHTGR began in October 1986, and was completed in March 1988. The NRC review focused on three key documents prepared by the MHTGR Program. The documents were the Preliminary Safety Information Document (PSID), the Probabilistic Risk Assessment (PRA) document, and the Regulatory Technology Development Plan (RTDP). The RTDP includes the portions of the Technology Development Program (TDP) which focus on safety issues. The issuance of an SER by NRC is imminent, summarizing the results of the review. A preview of the SER content is contained in three papers presented in August/September 1988 (Williams 1988,3 King 1988,* Rogers 19885). Key points from these papers are outlined below: o The MHTGR design reviewed by NRC was conceptual. Therefore, the NRC review concentrated on features and issues related to safety and viability. The General Design Criteria (GDC) used for LWRs were considered by NRC to insure that the MHTGR provided equivalent protection to the public. Several key policy issues arose during 19

the review, due to the different approach by which the MHTGR proposed to meet the criteria.

The NRC staff made recommendations on the four policy issues listed below. Their recommendations are conditional on completion of needed R&D programs, successful resolution of safety issues identified, successful prototype testing at an isolated site, and favorable completion of the deferred review items.

(1) Selection of design basis events

(2) Siting source term calculation and use

(3) Adequacy of containment

(4) Adequacy of emergency planning

Several important areas were not reviewed (see Table 1 [Williams 19883]). Review of these areas has been deferred to the future.

NRC expects, and DOE committed to, timely submission of a revised RTDP and a document describing DOE's plans for prototype testing.

Final determination of the licensability of the MHTGR is contingent upon the following:

Satisfactory resolution of issues identified in the SER. Table 2 (Williams 19883) is a listing of areas where analysis, research and development, and testing results will be required.

Completion of final design and licensing review by NRC

Successful design, construction, testing, and operation of a prototype reactor prior to design certification

TECHNOLOGY DEVELOPMENT

The MHTGR TDP was formulated during 1986 and 1987 to support the design effort. Top-level regulatory (see Table 3 [Cunliffe 19886]) and user requirements (see Fig. 3 [Millunzi 19881]) were established, and linked to design selections through a detailed functional analysis (see Fig. 4 [Millunzi 19881]). Data accumulated during the past 25 years from the LHTGR technology program were used to develop the MHTGR conceptual design. It was determined that additional design data, beyond what is available from the LHTGR program, are required for MHTGR-specific conditions. When suitable data were not available, designers used assumptions from data developed under LHTGR conditions, and then defined Design Data Needs (DDNs) and Technology Development Needs (TDNs) to validate or modify the assumptions (see Fig. 5 [Homan 19887]). The 1987 TDP describes programs 20

in fuel manufacturing, fuel performance, fission product behavior, graphite behavior, and structural materials performance. The 6-year program formulated in 1987 has been stretched somewhat in time due to funding constraints. In addition, several additional TDP areas have been added, including shielding analysis, physics methods validation, and thermal hydraulics validation.

UTILITY INTERFACE

The design and technology development teams have interfaced extensively with the nuclear utility organizations in several areas: o Identification and development of top-level user requirements o Participation in Analysis and Trade Studies (see Fig. 5) o Economic analyses (LaBar 19888) o Evaluation of existing experience base (Dilling 19889) o Evaluation of evolving technology (Gray 198810) o Review of design and design concepts (EPRI 198811)

PROGRAM PARTICIPANTS AND THEIR ROLES

DESIGN

Two design teams have been formed. The NI design team consists of three prime contractors: General Atomics (GA), Combustion Engineering (C-E), and Bechtel National Incorporated (BNI). Stone and Webster Engineering Corporation (SWEC) is the prime contractor for the ECA team, and C-E is a subcontractor to SWEC in this area. The Plant Design Control Office (PDCO) serves as design integrator.

GA has responsibility for design of the core, the control system, and several major components. GA is responsible for fuel fabrication, including fabrication technology development.

James Howden Company of Scotland is a subcontractor to GA for design of the main circulator,

C-E is responsible for the design of the pressure vessel, the steam generator, the crossduct vessel, the steam generator vessel, and the shutdown cooling heat exchanger. In addition, C-E is also responsible for the plant supervisory control system.

BNI is responsible for the RCCS design, shielding design, and all seismic analysis. 21

SWEC is Che prime balance-of-plant (BOP) contractor, and is responsible for designing the ECA, the operations center, and the Plant Control Data and Instrumentation System (PCDIS). Since many of the reliability/availability issues are related to BOP functions, SWEC has primary responsibility for reliability/availability assessments. SWEC is responsible for constructability evaluations, and contributes to economic analyses.

PDCO coordinates and integrates the NI and BOP design efforts. PDCO is responsible for schedule preparation and maintenance, configuration management, maintaining design discipline, and quality assurance. PDCO organizes and coordinates all design reviews.

LICENSING

Gas-Cooled Reactor Associates (GCRA) provided coordination from the MHTGR program side during the two-year review of the conceptual design by NRC. Technical support was provided by all other program participants. PDCO now has the lead for licensing coordination and NRC interactions.

TECHNOLOGY DEVELOPMENT

Oak Ridge National Laboratory (ORNL) has the technical lead in technology development. The technology development program supports only the NI phase of design. DDNs and TDNs are prepared by the design organizations (primarily GA and C-E) . ORNL responds to the DDNs and TDNs with experimental plans, which are reviewed by the designers. Several subcontractors are.involved in the technology program: o GA performs specific tasks associated with fuel fabrication development, graphite testing, and application of design methods to predict the results of experiments. The design methods validation areas are primarily in the area of core physics, fuel performance, fission product behavior, and thermal hydraulics performance. o Massachusetts Institute of Technology (MIT) is involved in three technology-related tasks:

(1) Building a high-pressure fission product loop. This loop will be used to measure the behavior of fission products under temperature and pressure conditions which will be present in an operating MHTGR. The data from the loop experiments will be used to complete formulation of MHTGR fission product behavior models.

(2) Control system design for the Direct-Cycle HTGR (HTGR-DC). This is an advanced application of the MHTGR, expected to offer significant cost and performance advantages over the Steam-Cycle HTGR (HTGR-SC). MIT investigators have recently 22

proposed solutions to two technical problems (recuperator performance and turbomachinery size) which contributed to the demise of the HTGR-DC program in Germany, Switzerland, and the U.S. in the mid-1970s.

(3) Thermal hydraulics code verification and validation.

Commissariat A L'Energie Atomique (CEA) is refurbishing the COMEDIE Loop in the Siloe Reactor at Centre D'Etudes Nucleaires De Grenoble (CENG). The COMEDIE Loop was used during the 1970s for cooperative (GA/CEA) research in the area of fission product behavior for the LHTGR design. This loop will be used to validate the MHTGR fission product behavior models.

C-E also has three tasks in the MHTGR technology development program:

(1) Evaluation of elevated temperature properties of the MHTGR pressure vessel materials. This task includes development of an elevated temperature code inquiry to be submitted to the ASME code committee, interface with ASME subcommittees, and participation in the US/FRG materials subprogram.

(2) Elevated temperature testing and irradiation testing of MHTGR pressure vessel materials. This task includes evaluation of results from the ORNL creep program, and results from the Nil Ductility Transition Temperature (NDTT) shift program. This work is discussed in more detail in a separate paper at this workshop.

(3) Steam generator seal development and testing. This task includes evaluation of steam generator seal concepts and requirements, completing trade studies, selection of a reference seal design, update of shroud seal test specifications, and initial design of test models and rigs.

UTILITY INTERFACE

GCRA manages and coordinates the utility interface. GCRA is funded directly by participating utilities. GCRA has direct access to the Electric Power Research Institute (EPRI), utility executives, and operations managers. These contacts are a valuable source of information and data which are used to accomplish the tasks listed earlier.

CONCLUSIONS

The U.S. MHTGR program is now in the preliminary design phase. The program has utilized a disciplined, functional analysis approach to develop a conceptual design. The conceptual design has been reviewed by 23

the NRC, with favorable results. A well-integrated design team is in place, supported by a technology development organization with an international flavor. When the design, licensing, and technology development programs were formulated in 1986, a six-year program was planned. Funding constraints have stretched the program schedule since 1986, but these three major activities still are closely coordinated and can deliver the first MHTGR by the mid- to late-1990s.

REFERENCES

1. A. C. Millunzi and S. R. Penfield, Jr., "Developing a Reactor for Today's Realities and Tomorrow's Needs," Proceedings of the 23rd Intersociety Energy Conversion Engineering Conference (Volume 1), August 1988, pp. 479-482.

2. A. J. Neylan, D. A. Dilling, and R. Ng, "Designing a Reactor for the Next Generation," Ibid... pp. 483-488.

3. P. M. Williams, T. L. King, and J. N. Wilson, "Results of a Preliminary Safety Review of the MHTGR," paper presented at Tenth International Conference on the HTGR, San Diego, California, September 20, 1988.

4. T. L. King, "Safety Evaluation of the Modular High-Temperature Gas- Cooled Reactor," Proceedings of the 23rd Intersocietv Energy Conversion Engineering Conference (Volume 1), August 1988, pp. 495-497

5. K. C. Rogers, "Key Questions Facing the NRC on the MHTGR," paper presented at the Tenth Annual International Conference on the HTGR, San Diego, California, September 19, 1988.

6. J. C. Cunliffe and F. A. Silady, "The Challenge of Licensing a Reactor With Passive Safety Characteristics," Proceedings of the 23rd Intersociety Energy Conversion Engineering Conference (Volume 1), August 1988, pp. 489- 494.

7. F. J. Horaan and A. J. Neylan, "MHTGR Technology Development Plan," Ibid.. pp. 511-514.

8. M. LaBar and H. Bowers, "Economic Characteristics of a Smaller, Simpler Reactor," Ibid.. pp. 449-504.

9. D. A. Dilling and G. Jones, "The Potential for Modular Construction/Zone Outfitting of the MHTGR," Ibid.. pp. 531-535.

10. S. Gray and G. Jones, "Magnetic Bearings: A MHTGR Design Selection With Broad Industrial Potential," Ibid, pp. 521-524.

11. EPRI Review Meeting on the Modular High Temperature Gas-Cooled Reactor Program, July 12-14, 1988, San Diego, California. 24

Table 1 SUMMARY OF AREAS CONSIDERED IK NRC REVIEU OF MHTGR

Areas Reviewed Fuel design Reactor physics Reactor vessel Passive heat removal systems Safety analysis Heat transport equipment Components of the primary system boundary Instrumentation Control Electrical systems Selected auxiliary systems Occupational exposure Human factors Safeguards and security Some balance of plant items

Areas Deferred Seismic design Radioactive waste handling systems Mechanical equipment design Structural graphite components Modeling of fission product transport Nuclear design Phenomena involving chemical processes Fluid flow design Reactor internals Vessel system and subsystems Heat transport system and subsystems Shutdown cooling system and subsystems Reactor cavity cooling system Reactor building Plant protection and instrumentation system Plant control, data, and instrumentation system Miscellaneous control and instrumentation group Electrical systems Service systems Steam and energy conversion systems Operational radionuclide control Occupational radiation protection Emergency preparedness Role of operators Safeguards and security Prototype plant testing Safety analysis Technical specifications and administrative controls Quality assurance 25

Table 2 LISTING OF MAJOR AREAS REQUIRING SUPPORTING ANALYSIS, RESEARCH, OR TESTING

Fuel Design Fuel performance models Fuel performance statistics from laboratory testing Manufacturing quality control Fuel performance under accident conditions Effects of fuel composition on performance Effects of external chemical attack on fuel performance

Nuclear Design Methods and data validation Uncertainties in negative temperature coefficient of reactivity Control materials

Thermal and Fluid Flow Design Core flow distribution Not streaks .Differential pressures and shear forces during depressurization events Flow-induced vibration on control rod guide tubes

Reactor Internals Seismic design and fragility data In-service deterioration of materials

Vessel System ASME and staff approval for elevated temperature service Catastrophic failure probability Neutron irradiation effects Seismic design, including support system

Reactor Cavity Cooling System and Reactor Cavity Heat transport design Vessel hot spots ln-vessel conduction Emissivities Effect of water vapor Repair and recovery Modeling conservatism and sensitivities to uncertainties Seismic design and fragility data Reactor cavity temperatures Duct and chimney design Heat transmission to the earth

Radionuclide Control (Source Term, fission Product Transport) Assumptions and model for back calculation from site boundary

Operations Advanced control system development Human factors analysis Crew size and training Worker exposure during maintenance

Prototype Plant Testing 26

Table 3

SWtWRY OF TOP-LEVEL REGULATORY CRITERIA FOR THE ttfTGR

Criteria Type Criteria Public Risk NRC Policy Statement on Safety Goals Prompt fatality risk 5 re« WB;>25 rem thyroid 27

US-DOE MHTGR PROGRAM

^RADIOACTIVE WASTE /MANAGEMENT BUIIOING

SWITCHYARD.

1-125(2) GENERAL ATOMICS 7-29-88

Fig. 1 MHTGR Site Plot Plan 28

US-DOE MHTGR PROGRAM

t HANSMISSION SYSTEM NUCLEAR (SIAND t L

FW PUMP \ rEE0WATER CONDENSATE

REACTOR MOOUIE NO. 2 M - FROM: • 1 POWER TO: CONVERSION REACTOR TRAIN Nai -125(1) MOOtftES GENERAL ATOMICS 10-7-87 NO. 3. AN0 4

Fig. 2 4 x 350 MW(t) MHTGR Energy Conversion Plant 29

SAKE ECONOMIC POWER

MAINTAIN MAINTAIN MAINTAIN MAINTAIN PLANT PLANT CONTROL OF EMERGENCY OPERATION PROTECTION RADIONUCLIDE PREPAREDNESS RELEASE

Fig. 3 Top-Level Functional Structure

US-DÜE MHTGR PROGRAM

USER TOP-LEVEL REQUIREMENTS REGULATORY CRITERIA

INTEGRATED APPROACH

ENGINEERING PRODUCT PLANT DESIGN, ETC.

Fig. 4 MHTGR Design Approach 30

ORNL-DWG 88-8737

DESIGN PROCESS REQUIREMENTS 1- 2-

ANALYSES AND TRADE STUDIES 1- 2-

• ASSUMPTIONS i- • 2-

DESIGN SELECTIONS 1- 2- RECYCLE TRADE STUDIES/ANALYSES EVALUATION: DOES DESIGN MEET ALL REQUIREMENTS REEVALUATE WITH NEW DATA YES END PLAN AND DO ANY ASSUMPTIONS NO DDN REQUIRED CONDUCT REQUIRE VERIFICATION NO TECHNOLOGY PROGRAM YES PREPARE DDN(s)

Fig. 5 Relationship of Design Data Needs to Design Process 31

Discussion of the presentation:

Present status of MHTGR program in USA

Nickel, KFA Julien, FRG: Do you provide input to the design and the licensing procedures of the four 350 MWth reactor modules for commercial applications by the modular system for tritium production, the so called New Production Reactor (NPR)?

Rittenhouse, ORNL, USA: The current design of the MHTGR NPR utilizes fuel of Fort St. Vrain quality and a containment. Questions have been asked by licensing authorities as to "if you need containment for the NPR why not for the commercial plant?". Part of the answer is that fuel of much improved quality will be employed in the latter. In general, NPR technology is today's technology >• the commercial MHTGR will utilize tomorrows technology. This is the reason for Technology Development Plans discussed in the paper.

Schuster, KFA Julien, FRG: Is the power of a module in the MHTGR (350 MW-t) related to the maximum fuel temperature which would be reached during accident conditions?

Is it possible to maintain or repair one of the 4 modules whilst the other modules still operating?

Rittenhouse, ORNL, USA: Yes, through considerations on power density and care and cooling systems geometries, maximum fuel temperature limit is still 1600 °C.

Yes, although during normal full power operation the four modules are headered in pairs to two power conversion trains, each module can feed into either of the power conversion trains. Thus, operation is possible with one, two, three or four modules matched appropriately to the power conversion trains. 32

Kirch, KFA Jülich, FRG: What is the latest position of your licensing authorities concerning the containment question for MHTGR?

Rittenhouse, QRNL, USA: Licensing authorities have not yet agreed that a containment is necessary. However, neither have they decreed that a containment will be required. The question is still open but the proof will be the responsibility of the MHTGR Program. The design proceeds at the moment without a containment. Use of a containment would make very difficult the application of the current RCCS-air cooling. (See also answer to question by Nickel). 33

PRESENT STATUS OF HTTR PROJECT IN JAPAN

Toshiyuki TANAKA and Shinzo SAITO

HTTR Designing Laboratory Japan Atomic Energy Research Institute Tokai-mura, Naka-gun, Ibaraki-ken, 319-11 Japan

Abstract

. In Japan, the research and development on the HTR had been carried out for more than fifteen years as the multi-purpose VHTR program for direct utilization of nuclear process heat such as nuclear steel making. Recently, reflecting the change of social and energy situation and with no incentives for industries to introduce such in the near future, the JAERI has changed them for more basic "HTTR program" to establish the HTR technology basis and upgrade them.

The HTTR is a test reactor with thermal output of 30 MW and an outlet coolant temperature of 950°C, employing a pin-in- block type fuel block, and has the capability to demonstrate nuclear process heat utilization using an intermediate heat exchanger. Since 1986 detailed designs, in which major systems and . components are determined in line with the HTTR concept, have been made for safety review of the Government which will start early in 1989. At the request of the STA the Reactor Safety Research Association has reviewed the safety evaluation guideline and started the work on the establishment of general design criteria and design code or guide for graphite and high- temperature structure of the HTTR. Construction permit of the HTTR will be issued by the Government early in 1990. 34

1 . Introduction

In 1969 the Japan Atomic Energy Research Institute (JAERI) started the development of a very high-temperature gas-cooled reactor (VHTR) which can produce gas of nearly 1000°C at the outlet and research and development (R&D) on various subjects such as fuels, materials and components have been carried out. The social situations since the start of R&D have changed, and presently the demand of nuclear heat in Japan is not strong. In June 1987, the Japanese Atomic Energy Commission issued the revision of the Long-Term Program for Development and Utilization of Nuclear Energy, recommending that Japan should proceed with the development of more advanced new technologies for the future, parallel to the existing nuclear systems. It also emphasizes that the HTR is one of the most promising reactors with high efficiency and inherent safety, therefore it should be explored for the broader use of nuclear energy, not only for power production. Then the early construction of a High-Temperature Engineering Test Reactor (HTTR) by JAERI was proposed. According to this program, the JAERI has changed the VHTR for "HTTR program" to establish the HTR technology basis and upgrade them. The application for the construction of the HTTR will be made by the JAERI to the Science and Technology Agency (STA) in February 1989.

2. Outline of HTTR

The HTTR consists of a core of 30 MWt, a main cooling circuit, an auxiliary cooling circuit and related systems. The reactor pressure vessel is 13.2 m high and 5.5 m in diameter and contains the core, graphite reflectors, core support structure and radial restraining devices as shown in Fig. 1 . A major specification is listed in Table 1. The reactor building contains a containment vessel, sub-systems for cooling systems, 35

ventilation and air conditioning systems, a reactor control room, a spent fuel storage pool, etc. as shown in Fig. 2. The reactor vessel is placed at almost the center of the reactor building. The containment vessel is rather small and it has a big nozzle above the reactor vessel. The nozzle cover is removed during refueling. The reactor core is graphite moderated and cooled by helium gas, and prismatic fuel elements in the form of hexagonal blocks are used. The active core consists of 30 fuel columns and 7 control rod columns, each column being 5 blocks high 2.9 m. The active core of 2.3 m in diameter is surrounded by 12 replaceable reflector columns, 9 reflector-zone control rod columns, and 3 irradiation test columns1. The permanent reflector surrounds the replaceable reflector and consists of large polygonal graphite blocks fixed by restraining devices. Each hexagonal graphite block has three dowels on the top and three associated sockets at the bottom, and the blocks are fixed by the dowel sockets. This method has been proved sufficient against the design earthquake. Reactivity control is provided by control rods which are individually supported by mechanisms located in stand-pipes connected to the hemispherical top head of the reactor vessel, and inserted -into channels in the core and replaceable reflector region. The reactor shutdown under the high-temperature condition is made by inserting 9 pairs of control rods into the reflector region at first", while 7 pairs of control rods' in the core region are added for the shutdown at low-temperature. Back-up shutdown capability is provided by insertion of boron carbide/graphite pellets into separate holes in -the control rod blocks. Refueling is accomplished with the reactor shutdown and depressurized.

The reactor core is cooled by helium gas of 395 °C at the reactor inlet temperature which flows downward through- the core. The maximum fuel temperature in the normal operational condition is approximately 1 500°C. Reactivity 'power coefficient is largely negative and each reactivity temperature coefficient is also negative. Major nuclear and thermal- 36

hydraulic characteristics are tabulated in Table 2. A fuel element assembly, 36 cm across the distance between side surfaces and 58 cm in length, is made of fuel rods and a hexagonal graphite block as shown in Fig. 3. The fuel consists of TRISO coated particles of low enriched uranium oxide whose average enrichment is about 6 % and the kernel diameter is 6 00 pm. The particles are scattered in a graphite matrix and sintered to form fuel compacts which are contained in a sleeve to form a fuel rod. The fuel rods of 3.4 cm in diameter are contained within vertical holes of the graphite blocks. Helium gas flows through the gap between the vertical hole and the fuel rod to remove heat produced by fission and gamma heating. The associated fuel performance tests for domestic fuels have been conducted in the in-pile gas loop (OGL-1 ) and other out-of-pile facilities, and thermal-hydraulic tests have been performed in the Helium Engineering Demonstration Loop (HENDEL). The reactor cooling system is composed of a main cooling system (MCS), an auxiliary cooling system (ACS) and two reactor vessel cooling systems (VCSs). The reactor cooling system is. schematically shown in Fig. 4. The ACS is in the stand-by condition during normal reactor operation and operated to remove, residual heat from the core when there is a trouble in the MCS but the flow in the primary cooling circuit is still kept. Both VCSs are operated at 100 % flow rate during the normal operation in order to cool biological shield around the reactor vessel, and they serve to cool the reactor vessel and the core in an accident such as a pipe break of the MCS in which the flow in the primary cooling circuit is not kept. The MCS is, therefore, non-safety class but the ACS and the VCS are safety classes.

The MCS is separated into two lines outside the reactor vessel. The heated helium gas is cooled by a He/He intermediate heat exchanger (IHX) in one line or cooled directly by a pressurized water cooler (PWC) in the other line. Heat is finally removed by an air cooler in both lines, although another PWC is necessary after the IHX in the first line. When 37

the first line with heat transfer capacity of 10 MW is operated, the second line, which has heat transfer capacity of 30 MW, is operated at 20 MW. The IHX is a heat exchanger of vertical helically-coiled counter flow type as shown in Fig. 5. Primary coolant flows on the shell side and secondary coolant on the tube side. Materials for the cold pressure boundary and very high- temperature structures are 2 1/4Cr-1Mo steel and Hastelloy XR, respectively. The PWC is very similar to the IHX, which is a vertical U-tube type heat exchanger as shown in Fig. 6. Primary helium gas coolant flows outside the heat transfer tubes and pressurized water flows inside the tube. Coaxial double pipes as shown in Fig. 7 are used for transferring hot helium gas. Cold gas of about 395°C flows toward reactor inlet through the annulus between the inner and outer pipes and hot gas from the reactor flows inside the inner pipe. The inner pipe is almost in the same temperature as that of the outer pipe because the inner pipe is internally insulated with Kaowool and lined with Hastelloy XR. The outer pipe has a function of the pressure boundary, while the inner pipe is designed only to withstand the differential pressure between the inner and outer pipes. Difference in thermal expansion between these pipes is absorbed by flexibility of the inner pipe. Therefore, a compensator such as bellows joint is not used. The ACS consists mainly of an auxiliary heat exchanger (AHX), auxiliary gas circulators (AGCs) and an air cooler. The heat transfer capacity of the ACS is about 3.5 MW and these components are similar to those of the MCS.

The VCS has the cooling function mentioned above and each VCS has heat removal capacity of 100 % to cool reactor vessel and core in emergency. Figure 8 shows irradiation holes in the core and the reflector .region. Thermal and fast neutron fluxes in these regions are of the order of 10^ n/m^ • s, and the temperatures are between 400°C and 1100°C, depending on the axial and radial positions in the core as shown in Table 3. Irradiation of large-sized samples are also possible by using a graphite 38

basket as shown in Fig. 9. The JAERI is planning irradiation tests of spherical fuels by using the basket, and capsules are intended for irradiation of small samples.

3. Present status of HTTR project

Table 4 shows a current schedule of the HTTR project. The final stage of detailed design started in December 1988, with start of construction scheduled early in 1990 and first criticality 1995 Fiscal Year.

The safety review inside JAERI and preliminary hearing by the STA officials have been carried out and the safety review of the Government will start in February 1989, while the part of construction budget of HTTR, for construction of the building foundation and manufacture of the reactor core support structure, was approved this January by the Ministry of Finance.

At the request of the STA the Reactor Safety Research Association has reviewed the safety evaluation guideline and started the work on the establishment of general design criteria and design code or guide for graphite and high- temperature structure of the HTTR. The related R&D has been almost finished to validate the design and safety of the HTTR and now redirected for upgrading high-temperature technology.

HTTR is the sixth "graphite-moderated and helium-cooled reactor" in the world and will serve as the test facility to establish basic technologies for advanced HTRs and also to conduct innovative researches in high-temperature technologies. Demonstration of nuclear heat process utilization using the HTTR might be planned as the international cooperative program in future. 39

Table 1 Main parameters of HTTR

Reactor thermal output 30 MW Reactor outlet coolant temperature 850'C/950°C Reactor inlet coolant temperature 395'C Fuel Low enriched UO2 Fuel element type Prismatic block Direction of coolant flow Downward-flow Pressure vessel Steel Number of main cooling loop 1 Heat transmission IKX and PWC (parallel loaded) Primary coolant pressure 4 MPa Containment type Steel containment Plant lifetime 20 years

Table 2 Major nuclear and thermal-hydraulic parameters

Main specification Thermal power 30 MW Core diameter 2.3 m Core height 2.9 m Average power density 2.5 W/cm* Fuel loading off-load, 1 batch Nuclear Excess reactivity 15% Jk Uranium enrichment 3~10 w\% average about 6wt% fuel burn up (overage) 22 GWd/t Reactivity coefficient Fuel temperature coefficient -(4.6 to 1.5)X10-S Jlk/kfC Moderator temperature coefficient {-17.1 to 0.9?) X 10-5 Jk/k/"C Power coefficient -(2.4 to 4.O)xio-3 /fk/k/MW Prompt neutron lifetime 0.67~0.78 ms Effective delayed neutron fraction 0.0O47~0.0065 Thermal-hydraulic Total coolant flow 10.2 kg/s Inlet coolant temperature 395 *C Outlet coolant temperature 950 'C (max.) Power peaking factor Radial 1.1 Axial 1.7 Effective core coolant flow rate 88 % Max. fuel temperature 1495*C 40

Table 3 Irradiation test conditions

Max. specimen Max. thermol Envifonmenial Irradiation size nuetron flux temperature Example of irradiation teil region (mm) (n/m'-S) CO

* Melal and ceramics irradiation tests Center column 1 block sue 7X 10" 400 — 1100 * fuel failure test in block

Fuel column 1 block size 5X1O17 400—1100 • Fuel perform on« test in block

* Metal and ceramics irradiation tests Replaceable 250 (Dia.) * Fuel failure test in capsule reflecior 4X JO17 400—800 X500 (L) ' Continuous loke-oul el tritium (hole A) ' Irradiation creep test

Replaceable 130 (Dia.) * Development of tritium production reflector 4X1017 400—800 XS00 (I) end take-out techniques (hole B)

Permanent 100 (Dia.) 3X1017 400 —BOO * Metal and ceramics irradiation tests reflector X3000 (L)

Table 4 Schedule of the HTTR program rT87 88 89 90 91 92 93 94 95 —

Design 1 icensing j i

1 Construction j

• j Oporalio n, Tests [ 41 'ODD

Stand pipe m

— Pi s Reactor pressure vessel

Fuel region

Permanent reflector

Replaceable reflector

Core restraint mechanics

Hot plenum block

Outlet pipe

Fig. 1 Bird's-eye view of the* reactor vessel and core

Air Cooler

Intermediate Heat Exchanger

Reactor Containment Veisel Reactor Vesiel

Fig. 2 Bird's-eye view of the reactor building. 42

Fuel handling hole

Fuel kernel Dowel High density PyC

SIC fuel low density PyC- ^compact Ctoptiilc

8 mm

39mm

Fig. 3 Bird's-eye view of block type fuel

Containment veit«l lnUrm«di*U Kail «icKi Prlmarji prctaurltad Vend cooling panel wiUr caolvr PCC Prlmarjr |ai cl reutet er SPWC wiiar c«ot*r sec S*c»ndarr |a« circulator AKX AualHar]' haat aicKanfar AGC Auilli

Air cooler

Waltr pump

Fig. 4 Flow sheet of cooling system 43

Secondary Ho to SPWC Secondary Ht from SPWC

Primary He to Gas circulator

Primary He from gas circulator

Inner shell Outer shell Tube support Manifold Insulator

Helical tube Primory He to gas circulator Header (independent PWC operation)

Primary He to gas tirculator (combined operation with PWC) Outer shell

Primory He from [[ ty Primary He to RPV Buffie plate Primary He (rom RPV gas circulator

Fig. 5 Schematic, drawing Insulator Tube of the IHX Inner shell

Primary He to Thermal shield RPV

Primary He from RPV

Tube sheet

Water outlet \ Separator wall

Fig. 6 Schematic drawing.of the PWC 44

High temp. He from reactor

Noiile

irmc PWC 7 Liner Inner pipe Insulator Outer pipe

Fig. 7 Schematic drawing of coaxial double pipes

Center column test region

Fuel irradiation test region

Control rod channel block

Permanent reflector Irradiation hole 8 in replaceable reflector Replaceable reflector

Neutron detector hole

Irradiation hole in permanent reflector

Irradiation hole A in replaceable reflector

Fig. 8 Cross sectional view of irradiation regions 45

Flonge

Spherical Fuel

Coolant Hole — Boron Ball

Fig. 9 A graphite basket for spherical fuel irradiation 46

Discussion of the presentation:

Present status of HTTR project in Japan

Kirch, KFA Julien, FRG: We congratulate JAERI for the positive decision on HTTR. The design does not take much advantage of typical HTR safety features. Is there a chance to change this for future HTR's in Japan?

Tanaka, JAERI, Japan: Though the discussion on the safety evaluation guideline and also based on the safety review inside JAERI and preliminary hearing by the STA officials, we are obliged to follow the LWR practice for the time being. This is also true in the case of the FBR. In the future we hope a more reasonable practice might be established in Japan.

Schuster, KFA Julien, FRG: Do you plan to connect power consuming processes to the HTTR, especially to the IHX secondary circuit?

Tanaka, JAERI, Japan: The reactor is cooled by IHX and PWC and heat is finally removed by an air cooler. In the future we will connect a thermal utilization system, such as steam reformer or a hydrogen production unit. Now we are performing fundamental research in this area.

Nickel, KFA Julien, FRG: The proposed outlet temperature of the coolant will be 950 °C, this leads to a maximum fuel temperature of > 1500 °C. Can you say what you are doing against the Amoeba effect in your fuel element design?

What are the specific design data against earthquake? 47

Tanaka, JAERI, Japan: Present design of fuel and core limits operating time to 330 days at 950 °C outlet temperature operation. To overcome the Amoeba effect problem we need to increase the thickness of the first layer of PyC.

We have accumulated such data with one dimensional and two dimensional graphite block vibration tests. Now we make seismic experiments on the full-sized core bottom structure to validate computer codes and also reliability and confirmation tests of CRD insertion during an earthquake have been carried out.

Groß, HRB Mannheim, FRG: You have a relatively high core inlet temperature of about 400 °C. What about the temperatures in the upper plenum of the reactor vessel in the case of shutdown of the helium-blowers, that is in the case of interruption of the gas flow?

Tanaka, JAERI, Japan: We plan to instal thermal insulation layers of metal foils inside the reactor vessel top cap. So maximum temperature of each vessel under abnormal conditions is lower than the design temperature of 450 °C. 48

HTR-situation in China

D. Wang, S. Xu

Institute of Nuclear Energy Technology P.O. Box 1021, 102201 CN, Beijing, China

Abstract

Research on HTR technology can be divided into two stages in China.

The first stage is from 1975 to 1978. The main aim was to look into the possibility of using HTR as a high convertor as well as a power producer. Various aspects of HTR technology have been invescigated.

The second stage started in 1984. At the beginning of the stage, the main activities include:

-- feasibility study of using HTR in China, -- R & D of HTR technology, -- establishing international exchange and cooperation.

On the basis of these activities, HTR technology has been put into the state's high technology programme and is earring out by the institutions from the whole country.

International cooperation with KFA, and German industry is widened and deepened. A joint programme has been set up to build a 10 MW test module HTR at INET, China. 49

1. Introduction

China is a large country with vast territory and great population; There is a big demand tor energy for its economic development and rising the living standard of its people. • In general, China has rich energy resource to meet the demand, but the seperation of energy resource area and the industrial area makes great difficulties in energy supply. Nuclear energy will play more important ,role in China than in some other countries.

There are two LWR power plants under construction, however, nuclear energy business is still in its infancy stage in our country. We have more scope and flexibility in choosing nuclear•energy system which suits our situation and demand best.

For its specific design, HTR- has the advantages of high working temperature which means higher thermal efficiency and wider application area, high fuel utilization capability, inherent safety feature and high flexibility in fuel cycle. All these make HTR one of the most attractive nuclear energy systems. R & D, and commercialization of HTR have been actively pursued in the world^

In China, ft & D programme was first set up in the middle of 1970's at INET, Beijing. The aim of the programme was to look into the feasibility of building a HTR with high fuel conversion ratio and the application potential of such reactor. The answer was possible but difficult. The work was stopped in 1978.

To find the way out for HTR development in China, INET started R & D work on HTR again in 1984. Work on some key 50

item of HTR technology was carried out in order to show the reality and benefit of building a HTR system in China. Atthe same time, contact with the international HTR community was established, close cooperation between INET, KFA and. German industry has been developed. On the basis of these activities, HTR technology has been put into state's high technology programme, and a joint programme ot building a 10 MW test module HTR at INET has been signed by Interatom GmbH. INET. and KFA GmbH.

2. Feasibility study on high conversion ratio HTR power plant

The main aim of the study was to fully use - of the potential of HTR. Because there was no demand envisaged using HTR as high temperature process heat supplier, the work was concentrated on using HTR as a high convertor and as a high efficient power producer.

The conceptual design of a 10 MW reactor was carried out. All aspects related to build such reactor was investigated. The emphasis was put on the following items. .

1) Reactor design

2.) Reactor components design and experiment, e.g. reactor structure, pre-stressed concrete pressure vessel, fuel element load and unload system, etc.

3) Fuel technology, include fresh fuel element fabrication spent fuel element reprocessing, and fuel element refabrication. 51

4) Enviroment effect and. safety analysis.

It was made clear that, to get high conversion ratio,

the following conditions have to be fulfilled.

a) High speed fuel cycle is favoured to minimise the cummulation of U232 and nuclear transmulation of Pa233. To meet this demand, pebble bed type of reactor design was selected. Biso coated fuel particle with resine kernel and loosely packed.in graphite ball is a suitable fuel element design;

. b). The use of Th-U233 cycle is necessary, therefore the fuel element and breed element was seperately designed, fabricated, cycled, reprocessed and refabricated;

c) High recovery ratio of nuclear fuel during reprocessing and refabrication is highly hoped.

The conclusions are .

-- Highconvession ratio could be realized under certain conditions.

-- There is no serious - difficulty in civil engineering aspect for building a HTR in China.

• -- A great deal of R & D work in nuclear -field, is needed, especially when safety and high conversion ratio are concerned. 52

3. R & D of" HTR technology.

As mentioned above, there are a lot of R & D work on HTR technology which have to be done before erection of a HTR in China. Feasibitily study on application of HTR, and fuel fabrication technology was started again in 1984. After years preperation, international cooperation is established and developed, especially cooperation between INET and KFA is widened and deepened. This has promoted our work greatly.

On the basis of these efforts, HTR technology has been put into the state's high technology programme since 1987. The first stage of the HTR programme will concentrate on the feasibility study of using HTR as efficient power plant and process heat supplier at intermediate temperature. The items of R & D work are more or less same as previous, but the contents of them are greatly changed, for example,, the fuel cycle is changed to low enriched uranium cycle, Triso coated fuel particle will be choosed and AVR/THTR type spherical fuel element is to be used. Spent fuel element will be stored. In consideration of the difficulty of fuel reprocessing technology and long term object of using Th in HTR, R & D of fuel reprocessing technology is included in the programme, even so it is not necessary for the first stage of HTR programme. The programme is now being carried out by the institutions f rom -tAe. whole country.

The programme was put into operation in 1986. Conceptual design has been carried out; A laboratory for fuel technology R & D is under construction, samples of cp and spherical fuel element has been produced; key items of R & D of fuel reproccessing has been investigated; R & D of He technology proceeds smoothly. 53

International cooperation has been expanded to. industry abroad in recent years. A joint programme of developing HTR technology in China has been signed at the end of 1988. The partners are Interatom GmbH. IWET and KFA GmbH.

The key point of the joint programme is to build a 10 MW test module HTR to provide a nuclear test facility with which relevant and unique features of the HTR module can be demonstrated. The main items are as following

1) Froduct application : electricity generation . : process heat generation district heat generation methane reforming

2) Components testing - • graphitic core stuctures steam generator helium blower fuel handling

3) Verification of HTR-Module inherent safety featurest negative temperature coefficient of reactivity^ temperature limitation due to passive decay heat removal, limitation of power excursion due to water ingress

4) Fuel element mass test for temperature up to 1600°C

The Test-Module will be located at INET, Beijing. The time schedule of the programme is shown in fig 1. 54

The conceptual design has been completed, the main data of the lü MW test module are as following

Maximum thermal power 20MW Average thermal power 10MW Primary helium pressure 30Bar Secondary steam pressure 35Bar Cold helium temperature 250°C Average hot helium temperature 700°C steam temperature 435°C core volume 5M3

The structure of the test module is show in figure 2.

Training programme is now being carried out, the feasibility study on the application of HTR module in China is performed smoothly.

4. comclusion

Good enviroment for the development of HTR technology has been created in China. A prosperous future of HTR is envisaged.

r.n

»1000

msoo

FIG. 2 CROSS SECTION OF THE PRIMARY CIRCUIT OF THE TEST MODULE 57

Reference

1) Lu Yingzhong, Wang Dazhong, Zhong Daxin, Gao Zuing; Qin Zhenya. The study of the HTR technology and its industrial application Proceedings of the 6th PBNC. Sept. 1987, Beijing.

2) Xu Yuanhui Status of HTR R & D programme in China. Report of INET, Beijing, 1988.

3) W. Frohling, K. Kugeler, R. D. stoll, Zhong Daxin. Ye Daxin Combining steam drive technology and ' small nuclear plants with reference to shengli oilfield/China. International conf, on heavy crude and tar sands Aug. 7-12. 1988. Edmonton, Canada,

4) Zhong Daxin The application studies of HTR in PRC. IAEA, Technical committee meeting on design requriements for future applications of advanced concepts in developing countries. 6-11 Dec. 1988, Vienna

5) Chinese-German R & D cooperation HTR Test-Module China final report for the concept phase Sept. 1988. 58

Discussion of the presentation

HTR situation in China

Schuster, KFA Julien, FRG: Is there competition between the HTR project line and the LWR development in China? Do you intend to use the HTR for applications which cannot be covered by the temperature of a LWR?

Xu Shijiang, University Peking, China: As mentioned in the presentation, China's nuclear energy is still in its infancy period. We have the flexibility to choose an advanced reactor type as the main line for our long term nuclear energy system. In this sense, there is competition between different reactor systems. HTR is the only reactor system which could provide high temperature process heat, therefore even if HTR could not become the main line of electricity production, we do intend to use the HTR for applications which cannot be covered by LWR, for high temperature thermal energy represents a large proportion of the total energy consumed.

Wolters, KFA Julien, FRG: What is the reason for choosing the low outlet gas temperature of 700 °C instead of 750 °C and the low steam temperature of 430 °C instead of 530 °C?

Xu Shijiang, University Peking, China: The main aim of building the 10 MW test module HTR is to provide a nuclear test facility with which the unique features of HTR module can be demonstrated. Electricity is just a by-product, otherwise the energy will be discarded as waste heat. Therefore the parameters of the reactor design are not optimal, they are designed according to what is already easily available and economical. 59

'Present Status of the High Temperature Reactor in the Federal Republic of Germany"

H. Nickel

Kernforschungsanlage Julien GmbH, Institut für Reaktorwerkstoffe, Fed. Rep. of Germany

Abstract

In the Federal Republic of Germany, the helium-cooled high-temperature reactor (HTR) and the fast breeder reactor are • being promoted as advanced nuclear reactor systems. The HTR development based on spherical fuel elements has led to the construction of a 300 MWej prototype reactor, the THTR 300, following favourable experience with the operation of a small experimental reactor, the AVR. The THTR 300 became operational in the mid-eighties. It has exhibited good service behaviour and has confirmed the favourable safety characteristics of the system.

At the present time, two HTR concepts are being followed up by German companies, a small HTR (approx. 200 MW-tn) with a steel pressure vessel and a medium-sized reactor (approx. 550 MWei) with a prestressed concrete pressure vessel for electricity generation.

The special safety features of the HTR originate in the core construction of ceramic materials resistant to high temperatures and the" low power density of the core. An analysis of the technical safety features shows that the general activity containment concepts developed for light-water reactors (LWR) are not transferable to the HTR. It was necessary to develop technical safety concepts especially for the HTR, and the result is the HTR integrity concept.

Licensing procedures have been established for the design, construction and commissioning of the prototype THTR 300 and invaluable experience gained. A concept for licensing approval of the HTR module plants, independent of the siting of the nuclear plant, is now being investigated. 60

1. Introduction

The high temperature reactor (HTR) is a development of the graphite- moderated, gas-cooled reactor class. Early reactors of this type are the Magnox and the advanced gas reactors (AGRs) developed mainly in the United Kingdom during 1950s and 1960s, of which to date 50 plants have been operated. The increase in coolant gas exit temperatures from 675°C in the AGR to temperatures in the range 750-950°C in the HTR was achieved by

the use of a completely ceramic reactor core; the use of helium instead of carbon dioxide as the coolant medium; the use of ccated particle fuel.

Table 1 summarizes the experience gained world-wide with gas-cooled reactors. The Magnox and AGR have provided about 1000 reactor-years of operation and the HTR about 50. Gas-cooled reactor experience is about 25 % of the total experience with nuclear reactors, but because of the small size of such plants compared with water-cooled reactors, the proportion of electricity produced by gas-cooled reactors is only just over 10 % of the total nuclear electricity generation.

Medium-sized prototype steam-cycle HTRs have been ccnstructed in the USA and the FRG; in USA the Fort St. Vrain Reactor which uses block- type fuel elements has been in operation since 1969 and in Germany on the basis of the AVR1) experience the Thorium High Temperature Reactor (THTR-300)2) with a pebble-bed core design and spherical fuel elements began electricity production in 1985. There are two recent designs for the next stage in the HTR development; one is the medium-power plant HTR-500 with a rating of 1390 MWtnermal and the second, a small plant (HTR-Modul) with a rating of 200 MW^hermal* In the HTR-Modul concept, active safety systems are to a large extent unnecessary because of the inherently safe design /1-3/.

1) AVR = Arbeitsgemeinschaft Versuchsreaktor

2) THTR 300 = Thorium High Temperature Reactor with 300 MWej 61

The construction of HTR test reactors is planned in Japan and the People's Republic of China and a demonstration plant (German HTR-Modul concept) in the USSR. • Both China and the USSR have recently signed contracts with the FRG for the development of the HTR.

2. Characteristic Features of the German HTRs

A unique feature of the German HTR designs is the form of the fuel elements. They are spheres of 60 mm diameter and consist of up to 40.000 coated fuel particles of (U,Th)02 or UO2 embedded in a graphite matrix. Each fuel particle is coated with either three layers of pyrocarbon (BISO particles) or three layers of pyrocarbon with an additional SiC layer (35 /urn thick) (TRISO particles). The pyrocarbon coatings are more than a hundred /urn thick. The coatings on the particles provide the first fission product barrier (see Figure 1), whose integrity is maintained even at very high temperatures. Metallic fission products are retained up to approx. 16C0°C and the remaining fission products up to "* approx. 2000°C. In contrast, the maximum operating temperature of the fuel elements is 1250°C.

The graphitic matrix material protects the fuel against mechanical influences. Furthermore, the intimate contact between the fuel particles and the matrix material, which has a high heat capacity and good thermal conductivity-, provides protection against overheating.

Regarding the fuel itself, the reference design up to 1979 was (U,Th)02 with 93 % enrichment of 235y ^\j (high enriched uranium) fuel, using the BISO coating concept. As a consequence of uncertainties in long- term HEU supply following the INFCE (International Fuel Cycle Evalua- tion) discussions, the design was switched to UO2 fuel of 10 % enrichment (LEU: low-enriched uranium fuel) using the TRISO coating concept.

The high total heat capacity of the pebble-bed core and the graphite reflector surrounding it and the low power density permit only a very slow temperature rise in the case of cooling failure. Many hours therefore pass before the individual fuel elements are heated to such an extent that a measurable release of fission products occurs. 62

The graphitic spherical fuel elements and the reflector retain their shape in a helium atmosphere up to temperatures of 2800°C. Such' temperatures are not reached even during a long-term failure of the cooling. A core melt-down can be ruled out.

Other characteristics of the HTR design are:

The temperature coefficient of the reactivity is negative throughout the entire temperature range. An increase in reactor core temperature, therefore, reduces the reactivity. This effect can be used for a self-shutdown by reducing reactor cooling.

The helium coolant does not react with the fuel particles, the graphite or the structural materials; it is chemically inert. Chemical reactions are only possible with the impurities H2O, H£, CO, CO2 and CH4 in the reactor helium. Furthermore, helium has no influence whatsoever on the neutron economy.

In the case of depressurization, the density of the coolant merely decreases. Cooling of the reactor can be ensured by residual removal systems, even in the depressurized state.

3. HTR concepts in the Federal Republic of Germany

The HTR with pebble-bed core and the fast breeder reactor are the most important advanced reactors systems in the FRG. The HTR has been especially developed and constructed in our country since the early sixties, supported by extensive research funds. The HTR has a variety of potential applications: apart from pure electricity generation, it has favourable features for process steam production and district heating and is also suitable for producing chemical process heat which can be supplied at a temperature level of up to 950°C for the refining of fossil energy carriers (e.g. coal gasification).

3.1 AVR, the origin of the German HTR development

The experimental AVR facility (46 MW^h» 15 MWej) became operational in the late sixties. After more than 21 years of successful operation this 63

test reactor was shut-down at the end of 1988, having produced about 1.7 Twh of electricity. In addition to the electricity production with an excellent availability, the AVR has been used as a fuel element test bed. More than 15 different fuel element types have been successfully tested. The AVR has also provided a research facility for fission product release. The reactor has operated with the helium gas outlet temperature of 950°C for more than 15 years and has thus demonstrated the suitability of the advanced HTR for process heat applications. The current coolant gas activity of the AVR before shutdown was been approximately 5.5 x,1011 Bq ( 15 Ci) for all nuclides. During normal operation 7.4 x 1011 to 1.1 x 1012 Bq (20 to 30 Ci) of noble gases, a few tenths of a megabecquerel (a few microcuries) of. aerosols and very low non-measurable amounts of iodine have been released annually into the environment /4/. The AVR experiences are of great value for the further development of modern HTR designs. Note that safety problems have never occurred, although operational disturbances have frequently resulted in- prolonged shut-down periods. The principal disturbances were ingresses of water into the core initiated by steam generator leaks or sealing failures in the water-helium system. Damage caused by considerable amounts of water entering the core remained minor and was not of any significance regarding safety. In the AVR, for example, 25 to 30 m3 of water entered the primary circuit via a leak in the steam generator, but no appreciable corrosion on the fuel elements or the reflector graphite occurred.

Before shut-down, the AVR had been used for a comprehensive test programme including simulation of depressurization accidents and validation of nuclear and thermohydraulic codes.

3.2 THTR 300, a steam-cycle prototype reactor

Based on the favourable operational experience with the pebble-bed core concept as demonstrated during the successful operation of the AVR, the THTR 300 was designed and constructed by BBC3)/HRB4) (Fig. 2). It has a power rating of 750 MW-th providing about 300 MWei output /5/.

3) BBC=Brown Boveri & Cie 4) HRB = Hochtemperatur-Reaktorbau GmbH 64

The heat is generated, as in the AVR, in a pebble-bed of spherical fuel elements, with an average power density of 6 MW/m^. The dimensions of the cylindrical core are 5.6 m diameter, 5.1 m height.

The coolant helium flows through the reactor core from top to bottom and is heated up from 250 to 750 °C. Six steam generators, arranged around the reactor core, produce superheated steam at 545 °C and 190 bar, as in conventional coal-fired power stations. The steam generators are also used for residual heat removal from the coolant circuit when the reactor is shut down. The whole primary circuit of the THTR 300 is contained within a prestressed concrete pressure vessel. The operating pressure is 40 bar.

The commissioning of the THTR 300 was carried out in several steps, in which the core physics and thermohydraulic design as well as the function of all operating and safety systems were intensively tested. All the significant design parameters were thereby verified. In particular, the tests carried out at the high operational temperature, such as cyclic stressing, failure of individual components, rapid shutdown and residual heat removal procedures for different operating conditions, confirmed the calculated and approved design values.

Special safety tests for power stabi1ization using the negative temperature coefficient of the reactivity and for the investigation of residual heat removal by natural convection have confirmed the inherent safety characteristics of the THTR 300.

Electricity generation began in 1985, and following extensive commissioning trials the plant was handed over to the utility on June 1, 1987. A total electrical energy of 2.9 Twh had been generated up to the end of 1988.

Turning to the operating experience during the commissioning phase and operating time of the THTR-300 the following events were discussed in the RSK5 /6/:

5) RSK = Reactor Safety Commission 65

Relatively frequent automatic shut-down with residual heat removal via the residual heat removal system. At the very beginning this event occurred rather frequently. It could be avoided by resetting the limits on the basis of commissioning experience.

Increased occurrence of damaged fuel element spheres beyond what had been anticipated. Because of trials during the commissioning phase of the reactor, the core rods were inserted into the core more frequently than is normally the case, and the core had undergone a certain densification as an insufficient number of spheres had been circulated in the meantime. This meant that some spheres had been subjected to unusually high stresses. The withdrawal of the damaged elements took more time than had originally been anticipated since the operating elements traverse the reactor core more quickly in the centre, and more slowly towards the rim, than had been calculated. The number of damaged elements is at the moment less than 0.6 % per pass.

Difficulties in withdrawing spheres from the reactor. Because of flow-induced forces, the removal of fuel elements did not operate as expected especially under full power conditions. By changing some construction details this effect was eliminated. Meanwhile it is possible to remove the anticipated number of fuel elements at any reactor power condition..

Rupture of screw heads on some cover plates of the hot duct insulation system. Within the scope of the current phase of inspection, the licensee has examined the six hot gas channels and parts of the hot gas plenum by means of a probe. In doing so, it was found that the heads of some INCOLOY 800 central bolts and of a few corner bolts, which are used for the fixation of metallic insulation packages, were missing. This effect is based on high design stresses in some bolts and the reduced ductility of these INCOLOY 800 material caused by thermal neutron fluence. While this phenomena is still under consideration, it can be concluded that the failure of some bolts is not safety relevant. 66

All the problems to date have been related only to operational aspects and have not led to additional, more extensive, safety related measures. With regard to future HTR plants the operation of the THTR is being considered in the context of optimizing detail constructions and operational procedures.

3.3 HTR-concepts under planning in FRG

In the Federal Republic of Germany there are two industrial companies developing and offering HTR plants, Hochtemperatur-Reaktorbau GmbH (HRB) belonging to the ASEA Brown Boveri (ABB) group and the INTERATOM company which is part of the SIEMENS/KWU group. The two companies intend to form a "HTR-GmbH" for the marketing of HTR plants. At the present time, two HTR concepts are being followed up, a small HTR (approx. 200 MW^h) with a steel pressure vessel and a medium-sized

reactor (approx. 550 MWei) with a prestressed concrete pressure vessel (PCRV) for electricity generation.

3.3.1 HTR 500 follow-up to the THTR 300

The HTR 500 has a thermal power of 1390 MWth at about 550 MWej /7-9/. It is planned as a plant for electric power generation as a logical further development of the THTR 300.

In 1983, the HRB company began work on the technical design. The safety analysis report is currently in preparation. Fig. 3 shows a schematic view of the reactor, whose major features are: pebble-bed core 7 m in diameter and 5.4 m in height power density 6.5 MW/m3, primary circuit pressure 55 bar U235 enrichment approx. 9 %, burn-up ~100000 MWd/t helium temperatures inlet 260°C, outlet 700°C two independent shut-down systems with simplified drive mechanisms are provided, one freely dropping shut-down rods arranged in the reflector for control and hot-shut-down, the other rods inserted into the reactor core for long-time shutdown 67

prestressed concrete reactor pressure vessel (PCRV) with a gastight insulated steel liner, water-cooled by the Liner Cooling System (LCS) six steam generators integrated within the vessel - • two-train independent auxiliary system for after-heat removal (AHRS) making use of natural convection under operating pressure; producing adequate cooling of the core even with a depressurized primary circuit. The allowable failure times for the AHRS and LCS are estimated at 5 and 10 hours, respectively.

As for all modern HTR concepts, the reactor pressure vessel is accommodated in a reactor building which provides protection against external impacts. The reactor system was designed for safety reasons in such a way that the fuel elements, core geometry and reactor pressure vessel cannot be destroyed even in the case of the worst hypothetical accident. In the case of failure of all active cooling measures and with a, depressurized reactor, a small fraction of the fuel elements reach temperatures of near to 2500°C. The structure of the fuel elements remains stable at these temperatures. The fission products discharging into the primary circuit upon failure of fuel particles within the graphite matrix may in part escape into the reactor containment through leaks. However, since this release takes place in the depressurized state, the leak rates are small. It is possible to retain fission products in the containment by suitable filter systems so that environmental pollution can be kept within strict limits.

3.3.2 The HTR Modul, a reactor system base on inherent safety properties

The development of the HTR-Modul with a thermal power of 200 MW^h (or 170 MWth if used for process heat) was begun by Siemens/Interatom in 1979 /10-12/. A detailed report on the technical design and safety features was presented in 1983. Fig. 4 shows a schematic view of the HTR-Modul. The main features of this reactor concept are: pebble-bed core, 3 m in diameter and 9.4 m in height power density 3 MW-^h/m^, primary circuit pressure 60 bar 68

U235 enrichment approx. 8 %, burn-up 80000 MWd/t - helium temperatures inlet 250°C, outlet 700°C (or 300°C and 950°C if used for process heat) two independent shut-down systems are exclusively acting within the reflector region to avoid the need for absorber rods to penetrate the pebble bed. the core and steam generator are housed in ferritic steel pressure vessels in a side-by-side arrangement connected by a coaxial duct of steel, all with the same high standard of quality assurance, operational residual heat removal via start-up and shut-down circuits in the case of failure of the steam generator, passive residual heat removal from the reactor by convection, thermal conduction and radiation to surface coolers (3 x 100 %) positioned outside the pressure vessel. This system is not of relevance for safety, but only for the protection of components.

The vessels of the primary system are accommodated in a reactor building for protection against external impacts.

The major feature of the safety concept is the limitation of the maximum fuel temperatures even in the case of hypothetical accidents to about 1600°C by means of the selected design and geometry. The low-level positioning of the steam generator prevents a direct ingress of water into the reactor core in the case of any leakages.

4. Behaviour of HTR under severe accident conditions

The behaviour of HTR after loss of core cooling was carefully analysed in recent years. The fission product release during core heat up into the environment was of great interest. The behaviour of irradiated spherical fuel elements at elevated temperatures was investigated by annealing ramp tests (50 °C/h) in the range of 1200 to 2500 °C and by transient release tests with high precision in the range of 1400 to 1800 °C /13/. From the tests of 13 TRISO fuel elements (16000 coated particles each) up to 2500 °C and another 9 fuel elements at 1600 °C for up to 1000 h or 1800 °C for up to 100 h (Fig. 5) it has been shown that 69

for temperatures *^ 1600 °C all safety relevant fission products are retained completely; for temperature p- 2100 °C the SiC layer begins to decompose thermally while pyrocarbon layers remain intact; at temperatures > 1700 °C the SiC layer can be damaged by corrosive attack of fission products.

The retention capability of different kinds of graphite was investigated mainly for caesium and strontium by static (adsorption isotherms) /14,15/ and kinetic experiments (vaporization and sorption by graphite samples in the range of 800 to 1000 °C) /16/. From these experiments it is known that graphite surfaces of cooler fuel elements and reflectors would almost completely retain strontium, if released from fuel elements under severe accident conditions. Caesium is also significantly retained, while the retention of iodine and of noble gases is marginal.

The removal of after-heat from the core is, in the first instance, undertaken by the main cooling system; if this fails, then a separate residual heat removal system and, in part,, also component cool ing systems (e.g., PCRV liner cooling system) can take over this function as active safety devices. The reactor is or can also be designed in such a way that natural convection, thermal conduction, and radiation are sufficient for heat removal.

The first kind of heat removal is. planned for HTR 500. The second concept is verified for the HTR-ModuI. Should the after-heat removal systems fail, a core heat-up accident would result. Because with increasing core temperature heat removal is improved and after-heat production continuously decreases, the core temperature would go through a maximum, and a subsequent slow decrease. The magnitude and the duration of the maximum temperature stage largely depend on the pressure level and the design of the plant. The dimensions of the core are essential factors because they determine the surface- to-volume ratio of the core. Fig. 6 shows the hot-spot temperatures calculated for current German HTRs /17/: 70

For pressurized reactors, the temperatures remain below 1600°C even for medium HTRs (HTR 500) as a result of core-internal natural convection. The highest values are to be expected for depressurized reactors, remaining in the range of 1600°C for all small HTRs but possibly rising to 2600°C for medium-sized HTRs. The cooling of the reactor cell for the small HTR, as well as liner cooling for the medium HTR, ensures that serious damage to the vital components can be ruled out and that the reactor can gradually be cooled down again in a controlled manner. The influence of these cooling systems on the maximum hot-spot temperatures is negligible.

The HTR therefore offers the possibility of limiting the core temperatures, even without (active) cooling systems, in such a way that, in comparison with normal operation, the region of significantly increased fission-product release is not reached at all or at least not significantly exceeded. This has been realized in the HTR-Modul. However, for the assessment of the consequences of a core heat-up, not only are the expected maximum hot-spot temperatures of significance but also the temporal sequence of temperatures and their radial and axial profiles. In the HTR-500, 5 h after the failure of after-heat removal, less than 10 % of the fuel elements would exceed the maximum operating temperatures of 1100°C; more than four days after the start of the accident, only about 50 % of the fuel elements exceed temperatures of 1600°C and about 15 % temperatures of 2000 °C. A greatly delayed release of fission products from the core results from the slow development of the temperatures and their distribution. On the basis of their physical properties, noble gases and halogens are released first; metallic fission products follow after a time lag.

5. Licensing procedures

At the moment there are no firm plans to erect a HTR plant in FRG, but the Siemens/Interatom consortium have initiated licensing steps for the HTR-Modul. 71

The formal licensing procedure is based upon the German Atomic Energy Act /18/. It describes the necessary requirements for granting a licence so that persons and property are protected against the hazards of nuclear radiation.

The key part in the licensing procedure is' played by the nuclear regulatory commission of the State in which the nuclear plant is to be constructed. However, the individual States (Bundesländer) do not implement the licensing procedure autonomously but rather on behalf of the Federal Ministry of the Environment, Nature Conservation and Reactor Safety (BMU) of the Federal Republic of Germany. This means in essence that the BMU ensures a uniform implementation of the licensing procedures in the various States. The authorities of the States and the BMU engage experts to assess the application documents.

As the most important advisory bodies, the Reactor Safety Commission (RSK) and the Radiation Protection Commission (SSK) act on behalf of the BMU. Licenses for construction and operation are generally granted in separate stages as so-called partial construction or partial operating licenses. The first partial construction licence according to § 7 of the Atomic Energy Act normally approves the decision of the site for construction of the nuclear power station after thorough inspection and investigation. A positive decision about the overall concept of the plant should have already been made at this point in time. However, pursuant to § 7a of the Atomic Energy Act, an interim decision may be issued by the licensing authority on application with respect to the individual issues decisive for granting a licence for a facility.

In April 1987 the procedure was begun with an application to the Ministry of the Environment of the State of Lower Saxony, as the competent licensing authority, for a site-independent concept licence for the HTR-Modul.

It should be clearly stated that the granting of the concept licence cannot replace or by-pass the 1icensing procedure necessary to construct and operate a plant at a given place. The presence of the concept licence can only facilitate this procedure. 72

References IM Kroger, W.; Nickel, H.; Schulten, R. Safety Characteristics of Modern High-Temperature Reactors: Focus on German Designs, Nuclear Safety 29 (1988) 36-48 111 Kroger, W.; Schulten, R. Technology and Safety Aspects of German HTRGR, Proc. ANS Int. Topical Meeting on Safety of Next Generation Power Reactors, Seattle, May 1-5, 1988 131 Nickel, H.; Hofmann, K.; Wachholz, W.; Weisbrodt, I. The Helium-Cooled High-Temperature Reactor in the Federal Republic of Germany: Safety Features, Integrity Concept, Outlook for Design Codes and Licensing Procedures, Symposium on Regulatory Practices and Safety Standards for Nuclear Power Plants, IAEA-Symposium, München, Nov. 7.-10., 1988, IAEA-SM-307/31 /4/ Marnet,C; Ziermann, E. AVR Operation and Experience, Proc. Seventh Int. Conf. on High- Temperature Gas-Cooled Reactors, Dortmund, September 1985, VGB-Kraftwerkstechnik GmbH, VGB-TB 111 (1985) 174-181

15/ Wittchow, I.; Baust, E.; Scheming, J. Status of the Construction of the THTR-300 MW and the Design of a 500 MW Follow- on Plant, Nucl. Eng. and Design 78 (1984) 109-117

/6/ Results of the Safety Review of Nuclear Power Plants in the Federal Republic of Germany, Recommendation by the Reactor Safety Commission (RSK), Nov. 23, 1988

111 Baust, E.; Schöning, J. The Basic Design for Commercial HTR Power Stations, Proc. IAEA TC-Meeting on Gas-Cooled Reactors and their Applications, IAEA-TEC DOC 436, Oct. 1986, 235-252

181 Wachholz, W. The Safety Characteristics of the HTR 500 Reactor Plant, Nuclear Engineering and Design, 109 (1988) 307-312 191 Arndt, E. et al. HTR 500 - Design, Safety and Project Status fo the Medium-Sized Commercial High-Temperature Reactor, Nuclear Engineering International, 33 (1988), 20-28

/10/ Reutler, H.; Lohnert, G.H. Advantages of going Modular in HTRs, Nuclear Engineering and Design, 78 (1984), 129-136

/11/ Frever, H.; Keller, W.; Pruschek, R. The Modular High-Temperature Reactor, Nucl. Science and Engineering, 90 (1985) 411-426 73

IMI Weisbrodt, I. Inherent Safety Design Features of the HTR-Module, Int. Conf. on Nuclear Power Performance and Safety, IAEA, Vienna, September/October 1987 /13/ Schenk, W.; Pitzer, D.; Nabielek, H. Fission Product Release Profiles from Spherical HTR Fuel Elements at Accident Temperatures, Report Jül-2234, 1988 /14/ Hilpert, K.; Gerads, H.; Kath, D.; Kobertz, D.; Moormann, R.; Verfondern, K. Retention of Metallic Fission Products by Grahite Materials in the HTR-Core-A Key Effect for the Reactor Safety, Proceedings, German Nuclear Society, Karlsruhe, June 1987, 319-322, Deutsches Atomforum e.V. Bonn 1987 /15/ Hilpert, K.; Gerads, H.; Kath, D.; Kobertz, D. Sorption of Caesium and its Vaporization from Graphitic Materials at High Temperatures, High Temp.-High Pressures 20 (1988) 157-164 /16/ Iniotakis, N.; Bartheis, H.; von der Decken, C.-ß. The Retardation of the Release of Fission Products deriving from Graphite Structures during a Hypothetical Accident of HTR-500, Seminar on Small and Medium-Sized Nuclear Reactors, Lausanne, August, 1987 /17/ Rehm, W.; Jahn, W.; Verfondern, K. Safety Analysis of Core Cooling Accidents in Small and Medium-Siced HTRs, Brennst.-Wärme-Kraft 37 (1985) 279-282 /18/ Gesetz über die friedliche Verwendung der Kernenergie und den Schutz gegen ihre Gefahren (Atomgesetz), Act on the peaceful use of nuclear energy and the protection against its hazards (Atomic Energy Act), Publication of the revised text of the Atomic Energy Act of 15.07.1985 with the amendment on the Federal Law Gazette I no. 8 of 21.02.1986

75.

Graphite

shell

motrix Primary circuit |

coated porticie

1 Fruthon of defective particles SIO"* operational tempe- pyro- . rature < 1250 °C « • carbon 1 Increasing perme- ability >I600°C total particle failure > 2500 °C

Mechomcal stability

Fig. 1: Fission Product Barriers and Retention Properties of Fuel Elements for HTGR

water ^ steam

core rods coolant gas

helium

reactor core prestressed steam generator concrete vessel

graphite reflector

sphere discharge tube

Fig. 2: Sectional View of the THTR 300 Reactor Pressure Vessel (schematic)

Discussion of the presentation

Present status of HTR in FRG

Huber, Coionco, Switzerland: Some years ago a discussion about "walk away" reactors started. This is a reactor which can be left to itself even under severe accident conditions. No action of the operators is necessary e.g. to start some active safety systems. Can the HTR-Modul be designated as such a "walk away" reactor?

Nickel, KFA Julien, FRG: The idea behind the HTR Modul concept is the limitation of maximum core temperature to 1600° even in hypothetical core heat up accidents. By experiments and tests at AVR and at our institute it is well established that the fuel elements, especially the coated particles, retain the fission products. So after a core heat up accident not only does the core geometry remain intact but also the fuel elements do not lose their retention properties. So we can certainly say that we have in some sense here a "walk away" reactor.

Alder, PSI Wiireniinqen, Switzerland: Could you give some more information about the water penetration into the core of the HTR-Modul?

Nickel, KFA JÜlich, FRG: Due to the staggered positioning of reactor core and steam generator a very large ingress of water into the core is inherently ruled out. Hence, a flooding of the core is not possible.

For the design basic accident a guillotine rupture of one steam generator tube is assumed. After detection of moisture the steam generator is dumped in approximately 60 seconds. The total steam 79 ingress into the primary circuit was calculated to be less than 600 kg. If the gas purification plant works properly the ingressed steam will be removed from the primary circuit in less than 2 hours. If, however, the gas purification plant fails completeley the ingressed steam will be transformed into watergas. Due to the pressure increase, the safety valve will open after approximately 5 hours and will relief directly into the reactor building. It can easily be shown that the mixture of helium, air and watergas is not inflammable in any circumstances. If the filters of the reactor building are available the dose rate to the environment (tyhroid, infant) had been conservatively calculated to amount to 1.26 mrem. If the filters are postulated to be failing the dose rate increases to 170 mrem (best estimate: 17 mrem). 81

Section II:

Technical safety boundary conditions 83

HTR SAFETY FEATURES AND THE INTEGRITY CONCEPT

by J. Wolters Institute for Nuclear Safety Research Nuclear Research Centre Jiilich, FRG

1. INTRODUCTION

The objective of this paper is to delineate briefly the accident behaviour of the HTR in the case of the structural failure of those components which serve for the safe enclosure of the radioactive substances within the plant. This is the basis for the formulation of the general requirements for the structural integrity of these components. The requirements together with the measures necessary to ensure an appropriate quality of the components form what we understand as the integrity concept /I/. It is part of the HTR-specific safety concept.

2. ACCIDENT BEHAVIOUR OF THE HTR IN THE EVENT OF STRUCTURAL FAILURES

The measure for the decision whether a structural failure of a component serving for the safe enclosure of radioactive substances is tolerable in the framework of design-relevant accidents is the radiological consequences resulting from such a failure. In order to understand the requirements for the structural integrity it is thus necessary to explain what radiological consequences are to be expected in the event of a postulated structural failure of the components in question. For this purpose . it is expedient to distinguish between the following three groups of components (analogous with the integrity concept): 84

- components with a barrier function,

- components whose failure would impair the barrier function, - components whose failure would result in an ingress of reactive gases into the intact primary circuit.

2.1 Components with the Function of Fission Product Barrier

The three 'classical' fission product barriers of the HTR are: the coating of the fuel particles, the pressure boundary of the primary circuit and the reactor protection building, which is designed as a confinement. Fig. 1 shows schematically the arrangement of these components for the HTR-500 121, a medium- sized HTR with a prestressed concrete pressure vessel (PCRV), and Fig. 2 for the HTR-Modul /3/, one of the two German concepts for a small HTR with a steel pressure vessel.

The first fission product barrier is the most significant barrier since it encloses the total fission product inventory. About 15,000 particles are contained in one fuel sphere. They are embedded in a graphite matrix, which itself is encompassed by a fuel-free shell of graphite of 5 mm thickness. The particles are thus well protected against mechanical and chemical attacks from the outside. The coating of the particles consists of two layers of pyrocarbon with an intermediate layer of silicon carbon. The fabrication process of this triso particle and the fuel element has meanwhile reached such a high standard that less than 5 particles with a defective coating are to be expected in ten fuel spheres. To this figure a maximum of 8 particles whose coating may be damaged by irradiation are to be added. Together with the uranium contamination of the particle surface the fraction of uranium not enclosed by the coating may vary between 6-10"5 and 1-10"*.

The uranium not enclosed determines the quantity of fission products released from the fuel elements during operational states. Since the fraction itself is low and since there is an effective retention mechanism in the kernels of the particles and in the graphite the release rates during operational sta- 85

tes are very low. In addition, all the fission products except the noble gases are deposited on the cold surfaces of the primary circuit on the way from the core via the heat exchangers to the cold gas side. Consequently, the contamination of the coolant, particularly the cold gas which encloses the hot gas, is very low. Apart from the small content of He-3 the coolant itself is not activated. It is therefore possible to release this gas into the environment in the event of a structural failure of the pressure boundary without reaching the limits of the German Radiation Protection Ordinance /4/ for design- relevant accidents. The release of contaminated dust may contribute significantly to the radiation exposure.

An increase of the failure fraction of the first fission product barrier on its own does not influence the integrity of the other barriers or the function of safety equipment such as the shut-down system and the afterheat removal system. But the higher contamination of the coolant and of the primary circuit components would result in an increase of exposure values of the site personnel and the environment during operational states and in accident conditions. It is therefore not only a question of safety but also of plant performance and availability to maintain the structural integrity of the first fission product barrier during operational states.

The significance of a structural failure of the second fission product barrier, the pressure boundary of the primary circuit, depends on how this would affect the integrity of the first and third barrier. A direct effect on the first barrier can be excluded since the thermal and mechanical loads of the fuel elements associated with the depressurization of the primary circuit are very low. This is a consequence of the low power of a fuel sphere compared with a high heat capacity, the low temperature gradient in the sphere and the use of a phase- stable and compressible coolant. Even a loss of forced cooling would result only in a slow increase of fuel temperature. Since the coolant is not lost but only reduced in, density afterheat removal by forced cooling is still possible. 86

The HTR-500 relies on forced cooling. It is therefore equipped with two separate afterheat removal loops which are designed such that even at atmospheric pressure one loop is sufficient to cool down the core (Fig. 3). In this case the fuel temperatures do not exceed the design values of the operational states and the structural integrity of the first barrier is completely maintained. But this cannot be guaranteed for large-scale failures of the pressure boundary because this may result in a loss of forced cooling due to damage of components inside the primary circuit. A core heat-up would follow during which the temperatures of the fuel elements reach very high values (Fig. 3). In spite of the high temperature stability of the coating this barrier becomes partially defective or per- meable for metallic fission products so that significant quantities of fission products are released from the fuel (Fig. 4) . Although there is a considerable reduction of the quantities on the way from the core via the primary circuit and the reactor building /6/ the release values are too high to be tolerable in the framework of a design relevant accident.

A second reason that a large-scale failure must be avoided is the hazard of a massive ingress of air by natural convection. In spite of the cooling of the core this may result in corrosion-induced damage of the first barrier by a self- sustaining chemical reaction of the air with the hot graphite of the core. Of course this is a question of the ingress rate which is a function of the size and the orientation of the leak. From the experimental study of the counter-current flow in a leak of different sizes and orientations /If we know that only the failure of the steam generator penetration in the PCRV ceiling could lead to ingress rates with the potential of corrosion-induced fission product release from the fuel (Fig. 5). Therefore such a structural failure is not tolerable.

The third reason for the prevention of large-scale failures of the pressure boundary is the resulting thermal and mechanical load on the reactor protection building and its inner struc- 87 ture. The building and its relief system which protects the building from excess internal pressure and which blows off directly into the stack (Fig. 1) are designed to cope with a leak size corresponding to the cross-section of a fuel sphere charge pipe. Although this is a question of design one very quickly encounters limits of economy and feasibility, not to mention the problems associated with the impact of gas jets and missiles. The HTR-Modul does not need forced cooling for afterheat removal to protect the first barrier from being damaged by excess core temperatures. It is designed such that the .heat removal by heat conduction and radiation via the surface of the pressure vessel to a surrounding surface cooler is sufficient to prevent excess fuel temperatures under all accident conditions /5/. In the case of a depressurization accident which results in the highest core temperatures the maximum fuel temperature does not exceed 1600 °C. This high temperature may cause an increase in the number of particles with defective coating but due to the fact that less than half of the core reaches a temperature above the design value of the operational states the total fraction of defective particles remains well below 1-10"3. This in combination with the retention of metallic fission products on cooler surfaces of the core and in the primary circuit yields, release quantities which can be released into the environment via the stack without exceeding the exposure limits..

Nevertheless a large-scale failure of the pressure boundary must also be avoided with the HTR-Modul .The reasons are practically the same as for the HTR-500. Although for instance damage to the surface coolers would not influence the level of the core temperatures such an event is not acceptable as a design bases accident.

The requirements for the structural integrity of the first group of components are listed in Table 1. The reactor protection building is only demanded as a confinement in the case of a depressurization accident. A structural failure need 88

not be postulated in the framework of the single failure concept if the necessary proofs are provided. The strength of the building is determined by its function of protecting the nuclear part from external impacts.

2.2 Components whose Failure would Impair the Barrier Function

A second group includes the shut-down rods, the blowers, the heat exchangers of the afterheat removal system and the graphitic structures. It is clear that they must remain operational and structurally stable in normal operation and under accident conditions. A random structural failure of a safety equipment must be and is covered by redundancy.

2.3 Components whose Failure would Result in an Ingress of Reactive Gases into the Intact Primary Circuit

The third group of components include the steam generators of the HTR-500 and the HTR Modul. The implications of a structural failure of such a component depend among other aspects on the quantity of reactive gases entering the primary circuit and on the volume of the circuit. The quantity is determined by the leak size and by the countermeasures initiated by the plant protection system. In general, the safety concept is based on the postulated double-ended fracture of a single heating tube in such a component. Even with the small volume of the HTR-Modul primary circuit no problems are experienced in coping with such an event /8/.

The total amount of water entering the circuit is limited to 600 kg/s by the isolation and dump of the steam generator. This amount of water increases the pressure in the primary circuit by 3 bar. A pressure build-up to the set-point of the relief system is only to be expected if the gas purification plant which controls the pressure and a special gas circulation system which extracts the steam by condensation are not 89

available (Fig. 6). This event sequence is considered as a design relevant accident. It would result in the release of 10 % of the gas inventory of the primary circuit via the stack into the environment. Besides the normal contamination, the gas is laden with the fission products stripped from the surface of the steam generator by liquid water and steam. There- fore the surface contamination of the steam generator is a significant aspect. Yet an open question is whether there is an increase . in iodine release by the hydrolysis of defective particles. But even if we take this into account the exposure values remain below the limits though without filtering the thyroid dose of a child comes very close to the limit.

The corrosion degree of the fuel spheres reaches a maximum value of 1,5 w/o and is thus negligible.

Larger leaks would occur in the case of a structural failure of the steam or feed water manifold integrated into the primary circuit. In this case the pressure may directly reach the set-point of the safety valve (second relief system, Fig. 2) but will not exceed the design pressure of the vessel /5/. But the main hazard does not originate from the large quantity of water in the primary circuit as revealed by a worst case consideration of a leak in the steam manifold of a size corresponding to the cross-section of all the heating tubes /5/. The local effects on the pressure boundary caused by jets and whipping tubes are more serious. It is therefore obvious that such a structural failure must be excluded.

The requirements for the structural integrity of these components are also listed in Table 1. The accident conditions for which they must be designed include pipe ruptures on the secondary side. 90

3. QUALITY AND QUALIFICATION MEASURES

In accordance with the requirements made, measures must be taken to ensure an appropriate quality of the components. Where applicable the measures of the KTA rules 3201 and 3211 are used. This is for instance the case for the steel pressure vessels of the small HTR (HTR-Modul and HTR-100) and for the penetrations of the Prestressed Concrete Pressure Vessel (PCRV) of a medium-sized HTR. Conversely special measures are defined for the concrete body of the PCRV, the prestressing system and the steel liner with its thermal insulation. A KTA rule and a DIN-standard for the PCRV are in preparation.

Special measures are also defined for the components of the third group when they are operated at a higher temperature than 360 °C. An optimized heat-resistant steel according to HTR specification is to be used for these components (X20CrMoV 121, XlONiCrAlTi 3220). The latter has also been qualified together with NiCr22Col2Mo for the heat-exchanging components of a nuclear process heat plant. The determination of dimensions is based on time dependent design values. For specific parts thermal stresses are already taken into account in dimensioning. The main principle is the limitation of plastic strain and the avoidance of local hardening in order to prevent crack initiation. For some components it should be confirmed that a postulated flaw of a size at the detectable limit cannot grow to a critical size where spontaneous failure must be expected.

More details are included in the special papers which will be presented later. 91

4. REFERENCES

tit Auslegungskriterien für hochtemperaturbelastete metallische und keramische Komponenten sowie des Spannbeton-Re- aktordruckbehälters zukünftiger HTR-Anlagen. Band I, Teil A: Sicherheitstechnische Randbedingungen Kernforschungsanlage Julien, Institut für Reaktorwerk- stoffe, August 1988.

/2/ Wachholz, W. • Das Sicherheitskonzept des HTR-500. Tagungsbericht der Fachtagung 'Sicherheit von Hochtempera- turreaktoren' am 19. und 20. März 1985, Jül-Conf.-53, Juni 1985.

/3/ Hübel, H.; Lohnert, G. Das Sicherheitskonzept des HTR-Modul, veranschaulicht am Beispiel des Wassereinbruchs in den Primärkreislauf. Tagungsbericht der Fachtagung 'Sicherheit von Hochtempera- turreaktoren' am 19. und 20. März 1985, Jül-Conf.-53, Juni 1985.

/4/ Radiation Protection Ordinance (Verordnung über den Schutz vor Schäden durch ionisierende Strahlen (Strahlenschutz- verordnung) . Das Deutsche Bundesrecht, 9. Auflage, Nomos Verlagsgesellschaft, Baden-Baden, 1982.

/5/ KFA/ISF • Zum Störfallverhalten des HTR-500. Eine Trendanalyse. Jül-Spez-260, Juni 1984.

/6/ KFA-ISF Zum Störfallverhalten des HTR-500. Eine Trendanalyse. Jül-Spez-220, September 1983. flf Breitbach, G.; David, H.P.; Nickel, M.; Wolters, J. Ausström- und Gasaustauschvorgänge nach Lecks im Primärkreislauf von Hochtemperaturreaktoren. Jül-Spez-469, Oktober 1988.

/8/ Wolters, J.; Bongartz, R.; Jahn, W.; Moormann, R. The Significance of Water Ingress Accidents in Small HTRs. Nuclear Engineering and Design 109 (1.988) 289-294, North- Holland, Amsterdam. 92

GENERAL REQUIREMENTS FOR THE STRUCTURAL INTEGRITY

GROUP I: COMPONENTS WITH A BARRIER FUNCTION

. COATED PARTICLES: - THE BARRIER FUNCTION MUST BE MAINTAINED IN OPERATIONAL STATES AND UNDER DESIGN- RELEVANT ACCIDENT CONDITIONS

. PRESSURE BOUNDARY: - THE GAS TIGHTNESS MUST BE MAINTAINED IN OPERATIONAL STATES

- LARGE-SCALE FAILURES MUST BE EXCLUDED

- LEAKS MUST BE RESTRICTED IN SIZE SO THAT OTHER COMPONENTS ARE NOT UNDULY INFLUENCED

PROTECTION BUILDING: - CONTROLLED RELEASE OF LARGE LEAKAGES AND FILTERED RELEASE OF SMALL LEAKAGES OF THE PRIMARY CIRCUIT MUST BE ENSURED:

- STABILITY AND INTERNAL INTEGRITY (NO MISSILE PENETRATION, NO INNER BURST-OFF OF CONCRETE) MUST BE MAINTAINED UNDER EXTERNAL IMPACTS (AIR CRASHES, EXPLOSION SHOCKWAVES ETC.).

GROUP III: COMPONENTS WHOSE FAILURE WOULD RESULT IN AN INGRESS OF REACTIVE GASES INTO THE INTACT PRIMARY CIRCUIT

- EXTENSIVE FAILURE MUST BE RULED-OUT

- SMALL LEAKS MUST BE KEPT UNDER CONTROL

- STRUCTURAL STABILITY AND TIGHTNESS MUST BE MAINTAINED UNDER DESIGN-RELEVANT ACCIDENT CONDITIONS.

TABLE 1

S u. 9 s

«I x o ^ * 51 W ». 2 » « o m^ x u. *. & *- £

> x Si *- Ü c< 2

la ? a« U. Q. s U- Q. •» Q)O XOOl %00l

Maximum No Forced Cooling Average

Design Temperature of Coated Particles

Maximum, Forced Cooling 0 Time Ihj

Fuel Temperatures of the HTR-500 in the Case of a Depressurization Accident with and without Forced Core Cooling Fia-3 95

m— 4" __-—^^^

i 10° - Air

"D H = 2.5xD ^s^ Theory ij/l 1 \ 10 *• Vessel o> Ü x Buoyancy - "Body •* -2n A) « 10 - Helium /, / Experiment c(c0 / (0 / / © at Scales *

* / ixperimental Facility r 10""-I i iiiii 0 0.4 0.8 1.2 1.6 2.0 uiamexer imj ^ Air Ingress Rate versus Leak Diameter for a Konstant Height to Diameter Quotient of 2.5 (typical for the Steam Generator Penetration of the PCRV) Fig. 5

70- Set Point of Relief Valve Condenser only without GRA Natural Circulation without GRA Stagnation

Gas Purification Plant (GRA) 0 10 Time Ihl Time Dependent Pressure in the Primary Circuit of the HTR-Modul after an Ingress of 600 kg of Steam for Different Plant Conditions Fig.6 96

Discussion of the presentation:

HTR Safety Features and the Integrity Concept

Schulz, GRS Köln, FRG: The safety concept of the HTR relies to a large extent on the fission retention capability of the coatings. What are the testing procedures to ensure the integrity of the coatings?

Nickel, KFA Julien, FRG: First of all I should mention that since the 1960s extensive tests with fuel elements and coated particles have been carried out. In the AVR more then 15 fuel elements types have been tested and statistics about defect particles exist and quality assurance is based on the long experience gained. It is possible to produce the particles with a high quality. The test procedures are well established by NUKEM, which manufactures the fuel elements. 97

Classification, of systems and components into safety classes and

quality standard classes

M. Dette, Rheinisch Westfälischer Technischer Überwachungsverein e.V., Essen.

The objective of every component classification for nuclear power plants is to prevent accidents during the power plant's erection and operation by means of specific quality measures. Classification establishes a close linkage and weighting of the following points:

safety relevance of a component its failure probability and the effect of a possible failure in the framework of plant protection, environmental protection and work protection

The following internationally adopted procedure can be taken as a basis for licensing:

USNRC licensing decree 10CFR 50 for LWR requirement category 1 and USNRC licensing guideline 1.26 for LWR requirement categories 2 and 3 and ANSI N-182 as safety criteria for PWR plants . ANSI N-212 as safety criteria for BWR plants ANSI N-213 for HTR plants ANS-54 for SNR plants furthermore the

IAEA Safety Guide on Safety Functions and Component Classification for BWR, PWR and PTR 98

A draft KTA rule 3202.1 or, as the case may be, 3200 has been drawn up in the FRG which describes the basis for a classification of systems and components and their correlation with the type-and scope of the quality assurance measures for pressurized water reactors.

None of these regulations or drafts adequately covers the specific demands for the classification of HTR plants and so a recommendation has been drawn up by a working party for the research project "HTR design criteria" with the following aims

Special HTR features are to be taken more into account

It is to be applicable to different HTR plants, e.g. HTR 500, HTR module, HTR 100, PNP) and for different design principles.

proven procedures for classification are to be maintained.

The aim is to provide a balanced safety-related design of systems and components in which the greater the effects of a failure may be, the less the occurrence probability is to be assumed.

Economic objectives are also included in the classification scheme.

When putting these points into practice it was found advisable to evaluate the two aspects

safety technology and requirements separately and successively.

The first stage is at the same time the starting point for a specifically HTR-plant-related classification of all components into safety classes. The second stage will then consist of the establishment of requirements. 99

On the basis of a meaningful classification of the systems, components, component groups (e.g. valves, tanks etc.) or parts with a special function (e.g. liners, vessel closures, RPV internals etc.) for the entire plant (see column 1 in annex 1), every system or component is allocated a safety class (SC) in accordance with this classification. Every allocation is based on the performance of a safety analysis in which the radiological effects of a component failure for plant personnel and the environment including the possible consequential failure together with the failure of a required safety function are evaluated. The subdivision into five safety classes is therefore also based on the related limit values of the Radiation Protection Decree (StrSchV; column 2 in annex 1). It is conceivable , in particular for HTR plants of lower capacity, however, that the number of safety classes be reduced to three or even only two, and this can be concluded from these procedures.

The HTR 500 metal pressure vessel closure with simple cover head was selected as an example for the procedure for such a classification (compare safety-related tests on the accident behaviour of the HTR 500, Jül - Spez. - 240, Jan 1984). Here, a non-isolatable leak in the primary circuit is structurally . . . 2 limited to a maximum cross section of 33 cm . A consequential failure of other components was not assumed in this analysis. There were the following effects from the assumed failure of the required safety functions (shutdown, decay heat removal systems (NWA) or filtering of the reactor protection building (RSG)):

1. Shutdown

The reactor does not reach the critical temperatures for release of fission products (after-heat removal functions, recriticality is only possible in a relatively cold condition as a result of negative temperature coefficients) D < § 28.3 100

2. Decay heat removal system (NWA)

Reactor unpressurized, primary circuit open, core heating accident, release of fission products into the reactor containment, problems in the long term with strontium (Sr) even with a functioning reactor building filter section. D> § 28.3 3. Reactor building (RSG) filter

After-heat removal system functions, no increase in core temperatures with a relatively small leakage cross section as compared with a), only a slow gas exchange, filtering of the fission products. D < § 28.3

In the case of HTR-500, and under these boundary conditions, the radiological effect with a non isolatable leak in the vessel end combined with the after-heat removal failure is a major factor for classification into safety class 1 (see column 2 in table 1).

Comparable to this procedure is the classification of every system or component of the entire plant into a single safety class.

Of course the components with the highest safety relevance such as the above mentioned vessel closure, the reactor pressure vessel of steel, the main steam headers or decay heat removal heat exchangers.

Due to the wide range of permissible limits (factor greater than 100), SC 2 was subdivided into a and b; depending on the concept, the fuel element refuelling plant could be placed into this class.

For SC 3 the doses for the environment are below 0.03 / 0.09 rem. In the case of SC 4 components, the safety analysis does not reveal any radiological effects on the environment or personnel, for example the engine house. tO1

Following the procedure described at the beginning, a transition has to take place from safety classes to quality standard classes. The principle: the greater the safety relevance of a component, the more stringent the requirements for quality assurance. That means adequate steps must be taken to prevent damage. These preventive measures are quantified in the quality standard classes. In this concept the type and scope of these requirements are governed, among other things, by the following input factors:

the classification into a safety class from the safety analysis

the component safety aim, e.g. from the integrity concept

the function of the components for example

pressure bearing or unpressurized systems.

metallic or ceramic internals in the reactor pressure vessel

measuring and control systems

ventilating installations

electrical power supply

significance of the parts of a component for the safety of the component as a whole.

This means that with a pressure bearing system completely different requirements have to be set from those, for example, for ventilation installations.

Furthermore each system consists of individual components; the pressure-bearing systems, for example, piping, vessels, fittings/valves and headers. In principle the same requirements for the related quality standard class must also be applied to these components (annex 2) . 102

Finally most components consist of a functional group of several individual parts with various tasks and safety significance, so that a distribution into groups of single parts would be practicable; e.g. one divides the pressure-bearing systems into four groups of single parts: pressure-bearing walls/ supports/functional parts/small parts and standard parts (Annex 2).

This can be best described in relation to a fitting/valve where there are four groups of parts when it comes to considering aspects of safety and functioning:

the body as the pressure-bearing wall the neck of the fitting as the support or load-bearing part the gate valve as the functioning part the seals as the small and standard parts

The classification aspects described here are taken into account in the national and international codes in the same way when it comes to drawing up requirements. But in these codes a quality standard class is allocated directly, that is in a fixed and rigid fashion, to the safety classes, e.g.:

safety class (SC) 1 > quality standard class (QSC) A safety class (SC) 2 > quality standard class (QSC) B safety class (SC) 3 > quality standard class (QSC) C safety class (SC) 4 > quality standard class (QSC) D

In order to differentiate between the safety classes which are designated by numbers (SC 1 to SC 4) the quality standard classes are designated by letters (A,B,C or D).

The number of quality standard classes must be laid down for every functional or component group where, in addition to the different requirements for the safety goals, manufacturing, inspection and economic aspects also have to be considered. 103

The following classification is recommended:

four QSCs for ventilation equipment three QSCs for pressure-bearing systems, measuring and control systems, metallic internals in the reactor pressure vessel two QSCs for power supply systems one QSC for reactor pressure vessel, liners

As a rule there is a direct correlation between the safety class (SC) and the quality standard class (QSC). However there are a number of reasons why this should be handled flexibly.

Thus it is conceivable, with different safety concepts, designs or degrees of utilization, to regrade a certain component up or down. e.g. for SC 2.

SC 1 QSC A

SC 2 ^~- ^. QSC B

SC 3 . QSC C

SC 4 QSC D

Reasons for a lower classification are:

low load level easy replaceability additional in-service inspections or operational monitoring additional redundancy or diversity

On the other hand higher classification can be considered where:

there are stringent requirements with regard to availability if inservice inspections are largely dispensed with possible effects on the safety systems 104

The quality assurance measures recently laid down in the individual component specifications are based mainly on the component safety goals as a general requirement. These can be described as follows, taking the pressure-bearing systems as an example:

QSC A: No failures due to fractures or major leakages, full serviceability must be ensured; e.g. through requirements beyond those laid down in AD, TRD and DIN. (Compare "integrity concepts")

QSC B: No large-area failure, low outage probability, restricted functional losses permissible; through requirements sometimes beyond those laid down in AD, TRD and DIN, downgraded vis-a-vis QSC 1.

QSC C: Functional losses of no safety relevance; requirements laid down in conventional rules and standards.

This has the following significance for the allocation of SC and QSC:

In future HTR standards no rigid allocations are to be specified The allocation should rather be suggested in the component specification and will be laid down in the licensing procedure, taking all safety aspects relating to the plant into account.

The requirements for the quality standard classes from the component specification are to be implemented in detail during design approval, e.g. in single specifications. These single specifications, e.g. for calculation, design, manufacture, fabrication and documentation, are to be drawn up for each component and quality standard class. 105

Summary

This suggestion to divide up the components into safety classes and quality standard classes is to be regarded as a a basis for a discussion and not as a fixed set of rules.

On the basis of the internationally established procedure for classification, the following items were also taken into account:

the special features of HTR such as safety concepts, design principles, degree of utilization

the applicability for all conceivable HTR plants (e.g. HTR 500, HTR module, PNP etc.).

After the description of the systems and components for which safety aims have to be set up, one proceeds in two stages (annex 3):

1. A safety analysis is carried out with enquiry as to component failure, consequential failure, and the failure of a required safety function. The safety class is obtained with the limit values in the radiation protection decree.

2. . Now the classification into the quality standard classes takes place with the possibility of regrading either up or down.

Furthermore requirements are to be set up which guarantee that the safety aims are met. A keyword here is the integrity concept. 106

Annex 1

Components and systems for Evaluation of the radiological which component effects, talcing the design specifications to be drawn measures and system-specific up are (lJ charactistics into account f2)

Examples

Safety analysis for the Reactor pressure vessel classification of components

RPV internals

Shutdown system

Radiological effect of a Gas conduits component failure including the possible consequential failures Hain fan Pressure relief and the failure of a required system Steam generator safety function for

After-heat removal system - the operating personnel and - the environment Refuelling equipment

Operating cooling water system in primary area

SKJNRCANCE OF THE FALURE OF A Compressed air system SAFETY THE BASIS OF THE CLASSES SAFETY ANALYSS for Waste water systems sc ENVHOWÄNTAL PROTECT. WOW SAFETY Doses coutd exceed Dose« cotid exceed Ancillary gas circuits SC 1 5/15 rem 3/3Orem* Mm* acccrdkig (hr* accorotog Steam circuit tog A3 SfrSchV) to§3O2 SbSchV)

Doses coiid reach Doses could reach Cooling water systems and SC 2a water supply 5/15 rem 5/30 rem*

Ooees cotid exceed Dose« cotid exceed Cooling gas temperature SC 2b 0.03/0.09 rem 2,5/15 rem* measurement (*nt •ccoroma, cordetg to $ «5 SSrSchV) to g 48 SfSchV) Emergency power system Doses cotid reach Doses cotid reach SC3 0.03/0.09 rem 2,5/15 rem* Pressure relief system SC4 none Reactor containment StrSchV = Dma— on Protection from Damage by lonähg R*Mtan, FRO 2001.197« / K01.1987 t AI Jain rtm», by way ol «lampl«, to th« «hot» body

* Hau - y*ar IV«« to StrSchV. laid down tar on« *v«rt 1 Sv

vessels COMPONENTS headers

pipings

groups of pressure - support functioning small and PARTS bearing parts standard wall parts

e.g. eg. e.g. e.g.

PARTS body neck gate valve seals bonnet spindle bolts gland Devision of systems, components and groups of parts; for example for fittings /valves in pressure- bearing systems 108

Annex 3 DESCRIPTION OF THE SYSTEMS AND COMPONENTS FOR WHICH SAFETY GOALS ARE TO BE DRAWN UP i SAFETY ANALYSES DESCRIPTION OF SAFETY CLASSES SC 1 " 4 (N = 5) FAILURE CONSEQUENTIAL FAILURE REQUIRED SAFETY FUNCTION i REQUIREMENTS WITH REGARD TO QUALITY STANDARD CLASSES IN ORDER TO ACHIEVE THE SAFETY GOALS

FOR SYSTEM, COMPONENT AND PART i CLASSIFICATION INTO QUALITY STANDARD CLASSES QSC A, B...(N = 1 to max. 4)

GRADING UP AND DOWN IS POSSIBLE

Procedure for the suggested classification (summary) 109

Discussion of the presentation:

Classification of systems and components into safety classes and quality standard classes

Helmers, TÜV Hannover, FR6: Please show us again the picture in which you define the safety classes by a corresponding ranking of dose rates. You divide the safety class 2 in two parts with different consequences for the components in question. Part 2a corresponds with dose rates less than 5 rem and part 2b with those higher than 0.03 rem. What is your schedule to decide whether to choose class 2a and 2b?

Dette, RWTÜV Essen. FRG: The work group has been wrestling with this point for a long time. As previously mentioned, the range of permissible limits - the difference being greater than 100 - appeared so considerable to us, that we consider it sensible to sub-divide them. A clear separation, of, for example, half a band-width was not made either as the whole concept was deliberately designed to be flexible. With the transition from the 5 safety classes to the 3 quality standard classes of the pressure bearing system the following is obtained with the normal allocation SC 1/2 a - QSC A and SC 2b/3 - QSC B and SC 4 - QSC C, and so the components with comparable risk potential can be grouped together. 111

Section III:

Metallic high temperature components 113

Metallurgical and physical fundamentals for the design of high temperature components

F. Schubert Kernforschungsanlage Julien GmbH Institut für Reaktorwerkstoffe

Abstract:

Metallic components in advanced reactors high temperature (HTR) are exposed to different temperatures and stress levels. At ambient temperature the allowable stresses are limited by time-independent properties and stress analysis can be carried out using elasto-plastic stress-strain relationship. At elevated temperatures, the allowable stress must be limited against time-dependent properties and inelastic analysis has to verify residual strains and stresses by using relationships between stress and strain rate.

Thermally induced, strain-controlled fatigue exposure produces, depending on temperature and strain rate, kinematic hardening or softening, whereby some proportion of the strain is creep or relaxation controlled. For specific parts of the steam generator, the live steam circuit, heat exchangers and hot ducts, creep resistant steels and alloys are required to prevent time-dependent failure modes, such as creep deformation and damage, creep-fatigue, exhaustion and damage, microstructural instabilities, excessive high temperature corrosion and, for some components, loss of deformability due to neutron irradiation.

The metallurgical understanding behind all the failure modes are discussed. The principle demands for the design of all components operating at elevatedand high temperatures are the limitation of total remaining creep strains and the avoidance of localized plastic and creep strains. 114

1. Introduction and principal regions of behaviour

The metallic components in advanced gas cooled HTR are exposed to different stress levels in the temperature region from ambient up to 1000°C. The loadings which lead to the different stress levels may be classified as:

load-controlled primary stresses thermally induced, strain-controlled secondary stresses local or strain-controlled peak stresses.

The reaction of a metallic component to these loadings depends on the temperature of the exposure and on which transient or which strain rate the load is put on the component. Within the stress-temperature-parameter field there are three principle areas as shown in Fig. 1.

Region A: At ambient temperature, an axial loaded test bar, the simplest model for a component, reacts by elastic deformation if the stress level is kept below the yield-strength of the material. The stress-balanced elastic strain " ee}" remains practically constant during the total load period. Thermally induced strains are balanced with stresses, fairly no change in the stress level occur. The allowable stresses, therefore, are limited by time-independent mechanical properties, such as the yield strength (or 0.2 % proof stress for austenitic steels and high temperature alloys). The equilibrium between stress and strain is calculated based on linear elastic materials behaviour and the effect of cyclic stresses can be derived by Wöhler-type design curves. HTR components with operational temperatures and stresses within the region A can be handled and designed by rules similar to those for LWR components, e.g. KTA 3201. 115

Region C: , At high temperatures (above about half the melting point), metallic materials deform continuously with increasing exposure time, a process known as creep, even at stress levels below the yield stresses. The balance of stress and strain is not static. For a given stress the rate of deformation, the creep rate decreases in the first stage, becomes more or less constant in the second stage and then increases rapidly in the third stage until rupture occurs. It is only in the steady state of the creep behaviour that a kind of quasi-equilibrium between stress and strain rate exists. In a hot tensile test, the stress-strain curve can satisfactory approximated by an elasto-viscous constitutive equation and the. classical plasticity becomes unimportant. The allowable stress levels have to be limited against temperature and time-dependent creep or creep-rupture properties, such as 1 % creep-strain limit and the stress to rupture.

Region B: In the temperature region between ambient and high temperature, the primary stresses must also limited by time-dependent mechanical properties, 1 % creep strain. In a tensile test, however, the stress-strain curve can be separated into an elastic part, a plastic part and a creep part of the total strain.

In both cases B and C, thermally induced, strain-controlled fatigue shows either kinematic hardening or softening effects. With increasing temperature and decreasing strain rate, the creep-controlled mechanisms become more importance and the thermally induced stresses relax.

In the development of structural design codes for metallic structural components of advanced HTR the regions B and C are covered.

2. Creep and creep rupture behaviour

Most creep tests are carried out at constant load, which introduces complications for the evaluation of measured data as the true stress and strain rate vary. Nevertheless, in addition to providing useful and 116 practical design data, this kind of test can also give useful information regarding strain rate sensitivity and some plastic instability events. The creep tests are required for

the derivation of appropriate design' data; the verification of constitutive equations for describing high temperature deformation behaviour.

2.1 Cree£ desiaji_data

Constant load creep tests at different load levels and test temperatures provide the basic data for deriving the time-dependent design stresses (the St values).

In Fig. 2 creep strains are plotted against exposure time (in logarithmic scaling) to obtain creep strain curves. From such curves the time to 1 % strain and the rupture time are obtained. The initial stress is then plotted against time to 1 % strain and time to rupture to provide creep rupture plots. Using appropriate safety margins, the time- dependent stress intensity design value St can then be obtained. In the first step of design, the St values are used for limitation of the primary stresses and for definition of the thickness and from of pressure or deadweight bearing components.

2.2 Stress^strain rate_relat_k)nshi£

For inelastic analysis and calculation of relaxation behaviour and accumulated remaining creep deformation, the relationship between stress and strain rate must be mathematically expressed. The first approximation is concerned with the concept of steady state creep deformation. Contrary to the common and very popular interpretions, constant load creep curves seldom show a period of even approximately constant creep deformation rate. In order to understand the metallurgical mechanisms behind the creep controlling parameters, a plot of true creep strain rate (logarathmic scale) versus true creep strain (linear scale) may be used. 117

A pure solid solution hardened alloy, in which deformation is controlled by nucleation and motion of dislocation, by the equilibrium of initiation and annihilation of dislocations and by formation of creep-voids, such a plot shows a very strong decrease of creep rate with increasing creep deformation down to a minimum creep rate, a constant creep rate, or in the case of load controlled tests, a slightly increasing creep rate and a rapid accelaration of creep rate (Fig. 3a), In the case of load-controlled tests, the slope of the steady state regime helps to define the parameter of the well known Nortons creep equation

= k

In reality, we find especially in high stress level and short term creep tests for solute solution annealed austenitic materials such as Alloy 800 and Alloy 617 a large deviation from this ideal behaviour. Also with the variation of stress level, different characteristic curves are found.

Dramatic examples are the curves obtained for axially loaded tubes of Alloy 800 (Fig. 3b) and specimens for machined from bar stock of the same master heat (Fig. 3c) with the same heat treatment. All these tests were carried out at relatively high stresses, which means short term creep- tests. The log e - e curves show a fairly long duration with decreasing strain rate. After reaching a minimum creep rate, a transient creep stage is found before a steady state creep region is reached (Fig. 3b) . A second type of curve exhibits a continuous change in. the creep rate with no clear steady state creep (Fig. 3c).

From these observations, we can conclude:

short term creep curves of solution annealed material are very sensitive to the microstructure at the start of'the test an extrapolation to long term creep is nearly impossible. 118

Therefore, for the definition of long term design data there is a need of long term experimental tests with test times at least one third of the requested operational time . The reason for the behaviour of the technical alloys is the fact that several creep mechanisms have to be considered:

the mechanism of dislocation motion alone; the interaction of dislocation motion with precipitates, which may change with increasing exposure duration, as precipitates coarsen; diffusion-controlled deformation due to subgrain boundaries or grain boundaries effects.

The knowledge of the kinetics of microstructural changes in technical alloys helps us to understand the large scatter in the creep behaviour and to develop adequate inter- and extrapolation methods.

3. Microstructural instabilities

In fig- 4a the time-temperature-precipitation diagram of Alloy 800 is given.

In the solution annealed microstructure only some primary carbides and a very limited amount of secondary carbides are present. During heating up, secondary carbides are precipitated. The amount, the size, the morphology and the distribution are very dependent on the duration and temperature. The thermal loading may cause a change of the above described microstructural features. These changes in microstructure are accompanied by a measurable change in the density of the alloy, represented in the Figure 4b by the change of elongation of a specimen as a function of ageing time.

The next pictures should demonstrate the change in the microstructure during service. 119

The comparison of the microstructure of heat resistant ferritic steels 13 CrMoV 44 in the as heat treated condition and after an operation time of about 220 000 h at 530°C demonstrates the change in microstructure during service (Fig. 5). After the long term exposure no bainitic structure remains and the microstructure consists of a ferritic matrix with stable carbide precipitates. The change in microstructure of austenitic steels and nickel-base alloys is not as marked as in ferritic steels.

In solution annealed Alloy 800 H, the grain boundaries and the grains are almost free from precipitations, but after long term exposure the grain boundaries are decorated with carbides (Fig. 6). It is obvious, that these different microstructural features, especially the differences in the morphology and distribution of precipitates, must have an effect on the creep behaviour.

The motion of dislocations, the principal mechanism of deformation, reflects these microstructural features. Fig. 7a illustrates the distribution of dislocations around a primary carbide in the grain of solution annealed Alloy 800 H. Fig. 7b gives an example of the beginning of subgrain formation, which is typically for low or medium stressed, long term creep exposed specimens of Alloy 617. These examples may provide an frankly understanding of the complicated changes in microstructure of a metallic component. Even different loading sequences introduce a slightly different deformation behaviour in short term and/or transient loading, due, to microstructural effects.

4. Fatigue properties

In strain-controlled, low cycle fatigue tests, the stress-strain behaviour is similar to that in hot tensile tests. In region B, where we had to account the plastic deformation, the saturation stresses reflect kinematic hardening. 120

In order to obtain the LCF design curve, LCF tests with different strain ranges are performed and the number of cycles to technical crack initiation or to failure are evaluated (Fig. 8). The resulting curves are weighted with safety margins.

Having in mind what has been discussed for the creep, this procedure often reflects only the short term behaviour of the solution heat treated material and the stress-strain behaviour of aged material will differ with respect to the type of hardening. The saturation stress per cycle will become less than that of the solution-treated material.

With increasing temperature and decreasing strain rate, kinematic hardening changes to kinematic softening, and the failure mechanism becomes partly creep-controlled. In order to overcome the problems of different elasto-plastic-creep response for aged material, low cycle fatigue tests with holding time are performed to simulate a creep-fatigue interacting loading.

The problems of the interpretation of these tests for life-time fraction rules will be a special topic of this workshop. In principle, as shown in Fig. 9, for the region B and region C, it has been observed that transgranular fatigue crack or damage do not occur simultaneously with intergranular creep damage. In the region of operation temperature around 550°C, where creep does not necessarily predominate the exhaustion due to fatigue, both creep and fatigue exhaustion can be handed additionally without a synergetic effect. At temperature around 950°C, in addition to transgranular fatigue crack initiation and propagation, crack initiation, crack- and propagation can become creep-controlled i.e. of intergranular type. A mixed mode is observed.

The available experimental results of LCF tests with hold time for the favoured materials for HTR components belongs more or less to the strain range which is controlled by fatigue failures; only a very limited number of experiments are in the creep-fatigue interaction area. Therefore, life-time fraction rules derived from this type of experiment must be treated very seriously. 121

5. Environmental effects

Some of these experiments indicated an influence of the test environment. There are also indications that the crack growth in air, in simulated HTR helium and in vacuum differs.

5.1 Gas-MetalJ_interaction_i£ HTR helium

The coolant helium of a HTR contains impurities of methane, carbon monoxide, hydrogen and water. To describe the behaviour of a metal surface in respect to oxidation, carburization or decarburization, the temperature dependence of carbon activity and oxygen partial pressure of the gas must be deduced. The gas-metal interaction is determined by the kinetics of different carbon and oxygen transferring reactions between the metal surface and the atmosphere. In a schematic stability diagramm (Fig. 10)1 carburization, oxidation or decarburization are obtained depending on carbcn activity and oxygen potentialy. (Fig. 11 gives typical microstructure of the surface of specimen of Alloy 617 after these different gas/metal reactions) controlled by the most active alloying chromium. Furthermore the partial pressure of carbon monoxide is critical for the occurrence of oxidation with small amount of carburization or decarburization.

Form an evaluation of all corrosion, tests in simulated HTR helium, we may conclude that for Alloy 800 H and Alloy 617 up to 900°C that the gas-metal interactions lead to the formation of oxide layers and slight capbon.; •.uptake.! and have no effect on design data and deformation controlling mechanisms for strains of less than one to two percent. At temperatures above 900°C, the environmental influence on mechanical properties can be neglected provided that the helium contains sufficient carbon monoxide to keep the oxygen partial pressure and the carbon activity high enough to avoid decarburization. 122

5.2 Neutron irradiation effects

The impact on neutron irradiation must be taken in account for both fast (E > 0.1 MeV) and thermal (E < 0.025 eV) neutrons. The control rods of a pebble-bed type high temperature gas cooled reactor are one of the metallic components are exposed to a neutron fluence which can not be neglected. The heat exchanging components on the other hand are far enough from the core, for the expected fluence to be smaller than that, which could be harmful for the steels and alloys. A great effort concerning the tensile and creep properties of neutron irradiated specimens of a certain number of steel and nickel-base alloys, together with an optimization of the microstructure of the favoured steel X8CrNiMoNb 16 16, is summarized in the following statements:

The impact of neutron irradiation is dependent on the accumulated fluence, on the proportion of thermal neutrons, on irradiation temperature and post-irradiation test temperature (Fig. 12).

The thermal neutron induced embrittlement starts at thermal fluences of about 1 • 10*8 cm~2 and is wery strongly dependent on the test temperature; below 600°C, there is no marked effect. At higher temperatures, however, these irradiation condition, although the content of (n -a;)- reaction products is very limited, the deformability is markedly reduced. The specimen failed by intercrystalline brittle cracks. These observations forced the designer of HTR plants to be very careful with metallic structures which are exposed to thermal neutron irradiation. Ferritic steels are less sensitive to thermal neutron irradiation than austenitic alloys.

6. Final remarks

For the structural design of high temperature components, the designer should be aware that a high temperature exposed metallic structure is like a living body:

It suffers under temperature and stress, it ages, it reacts 123

with the environment and can become very sensitive to thermal neutron irradiation with increasing temperature (namely the austenitic steels and alloys). The task of the designer must be to limit the loading conditions in order to avoid the following failure modes (table 1):

Table 1:

Metallurgical and physical fundamentals for the design of high temperature components

loading reaction of the material

temperature microstructural instabilities mechanical loading creep deformation and creep damage transients fatigue exhaustion and crack initiation environment corrosion products (crack initiation) neutron irradiation loss of deformability

failure modes

ductile and brittle fracture due to short-term loadings creep rupture due to long-term loadings creep fatigue failure due to cyclic loadings excessive strain due to incremental deformation or creep ratcheting loss of stability due to short-term loadings and long term loading loss of stability due to long-term loadings environmentally caused material failure (excessive corrosion) fast fracture due to instable crack growth

The work on structural design rules and method for the high temperature regimes B and C, the central discussions and work is concerned with the avoidance and limitation of these failure modes. 124

according to RABOTNOV

A B elast + plast ptast+creep

RT 500 1000 temperature /°C

£ = const., T = const. 6

6 = const., T =

Fig, 1: Principle Temperature Stress-Strain-Behaviour of Metallic HTR-Components 125

100 - creep rupture elongation . c 115MPa co 10

0.1 10 10 102 103 10 time /h

CO 200 a. 100 creep rupture strength

CO 1 %- creep starin fanit CO 50 a> T = 850 °C CO 20 NiCr 23 Co 12 Mo 10 10 0 10 10 103 10 time / h

Fig. 2: Scheme for Evaluation of creep limits 126

wahre Dehnung £w Fig. 3a: True strain versus true strain (loge-e) schematic creep curves

10 wahre Detinung/ X Fig. 3b: log e/e -diagram for IHX-tubes of Alloy 800 axial loading (950°C,CT= 30 N/mm2)

te AVL-M AtL-30 »VL-7

X N W

-2 te I I 18 12 14 10 wahre Dehnung/ % Fig. 3c: log e /e-diagramm for standard creep specimens (bar materials) of Alloy 800 under axial loading (950°C, o = 30 N/mm2) 127

time temperature precipitation diagram

change in length due to ageing

0.06

10 104 10

time/h 4: Example of microstructure instabilities in XlONiCrAITi 32 20 (Alloy 800) 128

50 pm

a} a» heat treated b) as exposed: 535°C 220000hre.

Typical Mtoroatnicture of the Steel 13CrMo44

Fig. 5: Typical Microstructure of the Steel 13CrMo44

20um

a) as solution annealed b) as exposed: 900°C 39 510 hrs.

Typical Microstructure of XIONiCrAITI 32 20 (Alloy 800H)

Fig. 6: Typical Microstructure of XlONiCrAiTi 32 20 (Alloy 800H) 129

Fig. 7a: Dislocations arrangement around a primary carbide in an as solution heat-treated Alloy 800

Alloy 617. O.OB : : C. creep ' tesled in air. 9SO C. IB MPa. * 7.6 :: strain

_ ^,V pinning of subgrain boundaries * -" by carbides

Fig. 7b: Dislocations and Carbide arrangement in a solution heat treated and creep-exposed specimen of Alloy 617 ( 950°C, 18 MPa, 7,6 % strain) 130

max. Mm*

numbar o( cyctos

Fig. 8: Evaluation of LCF-Design Curves

Fatigue initiation Fatigue failure en G O .C-F interaction c "o Creep initiation "in Creep failure O o

Log (Number of cycles)

Fig. 9 Creep-Fatigue Interaction Diagram (schematic) 131

Fig. 10 Corrosion areas in the stability diagram for chromium

Typical Corrosion Effects in INCONEL 617

surface layer

bulk

Exposure at 950" C in HTGR helium with different Impurity contents

Fig. 11: Typical microstructure in the surface of Alloy 617 after different gas-metall reactions. 132

creep strain

100 —i ^..

50 I"--- - 40 i .L4981 KA2 (Irradiated) 30 20 1.4981 KA3(lrr idlated) 10 r^"^^^— **fc^_" 1 5 4 3

10' 10' 10' 10" 10" 10' thermal fluence/m"*

Fig. 12: Microstructurally optimized variant KA 2 of the austenitic steel 14981 exhibits superior creep rupture ductility at 1123 K 133

Workshop on Structural Design Criteria for HTRs 31 January - 1 February at KFA Jülich

Load Levels, Stresses, Failure Modes and Design Criteria

Bieniussa, K. and H. Reck

Gesellschaft für Reaktorsicherheit (GRS) mbH Köln 134

1. Introduction

The present paper discusses the design against failure of metallic components to be used in high-temperature service. For this purpose, it is demonstrated within the scope of the design of components that no failure of a component to be assessed is to be anticipated under the loads to be applied. This demonstration requires knowledge with respect to the following aspects (Fig. 1):

possible failure modes

loads to be applied

selected design concept

given design criteria.

2. Failure Modes

2.1 Material Behavior

Load applications cause stresses and strains in the material and may lead to material exhaustion. Special cases of stresses which are also used in the testing of materials are (Fig. 2)

variable stresses occurring during the application and removal of loads (-• tensile test)

constant stresses occurring under constant loads (-* creep test)

cyclic stresses occurring under alternating loads (-> fatigue test). 135

Starting out from the special case of the constant load, the material will behave differently, depending on whether its application temperature is below or above

T = 0.5 . Tg, where T„ - melting point in K. In the temperature range of

T < 0.5 . Ts - which is hereinafter referred to as low-temperature range - the materi al behavior is independent o f time, as a constant load also results in a constant stress. On the other hand, the material is characterized by a time- dependent behavior in.the high-temperature range even if the load is constant, as is shown by the following two special cases (Fig. 3):

creep test with stress = constant and strain ^ constant (rising)

relaxation test with strain = constant and stress ^ constant (falling).

In addition, the creep tests show that both the strain rate and the attainable elongation at rupture will increase with increasing stress, while the time to rupture decreases with increasing stress. This means that a rupture will not occur under short-term but only after long-term load applications.

In the cases of cyclic .stresses, the lifetime of the material is • also influenced by the time history of the stress: apart from the range of the stress, the rate of the stress and the hold time under stress are the decisive factors influencing failure.

Furthermore, it should be remembered that, as a rule, . welds in the high-temperature range do not reach the strength properties of the base material. 136

2.2 Failure Mechanisms

The failure modes of components, namely

rupture or leak,

structure instability,

loss of function

may be caused by the following failure mechanisms as a result of short-term load applications (stla) or long-term load applications (ltla) (Fig. 4):

excessive plastification of the material involving . ductile rupture (stla) . tough creep rupture (ltla)

missing material ductility involving • brittle rupture (stla) . non tough creep rupture (ltla)

material exhaustion involving . fatigue failure (stla) . creep fatigue failure (ltla)

component exhaustion involving . sudden loss of stability (stla) . creep-related loss of stability (ltla)

gross distortion due to . ratcheting (stla) . creep ratcheting (ltla)

loss of function due to . excessive deformation (stla) . excessive deformation (ltla) 137

material deterioration in service due to . environmental effects (ltla) . friction welding (ltla).

The damaging effect of the conditions of application cannot be ruled out by design" criteria alone, but requires additional measures with respect to the selection of materials and the structural design of components.

2.3 Consideration of the Influence of Time

From the description of the behavior of the materials and from the list of the various failure mechanisms, it can be gathered that time is an essential factor with respect to design in the high-temperature range. Thus, for the purpose of the design of components,

the load applications and their time histories have to be quoted and

the material data have to be made available considering their dependence on time.

Moreover, as several temperature levels may cause different levels and durations of loads, certain agreements must be made for design purposes with respect tö

the beginning of time-dependent material behavior,

- the evaluation of different stresses, and

- the evaluation of different stress cycles.

The -high-temperature range is characterized by time-dependent material behavior. The transition from low-temperature to high-temperature range - • which is a gradual one and also depends on the material - can be defined by means of a temperature/time curve. Fig. 5 shows the transition as-well as 138

three ranges which have to be covered by different rules:

Range I : low-temperature range without creep.

Range II : transition range without any significant creep, Range III: high-temperature range with creep.

The determination of such boundary curves will be explained in one of the following papers.

In the high-temperature range, different stresses which may vary in both level and duration have to be assessed as a whole because of their individual creep contributions. For this assessment, a linear relation, the Robinson Rule, is used which adds up the individual degrees of time-use-fraction (Fig. 6):

D c = I t./a/t mi• where t. = available time for the stress t • = allowable time for the stress

When evaluating different load alternations, the influence of possible hold times is considered by adding up in a linear approach the existing degrees of time-use-fraction to the degrees of fatigue-use-fraction.

3. Design Concept

3.1 Preconditions

For the design of components aiming at the prevention of failures under load, experiments, rules or analyses may be used. The design concept to be presented here is based on analyses and the material data to be used for this purpose as well as the design criteria to be considered were verified by experiments. 139

Generally, the design concept (Fig. 7) requires a suitable selection of materials and thus excludes any use of brittle or embrittling materials. For the materials selected, the necessary time-independent and time-dependent material data for both the base material and the welds shall be made available, in. addition, it is postulated that, when determining the material properties the ambient influence for the planned application is ascertained correctly and that the decisive loads such as temperature and pressure, as well as the conditioning of the process fluid, will be monitored during the later operation of the component. Quality assurance during manufacture and assembly of the components makes it possible to do without fracture-mechanical demonstrations within the scope of this concept.

As is also the case in other codes,' the design of components in the high-temperature range is done in two steps:

dimensioning of the load-bearing wall,

analysis of the mechanical behavior.

3.2 Dimensioning

Within the scope of dimensioning (Fig. 8), the load-bearing wall thicknesses of the components shall be laid down in such a way that the load capability of both the pressure loads and the external forces will be possible over the entire time of components application. '

However, this aim cannot always be reached, since

the entire time load spectrum has to be taken into account (also see 4.3, Dimensioning Condition),

' - • . • . . . thermal loads may require enlarged wall thicknesses for compliance with the design criteria, 140

wall thickness additions have to be provided in order to cope with environmental effects such as corrosion and erosion.

Therefore, when designing components in the high-temperature range, the dimensioning is often only a first establishment of the wall thicknesses which may have to be corrected within the scope of the following stress analysis.

3.3 Stress Analysis

Within the scope of the analysis of mechanical behavior (Fig. 9), it shall be demonstrated, in order to avoid failure, that the components will be able to withstand all the anticipated stresses during their specified lifetime. For this purpose, it is demonstrated that the possible failure modes will not occur, since suitable design criteria are met in the high- temperature range.

Within the stress analysis, amount, frequency and time history of the loads as well as the resulting stresses shall be determined and assessed. For the assessment of the time histories of loads and stresses, it is often not possible to reach the aim by means of stress analysis in which an elastic behavior of the material is postulated for simplifica- tion purposes. In such cases, the anticipated time history of deformation as well as the stress/strain time history has to be determined, and subsequently assessed, by means of simplified inelastic or detailed inelastic methods of calculation. The following papers will discuss the necessary comprehensive preparatory work which is necessary in particular for the latter methods of calculation in order to

idealize the geometry of the components, :

select representative load histograms,

formulate the material constitutive equations, and

perform the inelastic calculations. 141

4. Determination of the Loads

4.1 Events and Loads

All the events or sequences of events - hereinafter referred to as "events" - in the entire plant whose individual components have to be designed, are determined with respect to their operating and safety-related importance and allocated to five event categories.

Each of the events is allocated to a certain load level in a component-specific approach (Fig. 10), and the associated stationary and instationary loads are determined for each event. Subsequently, the events, the event categories, the load levels and the associated loads are summarized in event tables (= load case tables).

Loads are all the controlling influences acting on the component as a result of mechanical and thermal loads, corrosion, erosion and irradiation. They have to be taken into account in design, construction and stress analysis in a component specific approach using precise or conservative values.

4.2 Load Levels

The individual load levels differ with respect to the size of the safety margin to failure concerning the individual design criteria. For differentiation purposes, four operating levels (Levels A, B, C and D) as well as the dimensioning level (Level 0) are used. The operating levels comprise three levels deviating from normal operation (Level A).

The classification criteria depend, on the one hand, on the respective requirements which the component has to meet when required and, on the other hand, on the measures to be taken to ensure continued operation after occurrence of the event (Fig. 11): 142

Level A: Normal operation: continued operation

Level B: Restricted operation: removal of causes Level C: Interrupted operation: inspection/repair Level D: Interrupted operation: repair/replace.

Another difference between the individual operating levels relates to integrity, i.e. in how far deformations are permitted. The loads to be taken into account in each case are aimed at this integrity concept for the components.

The dimensioning level is used for the first dimensioning of the component wall thicknesses. In this context, all the loads from the operating levels A through D that result from mechanical loads are taken into account.

If several pressures and temperatures apply for various times during the specified duration of load application, these have to be summed up in a collective referred to as the dimensioning condition and the dimensioning of the component wall thicknesses has to be based on this condition.

4.3 Dimensioning Condition

The loads of the dimensioning condition are the following:

dimensioning pressure, dimensioning temperature, additional mechanical loads, dimensioning time.

As in the design in the low-temperature range, the dimensioning pressure and the dimensioning temperature are the maximum values of the pressure differentials and mean wall temperatures, respectively, which occur at the operating levels A through D. In this context, the additional mechanical loads are transferred into equivalent pressure loads for dimensioning purposes. 143

The dimensioning time which is still a free parameter is laid down in such a way that, considering the other dimensioning factors, an equal material use-fraction exists in relation to the degree of time-use-fraction as exists under the operating load conditions (Fig. 12):

The pressure load p is transferred into stresses a by means of a factor f.

o = f . p

For this purpose, factor f is chosen in such a way that a time-use-fraction of D"c = 1 is reached when all operating loads are taken into account.

Using the temperature and time-dependent stress diagram which is decisive for the design, the time-use-fraction is determined for all operating loads i:

Dc = where t. = operating load application time t . = allowable load application time, mi both depending on temperature T. and stress a..

Using the temperature and time-dependent stress diagram which is decisive for the design, the dimensioning time is determined as the allowable loading time as a result of the stresses from the dimensioning pressure and the additional mechanical loads at the dimensioning temperature. 144

5. Design Criteria

5.1 Safeguarding Concept

The safeguarding concept for the design of components assumes the methodical approach of the low-temperature range which it supplements, if required, by necessary demonstrations in the high-temperature range.

The essential items of the methodology are (Fig. 13):

- Classification of loads regarding failure effects: primary values (stress-controlled), secondary values (deformation-controlled), peak values.

Classification of stresses regarding the distribution across the section

equivalent membrane forces, equivalent bending moments, peak values.

Safeguarding of loadings against allowable values: primary stress limits, strain and strain cycle design limits and secondary stress limits, respectively, safety against failure loads and times.

Stepwise safeguarding of the total stress state: limitation of primary stresses, limitation of strains and strain cycles and secondary stress limits, respectively, limitations of loadings and times.

Supplementary demonstration in the high-temperature range:

safeguarding in the time range. 145

assessment of welds, identification of environmental effects.

5.2 •"* Design Limits

The design limits and aspects of assessment elaborated for the design of components in the high-temperature range can only be presented by way of general examples here, since any individual assessment will require a great number of diffe- rentiations to be made, e.g. with respect to kind of material,

product form, stress distribution,

operating level. Thus, this presentation is restricted to membrane stresses at operating level A in the case of austenitic forgings.

5.2.1 Limitation of the primary stresses (Fig. 14):

Primary stresses can be determined by means of elastic calculation methods. The safeguarding of the primary membrane stress P in a component always requires a comparison with the allowable stress as well as, in the case of more than one operating state,, the additional demonstration of the time-use-fraction.

1. Limitation of stress P ^ S m m s st where S = f (short-term strength) S. = f (long-term strength for time t).

2. Degree of time-use-fraction

D =1 (t./t .) ^ 1-0 c . v r mi' where t. and t . are as already described, l mi J 146

5.2.2 Limitations of strains and strain cycles (Fig. 15):

As a rule, the calculation of strains and strain cycles requires inelastic methods of calculation which supply statements on the time history of strain for given stress histograms. The accumulated strains e to be expected towards the end of the lifetime of a component as well as the strain cycles Ae accumulated until then, including the effects of the hold times t that have occurred have to be safeguarded.

1. Limitation of strain

e % 0.01 m

2. Degree of creep fatigue use fraction

DCF = Z ni / dt/tD + Z ni/Ni = D i 0 i where

n- is the existing number of load cycles for strain cycle Ae N. is the allowable number of load cycles for strain cycle Ae t is the existing hold time for stress CT = f (t) t is the allowable hold time for stress a D is the allowable degree of use-fraction

D = f (DF,DC).

In the transition range (range II in Fig. 5), the approach presented here can be simplified: the strain limitation is replaced by a modified secondary stress limitation and the : degree of creep use-fraction results as a global value of Dc 0.1 because of non-significant creep. 147

5.2.3 . Limitation of the loads (Fig. 16):' Loads P which may lead to instability failure, such as the time-independent instability load in the case of critical load problems P, and the time-dependent instability load in the case of collapse load problems P, . require, when being calculated, that the inelastic material behavior as well as all' loads which act simultaneously are taken into consideration. In these cases, a limitation of the loads and/or the load application times has to be made, since no definite stress limitation with allowable stress values is possible.

1. time-independent stability limits

P = P,km /3.00 load-controlled (P = P, /1.67 strain-controlled)

2. ^time-dependent stability limits

t = ^t/5 where t is the allowable and t, is the col laps time under load P. With n as Norton's creep exponent, the time safety 5 corresponds to a load safety of 5 ' when applying Hoff's theory.

5.3 Example of Application The application of the design criteria is to be demonstrated using the time-dependent instability failure as an example.

Let us assume a pipe made of Inconel 617 and subjected to an external pressure of p = 35 bar and a temperature of T = 950 °C. The pipe has an initial ovality of 1.6%. The initial ovality provides a bending in the pipe which leads to an increase in the ovality and a simultaneous increase in the stresses as a result of creep deformations. Until a load application .time of about • 800 h, the ovality increases linearly in first approximation,• and after 930 h there is a great increase in ovality until the pipe fails (Fig. 17). 148

In accordance with the design criteria, the allowable lifetime of this pipe is

t = (930/5) h = 186 h.

Within this time, there is only a slight increase in the ovality of the pipe so that this would not affect the operating aspects of the plant.

The time-independent instability failure in the case of this pipe is anticipated well above p = 1000 bar, i.e. the time- dependent kind of failure determines the design.

6. Summary and Outlook

The design criteria for metallic components which are to be used in the high-temperature range were described by way of example: the time-dependent material behavior in this temperature range requires (Fig. 18)

safeguarding against time-dependent kinds of failure,

provision of time-dependent material data,

consideration of the time-dependent loads.

The design criteria elaborated as well as the available material data permit a design of the components that can withstand the loads for the planned operating lifetimes. However, further developments in individual areas seem to be desirable, such as

assessment of welds,

simplified inelastic calculation methods, safeguarding of deformation-controlled loads.

When preparing the design criteria, the approaches used in other codes, e.g. the ASME Code or the RCC-MR Code, were taken into consideration, but not always followed. Differen- ces exist, for example, regarding the 149

treatment of design conditions,

formation of allowable stresses,

assessment of welds,

safeguarding against loss of stability.

We are quite prepared to enter into an exchange of experience in the field of design criteria.

Workshop on Structural Design Criteria for HTR

Load Levels, Stresses, Failure Modes and Design Criteria

by . K. Bieniussa and H. Reck

Aspects: -failure modes, -load types -design concept, —design criteria

Fig. 1 : Introduction 150

-variable

-constant

Time -cyclic

Time Fig. 2: Special Load Cases

7\me Strain Stress Strain Stress

Time Time creep relaxation Fig. 3: Creep and Relaxation 151

time independent failure modes: time dependent failure modes: ductile short term rupture tough creep rupture brittle short term rupture non tough creep rupture fatigue failure creep fatigue failure buckling creep buckling gross distortion due gross distortion due to ratcheting to creep ratcheting loss of function due loss of function due to excessive deformation to excessive, deformation environmental effects environmental effects

Fig. 4: Failure Modes

Fig. 5 : Creep-Cross-Over Curve

id 10 10 Time (h) 152

D c = I mi

tj = availiable time for the stress,

tmi = allowable time for the stress

Fig. 6: Robinson Rule

Objectives: -material selection, determination of: -wall thickness.- . ' . .-diameter, J-)£ \ •; •'•/..'jj';?| -material properties, -geometry, • •'-? -V '' _; '• yfi

• '••(,:_ • • \,.y Considerations: . ,; -environmental effects, influence of: ;-mechanical loads, * -^ -quality assurance -thermal loads^ * '. :,> —environmental effects, . •;

* loads = f(time) . . => load histogram -

Flg.7: Basic Requirements Fig. 8 : Dimensioning of the Structure 153

Objectives:

-proof of load capacity, —consideration of design criteria, —exclusion of failure modes

Considerations: —possible calculation proceedure, —idealizing of the geometry, —representative load histogram, —formulation "of material constitutive equations, —performance of calculation

Fig. 9 : Analysis of Mechanical Behaviour 154

PLANT COMPONENTS event classes: load levels: 1 . A

2 B yes yes

no/6pe ration yslem *

special consideration * restricted operation/inspection ** shut down/inspection

Fig. 10: Plant Event Classes and Component Load Levels

operation modes load integrity: levels: of components: loads:

normal operation: no plastic all loads A continued operation deformation

restricted operation: no plastic all loads B removal of cause deformation local plastic shut down: all loads C inspection/repair deformation

shut down: gross plastic . primary D repaii/replace deformation. loads

Fig. 11: Load Level Aspects

156

Austenitic Steel

-rolled or forged-

R 27j R 5 R 11 ' m,T/ P>0.2,RTA ' PI0.2,T/

R 1 0 R ' l.O1T,t/ ' ' m,T.3xtA°' f^

m " ~m'

Fig.14: Primary Stress

1 o

Fig.15: Creep Fhtigue Usage Fhctor

-time independent:

P <= Pkm/3.0 load controled

P <= Pkm/1.67 strain controled

-time dependent:

t<= tkl/5.0 Fig.16: Stability 157

. •

j J• 0 —e 0.0 200.0 400.0 600.0 800.0 1000.0 . 1200.0 Time(h) T= 950 C, P= 35 bar, D- 40 mm. t= 3.3 -mm, a= 1,6 % Fig. 17: Ovality as Function of Time

Summary: -safeguarding against time dependent kinds of failure, -provision of time dependent material data, -consideration of the time dependent loads

Outlook: -assessment of welds, -simpified inelastic calculation methods, -safeguarding of deformation controlled loads

Fig. 18 : Summary and Outlook 158

Discussion of the presentation:

Load levels, stresses, failure modes and design criteria

Muto, JAERI, Japan: In order to determine a loading cycle for the stress analysis, we must superpose all imposed loads such as dead weight, seismic load and thermal load. How do you deal with this, particularly with a superposition of a seismic load and a thermal load?

Bienussa, GRS, Köln: We have the following procedure: seismic load is considered a low frequency short-time loading; for this reason we can use elastic calculations for the fatigue analysis. If seismic load can occur often in the life-time of a plant, as it might be in your country, this event should be included in the load histogram of a inelastic stress-strain analysis.

Muto, JAERI, Japan: The problem is that it is very cumbersome to superpose many loads in the practical inelastic stress analysis, though it may be correct. 159

Basic Requirements relating to Quality Assurance of Safety Related HTR Materials and Components

Dr.-Ing. Jürgen Just RWTÜV Essen, FRG

Introduction

The QA system for components in future high temperature reactors basically corresponds to the systems introduced in the FRG for LWR and other types of reactor.

Fig. 1 gives a schematic picture of this system with the most important QA elements. A distinction is to be drawn here between the elements of the system-related quality assurance (=system tests), shown on the left and the elements of the component- related quality assurance on the right. Here again a distinction is to be drawn between' the software tests and the hardware tests (component tests).

Tradionally the component-related hardware tests by the Independent authorized inspector are more highly rated in the German regulations as compared with the ASME code. This principle will also apply to the KTA rule 3221 for future HTR plants.

The following remarks will deal mainly with the quality assurance elements for:

material assessment the manufacture of materials and semifinished products and the manufacture of components

Special emphasis will be placed here on the HTR's special features 160

Material assessment

The aim of the material assessment is to demonstrate that the material in question is suitable for its intended use.

The principle that materials for use in conventional pressure bearing installations subject to mandatory inspection must be assessed by the authorized inspector is firmly anchored in the conventional regulation /I,2/. The procedure to be adopted and the test plans for certain types of products are laid down in VdTÜV specification sheets /3/. The results of the material assessment are drawn up in VdTÜV material sheets. At present there are approximately 500 such material sheets in existence.

Additional requirements have to be fulfilled in the case of materials used in nuclear components with a high safety related significance, e.g. with respect to neutron embrittlement, relaxation embrittlement and corrosion behaviour. General guidelines for the assessment of nuclear materials are laid down in the KTA draft 1406 (for base metals /4/) and in KTA 1408.1 /5/ (for weld filler metals) and special attention is paid to the requirements of light water reactors.

Within the framework of the Federal Ministry of Research and Technology (BMFT) research project "design criteria for future HTR plants" an investigation was conducted into how far these regulations could be applied to HTR materials that had to be modified.

Basically HTR materials are also tried and tested in conventional technology which have had to be modified to meet the specific HTR requirements. 161

Under the established procedure in the conventional and nuclear fields, a distinction is drawn between:

general material assessment unrelated to the manufacturer and material assessments specific to the manufacturer and/or semi-finished product.

The general assessment of the HTR reference materials X20 CrMoV 12, X10 NiCrAITi 32 20 (Incoloy 800 H) and NiCr 22 Co 12 Mo /Inconel 617) has been largely completed and can be recorded in data sheets. The assessments in the conventional field were taken into account in connection with the manufacturer and semifinished product-related assessment here.

Certain examinations such as fracture mechanics tests, irradiation tests, fatigue and corrosion tests can be performed jointly by several manufacturers or within the framework of a research project. They must have been completed by the nuclear commissioning.

Special features in connection with the assessment of HTR materials include especially the high operating temperatures over long periods. From this it follows that the creep behaviour within the scope of the assessment is checked in order to

.-= obtain a more solid and secure determination of the creep data and

*• to ensure that all manufacturers together with their materials comply with the data relevant to the design.

This also applies to the supplementary creep tests on welded joints which are not performed on pure weld metal according to DIN 32525, but on test pieces relevant to the assessment. 162

One special which should be mentioned is finally the age hardening tests in the area of the embrittlernent temperatures on welded test pieces with and without prior cold deformation.

Table l gives an overview of the typical tests performed for the assessment of reactor materials.

Quality assurance of materials and semi finished products

An initial condition for the manufacture of materials and semi finished products for HTR plants is, as mentioned, the material assessment related to the manufacturer. This does not necessarily have to be completed before commencement of manufacture, but can also take place entirely within the framework of the first deliveries.

A second condition is provision of evidence for the manufacturer's qualification. Since the HTR materials are produced without exception by experienced material manufacturers in the meaning of the conventional codes and standards (AD,TRD), known to both plant erectors and authorised inspectors, this is, generally speaking, not a difficult hurdle.

Then again, contrary to the ASME code, German codes and standards do not require an auditable quality assurance manual. Nevertheless most manufacturers fulfill ASME requirements in the interest of international consignments.

Whereas the the ASME-approved material manufacturers are almost entirely directly responsible for their products, the important safety-related HTR components are basically subject to an assessment by an authorised inspector. Therefore the responsibility of the inspector during the final inspection and pressure test of the component also covers material acceptance, in particular in the case of highly stressed, sensitive or modified materials which have been developed further and not because the manufacturers are not trusted. The interaction of the manufacturer,inspector and,where applicable, the main contractor. 163 during the final inspection however also facilitates a simplified consultation procedure in the event of minor non conformities in vis-a-vis the specification: the ASME regulations are far less flexible here.

For the acceptance test of the most important semifinished products i.e.

- sheets forgings, plates, bar steel seamless tubes for piping heat exchanger tubes of the following two material groups

austenitic steels, nickel based alloys and ferritic steels

there are now agreed drafts for the test sheets. These contain test plans (number of samples, sampling locations and directions, type of tests and type of test certificates). The requirements to be fulfilled can be found in the material sheets also available in draft form or from other material sheets. Table 2 gives an extract from a test sheet as an example.

The reason for this division into test sheets and material sheets is, among other things, that no uniform toughness concept is available for HTR materials such as is the case for LWR materials, because the verification of toughness at room temperature only has the significance of a uniformity test and evidence of delivery according to specification.

The verification of material properties in the time-related temperature range is a feature peculiar to HTR.

It is agreed that long-term properties established within the 164

framework of the material assessment must be confirmed to a certain extent on the semi-finished products for especially safety related components. It was agreed that in addition to the hot tensile test, creep tests should also be performed if necessary. This solution enables the necessary stipulations specific to the component to be met, either in the licensing procedure or within the framework of the design approval.

Quality assurance during component manufacture

The processes employed during component manufacture -bending, welding, stress relief, machining etc.,- are comparable for all reactors. Therefore many tried and tested quality assurance measures can be adopted in unchanged form.

The proposed procedure for welding procedure qualification, batch and production tests for HTR components of the highest requirements category has its own particular character. While the accent in the light water line is clearly on procedure qualification, i.e. preliminary qualification unrelated to specific components, in the HTR line the philosophy developed for the prototype THTR has been adopted. Table 3 shows a comparison of the procedures adopted for LWR and HTR component manufacture.

According to this the basis for procedure qualification is the general qualification verification of the manufacturer as provided in conventional vessel manufacture, i.e. the welding procedure qualification according to AD-HP 2/1.

On this basis, however, the manufacturer must provide two further component related verifications prior to actual fabrication

the batch test (welding material test) on all weld filler metal batches intended for use. It is done under the processing parameters of the component manufacturer.

the preliminary production test as an addition to the 165

welding procedure qualification test performed under the component related welding conditions and base metal and filler metal batches identical to the component.

In order to monitor ongoing fabrication and to renew the welding procedure qualification and the preliminary production test with their period of validity of 24 months, parallel production tests are still required.

The reason for this more component related QA procedure is, among other things, the fact that the mixed joints, which occur more frequently than in the LWR, demand more careful matching of the weld filler metals in order to minimize, for example, differences in strength in the welded joints, taking account of the heat treatment.

A further advantage is that types of defect not found or not reliably found in the non destructive examination - in particular systematic hot cracks or pores in the weld metal - can be prevented by optimizing the welding parameters or the weld filler metals.

The person to follow me will deal in greater detail with the problems related to the nondestructive volume examination which are of particular importance with the HTR.

Summary

All the quality assurance elements shown in the first figure must not be regarded separately from one another. Only their planned and harmonized interaction will achieve the objective that highly stressed high temperature components will be fabricated with just as great a freedom from manufacturing defects and will be operated with as little danger of catastrophic failure as LWR components which are comparable with regard to safety and which have proven themselves over an accumulated period of more than two hundred operating years. 166

References

/I/ Technical Rules for Steam Boilers TRD 100 General Principles for Materials (April 1975)

/2/ AD Specifications AD-WO General Principles for Materials (June 1986)

/3/ VdTÜV specifications VdTÜV specification 1255 Principles for Assessment of Materials for Plants subject to Mandatory Inspection by TÜV (September 1983)

/4/ Rules of the Nuclear Committee (KTA - Kerntechnischer Ausschuß) KTA 1404 Principles for Assessment of Materials for Use in Nuclear Power Plants (Draft - September 1984)

/5/ Rules of the Nuclear Committee KTA 1408.1 Quality Assurance of Welding Filler Metals and Auxiliary Materials for Pressure- and Activity-Bearing Components in Nuclear Power Plants; Part 1: Suitability Test (June 1985)

/6/ Rules of the Nuclear Committee KTA 1401 General Requirements for Quality Assurance (December 1987)

/!/ Rules of the Nuclear Committee KTA 3201.3 Primary Circuit Components in Light Water Reactors; Part 3: Manufacture (December 1987)

/8/ AD specifications AD-HP 2/1 Procedure Qualification Test for Welded Joints (July 1984) 167

Type of test Type of material

_. Ferritic Austenitic

Base material

Chemical analysis + + Tensile tests RT + + Tensile tests ET + + Long-term embrittlement + + Impact tests + + Drop weight tests + Technological tests (pipes) + + Hardness tests + Microstructure + + Tempering behaviour , + + TTT diagram + Stress relief behaviour + Cold workability + + Aging behaviour + Aging behaviour of cold worked material +2) +2« Flame cutting behaviour + Physical properties + + Fracture toughness + ' Radiation embrittlement + ' . + Fatigue behaviour + ' Stress corrosion cracking behaviour 1) + Ultrasonic testability + + 2) 2) Creep rupture tests + ' + 1) to be performed together with other manufacturers or in a research program

2) HTR specific test

Table 1 Material evaluation for nuclear components Scope of typical tests for base material 168

Tests, scope of testing and test certificates for sheets, plates, bars and forgings of austenitic steels and nickle base alloys

Test Scope Test certificate (see section 336) ace.to DIN 50049 Ladle analysis 1 per heat 2.3 Check analysis 1 piece ' per heat 3. IB Check of heat treatment per heat treatment 3. IB Tensile test at RT 1 per sampling location 3.1 A/C Tensile test at outside temperature 1 per sampling location 3.1 A/C Creep tests 4) 3.1 A/C Notch bar impact tests2) 1 set ISO-V per sampling 3.1 A/C location at RT

Metallographic 1 part per heat and 3.1 B examinations heat treatmentl) a) microstructure b) grain size c) Delta-ferrite"3) Intergranular corrosion 1 part per heat and 3.1 B resistance heat treatment Nondestructive examinations each part see (UT,surface crack) Section 4.3.3 3.1 A/C

Visual inspection every part 3. 1 A/C Identity check every part if no 3. 1 B check analysis available Dimensional check every part 3. 1 A/C 1) at raw weight = 5000 kg on both ends of each part 2) only with thickness = 5 mm 3) only with austenitic steels 4) with semi-finished products used in the creep range it can be agreed with the authorised inspector that creep tests (500-1000 h) can be used to replace or to supplement the hot tensile test

Table 2 extract from a test sheet 169

Type of LUR primary circuit components LUR components class 2 HTR components with examination KTA 3201.3 KTA 3211.3 highest safety conditions

Welding proc- Aim: Qualification of the intended welding procedure General veri ficat ion edure qua i - Welding procedure qualification is performed with the of fabrication e>- i fication test welding procedure, welding parameters, heat treatment perience of manu- parameters and materials intended for the component. facture during No base metal or filler metal batch uniformtta with pressure vessel and component batches required steam boiler manufacture by welding procedure qualification according to AD-HP 2/1 or TRO 201

Validity: 12 months Validity: 24 months requalification, requalificationcan be by restricted to HTR production test application, by HTR-AP

Batch test Aim: qualification of filler Qualification of filler Aim: As with LUR metal batch at filler metal metal batch "at filler primary circuit com- processors under the metal manufacturers. ponents. Additional conditions scheduled under standardized para- reservation of release for the component (KTA K08.3) meters (no relation to by the plant supplier components; KTA 1408.2) with the aim of "batch adjustment" (avoidance of strength tests)

Production Aims: - qualification of component welding Aims: test - requalification of welding procedure - HTR specific addition qualification, where relevant extension of to the available con- scope of welding procedure qualification ventional welding • qualification of welding parameters performed procedure qualifi- with procedures other than those employed for cation the initial welding and for which there is no - as for LUR valid production test production test

Number: according to type of Number: according to A distinction is drawn weld, per seam, per component type of weld based on between .or component set, min. I/year convent ionaI arrangement (AD-HP 2/1), minimum (1) preliminary prod- I/year uction test Performance: Preliminary (simulation heat treated) and Performance as simula- and in addition parallel part of test piece tion or parallel heat- are separated after welding treated test piece (2)paraltel production test

number based on KTA 3211.3

performance of preliminary production test as simulation heat treatment, of the parallel production test as simulated or parallel heat treated test piece

Base metal and filler metal batchesused for the component should be employed. Easements possible in consultation with the authorised inspector

TABLE 3 Bases for welding procedure, batch and production test for LUR and HTR

171

Non-destructive detection of flaws during manufacture and operation of components

F. Walte *)

Fraunhofer-Institut für zerstörungsfreie Prüfverfahren D-6600 Saarbrücken

SUMMARY The nondestructive testing (NDT) in the field of quality assurance of components in the light water reactor (LWR) technology is today very well established, nevertheless not all NDT methods are transformable to the case of high temperature reactor (HTR) components. The reason is the coarse grain structure of the austenitic material used in the HTR-technology in opposite to the fine grain structure in the case of ferritic material in the LWR-technology. Mainly the ultrasonic (UT) testing, which plays the dominant rule in the LWR inspection, is influenced by the coarse grain austenitic or nickelbase alloys structure. The present article analyses the influence of the coarse grain especially the dendritic structure in welds of austenitic and nickelbase alloys, discusses methods and ways to detect flaws in austenitic and dissimilar welds and gives a practicable rule for NDT in the field of HTR components with a combination of ultrasonic and X-ray inspection techniques.

INTRODUCTION

In order to detect flaws in primary components of .light water reactors a large number of NDT methods has been established in the last years. Table 1 shows - divided into pre-service and in-service inspection - the different inspection methods:

TABLE 1: NDT Methods in the field of LWR-inspection

pre-service inspection in-service inspection surfaces magnetic particle test dye penetration test ultrasonic inspection ultrasonic inspection eddycurrent inspection volume X-ray inspection ultrasonic inspection - ultrasonic inspection The main differences between LWR and HTR components relative to the nondestructive testing (NDT) are a smaller wall thickness and the coarse grain austenitic structure in opposite to the fine grain structure of the ferritic material used in the LWR technologies. This austenitic and nickelbase material influences the inspection methods mentioned in Table 1.

*) presentation by W. Kappes 172

The magnetic particle test, used for the surface inspection, is not applicable because of the non-magnetic material. The ultrasonic method especially used for the volume inspection of components (mainly of welds) is strongly influenced by the course grain and particularly the dendritic structure. This behaviour leads to a modification of the inspection concept compared to the LWR-component inspection.

DETECTION LIMITS IN THE CASE OF THE LWR-COMPONENTS INSPECTION

For the surface inspection 4 different methods are available. For thin-walled components with a smooth surface (mean surface roughness - 30 urn) the eddy current test /i/ is able to detect surface cracks down to a minimal crack depth of 0.1 mm. For austenitic weld material especially for austenitic claddings the detectability limit is a crack depth of 0.8 mm (fig. iA).

The dye penetration test /z/ is also able to detect cracks down to a minimal crack depth of 0.1 mm, if the ratio of the crack gap width to the crack depth is smaller than 0.2 (fig. iB).

The surface inspection with ultrasound /}, 4/ shows for the inspection of the non- accessible surface with free waves and the inspection of the accessible surface with guided waves (surface waves) a detectability limit of d . • i V20 (X = ultrasonic wave length). For a wave length of 3 mm assuming'a'surface roughness of S 3oum, dmjn ^ 0.15 mm (fig. 2). All these surface inspection methods - except the ultrasonic corner mirror effect - are not influenced by the austenitic and nickelbase material. For the volume inspection, up to a thickness of 80 mm, the standard X-ray inspection (300 kV-X-ray tube) yields good results with regard to the requirements of the German KTA-rules /s, 6/. For thicker walls it is necessary to use linear acce- lerators (fig. 3 upper part). This technique is expensive and unwieldy and also not useable for in-service inspection. Therefore an alternative method is the ultrasonic inspection.

Depending on the ultrasonic attenuation the ultrasonic test hi is able to inspect thick-walled components in agreement with the German KTA-rules /$/ (fig. 3 lower part); in addition this technique can be applied during in-service inspection.

FLAW DETECTION IN HTR-COMPONENTS

The base to establish a rule for NDT of HTR-components was a collection of documented flaws in HTR-components /8/. This collection has been carried out in the course of the HTR-project with a participation of all HTR project partners and also with important HTR component producers in Germany, Switzerland and Austria.

As expected due to the present welding technology the data collection was not very extensive. Twenty six welding defects were observed and documented; for all hypothetically assumed defects a real example was found.

Figs. 4 and 5 show two examples: lack of fusion in the weld flank and lack of fusion in the weld layer. In the first example the ultrasonic inspection was able to detect the flaw. In the second example an inspection has not been carried out. 173

Fig. 6 shows in the upper part the assumed defects in HTR-components and in the lower part the flaw depth versus the wall thickness. The indications marked with a circle have not been detected during the ultrasonic inspection. These not detected flaws are hot cracks and intergranular stress corrosion cracks. The former are not relevant due to the small crack length of about o.i mm and as for the latter it must be mentioned that the ultrasonic inspection was a conventional technique u- sing 45* shear waves which is not the optimal technique. An accumulation of hot cracks will be avoided through mechanical tests before welding.

In summary one can say, that the essential problem of HTR-component inspection is the volume inspection (weld-volume) of thick-walled components. At increasing wall thickness the detectability of the X-ray inspection becomes more and more worse and the traditional ultrasonic inspection is strongly influenced by the dendritic structure of the material (Fig. 7).

IMPROVEMENT OF THE ULTRASONIC INSPECTION

The austenitic and nickelbase material influences the ultrasonic propagation in a twofold manner Ig, 10/: at first the coarse grains influence the ultrasonic amplitude and the signal-to-noise ratio (scattering at the grain boundaries); on the other hand the dendritic material structure causes a sound beam distortion and fluctuation of the sound velocity (Figs. 8,9).

In the case of the coarse grains, special transducers /n, 12, 13, 14, 15/, the use of longitudinal waves and modern signal processing methods like - signal averaging (Fig. 10) /16/ - amplitude and time-of-flight locus curves (ALOK) /17/ - synthetic aperture focussing technique (SAFT) /18/19/ which were developed in many countries in the last years are helpfull to overcome the bad signal to noise ratio.

With these tools the austenitic and nickelbase material around the weld is inspec- table like a ferritic material.

The case of the weld volume inspection is difficult (Fig. 9, left part). Nevertheless a careful evaluation has shown, that the use of horizontally polarized shear waves (SH-waves) is helpfull to avoid a strong ultrasonic beam distorsion and allows to detect flaws in the weld volume /20, 21A

This special wave mode can not be generated with the traditional piezoelectric transducer concepts, but with electromagnetic transducers (EMAT) /22A Fig. 11 and 12 show an example.

Today with the exception of the ALOK-technique (used in the German LWR inservice inspection) all other above mentioned methods are more or less still in the application test phase. Therefore all these techniques are not directly mentioned in the here recommended rules for austenitic inspection. Nevertheless it is possible to use these techniques if the detectability can be demonstrated using a special test specimen. 174

RECOMMENDED RULES FOR THE INSPECTION OF AUSTENITIC AND NICKEL- BASE HTR COMPONENT MATERIAL

The complete details for NDT of HTR-components subdivided in semifinished products (plates, rods, bars, tubes and forged pieces) and components (mainly austenitic and nickelbase welds and dissimilar welds) are described in /z$/.

For the difficult NDT inspection of austenitic and nickelbase welds the following rules will be recommended:

Surface inspection

In this case it is possible to choose:

A: dye penetration test B: eddy current test C: ultrasonic surface wave inspection Weld volume inspection

For the volume inspection (fig. 13) it is also possible to choose: A: complete ultrasonic inspection B: X-ray inspection and ultrasonic inspection of the weld flanks and ultrasonic inspection of the weld layers C: X-ray inspection of the weld root and dye penetration test during the welding process and ultrasonic weld flank inspection.

Before the inspection, the selected procedure has to be checked on a special weld specimen. A typical test specimen for the ultrasonic inspection is shown in Fig. 14. Any further details are shown in Fig. 15 and 16.

By application of this recommended rules it is possible to detect defects2 in the surface and in the volume as described in Fig. 17. IN-SERVICE INSPECTION

The X-ray and dye penetration tests can not be used in the LWR in-service inspection because of the radioactive contamination. The same restrictions hold for the HTR in-service inspection if such an inspection is necessary.

More details especially for thin {i 10 mm) and medium wall thickness (£ 40 mm) and f°r dissimilar welds are given in the final report "Design Criteria for HTR-Components" This statement is only valid, if the defect has ideal surface conditions (stress free and not filled with corrosion products and other material). 175

As in the case of LWR in-service inspection, only the following methods can be used:

- visual inspection by TV-camera - eddy current inspection - ultrasonic inspection.

If the inspection area is accessible for an automatic scanning system, a NDT inspection is possible with the same restrictions of the ultrasonic inspection described above.

References III Metals Handbook, 8th Edition, Vol. II Nondestructive Inspection and Quality Control American Society for Metals Metals Park, Ohio, 44073, pp 75-92

I2I Metals Handbook, 8th Edition, Vol. II Nondestructive Inspection and Quality Control American Society for Metals Metals Park, Ohio, 44073, pp 20-43 A. Klein, H.J. Salzburger Characterization of Surface Defects by Rayleigh Waves New Procedures in Nondestructive Testing Springer, Berlin, 1983, pp. 193-202

P. Höller, G. Hübschen, H.J. Salzburger Nondestructive Testing of the Inner-Surface Zones of Nozzels, Vessels and Pipes from the Outside Nuclear Engineering and Design %j_ (1985) pp. 193-205 Sicherheitstechnische Regel des KTA, KTA 3201.3 Komponenten des Primärkreises von Leichtwasserreaktoren Teil 3: Herstellung, Kapitel 13

/6/ Deutsche Industrie-Norm (DIN) 54109, 54111, 54112 K. Fischer, H. Wüstenberg, W. Kappes, A. Waas More than 10 Years ISI-NDE on Primary Circuits of German Nuclear Power Plants; Experiences, Results, Developments and further Consequences Proceedings of the 9th Int. Conf. on NDE in the Nuclear Industry, ASM International 1988, 29-34 /8/ F. Walte, IzfP Saarbrücken HTR-Auslegungskriterien, BMFT-Forschungsvorhaben: Auslegungskriterien für hochtemperaturbelastete metallische und keramische Komponenten sowie des Spannbetonreaktordruckbehälters zukünftiger HTR-Anlagen B2-20 - Fehlererkennbarkeitskatalog für austenitische und nickelbasis- legierte Schweißnähte 176

/9/ J.A. Ogilvy The Influence of Austenitic Weld Geometry and Manufacture on Ultrasonic Inspection of Welded Joints Brisith Journal of NDT, May 1987, pp 147-156 /io/ K. Goebbels, G. Deuster, S.E. Greter Ultraschallschwa'chung in Stahl unter Berücksichtigung der Austenite Statusbericht IzfP, Nr. 740209 (1974)

/11/ Handbook on the Ultrasonic Examination of Austenitic Welds The International Institute of Welding, published by the American Welding Society

/12/ X. Edelmann The Practical Application of Ultrasonic Testing of Austenitic Weld Joints Material Evaluation 37, (1979) No. 10 pp 45-51 /13/ H. Wüstenberg, T. Just, W. Möhrle, J. Kutzner Zur Bedeutung fokussierender Prüfköpfe für die Ultraschallprüfung von Schweißnähten mit austenitischem Gefüge Materialprüfung 19 (1977)» Nr. 7, S. 246-251

/14/ K. Goebbels, H. Kapitza Methods for Nondestructive Testing of Austenitic High Temperature Gas-Cooled Reactor Components Nuclear Technology Vol. 66, Sept. 1984, pp. 695-702 /15/ 0. Gangelbauer, F. Wallner, R. Frielinghaus Contribution to the Ultrasonic Testing of Austenitic Welds using Longitudinal Angle-Beam Probes 9th International World Conf. NDT, Melbourne 1978 /16/ V. Schmitz, K. Goebbels Improvement of Signal-to-noise ratio for the Ultrasonic Testing of Coarse Grained Materials by Digital RF-Signal Averaging 1982 IEEE Ultrasonic Symposium Proceedings, pp. 950-953 /17/ O.A. Barbian, B. Groß, R. Licht Signalanhebung durch Entstörung von Laufzeitmeßwerten aus Ultraschall- prüfungen von ferritischen und austenitischen Werkstoffen - ALOK - Teil 1 Materialprüfung 23 (1981) Nr. 11, pp. 379-383

/18/ B.J. Smith Zip-scan - A New Concept in Ultrasonics Instrumentation British Journal of NDT, January 1986, pp. 9-15

/19/ W. Müller, V. Schmitz, G. Schäfer Reconstruction by the Synthetic Aperture Focussing Technique (SAFT) Nuclear Engineering and Design 94 (1986) 393-404 /20/ G. Hübschen, H.J. Salzburger Recent Results in NDT with Electromagnetic Ultrasonic Transducers Review of Progress in Quantitative NDE, Vol. 5B, Ed. by D.O. Thompson and D.E. Chimenti, Plenum Publishing Corporation 1986, 1687-1695 177

/2i/ H. Bohn, M. Kröning, W. Rathgeb, W. Gebhardt, W. Kappes, O.A. Barbian Nachweis zur Leistungsfähigkeit der phasengesteuerten Gruppenstrahler ALOK-Prüftechnik 12. MPA-Seminar, 9./10. Okt. 1986, Stuttgart, 20.1-16 izzf A. Wilbrand Quantitative modeling and experimental analysis of the physical properties of electromagnetic-ultrasonic transducers. in: D.O. Thompson, D.E. Chimenti (Eds); Review of Progress in Quanti- tative Nondestructive Evaluation, Vol. 7A, Plenum Publishing Corp., 1988, 671-680

/23/ Auslegungskriterien für hochtemperaturbelastete metallische und . keramische Komponenten sowie des Spannbeton-Reaktordruckbehälters zukünftiger HTR-Anlagen Endbericht zum Verbund-Forschungsvorhaben des BMFT Band Ha, Teil B: Metallische Komponenten, Aug. 1988 Kernforschungsanlage Jülich GmbH, Institut für Reaktorwerkstoffe 178

limit of Eddy current- corner effect detectability (ultrasonic)

coil Tube inspection S 0.1 mm limit of austenitic cladding detectabitity

dminä0.8 mm X= wavelength for \=3mm PUfe inspection B dmin=0,15mm B ultrasonic surface Dye penetration test waves

OQO — •. -T T_i °min {if th ratio g/diO.2)

Fig, i Surface crack detection Fig. 2 Detection of surface cracks with eddy current and dye with ultrasound. A) non accessible penetration test surface B) accessible surface

t 300 KV Require - recogniz- X-Ray tube mentsof ability german wire i diameter KTA rules [mm] 5 MeV linear accelerator] CONTROL ROD GUIDE ASSEMBLY

100 200

LACK OF Require - recogniz- 1MHz 45° puls echo FUSION abili'ty 3 mentsof german IN THE disk [ike WELD RAM: reflector z rules diameter 1 •• '.

Fig. 3 Limits for the volume Fig. 4 Collection of observed de- inspection of thick-walled LWR-components fects in welded joints of austenitic and nickel base alloys Defect No. Q{ 179 ALack of fusion in AIGSCC the weld flank B incomplete fusion • cracks in the weld root in the welding layer base metal NICROFER ohot cracks • porosity and inclusions 4221 ^cracks in the heat • cracks 1 to the surface weld metal NICROFER affected zone $4225 50 ® Inspe :fed,butnot- detected defect type:lack of fusion in the welding layer 10 a w • s 16mm i '•'•*&*••••" z 1 e 99

[mm] U

0.1 20 40 60 wallthickness [mm]

Fig. 5 Collection of observed defects Fig. 6 Summary of the collection in welded joints of austenitic and nickel of observed defects in welded joints base alloys of austenitic and nickel base alloys Defect No. D -n

accessible surface ultrasonic wave medium '< I volume in the weld weld • \ area of weld flanks

volume in the non accessible weld roof surface

Detect - Dye Eddy X-Ray Ultra- dendritic coarse ability penetration current sound structure grain accessible surface well well - well sound beam distor- strong ultrasonic non well tion scattering äccsessiblc - thin wall well •fluctuation of the low signal fonois; surface components sound velocity ratio weld . : " - well • well •IMPR0VEMENT- flanks in flank direction • use of horizontally • special transducer weld well strong - polarized shear roof - restrict • signal averaging waves (sh-waves) medium well strong • modern signal weld - up tomed. processing wall restrict. volume thickness

Fig. 7 Detectability of NDT tech- Fig. 8 Problems of NDT inspecting niques for the inspection of HTR . austenitic arid nickel base alloys components using ultrasound 180

H -xs. ultrasonic transducer straight beam angle beam f plate-I incidence thickness AS/N=/n n=number of averaged signals

* «

0-1 o : * B s to 0 1*6 3 10 sound path sound path in units of the plate in units of the plate shear horizon fat (sh) wave thickness thickness

Fig. 9 Sound field distortion Fig, io Improvement of signal- caused by dendritic structures to-noise (S/N) by spatial signal averaging

AldSl a «55 mm I «790kH* weld notch 5>t mm

S/N > ROB | natch/\ I

ferritic base austenitic metal base metal

50 tOO ISO 200 250 300 350 £.00 (.50 500 t\ma)

Fig, i r Amplitude locus curve of longitudinal surface defects at the interface, weld metal/ferritic base metal (SH-waves) 181

AldBI ausfenitic ( =790kHz weld edm notch S.trwn notch \S\

ferrittc base austenihc metal base metal

SO 100 ISO 200 250 300 150 UM tSO 500 x[mm|

Fig. 12 Amplitude locus curve of longitudinal surface defects at the interface weld metal/austenitic base metal (SH-waves)

complete ultrasonic inspect. or B B a X-Ray b ultrasound n (weld flanks) 02mm c ultrasound side-drilled hole (weld layer} or the use of ultrasound is possible only a X-Ray (root) if the signal-to-noise ratio (S/N) is : b dye penetra- inspection tion through a.b S/N * 6dS c ultrasound base metal (weld flanks) inspection through c.d.e S/N h 6 d3 weld metal

Fig. 13 Recommended NDT procedures Fig. 14 Requirements for the ul- for the inspection of HTR components trasonic inspection of HTR- ., components 182

defection of cracks in the weld root Angle beam scanning with longitudinal waves primary creeping wave - i1 -, ' shear* / I volume in the 33° wave \ / area of weld \ w flanks secondary creeping wave Longitudinal defect _ r detection defection of incomplete fusion in the weld layers m Transversedefect r1 detection '

Fig. 15 Defect detection in Fig. 16 Longitudinal and trans- austenitic welds using ultrasound verse defect detection in the area of weld flanks using ultrasound

lmin (mm) [mm] Dye penetra- 0.1 1.5 tion test Eddy current 0.8 5 of the watlthickness X-ray min Ultrasound surface

volume (2 MHz long wave! <»5o,areaZ0x20mrr flaw distance

Fig. 17 Detection limits for the inspection of HTR components 183

Discussion of the presentation:

Non-destructive detection of flaws during manufacture and operation of components

Schubert, KFA Julien, FRG: ' Mr. Kappes, I have to thank you very much that you replaced at short notice your colleague Mr.. Walte. Now a short comment: I would like to clarify that if we talk about HTR material, we always mentioned austenitic-steels or Ni-base alloys. Non-destructive test methods for ferrit,ic: steel, especially for the pressure vessel steels are well established within the practice for LWR.

Nickel, KFA Julien, FRG: .

YouTgave some information of the on-going activities concerning the improvements of the current NDT methods for Ni-base alloys. Can you give us an idea about the long term activities in this field. Which results are to be expected to overcome the problems of examination of dendritic structures in weldments and of coarse grained austenitic steels and in Ni-base alloys. What is the overall situation?

Kappes, IzfP Saarbrücken, FRG:' A big=i effort is on-going" in the development of advanced signal processing methods, using the potential of improving the signal to noise ratio. However the spatial signal averageing methods for example ask for a great number of individual ultrasonic-shots, which .have then to be evaluated by electronic processing data. This is, as in the case of LWR components, very time consuming. Therefore these methods are not yet practicable for an in-service inspection. Other methods that have potential for the inspection of coarse grained materials are for example SAFT (Synthetic Aperture Focussing Technique) and ALOK (Amplituden und Laufzeitortskurven-Verfahren); especially the latter procedure is already well established in the field of ISI of LWR components; these procedures can also be used in combination with electromagnetically excited shear horizontal waves that are less influenced by coarse grain structures than other wave types. In this context is must be mentioned that a big problem is the scarcity of specimens with realistic flaws for the austenitic alloys. 184

Hoffmann, CE, USA: Do existing German Codes address differences in the number and interval of examinations for in-service inspection of steel HTR vessels in comparison to LWR vessels?

Kappes, IzfP-Saarbrücken: The recommendation concerning the registration levels for HTR-components are slightly different from those for LWR-components of similiar material, but in principle the standards are the same.

Trumpfheller, Essen, FRG: UT-technigues have the best potential for ISI-testing with remote testing systems. But the noise indications caused by coarse grained structure of austenitic welds or castings decrease the detectibiliby of real defects the indications of which are not considerably higher than the noise indications. Therefore the problem cannot be solved by lowering the sensitivity in elevating the recording level; that would mean, there is no real examination. The present investigations try to find distinguishing properties and behaviour between flaw and noise indication, e. g. by statistical methods and other modes of ultrasonic waves. But there is no sense to have different sensitivity requirement for UT of ferritic and austenitic welds or castings.

Kappes, IzfP-Saarbrücken:

The reason for the problems with austenitic steels is quite simple; the grain boundaries reacts like realistic reflectors. A way to overcome this problem is the use of "pattern recognization algorithms" to do automatic inspection. In this way, many parallel inspection tracks are required. 185

Material Data and Constitutive Equations

H. J. Penkalla Nuclear Research Centre, Julien Institute for Reactor Materials

Abstract

Material data and constitutive equations form the basis for the determination of design values and for the inelastic analysis of the component behaviour under complex loading conditions. For metallic HTR components the materials NiCr 23 Co 12 Mo (Alloy 617), X 10 NiCrAlTi 32 20 (Alloy 800) and X 20 CrMoV 12 1 are selected. The material data are obtained from the test results of different material investigations programmes.

Due to the high application temperatures for metallic HTR components, the main part of the material data consists of creep and fatigue properties. Additional material data are the physical properties, short term properties and fracture mechanics properties. The evaluated data are presented in material data sheets

Constitutive Equations are considered under the aspect of creep under multiaxial and complex loadings. The transference of creep laws derived from uniaxial testings to multiaxial loading conditions is carried out using the von Mises theory and Norton's creep law. 186

1. Introduction

Material data form the basis for the determination of design values and are derived from material test results. The definition of material data which characterize material properties is given by a catalogue of failure modes for a given component loading condition. Therefore material data for design values are upper limits of allowable stresses or strains to avoid

rupture unallowed strain '

Constitutive equations are the basis for the inelastic analysis of the component behaviour under complex loading and for more complex component geometry. Particularly the inelastic analysis at very high temperature is a large field of theoretical and experimental ' investigations today. Two points of view are important for the development of constitutive equations:

the engineering aspect, that means an acceptable effort invested in inelastic analysis

the material understanding as an aspect of material qualification for a specific application.

In this report the material data generation is presented in respect to the procedure of evaluation, the presentation and the " statistical confidence limits. The main aspect of constitutive equations is the transfer of uniaxially determined creep behaviour and creep equations to multiaxial and complex loading conditions. 187

2. Material data

2.1 The materials

For the determination of design values the three materials selected are:

NiCr 23 Co 12 Mo (Alloy 617) X 10 NiCrAlTi 32 20 (Alloy 800(H)) X 20 CrMoV 12 1

NiCr 23 Co 12 Mo is a Ni base alloy with Cr, Co and Mo as solution hardener. Its high creep resistance allows an application temperature up to 1000 °C

X 10 NiCrAlTi 32 20 is a Fe base alloy with a good creep resistance up to 800 °C. Dependent on the heat treatment and the carbon content, three different specifications of this material are discussed /I/.

X 8 NiCrAlTi 32 21 for application temperatures > 500 °C X 5 NiCrAlTi 31 20 DE for application temperatures between 500 °C and 700 °C for superheater tubes of the steam generator X 5 NiCrAlTi 31 20 Rk for application temperature of about 550 °C .

X 20 CrMoV 12 1 is a ferritic steel and well known for application temperatures up to 550 °C

2.2 The material data generation

The material data determined in the project of HTR component design criteria come from test results of different material investigation programmes /2/, from literature and from other 188

sources which work in cooperation with the HTR development partners. The investigation programmes include

tests of physical properties tensile tests Creep and creep rupture tests fatigue tests fracture mechanic investigations corrosion tests

The greatest part of the investigations has been carried out with as-received material, but weldments and aged materials were also investigated.

The test data and test results coming from the different test laboratories were stored and evaluated in special material data bank systems, created for high temperature materials behaviour /3/. Figure 1 shows the typical data flow. A connected system of evaluation programs is developed for technical and statistical analysis for different metallurgical boundary conditions.

data acquisition and documentation

output (data): data bank - tables systems - Plots - statistics "

definitions: output (results): evaluation ' material laws " extrapolation - material properties methods evaluation - Isochronous stress strain curves - criteria for Programms - constants for evaluations material laws - metallurgical =1 - extrapolations boundary conditions etc. etc.

Figure 1: Data acquisition an evaluation" by material data bank 189

2 . 3 The material data presentation

After the evaluation of the material properties the data are summarized as drafts of material data sheets which are part of the final report of the project /4/. Besides general remarks about the application and about further specifications, the data sheets contain the following data in dependence on temperature up to the maximum application temperature:

metallurgical specifications Chemical composition, grain size, heat treatment etc. physical properties E- and G-modulus, specific heat, thermal expansion coefficient, thermal and electrical conductivity, magnetic susceptibility and density mechanical short term properties 0.2% and 1% proof stresses, ultimate tensile strength, rupture elongation and reduction of area, V-notch impact energy (also data for aged material and weldments) creep and creep rupture properties 1% creep strain limit, creep rupture strength, stationary creep rate in dependence on stress, constants for Norton's creep law, isochronous stress strain relationship for creep strain and total strain fatigue properties cycles to failure and crack initiation in dependence on strain range in air and HTR He, maximum stresses in the first cycles of LCF tests fracture mechanics properties constants of the Paris equations for the fatigue crack growth with Kj concept and for the creep crack growth with C concept, fracture toughness by J integral. corrosion behaviour general remarks for corrosion in air, HTR-He and reformer gas 190

welaments general remarks about the decrease of creep rupture strength, detailed data for short term properties and fracture mechanics properties statistics data statistics for different material properties

The drafts of material data sheets are an important advance for the creation of design rules for metallic high temperature components. The data statistics of most important data of short term, creep and fatigue properties, which form the basis for design values, are proofed in a satisfied expense. '

Figure 2 gives a general view about the data available for design. The temperature range of the data supply corresponds with the maximum application temperature of each material.

temperature / C 400 600

physical properties

short term properties — 60 000 h creep and creep — 105 000 h rupture properties — 100 000 h

fatigue properties

fracture mechan. properties

NiCr23 Co 12 Mo X 10 NiCrAITi 32 20 X20CrMoV12

Figure 2: Data supply for the considered materials and different material properties. 191

Of the three materials, X 10 NiCrAlTi 32 20 is the most investigated material. Therefore the evaluation of test results for this material is based on the largest data pool. The alloy X 20 CrMoV 12 1 is in Germany a well investigated material in respect to short term behaviour and creep properties. Data for the fatigue behaviour have a smaller basis. The material NiCr 23 Co 12 Mo has been investigated in a large programme particularly in respect to the HTR development.

2.4 Material data and statistics

As an example of the results of data evaluation the material properties of NiCr 23 Co 12 Mo are presented in this report.

2.4.1 Creep rupture strength

For design values at high temperatures the creep behaviour is very important. Material properties for the determination of the time dependent stress intensity limits are

1% creep strain limit creep rupture strength

In figure 3 the creep rupture strength of NiCr 23 Co 12 Mo is shown for temperatures range between 700 °C and 1000 °C. The drawn curves are mean isothermal values and the marks indicate the real measured times to creep rupture. The mean value has been calculated by MCM time-temperature parameter method.

Figure 4 shows the data statistics of the evaluated semifinished products. 38 different semi-finished products or melts have been tested and evaluated with a total number of 813 specimens. In Figure 5 the distribution of the test parameters time and temperature is indicated. The temperature 192

-h- »_ I700ttC

- 750°C •^ - B00flC 350 °C 300 "C^ 9 CO 350°C" J\ On o ( 300 *C' .- 4 •V o a. J< • if« 102 : 'S "S "H)50"c: • ^i •«. r to —Q -^ o S s* • \, c S , •

fV Mi»« a) * k —4 99U • • \ a. CD \ -••, — CD l_ 05 's S Ü \ s s \ \

10° l 10-1 10° 101 102 103 104 105 time / h Figure 3: Creep rupture strength of NiCr 23 Co 12 Mo

250

0 5 10 15 20 25 30 35 40 heat Figure 4: Data statistics for the Creep rupture strength of the evaluated heats 193

range is 550 °C to 1050 °C and the range of measured creep rupture times is 1 h to 60000 h.

To calculate a minimum va lue of the creep rupture strength, a statistical analysis over the deviation of the real 103 time measured rupture times, related to the calcula- 500 700 900 1100 ted meanvalues has been made (figure 5). The temperature /°C distribution of the Figure 5: Data statistics for scatter- deviations is shown in band analysis of the creep rupture strength figure 6. The minimum value is calculated from this distribution, in the course of which the probability of a single test result being larger than the minimum value is more than 95%.

number of tests 200

150" minimum 100 value 1 50- i—i

0 r—l I 1 :n n'. -0.75 -0.5 -0.25 0 0.25 0.5 0.75

•mean

FiQure 6: Scatterband analysis of the creep rupture strength of NiCr 23 Co 12 Mo 194

2.4.2 Low cycle fatigue properties

Figure 7 shows the number of cycles to failure in dependence on the applied strain ranges and temperatures. The drawn curves represent the mean values which are interpolated by a graphical method and the marks are given by the really measured test results. The isothermal curves form the basis for a fatigue design curve.

03 c 03

0.1

cycles to failure

Figure 7: Cycles to failure in LCF tests of NiCr 23 Co 12 Mo

Figure 8 shows the data statistics for the different semifinished products with a total number of 81 test results. Figure 9 represents the distribution of the number of evaluated tests over the ranges of strain range and temperature. 195

numbe • of tests 40

30- 20-

10- M Ü —. —...... IT"? n y- 1 2 3 4 5 6 7 heat

Figure 8: Data statistics for the cycles to failure of the evaluated heats

number of tests In figure 10, the devia- 4 tion of the- actually- measured number of cycles to failure from its mean value is shown. The definition of the deviation is the same as that for the creep rupture strength. The statistics of the deviations and temp. 0.1 0.3 0.6 1.0 1.5 3.0 the calculated minimum / C strain range / % values are shown in figure 10. * ' Figure 9: Data statistics for scatterband analysis of the cycles to failure 196

number of tests 40

30 J

20 _ minimum value i 0 -1.5 -1 -0.5 0 0.5 1 1.5 N F.meas log F.mean

Figure 10: Scatterband analysis of the cycles to failure of NiCr 23 Co 12 Mo

3. Constitutive equations

In the project of design criteria for HTR components different constitutive equations were discussed and partly verified- For constitutive equations many estimates exist today. They are divided in two principal categories:

models separating time independent plasticity and time dependent creep

unified models which define only an inelastic strain.

One of the discussed unified models is the Interatom model /5/. An important model separating creep and plasticity is the ORNL model /6/, developed for the steels of type AISI 304 and 316. To apply this model to the materials intended for HTR components, verification of the model is required. 197

In this report the aspects of creep, particularly the transfer of creep laws, determined in uniaxial tests, to multiaxial and complex loading conditions is considered.

3.1 General basis

To transfer uniaxial developed creep -laws to multiaxial .loading conditions some assumptions are given. The main postulate is the analogy of Hoff flI which defines a similarity between plasticity and creep. , The formalisms of plasticity .are applicable for creep prediction if the time independent values are replaced by their time deviations. The plastic strain must be replaced by the creep strain rate and the dissipation energy by the dissipation rate.

Hoff's analogy can be applied on the von Mises theory /8/ to define a mean stress. In this theory each stress tensor leading to a definite dissipation energy has the same mean stress. For creep the mean stress is defined by the dissipation rate, where each stress tensor leading to a definite dissipation rate is represented by the same mean stress.

Further postulates are:

constant volume This requirement implies that creep deformation leaves the volume unchanged. This is equivalent to the use of the stress deviator for determining the permanent deformation.

compatibility This implies that the strain rate parallel to a plane immersed arbitrarily in the material does not exhibit any discontinuity, i.e. that two material zones do not slide upon each other. The stress distribution establish under creep adapts to this requirement. 198

isot ropy Creep behaviour is identical in each direction. This requirement is not satisfied for every semi-finished product under consideration, but it serves as a first approximation.

creep law This represents here a mathematical function of the relation 6 = f(a). There has been no standard formulation to date for the entire stress region since even the creep mechanism are different, depending upon the stress involved. For the stress to be studied, Norton's creep law for stationary creep /9/ has proved to be generally applicable, the general formulation being:

es,eff = kavn

where av is the von Mises deviatoric stress.

Cylindrical coordinates are introduced for the tubes under discussion in the following. If for a given component load OJJ is the associated stress tensor, and 0^1* is its deviator, the application of the above mentioned postulates leads to the following equations for the compatibility of stresses and the strain rate tensor

- = -(<7 " CT ) dr r U r

1+v. * l-2v- 3 - i * a-. + S6- . + -ka""1 a E XJ E 1^ 2 v X3 M

3.2 Verification of the model

For the verification of the transfer model mentioned here, as a first step Norton's creep law for stationary creep has been fitted to results from uniaxial creep tests. Figure 11 shows 199

10 NiCr23Co12Mo 5- CO heat ADL 2- CO 2 i - CO 5-

1 10 5-

0 10 10"8 io7 ^(f io5 10"3 io'2 minimum creep rate / %/h

Figure 11: Stationary creep rate of NiCr 23 Co 12 Mo dependent on the uniaxial stress

the relation between the stationary creep rate and the applied stress for one heat of NiCr 23 Co 12 Mo. In the main programme different loadings for tube material were defined as combination of

internal pressure tensile load torsion load

A complete survey of selected loading conditions is given in table 1. The internal pressure is always a static load, tensile and torsion, however, can be applied as static, cyclic or relaxing or strain controlled load. The resulting multiaxial creep behaviour under these loading conditions were 200

internal tensile torsion pressure stat. cycl. relax. stat. cycl. relax. X X X X X X X X X X Y X X y X X X X X X X y X X X X X X X X X X

Table 1: Table of loading conditions of tube components for testings and calculations

calculated with the mentioned formalism and tested on IHX and reformer tubes.

The loading conditions selected for this purpose were orientated towards emergency situations with correspondingly high stresses and temperatures. The maximum temperatureselected was 950 °C and the stressing is chosen so as to produce deviatoric stresses of the order of 20 to 30 MPa.

3.3 Results

As an example of the results of the verification /10-12/, a loading condition consisting of a constant tensile stress and a constant shear strain is considered. This condition leads to the situation that the shear stress relaxes while the axial strain rate decreases. In figure 12. the relaxation of the shear stress in dependence on the time is shown for two experiments. In one experiment the axial tensile stress is 201

a = 0 and in the other experiment the tensile stress is a = 21 MPa.

shear strain const. shear stress relaxing tensile stress a const.

,0 -

a = 21 MPa

theoretical value foric= o

10 20 30 40 time

Figure 12: Influence of a constant tensile stress on the relaxation of the shear stress

In the experiment with a superimposed tensile stress the relaxation rate is higher than in the experiment without tensile stress. Otherwise the superposition of a torsion load to a tensile load leads to an increase of the axial strain rate. If the shear stress relaxes form an initial value the axial strain rate decreases asymptotically to a level given by the strain rate under pure tensile stress.

The interaction between torsion and tension is caused by the deviatorideviator:c stress av. The relaxation of the shear stress is given by

i = -(-) -/a+ 3T2' n"1 kET where T increases with a. The axial strain rate is given by 202

2 2 n 1 Gz = /a + 3t ~ ka

The axial strain rate decreases with the shear stress 7".

Experiments and calculations show in all considered loading conditions a good agreement. Nevertheless, the results are strongly dependent on the structural stability of the semifinished material. Changes in grain size and second phase precipitation in the initial stage of an experiment have a influence of the test results.

4. Conclusions

The evaluation of experimental test results for the candidate materials for HTR components are summarized in drafts for data sheets. They contain all relevant data for determining design values and for stress-strain analysis of a component. The data sheets contain values and statements concerning

physical properties mechanical properties creep and creep rupture properties fatigue properties fracture mechanics properties

The amount of data allows a reliable design by rules.

Constitutive equations are being developed in further investigations all over the world. The application of von Mises theory for multiaxial creep and of Norton's creep law for high temperature is verified with a satisfactory agreement between calculations and test results. 203

5. References

/I/ H. Diehl Langzeit-Festigkeitskennwerte des Werkstoffs 1.4876 - Auswertung der Ergebnisse aus Zeitstandversuchen am Werkstoff X 10 NiCrAlTi 32 20 zur Ermittlung der Zeitstandfestigkeit und 1%-Zeitdehngrenze BBC-HRB-Bericht GHRA 001760, 1986

/2/ H. Nickel, T. Kondo, P.L. Rittenhouse "Status of Metallic Materials Development for Application in Advanced High-Temperature Gas Cooled Reactors" Nucl. Techn., July 1984, Material Selection I, p. 12-22

/3/ H.J. Penkalla "Die Werkstoffdatenbank der Entwicklungsgemeinschaft Hochtemperaturreaktor - Ein Instrument zur Erstellung von Auslegungsdaten und zur inelastischen Analyse" VDI Berichte Nr. 600.4, 1987, p. 283-309

/4/ Endbericht zum Verbund-Forschungsvorhaben des BMFT "Auslegungskriterien für hochtemperaturbelastete metallische und keramische Komponenten sowie des Spannbeton-Reaktordruckbehälters zukünftiger HTR-Anlagen" Band II,. 1988

/5/ B. Böcke, F. Link, G. Schneider, D.B. Schneider "New Constitutive Equations to Describe Infinitesimal Elastic-Plastic Deformation" . PVP Conf. Orlando, Florida, ASME 82 TVP - 71, 1982

/6/ J.M. Corum, W.L. Greenstreet, K.C. Liu, C.E. Pugh, R.W. Swindeman "Interim Guidelines for Detailed Inelastic Analysis of High Temeprature Reactor System Components" ORNL-Report 5014, 1974 204

111 N.J. Hoff "Approximate Analysis of Structures in the Presence of Moderately Large Creep Deformations" QAM 12, 1954, p. 49-55

/8/ R. v. Mises "Mechanik der festen und flüssigen Körper im plastisch deformablen Zustand" Königl. Ges. der Wiss., Göttingen, 1913

/9/ F.H. Norton "Creep of Steels at High Temperature" Me. Graw Hill, New York, 1929

/10/ K. Franzke, H.J. Penkalla, M.Rödig, F. Schubert, H. Nickel "Untersuchungen zum Kriechverhalten von Rohren aus X 10 NiCrAlTi 32 20 (Nicrofer 32 20) im Anlieferungszu- stand bei mehrachsiger Belastung" Jül-Bericht 2127, 1987

/ll/ H.J. Penkalla, H. Nickel, F. Schubert "Kriechverhalten von Rohren aus X 10 NiCrAlTi 32 20 und NiCr 23 Co 12 Mo unter mehrachsiger statischer und zyklischer Belastung" 13. MPA-Seminar, Stuttgart, Oktober 1987

/12/ H.J. Penkalla, F. Schubert, H. Nickel "Alloy 617 Tubes at High Temeperature" Conf., Proc. of Superalloys 1988 in Seven Springs, Met. Soc, 1988, p. 2-12 205

Discussion of the presentation:

Material data and constitutive equations

Dahl, TU Aachen, FRG:

What is the meaning of %TV ? Mean stress or effective stress?

Penkaiia, KFA Julien, FRG: The mean stress is identical to the deviatoric stress or effective stress under the assumption of constant volume. With the v. Mises hypothesis £*v is only a function of the second invariant of the stress deviator. 206

Methods for Very High Temperature Design

J. J. Blass, J. M. Coruro, and S.-J. Chang Engineering Technology Division Oak Ridge National Laboratory Oak Ridge, Tennessee, U. S. A.

Abstract

Design rules and procedures for high-temperature, gas-cooled reactor components are being formulated as an ASME Boiler and Pressure Ves- sel Code Case. A draft of the Case, patterned after Code Case N-47, and united to Inconel 617 and temperatures of 982°C (1800°F) or less, will be completed in 1989 for consideration by relevant Code committees. The purpose of this paper is to provide a synopsis of the significant differences between the draft Case and N-47, and to provide more complete accounts of the development of allowable stress and stress rupture values and the development of isochronous stress vs strain curves, in both of which Oak Ridge National Laboratory (ORNL) played a principal role. The isochronous curves, which represent average behavior for many heats of Inconel 617, were based in part on a unified constitutive model developed at ORNL. Details are also provided of this model of inelastic deformation behavior, which does not distinguish between rate-dependent plasticity and time-dependent creep, along with comparisons between calculated and observed results of tests conducted on a typical heat of Inconel 617 by the General Electric Company for the Department of Energy.

1. INTRODUCTION

In the U.S.A., components of nuclear reactor systems are designed in accordance with applicable provisions of the ASME Boiler and Pressure Vessel Code. For Class 1 Components at relatively low temperatures

*Research sponsored by the Office of Advanced Reactor Programs, U.S. Department of Energy, under contract DE-AC05-S4OR21400 with Martin Marietta Energy Systems, Inc.

"The submnad manuKfipt tot tow auttiored by i contractor of the U.S. Government undo« contract No DC- ACO5-84OR214O0 Accordingly, the U.S. Government retains a nonancluaiva. rovany-rree keen*« 10 Dubhth or reproduce trie published form of thts contribution, or Mow others to do so. lor U.S. Gov«rnm«nt purpos«*.' 207

[371°C (700°F) or lower for ferritic steels and 427°C (800°F) or lower for austenitic steels and high-nickel alloys], these provisions are in Subsection NB of Section III of the Code. For Class 1 Components at higher temperatures, Code Case N-47 is applicable. The alloys permitted by the current edition of this Case1, and the corresponding temperature limits for a service life of 300,000 h or less are listed below.

Temperature Limits

2.25Cr-lMo 593 (1100) 800 H 760 (1400) 304 or 316 816 (1500)

In response to needs expressed by the General Electric Company and the U.S. Department of Energy, an ad hoc committee of the ASME Code was established in 1983 to formulate design methods and procedures for certain gas-cooled reactor components (steam-methane reformers) to operate at temperatures of 950°C (1742°F) or less. This committee, known as the Task Force on Very High Temperature Design (TF-VHTD), reports to the Subgroup on Elevated Temperature Design of the Subcommittee on Design.

After a period of extensive review and careful cons 5 derat ion of relevant issues, the committee decided to propose a new Code Case, patterned after relevant portions of Case N-47, and limited to Inconel 617, temperatures of 982°C (1800°F) or less, and service lives [total time at temperatures above 427°C (800°F)] of 100,000 h or less. Central to the new Case is a design-by-analysis concept addressing the following possible failure modes in high-temperature service: (1) ductile rupture from short-term loadings, (2) creep rupture from long-term loadings, (3) creep-fatigue failure, (4) gross distortion due to incremental collapse and ratchetting, (5) creep buckling due to long-term loadings, (6) loss of function due to excessive deformation, (7) buckling due to short-term loadings, and (8) non-ductile rupture. A draft of the new Case has been in preparation for two years, and is expected to be completed in the first half of 1989 for consideration by relevant Code committees.

A synopsis of significant differences between the draft Case and Case N-47 is provided in Section 2. More complete accounts are given in 208

Sections 3, 4, and 5 of three related developments for the new Case in which ORNL played a principal role. Section 3 concerns the allowable stress and stress rupture values which are the bases for limits on load- controlled (primary) stresses. Section 4 describes the unified constitutive model of inelastic deformation behavior which was used as a basis for isochronous stress vs strain curves. Section 5 covers the isochronous stress vs strain curves which are used in simplified evaluations of ratchetting and creep-fatigue. Concluding remarks are given in Section 6.

2. DIFFERENCES BETWEEN DRAFT CASE AND N-47 •

As stated• in the Introduction, the draft Code Case is patterned after relevant portions of Code Case N-47, and thus consists of three articles and two appendices. The numbering system follows that of Sub- section NE of the Code. Articles -1000 Introduction and -2000 Materials are each relatively brief, the former consisting of one subarticle defining the scope of the Case and the latter consisting of three para- graphs supplementing article NB-2000 Materials of Subsection NB. Arti- cle -3000 Design is quite extensive and mostly self-contained. Appendix I provides relevant material specifications and properties for Inconel 617 and Appendix T provides certain evaluation procedures and criteria, including those related to ratchetting, creep-fatigue, buckling, and welds.

Since the intended application of the draft Case is design of steam-methane reformers, no component-specific rules are provided for piping, pumps, and valves, such as those of Case N-47. Thus draft Article -3000 Design consists of only three subarticles: -3100 General Requirements for Design, -3200 Design by Analysis, and -3000 Vessel Design. As consequences of the intended application, there are many differences in the content of corresponding portions of the draft Case and Case N-47. Most of these differences are material and temperature related, and the most noteworthy are in Subarticle -3200 Design by Anal- ysis or its related Appendices I and T. These are summarized below. 209

There are a number of differences related to the unique inelastic behavior characteristics of Inconel 617 at temperatures close to 98 2°C (1800°F). These characteristics include: (1) lack of a clear distinction between time-independent (elastic-plastic) behavior and time- dependent (creep) behavior, (2) great dependence of flow stress on strain rate, and (3) softening with time, temperature, and strain. Thus provisions or limits of Case N-47 that are based on time- and rate- independent, or strain hardening, idealizations of material behavior are not included or are notably altered in the draft Case- The constitutive models of inelastic behavior employed in any required inelastic analyses should also reflect the above characteristics.

There are sorae notable differences in the limits of -3220 applied to load-controlled (primary) stresses that are obtained from elastic analyses. In the application of these limits certain terms are used whose definitions are given in the draft Case and in Case N-47, and are generally consistent with those of Subsection NB. Thus P and P^ are general and local primary membrane stress intensities and P. is primary bending stress intensity. Five categories of loadings are considered: Design and Service Levels A, B, C, and D. The Units are based in part on values of S , S t and Sc given in Appendix I. SQ is based on short- term tensile properties, and on 105-h creep and creep-rupture properties. S is the lesser of S , based on short-term tensile properties, and S , based on data from constant-load creep tests.

The differences in the limits of -3220 between the draft Case and Case N-47 are in the definition of S and in certain of the limits applied to Design Loadings and to Service Level D Loadings. In Case N-47, the stresses corresponding at time t to a total strain of 1%, to the initiation of tertiary creep, and to rupture are considered in Sc (and hence also in S ). In the draft Case, the stress corresponding to the initiation of tertiary creep in time t Is not considered. In Case

N-47, the Design Limits are given by P _< S and ?L + P. <_ 1.5 SQ. In the. draft Case, SQ is replaced by the greater of SQ and SmC for the design life. In Case N-47 the Service Level D limits are obtained in part from Appendix F of Section III for P and PL + Pfe. In the draft 210

Case the limits from Appendix F are replaced by 70% of the lesser of the collapse load and the plastic instability load. The draft Case also cautions that instability can arise from material instability associated with strain softening as well as from structural instability.

It should be noted here that, because of the intended application of Inconel 617, the draft Case does not address cladding in Paragraph -3227, does not permit bolts in -3230 unless they satisfy the limits of Case N-47, and cautions against significant reductions in fracture toughness from extended exposure to elevated temperatures in -3240.

The draft Appendices I and T are analogous to those of Case N-47. Appendix I consists entirely of information specific to Inconel 617. Since bolts are not permitted by the new Case, material specifications and allowable stresses for bolts are not included in Appendix I. Stress rupture factors, for weldnents are being developed and are expected to be available for inclusion in the first draft of Appendix I. Appendix T also includes information specific to Inconel 617, such as fatigue design curves (Fig. T-1420-1) and isochronous stress vs strain curves in T-1800. The former are used in creep-fatigue evaluations based either on inelastic analysis (T-1420) or elastic analysis (T-1430). The latter are used in simplified evaluations of ratchetting (T-1330) and creep- fatigue (T-1430). A creep-fatigue damage interaction diagram for Inconel 617 (Fig. T-1420-2) is being developed and is expected to be available for inclusion in the first draft of Appendix T. In accordance with the coverage of the draft Case, Appendix T will contain no specific provisions for piping and bolting.

3. ALLOWABLE STRESS AND STRESS RUPTURE VALUES

In -3220 of the draft Case, limits are placed on load-controlled (primary) stresses that are obtained from elastic analyses. The limits are based on values of SQ, SmC, Sm, Sc, and Sr, which are provided in Appendix I. S is based on criteria given in Division 1 of Section VIIt (Pressure Vessels). The values of S were obtained from Code Case 1956- 1 and from another Case presently under development, which provide for 211

use of Inconel 617 in Division 1 of Section VIII. The development of Che other values is described below.

The criteria employed for the draft Case are as follows: , ( • S rat

2/3 of specified minimum tensile strength at room temperature [655 MPa (95 ksi)] 11/30 of minimum tensile strength at temperature s.£ 2/3 of specified minimum yield strength at room temperature [241 MPa (35 ksi)] 90% of minimum yield strength at temperatures up to 649°C (1200°F); 2/3 of minimun yield strength at temperatures above 649°C (1200°F) 2/3 of minimum stress to rupture in tine t

Minimum stress to produce 1% strain in time t Minimum stress to rupture in time t Sr "

The above criteria are similar to those employed for Case N-47 except that the above criteria for Sc do not include 80% of minimum stress for onset of tertiary creep In time t because increasing strain rate is observed very early in constant-load creep tests of Inconel 617 at temperatures close to 982°C (1800°F).

For Sm the minimum yield and tensile strength values up to 760°C (1400°F) were consistent with values used for S in Code Case 1956-1. Above 760°C (1400°F), average values fron Huntington Alloys2 were multiplied by the corresponding ratio (0.64 for yield strength and 0.86 for ultimate tensile strength) of Code minimum to Hunt ington average at temperatures below 760°C (1400°F). Smoothed values of S were obtained m from a cubic equation fitted by least squares to values obtained from the above procedure for temperatures in the range 21°C (70°F) <_ T <_

982°C (1800°F). For Sm in ksi and T in °F*, the equation may be written

*To convert temperature from °C to °F multiply by 1.8 and add 32. To convert stress from ksi to MPa multiply by 6.895. 212

5 2 8 3 Sa = 33.06 - 0.0313T + 3.034 x 10" T - 1.168 x 10" T

For Sc Che minimum stress (Sr) Co rupture in time t and Che minimum stress to produce 1% strain in time t were obtained from plots of the Larson-Miller parameter from Huntington Alloys.2 In each case, the minimum was taken to be 1.65 standard deviations below ehe mean. Isochron-

ous curves of Sc vs temperature were drawn to converge on a value of 248 MPa (36 ksi) at 316°C (600°F), which is an estimate of the (tine- independent) minimum stress corresponding to 1% strain at this temperature. This estimate is based on the average stress vs strain curve for a strain rate of 0.005/min and a temperature of 649°C (1200°F), the variation in yield stress with temperature, and the ratio of minimum to

average yield stress. Figure 1 is a plot of S , and of St for given times» vs temperature; and Fig. 2 is a plot of S vs time for given

temperatures. Sr is plotced vs time for given temperatures in Fig, 3.

60C 1600 1800 2000

TEMP.(F)

Fig. 1. Sm and St for given times vs temperature 213

100 1000 10000 100000 1000000

LOAD DURATION, t, IN HOURS

Fig. 2. S vs time tor given temperatures.

1 00 1000 10000 100000

TIME (h)

Fig. 3. S vs time for given temperatures 214

4. UNIFIED CONSTITUTIVE MODEL

Limits on deformation-controlled quantities in -3250 of the draft Case and Case N-47 may be satisfied by criteria in Appendix T based on elastic, simplified Inelastic, and detailed inelastic analyses. Consti- tutive equations are used in detailed inelastic analyses of structural components to accurately model the inelastic response of structural materials to nultiaxial loading histories. In uniaxial form, the equations may also be used to develop some of the design tools that are used in simplified methods of analysis. The equations given below were used to generate isochronous curves of stress vs strain for a typical heat of Inconel 617. These curves were used as a basis for the average isochronous curves in T-1800 of the draft Appendix T.

Constitutive models of inelastic behavior are typically based on classical concepts of time-independent plasticity and time-dependent creep. In such models, inelastic strain is considered the sun of separately evaluated plastic and creep strains. However, for Inconel 617 at temperatures close to 982°C (1800°F), this distinction between plasticity and creep is unrealistic; inelastic behavior is always significantly time- or rate-dependent. To properly characterize such behavior, a so-called unified, or viscoplastic, constitutive model is much more appropriate. Such a model can provide a useful description of both short- and long-term behavior as a function of loading rate.

The unified constitutive raodel developed by Robinson3 at ORNL has been used to represent important behavioral features of structural alloys at high temperature. In this model, nultiaxial flow and growth equations are derived from a potential function of applied and internal stresses. The growth equation is of the widely accepted Eailey-Orowan type, and thus contains two terms corresponding to the competing mechanisms of hardening and recovery.

This mathematical framework, of coupled partial differential equations was used as the starting point for development of unified equations for a typical heat (XX63A8UK.) of Inconel 617 based on the results of tensile and creep tests conducted by the General Electric Company for the Department of Energy. 215

The resulting uniaxial equations for monotonic loading may be written as follows:

2 ,n/2 Flow Equation -e. = F[(a - a) - R2 ] (a - a) Growth Equation Ö = He /a1"5 - Ro1"5 Total Strain Equation e = a/E + e. where F = 1.704xl0"Cö + 5). m H = 10 R = 0.46*10(ra " 6 0.35 $ » 1 + 3.16xl03e° n 0 o — o n = { n + 0.08(a - 0 and the parameters E, K., m, n 6 and a are given in Table 1 for the temperature range 649°C (1200°F) to 982°C (1800°F). In the above equations, e. is inelastic strain in %, a iis applied stress in ksi, a is internal stress in ksi, and time is in h.

Table 1. Material Parameters for Inconel 617

Temperature E K a °C <°F) (ksi) (ksi) m no 6 (fisi) 649 (1200) . 246.000 10.000 3.800 4.000 5.000 10.000 677 (1250) 242.844 9.250 3.353 3.750 4.486 12.500 704 (1300) 239.625 8.000 3.000 3.500 4.000 15.000 732 (1350) 236.344 6.250 2.743 3.250 3.542 17.500 760 (1400) 233.000 4.000 2.585 3.000 3.113 20.000 788 (1450) 229.594 ' 3.369 2.286 2.875 2.568 20.625 816 (1500) 226.125 2.825 2.009 2.750 2.067 22.500 843 (1550) 222.594 2.369 . 1.756 2.625 1.611 25.625 871 (1600) 219.000 2.000 1.526 2.500 1.200 30.000 899 (1650) 215.344 1.719 1.319 2.375 0.833 35.625 927 (1700) 211.625 1.525 1.135 2.250 0.511 42.500 954 (1750) 207.844 1.419 0.974 2.125 0.234 50.625 982 (1800) 204.000 1.400 0.836 2.000 0.001 60.000

These equations differ from some other applications of Robinson's model in two respects. The softening function 4> was introduced into the 216

flow equation and into the recovery term of the growth equation to effectively model both the stress drop after yielding observed in constant-strain-rate tensile tests and the accelerating strain vs time observed in constant-load creep tests. The exponent n in the flow equation was given a dependence on applied stress to more effectively model the inelastic response at higher values of applied stress.

The form of the above equations was guided to a large extent by current understanding of the mechanisms governing inelastic deformation on the microscale. The values of various parameters were obtained by a process of trial and error guided by experience. The differential equations for inelastic strain and internal stress were numerically integrated for tensile and creep loading conditions and the results visually compared with appropriate test data. Adjustments were made in values of parameters until satisfactory agreement was obtained. Comparisons of experimental and analytical results for tensile loading conditions are shown in Figs. 4-6, and for creep loading conditions in Figs. 7 and 8.

5. ISOCHRONOUS STRESS VS STRAIN CURVES

As stated above, the unified model was used to generate typical isochronous curves of stress vs strain that served as a basis for the average isochronous curves in T-1800 of the draft Appendix T. The latter curves are required for simplified evaluations of ratchetting (T-1330) and creep-fatigue (T-1430).

Development of the unified constitutive model presented above was based on the results of tensile and creep tests conducted on a single, typical heat of Inconel 617. However, the hot tensile (0.005/inin) and isochronous stress vs strain curves in the draft Appendix T were required to represent average behavior for a large number of heats of Inconel

*The subroutine RKF45 was used for integration. As described in Chapter 6 of Ref. 4, it is based on the Runge-Kutta formulae of Fehlberg, which are of fourth and fifth order. These provide for a step-by-step solution of the initial value problem, with automatic control of step size to achieve the required accuracy. 217

PREDICTION EXPERIMENT

U.U3 — ^— tn — —~ ------— o 1- 0.003 — —, in — Q.0003 — aLJ: o If— . . . —- — • T-TZ i—• 1 > 1 0.0 0.1 0.2 0.3 0.1 0.5 0.6 0.7 O.B 0.9 1.0 TOTHL STRflIN iV.) Fig. 4. Calculated and observed stress vs strain at 75O"C (1382°F) for given strain rates (per min). . 0 6 in PREDICTION o EXPERIMENT

o (KS I 0 3 f

/;•• . 0 3 TRES S in Ä • 0.003 a ' i ------0.0003 — — o -

o u>

o o / 0.0 0.1 0.2 0.3 0.1 0.5 0.6 0.7 0.8 0.9 1.0 TOTflL STRfllN iV.) Fig. 5. Calculated and observed stress vs strain at 850°C (1562°F) for given strain rates (per min). 218

o (£1- PREDICTION EXPERIMENT

. 0 28. 32. "— f ~ • - -~ — (K b I / tn M3. 0 2 in LJ v :r o

0.005 o \- fM-

o 0.0005

f I.—

o 0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 TOTflL STRflIN I'A) Fig. 6. Calculated and observed stress vs strain at 95O~C C 1742"F) for given strain rates (per min). 1 i / PREDICTION li EXPERIMENT ! f / / / j / y a: 7 a: if /V •'e.y in • 7 / , LJ / ; / y 1 •••

0.0 1.0 2.0 3.0 1.0 5.0 6.0 7.0 8.0 9.0 10.0 TIME (100 HRSJ Fig. 7. Calculated and observed strain vs time at 850°C (1562°F) for given stresses in ksi. 219

ft 1 / / — PREDICTION I / — EXPERIMENT I / 5 z4.5,' 1 / /I // in 1 CJ 1 / j / LJ 1 7. / / I -----

O.O 0-5 1.0 1.5 2.0 2.5 3.0 3.5 1.0 VS 5.0 TIME (100 HRS)

Fig. 8. Calculated and observed strain vs time at 950°C (1742°F) for given stresses in ksi.

617. Accordingly, tensile and isochronous stress vs strain curves generated by the unified model for given temperatures were adjusted to fit average behavior for those temperatures based on average values from Huntington Alloys2 for 0.2% offset yield strength and stress to produce 1% strain in a given time. This was accomplished by translating a model tensile curve along the elastic line until it passed through the corresponding average yield stress and by translating a model isochronous curve along the elastic line until it passed through the corresponding average stress for 1% strain. The resulting average curves for given temperatures and times are shown in.Figs. 9-15., . -

6. CONCLUDING REMARKS

t A draft ASME Code Case for Class 1 Components of nuclear reactor systems, patterned after relevant portions of Case N-47 and limited to Inconel 617 and temperatures of 982°C (1800°F) or less, will be completed in 1989. Accounts were provided in this paper of the differences between the draft Case and Case N-47, and of the development for the draft Case of allowable stress values, a unified constitutive model, and isochronous stress vs strain curves. 220

00 - i 1 1 1 1 J 1 1 „_

MATERIAL - Ni-Cr-Co-Mo (ALLOY 617) 90 - e TEMPERATURE - 1200 F

80 - -

70 -

•i\\ —

60 -

10Mi_^ 50 -

40 -

30 - 10* h — 20 - 10s h "" 10 - r — 0 - 1 -\ 1 T— " 6 0.21 0.41 0.61 0.81 l'.O1 1.21 1.4 1 .'6 I'.S 2\0

STRAIN C*4 Fig. 9. Average isochronous stress vs strain at 649°C (1200°F) at given times.

The draft Case will not be without certain shortcomings. It is, after all, the recent work of a small committee, based on limited available information concerning the behavior of Inconel 617 at very high temperatures. It will be some time before the new Case will be as well established as Case N-47, which has evolved over about 20 years (taking the preceding Cases 1592 and 1331 into account).

The draft Case also inherits known shortcomings of Case N-47. For example, the creep-fatigue criteria in Appendix T of Case N-47 are generally considered inadequate, even though some improvements are currently being made in the rules for both elastic and inelastic analysis. Improvements are also being made in the simplified ratchetting procedures in Appendix T of Case N-47. However, the improved procedures will still require that the average temperature across the wall at one extreme of each secondary stress intensity cycle be less than the temperature for 221

100

MATERIAL - Ni-Cr-Co-Mo (ALLOY 617) 90 TEMPERATURE - 1300°F

80-

0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0

STRAIN (%) Fig. 10. Average isochronous stress vs strain at 704°C (13OO°F) for given times.

J i 1 L 50 MATERIAL - Ni-Cr-Co-Mo (ALLOY 617) 45 - TEMPERATURE - 1400JF

40 -

35 -

30 -

25 -

20 _

15 -

10 -

5 -

0 0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 Z.O STRAIN (%) Fig. 11. Average isochronous stress strain at 76O°C (I4UÜÜF) for given times. 222

MATERIAL - Ni-Cr-Co-Mo (ALLOY 617) 45 -l TEMPERATURE - T500*F

HOT TENSILE (0.005/rain)

0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0

STRAIN [%) Fig. 12. Average isochronous stress vs strain at 8I6°C (I5OO°F) for given times. 50 MATERIAL -Ni-Cr-Co-Mo (ALLOY 617) 45- TEMPERATURE - 1600=F

40-

35 - HOT TENSILE 0.005/min 30 -

25-

20 -

15-

10-

5 -

s 0 10 h 1 1 1 r —I 1 1 r 0 0.2 0.4 0.6 0.£ 1.0 1.2 1.4 1.6 1.8 2.0 STRAIN {%) Fig. 13. Average isochronous stress vs strain at 871°C (16008F) for given times. 223

26- __i 1 1 1 1 1 1— 1 i i_

/ V HOT TENSILE (0.005/min) • 24- / ^

22- 0.1 h ^—•-

20 • •

18 - X MATERIAL - NT-Cr-Co-Mo (ALLOY 617) / TEMPERATURE - 17OOaF 16 -

14 - 1 h .— 100 0 ps i 12 -

tESS . '• ^-— 10 - 10 h _ _- 8 -

J 6 - 10 h . 3 4 - V^-~— 10 h in- h 2 -|£— 10s h r 0 - 0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0 STRAIN (5f) Fig. 14. Average isochronous stress vs strain at 927"C (1700°F) for given times.

HOT TENSILE (0.005/min)

MATERIAL - Ni-O-Co-Mo (ALLOY 617) TEMPERATURE - 1SOO;F

0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0

STRAIN (X) Fig. 15. Average isochronous stress vs strain at 982°C (1800°F) for given times. 224

s which S = St at 10 h. A simplified procedure without this requirement is needed to avoid more costly detailed inelastic analyses. This need is even greater for the new Case, as a consequence of the intended application to steam-methane reformers.

REFERENCES'

1. Case N-47-27 > Class 1 Components in Elevated Temperature Service,

Section III3 Division 1, ASMS Boiler and Pressure Vessel Codey American Society of Mechanical Engineers, New York, 1988.

2. T. H. Bassford, Inconel Alloy 617, Huntington Alloys, Inc., Huntington, WV, June 1982.

3. D. N. Robinson, "Constitutive Relationships for Anisotropie High- Temperature Alloys," Nuclear Engineering and Design, 83 (1984) 389- 396.

4. G. E. Forsythe, M. A. Malcolm, and C. B. Moler, Computer Methods for Mathematical Computations, Prentice-Hall, 1977. 225

Discussion of the presentation:

Methods for very high temperature design

Schuster, KFA Julien, FRG: In the new code case for INCONEL 617, is there a specific consideration of preageing effects on structure, and in consequence, on mechanical behaviour? Is any consideration of the effects of corrosion on mechanical behaviour to be included?

Blass, ORNL, Oak Ridge: The draft of the new Code Case states that these factors are to be taken into account, but does not provide specific guidance on how to do so.

Seehafer, INTERATOM, FRG: You presented stress design limits. What about strain limits above all concerning weldment design philosophy (Reduction)?

Your presentation indicated that the new code case does not allow bolts; is this strictly the case or is it temperature dependent?

Blass, QRNL, Oak Ridge: Given that the new code case is patterned after Case N-47, the presentation concentrated on notable differences between the two. The new case, like N-47, will have strain limits in Appendix T and weldment stress reduction factors in Appendix I.

The new case does state that bolts are not permitted unless they satisfy the limits of case N-47.

Penkalla, KFA Julien, FRG: Two questions concerning the constitutive equation for Alloy 617. Is the model an empirical model or derived from hardening rules? 226

The variable«^ is declared as internal stress. In the initial ORNL-model **C is the centre of the yield surface for kinematic hardening. Are both definitions of the same character?

Biass, ORNL, Oak Ridge: As stated in the full text of the paper, the form of the equations was guided to a large extent by current understanding of the mechanisms governing inelastic deformation on the microscale. The values of various parameters were obtained by a process of trial and error guided by experience. The differential equations were integrated numerically given loading conditions and the results compared with test data. Parameter values were adjusted to obtain satisfactory agreement.

In the contexts of the models (the earlier one makes a distinction between plastic and creep strains and the present unified model does not), the variables designated by the symbol ©< are similar in certain respects.

Blumer, Sulzer Winterthur, Switzerland: The unified equation shows the necessity of a time-dependent material description, also for very short times, above 800 °C. What then is the value of retaining "plastic" approaches for primary stress limits, like

Sm or App. F. for faulted conditions? (e. g. it matters whether the faulted condition duration is 10 min. or 1 to 2 hours).

Blass, ORNL, Oak Ridge: In its consideration of these issues, the Task Force on Very High Temperature Design was faced with precedents established by Subsection NB and Case N-47 of the ASME Code. The Task Force recognized the unique inelastic behavior characteristics of Inconel 617 at temperatures close to 982 °C (1800 °F) but wanted the new Case to make use of the proven methods of Subsection NB and Case N-47. Thus a unified constitutive model is to be used in inelastic analyses and limits are placed on load-controlled (primary) stresses obtained from elastic analyses. As in N-47, the primary stress limits for Service Levels A, B and C are based on the time-independent allowable Sm and the time-dependent 227 allowable St- For Service Level D, the limits of N-47 based on minimum stress to rupture in time t are retained. However the limits of N-47 from Appendix F of Section III for Pm and P[_ + P5 are replaced by 70 % of the lessor of the plastic collapse load and the plastic instability load. The method used to determine these loads should account for rate-dependence, time-dependence, and material instability. 228

Life fraction rules

K. Maile

Staatliche Materialprüfungsanstalt Stuttgart

Evaluations for lifetime estimation of high temperature loaded HTR-components under creep fatigue load had been performed. The evaluations were carried out on the basis of experimental data of strain controlled fatigue tests with respectively without hold times performed on material NiCr 22 Co 12 Mo (Inconel 617). Life prediction was made by means of the linear damage accumulation rule. Due to the high temperatures no realistic estimates of creep damage can be obtained with this rule. Therefore the rule was modified. The modifications made at several institutions consist in a different analysis of the relaxation curve including different calculation of the creep damage estimate resp. in an extended rule, taking into consideration the interaction between creep and fatigue. In order to reach a better result transparency and to reduce data set dependent result scattering a round robin with a given data set was carried out.

The round robin yielded that for a given test temperature of T = 950 °C a realistic estimate of damage can be obtained with each modification. Furthermore a reduction of resulting scatterbands in the interaction diagram can be observed, i. e. the practicability of the rule has been increased.

Moreover new ways of further improving the rule regarding the material physics background were shown. 229

1 Initial situation and task

in literature a multitude of procedures for life assessment of components under creep fatigue load are mentioned. As examples can be.given:

- frequency-separation-method as per Coffin - Manson's strainrange partitioning method - linear damage accumulation rule

The linear damage accumulation rule is the most frequently employed procedure for lifetime analysis and is referred to in most guide lines, e. g. - TRD 508 FRG - ASME CC N 47-17 USA - RCC-MR F The reasons for the utilization of the linear damage accumulation rule are listed below: - wide international, employment - simple handling - high flexibility - simple tests for verification - determination of the dominant damage mode The aim of the works was thus to use resp. modify the linear damage accumulation rule with given data sets for HTR-components and HTR specific materials. The evaluations were carried out as a Round Robin /I/. The following institutions were involved: - ASEA BROWN BOVERIE ABB - HOCHTEMPERATUR REAKTORBAU HRB - INTERATOM IA - Gesellschaft für Reaktorsicherheit GRS - Rheinisch Westfälischer Technischer Überwachungsverein RWTÜV - Kernforschungsanlage Julien - Staatliche Materialprüfungsanstalt Universität Stuttgart MPA

The results of each institution should be compared with one another and evaluated.

•2 Problems with the employment of the linear damage accumulation rule The general formula of LDAR is like this:

D = Z. nL/Nit + LL tci/tRi = DF + Dc ni number of cycles under creep fatigue conditions :Ntf number of cycles to failure at same strain range at pure fatigue loading 230

tci time under creep load at creep fatigue tRi rupture time under creep load at D ASME CC N 47 nomenclature: total damage parameter in German usage: total damage estimation parameter or creep fatigue usage factor

The evaluation of damage due to fatigue damage DF and accumulated creep damage Dc will be separated by the rule. The total damage D is determined by adding the two terms. The evaluation is most frequently made on the basis of strain-controlled LCF-tests with hold time, compare Fig. 1.

The rule is based upon the assumption that the individual damage estimation terms are independent of one another. For specimen failure D-value 1 is defined. However, experiments have shown that D-values may considerably deviate from the theoretical value 1. Therefore, in design codes often values < 1 are used.

Due to possible interaction between fatigue and creep the rule can only be regarded as an approximate procedure. Nevertheless, the method is the most common procedure for lifetime analysis and is mentioned in most guide-lines. Investigations on ferritic and austenitic steels have shown the necessity for modification of the rule to obtain conservative results.

The problems arising in connection with the application of the rule, especially for the evaluation of creep loading, are the following:

- how are cyclic hardening and softening considered ? - how can creep during the relaxation phase be treated ? in the frame of the round robin, different proposals, which partly can be traced back to literature, had been worked out hereto, which subsequently will be discussed and presented.

Results of the Round Robin 3.1 Data material

As data material Interatom test results from LCF-hold time tests at NiCr 22 Co 12 Mo trademark Inconel 617 were employed. The tests were carried out at 950 °C. The creep fatigue tests were performed in helium, the creep tests in air. The strain rate at fatigue tests was 4-10'3 s~1. The hold times of 1, 3, 10, 30 and 120 minutes were implemented at strain maximum. The tests were carried out at a total strain range of 0.3 resp. 0.6 %. A maximum test time of about 500 h resulted. 3.2 Modifications effected at LDAR

In the frame of the round robin the necessity for modification of the LDAR became apparent. The calculation of realistic and accurate creep damage of specimens tested at high temperature is dependent 231

upon a qualified data set, especially on relaxation data. Fig". 2 schematically shows a relaxation curve resp. its change under consideration of the influence of various test parameters. A scatterband emerges, which can be depicted by means of a mean value curve. As to the determination of creep damage the question arises, which course of relaxation is to be fundamentally. The following possibilities, the participants of the Round Robin, mentioned in brackets had laid claim to, are at hand:

- not taking into consideration the mentioned effects by stipulating a mean relaxation curve, which is valid for a certain kind of material at a certain temperature (ABB /2/, HRB

- consideration of effects at cycle - resp. test-dependent evaluation, taking the experimentally determined relaxation curve as a basis (MPA /4/) - regard of relaxation curve at a fixed, cycle number, normally at n/Nf = 0.5 dependent on strain range (Interatom , RWTÜV For calculation of creep damage the following methods were chosen, Fig. 3:

- definition of an "equivalent" stress a9q for the total course of reaxation (ABB, HRB) (fig. 3a) - integration of the relaxation curve (ABB /I/, HRB /2/) (fig. 3b)

When testing the material NiCr 22 Co 12 Mo at 950 °C an additional temperature dependent problem arises, which doesn't exist at lower temperatures. As shown in /5/, the initial stress relaxation is very rapid starting with 200 Mpa. The creep rupture curve of this material covers only stresses lower than 80 Mpa with rupture times of tR > 10 h. Therefore, it is not possible to make an evaluation of stress relaxation within the first seconds due to the high stress magnitude. The following modifications therefore were proposed:

- differentiation between a plastic part, and -a creep part of the relaxation curve. For this it is necessary to set a time for separation. This time is now the basis for> calculation of creep damage. For the choice of this "starting time" no rule is available so far. There is the possibility to cut off after certain periods of time e. g. 1 min or 5 minutes (ABB /I/, HRB /2/, MPA /3/). A consideration of structural events in the material is possible on the basis of the deformation rate. In this way, the first part of the relaxation curve is interpreted as plastic deformation due to the fact that the deformation rate is much 232

higher in comparison with the stationary creep rate. The part of the relaxation curve with deformation rates in magnitude order of the minimum creep rate is regarded as decisive with respect to creep damage (INTERATOM /5/, MPA Diminishing of creep damage, achieved by cutting the relaxation curve, is compensated by increasing fatigue damage. This can be done by choosing a hold time related fatigue curve as reference curve (hold time corresponds to cutting off time (MPA, HRB, ABB)). Another possibility is to exend the strain range by extrapolation of the hysteresis loop in the tension hold time range (IA).

modification of the damage rule by introducing an additional parameter c (t). Fig. 3d (RWTÜV /6/). This parameter controlls the time-differential dt/t depending on the hold time. This way, a strain rate dependent evaluation of creep damage is possible, Fig. 3d (RWTÜV /6/). The parameter c(t) describes the interaction between plasticity (fatigue) and creep.

4 Comparison of the modifications

In. Fig. 4, the result of an evaluation on the basis of an equivalent stress and the consideration of a starting time after one minute is illustrated (ABB /2/). A mean relaxation curve independent of the strain range was chosen as a basis. The total damage factor D assumes an average value of 0.93. The integration of the mean relaxation curve doesn't reduce the result scattering to a large degree. The total damage factor D assumes smaller values due to the reduction of creep damage part Dc. Even the evaluation of HRB /3/, Fig. 5, is based upon a mean relaxation curve which is valid for all LCF-tests. The mentioned starting time as per Fig. 3c was determined at 3 resp, 10 min. In the reference curve for the determination of the fatigue damage factor Df hold times, adjusted to the starting times were taken into account. The values for D are close to 1.

The difference in results between the two procedures is negligible as to be seen in Figs. 4 and 5. This may, in the first place be traced back to the determination of a mean relaxation curve. MPA evaluations /4/ based on a similar procedure, using actual relaxation curves, taken from respective tests, prove that this leads to much bigger scattering at the values of the creep damage factor. The determination of a mean relaxation curve and also of a mean creep rupture curve hence presents a "manual" minimization of the scattering so that both the experimentally and structurally conditioned scatterings have no major impact on the evaluation.

Investigations by Interatom /5/ on round robin material have shown that the deformation behaviour during relaxation below a stress value of a = 50 Mpa, nearly corresponds to that of a creep test. A 233

stress dependent separation of the relaxation curve in a plastic fatigue range resp. a creep range applies to the evaluations of IA. Creep damage in the remaining part of the relaxation curve is determined through integration, whereby manual averaging resp. a manual adjustment of relaxation curves was carried out. As reference curve for the fatigue part a fatigue curve is used which is delineated over the total plastic strain range. The damage diagram thus received is depicted in Fig. 6. The determined values lie below D = 1. A reduction of the creep damage factor's scattering is also reached by fixing the starting time over the section point of tangents, which comprises the relaxation course, (MPA /4/), Fig. 7.

The procedure of RWTÜV /5/ which was carried out according to Fig. 3d presupposes the following dependencies:

- relaxation curve - cycles to failure vs. hold-time - curve - creep rupture curve

The evaluation can only be carried out numerically by means of integration. The relaxation course was averaged from test data at n/Nf = 0.5 depending on the respective strain range. In Fig. 8, the course of damage factors of both the modified and the conventional damage accumulation rule are shown. Increasing hold time resp. decreasing deformation rate from the right to the left side corresponds to the curve's course.

The calculated damage Dc and Dr as per the conventional and modified damage rule, show totally different courses in the damage diagram with regard to the magnitude of the values. The calculated total damage factor according to the modified rule remains under the defined damage limit D = 1, contrary to those of the conventional ones.

5 Conclusions The modifications effected by the participants of the Round Robin have revealed, that the LDAR has a high potential for development. With reference to - detailed data evaluations resp. data processing - reproducability of the particular procedures - consideration of strain rates at relaxation, i. e. interaction between creep and fatigue, the modifications carried out are exclusively related to the creep damage part. Here it turned out that as far as the data material at hand permits the determination of a material , specific and temperature-dependent relaxation mean value curve, the total damage factor assumes values of < 1. A requirement, however, is that the relaxation course is divided in a plastic and a creep part. The deformation rate during the relaxation phase can be considered as a 234

criteria for splitting up. Thus the magnitude of strain rate during relaxation is to be in order of magnitude of creep rate from creep tests. Regarding the material at hand this is accurate for stresses < 50 N/mm2 at a test temperatures of 950 °C. However, this value is dependent on structure physics. For specimens of different melts this value may vary. Another possibility is to consider the creep damage part of the relaxation course by using a hold time and strain rate dependent integration method. This method has the advantage that no empiric thresholds or mean values for fixing starting time have to be used. The works have proved, that the modifications effected have extended the area of application of the LDAR in high temperature regime. Moreover, new ways were shown to achieve improvements of the rule with respect to structural behaviour of the material. 235

Literature

Reports working package B5 - Schadensakkumulation

/I/ Maile, K.: Endbericht über die Ergebnisse des Arbeitskreises B5-Schadensakkumulation, 1988

/2/ Raule, P.: Ringauswertung von LCF-Versuchen mit Haltezeit an IN 617 bezüglich linearen Lebensdaueranteil-Regeln, 1987 B5-9.5

/3/ Diehl, H.: Ringauswertung von LCF-Versuchen mit Haltezeit bezüglich linearer Lebensdaueranteil-Regel Ringauswertung von LCF-Versuchen mit Haltezeit bezüglich linearer Lebensdaueranteil-Regel

/4/ Obst, V.: Anwendung verschiedener Auswerteverfahren für vorliegende Ermüdungsversuche mit Haltezeiten. Ringauswertung vorge- gebener Versuchsreihe mit dem Ziel der Darlegung der Aus- wertemethode, ihrer Besonderheiten und ihrer Genauigkeiten. Auswertungen nach dem "Tangentenkonzept" und nach dem "Stütz wertekonzept". B5-9.4b, BMFT-FV: HTR-Auslegungskriterien

/5/ Meurer, H. P.: Berechnung der Schadensakkumulation bei kombinierter Kriech- Ermüdungsbelastung: Beitrag zur Ringauswertung aus der Sicht der Werkstoffabteilung B 5-9.16, 1987 see also: Meurer, H. P; Breitling, N; Dietz, W; Influence of 'Hold Time and Strain Rate on the LCF-Behaviour of Alloy 617 at 950 °C 2nd Int. Conf. Low Cycle Fatigue and Elasto Plastic Behaviour of Materials, Munich, Sept. 1987

/6/ Lehmann, H.-J.: Ringauswertung zur linearen Schadensregel für den Werkstoff Inconel 617 B 5-9.3, 1987 236

«H V WTIME Ni cyclic hardening

TIME

FIG. 1 Low cycle fatigue tests

Influence of: 1 totalstrain range \ strain rate \ ? c/) temperature 1 < UJ cyclic hardening ) material H- I— cyclic softening ) , (/) if)

FIG. 2: Influence of test parameters 237

FIG. 3a: "Equivalent" stress

< ÜJ or a:

= /dt/tR (a) ..

FIG. 3b: Integration of the relaxation curve 238

if)

FIG. 3c: Separation of the relaxation curve

Interaction term c (t) Alloy 617, T=950 N 1.0 • f

0,8

0,6

0,4

oAf.»0.3 . 4 . 10"3s"' 3 1 0.2 „ Aft = 0.6 » 4 • 10" s" 1 o Aft 0.3 = 2 • 10" V 1 ..*••" * Aft 0.6 = 3.3 • io- V 10 10 10' 10 Hold time t u in min c (t) Dc (tH) • Df (tH) a D (tH)

FIG. 3d: Additional time dependent parameters 239

• 2 e, = 0,3 V. A 2e, = 0,6 V. öcq Beding / Conway

o 2e,=0.3V. A 2E, = 0,6 V. öcq Integration a> en holdtime £ •a Q.

0,2 0,*, 0.6 0,6 Fatigue damage Dp

FIG Results of ABB (mean relaxation curve, equivalent stress, starting time)

1.0

o t[| = 3 min Q 0,6 = 10 min Q) 0,6

CD average value 0 = 0,93 Oj2

0 0 0,2 0A 0,6 0,8 i.o Fatigue damage DF

FIG. 5: Results of HRB (mean relaxation curve, equivalent stress, starting time) 240

1.0 - " Act = 1,5 V. • • Ae, s 0,5 V.

- .30 • o A Ei = 0,3V. *«• e =

A* O £**2.10-Vi 0)

cn • E (0 0.5

o .10 • 3 Q3

.120 * 10 A3 ,30 A 1 A i t i 1 . 1 . 0.5 1.0 Fatigue damage DF FIG. 6: Results of IA (separating on the basis of creep deformation behaviour)

Hold time A (30)

A(30)

Q 10° 0) U) 03 E OJ X)

10-2 i i i i i i 11 1 1 1 1 ! 1 1 1 1 m 1 10' 10" 10° 10 Fatigue damage Dp FIG. 7: Results of MPA (actual relaxation curves) 241

Modified life fraction rule Conventionel life fraction rule

( Aet=0,37v

2 % 0.6 0.8 0,2 0/, 0,6 0,8 1 Fatigue damage Dp Fatigue damage DF

FIG. 8: Results of RWTÜV (additional time dependent terms) 242

Discussion of the presentation

Life time fraction rules

Schuster, KFA Julien, FRG: Would it not be better to call the "plastic" part of relaxation during hold period in fatigue loading "behaviour deviating from creep behaviour"? Plastic, i. e. time independent-deformation behaviour, is related to a change in total strain, but total strain is kept constant during the hold period.

Maile, HPA Stuttgart, FRG: I agree with this statement. The definition "plastic" was only chosen to show the difference to the "creep" part and was not referred to changes in the microstructure. 243

The present status of research and development works for the preparation of the high temperature design code

Y. Muto Y. Kaji Thermal Structure Laboratory

Y. Miyamoto ', . HENDEL Development Laboratory H. Nakajiraa ' , .... Material Performance & Testing Laboratory 0. Baba HTTR Designing Laboratory

Presented at the Workshop on Structural Design Criteria for HTR held January 31 to February 1, 1989 in Julien, Federal Republic of Germany 244

ABSTRACT

A structural design code for the HTTR components is prepared in JA ER I based on rules of both the Elevated- Temperature Structural Design Guide for Monju and the ASME Code Case N-47 an(i on material data of Hastelloy XR. Research and development works have been conducted to ascertain the validity of the rules for the material and loads in the service condition of HTTR. These works consist of creep tests, creep-fatigue tests, biaxial creep tests, weldment tests and component tests. Test objectives, test parameters, test apparatuses and time schedules for these are introduced. In addition, some test results are described. A tertiary creep behavior and creep-damage criterion are discussed based on a stress-controlled creep test of Hastelloy XR at 900°C. Creep-fatigue life-prediction methods are compared using fatigue-test results with both fast and slow strain rates or with both tension and compression hold times. A result of bending creep-fatigue test of heat transfer tubes is described. 245

Contents 1 . Introduction 2. Scope of research and development works 2.1 Tensile and creep strength (1) Tensile strength (2) Creep strength 2.2 Creep-fatigue strength (1) Creep-fatigue test in helium (2) Creep-fatigue test in vacuum (3) Fatigue test in air 2.3 Multiaxial behavior (1) Biaxial creep test (2) Biaxial creep fatigue test 2.4. Fracture toughness (1) Effect of thermal aging 2.5 Weldment property (1) Weld joint stress-rupture test (2) Inelastic behavior of weldment (3) Tube butt-weld stress-rupture test () Dissimilar weld test 2.6 Component tests (1) Heat transfer tube bending test (2) IHX structural model test (Tube) (3) IHX structural model test (Manifold) (i) Creep buckling test 2.7 Tests for an advanced design code (1) Creep-damage test (2) Crack-initiation test 3- Test results 3.1 Creep-damage test 3.2 Creep-fatigue test 3.3 Heat transfer tube bending test 246

1. Introduction

A structural design code for the High Temperature Engineering Test Reactor(HTTR) developed by the Japan Atomic Energy Research Institute (JAERI) was prepared based on both the Elevated-Temperature Structural Design Guide for Monju(ETSDG)[ 1 ] and the ASME Code Case N-4.7. A maximum service temperature is 950 °C and Hastelloy-XR is a candidate material for the high temperature components in this reactor which are not covered by the codes. Then, material data have been experimentally obtained. In addition, some structural tests and component tests are undertaken or planned to ascertain rules in the codes. In parallel with these efforts, some other works are undertaken to rationalize rules in the above-mentioned codes and to establish more advanced design code. The objectives, the parameters and the apparatuses for these tests are briefly described. The time schedules are shown in Table 1. In addition, recent results of the tests in a Thermal Structure Laboratory are also shown.

2. Scope of research and development works 2.1 Tensile and creep strength (1) Tensile strength Short-time tensile tests for solution annealed Hastelloy-XR were conducted at temperatures ranging from room temperature to 1000°C with different strain rates. Preliminary tests for the aged ones were also made. As to the shape of stress strain curves, no work hardening behavior was reconfirmed at and above 850°C, this nay be attributed to dynamic recrystallization during testing.

(2) Creep strength Long-term creep tests on Hastelloy XR at temperatures up to 1000°C have been conducted in JAERI- type B standard helium. Some of the tests are exceeding 2.9x10^ h as of December 1988. A test data base has already been prepared for the performance in air up to 2.4xtO^ h. One of the important observations is that no statistically significant diffrence has been recognized between the helium and air environments in the creep rupture performance of the alloy. Improved oxidation resistance of the alloy is believed to have contributed to the results, and such a feature is a special advantage in prediction of the long term performance in the service environment. More detailed creep rupture behavior on the alloy in several kinds of helium environments with 247

different impurity compositions will be presented in this workshop.

2.2 Creep-fatigue strength (1) Creep-fatigue test in helium To estimate creep-fatigue interaction quantitatively, fatigue tests were conducted focusing on the effects of strain rate and hold time at temperatures ranging from 700°C to 900°C. As to the effect of strain rate, decreasing the strain rate led to notable decreases in the fatigue life. Based on the shapes of stress strain curves and the crack morphology, a considerable contribution of creep damage was included at lower strain rates. In the experiments with the trapezoidal strain waveform with different hold types, the fatigue life was reduced most effectively in tensile hold-time experiments, while no effect was observed for merely compressive case. Preliminary calculations by the cumulative damage rule are underway for creep-fatigue damage evaluation. (2) Creep-fatigue test in vacuum Fatigue tests with triangular or trapezoidal waves for Hastelloy XR were conducted at 850°C mainly in vacuum. The tests are now continued for Hastelloy XR at 95O°C. This is a contract work assigned to Mitsubishi Heavy Industries, Ltd.(M.H.I) and is planned to be completed in September 1989. The program aims at developing or selecting the most appropriate creep-fatigue life prediction method. (3) Fatigue test in air Fatigue tests in air are carried out in order to get design fatigue curves at high temperature in high strain rate and at low temperature. The very high cycle fatigue tests at room, temperature is done in load controlled condition, and the strain controlled condition is applied for all other temperatures from 400°C to 1000°C. The strain rate is 10~-V-sec. The tests are assigned to the Kawasaki Heavy Industries, Ltd. (K.H.I). The tests are planned to be completed in March 1989- 2.3 Multiaxial behavior (1) Biaxial creep test Combined tension and torsion creep tests at 900°C in air were conducted- in the Toyohashi University of Technology. A 248

flow rule of Von Mises was confirmed to be valid for the range of experimental parameters[2]. A test machine will be installed at the Thermal Structure Laboratory at the end of March in 1989- A succeeding test aims at examining the effect of combined compressive and torsional stress on creep behavior and rupture.

(2) Biaxial creep fatigue test The tests consist of constant load creep test and combined load test of constant axial load and load-controlled cyclic torsional load where thin-walled tubular test specimen is used. The objective of tests is to examine the biaxial creep-fatigue behavior preliminarily. The tests are assigned to the Ishikawaj ima Harima Heavy Industries, Ltd. (I.H.I) under a contract with JAERI. Though there was some delay due to machine trouble, the test is planned to be finished in March 1989-

2.U Fracture toughness (1) Effect of thermal aging Tests up to 1x10^h have been carried out in connection with the influences of aging temperature, creep deformation under stress, and thermal cycles. Some of the heats have shown a serious loss of dynamic fracture toughness after aging. Detailed ana lysis of the fracture surfaces has revealed that the loss of toughness and ductility was attributed to precipitation of M/C carbide along the grain boundaries occurring during the aging. Preliminary results of JJQ tests were obtained.

2.5 Weldment property (1) Welded joint stress-rupture test Stress-rupture tests are conducted at 950°C in air. Test specimens were machined from TIG-welded rolled HastelloyXR plate of 20mm thickness. Rupture data up to 6000h are expected to be obtained by September 1989. Rupture strengths of welded joint showed equivalent strengths to those of base metal at high stresses, but they showed gradual reductions for the lower stresses. Rupture elongations of weld metal specimens were very small.

(2) Inelastic behavior of weldment Creep test and Creep-fatigue test for plate type specimens of TIG weldments are planned. A surface of the 249

specimen will be inspected by interrupting the test. Structural design method for the weldment will be examined by comparing the test result with an inelastic calculational result based on creep constitutive equations for both base metal and weld metal.

(3) Tube butt-weld stress-rupture test Test specimens were machined from TIG butt-welded joints made of medium boron HastelloyXR heat transfer tubes of O.D.31.8mm x t3.5mm and were used for the stress-rupture tests at 95O°C. Only short time tests were conducted hitherto, and a creep-rupture life was identical to that of the base metal and rupture occurred at base metal. The tests were conducted cooperating with I.H.I.

(4.) Dissimilar weld test Dissimilar welded joints of 2-1/£Cr-1Mo steel with Hastelloy XR are to be used in the intermediate heat exchanger (IHX) in the HTTR. The tests were planned to optimize weld conditions and to confirm the mechanical properties, mainly tensile strength and rupture elongation. The tests are assigned to M.H.I and are expected to be completed in March 1989. 2.6 Component tests (t) Heat transfer tube bending test A servo-controlled 4-point bending machine was developed. The test machine and test result are shown in Sec.3.3. (2) IHX structural model test (Tube) A model structure for the IHX as shown in Fig.1 consists of a manifold, 8 helical tubes and 8 connecting tubes, which are installed in an electrically heated retort. The test apparatus is shown in Fig.2. Each connecting tube is attached to the manifold in one end horizontally and the other end is vertically extended to the upper loading grid. These tubes are subjected to both a reciprocal load and an internal pressure of'/, MPa. These loads induce combined cyclic bending and steady'tensile stresses in tubes. Test temperature is 950°C and test environment is helium gas. Fabrication of the model was finished at I.H.I in February 1989. Photo 1 shows a tube header section of the manifold. The test apparatus is now under construction also at I.H.I and will be completed in July 1989. 250

(3) IHX structural model test (Manifold) The same test structure is utilized for the tests as for the tests in Sec.2.6(2). The objective of this test is to examine the creep deformation behavior in a cylindrical ligament structure like the manifold. Biaxial stress is generated by an internal pressure of helium gas.

(4.) Creep buckling test The tests aim at obtaining data concerning creep buckling time of Hastelloy-XR cylinder {O.D.31.8mm x t2mm) at high temperatures due to outer pressure. Diameter, thickness and ovality of the test specimens are varied. The test apparatus was completed in October 1983. A test specimen with a complete circular cross section was tested at 95O°C and 5 MPa and ruptured after 13 hours. A next specimen with some out of roundness is now tested. The result was compared with both the results of finite element analysis and simplified analysis.

2.7 Tests for an advanced design code (1) Creep damage test The tests aim at clarifying creep strain hardening behavior, a time to the onset of tertiary creep and a damage accumulation rule. Stress controlled creep tests were carried out. Test results are discussed in Sec.3-1 • (2) Crack initiation test The tests aim at clarifying a cracking behavior in high temperature alloy and characterizing the crack initiation. Notched thin plate specimen(t2mm) of Haste Hoy XR was used. Crack initiation and growth were observed by an optical method in real time. Both the frequency and the hold time dependences.of the creep-fatigue life were clarified and the relationship between a notch opening displacement rate and a crack growth rate was obtained[3l-

3. Test results 3.1 Creep-damage test Creep strain behavior and creep damage characteristic of medium boron Hastelloy XR at 850°C and 900°C were examined by stress-control led creep tests. First, creep constitutive equation was made based on the Garofalo-Kachanov formulate] [5] using creep data obtained under the constant stress, that is, 251

rt e = e0 + e^d - e- ) + grain t + etertiary CD and r 2>) 6 = £^re~ + 6min + tertiary ^ where, £1, r, £min and £tertiary are a primary strain parameter, an asymptotic parameter, a minimum creep rate and a tertiary creep strain respectively. Stress dependences of the former three parameters were easily determined by plotting the experimental strain against stress as shown in Fig.3, Fig.4 and Fig.5 though £1 data had a considerable scattering. For the tertiary creep, Kachanov formula was applied, which lead to the following equation.

tr - t (3) •min \tr " t where tr is equal to a time to rupture and t3 is equal to a time to onset of tertiary creep shown in Fig.6 and in Fig.7, respectively. An experimental data plot is shown in Fig.8. Consequently, the following creep constitutive equation is obtained.

0 < t < t3 U)

£ = et(1 _ e-rt) + emin t

2 0023 £-! = 5.^O33x1O" e-°- 6 • r = 1.2U1x1O"3 e0'0530^ = 1.9761x10-^ 66'0830

,£(t-t3)

11 5 0106 t3 = 1.1088X10 Ö" ' (6) These, equations can predict fairly well the creep strain 252

behavior for the constant stress condition. Secondly, examination of the creep strain behavior for the variable stress condition is under way. Thirdly, creep damage was examined. The following four types of variable stress tests were conducted. (a) Constant load creep test. The tests equal to gradually stress increasing test, and (b) Graduary stress descending.test, (c) Once stress increasing test and (d) Once stress increasing test. Summations of both the time fraction and the strain fraction up to rupture time were calculated. The results are shown in Fig.9 and Fig.10. Rupture elongation data used for the calculation is shown in Fig.11. From these figures, it is concluded that the strain fraction rule (ductility exhaustion rule) is more adequate compared with the time fraction rule (Robinson rule) in the range of experimental data.

3.2 Creep-fatigue test Fatigue tests with triangular or trapezoidal waveforms as shown in Fig.12 were conducted for the medium boron Hastelloy XR at 850°C. The test objective is to select or to develop the most appropriate creep-fatigue life prediction method. Fig.13 and Fig.14. show the creep-fatigue damage evaluation based on the time fraction rule. Creep damage was calculated using the creep rupture data under the constant load in Fig.13 and under the constant stress in Fig.14-^ The following are observed. (i) The evaluation using the constant stress data results in the better prediction of creep fatigue life. (ii) An environmental effect is observed. An air environment yields the shorter creep fatigue life compared with a vacuum environment. (iii) The range of variation of the calculated creep damage is larger than that of fatigue damage. (iv) Creep fatigue damage seems to depend on whether the load is symmetrical or non-symmetrical. Values of calculated creep damage range from 0.1 to 1 for the non-symmetrical loaded data shown as the group of triangular or semi-circular symbols. On the other hand, they are equal to 1 approximately for the symmetrical loaded data. This sujests that life is determined by fatigue phenomenon for the symmetrical load and an influence of creep is mainly effective in the case of non- symmetrical load. 253

Predictions of the cycles to failure were tried by three methods. Figure 15, Fig.16 and Fig.17 show the predicted cycles to failure versus experimental cycles to failure by the time fraction rule, a strain range partition method[6] and a damage rate equation[7] respectively. From these figures, the time fraction rule seems to be ostensibly inferior to the other two methods. However, the following comments should be pointed out. (i) In actual loading cycles, all cycles have non- symmetrical waveforms because down-trangent for the shut-down are different from an up-trangent for the start-up, (ii) There is a possibility of adhesion of cracked surfaces under compressive stress in the vacuum environment. This phenomenon is scarcely probable for the actual helium gas environment. (iii) If we observe the results except for data both under the symmetrical load and under the compressive load in vacuum, the variation of prediction for the time fraction rule is not worse than those for the other methods, (iv) Calculation practices for the actual complicated cycles are easier, and not ambiguous, for the time fraction rule compared with those for the other two methods. Consequently, we consider that more investigations are necessary in order to be able to select the most suitable one. Therefore, the creep-fatigue tests for the more complicated cycles are under way.

3-3 Heat transfer tube bending test. The aim of the tests is to obtain the failure data of actual heat transfer tubes for the IHX under the cyclic load. The test apparatus is a survo controlled bending machine as shown in Fig.18. A test tube is horizontally installed on two supports locating 800mm appart from each other and fixed by two grip mechanisms. Vertical load is applied to the ends of the grip mechanisms so that 4 point bending load is imposed to the test tube. The support structures can slide to absorb the axial displacement due to the bending deformation of the test tube. A vertical rod is provided on the upper surface of the test tube at its center to measure a vertical displacement at center (y). As the first test specimens, the medium boron Hastelloy- XR tubes of O.D.31.8mm x t3<5mm x L85Omm were used. Figure 19 shows the test result where a total strain range (AE) is converted to a displacement range (Ay) at the center by the following equation. 254

Ay = (6) where a and d denote a half length between the support points and a length between a support point and a loading point respectively. The cycles to failure seems to be somewhat shorter than those in reference [8][9]- In order to clarify this reason» a push-pull type fatigue test is under way for the same material.

Reference: [1] Iida,K.,et al.,11 Simplified analysis and design for elevated temperature components of Monju,"Nuclear Engineering and Design, Vol.98, No.3, pp.305-317 (January(II) 1987)

[2] Ohno.N. et al.,"Creep of Hastelloy XR in Steady and Non Steady Multiaxial Stress States at 900°C," Trans. JSME, Ser.A, Vol.53, No.491, pp.1191-1196 (July 1987) [3] Kikuchi,K. et al.,"Stress-holding time effects on creep- fracture cracking in a notched specimen,11 Engineering Fracture Mechanics, Vol.28, No.3, pp.345-367 (1987)

[£] Garofalo,F.,"Fundermentals of Creep and Creep-Rupture in Metals," The MacMillan Co., New York (1965)

[5] Rabotnov,Yu.N.,"Creep Problems in Structural Members," North-Holland Publishing Co., Amsterdam London (1969) [6] Manson,S.S. et al.,"Creep-fatigue analysis by strain range partitioning," Proc. 1st Symp. Des. Elv. Temp., San Francisco (1971)

[7] Majumdar,S. and P.S.Maiya,"A Mechanistic Model for Time- Dependent Fatigue," Journal of Engineering Materials and Technology, Vol.102, pp.159-167 (January 1980)

[8] Tsuji,H.," Effects of aging and test temperature on tensile and fatigue properties of Nickel-base heat-resistant alloy," JAERI-M 87-106 (October 1987)

[9] Kanazawa,K.,"Fatigue and uniaxial deformation behavior of Hastelloy XR at elevated temperatures," JAERI-M 87-1 H (August 1987)

Photo 1 Ligament section of the IHX manifold structural test model 258

Manifold Support

Tube Supports

Helium Gas Supply Nozzle Heat Transfer Tubes fc 1*31.8x13.5 Connecting Tubes x t3.5

Ligament Section Stubs

Manifold Lower Head Visual Inspection Nozzle

Connecting Tubes

Heat Transfer Tubes

Tube Supports

A-A Section

Fig. I I H X High Temperature Structural Test Model 259

Actuater

Load Cell Helium Gas Supply Nozzle Connecting Rods Bellows

Fig. 2 Test Apparatus for I H X High; Temperature Structural Test Model £60

U) 2 00023cr 1 = 5.4033 xi Ö e

Hastelloy XR I io2 Q- 900°C, stress constant

0 50 100 150 Stress a (MPa) Fig. 3 Primary strain parameter C| versus stress o~

Hastelloy XR 900 °C , stress constant

r = 1.2141 x103ea053cr a> E a IO o

Q. E

0 0 50 100 50 Stress

Hastelloy XR 500 900 °C Stress constant

H4 .6.0830 E'mminm= 1.9761 x 10"

tt 100 o>

10 3 2 \ff t cr io" !0

Minimum creep strain rate (l/h)

Fig.5 Minimum creep strain rote versus stress

500 o Stress constant A Load constant

-5.0801 tr» 2.4689x10" 0"

b 100 «/> Ä 50

53768 r = 3.2731xl0"a~ Hastelloy XR 900 °C, Stress constant 10 10° IO1 10Z 10'

Time to rupture tr ( h)

Fig.6 Time to rupture tr versus stress 0" 262

= 1.1088x10" a'50106

100

0 Hastelloy XR 900 °C

10° 101 102 10' Time to the onset of tertiary creep t3 h)

Fig.7 Time to the onset of tertiary creep *3 versus stress 0"

1 1 i 1 Hastelloy XR o 45.04 MPa A 58.02 MPa 900 C, D 77.9 MPa • 99.87MP3 A 119.9 MPa 10 - D D

- * D - •& 0 #D A P

l 1 p 1CP 10' 10.-1 tr-t

Fig. 8 Strain rate behaviour in tertiary creep region 263

1 1 v i 78*68 _ V A 100.0 7799

O 99.87 A 6802 O O o - o 77.89 5802 4504 1199 5795 520 4437 E a 90.15

Hastelloy XR o - 900 °C E 4 Stress increase

Numbers adjacent to data points 0 Stress constant " indicate initial stress in MPa Stress decrease

i i i 100 1000 Time to rupture (h) Fig.9 Creep damage characteristics (Time fraction

1 1 i i 1 Double necking 5795 A —~-" 520 ,

68.02 o 1 1 4QQT O 4504 99.87 5802 .2 o

Haselloy XR " 900*t A Stress increase

CO Numbers adjacen to doto points o Stress constont indicote initial stress in MR] v Stress decrease

0 i ! i 10 100 1000 Time to rupture (h)

Fig.1O Creep damage characteristics (Ductility fraction 264

70 Double necking

40 o Stress constant A Load constant a, 30

Q_ Er=-1.4644 +1.0748 logioO" 20 Hastelloy XR 10 900 °C 'Er=-1.3840+1.0907logioO"i

0 10 Initial stress (MPa) Fig,11 Rupture elongation versus initial stress

266

50 IIIITTT 1 I I I I I I I Closed Symbols i In Vacuum " Open Symbols : In Air

Haste Hoy XR o 850 °C o V Wave- Sym- A form bol E o PP O O PC o g- 10 CD O CP o o CC D T.H A

Df + Dc = 1 G.H V T&CH o

0.1 I I I I I I I 1 I 1 0.1 1.0 10 Fatigue Damage [Df 1

Fig.13 Creep damage vs. fatigue damage based on constant load stress-rupture lime 267

50 T I i FT T I I I I I I Closed Symbols : In Vacuum Open Symbols : In Air

Hastelloy XR o O 850 °C • Wave- Sym- a cP form bol PP O PC c± 1.0 V CP CC D o A° \O T.H A C.H V T&CH O

0.1 1 I I L L ) 1 I I I I I I 0.1 1.0 10 Fatigue Damage [Df

Fig. 14 Creep damage vs. fatigue damage based on constant stress stress-rupture time 268

t I I I I I I r r i i i 111 1 [7~r . Hostel loy XR _ 850 °C

Factor of 2

o Wave- Sym- o form bol •a PP O PC CP al CC D T.H A Closed Symbols : In Vacuum C.H V - Open Symbols : In Air T*CH O

i i i i i 11 I I I I I I I

Experimental Cycles to Failure

Fig.15 Predicted cycles to failure vs. experimental cycles to failure based on time faction rule using constant stress stress-rupture time 269

I I I I I I I 11 ] I I I I I I I | Hastelloy XR 850 °C

CO =3 = io3

Wave- Sym- form bol o >*» O PP O PC o CD 10' Factor of 2 CP o CC D CD T.H A QL Closed Symbols : In Vacuum C.H V Open Symbols : In Air It CM o

1 1 I Mil I I I M M I 1 f i 10 _ 10ö ' 104 Experimental Cycles to Failure [Nexp]

Fig.16 Predicted cycles to failure vs. experimental cycles to failure based on strain range partitioning method 270

T I I 1 I I I I I I I I I I I _ Haste Hoy XR - 850°C m 10 T A^/ y

•o CD

= 10'

Wave- Sym- CD form bol O o PP O ; PC •s io Factor of 2 CP CC a Q. T.H A Closed Symbols : In Vacuum C.H V Open Symbols : In Air TÄC.H O 10 i i i r i 1111 i i i ii i 111 i

Experimental Cycles to Failure [Nexp]

Fig.17 Predicted cycles to failure vs. experimental cycles to failure based on damage rate equation 271

1400 800 200 .. DISPLACEMENT METER

TUBE

CROSS HEAD

MOVABLE ARM

GIRDER

SUPPORT

Fig. 18 Tube-Bending Fatigue Machine

273

Discussion of the presentation:

The present status of research and development works for the preparation of the high temperature design code

Nickel, KFA Julien, FR6: Dr. Tanaka informed about the HTR-Testreactor conditions, in this case the He-outlet temperature is 950 °C. Your results, mostly for Hastelloy XR-R show a good stability at 850 °C. How are the creep data at higher temperatures for thii material or are you interested using other alloys e. g. INCONEL 617 for the highest temperatures?

Muto, JAERI, Japan: I have shown fatigue test results at 850 °C and creep damage test at 900 °C. However, many tests were conducted at 950 °C. The strength of Hastelloy XR is not sufficient for the long design-life at 950 °C. The HTTR will be operated at below 850 °C during the remained service time. In addition, I would like to point out that a large amount of creep strain occurs only in the first loading cycle, enabling the stresses in succeeding total creep-fatigue damage to relax. Therefore it is still possible for the total creep-fatigue damage to be below 1.0, even in the case of 20 years operation at 950 °C. The designs of HTTR components including material have already been completely determined. Therefore, it is impossible to change the candidate material for the IHX from Hastelloy XR to Inconel 617. We are planning to use a tungsten-strengthened nickel-base alloy for the commercial plant.

Schuster, KFA Julien. FR6: The SRP-damage accumulation rule is mostly applied for LCF-loadings with strain rates of about 10"4 s"1. The HTR components suffer loading rates of below 10"6 s'1. What is a typical loading rate in your experiments? 274

Muto, JAERI, Japan: The fast and slow strain rates are 5 x 10"^ s~l and 1 x 10"4 s"1 respectively. In the case of the strain-range partitioning method, we need all the experimental data of the pp, pc, cp and cc tests for the base data to predict creep-fatigue damages. Consequently, only the limited test results with hold time were available for examining the validity of this method. In addition, this method is not thought to be adequate for the case of a waveform with mean stress. Therefore, we are planning to conduct tests with more complicated waveforms, including the slower strain rate, to be able to simulate the actual loading cycles. The tests will be completed in November 1989. 275

Creep rupture characteristics in the HTGR simulated helium gas environment and their relevance to structural design

Y.Kurata, Y.Ogawa, H.Nakajima and T.Kondo Japan Atomic Energy Research Institute Department of Fuels and Materials Research

Abstract

Creep rupture characteristics in the HTGR simulated helium gas environment and their relevance to structural design are described for Hastelloy XR and Hastelloy XR-I I . versions of Hastelloy X modified for nuclear applications. The results of creep and corrosion tests in several kinds of helium environments with different iopurity compositions are presented. Corrosion data are analyzed to clarify the corrosion mechanism and estimate the long-term corrosion effect in the' HTGR heliun. Creep data obtained under refined conditions make clear the effect of decarburization, carburization and oxidation on creep behaviour. On the basis of the results obtained in this study the range of gas composition is defined where decarburization and rapid carburization cannot occur for the alloys at 950° C. A data base which allows high temperature structural design for HTGR has been obtained in a standardized HTGR simulated gas. 276

1 .Introduction Metallic materials for intermediate heat exchangers(IHXs) of high temperature gas-coo led reactors(HTGRs) which allow the use of nuclear heat for chemical process are used in high temperaturef800-950° C) coolant helium. Since the coolant helium contains small amounts of impurities, i.e.H,, H, 0, CO, CH, . CO, which react with the metallic materials, it is necessary to study the helium environmental effect on mechanical properties of the materialsfl]. As the estimate of creep damage is important to design high temperature components such as IHXs of HTGRs, the effect of helium environment on creep behaviour is studied in this paper. Hastelloy XR and Hastelloy XR-II are alloys modified for nuclear applications and are candidate materials for high temperature components of the high temperature engineering test react or(HTTR) being designed in Japan. Therefore creep and corrosion tests of the alloys in the HTGR simulated helium have been conducted. Extensive studies on corrosion mechanisms in HTGR helium have been carried out mainly for Inconel 617[2,3] . Main corrosion phenomena were carburization, decarburization and oxidation depending on the particular gas composition and temperature. Creep behaviour in HTGR helium has been estimated from the results of corrosion tests in most studies because of the experimental difficulty. In this paper both creep and corrosion tests were conducted under strictly controlled conditions to elucidate a corrosion model and the effect of corrosion on creep behaviour in HTGR helium. On the basis of the results obtained in the experiments the range of impurity gas composit ion where significant degradation of mechanical properties cannot occur is determined in this paper when the alloys are used for a long time at 950° C in HTGR helium. Long term creep data allowing the structural design of high temperature components of HTGRs have been accumulated in a standardized test gas within such range of impurity gas composition. On the other hand it is shown that some range of impurity gas composition causes 277

significant degradation of creep properties which makes difficulty in defining structural design criteria.

2.Theoretical background of corrosion Corrosion in HTGR helium depends on the kinetics of individual gas-metal reactions because of the high gas velocity. Oxygen partial pressure and carbon activity of helium environment containing small amounts of H,, Ha 0, CO, CH,, CO, are calculated from following reactions: Oxygen partial pressure 1 ,, JVo H2O = H* + — 0} /*%,= —'AT, m

CO, = CO + —0, /"'% = —— • A', (2|

Carbon activity P * P

C0 + H,= [C] + H,0 ae = — -»AT, (3)

at 2CO=(C)+C02 ac=—~-K, (4) Pco,

CH.-CO+2H, ac-—^-AT, (5)

P\2

Where PHr0 , PHf , Pca ; Pco , PCH are the partial pressures of H, 0, Ha, COj , CO, CH4 , ac carbon activity and

K, to K5 equilibrium constants of the reactions respectively. Since chromium is a main oxide and carbide forming element of high temperature alloys, it has been shown that the stability diagram for chromium is important to describe corrosion behaviour in HTGR helium[2,3]. On the basis of this idea the stability diagram for chromium (a Cr =0.8) was employed to analyze corrosion behaviour of Hastelloy XR and XR-II in this paper. Equation (5) can be ignored on account of the low reaction rate under oxygen partial pressure where chromia is' stable. Thus four locations(See A.B.C.D in Fig.l) are determined from oxygen partial pressure and carbon activity calculated from equations (1) to (4) . From equations (1) and (3) or (2) and (4) following relation is obtained p CO =O+4-0, ac= -—- • K (61 <• To, 278

Where K is ;m equilibrium constant of the reaction. From equation (6) the carbon activity of the gas depends on the partial pressure of carbon monoxide under given oxygen partial pressure. The borderline P» which divides corrosion regions is drawn on the stability diagram of Fig.l. Carbon transfer between metal and gas occurs depending on the difference in carbon activity between metal and gas or oxide-metal interface. It was shown that corrosion does not proceed in the way of model(I) and model(IV) of Fig.l[2-4]. Fig.2 shows carbon activity and oxygen partial pressure of helium environments whose impurity composition is shown in Table 1. The results of corrosion tests in He-3 and He-5 were compared to make clear which model of (II) or (III) in Fig.l was suited and which reactions from (1) to (4) controlled oxidation and carburization or decarburization . Since decarburization occurred in both He-3 and He-5 as shown in Fig.3, it was shown that model (II) of Fig.l is suited[4]. Moreover we can conclude that equations (1) and (3) control the reactions because oxygen pick up was almost equal for both He-3 and He-5 as shown in Fig.3. Decarburization will occur in the regions where partial pressure of carbon monoxide is

lower than Pco and chroraia is stable or where chromium is stable. It is considered that the reaction of methane decomposition defines the carbon activity of helium environment under low oxygen partial pressure where chromium is not oxidized.

3.Experimental procedure 3.1 Creep tests in JAERI Type B helium The material used in this experiment was Hastelloy XR whose corrosion resistance in HTGR helium environment was improved[5] . JAERI Type B helium which was mainly determined from Huddle's prediction[6] was employed as the primary coolant simulated helium of HTGRs. The impurity composition is shown in Table 1. Oxygen partial pressure and carbon activity calculated from equations (1) and (3) are also shown in Fig.4. Single-type creep testing machines in helium|7] were used which were specially made to examine the helium environmental effect.. Impurity controlled helium 279 was supplied in lJt/min per one specimen from the helium circulation loop which had purification and impuruty addition functions. Creep tests in JAERI Type B helium were carried out at 800 to 1000°C up to about 10, OOOhrs [8 ] . Parallel to the tests in the helium creep tests in air were also conducted for the same material[9].

3.2 Creep and corrosion tests in , decarburizing helium environments Materials used in the experiment were Hastelloy XR and XR-11,versions of Hastelloy X modified for nuclear applications. The chemical composition is shown in Table 2. Improvement of creep strength for Hastelloy XR-II was made by boron addition[10]. He-1 to He-4 in Table 1 were used in the experiment[4,11]. Oxygen partial pressure and carbon activity calculated from equations (1) and (3) or equations (1) and (5)(for He-4) are shown in Fig.4. The same type of creep testing machines in -helium as described in 3.1 were used in the experiment. Corrosion tests together with creep tests were conducted by suspending corrosion coupons near the creep specimen. The test temperature was 950° C, and initial applied stresses 31MPa and 26MPa.

3.3 Creep and corrosion tests in carburizing helium envi ronments Materials and experimental equipment were the same as described in 3.2. Not only creep tests but also corrosion tests were conducted by means of the creep testing machine in helium. Impurity composition in the helium environment were chosen to obtain higher carbon activity[12]. Table 3 shows impurity composition and Fig.5 carbon activity and oxygen partial pressure of each helium environment. The test temperature was 950° C, and initial applied stresses 31MPa and 26MPa.

4.Results and discussion 4.1 Creep behaviour in -JAERI Type B heJium * Fig.6 shows creep rupture times in JAERI Type B helium compared with those in air. Although the scatter band is wide because the data were obtained from all the product 280 forms of bar, tube and plate, it seems that there is no difference in rupture times between JAERI Type B helium and air. Data for the tube material in JAERI Type B helium and air are shown in Fig.7. The shaded regions show 95* prediction intervals for rupture time data in air. Although the prediction intervals are narrower than those of all forms of the products, rupture times in the helium are within the 95% prediction intervals. Impurity composition of JAERI Type B helium is in the region where oxide film of chromia forms and carburization occurrs slightly as indicated in Fig.4. Since the reaction rate of corrosion is low in the region, it is concluded that the effect of corrosion on creep strength is small after long service. The fact that there is no difference in rupture time data between JAERI Type B helium and air is caused by the fact that the impurity composition of JAERI Type B helium is in the protective corrosion region and that corrosion resistance of Hastelloy XR is improved. On the basis of the results obtained in the study it becomes possible to use creep data in air as well as those in the helium in order to design high temperature components for HTGR in helium environment such as JAERI Type B helium.

4.2 Creep and corrosion behaviour in decarburizing he 1iuD envirönnents The relation between stress and rupture time in He-1 to He-4 is shown in Fig.8. Although rupture times are almost equal in He-1 and He-2, rupture times are significantly shortened in He-3 and those in He-4 are halfway aiong them. There is little difference in not only rupture tine but also creep deformation behaviour between He-1 and He-2 as indicated in Fig.9. Accelerating creep occurs at the early stage and leads to failure in He-3. Fig.10 shows the relation between weight gain and exposure time in He-1 to He-4. Although the weight of the specimen increases in He-1 and He-2. weight change is not observed in He-3 and He-4. On the other hand rapid decarburization occurs in He-3 and slight carburization in He-1, He-2 and He-4 as shown in Fig.11. Creep rupture times are shortened in He-3 where decarburization occurs as shown in Fig.12 281 which represents the rupture time in each helium environment on the stability diagram for the alloys. Since decarburization causes a significant decrease in creep strength, attention should be paid to it from the viewpoint of structural design and plant operation[4,11]. The, value of P' which determines the protective corrosion region is about lOnatm(lPa) for Hastelloy XR and XR-II. This value is lower than 30uatra(3Pa) for Inconel 617.

4.3 Creep and corrosion behaviour in carburizing helium environments The relation between stress and rupture time in carburizing heliu« environments is shown in Fig.13. For Hastelloy XR rupture times in He-1 and He-2 are almost equal and those in He-6 and He-7 are longer. For Hastelloy XR-II rupture times in He-l.He-2 and He-6 are almost equal and those in He-7 are somewhat longer. As shown in typical creep curves of Hastelloy XR of Fig.14, creep rate is still low in He-6 and He-7 at the stage where the accelerating creep occurs in He-1 and He-2. The relationships between weight gain and change in carbon content with exposure time are shown in Figs.15 and 16 respectively. While weight gain is small and little protective oxide film forms in He-6, increase in carbon content is large. Although slight carburization was observed for corrosion specimens in He-7, it was considered that carburization would be increased for creep specimens in He-7 because of surface cracking. Therefore the fact that rupture times are long in He-7 iay be related to the crack of surface oxide film and carburization occurring continuously. Creep rupture tiies in carburizing helium environments are shown on the stability diagra« for the alloys in Fig.17. In the region where chromia was stable and partial pressure of carbon monoxide beyond Pi0 creep rupture times are almost equal or somewhat longer. He-4 in Fig.17 locates on the line of log a.=0, calculated from equation (5). However carbon increase was somewhat smaller in He-4 than that in He-6. Therefore the reaction rate of equation (5) is slow and effective carbon activity of He-4 is considered to be lower than the value calculated from equation (5). There may be some cases 282 where creep rupture tines are somewhat shortened depending on impurity composition in the region of chromium carbide as shown in Fig.17 even if decarburization does not occur. Although the effect of helium environment on creep rupture strength is complicated, the effect is considered to be composed of several corrosion factors as shown in Table 4 [ 12] . Since the formation of a protective oxide film is difficult and rapid carburization is apt to occur in the region of chromium carbide, it is desirable to avoid such a condition at the elevated temperatures. 5.Conclusion Creep rupture characteristics of Hastelloy XR and Hastelloy XR-II in HTGR simulated helium gas environment have been described. The reactions which control carbon activity and oxygen partial pressure of HTGR helium environment were defined through analysis of corrosion mechanism in HTGR helium. From the results of long term creep tests in air and in JAER1 Type B helium, a standardized primary coolant gas of HTGRs, there was no difference in creep rupture data between both environments. In order to design IHXs it would be possible to use data in air as well as those in the helium environment under the condition where protective corrosion behaviour is observed. On the basis of the results of creep and corrosion tests in helium environments with different impurity compositions, it was shown that decarburization causes significant degradation of creep properties and the boundary condition of decarburization region was defined. Partial pressure of carbon monoxide was the most influential in preventing decarburization for Hastelloy XR and XR-II at 950°C. The embritt]ement due to carburization was anticipitated because carburization became significant under conditions where a protective chromia film did not form and carbon activity was high. In this region there were some cases where creep properties were deteriorized. From the viewpoint of long term stability of material properties, additional work will be needed in order to use the alloys in this region. It is desirable to avoid such a condition that makes difficulty in defining structural design

criteria. 283

References

1)H.Nickel,T.Kondo and .P.L.Rittenhouse:Nuclear Technology 66(1984)12 2 )W.J.Quadakkers and H.Schuster:Nuclear Technology 66(1984) 404 3)W.J.Quadakkers:Werkstoffe und Korrosion 36(1985)335

4)Y.Kurata,Y.Ogawa and H.Nakajiraa:JAERI Research Report JAERI-M 88-176(1988) 5)M.Shindo and T.Kondo:Proc. Conf. on Gas-Cooled Reactors Today, Bristol/ÜK,1982(British Nuclear Energy Society) Vol .2,p.179 6)R.A.U.Huddle:Proc. Conf. on High Temperature Reactor and Process Applicat ions,November 26-28,1974,paper No.40, British Nuclear Energy Society 7)Y.Ogawa and T.Kondo:JAERI Research Report JAERI« 8801 (1980) 8)Y. Kurata ,Y. Ogawa and T . Kondo": Nuclear Technology 66(1984) 250 9)S.Yokoi.Y.Monma.T.Kondo,Y.Ogawa and Y.Kurata:JAERI Research Report JAERI -M 83138(1983) 10)Y.Kurata,K.Sato,T.Nakanishi,K.Sahira and T.Kondo:Proc. Int. Conf. on Creep (1986)p.97 11)Y.Kurata,Y.Ogawa and H.Nakajima:Tetsu to Hagane 74(1988) 97 12)Y.Kurata,Y.Ogawa and H.Nakajima:Tetsu to Hagane 74(1988) 2185 •284

Table 1 Impurity composition in helium environments

pato(Pa)

CO co H2O H2 2

He-1* 200(20) 100(10) 2(0.2) 5(0.5) 1(0.1)

He-2 50(5) 25(2.5) 2(0.2) 5(0.5) 1 (0.1)

He-3 500(50) 3(0.3) - 5(0.5) 1(0.1)

He-4 500(50) 3(0.3) - 5(0.5) 0.05(0.005)

He-5 440(44) 3(0.3) 0.5(0.005) G(0.6) 1(0.1)

kJAERI Type B helium

Table 2 Chemical composition of specimens(wt*)

Mn Cr Hutelloy XR 0.07 0.B3 0.33 <0.-»5 <0.'»5 21-W <0.0$ 9.10 0.48 IS.II <0.05 «J.05 P.0CW3S Hastelloy XR- 0 0.07 0.87 0.27 <0.«5 0.001 21-96 0.12 9.Z-10.46 18.33 0.03

Table 3 Impurity composition in carburizing helium environments

Hatm (Pa)

H2 CO C02 CH« H2O He-1* 200(20) 100(10) 2(0.2) 5(0.5) 1(0.1) He-2 50(5) 25(2.5) 2(0.2) 5(0.5) 1(0.I) He-6 SOOt50) 100110) 0.05(0.005) He-7 500150) 400UO) 10(1) :»(0.5) 2(0.2) "JAERI Type B helium

Table 4 Effect of environmental factors on creep rupture strength of Hastelloy XR and XR-JI in helium envi ronment

——-——^______Cjrburiud'on D«irfcortzalion D**r»da(iitn uniirr »•!•>• low oivgirn parti'ii pressure Strengthening 4 4 Hasldloy Xlt Weak«,,«« +

Strengthening f HastelloyXR-II Weakening + 285

o O C7» O

log P02 III Metol Oxide Gas (II) Metal Oxide Gas .

In ill P CM • 0l c Ü_ B CP o 1

# f 1 o 9 o "tl\v In«» B

lo g N C

(E) Metal Oxide Gas (V)Meraf Oxide Gas ^-- •—

2) CM FW A o — Q_ 0 / o» o

lo g /

A' o o o a o» Oc a \ o A o \ 0 N

Fig.l Schematic diagram of corrosion.mechanism showing the relationship between decarburization or carburization and oxidation in impure heJium.

Po(;Po calculated from equation Cl). <>i P0];Po, calculated from equation (2), (» a e ; a t calculated from equation (3) , (« ae ;a calculated from equation (4) 286 0 950°C

Cr2o3 -1 h

-2 P^\ /

-3 B > •<> Ht-I o> —4 o \ Hi-3 t \ +- ^t "Z -5 c Chromium

-6 Ht-5

-26 -24" -22 -20 -18 -16 -14

log Po2

Fig.2 Carbon activity and oxygen partial pressure of test

helium in the stability diagram for chromiua(a Cr =0.8) at 950°C. Carbon activity and oxygen partial pressure are calculated in the way shown in Fig.l except for He-4.

E-02 • o I I •e -OA --

0.2 -- Q. 3

c 5 0.1 -

He-3 He-15 He-I3 He-5

Hastellcy XR Hastelloy XR-1I

Fig.3 Results of corrosion tests in He-3 and He-5. Test ;4 86.6h(He-3).439.7h(He-5). 287

Fig.4 Carbon activity and oxygen partial pressure of different helium environnents io the stability diagram

for Cr(a Cr =0.8) at 950°C

-7 -26 -22 -20 -13 -14

log P02

Fig.5 Carbon activity and oxygen.partjaJ pressure of carburizing helium environments in the stability c diagram for Cr(aCr=0.8) at 950 C 288 14 8O0*C 10 !0G 8 6 9O0*C 5 50 4 1000'C 3

2 20 £ o Bor \ A Tube >n helium o Plate/ 10 '//// in air

1O1 102 103 to4 Time to rupture (h)

Fig.6 Comparison of creep rupture data in JAERI Type B helium and in air for Hastelloy XR sampled from bar, tube and plate materials. The shaded regions indicate 95* prediction intervals obtained in air for the same set of Hastelloy XR.

to 100 8 — 6 e 5 i* 1000°C o ^ 3 20 £ 2 A in helium 1 XU in air 10 as

10s 102 103 104 Time to rupture ( h }

Fig.7 Comparison of creep rupture data in JAERI Type B helium and in air for Hastelloy XR sampled from tubes The shaded regions indicate 95% prediction intervals obtained in air for the same set of Hastelloy XR. 289

40 50 Hastelloy XR Hastelloy XR-II 35 • 950°C 950°C 40 26MPa 30 He-1 X He-3 He-4 o He-1 30 • He-2 / 25 He 2 He-2 * x / ^-3He-4 - o V He-3 f Q. I > ' • ,', a He-4 Tz 20 HasteUoy XR-II 35 -- 950°C 10 • in 30 - • i • sz a. ^. 25 • 0 200 400 600 800 1000 120O Time (h) 20 Fig.9 5 10£ 2 5 103 2 ! Comparison of creep curves for Time to rupture (h) Hastelloy XR-II in different Fig.8 helium environments Comparison of creep rupture times in different helium environcents

1.2 0.06 HasteUoy XR HasteUoy XR 950°C He-1 0.8 - 950Oe ^nrnUPi ™ 0.03 0 0.4 - \ -0.03 He-i O H«-1 He-2 0 i a -V-. •-^:^rvV.Q^---- - He-4 A He-2 -0.06 He-3 He-3 He-3 c He-4 o D He-4 Hastelloy XR- He-1 Hastelloy XR-l 0.03 950°C 95O°C «0.8 - 0

-0.03 0.4 •

-0.06 He-3 0 <----,Q-„ 0 200 400 600 800 1000- 1200 140C -0.4 • Exposure time (h) 0 200 400 600 800 100C I2O0 1400 Exposure time (h) Fig.10 Fig- 11 Relation between weight gain • Change in carbon content of and exposure time in different Hastelloy XR and XR-II exposed helium enviroments to different helium, environments 290

~ l500r 950 *C Hastelloy XR- II .§ 1000 fp I Hastelloy XR 500 a. or

-26 -22 -20 -18 -16 -14

log Po2

Fig.12 Comparison of creep rupture times of Hastelloy XR and XR-II under 26MPa in different helium environments

indicated in the stability diagram for Cr(aCr=0.8) at 950eC

40 50 HasTelloy XR HasteUoy XR 35 • 95O°C 95O°C 30 • 26MPa

au7. o 25 \ \^- He-6 He-1 He-7 He-2 He-2 we-6 D H--7 HasteUoy XR-I! 35 - 950°C \ \ 3C - CCG. He-7 25 - He-1 200 400 600 aoo 1000 He-2 He-6 Trme (h) 20 i iii 4 102 5 10s 2 10 Time to rupture (ri) Fig. 14 Fig.13 Comparison of creep curves of Comparison of creep rupture times Hastelloy XR in carburizing in carburizing helium environments heii um envi ronments 291 0.06 \z Hostelloy XR 950°C He-i 0.8

0.4 *E ^ 0 ^E _c o o» Hosteltoy XR-ll 950°C I» 0.8 jj—

-<7—V H#-6 0.4

0 0 500 SCO 10O0 "200 1400 -0.4 Exprsure tire (h) 0 200 400 600 800 1000 1200 1400 Exposure time^ (fi) . F i g . 1 5 Fig.16

Relation between weight gain Chaage in carbon content of and exposure time in carburizing Hastelloy XR and XR-II exposed helium environments to carbarizing helium environaents

1500 r 950°C I 1000h fllHasteltoy Xfi-II {JpHastetloy XR a> 500 h a. 0 L

-26 -24 -22 -20 -18 -1A log Po2 Fig.17 Comparison of creep rupture times of Hastelloy XR and XR-II under 26MPa in carburizing heliuo envi roment s

indicated in the stability diagram for Cr(aCr=0.3) at 950*C 292

Discussion of the presentation:

Creep rupture characteristics in the HTGR simulated helium gas environment and their relevance to structural design

Schad, Lurgi, FRG: What did your specimen look like and how deep into the material did you notice composition changes?

Kurata, JAERI, Japan: The size of our creep specimen was 6 mm in diameter and 30 mm in gauge length. The corrosion specimen was a rectangular type of 10 mm in length, 5 mm in width and 2 mm in thickness. The depth of the decarburized region was estimated to be 400 - 500 urn for the corrosion specimen exposed to He-3 for about 200 h.

Schuster, KFA Julien, FRG: How did you find or determine the Peg-line in your stability diagrams?

Kurata, JAERI, Japan: We determined PQQ for Hastelloy XR and XR-II using 0.8 as the value of

acr. This value for Hastelloy X was obtained from the following reference: W.J. Quadakkers: Werkstoffe und Korrosion 36(1985)335 293

Assessment of Primary and Secondary Stresses for Component Design

E. Bodmann Hochtemperatur-Reaktorbau GmbH, Mannheim

Contribution to • . Workshop on Structural Design Criteria for HTR 31. January - 1. February 1989, Julien, KFA 294

Assessment of Primary and Secondary Stresses for Component Design

1. The background of the definition of primary and secondary stress

Primary and secondary stresses are determined by a calculation process. Practically they can not be measured at any point of a structure. Their initial definition - e. g. in the ASME-Code, section III - is based on the assumption of a ductile, elasto-plastic material behaviour as an imperative requirement. The 1 imitation of the primary and secondary stresses in a component is one of the most important elements of design, because it is a means against a number of component failure modes due to static and cyclic loads.

Therefore the following question is not trivial: Does the definition of the primary and secondary stress also hold, if creep is a determinant or a additional effect of the inelastic material behaviour?

We studied three design methods which use the ductility to see, if methods have to be changed if creep becomes significant:

the assessment of primary bending stresses using a section factor

the general method of the dimensioning of load carrying areas by limiting the primary stress

the assessment of primary and secondary stresses in the case of elastic follow-up.

In the following the results of these investigations are di scussed. 295

2. The section factor for the assessment of primary bending stresses

Due to inelastic strains and a subsequent stress-redistibu- tion a section area has a larger load carrying capability than suggested by elastic stress calculation. This

enlargement of capability is expressed by a factor fs. In an assessment of combined primary membrane and bending stresses the latter are to be reduced by the section factor before comparison with the allowable primary stress.

We assume for the determination of the section factor that the equation

£. = A • 6 n

for the stress-strain-relation at a certain time after loading is reasonably accurate both for plasticity and for creep. (Fig. 1).

Obviously this.approach also holds if the rate of creep strain is given by the Norton's law . .

I = B •

We assume as usual that sections remain plain after bending deformation :

— = const (x) flr* and get a stress distribution

as shown in Fig. 2. By integration of this stress 296

distribution for the different types of section areas (e.g. rectangle, circular tube) we get bending moments. The ratio of the bending moment to elastic calculated bending moment is the section factor.

It is remarkable that the section factor is only a function of the stress exponent n (Fig. 3). In the extreme case n —* * we get the well known solution for ideally plastic behaviour. In the case n = 1 there is no enlargement of capability, the solution for the stress distribution holds as well for an elastic as for a linear-visco-elastic behaviour. Section factors for 1 < n < oo are independent of the type of inelastic strain as plasticity or creep.

3. Dimensioning of load carrying section areas

The following considerations are useful for thick walled tubes or nozzles under internal pressure as examples for components with a complex stress distribution. Usually the load carrying section areas of such components are . dimensioned to cope with an allowable primary stress. Stesses and strains locally rise above the mean (primary) stress and related mean strain. Local inelastic strains are tolerated to a 1imited extent.

We believe that a "3-Beam-Model with Different Flexibilities" Fig. 4 is a good tool to study the inelastic strains due to plasticity and creep in such a cross section, or and X are the ratios of the two section areas and the two beam lengths respectivly. G" is the primary stress,fc, is the ratio of the strain in the higher loaded beam "1" to the mean strain £, produced by the mean (primary) stress 0" .

We analysed this model using

a) Norton's creep law c = B • Cn 297 b) two types of elasto-plastic behaviour.

Fig. 5 shows the mode Is properties for the stress exponent n = 1, this is also the elastic behaviour. The strain ratio is presented as a function of the two geometric parameters * and A . It is not so important to look for the absolute numbers in the figure, but to compare the figures for the different material models.

Fig. 6 shows the models behaviour for a realistic creep stress exponent n = 4, Fig. 7 the same for an unrealistic high exponent n = 25. There is nearly no difference between these two figures {6 and 7). The analysis of the 3-Beam-Model1 using a specific-elasto-plastic stress-strain-law

G" *•" £ - -^ + K-

We concluded from these results the following:

- The rules for dimension ing of load -carrying sections and for the geometrical designing of components developed mainly for the lower temperature range are valid and useful also in the creep range.

The elastic analysis is a useful method to optimise the geometry of. the component with respect to minimum local strains in the creep range also. 298

4. Elastic fol1ow-up

We look in this context into complex structures, mainly piping systems. When designed for the high temperature service they are usually loaded by the restraint of thermal expansion or by support movements. These restraints act to a certain amount as primary loads, as it can be demonstrated by Fig.10. In the case of the combination of primary and secondary

stresses generally a fictitions stress G*fe above the yield stress is permissible (point A). This stress Cfe is a result of an elastic analysis of the structure.

In the case of a rigid restraint of the movements the real stress-strain situation is that of point B. In the case that

Cfe is definitivly primary, the structure fails by large strains and deformations (D). In realistic cases the highly loaded parts of the structure undergo stresses and strains above that of point B, i. e. point C. The angle y of the line A-C is a measure of the "elastic follow-up" of the structure. If the angle

The line k in Fig. 12 is assumed to be the stress strain curve of the material from a stress-rupture test. We use in our thinking a fictitious material model with the following property: Initially after loading it has the ability to

undergo the stress C fe at point A. After that time we switch down the fictitions yield stress. When reaching the real yield stress the structure is at poit B^. In the case of creep at high temperatures now the relaxation process "switches" the yield stress down; we get a certain type of isochronous stress strain curves (1, m) by relaxation from 299 the curve k. In this diagram the "elastic follow-up" process can be visualized. As can be seen, there is no difference in principle between elastic follow-up caused by plasticity at low temperatures and by (additional) creep at high temperatu re.

Now we have a tool to quantify elastic follow-up without too much of an effort. The steps of a related elasto-plastic analysis are listed in Table 1.

Fig.13 shows in principle the step from the results of an inelastic analysis with increasing load and constant yield stress to the elastic follow-up process with decreasing yield stress and constant load.

In Fig. -14 results of the related analysis of a real pipe structure are presented.

Summary

We investigated three design methods which are based on the assumption of a sufficiently ductile material. It was possible to show in all three cases, that creep as a predominant or additional material effet does not influence the validity of these methods.

Moreover it has been shown that plasticity and creep can be treated uniformly. Creep can be considered as time dependent plasticity with respect to the macroscopic effect of inelastic strain. The time dependence however does not affect these design methods. 300

Table 1

Steps of an Inelastic (Elastoplastic) Analysis of a Structure to Quantify Elastic Follow-Up

1. Modeling of the Structure for an Elasto-Plastic Finite-Element Analysis

2. Elastic Calculation using a Unit Load; Searching the Position of Maximum Stress S in the Structure

3. Set the Fictitions Yield Stress equal to 6

4. Elasto-Plastic Analysis of the Structure by Stepwise Increasing Load greater than Unit Load from Step 2.

5. Identify the Over-Proportional Strain versus Load at the Point of Maximum Stress (see Fig. 13)

6. Drawing the Elastic Follow-Up Diagramm using the Result of Step 5 (see Fig. 13) 301 > • '/ a M — - -G £ fri 1 c - A •

B-G

Bending Stress Distribution for Nonlinear Stress-Strain-Law snd Basic Formulas for Section Factor Determination Fiq. -1

1 / 7

X a Q5 1

Stress Distribution in a Rectangular Section due to a Bending Motment as a Function of the Stress-Exponent n - 302

20 n

Section Factor F^ as a Function of the Stress Exponent n a) rectangular section b) circular tube section

Tig. 3

e. 5—£

3 - Beam - rodel with different Stiffnesses

FiQ. 303

s k, STRfllN-RRTIO flS fl FUNKTION OF cC flNO -X

CREEP":STRESS EXPONENT M=l =10

0-5

dL-10 0.2

0.1

0.20 0-30 0-4Q 0-50 0.60 0.90 1-00 X Fig. 5

Tk, STRfllN-RRTIO RS R FUNKTION

CREEP :STRESS EXPONENT N=4

0.10 0-20 0.30 0-40 0.50 O.GO .0.70 0.90 1-00 X Fig. 6 304

a h STRRIN-RRTIO RS fl FUNKT ION o o . a OF c< RND X

CREEP^STRESS EXPONENT N=25 «C =io 1.0 0 5 o o , i- 2

o o . 1

0 .5 e o . to 0 • 2 a *-

36 "»

o ^^

0 1. ( 1 ( 1 1— 1 —i —t . 0-10 0.20 ' 0-30 0.40 0.S0 0-10 0-90

Fig. 7

'0.00 0.30 0-40 0.61 a. Fig. 305

STRRIN-RflTIO flS fl FUNKTION OF

ELflSTIC-lOEfllLY PLflSTIC BEHOVIOR eC -

0.5

0.2

0.1

°0.00 0.10 0.20 0.3Q O.JO 0-S3 Q.60 0.70 Q.EO 0-90 1-0

Fig.9

/ f1

Elastic FollOW-Up

Fia 10 306

A j:

? .^ r

1 £~

Definition of Allowable Primary plus Secondary Stress 6^( In Case of Elastic Follow-Up

Fig. 11

Combination of Fictitious Yield Stress and isochronous Stress - Strain - Curves

Fig.12 307

K - TT #, -

Strain at Point of Concentration versus Load is a Type of Elastic Follow-Up Diagram

Fig.13

2 Y *

Elastic Follow-up Calculated for a Real Pipe Structure

Fig. 308

Discussion of the presentation:

Assessment of primary and secondary stresses for component design

Alder, PSI Wilreniinqen, Switzerland: What is the value of the exponent in Norton's law that you use most frequently? Is it based on experimental values only or do you add some margin?

Bodmann, HRB Mannheim, FR6: The realistic range of the stress exponent is 4 to 8, the lower end of the range for the high service temperatures (IN 617, Alloy 800), the higher end for the lower temperatures (550 - 700 °C). The values are derived from a lot of creep rupture tests of the HTR-materials. For the section factors we used n = 4, but they can be taken for a specific value of n depending on material and service temperature. 309

Elastic and inelastic analysis of component behaviour

H.-J. Seehafer

Interatom GmbH, Bergisch Gladbach, FRG

Abstract

Analyses of component behaviour in the high temperature field have to be orientated on the component loadings and on the corresponding failure modes. As far as time dependent material behaviour is relevant, inelastic considerations have to be performed. As complete inelastic analyses are however really expensive, a stepwise procedure by applying elastic, simplified inelastic and detailed inelastic analyses is generally suitable.

Criteria for defining the time and temperature dependent area of creep relevance are discussed by presenting corresponding creep- cross-over curves for three reference materials ensuring component design against creep -failure without performing any inelastic analysis. The application of these creep-cross-over curves is restricted to normal loading conditions and does not consider. . possible fatigue failures where hold time effects may become significant .

If fatigue failures become relevant, corresponding use fraction ratios can be evaluated on basis of elastic stress analyses by applying stress intensity factors in order to consider an increase of the strain range due to plastic effects. Stress intensity factors for selected conditions and for particular material behaviour have been determined, thus enabling the performance of simplified elastic-plastic analyses. Compared to ASME recommendations, significant differences depending on stress.level have been identified. The practical importance of this procedure is related to assessments of peak stresses due to notches and geometrical discontinuities. 310

Concerning creep-fatigue interaction the separate evaluation of both creep and fatigue use fraction and the combined assessment of these use fractions seams to be a suitable procedure. Both the development of adequate constitutive equations for describing the creep damage and the development of suitable damage rules are of basic importance. The procedure and the influence of relevant parameters describing the material behaviour are discussed on the basis of an inelastic analysis of a structural part of the Modular HTR steam generator.

If the combination of primary stresses and cyclic secondary stresses occur, the inelastic strain will cyclically increase up to an allowable limit. Parametric inelastic studies have been started with regard to the development of simplified methods and envelope requirements for avoiding ratcheting or creep ratcheting failure. These studies should be extended for different and more general applicability.

Summarizing the results, one can conclude that for components of the electrical power generation plant, operating up to temperatures of about 600 °C, comprehensive inelastic analyses will not become necessary. All loads and load cases can be considered most effective by a stepwise procedure applying elastic and simplified inelastic methods. Nevertheless the basis for performing inelastic analysis is largely available, the improvement of the data basis of constitutive equations however is desirable. 311

Introduction

Inelastic material behavior becomes more relevant with increasing temperatures. The inelastic material behavior can be considered in the component design by means of elastic as well as of inelastic methods. The elastic procedures can be performed fast and inexpensively, needing however raised safety factors for ensuring adequate conservatism of the assessment of the results. On the other hand side inelastic methods are very time and money consuming. These procedures are characterized by a realization of the envisaged utilization of the inelastic, material behavior by means of enhanced expenses for performing corresponding analyses. The grade of conservatism can be reasonably reduced compared to elastic procedures while the necessary expense must be enhanced by a factor of 5 or 10 up to 20 for particular conditions.

The basic question which arises by discussing elastic as well as inelastic design procedures is therefore:

Are there reasonable - meaning economical - methods enabling the consideration of inelastic material behavior without being forced to perform complete and very expensive detailed inelastic analysis?

Discussing this question basically, the corresponding situation for the most relevant failure modes will be examined closely. In accordance with the survey of all

possible failure modes in figure 1# the most relevant failure modes in high temperature loaded structures are:

- creep - fatigue - creep ratcheting

The aspect of simplified procedure for these three characteristic failure modes will be discussed in detail as follows. 312

2. Creep-cross-over curves

With regard to the attempt of saving money by reducing the expense one can state that the best way would be to renounce performing any inelastic analysis. Envisageing this, the decisive first step is to identify creep relevant conditions of corresponding structures.

Keeping in mind that creep relevant conditions are characterized by two parameters:

- time and - temperature

a set of suitable ciriteria enabling the definition of creep relevance could be developed for the materials

X20 CrMo V 12 1 X10 NiCr AITi 3220 NiCr 22 Co 12 Mo

These criteria, summarized in figure 2, are aimed at and based on well known high-temperature criteria for limiting:

- primary membrane stress (S = S.) m t

- local and bending primary stresses for level B and C conditions for avoiding creep ratcheting corresponding to Test Nr. A-3 of /I/

relaxation to 33% based on yield stress in order to avoid creep ratcheting

inelastic total strain to 1% and creep strain to 50% of plastic strain 313

Each of these criteria can be expressed by a, corresponding time-temperature correlation as given exemplaryly for the material X10 NiCr Al Ti 3220 in figure 3. This figure indicates more or less stringent restrictions by each of the criteria. The recommendation of the creep-cross-over curve as a lower bound is characterized by the points

t = 1 hour and 650 °C t = 3 - 105 hours and 500 °C

Divergent to figure 3 there is no further reduction of the creep-cross-over curve below 500 °C.

Following the procedure described by the above criteria, corresponding creep-cross-over curves as given in figure 4 for the three HTR-materials have been developed on.basis of the available material data.

The rather strong slope of the. material NiCr 22 Co 12 Mo is based on conservative assumptions due to lack of the data base in the low temperature area.

Below these curves creep effects are nearly negligible, enabling the selection of structural parts where inelastic analysis need not be performed. These creep-cross-over curves are applicable for normal and upset conditions but not for emergency conditions and creep-fatigue interaction failure modes. Identifying creep relevant conditions by creep-cross-over curves the number, of inelastic analyses can be reduced significantly.

The second step after having selected structures exposed to inelastic conditions is to look for a simplified procedure in order to avoid any failure mode that may become possible. Keeping in mind that above all, peak stresses are the reason for crack initiations leading at least to possible failures, the following discussion will be concentrated on fatigue failure modes in connection with peak stresses. 314

3. Strain concentration factors

If strain controlled loadings are evaluated on basis of elastic analyses, the fictitious stress-strain-state does not represent the real situation above all under high temperature conditions. This real situation is characterized - as indicated in figure 5 - by a larger strain range than elastically evaluated. The amount of strain range increase due to plasticity related to the elastical fictitious strain range is defined by the plastic strain concentration factor K . This factor is well known for notch effects /2/ but is e also defined in the ASME Code /3/, for peak stress assessments on elastic basis. Comprehensive investigations at Interatom /4/ have shown that divergent to /3/, significant effects should be considered by evaluating plastic K -faktors as indicated in figure 6. Main difference comparing the proposed procedure in /4/ to ASME is that it is based on peak stresses and not on primary and secondary membrane and bending stresses. As the K -factors are applied to peak stresses, this procedure is more reasonable and not unconservative for notches as it is, following the ASME procedure. As pointed out in figure 6, the basic effects which have to be taken into account are:

- geometry - notch configuration - multiaxial load cases - material hardening

As far as geometry effects are concerned, the most severe situation leading to peak stress concentrations is characterized by strong material discontinuities as it is given by cantilevers. Inelastic analyses have therefore been performed for cantilevers of the material Incoloy 800 H at 750 °C enabling the evaluation of strain concentration 315

factors K1 depending on peak stresses. On basis of the same material behavior, which is characterized by the elastic/plastic transition stress and by the plastic slope describing the material hardening, inelastic analysis have been performed for evaluating corresponding K„-factors applicable for notch conditions.

Finally multiaxial effects, arising for instance due to plastic strain concentrations in connection with different Poisson's ratios of elastic and elasto-plastic behaviour, have been investigated comprehensively leading to corresponding strain concentration factors K, .

The set of the three K -factors is well representing the strain concentration effects

- geometry (= K,) - notches (= K ) - multiaxiality (=K^)

The results of an inelastic parameter study on basis of the material Incoloy 800 H at 750 °C are summarized in figure 7 and compared to the corresponding results of ASME procedure. In order to get the basis for comparing ASME and Interatom procedure the peak stresses have been assumed to be equal to primary and secondary membrane and bending stresses. On basis of the results in figure 7 one can draw the following conclusions

- The envelope of the set of the K -curves gives the recommended K -factor depending on stress level

- For lower stress levels ASME K -faktors are not conservative e - For higher stress levels, ASME procedure is extreme over- sonservative 316

- The geometry effect, represented by the cantilever - is the most severe condition for high stress levels compared to multiaxiality and notch effects

Figure 8 shows as result of analogeous investigations the envelope of K -factors for the material AISI 316 for three temperature levels. Compared to ASME results, the above mentioned conclusions could be confirmed. Surprising, but explained by the particular hardening effect of the material, the K -factors evaluated for a temperature level of 400 °C are higher than they are for temperatures above 550 °C. As pointed out above, material hardening and geometry effects are of basic importance for evaluating K - factors. Additional investigations have been performed, trying to determine the potential for reducing conservatively evaluated K -factors by taking into account more realistic geometry and hardening effects. Figure 9 shows the results of corresponding inelastic analysis related to the stress ratio £", , / 6^ • Two aspects are of particular interest:

1. The cantilever, assumed as basis geometry, has been confirmed as very conservative condition compared to the real case of an elbow

2. Hardening effect of the material has an important influence on the strain concentration factors comparing full drawn curves with corresponding dotted curves.

In order to demonstrate the effect of strain concentration factors by a practical example, the structural transition of the heat exchanger tube to the main steam header has been investigated (see figure 10). The temperature transient which has been applied to this structure has been increased by a factor of about 10 related to real plant conditions in order to impose loading conditions causing relevant inelastic material behaviour. 317

The corresponding fatigue assessment, which has been based on the recommended design fatigue curve of the material Incoloy 800 shows significant differences of the allowable cycles dependent on the applied procedure. Summarizing the results of the investigations:

- ASME procedure is not suitable for predicting realistically strain enhancements on basis of elastic analysis by applying strain concentration factors

- The hereabove proposed procedure considers significant effects as geometry, material hardening, notches, peak stresses and multiaxiallity

- For practical application, K -factors should be evaluated on basis of limiting conditions {geometry, hardening, ...) In most of the cases, this will be an adequate simplified procedure for fatigue analysis. The remaining potential could than be evaluated by more sophisticated considerations (e.g. elbow compared to cantilever).

4. Ratcheting

Returning the question about time-saving simplified procedures in the field of inelastic analysis, the situation concerning the possible failure mode "ratcheting" is as follows:

- Due to combined primary and secondary cyclic stresses, strain enhencements can occur during each cycle. This phenomena is indicated in the upper part of figure 11.

- There are several parameters - summazized in figure 12 - which are important in view of ratcheting-strain-enhencement per cycle. 318

The procedure for predicting ratcheting strain per cycle can be recommended by generating correlations of on basis of corresponding parameter studies depending on the mentioned parameters.

As a first step this procedure has been performed for a heat exchanger tube taking into account primary/secondary stress- ratio and the substitution of an adequate mean temperature instead of considering the actual temperature distribution (lower part of Fig. 11).

This procedure seams to be very promising. It is basically comparable to the procedure of the strain concentration factors and enables to perform a vide range of practical conditions. 319

5 - Summary

Inelastic considerations become more relevant as teh structural temperature increases. In order to keep the expense of the analysis reasonable low, methods have been developed enabling simplified procedure.

As far as creep phenomena are concerned time and temperature relations could be established as creep-cross-cover-curves enabling the identification of creep relevant conditions and enabling the reduction of the number of inelastic analysis.

Concerning fatigue failure a procedure of developing strain concentration factors by considering strain range enhancement due to plastic effects has been developed. Most important parameters as for instance "geometry, material handling, multiaxiality" could be considered by corresponding strain concentration factors. Compared to ASME procedure it is characterised by a more realistic consideration with balanced conservatism.

Similar to strain concentration factors, most important parameters which have an influence on ratcheting effects have been identified. As a first step procedures have been developed enabling ratcheting evaluations on basis of simplified analysis.

Alltogether one can conclude that in spite of enhanced inelastic behavior, we are able to perform structural analysis comprehensively but economically. 320

References

/ 1 / ASME-Code Case N47-26

/ 2 / Neuber, H. "Theory of Stress Concentration for Shear- Strained Prismatical Bodies with Arbitrary, Nonlinear Stress-Strain Law", Transactions of the ASME: Journal of Applied Mechanics, 1961, pp. 544-550

/ 3 / American Society of Mechanical Engineers: "ASME Boiler and Pressure Vessel Code", Section III, 1984; The American Society of Mechanical Engineers, New York

/ 4 / H. Hübel: "A New Concept for Plastic Strain Concen- tration Factors", Second International Conference on Low Cycle Fatigue and Elasto- Plastic Behaviour of Materials held at . Arabella Conference Centre, Munich 1987 321

Time Independent Time Dependent

• Ductile short-term rupture • Tough creep-rupture • Brittle short-term rupture • Non-tough creep rupture • Fatigue failure due to load cycle • Creep fatigue failure • Buckling • Creep buckling • Ratcheting • Creep-ratcheting

• Loss of function due to excessive deformation • Environmentally effects and friction welding

Fig. 1 : Failure Modes

Sm = St

t = 0,1 -tB(1,65Sm,T) t = 0,1 •tßO.SSy.T)

RELAX ~3

.= 0,5 £p (£^ = 1 /o)

Fig 2 • CRITERIA FOR DEFINING CREEP - CROSS OVER CURVES 322

Temperatur [°C]

750 ^t = 0.1tB{1.65Sm,T) ©33% Relaxation

700 -<%•

^ ©vorgeschlagene Grenzkurve

650 ©',® 600

550

500

450 10° 10' 103 103 10* 105 Zeit [h]

Fig. 3: Temperatur-Zeit-Grenzkurven Werkstoff: Alloy 800 H

700- X10NiCrAITi3220 NiCf22Co12Mo X20CrMoV121

350 105 TIME[h]

Fig. 4 : CREEP - CROSS - OVER CURVES 323

STRESS

STRAIN £

Fig- 5 : PRINCIPLE OF PLASTIC STRAIN CONCENTRATION

ASME PROPOSAL

• STRESS LEVEL • EFFECTIVE STRESS (NOT PEAK STRESSES) RANGE

• 2 D1FFERENT TYPES • MATERIAL DEPENDANT OF MATERIAL ONLY HARDENING PARAMETERS

• GEOMETRY

• LOAD CASE (MULTIAXIAUTY)

Fig. 6: CONSIDERED. PARAMETERS FOR DEFINING PLASTIC Ke- FAKTORS 324

ASME-Ke

- CANTILEVER "NOTCH

MULTIAXIAUTY ENVELOPE

1.0 100 200 300 400 500 600 700 BOO

PEAK STRESS Sp[N/mmJ]

Fi 9- 7 : Ke . FACTORS OF INCOLOY 800 H AT 750 °C

1—T = 20 C. 2 T = 400 C. 3— T> = 550 C.

0 200 400 STRESS [N/mm2]

Fig. 8 • PLASTIC STRAIN CONCENTRATION OF AISI 316 FOR 3 TEMPERATURES COMPARED TO ASME- RECOMMENDATION 325

Ke' o -1 E-E, 3.0-

- j / 1 — CANTILEVER C = 0 / 1 •— CANTILEVER C=10000 2 — ELBOW C = 0 / 2 --- ELBOW C = 10000

1 ^^ ^- • / f 2 / — — -^ 2 --• ——--—— 1 .0 ' r ' , 4 ' 1 2 ' ' 2 '3 0 ' ' 3 4 ' '3.8

gi.el. STRESS RATIO

Fig. 9 ; COMPARISON OF CANTILEVER AND ELBOW

A£ ASME PROPOSAL ke [-1 3,33 1,49 A£[%] 0,72 0,32

NZU|[-I 110 1800

10 N CYCLES

Fig-10:FATIGUE - ASSESSMENT OF A HTR - COMPONENT

327

Discussion of the presentation:

Elastic and inelastic analysis of component behaviour

Nickel, KFA Julien, FRG: In your temperature versus time curve you show a linear decrease using the three alloys X10 NiCrAITi 32 20, NiCr 22 Co 12 Mo and X20 CrMoV 12 1 up to about 10^ h, then you have a stable line at the used temperature. What is the reason for going this way?

Seehafer, IA, Bensberg: The linear decrease of the curves in semi-logarithmic relation between temperature and time is in accordance with the corresponding material data fulfilling the set of criteria for generating these curves. As pointed out by the individual curves, there is no overall limit at 10^ hours for all materials, but there is an individual limit at a particular temperature for all materials, below which creep effects do not occur or cannot be identified by the corresponding material data. This statement is in accordnace with ASME knowledge, defining a temperature limit excluding creep at about 370 °C (700 °F).

Schuster, KFA Julien, FRG: One way you indicated to reduce the effort in inelastic analysis is to concentrate on those parts which are under the highest loads. In order to find these locations in the first place, would it not be necessary to use finite element analysis?

Seehafer, IA Bensberg, FRG: Stress concentrations always occur at characteristic structural discontinuities. These areas can be identified stepwise: - apart from any analysis, the feeling and engineering judgement is an important basis - caring for suitable constructions avoiding these areas as far as is possible, but also knowing where they will occur in reality. - highly loaded areas can be approximately identified by analytical methods. 328

- most loaded structural areas can be found out by elastic finite element analysis by using simplified models, a procedure which is not expensive.

Schad, Lurqi GmbH, FRG: Did you determine the creep cross-over curves beyond 700 °C and do you apply your simplified analysis to higher temperature ranges?

Seehafer, IA Bensberg, FR6: The creep cross-over curves are dependent above all on the material and the characteristic material data involved in the corresponding criteria for generating these curves. As we can see, for the investigated high temperature materials, the upper limit of 700 °C for the strongest high temperature alloy under consideration (Inconel 617) corresponds to a time of 1 hour only. It seems unreasonable from a practical point of view to define the creep cross over curve below 1 hour. Creep cross-over curves beyond 700 °C would therefore mean extending the discussion to other materials. 329

Significance of Fracture Mechanics

K. Schneider

ASEA BROWN BOVERI AG, Mannheim, FRG

1. Introduction

Security and availability of High Temperature Reactor power plants must be guaranteed by proving the integrity of the security relevant components. In addition to the proof of limited degradation and limited deformation during service fracture mechanics analysis must secure that catastrophic or limited failure of components will not affect safety precautions for incident prevention or control.

In the low temperature regime characterized by time independent materials properties fracture mechanics calculations are established in code cases. In the high temperature field with time dependent materials behaviour the procedure for evaluating defective components is not yet regulated. Recommendations and materials data for the evaluation of the defective components can be derived from results of recent and current HTR research programmes. 330

2. Requirements for exclusion of rupture

For the evaluation of the integrity of HTR components defects are postulated in the parent and weld metal resulting from fabrication. The incident analysis assumes failure of pressurized and loaded components starting from defects in the range of defect detectability.

These components are evaluated by fracture mechanics methods to prove that sudden failure due to these fabrication defects can be excluded.

The following requirements for the proof of exclusion of rupture are derived on the basis of the border lines for the evaluation of defective components in the low temperature regime established in "Rahmenspezifikation Basis- sicherheit" and considering the specific loading and design conditions of HTR components. The individual verifications are based on distinct postulated damage processes.

o A sufficient safety margin must be proven between the postulated original defect size within the limits of nondestructive testing and the critical defect leading to catastrophic failure of the component.

o Crack growth of the postulated defect due to service loading is negligible.

o Leak before break is guaranteed by flow stress criteria.

o The crack opening area of through-wall defects is only a small fraction of the cross section of the pressurized containment. 331

o The applicability of , the flow stress criteria is guaranteed by the proof of sufficient ductility even after long time service.

3. Procedure for the proof of integrity

According to fig. 1 a defect postulated in the high loaded area of the component will exhibit crack growth during service due to start up and shut down procedures, load changes, disturbances and continuous operation. Cracks originating from service loading must be avoided by proper design and therefore are not regarded with fracture mechanics methods. The defect dimensions at the end of service life is compared to the defect size leading to spontaneous failure. This critical defect size is determined at the most severe loading event and by the relevant materials data (toughness, flow stress at incident temperature). The variation of secondary stresses due to relaxation at high temperature must be taken into account.

3.1 Components and materials

Thickwalled pressurized containments and installations of HTR's must be evaluated by fracture mechanics methods if the failure affects safety precautions for incident prevention or control.

Regarding the power generating HTR these components are piping elements, understructures, shells, covers, thermo- sleeves, swage block headers and forgings (mouldings) out of ferritic steel X 20 CrMoV 12 land austenitic alloy 800 with service (incident) temperatures up to 535°C (600°C) and 700°C (750CC). 332

In the case of the nuclear process heat generating HTR in addition the hot gas header, mountings and hot gas ducts made out of alloy 617 in the temperature range of 900°C (950°C) must be regarded.

In the low temperature regime shells and supporting plates are fabricated with 10 CrMo 9 10 and 15 Mo 3 with temperatures up to 350°C. Small thinwalled tubes (heat exchanger) are excluded in this evaluation because limited failure is assumed by leakage minimization.

3.2 Defect size and location

Based on high quality and inspection standards for the fabrication of HTR components it is assumed, that semi elliptical surface defects with a depth of 3 mm and surface length of 20 mm remain unobserved in the weld metal of the component. Due to more favorable fabrication and test condition smaller defects are postulated for parent metals (2 x 15 mm). These dimensions cover by far the limitations of nondestructive testing (fig. 2). Defects are assumed in the high loaded region of a component (inner surface and stress concentration areas). For fracture mechanics procedure these defects are treated as sharp cracks thus incubation for crack initiation is neglected.

3.3 Loading conditions

The service loadings are composed of quasi-stationary stresses due to pressure, dead load and thermal restraint and instationary thermal induced stresses from start-up and shut-down, load changes and disturbances. 333

Decisive for the evaluation of defective components are the effective crack opening stresses perpendicular to the crack plane or shear stresses in the crack plane.

The virtual essentially strain induced stresses of the loading collective must be reduced according to plastification and relaxation at high service temperatures. At low temperature this stress relaxation has to be taken into account as additional loading.

Failure conditions are evaluated by ductile fracture criteria. Therefore primary global stresses from pressure, dead load, earth quake and loadings from adjacent piping determine ductile failure.

3.4 Crack growth consideration

According to the service conditions crack growth is caused by creep and fatigue loading. Crack growth accumulation therefore is calculated by consecutive creep and fatigue crack growth increments.

Comprehensive crack growth laws covering the relevant HTR-materials behaviour at specific temperatures as well as representative component configurations and loadings are considered in a first step to evaluate crack growth. In case of insufficient safety nargin material and component relevant calculations must be performed.

3.4.1 Fatigue crack growth

Fatigue crack growth calculations can be performed by using the Peris law with AK (range of stress intensity factor) as valid parameter to describe crack growth rates up to temperatures of 900°C: 334

da/dN = C x AKn

Depending on temperature and R-value (stress ratio} different slope factors n and constants C must be applied regarding the limitations of validity of the Paris law such as threshold value AK.. and instability.

Three characteristic temperature ranges for the integrity considerations can be stated in addition to the low temperature range (fig. 3):

550°C (X 20 CrMoV 12 1, alloy 800 parent and weld metal)

da/dN = 7 x 10"11 x AK2'8 [MPa/m, m/cycle) 700°C (alloy 800, alloy 617 parent and weld metal) da/dN - 2,84 x 10"9 x AK1'62 [MPa/m, m/cycle) 900°C (alloy 800, alloy 617 parent and weld metal)

da/dN = 1,2 x 10"9 x AK2'53 [MPa/m, m/cycle)

These Paris laws are valid for R-values less than 0,1 with threshold values exceeding 5 MPa/m. Depending on the test temperature there is a shift of the crack velocity with rising R-values (factor 2 for R = 0,5 at 850°C).

The crack growth of circumferential flaws is determined by the resulting collective of axial stresses including transients and relaxing stresses from thermal restraint.

The crack growth of axial flaws is influenced by circumferential stresses from internal pressure and transients. 335

3.4.2 Creep crack growth

At present no general valid load parameter can be stated for creep crack growth considerations for the relevant materials. Depending on the load level and expected lifetime until failure the crack growth of precracked CT-specimen is described with the load parameter K or C* (fig. 4, 5). The characteristic time t. determining the transition from K to C* controlled crack growth is defined by . K2 x (1-v2) C* (n+1) x E

v, E,C* are Poisson's ratio, Young's modulus and energy rate n is the exponent of Norton's creep law

€ = a x an

Creep crack growth laws depending on the energy rate C* derived from experimental values are determined for the temperature range defined in 3.4.1 (fig. 6). The energy rate related characterization is favored due to the relative low scatter band.

There are no tabulated values available for the determination of C* of surface cracks in complex components therefore C* must be approximated using data for circumferential and axial cracks with a/c = 0.

A comparison between calculated and measured crack growth values for the ferritic 12 % Cr-steel at 550°C (fig. 7) shows, that,the crack growth in a CT-specimen calculated with the C* vs da/dt relation (C* derived from net stress and Norton's constants) is by far nonconservative whereas the crack growth derived from the stress intensity factor K covers the measured values. 336

Comparing the complete period of crack growth with the crack initiation time it is obvious, that crack growth with cavity nucleation in front of the crack is dominating the creep crack growth process. An evaluation of the stresses causing crack growth of less than 0,1 mm originating from postulated surface defects of 3 x 20 mm in a component shows in comparison with the minimum stress rupture data and the 1 % creep strain limit, that the verification of negligible crack growth for the creep regime is covered by design (fig. 8). This still must be verified for the other materials and temperature ranges of interest.

3.5 Leak before break

The verification of rupture exclusion and leak before break must be demonstrated for the components at maximum and minimum service temperature. In both temperature regimes the materials of consideration exhibit good ductile behaviour so that ductile fracture criteria can be applied. In the case of X 20 CrMoV 12 1 weld metal Charpy V-notch tests with long time exposed material show, that a lowest service temperature exceeding ambient temperature must be defined according to available test results.

Several equations for the determination of critical crack sizes have been evaluated with pipe burst test results in the past. For ductile fracture only primary loads are taken into account (dead load, pressure, loading from adjacent piping, earth quake loading). The evaluation of axial flaws in pipes is performed by the empirical Battelle formula based on a multitude of pipe burst tests with axial flaws: 337

2 x aFl _ 12,5 xAv x E x n a = x arc cos e , I I "7 n H x n 4 xff«-, 2 x 2 x C with V

and M = (1 + 0,3798 x X2 - 0,00124 x X4)*5

The shell parameter X is determined by the crack size and pipe geometry:

X = 1,8178 x C/VR x t

Comprising on equivalent crack area the critical crack size for surface flaws can be derived for leak conditions. As an example fig. 9 shows the comparison of failure calculation for axial flaws in 12 % Cr-steel pipes at 550 C and experimental results. The shaded area indicates, that critical crack sizes exceed the pipe diameter for thin- walled pipes (t/R < 0,1) proving leak before break for low primary stresses due the design in the creep regime.

In the case of circumferential flaws thermal induced loads from adjacent piping must be regarded in addition to the primary loads. At high service temperatures these secondary loadings are reduced during service life due to relaxation thus increasing the critical crack sizes. Whereas for low service temperatures this stress relaxation causes additonal loadings reducing the safety margin between critical crack size and postulated initial crack (fig. 1). Critical crack sizes exceed 3/4 of the pipe diameter for reasonable pipe loadings using an equilibrium condition for plastic collaps based on the flow stress. 338

3.6 Leakage limitation

Despite the proof of negligible crack growth during service leakage is postulated. To ensure exclusion of pipe rupture it has to be verified that the leakage area is limited and thus resulting jet forces will not lead to pipe whip effects.

A multitude of leakage tests with different pipe dimensions and materials exhibit an empirical envelope curve which is used to prove leakage limitation for postulated leak crack lengths (fig. 10). It is supposed that the original postulated crack grows to twice the surface length until leakage occurs. The shaded area covers leak crack lengths of twice the wall thickness for thickwalled pipes (t/R < 0,3) or 10 % of the pipe diameter for thin- walled pipes minimizing the leakage area to less than 5 x 10~ of the pipe cross section.

Analytical approaches lead to nonconservative leakage areas compared to the envelop curve.

4. Design criteria for rupture exclusion

Examples in the procedure described above indicated that comprehensive fracture mechanics considerations may be successful on the basis of the special design criteria of HTR components.

Especially the proof of negligible crack growth for the relevant materials at corresponding reference temperatures as indicated before should be covered by design criteria in the creep regime. This must still be verified for the materials and temperatures of interest. 339

As indicated in fig. II a stepwise procedure for the proof of rupture exclusion is proposed.

In the case that the conditions for rupture exclusion are not met in the first step material or component specific considerations may be applied at reasonable expense.

If even individual verifications do not meet the demands for rupture exclusion either service restriction and service monitoring or design review is neccessary. 340

reduction of ductility increase of residual stress

crft \ mfn'

N safety margin S

crack growth La post, defect a

a ndt

Service Life Time fig. 1 : Proof of Integrity

highly stressed diss. weld joint areas

dimensions of postulated defects : X20 parent metal 2*15 mm X10 S-NiCr weld metal 3*20 mm X20 X—ray results of the diss. weld joint : • non uniform weld root • slag lines max. 1*5 mm • pores max. 1.5 0 fig.2 : Defect Sizes in the Water/Steam Circuit 341

10-»

o 10*-J

550 •o data 10-»-J r scatterband

10-10. .4 5 6 7 8 9 10 20 30 40 AK [MPaVm] fig.3 : Fatigue Crack Growth at Reference Temperatures

1U"' : da/dt = 2.74*1 ( m - 10-« i •• m •

•

to-« - .. 1 .

\ * : D w io- - • SG 24 MPaVm 1 + SG 29 MPaVm : : • GW 24 MPaVm ' K GW 29 MPaVm 10-"- •••I "i 10-» 10"» 10-* 10-' 10° 101 C* [W/m2]

fig.4 : Creep Crack Growth of X20 CrMoV 12 1 - at 550 «C 342

1 T 1 da/dt = 3.38*1 Q-***7** • •

CO E (—J 10-«- -M •D ;? O

• SG 24 MPaVm + SG 29 MPaVm • GW 24 MPaVm * GW 29 MPaVm 10-"- 10 20 30 40 50 60 70 80 90100 K| [MPaVm]

tig.5 : Creep Crack Growth of X20 CrMoV 12 1 at 550 °C

900 °C 700 °C 550 °C

C* = 2*anrt*v

1 10" r r i i i I i 1—I—T—r i i c 10-3 10-2 10-1 10 C* [W/m2]

fig.6 : Creep Crack Growth at Reference Temperatures 343

34-

32- X20 CrMoV 12 1 da/dt = 3.38*10-"*K.TJ7 550 °C

! j measurement

ü da/dt = 7.0 .10-^C*0-7"

20 —I > 1— 4000 8000 12000 16000 20000 Time [ h ] fig.7 : Evaluation of Creep Crack Growth Cclculation

200-

o Q-

100- 0) a

•*-> V) 80- 550 °C X20 CrMoV 12 1 60- a I I 111 I 1 I I I i.11 I I 1 I i III! I I I 1 IIII 10 103 104 10s 10« Ufe Time [ h ] fig.8 : Design Covers Creep Crack Growth Considerations 344

0.8- Experiment = 0.4 Calculation

0.6-

0.6 co CQ co CO fl) 0.4- CO Catastrophic Failure

O X 0.2-

* mtsmiE Service 2*c = 0 0.0 l111 1 1 1 1 1 ! 1 1— 0 1 2 3 4 5 6 7 8 9 10 11 12 13

Shell Parameter X

fig.9 : Verification of Leak Before Break

10°=r

X = 1.8i78*c/VR*t 10 2 3 4 5 7 Shell Parameter X

fig. 10 : Limitation of Leakage

347

Section IV:

Reactor pressure vessels 349

DESIGN CRITERIA

FOR PRESTRESSEO CONCRETE PRESSURE VESSELS

K. Sch immelpfennig

Stangenberg, Sehnellenbach und Partner Consulting Engineers, Bochum, FRG

1. SAFETY CONSIDERATIONS

Derived from the protection objective the demand is made on Pre- stressed Concrete Reactor Vessels (PCRVs) that the sealing and pressure retaining function under service conditions is ensured, extensive failure is excluded, the structural shape is preserved during accidents, and leakage is restricted to a rate which is acceptable for other components. Furthermore the thick concrete walls serve as shielding against radiation - which should not be neglected when talking about need of space and costs of a PCRV compared with other arrangements.

As is well known, PCRVs consist of different components: - the prestressed concrete structure (PCS), - the liner (core liner and penetration liners), - the closures and - the heat protection system (insulation and vessel cooling system), all of which have different functions. What is most important and a special feature of PCRVs, is the separation of load-bearing and sealing function, the former fulfilled by the PCS, the 1atter by the 1i ner. . .

Nevertheless, most of the demands.on a PCRV - if not direct task of the PCS - cannot be satisfied without participation of the PCS: •. • . . - Pressure-bearing capability is a d i.rect demand. 350

- Gas tightness of the liner during operation can only be guaranteed if the PCS fulfills its. load carrying function. - The same condition exists for restricted leakage in case of acci dents. - The design characteristics of the PCS make a decisive contribution to exclude extensi ve failure. - Preservation of structural shape in case of accidents is effected by the PCS. - The PCS also satisfies the radiation protection demands.

That is why in contrast with other PCRV components, a high safety importance is assigned to the PCS.

2. AIM OF PCS DESIGN CRITERIA

In comparison with other reinforced or prestressed concrete structures, there are some special features of PCRVs which contribute to safety without any more rigorous design criteria than for usual structures being applied:

- First, the geometry consisting of a cylinder and circular plates should be mentioned, whi ch is rather simp!e in compari son with sophisticated structural systems being built now and then for ordinary purposes.

- Second, the very thick concrete cross sections - cylinders of about 4 m, top and bottom slabs 5 to 7 m - make the structure unsensitive against local minor quality of concrete.

- Most important under this aspect is the extremely high number of single prestressing wires - about 120.000 in case of the THTR - as load-bear ing elements.

- As a further advantage, PCRVs are not exposed to the atmosphere like usual structures, but under controlled envi ronment. 351

- Furthermore the loads are strictly defined and not at random as for usual structures.

Because of the high safety importance of the PCS, design criteria have to be worked out with the aim to make the quality of a PCRV still better as it is due to its inherent features. That means desi gn rul es used for ordi nary structures have to be ex ten ded and supplemented, where necessary. Thi s mainly concerns the foil owing measures:

- use of a concrete mix geared to the special PCS demands, - extended evaluation of material data - especially for the indi v i d u a1 c o n c re te mix, - performance of detailed and realistic structural analysis for evaluating the safety against ultimate limit states and limits of serviceabili ty, - application of higher safety coefficients where necessary, - intensified supervision of materials and manufacture in comparison with common practice, - pressure testing in order to check the proper condition of the vessel, - monitoring of the mechanical behaviour throughout the whole PCRV life.

3. SYNOPSIS OF INVESTIGATIONS

The work of the group concerned with the PCS has therefore been focussed on topics in the above sense which are not sufficiently covered by the usual codes with respect to the special structure of PCSs and the -special demands on it, and different investigations yiel ding a basis for such specific design criteria have been carried out. All these are summari zed in Section Cl of

Endbericht zum Verbund-Forschungsvorhaben des BMFT "Ausle- gungskriterien für hochtemperaturbelastete metallische und kerami sehe Komponenten sowie des Spannbeton-Reaktordruckbehäl 352

ters zukünftiger HTR-Anlagen", Band III, Teil C: Spannbeton Reaktordruckbehälter. Kernforschungsanlage Jü1 ich GmbH, August 1988.

The main topics of this report - related to sections of a future code - are 1 is ted here: o Materials - Mechanical concrete data regarding actual test results - Thermal concrete data regarding actual test results - Prestressing steel data - Radiation effects o Design - Loads, load categories, design levels - Admissible concrete temperatures - Demands on structural analysis - Permissible stresses - Local stresses - Ultimate 1oad analysis - Structural details o Manufacture - Evaluation of experiences (THTR, abroad) o In-service survei1lance - Evaluation of experiences (THTR, abroad)

This listing - although even condensed - shows that it is impossible to give here a compiete presentation, and besides a lot of the results are not predestined for a presentation like this. Thus, only a couple of subjects being in the fore under the aspect of defining quality enlarging design criteria for PCSs are outlined in the following. 353

4. CONCRETE CHARACTERISTICS

The materials for a concrete to be used for a PCS shall be . care - fully selected with respect to the special requirements of PCRVs. In this sense the following criteria for concrete characteristies to be strived for should be mentioned in a code: - high compressive strength, - low sensitivity to continuous load, - low sensitivity to elevated temperatures, - low creep and shrinkage, - low thermal expansion, - low hydration heat, - good workability (pumping should be possible).

Some of these desirable characteristics are contrary. For instance high strength needs a low water/cement ratio. Combined with good workability this only can be achieved by use of a high cement content, which in turn causes higher creep and shrinkage and yields high hydration temperatures. That means, no concrete mix can be called ideal and hence be fixed in a code, and careful selection of concrete mix is an indispensable part of planning a PCRV project.

In connection with the German HTR 500 project, such a concrete mix development has been performed, which resulted in two definitely specified concrete mixes - one with basaltic, one with quartzitic aggregates. From the test program for selecting the mixes as well as from the following tests for these specific concretes extensive knowledge about material properties is at hand - especially with regard to temperature dependence. A part of these data has been put to the disposal of the "HTR design criteria" program; Fig. 1 shows some diagrams out of this as. examples. Although for these specific concretes sufficiently detailed investigations have been done and still are going to be done, it has to be asked, to what extent knowledge gained here can become part of a code. Because mixes cannot be prescribed, this 354

depends on - to what extent data are independent of the actual mix, - to what extent generalizations in connection with data from literature are possible - e.g. in the form of standardized curves or bands, - to what extent the structural behaviour is sensitive against possible deviations or scatters.

In this respect no final result could be gained. For a couple of properties it has been tried to defi ne temperature dependent guide values and band widths to be regarded, see e.g. Fig. 2. Recommendations have been made by which single tests calibration for a spec ific concrete mix would be possible. Besides, the sensitivity against variations of these values has been analysed, such as of variation of thermal conductivity from the lower to the upper bound on concrete temperature development during a typical accident, see Fig. 3. From this, certain recommendations have been derived.

As mentioned, investigations like this were only done on a small scale. The code setting commi ttee should use them as a basis for the decision, to which extent a code should comprise such guide values and whether possibly studies like this should be promoted.

5. DESIGN ANALYSIS

As explained earlier, a more detailed and realistic design analysis than for usual civil engineering structures is an essential part of quality enlarging measures for PCRVs. Thus, requirements in this connection have to be defined in a future code. It has therefore been a task within the program, based on existing stand ards and project related experience to settle the question in which manner and how distinctly formulations on this behalf must be or may be wri tten down in a code. Concern ing the demands on the analysis of the stand pipe zone of the top cap definite com- putational investigations have been carried out. 355

Some appropriate points of view shall be outlined here - first with respect to geometric modelling.

It is a strict requirement because of the thick-walled structural shape to perform the design analysis based on the continuum theory, which reasonably should be done by numerical methods.

The requi red refi nement of discretisation cannot (and should not) be definitely prescribed - for instance by stipulating numbers or maximum dimension of elements. Criteria for element size to be chosen (regarding the meaningful ness of the respective element) are given by the exspected steepness of stress or strain gradients in relation to the importance of the individual result. Analysing 1 oca! problems by model!ing separate substructures should always be considered.

On the other hand a code should also specify admissible relieves, e.g. the possibility to use an axisymmetric model if justified, such as for analysing the overall behaviour of a usual single cavity vessel.

In this sense it has been estimated important to check whether it is generally possible to refuse from explicite modelling of the stand pipe zone because of the regular arrangement of penetrations (see Fig. 4). Suitable calculations have shown that it is not only possible to describe stiffness and thermal properties of this region by a substitute fictitious homogeneous material when calculating the overall vessel or the whole top cap, but also to calculate stresses of steel tubes and confining concrete in a simple way starting from the "smeared" stress in the homogeneous material. This is also largely valid for ultimate load and accident analyses.

The liner should only be included in the PCS analy.ses where it is of significant influence on the structural behaviour, such as in case of detailed analysis of a penetration zone or in the ultimate 1imi t state. 356

Concerning the efficiency of calculation techniques in modelling material properties, it should bedemanded accord ing to the present state of knowledge that cracking and crushing of concrete in triaxial condition as we 11 as non-linear triaxial stress-strain relationship must be included in the algorithm. It will always be a problem here, how to verify the efficiency of the computer code used with respect to the individual task. Also here, no definite criteria can be layed down, and anyhow the manufacturer is requested to perform this verification.

Obviously, consideration of concrete creep - steady state in case of long term calculations, transitional in case of accident analyses - must be provided by the program used, in fact - as for most of the other character!* sties - in dependence of temperature.

5. LIMIT STATE DESIGN

The work concerning design rules has been based.an the principles used for PCRVs up to now, which means structural behaviour under service (and upset) conditions being analysed in detail with stresses being checked for permissibility, whereas ultimate load conditions are analysed only with respect to the overall structural behaviour with the open question how to treat the case of assumed pressurized cracks.

Foilowing these principles a system of permissible stresses for the different event categories, materials and if need zones of the structure has been worked out, which is shown e.g. for concrete in Fig. 5.

With respect to ultimate load analysis it has been investigated by simplified calculations under which conditions - the analysis with liner assumed to be tight or - the analysis with liner assumed to be leaking becomes the ruling case. In these considerations, the safety factor for the first case was set to 2.25 according to the former 357

HTR 1160 regulations, and to 1.5 for the second case following the British standard, both related here to the operating pressure. The re suit was that for usual wall and slab thicknesses the analysis with intact liner is more unfavourable, see Fig. 6. (The figure also shows that the design for the reference vessel carried out for satisfying the service load requirements may not meet the ultimate 1 oad demands; but the conservativity of the simpli fied calculation method may also be the reason for thi s.)

When converting all these considerations into a code it has to be (and has already been) questioned, how these design principles fit to those of the future Eurocode» the basic idea of which is that the actions multiplied by partial safety coefficients must be not greater than certain conditions described by different limit states and divided by partial coefficients related to these individual 1 inn t states. It is de fined that "a limit state is reached by the structure when a specified performance criterion is infringed". Limit states defined in the Eurocode are as follows:

o Ultimate 1imit state: - Loss of equilibrium - Gross deformation or di splacement - Attainment of maximum resistance

o Serviceability limit state: -Deformations or de fleet ions {appearance, use, other damage) - Vibrations (discomfort, damage) - Cracking of concrete (appearance, durability, tightness) - Microcracking of concrete in compression (durability, level of creep)

Obviously, it is possible to fit the PCS analyses practised up to now into this concept. Attainment of maximum resistance is the predominating ultimate limit state criterion for PCRVs, the 358

service load analyses of the previous kind may be considered as precaution against untolerabie deformations - especially due to creep -, that means loss of serviceability. To what extent they also may be interpreted as a kind of local ultimate limit state design is still to be discussed together with the levels of safe ty coefficients or permissible stresses, respectively. This includes the question how to take into account the enlarged safety requirements of a nuclear component on one hand, and on the other hand how to take advantage of the perfectly fixed loading conditions. Whatever will be defined about this in detail: The level of safety used in PCRV design up to now is obviously high enough.

6. IN-SERVICE SURVEILLANCE

Concerning the question of PCRV surveillance, published experiences with foreign vessels (Great Britain) and with the THTR vessel have been gathered.

It turned out that it is no problem to monitor the proper vessel behaviour during the pressure test. Here, a combination of vi- brating wire gauges embedded in undisturbed regions, which are easily comprehensible by calculation, with suitable deflection measurement devices is fully satisfactory. Fig. 7 shows some results like that from the THTR pressure test.

Still to be discussed is the question in which way the in-service surveillance should be best carried out. In case of non-grou ted prestressing tendons, direct and complete checking of prestressing forces is possible by using the tendon-lift-off technique. This is undoubtedly the best method to convince oneself of the proper state of a PCS. (This method is used in Great Bri- tain.)

Regarding corrosion protection most experiences exist with mortar grouting, although in a few cases not without problems. This ex- 359

eludes, however, direct measurement of prestressing forces. Use of organic grease, which has been the alternative for PCRVs up to now, is a rather you ng method. Experience for time periods of 40 years is not existing, and also here some problems are known The safer a corrosion protection system works, the less prestressing force control is necessary and the more substitution by other monitoring methods becomes possibl e. It is therefore necessary to focus further research activities on the question which combination of corrosion protection system and monitoring is the safest and most economic - even including a combination of pure wires or strands and suitable insoection methods. 360

FIG. 1 SOME TEST RESULTS

SPECIFYING PROPERTIES OF "HTR 500 CONCRETE"

1.2

1.0

u 0.8 ß I T I 1.0 puo-ci &| i— 0.6 -V i 0.4

70 \7Q ?00 500 0.2 ^rsiegeln? Pfobekörper in *C 0 I Residual strength of gravel 90 concrete after 28 days temperature 2eil Young's modulus versus time under different temperatures 1.2 (unloaded) 1,0 I=2UUUC 1.2 0.8 T-12O0 1.0 T= ?nno ______3 0.8 |0.A J|0.6 unversleg 0,2 C|^ 0.4 0 0.2 unversiegelte Pro Jekorper Zeit 90 0 Compressive strength versus time 25 30 50 70% under different temperatures Loststufen in%der Nenniesligkeit (unloaded) Compressive strength under sustained 1 oad and temperatures

1.2 1.2 T=i20°.u 10 -J 1.0 1= \ 0.8 0.8 •T=200°.— \ versiegelt :=l2C°.v £ 0.6 £0.5

unversieg-c-lt 5 0.4 "0.4 Protiekcrper 0.2 U.i u= unversiegelte

0 0 20 70 120 200 300 °C 25 30 50 70 «vSo LdStstuffin in % öer Young's modulus versus sustained Compressive strength under sustained tempera tu res (unloaded) load and temperatures 361

FIG. 2 GUIDE DATA FOR HTR CONCRETE PROPERTIES (Examples) 20

upper bound 15

mean value 10 x lower bound

«3

0 200 400 600 temperature —- Guide values of thermal expansion (limestone concrete)

3,0 W \ mK \ \ \ i *..

1,5 upper bound

•o mean value o 1 ,0 u 1ower bou nd

0,6

0 0 200 400 600 tempera ture —- Guide values of thermal conductivity (quartz it ic concrete) 362

FIG. 3 INFLUENCE OF VARYING A AND

200 [ T| 1— 1 —^—

•c O • L«neo» OOEhf ChtMZ( -

O ! L^'U'OK IWIfHf CBPUit 150 V < UnpO* OßEPt CdEKZE

fc-1 • tP P6ERC CPtUit < • LMVOA Uti'ERE CPEVJE u t» • l_*OFn* PPEBf CBtNZE

13 ^ ~^^^ tO u 100 - o. E (U

• ' • 1 • time — C o nc re te behind liner (quartz l tic concrete)

200

X • L.inBOA uMCft CREUIt « • uneo oatpt CREM2C

PMO « C» o i unto* 150

100 tu Q.

0 1 ti me Cone re te beh ind liner (1 i me stone cone re te 363

FIG. 4 IDEALIZATION OF STAND PIPE ZONE 364

FIG. 5 PERMISSIBLE COMPRESSIVE STRESSES IN CONCRETE

Load categories Permissible compressive stresses in concrete

Field stresses Local stresses

Construction *) ßr/2.0 ßr/1.65

Commissioning ßr/1.9 ßr/1.6

Pressure test ßr/1.85 ß /1.55 - i-2 ßc r * 1.45 ßc

Normal service conditions ßr/2.1 ßr/1.75

Upset service conditions ßr/2.0 ßr/1.65

Event category 3 accidents ßr/1.7 * 1-5 ßc ßr/1.4 " 1-8 ßc

Event category 4 accidents ßr/l-2 ßr/l.l **)

Multiaxial concrete strength - in case reduced due to temperature • at age of loading, at most 90 days. Uniaxial compressive strength - in case reduced due to temperature at age of loading, at most 90 days. *) If design loads cannot be exceeded or in case of short time load, field stresses up to ßr/1.8 and local stresses up to ßr/1.5 are permissible. Local stresses need not be established, if admissible with respect to PCRV safety demands. 365

FIG. 6 PARAMETRIC STUDIES OF ULTIMATE LOAD CAPACITY {related to minimum prestress for service conditions)

VARIATION OF CYLINDER WALL THICKNESS

1.5 ultimate load wi!h liner intocl pu„ = 2.25 working pressure ultimo!e lood with leoking liner pun = 1.50 working pressure

1.0

0.5

2.7 3.6 £.5 5.i 6.3 7.2 8.1 m 9.0 cylinder wall thickness •

VARIATION OF TOP SLAB THICKNESS

1.5

^^ 1.0

0.5

pwit = 2.25 working pressure ultimote lood with leoking liner 1.50 *vor king pressur e

63 72 81 90 9 9 106 m 117 top cap thickness 366

FIG. 7 THTR PRESSURE TEST COMPARISON MEASUREMENT - CALCULATION

-100

•100 10 \

•too A

-?OQ .IS" •3C0 JC .too / c '5 D --500

-600

0 10 mm 20 deflection — Key • section 90° calculated from % section 210* meosured hoop strain calculated o section 330° — calculated

5.60 * &o y \

* / — 2 _ to JC / \ 120 * 1 \ — 3 ••6 / \ •.so -/ 160 N — 4 O UC

-60 -10 -20 0 W «I M 80 O3 IM U0 .106 WO 0 » tO 60 «JO* KM + u 100 • strain cr — strain c^ — X Key . meosured. section 90° n * meosured. section 210° — colculoted

/n * M / / \ i sec I ion 90° • meosured, V section 210° eo so wo no o meosured, '1:.is seclion 330° rodius — — colculated 367

Discussion of the presentation:

Design criteria for prestressed concrete pressure vessels

Breitbach, KFA Jüiich, FRG: Is there some remarkable effect by irradiation on the concrete, perhaps some kind of weakening or ageing?

Schimmelpfennig, SSP-Bochum. FRG: Irradiation effects on concrete do exist, but in HTR designs performed up to now the amount of irradiation is within the range where irradiation effects are negligible. These limits and the procedures to be used if exceeding them are laid down in one of the documents worked out within the research programme.

Nickel, KFA Julich, FRG: You mentioned how necessary it is to protect the cables against corrosion attack. It was also shown in your requirement how important the monitoring of cables during the whole life time is. How many cables must be monitored (Percentage?) and how should this be done with these cables which are protected against corrosion in the proposed manner? What kind of research is be done in connection with the requirement to investigate or control the creep behaviour under steady state and transitional conditions? What is the requirement for accident conditions?

Schimmelpfennig, SSP-Bochum, FRG: The aim of PCRV monitoring is to obtain information on its proper behaviour through the whole lifetime without performing periodic pressure tests. If this is done by measuring prestressing forces, which is only possible in case of non-grouted tendons, the number of tendons to be monitored is based on the condition that the information gained - in fact together with information from other monitoring devices - must allow estimation of the overall vessel behaviour. A fixed percentage, 368 which is generally valid, cannot be defined, and finding the optimal constellation of corrosion protection, amount of prestressing force control and other monitoring systems needs further research. It is common practice to monitor all tendons if grease is used as corrosion protection. Creep characteristics of PCRV concretes must be known in detail, because creep causes undesired deformations under service conditions and desired decrease of restraint stresses under accident conditions. The latter is effected by the so - called transitional creep, which yields considerable stress reduction under increasing temperatures within a few hours. Thus, the main creep parameters should be investigated for each individual PCRV concrete mix.

Trumpfheiler, RWTÜV-Essen, FR6: With regard to the corrosion protection of prestressing wires there are two different opinions. One of them prefers the wires to be fitted directly into the concrete; the other opinion wants the cables to be fitted separately in pipes or channels where it is easier to perform in-service inspection of the wires. What is the present dominating opinion?

Schimmel pfennig, SSP-Bochum, FRG: Prestressing elements of PCRVs are always placed in pipes or channels. Mortar grouting in case of pipes provides corrosion protection and bonding. Since bonding is not necessary for PCRVs, the use of grease as practised in Great Britain is also under consideration here, since it allows prestressing force monitoring. Whether this advantage compensates the long time experience with mortar grouting gained with the numerous non-nuclear prestressed concrete structures (including also some negative cases) has still further to be discussed. In general there may be tendency towards non-grouting.

Paramasivam, Indian Institute of Technology, Madras, India: Is not the protection against radiation one of the criteria for designing the wall thickness of PCRV in addition to structural strength?

Can you kindly give some values for the strength of concrete and prestressing steel used in PCRVs? 369

Schimmel pfennig, SSP-Bochum, FRG: Protection against radiation is in principle one of the criteria for designing the concrete wall thickness, but since for HTRs built or designed up to now about 1 m concrete thickness is sufficient for irradiation protection purposes, this criterion does not become decisive for the design.

In actual PCRV design, the concrete compressive strength (uniaxial) is about 500 N/rnm^, the tensile strength of prestressing steel about 1700 N/mrn2. 370

DESIGN CRITERIA

FOR LINERS OF CONCRETE VESSELS

R. Oberpichler

Stangenberg, Schnellenbach und Partner Consulting Engineers, Bochum, FRG

1. GENERAL

This paper deals with the design of composite liners of prestressed concrete pressure vessels (PCPV) for future HTRs (see Figure 1).

The composite liner consists of welded steel plates and welded studs and cooling pipes working as flexible shear and/or tension anchors connecting the steel liner with the concrete vessel. On the inner surface of the liner plate the insulation is installed (see Figures 2 and 3).

Based on many years of experience in liner design a working program on devel oping design criteria with special regard to the composite liner of future HTRs was outlined by competent institutions. As already mentioned in a former paper the results are presented in the final report of a research work. In the foil owing out of this final report is referred in extracts which mainly show the design philosophy.

In the Federal Republic of Germany (FRG) up to now two 1i- ners have been built, on one hand the liner of the PCPV of the THTR-300 MWe-plant and on the other hand the liner of the concrete containment of the Gundremmingen-II-pl ant. These liners are constructed and designed in different ways, 371

so they are to be considered as prototypes. The design requirements and design criteria therefore are only valid for these special pi ants and they are only to some extent suitable to future HTRs. In the FRG no standards like those from KTA or DIN are existing on this special field of future HTRs which are describing the requirements on a composite . liner taking into consideration its safety-related importance and its constructional realization as an element of a composite structure.

Based on the concept of a liner only with sealing function the single steps for reaching this should be worked out. Especially the constructional and bearing behaviour of the composite liner as a structural member supported by the rigid PCPV should be taken into consideration for all the steps.

2. SAFETY-RELATED IMPORTANCE OF THE LINER

The main statement on the safety-re!a ted importance for liners of future HTRs is their minor safety function - although part of the activity-enclosure - compared to that of the prestressed concrete structure with its pressure bearing function and its integrity requirements. The composite liner is part of the concept guaranteeing the integrity of the safe enclosure of the cooling gas medium. Its main task therefore is 1eak-tightness during the whole lifetime of the plant. Both structural elements - the prestressed cone rete vessel and the composite liner - with their extremely different functions at least are behaving as only a unique structure (see Figures 1 to 3). The main differen- tiating features of the concrete structure and the compos i te liner are pointed out in Figure 4. 372

3. REQUIREMENTS TO THE DESIGN AND QUALITY CONCEPT OF THE COMPOSITE LINER

3.1 ANALYSIS OF THE MECHANICAL BEHAVIOUR

The special mechanical behaviour of the composite liner has been discussed in great detail because this in particular has to be the basis of defining requirements. The composite liner is mainly 1oaded by secondary or indirect loads, e.g. concrete strains at the inner surface of the vessel and temperature. Both are leading to restraints in the liner plate and the anchors. The hereby caused mechanical behaviour of the liner is called "not seifsupporting". In this case imperfections in the liner plate like initial deformations (see Figure 5) or differences in strength and/or thickness of adjacent panels are of great importance. They are influencing the local behaviour of the liner plate and result in i ncreased anchor forces and i ncreased deformations of the liner plate. Under the special boundary conditions of a not seifsupported 1 iner sudden and 1 arge sized fai1ure of the liner can be excluded if the structural integrity of the concrete vessel and consequently the composition between the liner pi ate and the cone rete vessel is maintained. Hereby the close relation of designing the composite liner and the cone re te vessel is pointed out very clearly.

Analysing the composite liner non-linearities of the liner material as well as of the embedded anchors have to be taken into account. Sufficient results are gained by analy- si ng the liner anchor system on the basis of an uncoupled two-dimensional model wh ich is not directly interacting with the concrete structure. Investigations with regard to the compl ex interacting of the liner anchor system and the concrete structure usi ng a three-dimensional model so far have shown to be unsuitable and expensive. 373

3.2 SPECIAL INVESTIGATIONS

In the course of the research program some special problems concerm" ng the mechanical behaviour of the liner anchor system have been investigated. One of these problems dealt with a more exact analysis of uniaxially stressed and initially deformed liner panels anchored only by studs. This was done by the finite element method using-2-D-soiid-elements. Because of the necessity to apply a refined element structure a large amount of computer capacity was needed. The results were shear and tension forces of the anchors and liner strains. Anchor shear forces and 1iner strains were found in good agreement with results obta ined by simplified calculation methods using plane stress mode Is. As new results anchor tension forces of not negligible quantity were gained from this analysis especially related to the adjacent anchors of the deformed liner field. As a conclusion of this further investigations of biaxially stressed and initially deformed liner panels under realistic supporting conditions should be done. However, with regard to the expen- siveness instead of highly theoretical investigations simplified calculation methods should be developed further as a tool for future design analysis.

Leak-tightness of the liner plate depends not only on its stress state but just as much on its quality of fabrication. That means, remaini ng defects in the liner pi ate after testing are to be treated with a view to their stabi1i ty under 1oadi ng conditions. For this reason an inelastic, fracture mechanics analysis was carried out in order to define criteria for preventing leakages of liner plates under real istic stress conditions and concerning the constructional and the mechanical behaviour of a composite liner. First investigations have shown, that stable crack propagation can be expected assuming usual constructional conditions for the liner p1 a te regarding initial deformations and initial cracks 374

which are not detected by testing. Stable crack propagation means in this case, such cracks will not pass through the thickness of the plate. In this field further investigations should be done possibly having an effect on the testing concept.

Further investigations were carried out'on penetration liners, for example the fuel element discharge pipe (see Fi- gure 3). Contrary to former design with thick-walled and pressure bearing pipes future HTR penetration liners will also be constructed on the basis of the composite liner concept with thin-walled and not seifsupporting liner plates. Usually the penetration liners consist of two different areas. The origi nal 1 iner area begins in si de the penetration of the "concrete slab or wall and is secondarily or indirectly loaded as mentioned in Figure 4. There is another area which in the axial direction is 1oaded directly by the inner pressure. This area is called the transition area between the steel closure and the liner. Calculations con- si der i ng sliding friction between the 1 iner as we!1 as the transition area and the concrete show a remarkabl e diminishing of the anchor forces. Further, different anchor arrange ments in the above called transition area have been investigated. This area was calculated with regard to non-linearities of steel and anchorage for single and combined load cases like concrete strains, temperature and primary forces caused by the inner pressure. The result was that in fact nearly 80 % of the whole anchor forces are only due to restraints by temperature. Consequently the inner pressure leads to a very small portion. Nevertheless the evaluation of the combined loading - primary and secondary - has to be taken into account in any case. Dependent on the state of the fuel el ernent discharge pipe the primary load may act at the inner or the outer section. Derived from the results the transition area follows up to those anchors which just reach their allowable forces. 375

As a 1 ready mentioned above sufficient results can be gained by analysing the liner anchor system on the basis of an uncoupled two-dimensional model. For th is it is necessary to know more about the coupling conditions between the liner plate and the concrete, defined by the load-displacement relationship of the anchors. Those are given by experiments and have to take into con si deration the 1iner-spec i fic boundary conditions as there are anchor spacing and liner pi ate thickness, the most important ones besides of the material properties. For this purpose requirements have been specified. Some characteristic features are sufficient variation of the main parameters like type, size and arrangement of the anchors as well as liner plate thickness, quality of the materials and kind and size of the loadings. Further there are special demands on the test pieces, the load application and the test amount. The requirements also deal with testing single and grouped anchors.

3.3 DESIGN DETAILS, FABRICATION AND TESTING

General requirements and design details have been worked out with regard to the special mechanical behaviour of a composite liner and i ts expected 1oadings as wel1 as to its safety re 1 ated importance. They have been discussed in detail and are outlined as standard proposals, containing general requirements for choosing material, for analysing the liner anchor system and for fabricating and testing it. Figure 6 shows the headlines of the design concept.

Wi th regard to the HTR-spec ific demands, criteria for choo sing suitable ma ten" als have been worked out. The criteria are related to the mechanical, thermal, radiological and chemical loads. Concerning the mechanical behaviour on materials there must be paid attention to the following requirements : 376

o Liner: - Strength - Ultimate strain - Ductility - Aging stability - Processibility - Testing o Anchorage: - Compatible with liner material o Welding filler: - High ducti1ity - Free of cracking - Compatible with base material

For this the demanded tests are spec ified which are necessary for assuring the quality of the liner plate, the studs and the cool ing pipes.

With respect to the special features of the composite liner a testing concept has been worked out. Although liner failure can be excluded practically, yet the possible consequences and necessary proceedings have been discussed for limiting damage presumed hypothetical liner leakages. The main goal s are:

o Prevention of manufacturing dependent disconnections of the material in the direction of the thickness.

o Limitation of crack propagation by guaranteeing - uncritical initial cracks and - a regular and safe composition between the liner plate and the concrete structure. 377

Kind and number of the demanded tests are described. These are : o Appraisal and design review o Requirements to the manufacturer o Manufacturing and assembly te sting - Material testing - Measurement testing - Nondestructive testing - Fabrication testing - Final assembly testing - Integral leakage testing - Operational testing

The qualification of the manufacturer is of great importance within the testing concept. Testing is to be realized by the manufacturer and by independent supervising institutions.

At least the requirements on the fabrication testing of the composi te 1 iner have been spec i fied. The results are outlined in detail in a standard proposal. Especially testing amount and the engaged supervi si ng institutions are specified in tables. Before concreting, fabrication and testing of the liner anchor system together with the documentation must be finished. The documentation consists of the design review documents, the acceptance tests of the material, records and minutes, test certifications and deviation attestations. After completion of the PCPV the integral leak-tightness of the composite liner as well as of the closures of the vessel is tested duri ng the pressure test. 378

3.4 SUMMARY

The composite liner can be expected highly safe against failure as long as the global integrity of the composite structure "steel liner + anchorage + concrete structure" can be guaranteed. Therefore the reliability of the composition between the liner pi ate and the cone re te structure is of great importance for designi ng the liner anchor system. That means the anchorage concept with regard to type, size and arrangement of the anchors, the connection of the anchors to the liner plate and the embedding of the anchors themselves in the concrete are responsible for the composition between liner and conrete. Thus, design criteria for analysing and fabricating a composite liner have been worked out which predominantly have to guarantee a safe anchorage. Standard proposals for several design details concerning material, analysis, construction and testing of the composite liner have been worked out. Beyond that, several questions remain open which should be treated for completing the design criteria. 379

Concrete closure

Penetration liner

Penetration linewr

FIGURE 1 Survey of the liner in a prestressed concrete pressure vessel 380

Core liner- insulation

FIGURE 2- Cross section of a composite liner with insulation 381 Core Core liner

c o

QJ c QJ

E O L_ l_ QJ

V) OJ

L_ V) •4- O

OJ QJ

I/) o

FIGURE Penetration liner and steel closure ot the fuel element discharge pipe 382

Concrete structure Composite liner

Requirements

Stability Leaktightness

Mechanical behaviour

Selfsupporting Not selfsupporting

Kind of loads

Primary or direct Secondary or indirect

Analysis

Allowable stresses Allowable displacements Ultimate loads Allowable strains Allowable deformations

FIGURE 4 Main design features of the concrete structure and the composite liner

384

1. Pre 1 i mi nary remarks

Scope General requirements Terms

2 . Loading s and loading combinations

3. Materials

Requ irements Properties Testi ng

4. Analysis of the mechanical behaviour

Requ irements Analytical Methods Design de taiIs Al1owables

5 . Construetion

6. Fabrication and Testing

FIGURE 6 Headlines of the design concept of a compos ite liner 385

Special Features of the Design of Pressure Vessel Closures and Heat Insulations

J. Pschowski, HRB Mannheim, FRG

Member of the ASEA BROWN BOVERI Group

1. Introduction

The activities of the working group "^g-tgJ^ZgJ1 fil-j:_91°s-u-es" and "Heat Protection System" started from the situation that up to that date design criteria had been established only for specific projects. These design criteria are contained in the licensing documents and can be evaluated only in connection with the experts opinions. A collection of design criteria, independent of the specific project, was not available in the form of a Code.

Today after four years of work of the working groups the bases for such a Code of regulations have been established. Thus for the metallic vessel closures reference can be made to a catalog of requirements for design, material selection, fabrication and testing, which meets the specific requirements of an HTR. For the heat protection system requirements to the materials, to the thermal and mechanical design as well as to manufacture and testing are available, which have been established under consideration of the safety relevance of this system. 386

From the wide scope of subjects handled by the working groups the following subjects will be presented in this report:

1. Between the penetration liner which is subjected to inner pressure, and the vessel closure there is a transitional range to which special laws and design criteria have to be applied. Based on realistic design data different anchor concepts were analysed for these transitional areas. The results can be considered as part of the design concept.

Different opinions were brought forward by the working group regarding the question, whether the metallic vessel closures were to be designed to the maximum quality assurance standards, such as claimed for the reactor pressure vessel of LWRs, or wether due to the redundant vessel closure design different design concepts can be used. In the course of the activities it became, however, evident that by making use of the basis safety criteria an integrity concept has been developed in connection with the double closure design, which meets the specific HTR conditions considering also the reduction of inservice insDections.

The scope of quality assurance measures depends on the safety relevance of the heat protection system. For this purpose the effect of its failure on the load-carrying behaviour of the PCRV and on the leak tightness of the liner have to be evaluated. The evaluation shows that the safety relevance of the heat protection system cannot be assumed to be higher than that of the liner itself. 387

Assuming that the design of the attachment fixture of the thermal insulation rules out inservice inspection, the conditions for elimination of failure had to be discussed for this purpose. Strength analyses were carried out and the fracture mechanics of the failure mechanisms was evaluated. The results permit to derive the boundary conditions for a concept of failure elimination.

2. Vessel Closures

2.1 Design Concept

The medium size and smaller penetrations of the prestressed concrete reactor vessel (PCRV) are closed using metallic vessel closures (Fig. 1). While the liner has only a sealing function, the vessel closures have a sealing as well as a load-bearing function.

Fig. 2 shows a list of all metallic vessel closures for a PCRV of the size required for the HTR 500 plant. It can be seen that in addition to the load-bearing and sealing functions the metallic vessel closures also fulfill various safety relevant and operational tasks. In total approximately 170 penetrations are provided with vessel closures.

The vessel closures can be classified in to three different types.

1. The first type is a primary closure designed as a closure with flange and double seal and an inner flow restirictor, independently anchored in the concrete, having the effect of a secondary closure (Fig. 3). During operation the flow restrictor is not exposed 388

to noticeable loads. In the event of a postulated failure of the primary closure it is ensured by the flow restrictor that the resulting leakage does not exceed the design basis leakage. For the HTR 500, e.g. the leakage cross section is limited to 33 cm2.

2. Fig. 4 shows a different type of vessel closure which is mainly used for piping penetrations such as e.g. the main steam outlet piping. The primary closure is provided by the inner closure element which is welded to the supply line and the stand pipe. The secondary closure is formed by the stand pipe and an outer flow restrictor which is independently anchored in the concrete.

3. A third type of vessel closure is shown in Fig. 5. The vessel closure is provided by an inner self-contained thimble pipe welded or flanged to the stand pipe. On the outside the penetration is closed with an outer closure acting as a flow restrictor and with a special anchoring. These closures are generally used for measuring purposes.

2.2 Requirements to the Design and Quality Assurance

The PCRV penetrations and their leak-tight closures can principally be classified into three groups with regard to design requirements and quality assurance measures (Fig. 6). The parts and elements protruding from the prestressed concrete structure are typical elements of conventional pressure vessel construction with regard to loads, materials, design shapes etc. 389

A transition range is provided to accommodate the expul- sion forces from the inner pressure as well as the loads which are typical of the liner. The transition range is subjected to special physical laws which are accurately measurable by experiment. However, for determining theoretical design requirements special considerations are required which can only be made after completion of the anchor tests currently underway. Manufacture of the transition range can follow the requirements of pressure vessel construction.

The penetration range beyond the transition range is considered as liner and shall be designed according to the liner requirements.

In the beginning of the activities of the working group there was no agreement on the question which rules had to be applied to the design of the parts and elements protruding from the prestressed concrete structure as well as to the manufacture of the transition range.

The original opinion that the vessel closures have to be designed to the maximum requirements to pressure vessels e.g. KTA 3201 or ASME Sect. Ill NB-3000 could not be maintained. In the course of the work it was realized more and more clearly that these rules furnish suitable requirements only to a limited extent. Thus the standard KTA 3201 was preferably developed for thick-walled components such as e.g. the reactor steel pressure vessel. In addition, the specific HTR safety concept and the special conditions of vessel closures have not been taken into consideration to a sufficient extent.

It seems to be more appropriate to follow a procedure according to KTA 3211 which is being established for the external systems of LWR and is to replace the general 390

specification of basis safety. This standard summarizes requirements in a reasonable form, offering the possibility to select graduated test groups according to load and structural size. These requirements furnish an appropriate concept for the vessel closures, too. This concept, which is known under the term of "basis safety concept" is particularly suitable - combined with the double closure design - to provide a failsafe solution which takes into consideration the deviation from the usual scope of inservice inspections, resulting from design.

Fig. 7 shows the principles of vessel closure integrity. The protection objectives derived from the overall HTR integrity concept (see paper Dr. Wolters):

"Maintenance of leaktightness and load-bearing function during normal and upset operation. Limited leakages in the event of accidents and exclusion of large-scale failure" are achieved by

the quality-by-production principle, known from the basis safety concept (Fig. 7, left side)

the multiple parties testing principle during manufacture and commissioning

the experience from LWR technology in applying the basis safety concept

supplemented specifically for the HTR by

the principles of plant monitoring

as well as the principle of redundant structures. 391

All theses measures as a whole allowed to exclude major leakages. This is a statement which exceeds the statements on basis safety.

3. Heat Protection System

3.1 Design Principle

It was the objective of the activity of the working group for the heat protection system to establish the criteria

for the requirements to the materialsr the thermal and mechanical design, and the -testing procedure during manufacture of the heat protection system.

Out of all the activities of the working group, the discussion will be concentrated on the topics "safety relevance of the heat protection system" and on "first considerations on a concept of failure exclusion of the attachment fixtures".

Fig. 8 shows a schematic of the heat protection system structure in a region without, any interaction. It is composed of

the thermal insulation

in this case consisting of the cover plates, the sealing sheets and the compacted insulating fiber material 392

the attachement fixture

composed of the attachement fixture bolts which are fully welded at their foot and a srewed-on hexagonal nut which is spot-welded to the cover plate and the bolt for safety reasons

in addition, the cooling pipes welded on to the liner are part of the heat protections system from a thermal dynamic aspect

It is the task of the heat protection system (Fig. 9) to protect all inner surfaces of the PCRV against the high temperatures of the primary coolant (approximately 300 °C).

3.2 Safety Relevance of the Heat Protection System

For analysing the safety relevance of the heat protection system an evaluation was performed (Fig. 10) on

the influence of heat protection failure on the PCRV load-bearing behaviour

the influence of heat protection failure on the leaktightness of the liner

the effects of failure of the liner cooling system.

It was determined that neither in the event of failure cf liner cooling nor in the event of hypothetically postulated partial or total failure of the heat protection system any effects would occur which endangered the overall load- bearing behaviour of the PCRV. Furthermore neither the possibility of reactor shutdown nor the required decay 393 heat removal are impaired. The effects of heat protection failure on the leaktightness of the liner have been evaluated by the working groups "liner" and "safety relevant boundary conditions". In these evaluations a hypothetical liner leakage was assumed to determine the safety relevance of the liner. The evaluation covered the activity release into the reactor building and the environment resulting from a liner leakage. The evaluation shows that even on the basis of conservative assumptions from the labor protection aspect as well as from the aspect of environmental protection, it is justified to classify the liner into a safety class of lower safety relevance.

Under this aspect the safety relevance of the heat protection system cannot be assumed to be higher than that of the.liner itself. This is based on the assumption that as a result of hypothetical failure of the heat protection system any leak occurring in the liner cannot exceed the size of the hypothetical leak analysed in the above-mentioned evaluation.

This leads to the following result:

Compared to the high safety relevance of the prestressed concrete structure (pressure-bearing function and leaktightness function) the heat protection system as well as the liner have a lower safety relevance. 394

3.3 Development of a Failure Exclusion Concept for the Thermal

Insulation Attachement Fixture Bolt

The boundary conditions of a failure exclusion concept were developed for the attachement fixtures. The concept was developed under the particular aspect that inservice inspections are not envisaged.

For this purpose a representative strength analyses was performed for the attachement fixture concept. In addition, fracture mechanics analyses were carried out which were, however, concentrated for the time being on the foot of the attachement fixture bolt, e.g. the point at which the bolt is welded on to the liner. Fig. 12 shows a FE-model which was applied for the fracture mechanics analyses.

The following loads were taken into account

dead weight prestressing forces of the insulating material acoustic forces flow-induced forces forces resulting from excentricity and obliqueness.

The results of a strength analysis have shown that the primary stresses are far below the allowable stresses and that the cyclic stress changes occur at a low level so that the attachement fixture with a minor additional optimization can be considered to be safe against fatigue 395

The fracture mechanics analyses for the range of the bolt foot were verified using the Kj-concept and assuming a linear-elastic behaviour. For evaluating the crack propagation an analytical and numerical integration of the "Paris equation" was performed.

In summarizing the analyses performed, the following requirements for developing a fracture exclusion concept of the attachement fixture bolts can be derived:

1. The design has to meet the basis safety criteria

2. The following additional requirements have to be made

a) the toughness of the materials has to ensure a safety factor against failure under static load of Kxc/KI > VlO. For the weldment KIQ should be greater than 2000 N/mnT3/2.

b) the primary membrane stresses shall be limited to a value of 50 N/mm2.

c) The critical flaw size under static load must be higher by a factor of 5 than the flaw size remaining after non-destructive testing and repair.

d) For excluding fatigue crack propagation the safety margin between A k and A kth must not fall below 2.5. This means that flaws above 10 % of the bolt diameter cannot be tolerated. 396

e) Since inservice inspections are not envisaged, deviations from a), b) and c) are not tolerated.

f) The final verification of the assumptions remains to be confirmed by experiment.

3. The above requirements have also to be apllied to the connection between bolt and cover plate. 397

Start-up CoU-gas neutron source temperature Coolant s»s measurement

purification system

V—l 1 *-^ Fuel element I I L—i discharge pipe L*-J

General View of PCRV Penetrations Hg.1 with Metallic Vessel Closures S&57-1

Incore rod 72 Reflector rod 48 Refuelling pipe 4 Neutron flux start-up measurement 2 Start-up neutron source ' 2 Visual inspection penetration in vessel head 6 Cold-gas temperature measurement above core 1 Coolant gas temperature measurement at decay heat removal system 2 Main steam outlet 6 Feedwater inlet 6 Hot-gas temperature measurement at steam generator 6 Visual inspection of water inlet for decay heat removal system 2 Visual inspection of water outlet for decay heat removal system 2 Bypass to helium purification system 2 Fuel element discharche pipe . _ 3 Assembly penetration (manhole) 1 Cooling water supply (or decay heat removal 2 Cooling water outlet for decay heat removal 2

Ust of Metallic Vessel Closures in a PCRV of üüJ the Size of the HTR-SOO 398

Anchorage for Anchorage for secondary closure primary closure

Supply lines

inside Flow restrictor as outside secondary closure

Primary closure

Prestressed -glass cable penetration

Penetration with Removable Closure and Row Restrictor

Fig-3 Examples of Typical Vessel Closures (Schematic drawings) 88.57-3

Anchorage (secondary)

Primary closure , How restrictor as secondary closure

inside outside

Penetration for Supply Line

Rg.4 Examples of Typical Vessel Closures (Schematic drawings) 88.57-4 399

Anchorage Anchorage (primary) (secondary)

Primary Flow restrictor as closure secondary closure

Inside outside

Penetration for Instrumentation

Examples of Typical Vessel Closures Fig.5 (Schematic drawings) 88.57- S

penetration Oner

Fig.6 Fuel Element Discharge Pipe 68.57-6 400

Prtocäpfttof Muttiple parties Ezpcrienc* from Principle of pt*Mt testing principle LWR technology redundant stnictaKS

OptMntion, Independent by «uking neol Continoom FunctiOMAy «uafiAcMion. quality -RlOmxk irtonilonnj of •eparatad COntroi ••* - H« tofcaae -dMiga - manufacturer research cross aacoon -material -technical (flow rasvtctor) expert codes in the reactor - plant suppfier — fracture boBdng douM« mechanics Pressure ^ntf dosura teraperatiw IIIIMIIPIIIIIUJ External inscrvic« vtsoection (intemal Basis »My inspection if reqöred)

a torg*n«>tur«

Integrity concept of metaMc vesid ciosura

mitor leakage

Fig.7 Principles of Vessel Closure Integrity S8J7-7

lu—

Rg.8 General View of Heat Protection System etsr-9 401

Section B-B [waifs?

Concrete Section A-A

Hexagon nut

Cover plate - Sealing sheet- Seating sleewe- Insulating fiber- material i ^5 Attachment bo«

Uner-

CooGng pipe

Fig.9 Structural Design of Heat Protection System 8837-8

Evaluation of the Safety Relevance of the Heat Protection System with regard to:

Influence of failure of the heat protection system on the load-bearing capacity of the PCRV

- Influence of faäure of the heat protection system on the liner tightness

- Effects resulting from failure of the liner cooling system

Result: Compared to the high safety relevance of the prestressed concrete structure (load-bearing function and ensurance of integrity) the heat protection system, as the liner, is of low safety relevance.

Determination of Safety Relevance of the Heat Protection System 86.57-10 402

Rg.11 FE Model of Attachment Fixture Bolt

1. Design according to the criteria of basis safety

2. Additional requirements to

a) safety against failure under static load, e.g. K b) Primary membrane stress ^ 50 N/mm2

c) Fabrication flaws ^ 10% of bolt diameter d) Elimination of vibrational crack propagation AK/AKthS2,5 e) Deviations from items a, b, c inadmissible, since inservice inspections are impossible

f) Experimental verification

HTR Fig.12 Development of a Concept for Eliminating Rupture of the Attachment Fixture Bolts 8&57-12 403

Discussion of the presentation:

Special features of the design of pressure vessel closures and heat insulations

Breitbach, KFA Julien, FRG: On THTR there are problems with broken bolts of the hot duct insulation. Has this finding an effect on the fixation concept of the liner insulation?

Pschowski, HRB Mannheim, FRG: The broken bolts in the hot gas channel of the THTR have no direct influence on the introduced failure exclusion concept of the insulation fixture bolt. The concept is developed for ferritic material to temperatures of max. 300 °C. In the hot gas channel of the THTR we have temperatures of 750 - 800 °C. 404

The HTR-module pressure vessel unit, design criteria and safety philosophy

G. Neumann, K. Heidt Siemens AG ÜB KWU, Erlangen K. Dumm, H. Rothfuß Interatom GmbH, Bergisch Gladbach

Abstract

A modular, helium-cooled high temperature reactor system for the coyeneiTÄtion of electricity and process heat has been developed by Siemens-Interatom. The primary system internals are housed in an interconnected arrangement of pressure vessels, "the pressure vessel unit". Based on Siemens experience with light water reactors, this pressure vessel unit is made of the proven LWR steel 20 MnMoNi 5 5.

Design, manufacture and operation of the pressure vessel unit will conform to German nuclear codes and standards for LWRs, such as RSK Guidelines and KTA Safety Standards. The general applicability of these rules is justified by comparisons of forged product forms, dimensions and operational parameters. Since the above-mentioned codes and standards were drawn up specifically for LWRs, some deviations or peculiarities for their application to HTRs are unavoidable. These are for instance:

The main steam nozzle, through which the steam line at 530 °C penetrates the steam generator pressure vessel which has a nominal design temperature of 350 °C. A special thermal sleeve is needed.

The pressure test concept, especially for inservice inspections. The preservice pressure test of the pressure vessel 405

unit will be performed in complete, accordance with the codes and standards at 1.3 times the design pressure of 70 bar and using water. Afterwards, the presence of graphite structures, ceramic insulation and, of course, the pebble bed core in the primary system must be taken into account. Hydrostatic pressure tests are.prohibited in this case; instead pneumatic tests are performed at 1.1 times design pressure accompanied by more detailed ultrasonic examinations.

The position of operational material irradiation surveillance specimens has to be chosen carefully (max. fluence 9 x 10 N/cm2 at a temperature of 195 °C). Accelerated irradiation specimens in the reactor are not considered. Design postulates concerning the increase of the nil ductility transition temperature (^T^pm) wiH ^e confirmed in a separate irradiation program.

In general, the requirements of the assured safety concept, aimed to rule out catastrophic failure of the pressure vessel unit during its lifetime are fulfilled.

The described design and the application of existing rules are an important part of the HTR Module concept. It is interesting to note that an official procedure for a non-site-specific concept licence for this reactor concept in Germany has been in progress since summer 1987. 406

Contents

1. Introduction

2. Description of the Pressure Vessel Unit; Comparison with Light Water Reactor Technology 2.1 Design 2.2 Design Criteria 2.3 Materials 2.4 Manufacture and Inspection

3. Applicable Codes and Standards 3.1 Requirements on the Pressure Vessel Unit 3-2 Operational Monitoring 3.3 Inservice Inspections

4. Specific Features of the HTR 4.1 Pressure Test Concept for the Primary Side of the Pressure Vessel Unit 4.2 Monitoring of Radiation Embrittlement of the Reactor Pressure Vessel 4.3 Electrical Penetrations

5- Safety Concept 5.1 Consideration of Operational Aspects 5.1.1 Operating Fluid Helium 5.1.2 Accident Loadings 5.2 Preclusion of Catastrophic Failure of the Pressure Vessel Unit 5.2.1 Preclusion of Failure 5.2.2 Monitoring of Safety against Failure 5.2.3 Fracture-Mechanics Analysis of Safety against Failure

6. Summary 407

1. Introduction

A modular helium-cooled high temperature reactor system especially for the cogeneration of electricity and process steam has been developed by Siemens-Interatom (Figure 1). For inherent safety reasons, the thermal

output of each modular unit is limited to 200 MWtn. The typical arrangement is shown in the cross section of a modular unit equipped with a steam generator (Figure 2). Helium at 60 bar and 250 °C flows downwards through the pebble bed ana is heated up to 700 °C. Then the helium passes through a concentric hot gas duct to the steam generator in which it flows downwards through the tube bundle to produce steam at 190 bar and 530 °C. From the steam generator the helium flows upwards through an annular gap surrounding the tube bundle to a blower which conveys the helium via the cold gas annulus to the graphite reflector structure surrounding the pebble bed core. Flowing upwards through channels in the reflector, the helium returns to the core inlet plenum. The helium mass flow is 85.5 kg/s. The entire primary system is located in one steel pressure vessel unit. The reactor cell is equipped with a threefold redundant cavity cooling system. It removes heat losses during operation and residual heat in case of failures in the main heat transfer system.

Based on Siemens experience with LWRs, the proven material 20 MnMoNi 5 5 was selected for the pressure vessel unit.

Established German nuclear codes and standards for LWRs such as RSK Guidelines and KTA Safety Standards are used in design (Figure 3). On account of certain HTR-specific

409

.05900 mm. -|094OO mm 1 Pebble bed 2 Pressure vessel 3 Fuel discharge 4 Small absorber balls 5 Reflector rod 6 Fuel loading 7 Steam generator: Pipe assembly 8 Outer shroud 9 Feed line 10 Live steam line 11 Blower 12 Hot gas duct 13 Surface cooler 14 Insulation

if (5*=^

HTR-Module: Cross Section of a Modular Unit with Steam Generator (Primary Circuit) Figure 2

411

features, there are some slight exceptions in the applicability of these rules which are described below. It is important to mention that an official procedure for a non-site-specific concept licence for this modular reactor system in Germany has been in progress since summer 1987. Of course, special importance is attached to the licensability of the pressure vessel unit.

2. Description of the Pressure Vessel Unit; Comparison with Light Water Reactor Technology

2.1 Design

The design and arrangement of the pressure vessel unit, which forms the reactor coolant pressure boundary, are apparent from Figure 4. The pressure vessel unit consists of the components:

Reactor pressure vessel Gas duct pressure vessel Steam generator pressure vessel with blower pressure vessel section

The principal dimensions of reactor pressure vessel and gas duct pressure vessel are given in Figure 5. The reactor pressure vessel includes connections for

Fuel discharge with failed fuel separator block Fuel feed with valve bank Small ball shutdown system with valve banks

Figure 6 gives the principal dimensions of the steam generator pressure vessel with blower pressure vessel 412

25200 Blower pressure vessel section

Live steam nozzle

Upper tangential support

21629

Connecting pressure vessel Lower 5500 tangential support

Feedwater nozzle

Reactor pressure vessel Steam generator HTR-Module Pressure Vessel Unit Figure 4 413

•0 6700 1800

Penetration pad Tangential support

0 5900

25200 18975

Pressure equalizing nozzle Connecting pressure vessel Support bracket 01500

Fuel charge 01200 KLAK connection 5500 4500 Fuel discharge I

-4250-

HTR-Module: Reactor Pressure Vessel and Connecting Pressure Vessel Figure 5 414

Penetration Blower pressure 3479 pad vessel section

Main steam nozzle 1990

=0= 2400 Uppertangential support Connecting - pressure vessel 140- nozzle Support bracket 21629

11800

Lower tangential support

Feedwater nozzle

HTR-Module: Steam Generator Pressure Vessel Figure 6 415

section. The steam generator pressure vessel includes feedwater and main steam nozzles for connection to the water/steam cycle (secondary system).

The design of the pressure vessel unit draws on experience gained during the construction and operation of light water reactor pressure vessels by virtue of

Comparable design conditions Comparable dimensions

Figure 7 compares the design and dimensions of reactor pressure vessels for

1300 MWe PWR plant

1300 MWe BWR plant HTR Module plant

It is apparent that the same design principles are observed for all important design features such as

Product form (only seamless forging) Flanged joints Nozzles Closure head and bottom head

This applies especially to the arrangement and design of welds in the pressure boundary and also of attachment welds, e.g. for vessel supports.

Against the background of light water reactor experience it is evident that the requirements for

Proper functioning and loading

417

Proper selection and use of materials Good manufacturing practice and ease of examination

which are crucial to the safety of the pressure vessel unit are certainly fulfilled.

2. 2 Design Criteria

A distinction is made in the design of the pressure vessel unit between the primary gas envelope and the areas around the main steam and feedwater nozzles (Figure 8).

On the primary side the HTR Module concept rules out contact of the hot gas with the pressure boundary. Furthermore, the pressure difference between cold and hot gas protects the pressure boundary even in the event of leaks in the walls separating hot from cold gas.

The temperature range postulated in design of the primary system is governed by functional conditions in the core region of the reactor pressure vessel. Since the annular gap between reactor pressure vessel and core barrel is filled with stagnant helium, the reactor pressure vessel wall is only heated orcooleolby radiation and convection.

Consequently:

a) The RPV wall warms up very slowly during initial plant start-up and restarts after long plant shutdowns. Therefore the smallest margin to the

actual RTNDT temperature occurs during these service 1 Pebble bed reactor core 2 Graphite reflector 3 Reactor pressure vessel 4 Connecting pressure vessel 5 Steam generator pressure vessel 6 Small absorber balls shut-down system 7 Absorber rods 8 Steam generator heat exchanger 9 Feedwater nozzle 10 Main steam nozzle 11 Blower 12 Hot gas duct 13 Surface cooler 14 Fuel charge tube 15 Fuel discharge system

HTR-Module: Cross Section of Modular Unit with Steam Generator (Primary Circuit) Figures 419

conditions. They are therefore the design basis for protection against brittle fracture. b) As a result of the design basis conditions

Reactor scram with residual heat removal via the cavity coolers at full primary system pressure Depressurization accident with residual heat removal via the cavity coolers

The decay heat of the reactor core and heat removal via the cavity coolers cause a temperature increase in the RPV wall. This temperature of 350 °C is the maximum design temperature for the pressure vessel unit.

The axial temperature profile in the RPV wall is given in Figure 9 for nominal operation and design-basis conditions.

The operating, design and test data for the pressure vessel unit are summarized in Figure 10.

On the secondary side, the design criteria are governed by the water/steam cycle. The operating and design data for the feedwater nozzle are apparent from Figure 11, those, for the main steam nozzle from Figure 12.

The scope of nuclear codes and standards applicable to reactor coolant system components of light water reactors o is limited to design temper^turejs^fj^axjJ^400_ C^ Therefore these codes and standards can be applied in full to the primary side of the pressure vessel unit. 420

Hotstand by (75 h)

Upper bound ofcore 14715mm

Pressure relief accident (90 h)

Nominal operation

1 Core 2 Graphite reflector 3 Carbon brick 4 Core barrel 5 Reactor pressure vessel 6 Insulation

100 150 200 250 300 °C 350 • Temperature HTR-Module: Reactor Pressure Vessel Inner Wall Axial Temperature Profile Figure 9 421

Operation

Pressure 60 bar Temperature 250 °C RPV: local max. 262 °C (see Fig. 9 for axial temperature distr ibution)

Design

Pressure (gauge) 70 bar Temperature 350 °C

Pressure Tests

Preservice test Inservice test Pressure (gauge) 91 bar 77 bar Temperature min. 33 °C approx. 50 °C Medium Water Hel ium

Fig. 10

HTR Module: Pressure Vessel Unit (Primary Side) Operational, Design , and Test Data

424

Generally accepted technical rules (conventional codes and standards, ASME Code) are also available for component requirements governed by the higher secondary- side design temperature.

2.3 Materials

The base material for the pressure boundary of the pressure vessel unit is the same material as that used for light water reactor components, i.e.

20 MnMoNi 5 5.

Materials not covered by the codes and standards for light water reactors are used for the main steam nozzle for which design temperatures exceed 400 °C. Figure 13 shows the materials used. These materials have been qualified by many years' satisfactory performance in service in conventional power plants and by supplementary testing in connection with construction of the nuclear power plants THTR 300 and SNR 300.

2.4 Manufacture and Inspection

These statements on design and materials show that the nuclear codes and standards can be applied in full to light water reactors. It has been demonstrated that these rules ca\i also be applied for quality requirements and verifications relating to manufacture and inspection.

426

3. Applicable Codes and Standards

The availability of accepted codes and standards for the pressure vessel unit of the HTR Module, for design, manufacture, quality control, including inservice inspections, is substantiated in the following. Apart from a few exceptions which result from concept- related differences to light water reactor technology, these codes and standards are the "KTA" nuclear safety standards of the Federal Republic of Germany. These standards cover the requirements of conventional codes and standards too. For the few HTR-specific features of the pressure vessel unit, it is shown elsewhere that solutions have been found which are equivalent from the safety standpoint.

3.1 Requirements on the Pressure Vessel Unit

The preceding information on design, materials, manufacture and inspection has shown that, apart from the main steam nozzle region with a design temperature of more than 400 °C, the necessary quality requirements and verifications are identical to those for light water reactors. This means that KTA Safety Standard 3201, Parts 1 to 3 can be applied in full.

Supplementary stipulations for materials and calculation need only be established for the main steam nozzle. In this case, the type and extent of material testing and the material properties to be attained can be based on experience with conventional power plant construction and supplementary testing carried out in connection with construction of the nuclear power plants THTR 300 and SNR 300. 427

The conventional codes and standards and the specifications of the ASME Code, Sect. Ill, Code Case N47 provide sound references for calculation, i.e. dimensioning considering time-dependent strength values and for design by analysis.

3.2 Operational Monitoring

KTA 3201.4 defines the requirements on operational monitoring as follows:

"Continuous operational monitoring includes the recording of measured variables and data of importance to the integrity of primary system components. In principle, this means monitoring of compliance with limits on which the design is based".

Further, more general requirements are established for instrumentation and monitoring.

The monitoring concept of the HTR Module fulfils these requirements, taking account of the differences specific to the concept.

3 .'3 Inservice Inspections

HTR-specific criteria must be assessed on the basis of the requirements of KTA Safety Standard 3201, Part 4 for inservice inspection. A distinction is made between

Non-destructive examination Visual examination Pressure tests 428

A concept has been prepared for inservice non-destructive examination which fully complies with the requirements of the valid edition of KTA 3201.4 as regards planned extent of examination, examination procedures and times of examination. This applies completely to examination techniques for the reactor pressure vessel. For the other components of the pressure vessel unit, the volume examination is limited to internal and external near- surface zones in anticipation of the revised version of KTA 3201.4.

The inservice inspection concept includes visual examination from the outside for the entire pressure vessel unit and additionally from the inside for the secondary side of the feedwater and main steam nozzles. Visual examinations of the primary side from the inside are not necessary since damage to the pressure vessel unit wall as a result of corrosion/erosion can be ruled out since the operating fluid is helium and flow velocities are low. Therefore the requirements stated in KTA 3201.4 can also be considered as fulfilled with regard to visual examination.

Inservice pressure tests of the secondary system are performed at 8-yearly intervals at 1.3 times design pressure. These constitute hydrostatic pressure tests conform to the requirements of KTA 3201.4.

The pressure test concept for the reactor coolant system differs from that for light water reactors because of the different boundary conditions. It is discussed in the following. 429

4. Specific Features o£ the HTR

It has been shown that the associated nuclear codes and standards for light water reactors can be applied practically in full for the pressure vessel unit since the technical boundary conditions are similar. However, certain specific features of the HTR neces- sitate various plant-specific stipulations.

4.l Pressure Test Concept for the Primary Side of the Pressure Vessel Unit

The pressure test concept is governed by the inservice pressure tests of the reactor coolant system. As a result of the reactor concept, (graphite internals, loaded core) these pressure tests can only be carried out with the operating fluid helium. The applicable conventional codes and standards which, through the "Druckbehälterverordnung" (Pressure Vessel Code), have statutory status in the Federal Republic of Germany do not permit pneumatic test pressures above 1.1 times design pressure. This is for reasons of occupational safety. However, the nuclear codes and standards for light water reactors demand that inservice pressure tests should permit an evaluation of safety comparable to that possible with the preservice pressure test which is performed with the water as test fluid at 1.3 times design pressure. The following pressure test concept is planned to satisfy both criteria, i.e. occupational safety and the required safety evaluation:

During construction: 430

. Preservice pressure test of the installed pressure vessel unit with water at 1.3 times design pressure . System pressure test prior to commissioning with gas at 1.1 times design pressure . Baseline examination, i.e. non-destructive examination of the pressure vessel unit including comparison with the results of examinations during manufacture.

During operation, inservice pressure tests will be performed at 8-yearly intervals with helium at 1.1 times design pressure. Each pressure test will be preceded by a full-extent non-destructive inservice examination. Any findings will be assessed by fracture mechanics analysis. After the pressure test, highly-stressed areas and, as applicable, areas in which indications were detected will be subjected to a repeat nondestructive examination.

4 .2 Monitoring of Radiation Embrittlement of the Reactor Pressure Vessel

Nuclear codes and standards require monitoring of radiation embrittlement for items receiving an EOL fluence greater than 1.1017 N/cm2 (E > 1 MeV). Details of surveillance programs are established in KTA 3203. This also states acceleration factors for the irradiation of surveillance specimens. The axial profile of EOL fluence and temperature on the inside of the RPV are plotted in Figure 14. This shows that a surveillance program must be carried out for the HTR Module to satisfy the nuclear codes and standards. 431

O(E>1MeV) n/cm2 1018 -4

Upper bound Neutron fluence of core ! 0E>1MeV + 14715 mm

Nominal operation

Position irradiation Temperature sample

1 Core 2 Graphite reflector 3 Carbon brick 4 Core barrel 5 Reactor pressure vessel 6 Insulation 18J 300 °C 350 • Temperature HTR-Module: Reactor Pressure Vessel Inner Wall Axial Temperature Profile Neutron Fluence E>1MeV Figure 14 432

Since accelerated surveillance is not planned, the following concept was drawn up. It permits a. technical evaluation comparable to that required by KTA 3203.

Separate accelerated irradiation program to confirm

the design basis RTNDT = 10 K. Monitoring of the actual radiation loading of the reactor pressure vessel by means of irradiation surveillance specimens and comparison of the results with those of the testing program. The times at which the 4 sets of specimens are to-be removed are governed by measurement tolerances, by the safety-evaluation of the actual condition of the components and by the need for forecasts for subsequent operating cycles.

Figure 15 shows the arrangement of the specimen capsule in the annular gap between reactor pressure vessel and core barrel.

4.3 Electrical Penetrations

Cable penetrations through the reactor coolant pressure boundary of the pressure vessel unit (gas side) are necessary for the electrically powered controls, instru- ments and drives located inside the pressure vessel unit Prestressed glass penetrations tailored to this purpose are used. These penetrations consist of a steel support plate in 15 MnNi6 3, into which the conductors are sealed with special glass as insulator. Together with the steel support plate, these glass items form part of the reactor coolant pressure boundary. The steel support plate is flanged (Figure 16). 433

Section C-D Section A-B RPV Core barrel RPV Core barrel

4 Specimen capsules uniformly Peripherie distributed Content of each capsules: 12 Charpy-V specimens base metal 12 Charpy-V specimens weld metal 12 Charpy-V specimens heat affected zone of a representative welded joint Fluence detectors

Irradiation Program for Reactor Pressure Vessel Figureis 434

Electric Penetration Figure 16 435

The penetrations are provided for both the reactor - pressure vessel and the blower pressure vessel section

5 . Safety Concept

5 .1 Consideration of Operational Aspects

"It"has been shown that the requirements of the relevant nuclear codes and standards for light water reactors have been fulfilled with regard to design and manufacture, and also for surveillance and inspection during operation; furthermore it is demonstrated that equivalent safety provisions are made for special features of the plant concept. Before discussing the safety concept, two important examples are considered to show that the safety philosophy of light water reactors can also be applied since the loading_during reactor operation jire comparable.

5.1.1 Operating Fluid Helium

The operating fluid helium has the following advantages over water as a coolant with regard to loading of the walls of the pressure vessel unit:

Corrosion is ruled out. Therefore austenitic cladding of the pressure vessel unit is not necessary. The otherwise necessary monitoring effort is obviated.

As a result of the properties of the heat transfer fluid, all transients take place more slowly than with water as coolant. Consequently the thermal 436

loading of the reator coolant pressure boundary is basically lower. Temperature shocks are practically impossible.

5.1,2 Accident Loadings

Since the cavity cooler alone is adequate for residual heat removal in the event of accidents in the HTR Module, i.e. a passive emergency cooling system from the standpoint of the pressure vessel unit, the consequence will always be an increase in wall temperature. As a result, the safety margin to brittle failure always increases, fracture-mechanics analyses of the safety margin for accidents and, possibly, corresponding operational monitoring measures are not necessary.

5. 2 Preclusion of Catastrophic Failure of the Pressure Vessel Unit

As has been demonstrated, the characteristics of all individual features governing the safety of the pressure vessel unit are comparable to those for the corresponding components of light water reactors. Therefore a validated and acknowledged safety concept also exists to verify the preclusion of catastrophic failure of the pressure vessel unit. This concept is based on three independent groups of protective features (independent redundancies). Failure of the pressure vessel unit can be ruled out throughout the plant's lifetime by each of these groups of features. Figure 17 shows this "redundancy concept".

438

5.2.1 Preclusion of Failure

Compliance with the requirements of KTA 3201, Part 1 to Part 3 provides assured safety

"which rules out catastrophic component failure as a result of defects originating in manufacture".

(PWR RSK Guidelines 10/81, Section 4.1.2)

5.2.2 Monitoring of Safety against Failure

Regardless of the assured-safety design of the components, all factors affecting the safety of the pressure vessel unit against failure are monitored to such an extent over the entire lifetime that catastrophic failure can be ruled out at all times. Essential measures are:

Inservice inspections according to KTA 3201.4 Continuous operational monitoring in compliance with the principles of KTA 3201.4 Irradiation surveillance according to KTA 3203 Leakage monitoring based on fracture-mechanics leak- before-break analysis, i.e. reliable detection of postulated stable leaks in the walls of the pressure vessel unit

5.2.3 Fracture-Mechanics Analysis of Safety against Failure

Assuming the worst conceivable boundary conditions (worst case principle, lower bound principle), it has been verified by means of representative fracture-mechanics analyses that catastrophic failure of the pressure vessel unit is ruled out by mechanical properties of materials 439

as has been confirmed by tests. The following verifications were provided:

Negligible growth of conceivable manufacturing defects throughout the plant lifetime Attainment of the critical crack depth and/or formation of a stable leak only after many reactor lifetimes Fracture-mechanics verification of the design basis leak Determination of the necessary detection sensitivity for leakage monitoring systems Preclusion of brittle fracture for all conceivable operating conditions and accidents. 440

6. Summary

It has been shown that the pressure vessel unit of the HTR Module can make use of experience gained from the construction and operation of reactor coolant system components of light water reactors. Therefore proven techniques and accepted codes and standards are available for all aspects such as

Design Material Manufacture, inspection Operational monitoring, inservice inspection

At the same time, the safety concept of light water reactor primary system components against catastrophic failure, which is characterized by multiple redundancies, can be applied in full. This means that the integrity of the pressure vessel unit of the HTR Module is assured for the total lifetime.

All relationships and conclusions briefly presented here . have been examined in depth and from the basis for the concept licensing procedure mentioned in the introduction. Therefore, with reference to the pressure vessel unit, the applicants do not anticipate any fundamental problems concerning the licensability of the components under consideration. 441

Discussion of the presentation:

The HTR-moduie with pressure vessel unit, design criteria and safety philosophy

Alder, PSI Wtirenlingen, Switzerland: How long is the lifetime of the reactor pressure vessel in the HTR? Would it be possible to exchange the RPV at the end of life?

Neumann, Siemens UBKWU, FRG: The spezified value for the lifetime of the RPV is 32 year full operation time and 40 years lifetime. The exchange of the RPV is not foreseen and has not yet been investigated. 442

DESIGN PRINCIPLES FOR MHTGR PRESSURE VESSELS

C. L. HOFFMANN NUCLEAR POWER SYSTEMS MATERIALS & CHEMICAL TECHNOLOGY COMBUSTION ENGINEERING, INC. WINDSOR, CONNECTICUT, USA

PRESENTED AT THE WORKSHOP ON STRUCTURAL DESIGN CRITERIA FOR HTR JUELICH, FED. REP. OF GERMANY JANUARY 31 - FEBRUARY 1, 1989 443

DISCLAIMER

This report was prepared as an account of work sponsored by the United States Government. Neither the United States nor the United States Department of Energy, nor any of their employees, makes any warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed, or represents that its use would not infringe privately owned rights. Reference herein to any specific commercial product, process, or service by trade name, mark, manufacturer, or otherwise, does not necessarily constitute or imply its endorsement, recommendation, or favoring by the United States Government or any agency thereof. The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof. 444

DESIGN PRINCIPLES FOR MHTGR PRESSURE VESSELS C. L. HOFFMANN COMBUSTION ENGINEERING, INC. WINDSOR, CONNECTICUT, USA

ABSTRACT Design of the Modular High Temperature Gas-Cooled Reactor (MHTGR) is based on extensive utilization of passive safety characteristics. The MHTGR design incorporates an uninsulated, steel reactor pressure vessel to allow decay heat removal by conduction and radiation even with a total loss of coolant or coolant flow. The MHTGR design also maximizes the use of LWR pressure vessel technology for design and materials to take advantage of the extensive experience base in the licensing, fabrication and operation of the plant.

The normal operation and passive safety features of the MHTGR reactor vessel result in operating conditions which differ from typical LWR vessels. The MHTGR reactor vessel normally operates at lower temperature than that of a typical LWR. However, certain low probability off-normal events can raise metal temperatures in some regions of the reactor vessel above 370"C (7OCPF}, which is the maximum temperature allowed by Section_III of the ASME Boiler Pressure Vessel Code for the selected presstrre""veTseTlhateriaFl;. Thorder design rules and materials, an application specific ASME Code Case is being sought to allow limited exposure of the selected pressure vessel materials at temperatures above 370°C (700°F). An experimental materials test program is underway to support the approval of this Code Case. The test program is intended to characterize the time/temperature dependent behavior of the MHTGR pressure vessel materials.

Prediction of the irradiation-induced shift in the nil-ductility transition temperature (NDTT) for the MHTGR reactor vessel is based upon an extensive data base and accepted methodology for LWR vessels. Since the MHTGR reactor vessel operates at temperatures somewhat below the range over which the LWR data base and methodology was developed, additional conservatism is applied to the shift predictions in the design. The calculated NDTT shifts will be verified by an experimental irradiation program during the design phase. In addition, a reactor pressure vessel surveillance program will monitor the change in material characteristics during the operating life of the reactor vessel. 445

INTRODUCTION The Modular High Temperature Gas-Cooled Reactor (MHTGR) incorporates design features intended to produce an inherently safe, passively controlled system. The passive safety characteristics of the MHTGR include an uninsulated, steel pressure vessel to allow decay heat removal by conduction and radiation even with total loss of coolant or coolant flow. The MHTGR is the only reactor concept currently proposed that has the capability to sustain a total loss of coolant without significant damage to the fuel or vessel system (1). In addition to developing a passively safe plant, a design objective for the MHTGR vessel system is to maximize the use of proven LWR pressure vessel technology for design and materials. The extensive experience base of LWR pressure vessel design and materials is utilized to facilitate the licensability, fabricability and operability of the MHTGR.

Operating characteristics of the MHTGR differ from typical LWR systems, in . part due to differences between the normal operation of LWR and HTGR reactors and also as a result of conditions resulting from the passive nature of the reactor safety systems. The normal operating temperatures of the MHTGR reactor vessel are considerably lower than temperatures found in a typical LWR. In general, lower irradiation temperatures result in increased damage. However, the MHTGR flux is lower and the neutron energy spectrum is composed of a higher percentage of lower energy neutrons which cause less damage. The differences in MHTGR irradiation environment are qualitatively offsetting but additional reseach is required to establish the quantitative relationship between these variables.

In addition, certain low probability off-normal operating events in the design duty cycle are predicted to result in metal temperatures higher than 370°C (700*F), the maximum service temperature for which allowable stresses are provided by Section III of the ASME Boiler and Pressure Vessel Code for the selected materials.

The effects of both normal and off-normal operating environments on design requirements for the MHTGR reactor pressure vessel are discussed in detail in this paper. The principal factors to be addressed by the design of the MHTGR reactor vessel are: 1) Elevated-temperature service of reactor vessel materials during off-normal events where temperatures exceed 370°C (700°F), and osed to the

The following discussion provides a brief description of the MHTGR design, the background, designers alternatives, selected design approach and the technical issues associated with each of these factors. Discussion of the technical issues is focused on the metallurgical considerations associated with elevated temperature service and irradiation embrittlement of pressure vessel materials. The current status of ongoing programs intended to satisfy the design data needs of the MHTGR reactor pressure vessel is presented. 446

DESIGN PRINCIPLES FOR THE HHTGR VESSEL SYSTEM The vessel system comprises three distinct, interconnected steel pressure vessels (Figure 1): a reactor vessel which houses and supports the reactor system; a steam generator vessel which houses and supports the steam generator and main circulator; and a crossduct vessel shich houses an internal concentric duct and provides flow passages between the reactor and steam generator. The configuration of the flow paths isolates the vessels from the hot helium, thereby avoiding elevated temperature service during normal operation. The primary objective of the MHTGR design is to produce an inherently safe nuclear steam supply system (NSSS). One key feature of the design, intended to enhance safety, is the passive heat removal system which can prevent excessive fuel temperatures and consequently, fission gas release, during periods of complete loss of forced coolant flow. This requires an uninsulated reactor vessel in order to allow residual and decay heat removal from the fuel through the reactor vessel to the passive heat removal system by conduction and radiation. In addition to inherent safety and passive controls, a goal of the MHTGR design is to maximize the use of light water reactor (LWR) technology in order to maximize reliability and minimize risk. The intent of the MHTGR design is to utilize existing design procedures and materials of Section III of the ASME Boiler and Pressure Vessel Code. Utilization of both design and materials experience from LWR's will benefit the MHTGR program by providing a highly reliable vessel system.

MHTGR REACTOR VESSEL OPERATING CONDITIONS The MHTGR reactor vessel is uninsulated to allow passive removal of decay heat by conduction and radiation to the vessel cavity which is cooled by the natural convection of ambient air. The normal operating conditions of the MHTGR reactor vessel are less severe than the conditions found in LWR's. The MHTGR operates at lower pressures, lower temperatures, and accumulates considerably less neutron fluence than LWR reactor vessels. During normal operation, the cold inlet helium is channeled within the reactor vessel to an upper plenum where the coolant is directed downward through the reactor. The hot helium is collected in a lower plenum and the coolant flows into an internally insulated duct, concentric with the crossduct vvessel, conveying the hot helium to the steam generator. The inlet helium thermally radiates to the uninsulated reactor vessel which is passively cooled by the passive heat removal system. Metal temperatures of the reactor vessel during normal operation are in the range of 120" to 205°C (250° to 400°F) compared to PWR vessel temperatures of 288° to 305°C {550° to 580°F). The design duty cycle of the MHTGR contains several low probability events (Service Level C and D conditions) in which all forced circulation is lost. For the events where the helium remains pressurized, extremely hot helium collects in the upper plenum by natural convection. Insulation on the upper plenum structure prevents a severe overtemperature on the reactor vessel upper head. However, reactor vessel metal temperatures do exceed 370°C (7006F) for both pressurized and depressurized events. The time duration of the events is 447 between 250 and 400 hours. Since the reactor is cooled by passive heat transfer, with no active decay heat removal systems, reactor vessel temperatures increase as heat is removed from the core by radiation and conduction. Allowable stresses for the selected pressure vessel steels are limited to 370'C (700°F) by Section III of the ASME Boiler and Pressure Vessel Code. Procedures for predicting irradiation induced changes in material properties of pressure vessel steels are based on an extensive database of LWR surveillance data (12). Since the normal operating temperature of the MHTGR is lower than LWR operating temperatures, irradiation damage cannot be accurately predicted for the MHTGR vessel using the same methods. Corrections to damage predictions must be made to account for differences in the irradiation environments between the MHTGR vessel and LWR conditions. A design approach was developed to address MHTGR operating conditions where metal temperatures exceed the currently allowable temperature limits and to account for potential differences in the response of the pressure vessel materials to the MHTGR irradiation environment. The following discussions describe how each of these issues will be satisfied in the final design of the MHTGR vessel system.

DESIGN APPROACH FOR ELEVATED TEMPERATURE SERVICE SA533 and SA508 pressure vessel steels commonly used in the fabrication of LWR vessels were selected for the MHTGR reactor vessel. Preliminary design of the MHTGR reactor vessel identified low probability, off-normal, Level C and D duty cycle events where metal temperatures exceeded 370*C (700°F), the maximum temperature allowed by Section III of_the ASjjE Codejfor the selected materials. Section III designs are based only on the time independent strength properTTeT'oT'materials. Since time dependent material behavior must be considered at elevated temperatures, Section III does not provide allowable stresses beyond 370°C (700°F) for these materials. There is no provision in Section III for deriving and using allowable stresses in the elevated temperature regime for SA533 and SA508 steels.

DESIGN ALTERNATIVES The design alternatives to account for the elevated temperature service MHTGR were as follows: 1) Generate a sufficient materials property data base to qualify the reactor pressure vessel materials, SA533 Grade B Class 1 and SA508 Classes 2 and 3 for Code Case N47. 2) Select a Cojjg^Case _N47 approved material for the vessel system in lieu of the typicälsteels useä* in LWR vessels. 3) Incorporate active, high reliability reactor cooling systems which will insure that vessel metal temperatures will not exceed 370°C (700°F). 4) Develop a special Code Case for SA533 Grade B Class 1 and SA508 Classes 2 and 3 pressure vessel materials to allow limited use of these materials above 370°C (700°F) for the MHTGR reactor pressure vessel. 448

DESIGN APPROACH The four design alternatives were evaluated to determine the best approach for successfully completing the design of the MHTGR vessel system. Technical, cost and schedular considerations were made in evaluating the alternatives. Qualifying SA533 and SA508 steels for Code Case N47 would require the establishment of a very large materials data base including: time-dependent allowable stress intensities, creep strain rates, isochronous stress-strain curves, fatigue strain range curves, and creep-fatigue damage envelope. The required range of data would far exceed what is needed to support the HHTGR reactor vessel design conditions. Selecting one of the existing Code Case N47 materials would involve selecting a material which does not have an irradiation data base with regard to NDTT shift as a function of neutron fluence. The existing Code Case N47 materials are 2 1/4 Cr - 1 Mo, Alloy 800H, and Type 304 and Type 316 stainless steels. A wery extensive data base quantifying chemistry, temperature and fluence effects would be required. The alternative to incorporate active, high reliability cooling systems into the MHTGR would defeat the primary design objective of achieving a passive system with no active cooling systems required to maintain system safety and operability during all anticipated design-basis events. There would be a significant loss in the simplicity of the design and operation of the MHTGR and a substantial increase in cost of the system. Based on the deficiencies of the other design alternatives, a decision was made to develop and obtain ASME Code approval for a special Code Case to allow use of SA533 and SA508 steels at temperatures above 370°C (700°F) for limited times. The selected design approach is to apply the highly developed and reliable LWR technology to the MHTGR vessel system. This includes using SA533 and SA508 steels with their extensive irradiation data bases which need only an incremental expansion to account for MHTGR irradiation conditions. The ASME Code design procedures, supporting data and allowable stresses are applicable to virtually the entire duty cycle with the exception of a small number of Level C and D events. The selected design approach provides for establishing the necessary allowable stresses and supplementary design rules to be used in conjunction with existing ASME Code Section III data and rules for duty cycle events where temperatures are greater than 37O°C (700°F).

ASME CODE INQUIRY An Inquiry was prepared and submitted to the ASME Code Committees for consideration. The general form of the inquiry was as follows: May SA533 Grade B, Class 1 plate, SA508 Class 3 forgings and their weldments be used in Section III, Division 1, Class 1 construction of i MHTGR reactor vessel, where metal temperatures during Level C and D events are greater than 370°C (700°F) but less than 538°C {1000'F) for limited times of exposure not to exceed 1000 hours? 449

The scope of the inquiry envelopes the conditions for the anticipated Level C and D events in the MHTGR duty cycle. The elevated temperature events are classified as: 1) A pressurized conduction cooldown with a maximum metal temperature of 410°C (770°F), with a time duration above 370T (7OCTF) of approximately 250 hours, and 2) A depressurized conduction cooldown with a maximum metal temperature of 470'C (880°F), with a time duration above 370°C (700°F) of approximately 400 hours.

Two events of each type have been included in the MHTGR duty cycle. The number of events was restricted to avoid elevated temperature fatigue concerns, while maintaining the flexibility within the duty cycle to return the MHTGR plant to service following one of these events. If two events of either type were to occur the duty cycle for that event would be reduced to zero, therefore the maximum number of temperature cycles above 370°C (700°F) is effectively limited to a total of three. The maximum number of events permitting the return of the vessel system to operation is only two, since the vessel can only be returned to operation when the duty cycle provides for the possible reoccurrence of either event. A positive response by the ASME Code Committee will result in the establishment of a special Code Case for the design of the MHTGR vessel. The Code Case would provide the Code Committee's response to the inquiry in terms of design rules which would govern the limited elevated temperature service of the MHTGR vessel and a set of time-temperature related allowable stresses which would limit the vessel stresses at temperatures above 370°C (700°F). The Inquiry is currently under review by several committees of the ASME Code and is being evaluated in terms of both design and materials requirements. The MHTGR design is intended to adhere to the rules of Section III of the ASME Code for all other design and operating conditions. The design for the limited elevated-temperature service would be based on the supplemental rules and requirements of the Code Case. The applicable features of Code Case N47 for elevated-temperature design will be included in the new Code Case for the MHTGR vessel. One of the key features of the Code Case will be to establish limits for time-dependent strains. The other main task of obtaining approval for the Code Case is to establish the allowable stresses for the service between 370°C (700°F) and 538°C (1000*F).

RELATED TEST PROGRAM A test program was developed to supply the materials property information required to develop allowable stresses and design rules for the special Code Case. Testing is being performed by Oak Ridge National Laboratories (ORNL). The test matrix includes three heats of SA533 plate, one SA508 forging and three weldments (two submerged arc welds and one shielded metal arc weld). 450

The testing to be performed on these materials is as follows: 1) Tensile testing from 370°C (70CTF) to 593°C (HOOT), 2) Creep testing at several different stress levels from 427°C (800°F) to 593°C (1100'F) for test times up to 2000 hours, and 3) Evaluation of the potential for thermal embrittlement of these materials by aging base and weld metal test specimens at (850°F) and (950°F) for 2000 hours and comparing Charpy impact test results for the aged and unaged material.

The results of the testing will be provided to the ASME Code Committee for use in establishing stress allowables. Table 1 shows the criteria used to establish stress allowables provided in Code Case N47. It is anticipated that the same criteria will be used to establish allowable stresses for the MHTGR Code Case. Figure 2 shows the Larson-Miller parameter for the preliminary test data combined with previously published data (3-7) for 1% creep strain in 1000 hours as a function of stress. The minimum curve for the test data defines the time-dependent stress behavior. The other curve on Figure 2 is the time-independent stress behavior based on tensile properties as indicated in Table 1. Figures 3 and 4 show similar plots for the other time-dependent criteria, 80% of stress to onset of tertiary creep and 2/3 of the stress to rupture, as shown in Table 1. When all of the testing is completed these curves will include all test results and will become the basis for establishing allowable stresses for the Code Case. Temper embrittlement of pressure vessel steels was raised as a potential concern for the time-temperature range of the Code Inquiry. Some studies have shown that temper embrittlement can occur in pressure vessel steels (8-11). In general, the materials which show significant changes in Charpy impact transition behavior due to temper embrittlement effects are high in residual element content and are not typical of modern production pressure vessel steels and weldments. Thermal aging of pressure vessel steels has been observed to have little effect on the tensile properties or upper shelf toughness (8,10,11). The most significant effects observed have been in NDTT or FATT (Fracture Appearance Transition Temperature) shifts. The effects are also more pronounced for weld and heat affected zone (HAZ) materials than for base metals. The thermal aging study included in the current test program is intended to evaluate the potential for temper embrittlement and measure the magnitude of any changes in Charpy impact properties.

DESIGN APPROACH FOR IRRADIATION EMBRITTLEMENT Pressure vessel integrity of operating LWR reactor vessels is based on a large data base of material property evaluations, fracture mechanics evaluations, and engineering experience in vessel design and fabrication. The prediction of changes in material properties as a result of neutron irradiation is based on extensive experience and research to define radiation damage mechanisms (12). The majority of the data is from and applicable to typical LWR conditions, namely irradiation temperatures in the range 277° - 305°C (530° - 580°F); total neutron fluxes in the range 10"10 - 10~13 (n/cm"2); and a neutron energy spectrum of 10% fast, 25% epithermal and 65% thermal. 451

Analysis of the data base has shown an influence of the irradiation temperature on the irradiation induced shift in the nil-ductility transition temperature (NDTT) of pressure vessel steels. However, limited data are available outside the LWR temperature range to accurately quantify the effects of lower irradiation temperatures anticipated for the MHT6R. In addition, there is insufficient data to define the effects of lower flux level and different neutron energy spectrum in the HHTGR on the irradiation response of the materials.

DESIGN ALTERNATIVES The HHTGR designers have only two alternatives. The procedures in Regulatory Guide 1.99, Rev. 2 (12) may be used for the prediction of neutron irradiation damage to the reactor vessel when credible surveillance data from the reactor in question are not available. The prediction procedures are applicable to the selected vessel materials, but for irradiation temperatures in the range of 274'C to 310°C (525° to 590°F) and fast fluence in the range 10*17 - 10 20 (n/cm~2). Corrections for operations outside these ranges should be justified by reference to actual data. The alternatives were to: 1) Generate the experimental data necessary to make the correlations between prediction techniques based on LWR experience and MHTGR operating irradiation environment, or 2) Design a much thicker vessel to keep operating stresses very low, providing adequate margin to brittle failure.

DESIGN APPROACH The second alternative would require a much more extensive material surveillance program to monitor the vessel materials during operation. Furthermore, to reduce the uncertainty with regard to pressurized thermal shock, additional plant protection control modes might be required. This would defeat the objective for passive control of the MHTGR concept. The proposed approach is to obtain sufficient data for the expected MHTGR operating temperature range and bounding fluence. The resulting data will be correlated with an equivalent fast fluence. This will provide the designer with a conservative procedure to estimate vessel irradiation damage. In addition, the design effort will be directed at minimizing the neutron flux in all vessel regions.

DESIGN CONSERVATISMS In order to proceed.with the MHTGR design, NDTT shift predictions have been calculated for the projected vessel fluences by applying additional conservatisms to the methodology of Ref. 12. A comparison of the irradiation environments of LWR and MHTGR vessels is given in Table 2. The effect of irradiation temperature on NDTT shift is shown in Figure 5. Irradiation 452 damage tends to saturate at lower temperatures but are about twice as large as the shift at 290°C (550'F) (13). Based on the existing data a conservative temperature correction factor of 2.25 is applied to the shift predictions to account for the irradiation temperature differences. The neutron energy spectrum of the MHTGR contains higher percentages of epithermal and thermal neutrons in comparison to LWR spectrums as shown in Table 2. Because the epithermal neutrons also cause damage, a higher shift is expected to result in an MHTGR vessel than in a LWR vessel for the same fast fluence. Data for material irradiated in both LWR and graphite moderated test reactors at approximately 12TC (250°F) resulted in the following shifts (14): NDTT Shifts in A302B Plate Material Fluence Light Water Graphite Moderated (n/cm"2) 0.5 x 10"19 94°C (170°F) 114°C (205°F) 0.75 x 10"19 114°C (205°F) 133°C (240°F) Shift predictions of Ref. 12 are correlated with fast neutron fluence. To specifically account for the effect of neutron spectrum, the equivalent fast fluence approach is used with the procedures of Ref. 12. To account for differences in neutron energy spectrum, the total neutron fluence of the MHTGR is converted to an equivalent fast fluence by applying weighting factors to each of the neutron energy ranges. The equivalent fast fluence, weighted for epithermal and thermal neutron damage, is then used in the Ref. 12 procedures for NDTT shift prediction. Table 3 shows a typical calculation of the equivalent fast fluence and NDTT shift for the MHTGR vessel. The calculated NDTT shift for the MHTGR vessel using the correction factors is several times the shift predicted by Ref. 12 based on the estimated fast fluence.

RELATED TEST PROGRAM The corrected values for predicted NDTT shift shown in Table 3 represent an extrapolation from the LWR database to MHTGR irradiation conditions. A test program has been initiated to provide confirmatory data which will demonstrate that adequate conservatism has been added to the NDTT shift predictions for the MHTGR vessel. The test program consists of a matrix of irradiation capsules to be exposed in a test reactor. The irradiations will include several materials including SA533 plate, SA508 forging and weld metals that will be irradiated at temperatures from 120°C (250°F) to 288°C (550°F) and fluences from 10*16 to 10*18 (n/cm"2). Shielding of some capsules will be employed to achieve a neutron spectrum similar to that of the MHTGR. The test program has been designed to provide quantitative data on the individual and synergistic effects of material product form and chemistry, temperature, fluence and neutron energy spectrum. The adequacy of the correction factors applied to the existing shift prediction methodology will be demonstrated.The results of the test program will provide a direct correlation of the conservatism of NDTT shift predictions to experimentally measured shifts. 453

SURVEILLANCE PROGRAM In addition to conservative NDTT shift predictions that have been verified by comparison to experimental data, the HHTGR design will include a reactor vessel materials surveillance program. The surveillance program will monitor changes in the vessel materials over the operating life of the MHTGR. The program will be designed to be similar to LWR programs, incorporating the actual materials used in fabricating the reactor vessel. Surveillance capsules will be removed periodically during the life of the MHTGR to monitor actual response of the vessel materials to irradiation.

SUMMARY & CURRENT STATUS MHTGR ELEVATED TEMPERATURE SERVICE The design for limited elevated-temperature service of the MHTGR reactor vessel will be governed by the rules and allowable stresses of a special Code Case to the ASME Code. An Inquiry has been submitted to the ASME Code Committees for review and approval. The Committees are currently reviewing the design considerations necessary to ensure the safety and reliability of the MHTGR vessel at temperatures above 37O°C (700T). Elevated-temperature time-dependent materials properties are being generated in an ongoing test program at ORNL. The elevated-temperature materials properities will be used to establish allowable stresses for MHTGR design during the Level C and D events at temperatures above 37O°C (700°F) for limited times. The test program will be completed in approximately 1 year. It is anticipated that formal approval of the Inquiry in the form of a new Code Case will be obtained shortly after completion of the test program.

MHTGR IRRADIATION EMBRITTLEMENT The MHTGR design has recognized that the current irradiation data base and prediction techniques do not adequately define the effects of the MHTGR irradiation environment on NDTT shift of vessel materials. The MHTGR design accounts for possible differences in NOTT shift behavior by: 1) Minimizing neutron flux to the vessel, particularly at critical locations of the vessel. 2) Applying conservative correction factors to NDTT shift predictions for temperature and neutron energy spectrum effects. Multiplication factors for irradiation temperature and effective fast neutron fluence have been used in NDTT shift predicitions to provide extra margin. 3) Obtaining experimental data from a test program, to quantify the effects of the MHTGR irradiation environment on NDTT shift. The results of the test program will be used to update the analytical model for damage prediction in the MHTGR vessel. 4) Incorporating a vessel materials surveillance program for the MHTGR reactor vessel to monitor accumulated irradiation damage over the operating life of the MHTGR and to validate the design basis calculations. 454

Several of the capsule irradiations have been completed from the test program Evaluation of the capsule test specimens from these initial capsules should begin within the next several months. Since the testing can be performed quite quickly it was decided to accumulate several sets of specimens from the capsules to reduce the hot cell time required for testing the irradiated specimens. This test program should be completed in 3 to 3 1/2 years.

ACKNOWLEDGEMENT This paper summarizes work performed under DOE Contract DE-ACO5-84OR21400 and it is published with the permission of DOE.

REFERENCES 1) S. R. Penfield, "Outlook for Nuclear Power in the 1990's and Beyond", Proceedings of the 57th General Meeting of the National Board of Boiler and Pressure Vessel Inspectors, Las Vegas, Nevada, May 2-6,1988. 2) Criteria for Design of Elevated Temperature Class 1 Components in Section III. Division 1. of the ASME Boiler and Pressure Vessel Code, American Society of Mechanical Engineers, New York, New York, May 1976. 3) Evaluations of the elevated Temoerautre Tensile and Creep-Rupture Properties of C-Mo. Mn-Mo and Mn-Mo-Ni Steels. ASTM Data Series Publication DS-47, American Society for Testing and Materials, Philadelphia, Pennsylvania, November 1971. 4) A. W. Pense and R. D. Stout, "Characterization of Heat Treated Pressure Vessel Steels for Elevated Temperature Service," Symposium on Heat-Treated Steels for Elevated Temperature Service, American Society of Mechanical Engineers, New York, New York, 1966, pp. 8-26. !5) E. B. Norris and R. D. Wylie, "Analysis of Data from Symposium on Heat-Treated Steels for Elevated Temperature Service," The Metals Properties Council, American Society for Mechanical Engineers, New York, New York, 1966, pp. 1-66. 6) J. J. OeBarbadillo, A. W. Pense and R. D. Stout, "The Creep-Rupture Properties of Pressure Vessel Steels - Part II," Welding Research Supplement, Welding Journal, August 1966, pp. 357s-367s. 7) G. B. Reddy and 0. J. Ayres, "High Temperature Elastic-Plastic and Creep Properties for SA533 Grade B Class 1 and SA508 Materials," EPRI NP-2763, Electric Power Research Institute, Palo Alto, California, December, 1982. 8) S. G. Druce, et al., "Effect of Ageing on Properties of Pressure Vessel Steels," Acta Metallurgica, Vol. 34, No. 4, pp. 641-652, 1986. 455

9) S. G. Druce and B. C. Edwards, "On the Temper Embrittlement of Manganese-Molybdenum-Nickel Steels," Nuclear Technology, Vol. 55, pp. 487-498, November 1981. 10) W. A. Logsdon, "The Influence of Long-Time Stress Relief Treatments on the Dynamic Fracture Toughness Properties of ASME SA508 Cl. 2a and ASME SA533 Gr. B Cl. 2 Pressure Vessel Steels," Journal of Materials for Energy Systems, Vol.3, No. 4, pp. 39-50, March 1982. 11) R. A. Swift and J. A. Gulya, "Temper Embrittlement of Pressure Vessel Steels," Welding Research Supplement, Welding Journal, pp. 57s-68s, February 1973. 12) Regulatory Guide 1.99, Revision 2, Radiation Embrittlement of Reactor Vessel Materials, U. S. Nuclear Regulatory Commission, Washington, D. C, May 1988. 13) L. E. Steele, Neutron Irradiation Embrittlement of Reactor Pressure Vessel Steels, International Atomic Energy Agency Technical Report No. 163, 1975, p. 123. 14) C. Z. Serpan and L. E. Steele, "Damaging Neutron Exposure Criteria for Evaluating the Embrittlement of Reactor Pressure Vessel Steels in Different Neutron Spectra," NRL Report 6415, Naval Research Laboratory, Washington, 0. C, July 28, 1966. 456

TABLE 1

CRITERIA FOR ESTABLISHING ALLOWABLE STRESSES FOR ELEVATED TEMPERATURE SERVICE

TIME-INDEPENDENT ALLOWABLE STRESSES

The symbol Sm is used for the time-independent allowable stress value. These values are based on tensile and yield strengths of the material. The allowable stress intensity value S for ferritic steels is defined as the least of the following four values: 1) One-third (1/3) of the specified minimum tensile strength at ambient temperature; 2) One-third (1/3) of the short-term tensile strength at temperature, as defined in Appendix A of Reference 2; 3) Two-thirds (2/3) of the specified minimum yield strength at ambient temperature; or 4) Two-thirds (2/3) of the short-term 0.2% offset yield strength at temperature, as defined in Appendix A of Reference 2.

TIME-DEPENDENT ALLOWABLE STRESSES

The symbol Sf is used for the basic time and temperature dependent allowable stress value for load-controlled stresses. S, values are defined as the least of the three quantities: 1) Two-Thirds (2/3) of the minimum stress to cause rupture in time t; 2) 80% of the minimum stress to cause the onset of tertiary creep in time t; and 3) the minimum stress to produce 1% total strain in time t. 457

TABLE 2 COMPARISON OF MHTGR AND LWR IRRADIATION ENVIRONMENTS

MHTGR LWR REACTOR VESSEL REACTOR VESSEL

TEMPERATURE 121* - 204°C 288° - 304°C (250° - 400°F) (550° - 580°F) FAST FLUX 107 - 108 1010 - 1011 (n/cm"2 sec, E > 1 MeV) FAST FLUENCE 1015 - 1016 1019 - 1020 (n/cm"2, E > 1 MeV) NEUTRON SPECTRUM FAST (E > 1 MeV) 4% 10% EPITHERMAL (0.4 < E < 1 MeV) 28% 25% THERMAL (E < 0.4 MeV) 68% 65% 458

TABLE 3 ESTIMATION OF EQUIVALENT FAST FLUENCE AND NDTT SHIFTS FOR THE MHTGR REACTOR VESSEL

EQUIVALENT FAST FLUENCE Fluence Damage Equivalent Neutron Enerqy , E (n/cnf 2) Factor Fast Fluence E > 0.9 MeV 5.7xlO15 1.0 5.7xlO15 0..1 MeV < E < 0.9 MeV 4.0xl016 0.5 2.0xl016 3.,05 eV < E < 0.1 MeV 9.6xlO16 0.0025 2.4xlO14 E < 3.05 eV 2.7xlO15 0.07 1.9xlO14

Total Equivalent Fast Fluence (EFF) = 2.6xlO16

PREDICTED NDTT SHIFT Prodicted NDTT Temperature Shift per Ref. 12 Correction (x2.25) Base Weld Base Weld Metal Metal Metal Metal Fast Fluence 5.7xlO15 2 2 5 5 Total EFF 2.6X1016 5 5 12 12 459

CORE ELECTRIC BARREL MOTOR REACTOR DRIVE VESSEL MAIN CROSS ACTIVE CIRCULATOR CORE DUCT

STEAM GENERATOR

STEAM GENERATOR VESSEL

Figure 1 - General Arrangement of HHTGR Reactor Vessel, Crossduct and Steam Generator Vessel. 460

LARSON-MILLER PARAMETER PLOT FOR

STRESS TO I'A CREEP STRAIN IN 1080 HOURS

1000 . . ,

T^ ?x. Vis 7K TIME INDEPENIENT, Sft, ...V f i i i r LJ LJ"L

Q_ Viz TIME DEPENDENT, S 100 \ t

CO Vb£ co .. \ tu ...\ CO \ \

Cl ksi = 6.5S85 MPa) 10 i i i 1 i i i 1 i i i

12 14 16 18 22

P = T * (29 + LOG t)

LARSON-MILLER PARAMETER PLOT FOR

88/i OF STRESS TO ONSET OF TERTIARY CREEP

. . . j ... | ... 1 ' 1886

'r^äi. ! X^E ; \ ITIME INDEPENEDENT, S^ •• n"ri"i ; U U I *™***HM3' a. 2M/ TIME DEPENDENT, St 188

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<1 ksi = 6.9"85 MPa) : 18 j . . . i . . . i . . . 1 i t I

12 14 16 18 28 22

CX 1098) p = T x (28 + LOG t)

Figure 3 - Larson-Miller Parameter Plot for 80% of the Stress to the Onset of Tertiary Creep. 462

LARSON-MILLER PARAMETER PLOT FOR

2/3 OF CREEP-RUPTURE STRESS

. i . 1000 . . .

* k TIME INDEPEh DENT, sJ^Vi rj n r LJ LJ L

Q. \ 3S X TIME DEPENDENT, S 100 \. (S) ^

<1 ksi = 6.585 MPa) 10 | . . . . , ,

12 14 13 20 22

CX 1090) p = T » C20 + LOG t)

Figure 4 - Larson-Miller Parameter Plot for 2/3 of the stress to rupture. 463

\ 300 (57 ui A3O2-B \ LJ REFERENCE STEEL a \ u 19 2 2 $ =3.1-3.5 X I0 (n/cm >lMeV) Ui (ff = 68mbf*Mr , FISSION) tr <3 tr 111 £ 200 2 \\ Z o

\ 100

\\ O

1 1 1 1 100 200 300 400 500 600 700 eoo _! 100 200 300 400 IRRADIATION TEMPERATURE

Figure 5 - Effect of Irradiation Temperature on Transition Temperature Increase for an A302-B Reference Steel (13). 465

Section V:

Structural graphite components 467

MATERIALS BEHAVIOUR AND DESIGN VALUES

G. Haag Kernforschungsanlage Julien Institut für Reaktorwerkstoffe

Abstract: In structural design criteria for High Temperature Reac- tors graphite has to be regarded as a material of its own which is significantly different from other construction materials. To guarantee the quality of graphitic materials, it is necessary to verify the properties of raw materials, to survey the production process and to control the final product properties. Designing graphitic reactor components one has to take into account problems that may arise from corrosion and from fast neutron radiation effects. Consequently, physical properties such as dimensional stability, Young's modulus, tensile strength, thermal conductivity, thermal expansivity and coefficient of irradiation induced creep have to be observed as a function of neutron fluence in irradiation experiments.

INTRODUCTION

With respect to structural design criteria for future High Tempe- rature Reactors it has to be considered that graphite is significantly different from many other construction materials: - In terms of the mechanical behaviour of metals graphite has to be regarded as brittle, but its stress-strain behaviour is hysteresis type. Upon applying stress graphite strains non- linearly. If the maximum stress is remaining significantly below the ultimate strength, the material relaxes leaving a residual strain. - In an inert atmosphere, graphite may be exposed to very high temperatures. - Within the HTR relevant temperature range, its strength and toughness increase with increasing temperature. - In general, graphite exhibits structural anisotropy. 468

The whole core structure of High Temperature Reactors is made of graphite (fig- 1). The most serious problems with respect to fast neutron radiation damage arise for those parts being close to the pebble bed, that is for the top, the inner side and the bottom reflector.

Inner side reflector

Fig. 1: Core structure of the HTR-500 reactor [3]

MATERIAL PROPERTIES

Design Data

The design of High Temperature Reactors and the construction of graphite components are based on the knowledge of particular material data. Table 1 shows five different categories for the application of graphite data and the properties belonging to them 469

Application Properties

neutron physics density neutron-capture cross-section

thermodynamics thermal conductivity ) specific heat ) density

construction density thermal expansivity ) strength data oxidation behaviour

strength analysis Young's modulus Poisson's ratio thermal expansivity strength data oxidation behaviour

radiology contents of lithium, europium and cobalt temperature dependent Table 1

Quality Control

The determination of material properties of reactor graphites is based on DIN standard testing procedures being worked out during the last 15 years and differing in some details from ASTM and other national standards. (DIN is the German Institute for Stan- dardization) .

To guarantee the quality of graphitic materials, it is necessary to verify the properties of raw materials, to survey the production process and to control the final product properties. Raw materials such as filler coke and pitch binder provided by other manufacturers have to be checked for their relevant properties. After each production step such as forming or baking, appropriate 470 non-destructive control procedures are applied. After graphiti- zation, material properties are examined in detail using random sampling and including destructive testing. The final check of machined products is made by dimensional and visual examination.

The whole procedure of controlling the raw material properties, the production steps and the final products, the way of how to take samples as well as the number of samples and finally, how to make all this information available is put down in the design criteria.

General Treatment of Corrosion Problems

The use of helium as coolant in HTRs decreases the corrosion problems to a minimum. However, corrosion problems may arise due

1001 Degree of Corrosion (w/o)

Depth X

Fig. 2a: Corrosion profile of a graphite component

Depth X

Fig. 2b: Strength profile of a graphite component 471

to oxygen and water vapour penetrating into the coolant either through leakages during normal operation or in case of an accident. Corrosion rates are determined experimentally for relevant temperatures and gas compositions, allowing to calculate corrosion under reactor conditions. The corroding gas mixture diffuses through the pores of the graphite which leads to a corrosion profile decreasing from the outer surface to the centre of the component (fig. 2a). Since corrosion can be related to a loss of strength (fig. 3), the corrosion profile can be converted into a loss-of-strength-profile (fig- 2b), where the loss of strength due to corrosion is effectively limited to the outer parts of graphite components. Therefore, corrosion effects can be taken into account by designing components with extra wall thickness. 1

10 20 Degree of Corrosion (w/o)

Fig. 3: Correlation between degree of corrosion and loss of strength of a nuclear graphite

IRRADIATION BEHAVIOUR

Irradiation induced stresses resulting from Wigner effect (i.e. fast neutron radiation damage in the graphite crystal stucture)are most important for the selection of graphite grades for application at high neutron fluences. The graphite lattice is fundamentally damaged in a reactor environment by collision of high-energy 472 neutrons with carbon atoms in the lattice (fig. 4). The carbon atoms are displaced to interstitial positions, leaving behind vacant sites in the layer planes. Some portion of the vacancies and interstitials are immediately annihilated by recombination, but those remaining may concentrate depending upon the conditions of neutron fluence and irradiation temperature and form larger clusters. At low temperatures, the vacancies remaining in the graphite lattice are immobile and remain as point defects [1].

Fig. 4: Displacement of carbon atoms from the graphite lattice by fast neutron radiation

Most of the German irradiation testing has been done in the High Flux Reactor at Petten which provides excellent conditions. Capsules being equipped with thermocouples and containing about 90 to 120 specimens are kept at constant temperature by gas mixture and by shifting the capsule vertically as fuel is burned and control-rods are raised. Neutron fluence determination is based on flux monitor evaluation.

Dimensional stability

The easiest way to follow fast neutron irradiation damage is to look at the linear dimensional changes of regularly shaped graphite specimens, which may exhibit shrinkages of a few percent or expansions of up to 15 or 20 percent. Three criteria mainly determine the usefullness of graphite in reactor design: (i) the maximum contraction which for instance influences the stability of a graphite stack, 473

(ii) the fluence where net dimensional expansion starts at high fluence, and (iii) the difference between the changes in the two main crystal- lographic directions.

with grain orientitlon

5 19 15 20 25 30 35 40 Neutron Fluence (1021/cm2 (EDN))

5 10 !5 20 25 30 35 Neutron Fluence (IB21/cm2 (EDN))

Fig. 5: Dimensional behaviour of reactor graphite ATR-2E with grain (a) and across grain (b) orientation under fast neutron irradiation Fig. 5 demonstrates the linear dimensional behaviour of ATR-2E reactor graphite at different temperatures with grain (a) and across grain (b). 474

Young's Modulus and Strength

Just as linear dimensions, other physical properties are also influenced by irradiation induced damages. Young's modulus is one of the properties of major interest to reactor component designers. In case of iterative irradiation testing, Young's modulus has to be determined as dynamic Young's modulus derived from sound velocity or sonic resonance measurements. With increasing neutron fluence it exhibits a characteristic behaviour (fig. 6) showing a rapid increase at very low fluences followed by a more or less pronounced plateau and a further increase reaching a maximum value. Towards the end of life of graphite as a solid, Youngs modulus goes down to its original value and then tends to zero.

12 3 4 Fast Neutron Fluence/1022 cm"2 (EDN)

Fig. 6: Irradiation behaviour of Young's modulus as a function of fast neutron fluence for different irradiation temperatures (ATR-2E reactor graphite)

Obviously, the life history of reactor graphite is closely related to the development of its strength. There is evidence that while the modulus is increasing with neutron fluence, the tensile strength is also increased following a square root relationship 475

where 0**o and

Irradiation Induced Creep

From the stress-strain behaviour of graphite it is known that a strain of more than about 0.5 percent leads to destruction. How- ever, irradiation induced changes of 10%, 20% or even more are observable. Consequently, an irradiation induced mechanism must exist to reduce internally developed stresses. It is called "irradiation induced creep" which is an important phenomenon at temperatures as low as 300 °C, whereas thermally-induced creep is of significance only at temperatures in excess of 2000 °C [1].

Obviously, creep data are indispensable for the assessment of lifetimes of reactor components exposed to fast neutron flux. To perform a creep experiment means to irradiate graphite specimens under tensile or compressive load and to observe the dimensional changes being different from those of unloaded specimens.

These experiments are expensive and lengthy. For cost reasons it would be desirable to observe another physical property with strong correlation to creep coefficient but easily to be measured. Fig. 7 shows that there is a general behaviour of creep coefficient with fast neutron fluence, no matter which irradiation temperature or if tensile or compressive load is applied. Qualita- tively, it is clear that the ability to relax stresses goes down with increasing Young's modulus and strength respectively and vice versa. However, the quantitative approach to. describe the beha- 476 viour of creep coefficient in terms of Young's modulus is not quite satisfactory, particularly at higher fluences. This leads to attempts to correlate the fluence dependence of creep with that of volume change. Future experiments must show if preliminary results [2] will be confirmed.

^3.5 -3.5

o ~2ß 'a a. 52J> -2.5

•2.0ö yj u. o MPRESSIVE 500 *C u_ u. UJ UJ 13 o -1.58 o 0_ UJ 0. UJ ÜJ .A Ul ' J cc UJ o UJ m =J0.5 -0.5 w en z: 2: 2 LJ FLUENCE x / 10 ncm" EDN o 0 o 6 8 10 12 \L 16 18 20 22 Creep Coefficients of ATR-2E EURATOM JRC under 5 MPa Load Patten

Fig. 7: Irradiation behaviour of creep coefficients of ATR-2E reactor graphite [4]

Thermal Properties

According to table 1, the irradiation behaviour of thermal expansivity and thermal conductivity have also been investigated.

Fig. 8 gives the thermal expansivity results for a typical reactor graphite. There is some evidence that for higher neutron fluences there is no major change in the linear thermal expansion coefficient. Detailed investigation of the high fluence behaviour is left to further irradiation experiments. 477

Fast Neutron Fluence/iO22cm~2 EDN

Fig. 8: Irradiation behaviour of linear thermal expansion coefficient for different irradiation temperatures (ATR-2E)

2.5 Fast Neutron Fluence (1022/cm2(EDN))

Fig. 9: Irradiation behaviour of thermal conductivity for different irradiation temperatures 478

The relative changes of thermal conductivity shown in fig. 9 so far have turned out to be rather material independent. However, at very high fluences a further decrease is supposed to be correlated to the dimensional changes and therefore to some extent should be material dependent.

Conclusion

Since the irradiation behaviour of reactor graphite cannot yet be predicted from preirradiation properties and structural features, it is still important to irradiate graphites at temperatures and fluences that include and some times extend beyond the design conditions of current reactor projects [1].

REFERENCES

[1] G.B. Engle, W.P. Eatherly: High Temperatures-High Pressures, 4 (1972), 119-158

[2] C.R. Kennedy, M. Cundy, G. Kleist: Proc. Carbon '88, Newcastle upon Tyne, 18-23 September 1988, pp. 443-445

[3] J. Scheming, W. Theymann: Atomkernenergie-Kerntechnik, 43 (1983), 248

[4] M.R. Cundy, G. Kleist, D. Mindermann: Extended Abstracs Carbon Conf. CARBONE 84, Bordeaux 1984, p. 406 479

Discussion of the presentation:

Materials behaviour and design values

Trumpfheller, Essen, FRG: J Can you say something about the inspection of graphite blocks in the HTR?

Theymann, HRB Mannheim, FRG: The top reflector of the AVR was inspected some years ago. We could not see any changes on the inner side of the reflector blocks. This result was better than expected from some stress calculations done before. The THTR-reflector cannot be inspected on the core side, only some parts of the outer side are designed for inspection. The AVR test operation programme is continuing a re-analysis of 1 or 2 graphite blocks of the side reflector. The graphite components are designed for the whole life time of the reactor.

Nickel, KFA-Jiilich, FRG: Please give some information to the auditorium what is about the R&D work which is still required for the projects Modular Reactor and the reactor HTR-500 in correlation with the design data which will be used.

Haag, KFA-Jülich, FRG: One of the major problems is that the graphite specimens tested to the highest neutron fluences had been taken from preproduction lots whereas the production lots had been manufactured as the last step of an optimization process. Therefore, the irradiation testing of the final qualities has not yet been completed. In the meantime, problems have come up concerning the long-term availability of the raw materials which is not unusual in material production. In general, the use of new raw materials at least has to be followed by a verification of the final graphite properties including the irradiation behaviour. 480

Design Methods and Criteria for Graphite Components

A. Schmidt HRB-Mannheim

1. Introduction

The design criteria developed in the Federal Republic of Germany in recent years for the graphite internals of an HTR are mainly based on the experience gained during the construction and licensing of the THTR 300 and on the know-how obtained from a graphite research programme. The THTR experience.had especially influenced the general design criteria, while the results from the research programme offered a special procedure for the evaluation of the stresses in graphite structures.

The present contribution furnishes a survey of the structural design features of the graphite components with special emphasis on the design method and the stress limits to be observed.

2. HTR 500 Graphite Internals For a better understanding of the analysis given below, the structural design of the graphite internals is first illustrated by the example of the HTR 500 (Figure 1).

The reflector and the core support structure jointly form the container of the pebble bed core and serve at the same time as a flow guidance structure for the coolant gas. The fuel elements added to the core through refueling pipes pass downward through the core and are then discharged through three fuel element discharge pipes arranged in the core support structure.

The core structure accommodates mechanical loads (dead weight, silo pressure, forces due to insertion of control rods, pressure differences) and seismic loads transferring them into the thermal shields either directly or via support elements.

The coolant gas flows downward through the top reflector, the core and the core support structure, passes through the hot gas plenum and is forwarded through penetrations in the side reflector to the heat exchangers. From the nuclear physics aspect, the graphite internals have the additional, effect of a reflector and shielding: The neutron losses from the core are limited and adjacent metallic structures are protected against excessive neutron fluxes. •481

The core support structure consists of the bottom layers, the cylindrical columns and the core bottom. The latter is composed of hexagonal graphite columns, its surface is formed conically towards the three fuel element discharge pipes.

The side reflector consists of an inner and an outer cylindrical wall. The outer, wall is horizontally prestressed by spring packs which are supported by the metallic thermal side shield. Above the bottom reflector both cylindrical walls are vertically supported by rocking pillars allowing restricted horizontal movement to compensate differential thermal expansion of the bottom and side reflector.

The top reflector, composed of hexagonal columns as the bot-' torn reflector, is vertically subdivided into three layers. Three columns each form a reflector unit and are suspended from the thermal top shield by an anchoring rod. The top reflector is not prestressed horizontally.

The individual elements of the core are positioned relative to one another by dowel and key systems. In the range of high neutron fluence the reflector blocks are provided with slots and are cooled on the core-side face to limit the strains induced by neutron irradiation.

Classification of Components

The components fabricated from graphite and/or carbon material can be classified under different quality assurance classes according to their tasks and functions. Class I (load-carrying-function)

This class includes the components which mainly have a load-carrying-function and whose task is to ensure the stabi' lity of the structure.

Class II (neutron physics function)

This class mainly includes the components fulfilling neutron physics tasks such as moderation and reflection of fission neutrons and shielding of load-carrying components against fast neutrons.

Class III (thermal insulating or shielding function) . '

This class includes components, especially also components from carbon material, for thermal insulation and shielding against neutrons and gamma-radiation. 482

Load and Load Levels

The graphitic internals experience loads due to:

dead weight,

weight of the pebble bed core (silo pressure),

prestressing forces from spring packs,

rod insertion forces,

pressure differences of the helium coolant,

neutron irradiation,

steady-state and transient thermal loads,

movement under transient operating conditions,

rearrangement of forces resulting from prestressing in case of differential expansions and displacements,

seismic loads, oxidation.

The loads result in stresses in the individual component parts. These stresses are classified into load levels which are different with regard to their allowable stress limits.

Load Level A

Load level A includes

normal operating conditions, upset operating conditions, testing conditions, load events with a postulated occurrance of N > 1 per service life as far as they do not result in a noticeable loading to the graphite internals. Under the conditions of load level A the functional capabili ty of the graphite components has to be maintained over the entire design service life of the reactor. 483

Load Level B Load level B includes

- . load events with a postulated occurrance of N < 1 per reactor service life, hypothetical events

as far as they result in a noticeable loading of the graphite internals.

Under the conditions of load level B the integrity of the internals has to be maintained so that safe shutdown of the reactor and safe decay heat removal is ensured.

For this reason it is permissible to take into consideration only primary stresses in this load level. It is admitted that in major areas cracks may occur in internals fabricated from graphite or carbon material.

Following the occurrance of an event of load level B, further operation of a reactor is not permitted, unless after the event the required integrity of the reactor internals can be verified by calculation or an appropriate test, possibly involving the exchange of the internals concerned. If this cannot be done, restart of the reactor operation is not permitted.

5. Stress Evaluation Usually the calculated stresses are evaluated by

^eq - fallow. bgq indicates the equivalent stress determined from the component stresses or the principal stresses, and 6"allow is the allowable stress.

5.1 Determination of the Equivalent Stress

Graphite has only a low ductility, i. e. rupture is only preceded by small inelastic deformations. Thus the application of a yield condition as a hypothesis of strength, as for metal materials in normal ranges of application, is not adequate for this material. 484

A comparison of the measured data from biaxial tests with theoretical models for low-ductility materials such as cast iron, ceramic materials, concrete and rock has shown that the criterion of the maximum strain energy is most suitable to describe the experimental results. This criterion states that the elastic energy accumulated per volume unit in the respective material element in the event of rupture is equivalent to the energy accumulated in a test specimen subjected to the uniaxial ultimate load. In this way the complex stress state is clearly correlated with the strength measured by uniaxial tensile or compression tests.

5.2 Evaluation Concept oh the Basis of Safety Factors

If a graphite component is subjected to a tensile, compressive or bending stress the following simple evaluation concept can be applied: a minimum ultimate strength is defined and safety factors referred to this minimum ultimate strength are used to obtain allowable stresses in graphite components.

This evaluation concept can generally be applied only to components which are exposed to external loads. Steady and transient temperature fields as well as fast neutron irradiation induce complex stress distributions within the components. For evaluating such stress states an evaluation concept is required which takes into consideration the special fracture behaviour of the graphite material. 485

5.3 Evaluation Concept on the Basis of Failure Probabilities

A concept suitable for the safety evaluation of stresses in graphite material is the method used for ceramic materials. In this method the allowable stress is defined as the stress for which the failure probability is below a fixed value. This procedure is also suitable to establish rules, because it is possible to determine this failure probability from measured strength data using an appropriate strength distribution.

Assuming a normal distribution for the graphite strength data, the probability can be stated to obtain a smaller value than a specified value. This probability, called failure probability in this connection, is e. g. 50 % for the mean value and 2.3 % for the value reduced by 2 standard deviations.

When plotting the failure probabilities for different graphite types against a safety factor defined as the ratio of material strength and allowable component stress, it can be shown that with a given safety factor the failure probabilities can fluctuate by several orders of magnitude, i. e. there is no relationship between a safety factor and the failure probability independent of graphite type. Therefore only values for allowable failure probabilities can be indicated in a code of rules and regulations.

The strength distribution for failure probabilities below 2 % is described more exactly by a Weibull distribution than by a normal distribution (see Figure 2). Therefore the failure probability is determined on the basis of the Weibull distribution.

By using the Weibull theory - theory of the, weakest link of a chain - it is possible to calculate the failure probability of any component exposed to load. The fact that local peak stresses are "less dangerous" than for membrane stress conditions is taken into consideration.

The evaluation procedure, described above was experimentally verified on specimens and on complex structures simulating components. These tests have shown a very close agreement between experimental results and the prognosis of the rupture probability for given loads (Figures 3 and 4, /I/). 486

The evaluation of stresses in a component using the Weibull theory does not require classification of the stresses into stress categories; only the allowable failure probabilities for the membrane stress condition have to be stated for the three quality assurance classes mentioned above and the two load levels. The strength assessment is considered to be fur- nished, if the resulting failure probability is below the allowable failure probability.

In the following the allowable failure probabilities (FP) are indicated for graphite components of the various quality assurance classes (QAC) and load levels (LL).

The values quoted for the allowable failure probabilities are based on a confidence level of 95 %. QAC LL . FP

I A 10"* I B 10";? II A 10 4 t< II B 5.10 2 III A 10"2 III B 5.10"2

Note:

QAC-I/LL-A: The failure probability covers the entire service life and must not be exceeded at any time. QAC-I/LL-B: The failure probability refers to one loading event.

QAC-II/LL-A: The lower value indicated refers to the beginning of the reactor operation

QAC-II/LL-B: The failure probability must not be exceeded during any loading event.

QAC-III/LL-A: The failure probability refers to the entire service life. QAC-III/LL-B: The failure probability must not be exceeded during any loading event. 487

6. Summary

In recent years criteria have been established on the basis of the THTR licensing procedure and a material research and development programme for the design of graphite internals of an HTR. The components were classified into classes according to their special functions; the loads applied to the components were classified into load levels depending on their occurrence probability. For evaluating the stresses the Weibull theory was used. Allowable failure probabilities are defined rather than allowable stresses. This evaluation method can be applied to any stress state.

Reference /I/ J. E. Brocklehurst; M. I Darby Concerning the Fracture of Graphite Under Different Test Conditions Mat. Sei. and Eng., 16 (1974) 91 488 in 1 il1l Bl Figure 1 ml® Core Structure Hi m Hlfl Kernaufbau eines Hochtemperaturreaktors 85.44-1

Stres» (N/mm1] 20-

' — Exp«rim«ntal data 16- Normal distributions

Wvlbull distributions: Figure 2 S. - 6,8 S, - 21,47 m - 7.9 S.-0.0 S,-21,49 m-12,15 S.-0.0 S,-21.14 m-10.57 Y-0,95 Y: Confid«oc« l«v«l

10' Iff1 F«lluf» probability

Normal and Wetbull Failure Probability Plots HTR hRB of Tensile Strength Data 89.3-1 489

Max. str*M •t failure [N/mm*]

tnner diameter [mm] a o exp. ttMor. (Welbuil) 40- o 6J5 • « 10.1 • 0 30-

20- 0 3 * ' + . • o Ten«te Figure 3 ttrengm 10-

5 X

0.1 02 0,4 0,8 0A 1,0

Diameter ratio d,/d.

HTR Internal Pressure Tests on Ring Specimens 89J-2

Max. «tr«M at failure [N/mm*]

«* 60 Outer diameter [mm] •xp. (heor. (Welbuil} 50- * 12.7 a o 25,4 a 40- • 50.8 >

O 30- 8 a - 8 - . a3 3-Polnt Figure U • a 20- c bend ! vtrwiQth 10- 5

0 5 10 20 30 40 50 60 70

* Inner diameter [mm]

HTR hRB Diametral Tests on Ring Specimens 69J-3 490

Situation at General Atomics, San Diego, USA

Compiled by A. Schmidt, HRB Design criteria for graphite components of a Large HTGR (LHTGR) were generated by General Atomics in the early eighties. A document entitled: "Design Requirements for Graphite Core Supports" was provided for Section III; Division 2 of the ASME Boiler and Pressure Vessel Code, Subsection CE. Design criteria contained in this document were based on a deterministic analysis with definition of a minimum. ultimate strength and introduction of safety factors- These safety factors depend upon load levels and on the appropriate stress category (membrane, membrane + bending and peak stress).

Stresses in replaceable graphite components were evaluated similarly using different safety factors. Stresses, in excess, of the maximum permissable, were found in the fuel block centre, however. Therefor, local cracking had to be allowed for under certain boundary conditions.

For the Medium-size HTGR (KHTGR) the design criteria for replaceable graphite components were reviewed. The revision was based on the following requirements:

- limit cracking of fuel reflector prismatic blocks such that no loss of safety or reliability functions occurs

- limit wide cracks that could release fuel particles into primary coolant

- choose allowable stress levels consistent with limited cracking

The deterministic stress assessment was dropped and prooabilistic design criteria were introduced instead. The main reason were the uncertainties in the values of radiation induced stresses, which originate from uncertainties in the material data and boundary conditions together with strenght variations in the structural 491

component. These uncertainties are now analyzed separately and lead to the following assessment concept:

a) a structural component is evaluated with respect to the following boundary conditions:

- mean values for material properties - mean values for dimensions - best-estimates for oxidation

- worst-case expected values for operating parameters . inter-element gaps . inlet coolant temperature . power factors . fluence, etc. b) Assessment of stress in the replaceable graphite components is by comparison with the strength mean value (table 1), assessment of the shear stresses in the connecting dowels is by comparison with the minimum ultimate strength.

The assessment concept is being validated by means of a substantial program of investigation and verification.

REPLACEABLE GRAPHITE COMPONENT DOWEL SHEAR-FORCE-TO-STRENGTH LIMITS STRESS-TO-STRENGTH RATIO LIMITS SERVICE CONDITION LIMIT SERVICE CONDITION LIMIT LEVEL A 0.55 LEVEL A 0.*. LEVEL B LEVEL B 0.6 W/08E 0.60 • W/0 OBE 0.70

LEVEL C 0.90 LEVEL C 0.6 LEVEL D LEVEL D 0.9 W/SSE 0.90 W/0 SSE 1.10

Legend: OBE Operational Basis Earthquake SSE Safe Shutdown Earthquake Table 1 Table 2 492

Discussion of the presentation:

Design methods and criteria for graphite components

Altes. KFA Julien, FRG: For the graphite components you use a probabilistic design concept and for the metallic components one uses the deterministic design concept. Is it possible to use different design concepts for the same reactor?

Schmidt, HRB Mannheim, FRG: We have to do that because a concept based on a yield condition is not appropriate for the graphite material. It is not possible to separate a complex stress state and to evaluate each separated stress state independently.

Schuster, KFA Julien, FRG: Is the Weibull modulus m changed by neutron fluences relevant for reflector conditions? If yes, is this change taken into account in life estimation?

Schmidt, HRB Mannheim, FRGi We do not know the change of the Weibull m modulus during the irradiation of the graphite. But we know that the scattering of the strength does not increase. Additionally we know from experimental results that the strength is increasing up to high fluences and the same time the graphite behaviour becomes "more elastic", i. e. the stress-strain-curve shows almost a linear relationship. Therefore it can be assumed that the modulus increases during the irradiation. 493

Analysis of the Graphite Side Reflector Block of the HTR-Module P. Rathjen Interatom GmbH, Bensberg, FRG

Abstract Relations and characteristic quantities describing the material behaviour, the design methods and design criteria for reactor internals of graphite have been elaborated within the scope of a research and development programme in the FRG. As an example for the practical application of the design methods and criteria, a side reflector block of graphite for the HTR-Module has been analysed. The selected side reflector block contains an oblong hole which is located in the front part and separated from the pebble-bed core by a graphite ligament. This opening serves to take in the small absorber balls for long-term shutdown.

The material considered in the stress assessment is reactor graphite ASR-1RS. The side reflector is subjected to stress-controlled loadings due to the pebble-bed mechanics and to strain-controlled loadings due to non-uniform distributions of temperature and of neutron irradiation. The maximum fluence of fast neutrons within the design life of the reactor amounts to 1*1022 n/cm2 (EDN) with an associated irradiation temperature of <500 °C. This implies that the irradiation induced change of the linear dimension is within the shrinkage range of the graphite, even not reaching the "turn round" point.

The analysis of deformations and stresses, taking into account the irradiation induced changes of properties and creep, has been carried out using a three-dimensional model of the finite element programme IAGUAR developed by Interatom. The analysis considers reactor operation and reactor shutdown as loading conditions. The results are presented in the form of time and section profiles of stresses and deformations.

As is shown by the formation of the equivalent stress according to a modified criterion of maximum strain energy a reliable determination of the three principal stresses is important. As a consequence, the necessity of a three-dimensional description of the problem for the block geometry under consideration has been deduced.

The evaluation of stresses with respect to the strength of graphite has been performed by using two concepts. One evaluation concept is based on the definition of a minimum strength of graphite and the estimate of safety coefficients, while the other concept provides an evaluation using admissible failure probabilities. The experience gained with the trial of these two concepts of stress assessment is presented and the-results obtained are discussed. 494

Introduction

Relations and characteristic quantities describing the material behaviour, the design methods and design criteria for reactor internals of graphite have been elaborated within the scope of a research and development programme in the FRG. As an example for the practical application of the design methods and criteria, a side reflector block of graphite for the HTR-Module has been analysed.

The selected side reflector block contains an oblong hole to hold the small absorber ball shutdown elements (KLAK) for long-term shutdown. It is located in the front part of the side reflector and is separated from the pebble-bed core by a ligament 60 mm thick (Fig. 1).

The material considered in the stress assessment is reactor graphite ASR-1RS. The side reflector block is exposed to stress-controlled . loadings due to pebble-bed mechanics and to strain-controlled loadings resulting from non-uniform distributions of temperature and of neutron irradiation*

The maximum fluence of fast neutrons for the projected reactor life amounts to 1 x 1022 n/cm2 (EDN). With the associated irradiation temperature < 500 °C, the irradiation-induced change in linear dimension is within the shrinkage range of graphite, not even reaching the "turn- round" point. Apart from analysing deformations and stresses taking account of irradiation-induced property changes and creep, the main subject of the study is to test two methodical approaches for stress evaluation.

Analysis Method and Calculation Model Deformations and stresses resulting from strain-controlled loads are analysed with the FEM program IAGUAR developed at Interatom. All specific special features relating to the material and the loading of the reflector graphite, which is exposed to neutron irradiation, are taken into consideration in the graphite material model by IAGUAR.

On the material side these are the anisotropy of the physical properties, the neutron irradiation-induced property changes and material creep, which is also induced by neutron irradiation. The load side consists of non-uniform temperature distributions and gradients of the irradiation- induced change in linear dimension of the graphite corresponding to the fluence distribution of fast neutrons in the graphite block. The following calculation steps are performed to analyse the temporal development of deformations and stresses for reactor operation and reactor shutdowns: 495

A time-dependent analysis with an adequately large number of time steps which presents the stress/strain history of the reflector block. This calculation takes account of the effects of neutron irradiation-induced shrinkage strains and material property changes as well as the effects of creep.

Some elastic analyses to establish the temperature stresses, taking account of irradiation-induced property changes for selected time points in the reactor life. The superpositioning of these stress values with the stresses from reactor operation gives the shutdown stresses.

The side reflector block is modelled with 3D solid elements using the two planes of symmetry for 1/4 of the block geometry (Fig. 2). The selected model consists of 162 elements with 20 node points and 27 integration points per element (3x3x3 Gauss integration).

Loading and Material Data The investigations are based on the load conditions in the side reflector at the location of maximum fluence of the fast neutrons. Fig. 3 shows the distribution of the fast neutron fluence and the temperatures at the edge of the pebble-bed core during reactor operation as function of height of the reactor core.

The maximum fast neutron fluence for the projected reactor life amounts to approx. 1 x 1022 n/cm2 (EDN). The decrease of the neutron flux in radial direction "towards the outside takes place according to the law

0(v) = 0 -

The temperature distribution in the block is governed by the core edge temperature and the cooling gas inlet temperature. With reference to the radial temperature drop of approx. 170 K between core edge and first cooling gas hole. Fig. 4 presents the temperature distribution for the EOL condition.

The material data of the analyses performed are based on the unirradiated data and the irradiation behaviour of the graphite ASR-1RS developed for the high-fluence range.

The following table contains some typical values for the unirradiated graphite ASR-1RS. 496

Table: Typical material data ASR-1RS; material unirradiated, measurement parallel to the preferred grain direction

Young's Modulus, Edyn 1.02 X 104 [MPa] Coefficient of linear ther- 3.8 x: 10-6 [K-i] mal expansion, 20-200 °C Poisson's ratio 0.15 [-] Tensile strength 17.9 [MPa]

Weibull parameter, Rc 18.5 [MPa] Weibull parameter, m 9 1J [MPa] Bending strength 26.9 [MPa] Compressive strength 66.5 [MPa]

Estat/Edyyn (relations between 0.6 ( 0.8) [-] static and dynamic Young's Moduli; 0.8 for the irradiated condition) x* confidence level y = 0.95

The irradiation behaviour of reactor graphite is described in the paper "Material Behaviour and Design Values".

4 Results of Analyses

4.1 Deformation of Graphite Block

As a result of the irradiation-induced creep and the en- suing stress relaxation, most of the shrinkage strain of approx. 2 % is converted to a stress-free block deformation for the EOL condition. Corresponding to the radial flux profile, the block deformation for the EOL condition is as shown in Fig. 5. The necessary hydraulic separation between the small absorber balls opening and the cooling gas flow in the pebble-bed core, which is provided by design measures, must take account of the block deformations for the EOL condition.

4 . 2 Stresses in the Graphite Block

The temporal development of stresses for two selected positions in the block is illustrated in Fig. 6. While the stresses reach their maximum at approx. 0.3 x 1022 n/cm2 (EDN) in the rear section of the block (position 2) and then decrease again in the following time period, the stresses in the front section of the block (position 1) reach their maximum in the EOL condition. At both positions the shutdown stresses clearly exceed the operating stresses. 497

The stress distribution in the area of the small absorber ball opening is of particular interest from the viewpoint of assessment. Figures 7 and 8 shows the shutdown stresses in the EOL condition for the oblong hole ligament and the side ligament.

The stress distributions are characterized by distinct bending stress fractions, whereby the tensile stress regions are to the side of the oblong hole opening. The decisive principal stresses for the formation of the equivalent stress aeqU/ which is effected using the modified criterion of maximum strain energy

with

öi = 0

öi - (ÖZ/OD) * Oi for Oi < 0 Oi = Principal stress (ai. On, Dm)

az = Tensile strength

minimum strength s = Standard deviation

Rm = Average strength v a admissible stress adm SF f = Influence of reactor operation on graphite strength

SF = Safety factor

For the EOL condition with 1 x 1022 n/cm2 (EDN) at the core surface and 0.6 x 1022 n/cm2 (EDN) in the reference sections of the ligaments, the following admissible stress values are obtained for graphite quality ASR-1RS:

Admissible stress values oadm [MPa]

Tensile Bending Safety factor Condition stress 02 stress o B SF [ ] [MPa] [MPa]

Begin of Life 1.5 9.3 14.6 3.0 4.6 7.3 End of Life 1.5 14 .5 22.8 3.0 7.2 11. 4

Rm = 17.9 - 4.0 = 13.9 [MPa] tensile strength

Rm = 26.9 - 5.0 = 21.9 [MPa] bending strength

increase in strength due to neutron irradiation

400 °c = 1-56 for 0.6 x 1022 n/cm2 (EDN)

Stress Assessment Representative for a number of reference sections to be defined, two sections in the area of the oblong hole are examined: 499

Cracking may be tolerable, SF = 1.5

Equivalent pressure loading due to pebble- bed mechanics

Cracking that affects operation SF = 3.0

Location of Failure mode and reference sections consequences

The following safety factors are deduced on the basis of the potential crack consequences:

Section E-E -> SF = 3.0 Section B-B -> SF = 1.5

According to the diagrams in Figures 9 and 10, the bending stress fractions dominate in both reference sections. The calculated stresses are therefore compared to the admissible bending stress values for 0.6 x 1022 n/cm2 (EOL). The stress values are within the design limits in both sections.

5. 2 Stress Assessment using Admissible Failure Probabilities

The equation to describe the failure probability

o \m dV p = i _ H i j \ o V c r» max

adapted to the given FE element with 3x3x3 integration is evaluated numerically.

E = k IP = 27 P=\ - exp - I ± I //> E = 1

= Number of elements to describe the whole block or a part of the block

E = Element

IP = Integration point = Volume associated with the integration point 500

= Volume of the element under consideration = Equivalent stress at the integration point

= Maximum of the equivalent stresses occurring in the whole block or in a part of the block

n 0 Re = (E /E )* Rc

Rc = Weibull parameter, unirradiated Rc = Weibull parameter, irradiated

m = Weibull parameter

En = Young's Modulus of the irradiated graphite corresponding to the fluence and the irradiation temperature at the integration point

E° = Young's Modulus of the unirradiated graphite

If the total block volume is included in the numerical evaluation, a temporal development of the failure probability such as that shown in Fig. 11 is calculated for reactor operation. For comparative purposes, Fig. 11 also contains the calculated failure probability for some selected FE elements related to their respective element volume. The comparison clearly shows that the definition of structural areas, analogous to the reference sections in the traditional approach, is not particularly important for the assessment of the failure probability. The volume regions of maximum tensile stresses rather determine the failure probability value. And in contrast, volume regions with compressive stresses and moderate tensile stresses do not give any significant contribution to the failure probability value.

For random stress distributions in the reflector block with volume regions which are difficult to define for assessment of the failure probability, it is therefore possible in the case of a good approximation to include the entire block volume without concern for unacceptable results. The admissible failure probability for quality standard class 1 (load-carrying function) is defined as 10~4 and that for quality standard class 2 (neutron-physical function) as 10*4 - 10-2 (the latter value for the EOL condition).

As shown in Fig. 12, failure probabilities in the range of 10"4 are calculated for the shutdown stresses in the EOL condition. Within the scope of the investigations performed here, this is considered acceptable for the side reflector block. 501

6 Conclusions

Taking the side reflector block of the HTR Module as an example, the design methods and evaluation criteria for reactor graphite were tested. It was possible to demonstrate that with the available material-describing data, calculation methods, and assessment methods, successful concepts are available for the design of reflector blocks exposed to irradiation. The stress distributions in the two examined ligaments of the side reflector block in the form of distinct bending fractions permitted evaluation according to both the concept based on admissible failure probabilities and the traditional deterministic stress assessment.

With reference to ease of handling and interpretation of the results obtained, the numerical evaluation of the failure probability proved to be non-problematic. In contrast to this, the application limits of the traditional stress assessment are probably to be found in cases of stress distributions more random than those encountered in the ligaments examined here. 502

overall view side reflector block with oblong hole for small absorberballs shut-down elements (KLAK)

HTR-Module reactor internals Fig. 1 view of graphite components 503

1.0-10« investigated position

fast Huence

• 0.8-10»

free surlace oblong hole for KLAK

temperature at service condition

plane o( symmetry

top bottom height of the reactor core

Fast neutron fluence and temperature Three-dimensional finite element Fig. 3 model of graphite side reflector block Fig. 2 at the core surface (EOL)

values in (mm)

Block deformation at the end of Temperature distribution at operating Fig. 4 Fig. 5 condition (EOL) reactor life (EOL) 504

plane of symmetry

JO- tension area

Area of oblong hole ligament shut-down stress Fig. 7 distribution in circumferential direction (EOL)

plane at symmetry

tension area

1 1 1 1 1— —i 1 1 r- 0.2 0.4 0,6 0,8 1,0 1.1 Iluence (10« n/cm*) (EOL)

Equivalent stress versus fluence for two Area of oblong hole ligament shut-down stress Fig. 6 characteristic locations distribution in vertical direction (EOL) Fig. 8

stress (MPa} stress (MPa) allowable bending stress (SF=3,0, 0-0,6-10« n/cm*)

obtong

side ligament 80 mm -10J

Stress distribution across oblong hole Stress distribution across the side ligament, Fig. 9 Fig.10 ligament, shut-down condition (EOL) shut-down condition (EOL) 505

failure probability failure probability

10-6

1,0-10» n/cm* (EDN) 10-' 1 Ö side ligament 10-s H

10-8 H m - 7.0; Re- 18.5 N/mm» 10-e 4 -n Mc "c \£O ) (tor parameter sensitivity only) 10-9 -4

10-' H

10-10 H

10-a H

10-1' H fluence 0,2 0.4 0.5 0.8 1.0 1.1 10« n/cm* (EON! 0,2 0,4 0,6 0,B 1.0 1.1 10" n/cm? (EDN)

Failure probability versus reactor life Fig.11 Failure probability versus reactor life at Fig.12 for selected areas and total volume operating and shut-down condition 506

DESIGN CRITERIA FOR GRAPHITE COMPONENTS OF HTTR

Tatsuo IYOKU and Shusaku SHIOZAWA HTTR Designing Laboratory Taketoshi ARAI Materials Strength Laboratory Japan Atomic Energy Research Institute Tokai-mura, Naka-gun, Ibaraki-ken, 319-11 Japan

Abstract

The design criteria for the graphite core and core support components in the High Temperature Engineering Test Reactor have been developed by JAERI and will be authorized by the Japanese Government. The design criteria are mainly based on ASME Sec. Ill, Div. 2, Subsection CE Code (draft), however, the ASME CE Code is partially modified in the items of bi-axes failure theory, buckling limit and oxidation effects on the basis of test data. This paper summarizes the design criteria developed by JAERI and details the limits different from the ASME CE Code. A brief explanation is also made in this paper for quality control specified in the design criteria.

1. Introduction

The High Temperature Engineering Test Reactor (HTTR) is scheduled to be built at JAERI, Japan, and the fundamental design and preliminary safety review are finished. However, the design criteria for the graphite components are not specified in the Japanese regulations. Thus, the design criteria for the graphite components of the HTTR have been 507

developed by JAERI and will be authorized by the Japanese Government for the HTTR construction permission. The design criteria were drafted by the JAERI's internal committee and reviewed by specialists outside JAERI to be applied in the safety review of the HTTR core and core support graphite structures by the Government. This paper presents the design criteria with an emphasis on the differences from the ASME CE Code. Quality control provided in the criteria is also described in brief.

2. Description of graphite components

The reactor core is an array of graphite fuel, control rod guide and replaceable reflector blocks which provide the physical structure for arrangement and confinement of the fissile fuel materials, neutron moderation, heat transfer, and the positioning of control/shielding absorber materials . The core is supported by the core support structures and confined by the core restraint mechanism. The arrangement, of graphite components is shown in Fig. 1 . The graphite components are divided into two kinds, one for the permanently installed graphite components of the core support structures and one for the replaceable components of the reactor core.

2.1 Core components The fuel blocks are graphite hexagonal right prisms with array of fuel rod holes. The fuel blocks are 360 mm across the flats and 580 mm high. Three dowel pins are installed on the top face which engage with dowel sockets in the bottom face of the block above. The dowel arrangement ensures the correct orientation of fuel blocks within the column with respect to each other. They are fabricated from grade IG-110, isotropic fine-grade nuclear grade graphite. Control rod guide and replaceable reflector blocks are the same external shape as the fuel blocks. They are also fabricated from the IG-110 graphite. 508

2.2 Core support components The core support assembly is shown in Fig. 1 . The hexagonal hot plenum block array is made up of two axial layers. This structure provides lateral and vertical positioning and support of the core array. The hot plenum block assembly contains passages which collect the primary coolant flow from the outlet of the columns and distribute it into the high temperature plenum beneath the hot plenum blocks. Hot plenum blocks operate at core outlet gas temperature, 950 C. These blocks are fabricated from grade PGX, a structural grade, medium-to-fine grained molded graphite. The core support posts and seats are designed to structurally support the core and hot plenum block array while providing a flow plenum to receive the primary coolant flow exiting the core. The posts and seats are made from grade IG- 110. The core floor thermal insulation layer consists of three blocks: lower plenum block, carbon block and bottom block. The permanent reflector is a graphite structure immediately surrounding the replaceable reflector and control rod guide columns on the periphery of the core. The permanent reflector is an assembly of graphite blocks making 12 circumferential segments and 8 axial layers.

3. Design criteria

3.1 Overview The design criteria are mainly based on ASME Sec. Ill, Div. 2, Subsection CE Code (draft), with some modifications according to test data. The design criteria are characterized as follows. (1) Category of graphite components The graphite components are categorized into core component s and core support components, since the both are different in use as shown in Table 1 and the core support components are thought to be more important from a safety point 509

of view. Then, the stress limits are specified to be less severe in the core components than the core support components.

(2) Design analysis technique The maximum principal stress failure theory, partially introducing the modified Coulomb-Mohr theory, is adopted for the failure theory as shown in Fig. 2. Details will be described in the following subsection. The basis for determining stress-strain fields is linear elastic stress analysis for the core support components and linear viscoelastic (irradiation-induced creep) stress analysis. Since mechanical and physical properties of graphite change with temperature, irradiation and oxidation, the design criteria specify that the design analysis shall be made in consideration of these changes. The analitical techniques are proposed in the design criteria in evaluating the irradiation and oxidation effects. A visco-elastic material model is specified to be used in the analysis of the thermal/irradiation environmental effects for the core components. The analytical techniques for the oxidation effects will be detailed in the following subsection.

(3) Stress limit Although secondary stress is self limiting in character, the amount of stress redistribution possible in graphite is significantly less than in metals. Therefore, secondary stress limits are specified conservatively in the same manner as primary stress limits. As well, peak stress is also limited in order to prevent crack initiation and growth even for a single stress (static fatigue). Bending tests show that graphite bars subj ect to pure bending exhibit strength higher than companion specimens brought to the same maximum stress in uniaxial tension, which can be derived from the Weibull theory. This suggests that higher stresses may be allowed for tension plus bending stresses than for uniaxial tension stress. From the facts mentioned above, in the design criteria the stress limits apply to the three stress components; membrane stress, point stress {membrane plus bending stress) and peak stress (total 510

stress). Design fatigue diagram is expressed in terms of the ratio of minimum to maximum applied peak stresses on the basis of test data. Special limits of pure shear stress and buckling stress are considered in the design criteria. The details of the buckling stress limit will be given in the following subsection. The design stress limits (Hopper diagram) are presented for the core support components and core components in Fig. 3 and Fig. 4, respectively. (4) Specified minimum ultimate strength The specified minimum ultimate strength is determined from statistical treatment of strength data such that the survival probability is 99% at a confidence level of 95%, because the ultimate strength exhibits considerable statistical scatter. (5) Design data All the design data necessary for stress analysis and evaluation are provided in the design criteria.

3.2 Limits different from ASME CE Code The JAERI's design criteria are mainly based on the ASME CE Code, however, the ASME CE Code is partially modified in the items of bi-axes failure theory, buckling limit, oxidation effects on the basis of test data. The comparison between the both design criteria is presented in Table 2. The limits in the JAERI's design criteria different from the ASME CE Code are detailed in the following. (1) Bi-axes failure theory The bi-axes failure theory is determined on the basis of test data. The bi-axes failure stresses of PGX specimens are shown in Fig. 2. The maximum principal stress failure theory should apply in the first quadrant (tension-tension stress state). This theory can even extend partially into the fourth quadrant (tension-compression stress state). In the lower part of the fourth quadrant where the compressive stress component is high, a modified Coulomb-Mohr theory fits the data 511

satisfactorily. Based on this test data and the bi-axes failure test data of grade IG-110, the maximum principal stress failure theory, partially introducing the modified Coulomb-Mohr theory, is adopted for the failure theory in the design criteria.

(2) Buckling limit Buckling tests which simulates the loads imposed on the support post have been done to obtain the empirical data for assessing the buckling limit of the support post, because the buckling behavior of the support post is generally difficult to (?) be predicted only by analysis. The test results give the formula of critical compressive stress (ccrit ) which is of the RanKine-Gordon type, as shown in Fig. 5. Design critical stresses (°d ) are determined conservatively from the critical compressive stress, and the design limits are specified in the design criteria as shown in Fig. 5. The design limits are determined such that the safety factor in each operation condition is in accordance with that of the design stress limits for membrane primary plus secondary stresses.

(3) Oxidation effect The . graphite components in the HTTR are subjected to impurity reactants in the helium coolant during a normal operation and might react with O2 or H O-in air or water ingress accident, respectively. There are three kinds of reaction regimes, depending on temperature. At low temperature, in the "chemical regime", the reactions are so slow that reactant can penetrate the graphite in depth, causing rather uniform attack and thus reducing the graphite strength without changing apparent geometries. At high temperature, in the "mass transfer regime", the chemical reactivity so high that thinning of component occurs because of successive surface oxidation. Between these two regimes, in the "in-pore diffusion-controlled regime", the reactants diffuse in the pores of graphite with gradient of concentration, resulting in the reduction of 512

strength depending on the burnoff profile. Since reaction rates depends on the temperature, the kind of reactants and graphite grade, oxidation analysis shall be done in detail to estimate the burnoff profiles of graphite structures. The strength of oxidized graphite is specified in the design criteria to be evaluated in the following manner, (i) Geometry reduction The region where amount of oxidation exceeds the 80% burnoff shall be regarded as disappearance from the structure. Below the 80% burnoff, the oxidation effects are accounted for by reducing a corrosion allowance from the dimensions of components. The corrosion allowance shall be calculated from the burnoff profile. (ii) Strength reduction The tensile strength decrease of grade IG-110 is shown as a function of burnoff in Fig. 6. The stress evaluation shall be made according to Fig. 6. {4) Quality control Industry-wide standards for raw material formulations and processing of graphite and carbon have not been established. Therefore, graphite and carbon must be selected on the basis of the appropriate properties required by the reactor design. Their quality should be controlled during production and after machining with respect to their properties and specifications. Table 3 shows the inspection items specified in the design criteria, which are selected to meet the design specifications of the HTTR and to assure the material identification.

4. Summary

Design criteria for the graphite components to be employed in the High Temperature Engineering Test Reactor have been established by JAERI, Japan. The design criteria are mainly based on ASME Sec. Ill, Div. 2, Subsection CE Code (draft), however, the ASME CE Code is partially modified in the items of 513

bi-axes failure theory, buckling limit and oxidation effects on the basis of test data. The quality control and design data are also provided in the criteria, which are not completed in the ASME CE Code. Furthermore, the criteria for core components are also specified as well as core support components, while the ASME CE Code applies only to core support components. These criteria are hoped to be useful in the safety review and the construction permission for the HTTR.

References

(1) F. Ho, et'al.. Biaxial Failure Surfaces of 2020 and PGX Graphites, SMIRT 7 L4/6, Chicago (1983) (2) K. Kikuchi, et al., Compressive Fracture Test of Support Post, JAERI-M 9109 {1980)," (in Japanese) (3) S. Yoda, et al. , Effects of Oxidation on Tensile and Compressive Deformation Behavior for Nuclear Grade Isotropie Graphite, International Carbon Conf. 84, Bordeaux France (1984)

516

Table 3 Inspection items for graphite and carbon materials

o Material identification o Impurity Ash Boron equivalent (graphite only) Impurities (graphite only) o Mechanical and thermal properties Bulk density Specific resistance Thermal expansion coefficient Anisotropy ratio (graphite only) Thermal conductivity (carbon only) Bending strength Tensile and compressive strength o Microstructure inspection . o Dimensional inspection o Visual inspection o Non-destructive inspection (graphite only) o Assembling inspection 517

CORE GRAPHITE COMPONENTS

REPLACEABLE REFLECTOR BLOCKUG-11O)

CONTROL ROD GUIDE BLOCK( IG-11O)

FUEL BLOCK( IG-11O)

CORE SUPPORT GRAPHITE COMPONENTS

PERMANENT REFLECTOR BLOCK(PGX)

HOT PLENUM BLOCK(PGX)

SUPPORT POST( IG-110)

LOWER PLENUM BLOCK (PGX) CARBON BLOCK (ASR-ORB) BOTTOM BLOCK ( PGX )

FIG.1 THE ARRANGEMENT OF GRAPHITE AND CARBON COMPONENTS 518

0 O0O£> Ö a. MAXIMUM PRINCIPAL STRESS THEORY

-10

-20

MODIFIED COULOMB MOHR THEORY -30

-40 0 10 20

01 ( MPa )

Fig.2 BIAXIAL STRESS FAILURE DATA OF PGX GRAPHITE1

521 JGT H Lü CC TEST RESULT12' u; 1.0 8I 5 LLJ ^ • (T - > CO cr\

o VIPRE ! o N N i 1.0 - DESIGN CRITICAL STRESS \ i= (J. _ 0.9 Sue «==*^: •^^^a^H- 0.00448 (L/D)2 OO 3TRE S

IVE ! OPERATION CONDITION ^^-^^ 8 0.5 ~~~~~--~~^I\r^,DESIGN LIMTT^"^^ LU CC Q_ LCO K

IC A " ————

1 1 1 1 1 CR N 0 5 10 15 POST LENGTH / DIAMETER L : POST LENGTH D : POST DIAMETER : Suc SPECIFIED MINIMUM ULTIMATE COMPRESS IVE STRENGTH

FIG. 5 BUCKLING LIMIT OF CORE SUPPORT POST FOR EACH OPERATING CONDITION 522

1.0

0 0.8 bw

0.6

0.4 STRENGT H

LÜ

CO —r LÜ 1— 0.2 \LIZE D ORM /

0.1 0 10 20

BURNOFF ( Y. ) 6} : TENSILE STRENGTH OF OXIDIZED SPECIMEN

6\o: TENSILE STRENGTH OF UNOXIDIZED SPECIMEN

FIG. 6 DEPENDENTH OF STRENGTH ON BURNOFF (3) IN UNlFORMiLY OXIDIZED IGH10 GRAPHITE 523

Discussion of the presentation:

Design criteria for graphite components of HTTR

Schubert, KFA Julien, FRG: Can you give an example of components which tend to fail by buckling or do you understand by buckling a bending problem?

Iyoku, JAERI, Japan: Graphite core support posts are subject to both a dead load due to the core weight and a seismic load. Moreover, the posts are subjected to bending moments, because a small inclination can occur to absorb the difference between graphite blocks and a steel support plate. A vertical xompressive stress, a horizontal tensile stress, a shear stress and a bending stress are induced within the support posts. The posts can fail by these combined stresses.

Haag, KFA Julien, FRG: Why do you measure electrical resistance and why don't you measure thermal conductivity of graphitic (only of carbon)? How do you carry out microstructure inspection?

Iyoku, JAERI, Japan: In order to assure the material identification, we selected several inspection items. The specific resistance is one of them. The main function of carbon blocks is to keep the metallic support plate cooled below about 500 °C. Therefore, the thermal conductivity of the carbon blocks is selected to meet the design specifications of the HTTR. We inspect the microstructure by comparison with the standard samples.

Bodmann, HRB Mannheim, FRG: We can see large differences in the definition and limitation of the different stress categories between your contribution and that of A. Schmidt presented before. I believe that the stress categories: 524 primary, secondary, peak, are based on the assumption of ductile material. But graphite is not ductile in this sense. We made a loading test on a graphite block several years ago which failed at a very low level of "primary stress" due to a "peak stress" at a geometric discontinuity. 525

Section VI:

General summary 527

Statements on Current HTR Structure Design Criteria

R. Trumpfhe11er Essen

The papers presented in this workshop have demonstrated the status of high temperature reactor technology with regard to its realization in the nuclear power industry of various countries as well as to the development of safety rules in Germany. Standardized safety rules thoroughfully worked out under consideration of scientific and technical fundamentals form a solid base to ensure the quality and integrity of the nuclear power plant. Though it is justified to speak about the inherent safety of high temperature reactors with their high temperature resistent core and with the advantages of the inert gas coolant,the structure integrity of the safety relevant components must be given. The standardized safety rules are on the way to be.set up and will follow the principles of the integrity concept and include the design criteria as fundamental requirements.

Design, construction quality, and inspection of components are the practical and most effective measures to prevent severe accidents. Therefore I am glad that the design criteria for HTR could be presented in this workshop and that - as we have learned - these criteria already determine definitely and almost completely the relevant requirements of the component rules

The informations to our theme of the component design criteria have been presented in four sections after the introduction in- to the internationally existing HTR projects:

1 the technical boundary conditions with regard to safety 2 the metallic high temperature components . 3 a particular section dealing with the reactor pressure vessel, especially with the prestressed concrete vessel 4 the structural graphite components 528

The Technical Safety Boundary Conditions

The general safety goals as they are laid down for all kinds of nuclear plants in the regulations, e.g. in Germany the Ato- mic Law, the Radiation Protection Decree, the Safety Criteria for nuclear power plants issued from the competent federal ministry, are doubtlessly to be regarded as safety boundary conditions for design criteria and rules. Because further and more detailed safety goals specific for HTR and valid for the entire federal republic are not yet set up by the authorities, the authors of our HTR design criteria have looked through the regulations for light water reactors and those of other countries and - as far as applicable on HTR - derived special HTR requirements as safety goals from these regulations.

In the safety concept for light water reactors the so-called basic safety of pressurized metal lie components laid down in the guidelines of German Reactor Safety Commission (RSK) became one of the most important boundary conditions with regard to the integrity of these components.

As a further safety boundary condition the protection of the nuclear parts of the plants against external impacts like earth quake, airplane crash and explosion shock waves is to be considered .

The discussion about the assumption of defined design accidents, e.g. sizes and locations of leakages and other failures, as safety design criteria has not yet been finished. Certain- ly these assumptions must depend on the type of HTR.

The classification of components according to their safety relevance may be important for manufacturers and orderers, but safety relevant components operating in the same temperature range must be treated according to the same safety boundary conditions and the same design criteria. Perhaps, the density of external quality assurance measures may be classified, but 529

not the design criteria.

Metallic High Temperature Components

This workshop has demonstrated that the design criteria set up for safety relevant metal lie high temperature components make a safety concept named integrity concept that with regard to its safety fonction is quite similar to the basic safety concept for LWR.

Basic safety concept has been established for metallic components operating in the temperature range below 400°C. It requires that pressurized LWR components with nuclear safety relevance have to meet the following conditions:

- high purity and high toughness of materials - conservative limit of primary stresses - avoidance of high peaks of local stresses - application of optimized fabrication and testing technologies - knowledge and evaluation of flaw distributions, if existing - taking into account influences of the medium (and radiation)

If these conditions are carefully met point by point, large breaks of pressurized light water components may be excluded entirely. This break exclusion becomes the more reliable the better these conditions are fulfilled. In cases where the ful- filment can be easily surveyed the failure probability is exactly zero.

The criteria set up for HTR proceeded from the opinion that the LWR basic safety concept for pressurized components cannot be applied on HTR components in the temperature range above 400°C. There are only two reasons:

1 The material purity condition of basic safety has fixed 1i- mits for alloy and residual elements in the.chemical composition. For .materials in the higher temperature range certain 530

alloy elements must exceed these limits.

2 In the high temperature range alterations of material properties due to temperature are to be taken into account, which do not occur in LWR components under 400°C.

The basic safety condition requiring knowledge and evaluation of flaw distributions, if existing, includes in most cases the design condition that NDE of the components must be possible with a satisfying flaw detectibility. In the case of austenitic steels and nickel-base alloys the sensitivity of NDE methods could be not sufficient, if the wall thickness is too large for radiation testing techniques and coarse grained welding or casting materials decrease the sensitivity of ultrasonic testing by a high scattering rate in the sound beam. These problems in the field of NDE should not be regarded as a severe fault in the component safety concept, because this dis- advantage can be compensated by a particularly stringent quality assurance and control during fabrication and a dense inspection system in operation to prevent failures by quality decreasing influences, e.g. frequent visual inspections, measurements of temperatures, pressure, moisture etc., leakage control. Standardized rules, which now are on the way to be set up, will pay special attention to this matter. Furthermore it is to be mentioned that essential progresses in ultrasonic testing of coarse grain materials are to be expected, e.g. by use of horizontally polarized shear waves - as reported in this workshop.

The other items of basic safety can be fulfilled in HTR-tech~ nology as well as in LWR-technology, so the toughness requirements, conservative stress limits, avoidance of local stress peaks, optimized fabrication technologies, and consideration of operational influences from the medium etc., e.g. stress corrosion cracking, corrosion, radiation, heat insulation.

These reflections from the papers of this workshop emphasize that the safety concept of the metallic HTR-components, the integrity concept, is rather similar to the basic safety for 531

LWR-components and a very useful HTR design criterion. Never- theless, the two mentioned distinctive marks in comparison with basic safety, the use of high temperature resistent materials and the alteration of material properties in service, need particular consideration. Indeed, the papers of this workshop have dealt thoroughfully with this matter.

Of course, for HTR-types where the pressurized walls will not be exposed to temperatures above 400°C the basic safety concept can be applied as well as for LWR. But precautions to prevent high temperatures on pressurized wall and an effective temperature control of the wall are very important.

The specific design destining effect changing the material properties of metallic pressurized components in high temperature exposure is creep. By the limitation of creep strain the stress limits in general are determined lower than in the basic safety concept for temperatures under 400°C. In addition to creep deformation and damage the effects of creep fatigue, exhaustion, microstructural instabilities, high temperature corrosion, and, for components in neutron radiation areas, embrittlement due to irradiation including the influence of thermal neutrons are to be considered. From the investigations in these fields the working group derived two essential design requirements:

- strong limitation of total remaining creep strains - avoidance of local plastic and creep strains.

To meet these demands in the higher temperature range inelastic analyses are necessary. The constitutive equations for these analyses need a material data basis. These data are to be gained from material testing programmes. The data already available from investigation programmes in USA, Japan and Germany make inelastic analyses performable; nevertheless extension and partly additional improvements of the existing data basis are desirable. But it has to be stated that the present status of experiences, data bases and reviewing justifies the step from design criterion to design rule. I am glad to be able on this 532

opportunity to express my thanks and my appreciation to the working groups for their successfull efforts. We heard the opinion of the authors that comprehensive inelastic analyses will probably not become necessary in the temperature range up to 600°C. Such a result would correspond to many experiences in the operation of conventional plants, but certainly also in this area a few questions need to be solved. Some of them are on the way, e.g. the problem of stress, strain and creep damage in pipe bows; the research programme for the temperature range about 550°C is already running.

The knowledge of the mechanical load, its varation with time, the time depending temperature distribution in the structural wall material and the material data is a precondition for correct life-time predictions. Experiences with analyses, verification by experiments, control measurements and additional research programmes are necessary to improve the methods and tools of analyses. To cover the remaining uncertainties plant inspections are to be performed and evaluated, e.g. visual examinations of components, NDE, metal lographic tests, and strain and temperature measurements. With such measures the integrity of high temperature components can be ensured with the same reliability as the basic safety components of LWR.

Reactor Pressure Vessels

The section deal ing with reactor pressure vessels introduced prestressed concrete vessels and metallic vessels the walls of which are kept on lower temperatures by cooling measures. prestressed concrete pressure Vessels are characterized by the separation between the load bearing fonction of prestressed concrete and the sealing fonction of the liner. The reliability of the wall cooling system and the integrity of both the prestressed concrete wall and the sealing liner are to be maintained through the whole service duration. It is easy to see that this can be realized. Furthermore it is possible to examine these requirements recurrently by means of in-service inspections and measure- 533

merits during operation.

The pressure vessel unit of HTR modul is also operating with lower temperature of about 350°C in the pressurized wall. There- fore design rules of LWR components are applicable, the basic safety concept can be used without exception. But the protection of the pressurized walls against the higher gas temperature is an essential requirement, which has to be controlled during operation.

Structural Graphite Components

The papers dealing with structural graphite components reported about the experience that an effective quality assurance during fabrication can ensure that the components stay through the whole life-time and that damaged components, if recognized, can be replaced. The present knowledge makes it possible to set up rules for graphite components, at least as a first approximation.

Summary

Summarizing I dare to say: all the facts which have been report ed in this workshop justify to establish standardized rules for structural design requirements for HTR components now. Our Ame- rican and Japanese colleagues already have gone this way. I af- firm my statement that the preconditions for setting up the rules are given, though there are a few questions left and to be solved. The solutions are a matter of time only . The preliminary report for the German nuclear standardization committee (KTA) has been finished and will be submitted soon.

I think we have to congratulate the working groups and Profes- sor Nickel as their Spiritus rector and head for this success. I finish with my wish and encouragemant: Carry on and be successful in your future standardization work as you were until now in preparing the fundamentals. 534

Discussion of the presentation:

Statements on current HTR structure design criteria

Schubert, KFA Julien, FRG: One of the main principles in LWR basic safety is the toughness. The toughness for high temperature metallic components must be replaced by specifying a material, which does not show creep notch embrittlement sensitivity. 535

Comments concerning the "Workshop on Structural Design Criteria for HTR"

Werner von Lensa, Kernforschungsanlage Julien - Projektträger des BMFT für die Entwicklung von Hochtemperaturreaktoren

About 30 years of HTR development in the FRG represent 30 years of funding. More than 8 billion marks have been spent to reach the present technical level and to build and operate the AVR and THTR.

The funding of nuclear power plant development today needs clear and persuasive arguments and a strong motivation. What could be the motivative aspects for going nuclear as an additional possibility for energy production? as the first area - competitiveness, economics, compatibility with the existing energy supply structure and reliability secondly - safety and environmental compatibility thirdly - long-term fuel availibility and supply security and last but not least - public acceptance.

This is precisely the position in democratic economic and national systems and all these preconditions must be fulfilled. Safety cannot replace profitableness for instance or vice versa.

Further R+D must be devoted to achieving further progress in all these fields because competing technologies are also exposed to an ongoing optimization process and may further more experience no great problems in public acceptance, for example.

The erection time and feedback of operating experience in conventional technologies are relatively short-term processes compared to the extremely long project preparation and completion times in the nuclear field (THTR about 15 years from start of construction to start of operation). It is clear that the mere conservation of 536

proven technologies of "forerunner" projects does not lead to competitive modern designs. As 'trial and error' cannot be the guideline for the development of nuclear plants there is a permanent need for focused and engaged R+D to achieve a 'state of the art' basis for new projects which will once again go into operation under further developed or changed technical, economic and politi- cal environments and are built to face extended lifetimes of about 30-50 a. So there is an overwhelming claim to the quality of precursing R+D.

It is a stringent requirement that extreme flexibility is needed to cope with such long-term, quasi - historical frames.

The HTR system possesses exciting features for future tasks in the production of electricity, process steam and process heat as a supplement to other systems and alternatives.

This is due to the simple and ingenious principle of the coated particle, i.e. a small and robust pressure vessel that can retain the fission products at their place of origin under all reasonable accident conditions.

What other fuel element is already exposed in the manufacturing process to considerably higher temperatures than those to be expected under worst conditions?

The coated particle is the essential high-technology product of HTR development within the last decade and it is the linking element between the different HTGR developments all over the world.

The 'integrity concept' that has been formulated within this project is based on this still revolutionary message and indicates that we have 'invested' a great deal directly into the quality of the fuel element to reach the safety goals. Other reactor systems have to invest in multiple barriers and active systems that do not have the same importance for HTGR's. 537

On the other hand, we must be aware that limited operating experience . with HTR's is the main handicap compared to the thousands of operating years with other reactor systems. Half of the total HTR operating experience is based on the AVR reactor that was shut down at the end of last year after 21 years of successful operation. The long-term behaviour of typical systems will be thoroughly studied in the following programme which should be most interesting for the whole HTR community; also as a check for design criteria.

Concerning the "classical" discussion about active and passive safety systems I would like to hint at an important aspect relevant for R+D.

Whereas active safety systems can simply be checked in their function, passive systems cannot normally be controlled in that way. The reliability of a passive safety system needs a much higher degree of physical and technical understanding than being necessary for active systems. This involves a higher quality in research, testing and demonstration to obtain the acceptance by the public and the licensing authorities, as well as to eliminate residual technical risks or needless conservative assumptions.

It is also obvious that the quantity of R+D to be implemented exceeds the possibilities of a single company. It can only be done by a broader community of system engineering and manufacturing companies together with potential users, research institutes, licensing and funding authorities. I think that we can learn a lot from the so-called 'integrated approach' as performed in the USA on the basis of only one lead project.

To introduce a new reactor line that fits the said criteria may even be beyond a purely national scope - since it has not been achieved up to now. In the present situation there is a strong tendency towards collaboration and concentration in joint ventures by nearly all well known manufacturers all over the world, even 538

for estabished and proven reactors (for example KWU/Framatom) and this concentration of know how will create even better and more economical concepts to compete with.

Unter these circumstances, the potential for international cooperation should be fully used to establish and broaden the basis for commercially competitive HTR projects which also cover the acceptance criteria for future reactor generations and - simply - to reduce FOIK costs.

On the occasion of this international workshop where we presented the results of a 9 years project on proposals for design criteria, we would like to demonstrate that we are open for a fruitful information exchange with all interested partners.

The discussions and reports also showed the limits and unresolved problems as well as improvement possiblities in this field. I would like to encourage the participants to keep in contact by defining work-sharing collaboration projects and to avoid parallel activities wherever possible and where it makes sense.

The governments have fulfilled their function with all the partners that are relevant here, by umbrella agreements on international collaboration, but it is the task of the experts and companies to fill these frames in an effective manner.

Finally, I would like to thank the project leader, Prof. Nickel, for his enthusiastic commitment to HTR-specific design criteria. He is also the 'father' of the 'integrity safety concept' that should be the guideline for further activities.

My thanks are also due to the whole team for their excellent work. 539

Summary of the Final Discussion

There has already been good cooperation in the HTR sector with various countries for quite a number of years, comprising:

- Umbrella Agreement (since 1978) USA - FRG - Switzerland

- Japan - FRG Contracts e.g. KFA-Jiilich - JAERI (since ~ 1980)

- Contracts China - FRG (small experimental HTR, 1988) USSR - FRG (prototype HTR, 1988)

Cooperation should be seen in relation to the following HTR projects:

USA: . MHTGR Japan: HTTR - China: 10 MW experimental reactor USSR: modular prototype HTR FRG: HTR 500, HTR-Modul - Switzerland: district heating HTR 20 MW

The status of work in various countries concerning structural design codes for HTR components is as follows. In the Federal Republic of Germany, the data and fundamental methods required have been derived during the past eight years. Based on the final report, the preliminary reports for the following nuclear design codes (1989)

KTA 3221 "Metallic HTR Components" - KTA 3231 "Safety-Related Requirements for the Design of Prestressed Concrete Reactor Pressure Vessels of a HTR Plant" KTA 3232 "Ceramic HTR Core Internals" will be submitted to the Nuclear Technology Committee with the intention of publishing the first drafts in 1990. 540

In the USA, there is one task force revising parts of the CC N 47 and one task force for "Very High Temperature Design". In Japan, work on design codes is closely oriented to the requirements of the HTTR using as far as possible LWR rules and procedures from CCN 47. The data material required for Hastelloy XR is being revised. There was agreement among the participants that the activities for structural design codes must be coordinated at an international level. Care should be taken to ensure that stress limits and assurance concepts are possibly identical or must at least be harmonized to such an extent that no striking differences occur. There is no complete agreement on the procedure for determining the stress equivalent value "S." between ASME CC 47 and the German proposal. The methods for deriving lifetime fraction rules from LCF experiments with hold times are not uniform' either. A further open question in connection with all design codes is that of the best suitable constitutive equations. The continuation of work for deriving simplified inelastic methods remains a task for the future.

There are two basic methods for evaluating the operational behaviour of graphitic core internals:

— assurance by stress reference values with safety factors — probabilistic determination of failure probability and its limitation.

In the course of the workshop, it became apparent that, in particular, the Japanese design concept to be applied to the HTTR is very conservative. Extensive compliance with the LWR rules accounts for the fact that the specific safety—related properties of the HTR have hardly been taken into consideration.

The Japanese representatives underlined that the approach selected for the HTTR aimed primarily at facilitating the licensing procedure. In the longer run, however, more credit should be given to the safety—relevant features of a HTR. International cooperation may be very helpful.

The German side pointed out that extensive basic research results were available for the design and assurance of HTR component integrity. This also applies to the temperature range of a nuclear process heat facility. However, more in-depth studies are still required in certain areas. 541

Members of working groups and task forces of the German Design Criteria project 543

Responsible for the reasearch and development project: KFA/IRW, H. Nickel

Members of the general group "HTR-Design-Criteria"

KFA/IRW F. Schubert ABB*, Mannheim K. Schneider GRS, Köln K. Bieniussa (bis 1987), H. Reck (1987) HRB, Mannheim E. Bodmann, J. Pschowski, C. Elter, A. Schmidt Interatom, Bensberg H. Breitling, H.-J. Seehafer IzfP, Saarbrücken F. Walte KFA/IRW H. Over (bis 1987), G. Breitbach (ab 1987) H.-J. Penkalla MPA Stuttgart W. Gaudig RWTÜV, Essen M. Dette SSP** Bochum K. Schimmelpfennig, R. Oberpichler

Guests:

BMFT H. Pohl KFA/HBK D. Mindermann, H. U. Brinkmann KFA/HTA G. Scheidler, J. Terkessidis KFA/PTH W. von Lensa

* ABB = ASEA Brown Boveri, vormals BBC = Brown Boveri & Cie ** SSP = Stangenberg, Schnellenbach und Partner GmbH, vormals ZSP = Zerna, Schnellenbach und Partner GmbH 544

Section A: Technical safety boundary conditions

Members of the working group:

KFA/IRW F. Schubert GRS, Köln K. Bieniussa (bis 1987) H. Reck (ab 1987) HRB, Mannheim C. Elter HRB, Mannheim J. Pschowski Interatom, Bensberg D. Klein IzfP, Saarbrücken F. Walte KFA/IRW G. Breitbach KFA/ISF J. Wolters RWTÜV, Essen M. Dette, K. Hofmann, W. Trapp SSP, Bochum K. Schimmelpfennig

Guests:

BMFT, Bonn H. Pohl KFA/PTH W. von Lensa

Section B: Metallic structural components

Responsible for Section B: KFA/IRW, F. Schubert

Members of task force Bl (Examination of materials)

RWTÜV, Essen M. Dette ABB, Mannheim F. Brzoska HRB, Mannheim C. Elter, J. Pschowski Interatom, Bensberg E. Ohrt RWTÜV, Essen G. Gnirß, V. Weber 545

Members of task forces B2, BIO: (Fabrication of materials and semifinished products and formats)

RWTÜV, Essen J. Just ABB, Mannheim F. Brzoska, M. Bloch, H. Mücke, W. Ostgathe HRB, Mannheim J. Pschowski, H. von Woikowsky-Biedau Interatom, Bensberg W. Niemeyer, E. Ohrt MPA, Stuttgart H. Waidele IzfP, Saarbrücken F. Walte RWTÜV, Essen M. Dette, V. Weber, G. Gnirß

Members of task forces B3, B4: (Data of physical and mechanical properties, constitutive equations)

KFA/IRW H.-J. Penkalla ABB, Mannheim R. Bürgel, G. Raule, V. Detampel, R. Schubert GRS, Köln K. Bieniussa, H. Reck HRB, Mannheim E. Bodmann, H. Diehl, F. Kemter Interatom, Bensberg H. Breitling, H. Honeff, H. Hübel MPA, Stuttgart W. Gaudig, V. Obst RWTÜV, Essen M. Dette, H.-J. Lehmann, V. Weber

Members of task force B5: (Life fraction rules)

MPA, Stuttgart W. Gaudig, K. Maile ABB, Mannheim G. Gnirß GRS, Köln K. Bieniussa, H. Reck HRB, Mannheim H. Diehl Interatom, Bensberg H.P. Meurer, R. Rohdenburger KFA/IRW H. J. Penkalla MPA, Stuttgart M. Hoffmann, V. Obst RWTÜV, Essen H.J. Becker, H.J. Lehmann 546

Members of task forces B6, B7: (Stress categories and load levels, failure modes and design criteria)

GRS, Köln K. Bieniussa, H. Reck ABB, Mannheim H. König, K. Schneider HRB, Mannheim E. Bodmann, J. Pschowski Interatom, Bensberg H. J. Seehafer KFA/IRW H. Over, G. Breitbach MPA, Stuttgart W. Gaudig RWTÜV, Essen M. Dette, M. Lüdeke, W. Petruschke, M. Möller

Members of task force B8: (Construction elements and requirements)

HRB, Mannheim E. Bodmann ABB, Mannheim H. König GRS, Köln K. Bieniussa, H. Reck Interatom, Bensberg H.J. Seehafer IzfP, Saarbrücken F. Walte KFA/IRW H. Over, G. Breitbach RWTÜV, Essen M. Dette, M. Heil

Members of task force B9: (Stress analysis)

Interatom, Bensberg H. J. Seehafer ABB, Mannheim H. König GRS, Köln K. Bieniussa, H. Reck HRB, Mannheim E. Bodmann KFA/IRW H. Over, G. Breitbach, H. J. Penkalla MPA, Stuttgart W. Gaudig RWTÖV, Essen M. Dette, M. Lüdeke 547

Section C: Prestressed concrete pressure vessel

Responsible for Section C: SSP, K. Schimmelpfennig

Members of task force Cl: (Prestressed concrete structure)

SSP, Bochum K. Schimmelpfennig HRB, Mannheim A. Weber Interatom, Bensberg M. Bauermeister RWTÜV, Essen H.J. Wagler SSP, Bochum M. Borgerhoff

Members of task force C2: (Liner)

SSP, Bochum R. Oberpichler GRS, Köln K. Bieniussa, H. Reck HRB, Mannheim H. Fröning, 0. Pschowski Interatom, Bensberg M. Bauermeister LCS, Gummersbach R. Weber RWTÜV, Essen M. Dette TÜV Bayern, München E. Prucker

Members of task force C3: (Pressure vessel closures)

RWTÜV, Essen P. Wieczorek HRB, Mannheim H. Fröning, J. Pschowski IzfP, Saarbrücken F. Walte SSP, Bochum R. Oberpichler

Members of task force C4: (Heat insulation system)

HRB, Mannheim G. Groß HRB, Mannheim J. Pschowski RWTÜV, Essen M. Dette, H. Fullroth, S. Werner 548

Section D: Graphitic structural components

Responsible for Section D: ABB, A. Schmidt

Members of the working group:

HRB, Mannheim A. Schmidt HRB, Mannheim E. Römischer, P. Kubaschewski Interatom, Bensberg R. J. Schulze, P. Rathjen, J. Willaschek KFA/HBK H. U. Brinkmann, V. Maly, D. Mindermann KFA/IRW G. Haag RWTÜV, Essen G. Wintermann Sign*, Meitingen G. Kraus, M. Spann