Iter Blanket, Shield and Material Data Base

ITER DOCUMENTATION SERIES, No. 29

INlS-mf —13018

ITER BLANKET, SHIELD AND MATERIAL DATA BASE

INTERNATIONAL ATOMIC ENERGY AGENCY, VIENNA, 1991 ITER BLANKET, SHIELD AND MATERIAL DATA BASE ITER DOCUMENTATION SERIES, No. 29

ITER BLANKET, SHIELD AND MATERIAL DATA BASE

PART A (ITER BLANKET AND SHIELD)

D. SMITH, W. DAENNER, Y. GOHAR, T. KURODA, G.E. SHATALOV, A.B. ANTIPENKOV, H. ATTAYA, C. BAKER, M. BILLONE, I.V. DANILOV, L.EL GUEBALY, M. FERRARI, P. GIERSZEWSKI, V.I. KHRIPUNOV, V.G. KOVALENKO, P. LORENZETTO, S. MAJUMDAR, K. MAKI, S. MORI, L. PETRIZZI, V. RADO, A. RAFFRAY, F. ROSATELLI, A. ROSSANI, M. SAWAN, O.L. SHCHIPAKIN, A. M.SIDOROV, G. SIMBOLOTTI, YU.S. STREBKOV, I. SVIATOSLAVSKY, H. TAKATSU, V.N. TEBUS.V. VIOLANTE, H. YOSHIDA, F. ZACCHIA, S.A. ZIMIN

PART B (MATERIAL DATA BASE)

D. SMITH, I.V. ALTOVSKY, V.R. BARABASH, J. BEESTON, M. BILLONE, J.L. BOUTARD, T. BURCHELL, J. DAVIS, S.A. FABRITSIEV, M. GROSSBECK, A. HASSANEIN, G.M. KALININ, P. LORENZETTO, R. MATTAS, K. NODA, R. NYGREN, N.V. ODINTSOV, V.V. RYBYN, H. TAKATSU, V.P. VINOKUROV, R.WATSON, C.WU

INTERNATIONAL ATOMIC ENERGY AGENCY VIENNA, 1991 ITER BLANKET, SHIELD AND MATERIAL DATA BASE IAEA, VIENNA, 1991 IAEA/ITER/DS/29

Printed by the IAEA in Austria October 1991 FOREWORD

Development of nuclear fusion as a practical energy source could provide great benefits. This fact has been widely recognized and fusion research has enjoyed a level of international co-operation unusual in other scientific areas. From its inception, the International Atomic Energy Agency has actively promoted the international exchange of fusion information. In this context, the IAEA responded in 1986 to calls for expansion of international co-operation in fusion energy development expressed at summit meetings of governmental leaders. At the invitation of the Director General there was a series of meetings in Vienna during 1987, at which representatives of the world's four major fusion programmes developed a detailed proposal for a joint venture called International Thermonuclear Experimental Reactor (ITER) Conceptual Design Activities (CDA). The Director General then invited each interested party to co-operate in the CDA in accordance with the Terms of Reference that had been worked out. All four Parties accepted this invitation. The ITER CDA, under the auspices of the IAEA, began in April 1988 and were successfully completed in December 1990. This work included two phases, the definition phase and the design phase. In 1988 the first phase produced a concept with a consistent set of technical characteristics and preliminary plans for co-ordinated R&D in support of ITER. The design phase produced a conceptual design, a description of site requirements, and preliminary construction schedule and cost estimate, as well as an ITER R&D plan. The information produced within the CDA has been made available for the ITER Parties to use either in their own programme or as part of an international collaboration. As part of its support of ITER, the IAEA is pleased to publish the documents that summarize the results of the Conceptual Design Activities.

U CONTENTS

PART A. ITER BLANKET AND SHIELD CONCEPTUAL DESIGN

I. INTRODUCTION 15

II. SUMMARY OF CONCEPTUAL DESIGN 17 1. DESIGN GUIDELINES AND INTEGRATION ISSUES 17 2. BLANKET DESIGN 20 3. BLANKET ISSUES 25 4. SHIELD DESIGN 27

III. FUNCTION, DESIGN REQUIREMENTS AND CRITICAL ISSUES 29 1. FUNCTION AND DESIGN REQUIREMENTS OF BLANKET AND SHIELD 29 1.1 Blanket 29 1.2 Shield 31 2. CRITICAL DESIGN ISSUES FOR THE BLANKET AND SHIELD 32 2.1 Blanket 32 2.2 Shield 34

IV. MATERIAL SELECTION 37 1. STRUCTURAL MATERIALS 38 2. TRITIUM BREEDING MATERIALS 38 3. ELECTRICAL INSULATORS 39

V. BLANKET AND SHIELD SEGMENTATION 41

VI. BLANKET DESIGN DESCRIPTION 47 1. LAYERED CONCEPT 47 1.1 Outboard Section 48 1.2 Inboard Section 53 1.3 Copper Stabilizer Integration 53 1.4 Penetration accommodation 55 1.5 Fabrication and Assembly 55 2. PEBBLE CONCEPT 60 3. BIT CONCEPT 70 3.1 Poloidal BIT Concept 70 3.2 Toroidal BIT Concept 78 4. LITHIUM-LEAD EUTECTIC CONCEPT 81 5. SUMMARY OF BLANKET DESIGN PARAMETERS 88

VII. DESIGN ANALYSIS 93 1. NEUTRONICS ANALYSIS 93 1.1 Introduction 93 1.2 Poloidal Neutron Wall loading distribution 93 1.3 Tritium Breeding Analysis 94 1.4 Activation and Decay Heat Analysis 114 2. THERMAL/MECHANICAL ANALYSES 120 2.1 Breeder Temperature Control 120 2.2 Hydraulic Analysis 126 3. STRESS ANALYSIS 128 3.1 Normal Operation 128 3.2 Disruption Analysis 131 4. TRITIUM RECOVERY AND INVENTORY 133 5. POWER VARIATION CAPABILITY 144 6. SAFETY ANALYSIS 146 6.1 LOCA Analyses of the BIT Concept 146 6.2 LOFA Analyses 147 6.3 Safety Analyses for the Lithium Lead Concept 151

VIII. SHIELD 155 1. DESIGN LIMITS AND SAFETY FACTORS 155 2. NUCLEAR RESPONSES IN THE TF COILS 157 2.1 General Remarks 157 2.2 Comparative Analysis of all Relevant Results 160 2.3 Three-dimensional Analysis for a Specific Configuration 167 2.4 Conclusions 170 3. DOSE AFTER SHUTDOWN 171 3.1 Dose Equivalent from Bulk Shield 171 3.2 Dose Equivalent from NBI Duct Shield 172 4. RADIATION STREAMING ANALYSES 173 4.1 Assembly Gaps 173 4.2 Vacuum Pumping Ducts 174 4.3 Divertor Coolant Tube Penetrations 175 4.4 Neutral Beam Injection Ports 176 4.5 Diagnostics Channels 177 5. SUMMARY OF BENCHMARK CALCULATIONS 177 CONTENTS

PART B. ITER MATERIALS EVALUATION AND DATA BASE

I. INTRODUCTION 183 1. STRUCTURAL MATERIALS 184 2. TRITIUM-BREEDING MATERIALS 185 3. PLASMA-FACING MATERIALS 186 4. ELECTRICAL INSULATORS 187

II. STRUCTURAL MATERIALS 189 1. TYPE 316 AUSTENITIC STAINLESS STEEL: FIRST WALL AND BLANKET STRUCTURE 189 1.1 Selected Materials 189 1.2 Status of Existing Data Base 189 1.3 Main Key Issues 202 2. DISPERSION STRENGTHENED (DS) COPPER: DIVERTOR STRUCTURAL MATERIAL 203 2.1 Basis for Selection 203 2.2 Status of Existing Data Base 203 2.3 Key R&D Needs 206 3. NIOBIUM ALLOYS 207 3.1 Basis for Selection 207 3.2 Status of Existing Data Base 208 3.3 Compatibility with Water Coolant 209 4. MOLYBDENUM ALLOYS-PFC STRUCTURE 213 4.1 Basis for Selection 213 4.2 Status of Existing Data Base 214 5. BRAZING ISSUES 218

III. BLANKET MATERIALS 219 1. CERAMIC BREEDER 219 1.1 Basis for Selection 219 1.2 Critical Issues 221 1.3 Status of Existing Data Base 221 1.4 Key R&D Needs 228 2. BERYLLIUM 231 2.1 Basis for Selection 231 2.2 Critical Issues 231 2.3 Status of Existing Data Base 231 2.4 Key R&D Needs 234 3. LiPb BREEDER 237 3.1 Basis for Selection 237 3.2 Critical Issues 237 3.3 Status of Existing Data Base 238 4. AQUEOUS LITHIUM SALT BREEDER 241 4.1 Introduction 241 4.2 Critical Issues 241 4.3 Status of Existing Data Base 242 4.4 Key R&D Needs 245

IV. PLASMA FACING MATERIALS 247 1. CARBON BASED MATERIALS 247 1.1 Data Base 247 1.2 Thcrmo-mechanical properties 252 1.3 Critical Issues 256 2. TUNGSTEN 258 2.1 Surface Properties 258 2.2 Physical Properties 259 3. BERYLLIUM 262 3.1 Data Base 263

V. ELECTRICAL INSULATORS 267 PART A

ITER BLANKET AND SHIELD CONCEPTUAL DESIGN

12. CONTRIBUTORS

PART A (ITER BLANKET AND SHIELD)

M. ABDOU, V.YU. ABORIN, S. ABOUSAID, V.V. ALEXANDROV, G.A. ANTONOV, L. ANZIDEI, A. BADAWI, A. BERTRAM, J. BLANCHARD, E. BOGUSCH, V.P. BONDARENKO, M. CAIRA, M. CHAZALON, A.M. CHEPOVSKI, S. CHIOCCHiO, F. CLAVAREZZA, R. CLEMMER, M. DALLE DONNE, V.V. DEMIDOV, O.F. DIKAREVA, E.V. DMITRIEVSKAJA, Z.A. DURIGINA, S.A. EREMIN, F. FABRIZI, G. FEDERICI, P. FINN, R. FOGARTY, M. GALMNA, D. GALLORI, F. GERVAISE, P. GORANSON, Z. GORBIS, H. GORENFLO, M. GRATTAROLA, Y.M. GRIBKOV, G. HAHN, C. HAMMONDS, S. HIRATA, P. HUBERT, O. IOP, Y. ISHIYAMA, R. JAKEMAN, C. JOHNSON, G. JONES, LA. KARTASHOV, G.M. KALININ, V.P. KARKLINA, H. KHATER, S.E. KHOMYAKOV, V.Y. KIRPAL, YU.G. KLABUKOV, K. KOIZUMI, T. KOBAYASHI, V.YU. KOLGANOV, G.I. KONDIR, J. KOPASZ, A.E. KOPYEV, K. KOSAKO, G. KULCINSKI, A. LIDE, C. LIN, A.V. LOPATKIN, V.E. LOUKASH, S.V. LUKASHIN, H. MAEKAWA, S.A. MAKAROV, M.L. MALYSHEV, A.V. MARACHEV, V.G. MARKOV, A.D. MARKOVSKY, M. MARTONE, R. MATTAS, J. MAYHALL, C. MAZZONE, M.N. MEDVEDEV, A. I. MERENKOV, H. MIURA, E. MOGAHED, V.N. MOSKALEV, D. MUIR, S. MUKHERJEE, T. NAKAMURA, C. NARDI, B. NELSON, V.Z. NEPOMNYASHCHIY, S.B. NIKITINA, S. NISHIO, V.D. NOVIKOV, G. OTTONELLO, V.V. OVECHKIN, L.D. PANTELEEV, I.P. PAZDRY, B.V. PETROV, V.V. POLIKSHA, V.YA. PROCHORENKO, E. PROUST, L. RINALDI, A. SANTAMARINA, R.T. SANTORO, KE. SATO, KO. SATO, F. SECOLO, G.A. SERNYAEV, V.V. SHIPILOV, E.A. SHIVERSKIY, YU.N. SOKURSKY, A.V. SIDORENKOV, E.N. SINITSIN, K.S. SKLADNOV, L. SORABELLA, S.I. SOROKIN, D. STEINER, K. SUMITA, T. SUZUKI, A.N. SVETCHKOPAL, D. SZE, S. TAM, YU.M. TRAPEZNIKOV, L. TURNER, A.G. UKHLINOV, M. UNO, K.A. VERSCHUUR, G. VIEIDER, A.V. VINNIKOV, N.K. VINOGRADOVA, D. WILLIAMSON, L. WITTENBERG, S. YAMAZAKI, A. YING, M. YOUSEFF, T. ZABAN, V.S. ZAKHARTSEV, A.N. ZARYANOV, V. ZEMLIANKIN I. INTRUDUCTION

The terms of reference for ITER provide for incorporation of a tritium breeding blanket to supply most of the tritium for the fuel cycle of the device. The blanket and shield combined must be designed to operate at a neutron wall loading of about 1 MW/m and to provide adequate shielding of the magnets to meet the fluence goai of 3 MWa/m at the first wall. The blanket is integrated with the first wall and shield into a box-like construction with separate coolant manifolds for the blanket/shield and the first wall. The blanket/shield system is highly segmented to allow for vertical remote maintenance and to accommodate disruption loads. The system is designed to operate for the entire lifetime of ITER without replacement. An austenitic steel structure with low temperature, low pressure water coolant was selected for the blanket/shield and first wall. A ceramic breeder concept has been selected as the first option for the diiver blanket. A beryllium multiplier is used in the blanket to enhance the tritium breeding capability. A Pb- Li breeder concept is proposed as an alternate concept. The combined blanket/shield design provides for protection of the magnets and the possibility of personnel access to the cryostat boundary after shutdown of the reactor. The shield is constructed primarily of austenitic steel and water. Local shielding around major penetrations is designed to reduce radiation to permissible levels. Section II gives a summary of the blanket and shield design, key results of the analyses, and the R&D requirements to qualify the systems for ITER construction. Section III gives a description of the function, the design requirements and specifications, and the critical design issues. The candidate materials considered for the blanket and the basis for selection are discussed in Section IV. The materials data base for these materials is summarized in Part B jf this technical report. A discussion of the blanket and shield segmentation requirements is presented in Section V. Details of the design description for the first option and alternate option blankets are presented in Section VI. The design analyses conducted in support of the blanket designs are summarized in Section VII. Additional details have been presented in each parties design reports. The shield design description and the supporting analyses are presented in Section VIII. II. SUMMARY OF ITER BLANKET AND SHIELD CONCEPTUAL DESIGN

The terms of reference for ITER provide for incorporation of a tritium breeding blanket with a breeding ratio as close to unity as practical. The main function of this blanket is to produce the necessary tritium required to achieve the ITER operation and test program objectives. The limited tritium supply from the international market dictates this tritium breeding function. In addition, the use of an effective breeding blanket provides a substantial economic advantage based on the current unit cost for tritium. The primary design goals for the blanket are the following: - achieve a net tritium breeding ratio (TBR) of about one, - operate at an average neutron wall loading of about 1 MW/m , - achieve an average fluence of at least lMWa/m and up to 3 MWa/m2, - be compatible with an overall machine availability of at least 10% with a goal of reaching about 25% in the technology phase, - tolerate accident conditions with passive methods.

The criteria developed to select the driver blanket for ITER include the following: performance capability, safety and environmental aspects, cost considerations, R&D requirements, reactor relevance and benefits, and reliability considerations. Three blanket concepts were considered during the concept definition process: ceramic breeder (solid breeder) concept, lithium-lead breeder concept, and aqueous-salt breeder concept. All concepts incorporate an austenitic steel (Type 316) structure with low temperature (<100 C) water coolant. The ceramic breeder concept has been selected as the "first option" for ITER with a Pb-Li breeder concept proposed as an alternate.

11.1. DESIGN GUIDELINES AND INTEGRATION ISSUES

Figure II-1 is an isometric view showing the inboard and outboard blanket segments of a single sector. This level of segmentation is necessary for the remote maintenance scheme. The main operating conditions and design guidelines are summarized in Table II-1. The blanket must accommodate the change in the power levels between the two operating phases, significant variations in fusion power within each phase, and the poloidal variations of the neutron wall loading. Also, the blanket/shield system must: - incorporate the twin loop copper stabilizer in the upper and lower outboard regions, - accommodate magnetic diagnostic loops within the blanket/shield and other diagnostic penetrations through the blanket, um» PL uo

OUTBOARD CENTRAL UPPER BLANKET SEGMENT

INBOARD OUTBOARD 5JD£ BLANKET SEGMENT BLANKET SEGMENT

LOWER PLUG

OUTBOARD SIDE BLAIKL-T SEGMENT OUTBOARO CENTRAL LOWER BLAKE? SEGMENT Fig. II-1. Isometric view shewing the inboard and outboard blanket segments of a single sector.

- provide support for the upper divertor plates with minimum shift during operation, - protect the TF coils against radiation damage and excessive nuclear heating, - allow for personnel access to the cryostat boundary after shutdown - provide for sixteen ports in the outboard midplane. Operating temperature limits specified for different materials are indicated in Table 2-1. The minimum temperature limits for the ceramic breeder materials are based on tritium recovery issues while the maximum temperature limits are based on mass transfer and material sintering considerations. The blanket performance during the off-normal conditions (plasma disruption, loss of coolant, or loss of coolant flow) has also been considered in the blanket design process. The design philosophy was to accommodate such conditions with passive methods. For example the inboard blanket is segmented to accommodate the electromagnetic forces during the plasma disruption. Internal reinforcement ribs are used to provide additional support for the outboard blanket.

18 TABLE III. BLANKET OPERATING CONDITIONS AND DESIGN GUIDELINES

PHYSICS TECHNOLOGY PHASE PHASE

FUSION POWER, MW 1100 860

NEUTRON WALL LOAD, MW/mZ (MIN/MAX)

INBOARD 0.4/1.1 0.3/0.9 OUTBOARD 0.8/1.5 0.6/1.2

DT FLAT BURN TIME, s UP TO 400 2300

MINIMUM DWELL TIME, s 200 200

,4 4 NUMBER OF DT PULSES 10 5xlO

DT FLUENCE GOAL, MWa/m2 0.05 3

OPERATING TEMPERATURE LIMITS,°C

AUSTENITIC STEEL (316)

STRUCTURAL COMPONENT <400 SHORT TERM OFF NORMAL <800 AQUEOUS INTERFACE <150

CERAMIC BREEDER TEMPERATURE RANGE Li2° 370 -1000 LLAiO, 450-900 370 -1000

BERYLLIUM <600

Special attention should be given to avoid mass transport above 800oC.

19 II.2. BLANKET DESIGN

The ceramic breeder concept has been selected as the first option with either LuO or a ternary ceramic (e.g., LiAKk or LuZrO.,) as the breeder material. The design specifications for the first option are given in Table II-2. The Pb-Li breeder concept is proposed as an alternate concept. Austenitic steel (Type 316 solution annealed) was selected as the reference structural material on the basis of an extensive database and ease of fabrication. Structural steel temperature limits are < 400°C because of radiation induced swelling and < 150°C because of aqueous stress corrosion considerations. The design allowable stress intensity, S for annealed material is 110 MPa. Cold- worked material, which has a much higher allowable stress limit, is an alternate option. Annealing of the cold-worked structure during fabrication processes, e.g., brazing and welding must be considered. Low temperature (60 - 100cC), low pressure (<1.5 MPa) water is specified as the reference coolant. In addition to the lower primary stress requirements, the low pressure significantly reduces safety concerns with pressurized water. The desire to achieve a tritium breeding ratio of about one with limited breeding volume because of inboard shielding requirements, divertor requirements, numerous penetrations, and provisions for nuclear testing; requires the extensive use of beryllium as a neutron multiplier for the ceramic breeder option. Primary constraints associated with the use of beryllium are radiation induced swelling , embrittlement and tritium retention. Cost and chemical toxicity must also be considered in the design. The operating temperature limit for beryllium is < 600°C because of swelling consideration. The first wall, blanket, and shield are integrated into a single unit with a separate cooling system for the first wall. Poloidal and toroidal coolant flow were chosen for the inboard and outboard first wall, respectively. Both poloidal and toroidal cooling were considered in the blanket designs. The inboard blanket is divided into three submodules per segment with two segments per sector in order to accommodate electromagnetic loads predicted during a disruption. The outboard blanket is divided intc three poloidal segments per sector with the central segment divided into upper and lower modules to provide for the major penetrations. All blanket segments are manifolded at the top except the lower central outboard segment which is manifolded at the bottom. Twin loop copper stabilizers about 1.5 meter high and 5 mm thick are incorporated into the upper and lower outboard blanket segments. Two solid breeder configurations are considered for the detailed blanket design, viz., a multilayer configuration shown in Figures 2-2 and 2-3 and a breeder-in-tube configuration shown in Figure 2-4. With a low temperature water coolant a major design problem with the ceramic breeder concept is to maintain the breeder temperature sufficiently high (>400-450°C) to provide for efficient tritium recovery with a margin up to 100°C and a low blanket tritium inventory. The layered configuration utilizes the beryllium zones up to several centimeters thick to provide the desired temperature gradient between the low temperature (60-100cC) coolant and the thin breeder zone. The water coolant is separated additionally from the tritium in the breeder. The temperature control in the

20 TABI E H-2. TRITIUM BREEDING BLANKET DESIGN SPECIFICATIONS

FIRST OPTION BLANKET CERAMIC BREEDER STRUCTURAL MATERIAL AUSTENITIC STEEL (316) COOLANT WATER: 60-100°C, < 1.5 MPa

BREEDER MATERIAL LijO or TERNARY (LiAlO2, L^ZO)

6Li ENRICHMENT 50-95%

NEUTRON MULTIPLIER BERYLLIUM

BREEDER CONFIGURATION LAYERED or BREEDER-IN- TUBE

BREEDER AND MULTIPLIER CLAD AUSTENITIC STEEL (316)

BREEDER TEMPERATURE CONTROL GRADIENT IN BERYLLIUM OR HELIUM GAS GAP

TRITIUM RECOVERY METHOD CONTINUOUS IN-SITU PURGE GAS: He + (0.2-1%)

COOLANT FLOW DIRECTION INBOARD-FIRST WALL POLOIDAL BLANKET POLOIDAL or TOROIDAL OUTBOARD-FIRST WALL TOROIDAL BLANKET TOROIDAL or POLOIDAL

ALTERNATE BLANKET Pb-Li BREEDER STRUCTURAL MATERIAL AUSTENITIC STEEL (316)

COOLANT WATER: 60-100°C, < 1.5 MPa

BREEDER 83Pb-17Li EUTECTIC ALLOY

TRITIUM RECOVERY BATCH PROCESSING

COOLANT FLOW POLOIDAL

21 K>

Fig. II-2. Isometric view of multilayered ceramic breeder blanket design with toroidal cooling and both LUO breeder and Be neutron multiplier in the form of sintered blocks S-S. (.'I AD

L - . O ft SB! I 1068

BfttlD/NG RfGION COOLANf (H:Ol

SHIftDING RfGION

L IUM

Fig. II-3. Cross sectional view of multilayered ceramic breeder blanket design with L^O breeder and Be neutron multiplier in the form of small sintered pebbles u> BLANKET MODULE - DEUJL

he I iurn gap outer clodding beryllium pressure tube water 2nd clodding spacer 1st clodding

OUTBOARD BLANKET M/DPUNE SECTION

Fig. H-4. Cross sectional view of BIT ceramic breeder blanket design with poloidal cooling and both LiA102 breeder and beryllium in the form of sintered pellets breeder-in-tube configuration is achieved by varying the thickness ( 1-3 mm) of a helium gas gap. This gas gap is also used for the helium purge for tritium recovery. The coolant is contained in an annular region outside the breeder and inside the beryllium. Two forms of ceramic breeder and beryllium are considered: sintered product (blocks or pellets) and small (approximately 1 mm dia) spheres or pebbles. The ceramic breeder is highly enriched (50-95% Li). The mass of ceramic breeder in the entire blanket varied from 13-90 tonne. Tritium is recovered from the breeder by a helium purge (He + 0.2 to 1% HU)- A net tritium breeding ratio of 0.8-0.9 is achievable. The calculated tritium inventory in the breeder can be maintained at less than 100 g even with cyclic operation. The total beryllium inventory in the blanket is about 200 tonne. Based on very limited data and conservative estimates that include the chemical and irradiation-induced trapping of tritium in beryllium, the end-of-Iife total tritium inventory in the beryllium multiplier zones is <400 g at 1 MWa/m2 and < 1.2 Kg at 3 MWa/m . This assumes that all of the tritium generated in the beryllium is retained. The blanket would be designed with a separate helium purge loop for the beryllium multiplier. The preliminary calculations indicate that the lifetime of the blanket exceeds the design goal of 3 MWa/m . Primary uncertainties in the lifetime predictions relate to aqueous stress corrosion and irradiation embrittlment of the steel, swelling of beryllium, and electromagnetically induced loads that occur during a disruption. Internal support structure within the outboard blanket is required to accommodate the projected disruption loads. An alternate blanket concept with 83 Pb-17 Li eutectic alloy as the breeding material has also been developed as shown in Fig. 11-5. The lithium-lead blanket has poloidal breeding channels which follow the first wall geometry. Each channel consists of coaxial pipes where the breeder is separated in individual chambers. The mass of Pb-Li in the entire blanket is about 1000 tonne. During operation, the eutectic alloy is in the solid form. The alloy is melted and transported out of the reactor for tritium recovery during the reactor down time. The lead serves as a neutron multiplier; therefore, beryllium is not used in this concept. A tritium breeding ratio of 0.7-0.8 is attainable. The main issues with this concept relate to problems associated with melting the breeder and safety problems associated with tritium containment and generation of polonium.

II.3. BLANKET ISSUES

The analyses performed on the ceramic breeder blanket designs indicate that the design goals can be achieved. The R&D requirements to qualify the designs have been identified. Major R&D issues include: Austenitic Steel Structure: Primary issues include aqueous stress corrosion and irradiation effects on the low temperature fracture toughness of Type 316 austenitic steel. Additional data are also required on the effects of irradiation on mechanical properties including weldments and braze joints.

25 ELEVATION VIEV OUTBOARD INBOARD B- a A - A H 1:b

Ee 0

H 1:4 amuruc «oi saw

Purge gos inLet/ Euteccic overflow evacuation Water coolant or heating gas inlet

Eutectic filling/ Purge gos out Lee

Woter coolant or heatrng gos outlet

Fig. II-5. Cross sectional view of 83Pb-17 Li breeder blanket design with breeder- in-tube configuration and poloidal cooling Characterization of the ceramic breeder: Data on tritium release and irradiation effects on the mechanical properties are required to reduce the design uncertainties.

Characterization of beryllium: Additional data are needed on fabrication techniques, irradiation effects such as swelling, embriulcment, tritium solubility and transport properties, and compatibility with the. steel structure.

Characterization of ceramic insulator: Development of reliable insulators and radiation testing is required.

Demonstration of temperature control: The methods used to provide a thermal insulation between the coolant and the ceramic breeder require proof testing under relevant conditions.

Lithium-lead breeder: Additional data are required on the thermomechanical behaviour of lithium-lead breeder concept.

Integrated module tests: Both inpile radiation testing and out-of-pile thermal- hydraulic and mechanical testing of mockups are required. This includes simulations of disruption loads.

II.4. SHIELD DESIGN

The shield has been designed in an integrated manner with the first wall, blanket and vacuum vessel because of geometrical and structural requirements. These components have a shielding function, which is accounted for in the shielding analyses. The main function of the shieid is to reduce the neutron and photon leakage intensities at the outer shield surfaces to acceptable levels. This reduction ensures that (a) the different reactor components arc protected from radiation damage and excessive nuclear heating, and (b) the workers and the public are protected from radiation exposure during operation and after shutdown. The preferred shielding materials are type 316 stainless steel and water. Integration issues, structural considerations, materials and shielding data base, and fabrication experience are the main reasons for this selection. Lead and boron carbide are considered to enhance the shielding performance. Their use is limited to the back of the vacuum vessel within the last 5 cm. Tungsten suggested for selected areas instead of lead and boron carbide where the nuclear responses warrant. A set of requirements has been defined in the form of upper limits for the different nuclear responses to satisfy these shielding functions. The limits are given in Table II-3. The shielding performance has been calculated by a number of one-, two-, and three-dimensional analyses. Special analyses were performed to study the effect of shield discontinuities and penetrations. A set of safely factors was developed to account for the uncertainties in the nuclear data,

27 TABLE II-3 SHIELDING PERFORMANCE PARAMETERS (SAFETY FACTORS ARE INCLUDED)

Response Design Calculated Limit Value

Total nuclear heating in toroidal field coils, KW 55 57 Peak nuclear heating in winding pack, mW/cm 5 1.8 Peak dose to electrical insulator, rads 5xl0y 5xlOy Peak fast (E>0.1MeV) neutron fluence to superconductor, n/cm 1019 6xlO18 Peak displacement in copper stabilizer, dpa 6xlO"3 2.7xl0'3 Peak helium production in type 316 stainless steel of the vacuum vessel to permit rewelding, appm 1 11 Biological dose outside cryostat one day after shutdown, mrem/h < 0.5 0.5a

a) A total thickness of 175 cm is assumed for the outboard first wall, blanket, shield, and cryostat.

calculational models, and calculational procedures. The shield is designed to accommodate two phases of operation and to achieve an average fluence of 3 MWa/m . During the life of the machine, about 3.8 full power years (FPY) of operation are expected; 0.05 FPY in the physics phase and 3.7 FPY in the technology phase. The anticipated fusion power in the physics and technology phases are 1100 and 860 MW. Table II-3 gives a summary of the main nuclear responses. In general, the shielding criteria are satisfied except for selected areas in the divertor regions where the shield performance is marginal. For example,rewelding of the vacuum vessel will be limited in the divertor regions near end of life.

28 fll. FUNCTION, DISIGN REQUREMENTS AND CRITICAL ISSUES

III.l. FUNCTION AND DESIGN REQUIREMENTS OF BLANKET AND SHIELD

III.l.l. Blanket

The terms of reference for ITER provide for incorporation of a tritium breeding, or driver blanket in the ITER design. The main function of this blanket is to produce the necessary tritium required for the operation of ITER, particularly during the test program in the technology phase. The limited tritium supply from the international market dictates this tritium breeding function. In addition, the use of an effective breeding blanket provides substantial economic advantage based on the current unit cost for tritium. The blanket also serves a shielding function. The goals for the ITER blanket include the following:

- Achieve a net tritium breeding ratio (TBR) of one - Operate at an average neutron wall loading of 1 MW/m - Achieve an average fluence of at least 1 MWa/m up to 3MWa/m2 - Be compatible with an overall machine availability of at least 10% and up to 25% in the latter stages of the technology phase - Tolerate accident conditions with passive methods

Ah* blankets must be compatible with the basic machine configuration and maintenance scheme. The blanket is integrated with the first wall and shield into a box-like construction with separate coolant manifolds for the blanket/shield and the first wall. Figure 3-1 is an exploded view of one blanket sector which includes the following:

Outboard - One sector per TF coil - Three poloidal segments per sector - Central segments are split into upper and lower segments to accommodate central ports - a total of 64 segments

Inboard - Two segments per sector (TF coil) - Each segment is divided into three subsegments which are electrically insulated to reduce electromagnetically induced disruption loads - a total of 32 segments or 96 subsegments

29 INBOARD BLANKS 1 StGMEM

(.OkfR PLUG

OUTWARD sroe BLANKET SEGMENT OUTBOARD CENTRAL 81 .WET SEGMENT

Fig. Ill-3. Isometric view showing the inboard and outboard blanket segments of a single sector.

The blanket shield system must also be designed to:

- incorporate the twin loop copper stabilizer (~ lm wide) in the upper and lower outboard regions, - accommodate magnetic diagnostic loops within the blanket/shield and other diagnostic penetrations through the blanket, - provide support for the upper divertor plates with minimum shift during operation, - protect the TF coils against radiation damage and examine nuclear heating, - allow for personnel access to the cryostat boundary after shutdown, - provide for sixteen ports in the outboard midplane,

Vertical coolant manifolds are required at the top to facilitate maintenance. An exception is the lower central segment and the lower shield plug, which are manifolded at the bottom.

30 TABLE IH.-l OPERATING CONDITIONS AND DESIGN GUIDELINES

PHYSICS TECHNOLOGY PHASE PHASE

Fusion Power, MW 1100 860 Neutron wall load, MW/m Inboard 0.4/1.2 0.3/0.9 Outboard 0.8/1.6 0.6/1.2 DT Flat Burn Time, s up to 4000 2300 Minimum Dwell Time, s 200 200 NUMBER OF DT PULSES 104 5xlO4 DT Fluence Goal, MWa/m2 0.05 3

Operating Temperature Limits, C Austenitic Steel (316) Structural Component < 400 Short Term off Normal < 800 Aqueous Intc face < 150 Ceramic Breeder Temperature Range 370-1000* LiAiO2 450-900 Li2ZrO3 370-1000 Beryllium <600

* Special attention should be given to avoid mass transport above 800°C.

The main operating parameters are listed in Table III.-l. Operating temperature limits are specified for all candidate materials. The blanket must be designed to accommodate the change in the power levels between the physics and the technology phases and significant variation in fusion power within each phase. The blanket must also be designed to accommodate the poloidal changes of the neutron wall loading which are reduced to 0.38 and 0.50 of the values at the midplane for the inboard and outboard regions, respectively. The blanket performance during off-normal conditions (plasma disruptions, loss of coolant, or loss of coolant flow) must also be considered in the design process. The design philosophy is to accommodate such conditions with passive methods. For example, the inboard blanket is segmented to reduce electromagnetic forces during the plasma disruptions and internal reinforcement ribs are used to provide additional support for the outboard blanket.

III.1.2. Shield

The main function of the shield during operation is to reduce the neutron and photon leakage intensities at the outer shield surfaces to acceptable levels. This reduction ensures that (a) the different reactor components are

31 protected from radiation damage and excessive nuclear heating, (b) the neutron reaction rates in the reactor components outside the shield system are reduced to avoid high-biological dose values in the designated areas for personnel access, and (c) the workers and the public are protected from radiation exposure. After shutdown, the shield system has to attenuate the decay gamma rays so that workers are permitted access to designated areas in the reactor hall within one day after shutdown with all shield in place. A set of requirements is defined in the form of upper limits for the different nuclear responses to satisfy these functions. The limits include the following:

- Total nuclear heating in toroidal field coils 55 KW - Peak nuclear heating in winding pack 5 mW/cm - Peak dose to electrical insulator 5x10 rads - Peak fast (E > 0.1 MeV) neutron fluence to superconductor 10 n/cm - Peak displacement in copper stabilizer 6 x 10" dpa - Peak helium production in type 316 stainless steel to permit rewelding 1 appm - Biological dose outside cryostat one day after shutdown 0.5 mrem/h

The first wall, blanket, and shield should be integrated in a single structural unit which is attached to the vacuum vessel. The four components are also performing a shielding function. These components should be designed to operate without scheduled maintenance for an average fluence of 3 MWa/ra . The preferred shielding materials are type 316 stainless steel and water. Lead and boron carbide should be limited to the back of the vacuum vessel within the last 5 cm. Tungsten material should be used only if essential from the shielding point of view.

II1.2. CRITICAL DESIGN ISSUES FOR THE BLANKET AND SHIELD

The blanket and shield design concepts developed during the conceptual design phase have been analyzed in sufficient detail to show that the objectives of ITER are credible. However, several critical design issues identified during the design phase must be further resolved before a commitment for construction can be made. The R&D required to qualify the blanket and shield designs have been identified and an R&D plan to address these issues has been developed.

III.2.1. Blanket

Critical blanket design issues that require further R&D and analysis to qualify the blanket design for ITER include: (1) additional materials data base for the structural material, ceramic breeder, beryllium, and ceramic insulator; (2)

32 technology related to temperaiure control of the cenmic breeder with low temperature water cooling; and (3) fabrication and testing of blanket mockups.

II1.2.1.1. Austenitic steel structure

Primary issues for the austenitic steel structure include development of an improved data base on aqueous stress corrosion and the effects of irradiation on the low temperature fracture toughness of Type 316 stainless steel. The aqueous corrosion effort should focus on effects of water chemistry associated with radiolysis, crevice effects, and irradiation effects for both base metal and weldments. The low temperature fracture toughness of irradiated steel must be evaluated including appropriate transmutation effects. Additional data are also required on the effects of irradiation on the tensile and creep properties including weldments, braze joints, and cold-worked material.

I 11.2.1.2. Characterization of the ceramic breeder

Additional data are needed on tritium release and irradiation effects on the mechanical properties of candidate ceramic breeder materials. For Li^O , additional data are required to better define the operating temperature window, including cyclic temperature effects and the mass transfer of lithium in the purge stream.

111.2.13. Characterization of beryllium

Further development of fabrication techniques of beryllium is needed to determine more reliably the rang;; of microstructures and porosities that c? - us readily fabricated. Additional data are needed on irradiation effects such as swelling, tritium solubility and transport properties, and chemical compatibility in order to define temperature limits for acceptable swelling and to evaluate ways to minimize the tritium inventory in beryllium,

HI. 2.1.4. Characterization of ceramic insulators

Further development is required to assure reliable insulator performance. Of particular importance is the determination of extent of radiation enhanced conductivity in a high flux environment and the mechanical integrity of the insulators under projected loading conditions.

111.2.1.5 Demonstration of adequate temperature control

Two approaches have been proposed to provide the desired temperature control of the ceramic breeder in a water cooled system. The first, which is based on the temperature gradient through the beryllium neutron multiplier and associated interfaces, requires further demonstration of reliability under cyclic thermal conditions, mechanical loadings, and materials-related effects under

33 irradiation. The second approach, which is based on temperature gradients through a thin (~ 1 mm) gas gap, also requires validation under thermal cycling and radiation creep conditions.

II 1.2.1.6. Thermomechanical behavior of lead-lithium concept

Testing is required on the thermomechanical response of blanket modules under conditions of melting and solidifying the eutectic alloy to assure acceptable performance for the tritium recovery operation. Additional testing of the tritium recovery operation and handling is needed.

III.2.1.7. Mockup fabrication and out-of-pile integrated test

Development of procedures for fabrication of mockup blanket segments is required. An integrated thermal-hydraulic and mechanical test of a one-half size mockup of a blanket segment should be conducted to assure satisfactory performance under the loading conditions that will be encountered in ITER. This task will require development of a major test facility which can be obtained by modifying an existing test facility to provide the appropriate test conditions for the first wall/blanket mockup test.

III. 2.1.8 Inpile irradiation testing of blanket module

A separate test of a small blanket module in a neutron environment is required to qualify the blanket for operation in ITER. Since a nuclear test of a large mockup is not feasible, a small module with characteristics of a blanket segment will be tested in ar. existing fission reactor to validate tritium recovery and performance in a high neutron flux environment. This module will contain breeder, coolant, neutron multiplier, and structure in a configuration similar to that of the reference blanket design.

III.2.2. Shield

Critical design issues that require further R&D and analysis to qualify the shield design for ITER include: (1) bulk shield performance, (2) shielding performance with internal voids, geometrical irregularities, and penetrations (3) neutron Kerma factors and activation cross-sections evaluations and measurements for the isotopes used in the different reactor components.

111.2.2.1. Bulk shield performance

Shielding experiments are required for typical shield configurations including the use of lead, boron carbide, and tungsten. Processing of the most recent nuclear data files is needed to perform analyses and comparison with the experimental results. The results will be used to verify the nuclear data required for the shield design, the adequacy of the different approximation used in the

34 design process, and the calculation^ methods. Also, Nuclear data processing is urgently needed for the design process of the nuclear components.

IH.2.2.2. Shielding performance with internal voids, geometrical irregularities, an d penetrations

Streaming and penetration experiments are needed to qualify the ITER shield design around the midplane ports, the water coolant lines, and the electrical elements of the vacuum vessel. Plane gaps with and without steps, circular ducts of various dimensions, and void regions should be included. Results will provide a measure for assessing the safety factors and the overall accuracy in nuclear calculations or empirical correction factors used in the design process.

III.2.2.3. Neutron Kerma factors and activation cross-sections

Evaluations and experiments are required to measure the neutron Kerma factors and activation cross-sections. Results are required to verify the nuclear heating in the nuclear components and the build up of radioactive isotopes and decay heat. These data will guarantee the expected thermomechanical performance of the nuclear components. IV. MATERIAL SELECTION

The operating limits for ITER depend to a great extent on the performance limitations of materials in the unique environment of a fusion reactor. As a consequence an important aspect of the ITER conceptual design activity is the selection of appropriate materials for all major components.The primary objectives of this activity were: (1) to compile and assess the available materials, (2) to develop and recommend a single set of design curves or correlations for use by the designers, and (3) to identify critical materials data needs and to recommend the required materials R&D needed to provide an adequate data base for the construction of ITER. Candidate materials for each application are listed in Table IV-1. A database for the candidate materials was compiled and presented at two Specialists Meetings on Materials Data base held at the ITER site in Garching in June 1988 and February 1990. The materials database assessment and evaluation is reported in more detail in the ITER Materials evaluation and Data Base report [1].

TABLE IV-1.1. CANDIDATE BLANKET MATERIALS

Component Candidate Materials

Structural Materials First Wall/Blanket Type 316 austenitic steel (SA) Type 316 austenitic steel (CW)

Tritium Breeding Materials Ceramic Breeder

24423 Metal Breeder Materials 83Pb-17Li eutectic Alloy Aqueous Lithium Salt Breeder LiOH.L^NiO-, (solutions) Neutron Multipliers Beryllium Lead Ceramic Insulators Oxides Aipy g Nitrides AlON;Si3N4,

37 IV.l. STRUCTURAL MATERIALS

IV.2. TRITIUM BREEDING MATERIALS

This class of materials includes ihe neutron multiplier materials in addition to the lithium bearing tritium breeding materials such as the ceramic- breeders, Pb-Li alloys, and aqueous lithium salts. The primary considerations in the selection of the breeder materials include: tritium breeding capability; ease and reliability of tritium recovery; thermal transport properties; and thermal, chemical, and irradiation stability. Since the blanket is expected to last the entire reactor lifetime, these material will receive neutron fluences corresponding to a neutron wall load of up to 3 MWa/m . The operating temperature range will be set primarily by the tritium recovery scenario. Candidate ceramic breeder materials include Li2O, LiAK^ and Li^ZrOo and LLSiCK. The data base covers the baseline physical properties, baseline mechanical properties, chemical stability/compatibility, radiation effects, and tritium solubility/transport. Features of the ceramic breeder relate to relative safety associated with their chemical stability, potential for low tritium inventory, relatively good data base, and potential reactor relevance. Key issues are associated with reliable tritium recovery-transport- processing, thermal transport, and compatibility/mass transport; particularly for LLO. The 83Pb-17Li eutectic alloy is a candidate breeder material which inherently provides neutron multiplication by the lead. This alloy melts at 235°C. The current design proposal provides for normal operation in the solid phase with subsequent melting during off-periods for tritium recovery. Key features of

38 the Pb-Li alloy include low tritium inventory and potential reactor relevance. Major concerns relate to tritium containment and recovery, compatibility with the structure and mechanical problems associated with the melting/solidification process. Considerable date base on the alloy has been developed in the last few years. The aqueous lithium salts have been proposed as a candidate breeder material. Solutions of LiOH and LiNCs have been evaluated in more detail. The primary feature of these materials is the flexibility associated with a liquid breeder, such as the ability to remove or replace the breeder easily. Primary concerns relate to corrosion/compatibility which is partly associated with radiolysis, safety in particular for the LiNO-i, tritium recovery issues. The data base for these materials is generally quite limited. Beryllium and lead are considered as neutron multipliers to enhance the tritium breeding performance. Beryllium is the most effective neutron multiplier. Key features associated with the use of beryllium include its high thermal conductivity, low density, and low activation properties - both short term and long-term. Major concerns related to the use of beryllium include its chemical toxicity, tritium retention characteristics, radiation-induced swelling and embrittlement, and cost. A considerable data base exists for beryllium in various forms. Lead is currently proposed as a neutron multiplier only in the Pb-Li alloy and is discussed above. Key features are its low cost and ease of fabrication. Major concerns relate to its high density, low melting temperature and activation products. An extensive data base exist for lead.

IV.3. ELECTRICAL INSULATORS

Electrical insulators are required in both the blanket and the divertor to reduce disruption induced electromagnetic loads to acceptable levels. Ceramic insulators are proposed because of their superior radiation damage resistance compared to organic insulators. Even so, the radiation effects are a critical issue for the ceramics insulators. The oxides, e.g., ALO-i and MgAUO^, are generally proposed as candidates. The oxides exhibit excellent insulating properties, are highly stable and readily available. However, radiation effects, particularly swelling and radiation induced conductivity are major concerns.

REFERENCE

(1] Part A of this Report V. BLANKET AND SHIELD SEGMENTATION

Several nuclear components are integrated inside the vacuum vessel, the main components are the first wall, tritium breeding (driver) blanket, shield, divertor plates, plasma heating and current drive ports, diagnostics ports, maintenance ports, testing ports, twin loop copper stabilizer, and active control coils. In-vessel integration of these components is shown in Figs.V-1 and VI-2. The inboard blanket is divided into 32 segments; each segment has two radial electric breaks along the breeding zone to minimize disruption-induced electromagnetic effects. The inboard blanket segments are joined into a ring with help of adjustable screws welded between each other with access through the gap between the segments. This ring is supported on the vacuum vessel at the bottom. The outboard blanket is divided into 64 segments of two types: 32 poloidal side segments placed on either side of an equatorial port, and 32 shorter central segments placed above and below the port. The outboard blanket/shield segments also incorporate passive copper loops for stabilization of the plasma vertical position. The active control coils are integrated with the vacuum vessel. Blanket/shield segments are fixed to the vacuum vessel. Attachment locks are described in Ref. 1. The coolant is supplied to the blanket segments at the top region except the central lower one, to which two options of coolant supply can be used: from the bottom or from the equatorial port. The lay-out of the coolant pipes uses the laser-beam for welding/cutting from inside the pipes. The pipes layout is described in Ref. 2. The inboard vacuum vessel design features a steel/water body backed by a thin layer of Pb/B^C. The shield is attached to the 300 mm thick vacuum vessel, which provides additional shielding. Thinner shielding is provided behind the divertor plates and near the NBI ducts on the back side of the outboard TF leg. The thickness of the outboard blanket/shield/vacuum vessel is designed to permit personnel access outside the cryostat one day after shutdown. Also the different penetrations and assembly gaps are shielded to satisfy the same criterion. The shape of the divertor plate is optimized to minimize infringement on the shielding space behind it. The inclination of (he divertor surface to the separalrix is 15 degrees at the outboard strike point and 45 degrees at the inboard strike point. The poloidal distances from the X-point to the strike points are respectively 1.5 meters and 0.6 meters. To improve the divertor vacuum pumping performance at the bottom part, the blanket segments have "noses" shading the pumping duct. The ITER design provides 16 large equatorial ports. Five of these ports are occupied by the plasma heating and current drive systems. The balance are reserved for plasma diagnostics, plasma fueling, maintenance equipment and nuclear test modules. Maintenance of all in-vessel components will be possible

41 SECTION UNDERNEATH TFC INBOARD suteer

i.J OUfBOAftO SIDE 81AMIT SEGKN1

Fig. V.I Vertical cross-section of different nuclear components

42 SEC7ION BETWEEN TFC

OUTBOARD CENTRAL UPPER SEGHENT

VERTICAL PORT

HORIZONTAL PORT

OUTWARD CENTRAL LOWES SEOMENT

43 UPPER DLUG

OUTBOAPfj CFWTRAl. 'JPPfS

i NBiM RP BLANKET SEGMENT

CENTRAL LOWER BLANKET SEGMENT

LQUEP Df'/E^'OP PL4Ft L OWF v P| UG

Fig. V.2 Isometric view of First wall/blanket/shield sector using fully remote methods. The maintenance approach and time required for completion will, however, vary With the type of component. Regular maintenance and/or replacement of the in-vessel plasma-facing components (divertor plates and first wall armor) is planned. This type of maintenance operation will be accomplished without warm-up of the magnet systems by using in-vessel transporters and manipulators inserted through four equatorial ports. These in-vessel operations will be performed in an inert-gas atmosphere to preclude oxygen contamination of the first wall. It may also be necessary to replace entire blanket modules, but the frequency of such operation is expected to be low enough that warm-up of the

44 magnet system and venting of the torus and cryostat vacuums are acceptable. Removal and replacement of blanket segments will be accomplished through the vertical access ports at the top using dedicated handling devices from above the reactor. Minor in-situ repairs to the first wall/blanket/shield segments can also be accomplished, without removal, using the in-vessel service equipment. The same blanket and shield components are designed to operate in both the Physics and Technology Phases without replacement. The divertor plates will be replaced several times during the 3 MWa/m operating fluence.

REFERENCES

[1J INTERNATIONAL ATOMIC ENERGY AGENCY, ITER Assembly and Maintenance, IAEA/ITER/DS/34, IAEA, Vienna (1991)

[2] INTERNATIONAL ATOMIC ENERGY AGENCY, ITER Contain- ment Struct-ies, IAEA/ITER/DS/28, IAEA, Vienna (1991) VI. BLANKET DESIGN DESCRIPTION

Four designs for the ITER driver blanket have been developed and the performance evaluated. Three designs are for the first option ceramic breeder concept: a layered configuration, a pebble bed design, and a breeder in-in-tube configuration. One design has been developed for the alternate Pb-Li breeder concept. Details of these designs are described in the following sections.

VI.l LAYERED CONCEPT

A ceramic-breeder water-cooled blanket concept has been developed for ITER based on a layered configuration. Lithium oxide (Li2O) or lithium zirconate (LUZrOo) is selected as the primary and the backup breeder materials, respectively. Austenitic steel (type 316 solution annealed) is the reference structural material, which is selected on the basis of an extenstive database and ease of fabrication. Low pressure (~ 1.0 MPa), low temperature (60 to 100°C) water is the coolant. The desire to achieve a tritium breeding ratio close to unity with limited breeding volume because of inboard shielding requirements, numerous penetrations, and provisions for nuclear testing, requires the use of beryllium as a neutron multiplier. A detailed parameter list is given in section VI.5. The first wall, blanket, and shield are integrated into a single unit with two separate cooling systems. Poloidal and toroidal coolant flow are chosen for the inboard and outboard first wall and blanket, respectively. Two fabricated forms are considered for breeder and multiplier materials: sintered blocks and pebbles (about 1 mm diameter). The breeder is highly enriched (95% Li). The use of high lithium-6 enrichment reduces the breeder volume required in the blanket and consequently the total tritium inventory in the blanket. Also, it reduces the temperature gradient in the solid breeder material which increases the blanket capability to accommodate power variation. The developed design can operate continuously up to 150% the nominal power without violating the different design guidelines, in particular the temperature limits for the different materials. Also, it can operate up to 200% of the nominal power for a limited fluence at the beginning of life. Three versions of the layered blanket design have been developed for ITER; the three versions have the same mechanical design. The only difference among the concepts is in the fabricated forms of breeder and multiplier materials. All the concepts have beryllium for two functions: neutron multiplication and breeder temperature control. Helium gaps and insulator materials are not used to control the breeder temperature. The first version has sintered blocks for both the multiplier (60 and 85% dense) and the breeder (80% dense) materials. The second version uses breeder pebbles and beryllium blocks. The last version is similar to the first except for the first and the last beryllium

47 zones. A thin layer of beryllium pebbles is located behind the first wall and at the back of the last beryllium zone. The main improvements associated with this third version are reduced beryllium mass (about 25% saving^ and the capability to accommodate larger blanket and/or first wall deformation. The blanket design philosophy is to produce the necessary tritium required for the ITER operation and to operate at power reactor conditions as much as possible. Also, the reliability and the safely aspects of the blanket are enhanced by using low-pressure water coolant and the separation of the tritium purge flow from the coolant system by several barriers. The other criteria used to guide the design process are mechanical simplicity, predictability, performance, cost, and minimum R&D requirements. The inboard blanket has a single breeder zone embedded in a beryllium zone. The poloidal coolant of the first wall (FW) anc tlic shield behind the blanket are used to cool the breeder region by conducting the nuclear heating to these coolant zones. This results in a simple design. The outboard blanket has two (or three) breeder zones with toroidal coolant flow, which improves the performance and the mechanical design of the blanket. An additional coolant panel (or two coolant panels for the three breeder zones) is used in the beryllium zone between the two breeder zones to get the appropriate temperature profile for the blanket materials. The blanket is designed to accommodate the plasma disruption conditions without exceeding the stress limits for the Type 316 austenitic steel. Each breeder zone is purged by He with 0.2% H2 for continuous tritium recovery. The blanket is designed with separate helium purge loop for the beryllium mulitplier.

VI.1.1. Outboard Section

An isometric view of the outboard blanket (OL) is shown in Figure VI. 1- 1 with two breeder zones and seprate cooling loop for the first wall. Figure VI.1-2 shows a poloidal cross sectional view of a side module with cross sections at midplane and at the upper extremity. It should be noted from Fig. VI.1-2 that the multiplier zone thickness increases from m

48 Me

Fig. VI. 1-1 Isometric view of layered ceramic breeder blanket design with toroidal cooling with LLO breeder and Beryllium neutron multiplier in the form of sintered blocks. UUIL, I

Fig. VI.1-2 Poloidal cross sectional view of a side segment of the layered ceramic breeder blanket.

50 The breeder consists of Li~O blocks, 8 mm thick, clad in 1 mm thick SS sheets to form a panel. These panels are continuous inside the blanket module and have built in manifolds on the sides running in the poloidal direction. Purge gas flows poloidally through the manifold, then toroidally across the panel through machined grooves and finally back out through the return manifold. This purge gas carries with it the tritium which is bred in the breeder. The FW consists of a 1.4 cm thick plate with built-in rectangular channels, 0.4 cm x 2 cm separated by 0.3 cm walls. It is assembled from two extruded SS plates, each with one half channel embossed on one side. The two sheets are assembled with the channel wall separations making contact and continuously roll bended, thus forming the plate with the built-in channels. Section C-C, D-D and E-E of Fig. VI.1-2 show a cross section of the first wall. The plate orientation is such that the channels run along one side wall, then toroidally through the FW and back along the other side wall. This insures good cooling of the side walls and the FW. The FW also has blind holes drilled and tapped for attachement of graphite armors used in the physics phase operation. Section D-D of Fig. VI.1-2 shows the FW followed by a single blanket coolant panel and section E-E, by two blanket coolant panels. These coolant panels are made the same as the FW but are only 0.6 cm overall thickness and have 0.2 cm x 2 cm coolant channels separated by 0.2 cm thick walls. The panels are welded to oblong supply and return manifolds which extend poloidally the full length of the module and are reinforced by through studs to prevent transfer of water pressure to the blanket components. The next two steel plates behind the supply manifold play an important role in the blanket design. These plates which are 3.2 cm and 6 cm thick, respectively, and are separated by a 3 cm gap; are continuous top to bottom and are welded to the module side wall all around. They serve three main functions: (1) The first plate completely seals the breeding blanket zone from the FW coolant manifolds. (2) The 3-cm space between the two plates defines the supply and return manifolds for the FW coolant as shown in Fig. VI.1-2. (3) The two plates act as structural elements tying the two sidewalls of the blanket module together and effectively creating a box for the breeding zone materials. The poloidal extent of the OB breeding blanket is about 4.1 m above the midplane. The acfual first wall extends somewhat further. These zones, which extend beyond the breeding blanket, consist of steel plates and cooling panels. However, the sidewall and first wail of the sector are the same. Figure VI. 1-3 shows severai views of an upper central module.A side view with the sidewall removed does show the internal details of the blanket. Section A-A is a cross section near the lower end of the blanket (nearest to the penetration) and section B-B of the upper extremity. In contrast with Fig. VI.1-2 the back of the module extends straight out, between the TF coils. The side module is needed so that it can be inserted circumferentially into the bore of the TF coil.

51 to

5/DE HCW CROSS SECfJON

Fig. VI. 1-3 Poloidal cross sectional view of an upper central segment of the layered ceramic breeder blanket. The FW manifolds distribute the coolant poloidally from which individual channels carry water toroidally across the FW, collecting into the return manifold and exiting the module through the return connecting pipe. The OB blanket has similar manifolding; however, as the water exits the return manifold, it is directed through the coolant channels in the shield before coming out of the module through the return manifold.

VI. 1.2 Inboard Section

The inboard blanket (IB) has three differences from OB: (1) IB blanket has poloidal coolant flow, (2) IB blanket has only one solid-breeder zone, and (3) The IB blanket is divided into 32 toroidally equal segments; each IB segment is subdivided poloidally into three electrically insulated subsegments. Figure VI. 1-4 is a side view of an IB module with cross sections at midplane and at the top extremity (Z= 3.4 m). The FW, side walls and blanket coolant panels are fabricated the same as in the OB blanket; however, the coolant channels run in the poloidal direction. The radial build of the IB blanket is smaller at midplane than at the extremities. Water and purge gas connections are all at the top. To reduce disruption effects, each module is subdivided into three parts electrically insulated from each other as shown in Fig. VI.1-4. The insulated zone extends 27 cm at midplane and 53 cm at the extremities. The three parts of the module are then E-beam welded together in the back and then bolted to a common shield backing. The breeder and the Be zone are purged with He gas. The FW is 1.5 cm thick and has 0.5 cm x 3.48 cm coolant channels spaced 0.2 cm apart. These spaces are increased in some places to allow room for fasteners needed to attach graphite armors used in the physics phase operations. Side walls are 1.0 cm thick and have 0.3 cm x 3.8 cm channels spaced every 1.2 cm. The FW coolant is divided into two parts and each flows through one side wall to the bottom of the subsegments, where the water flows back through the first wall. The blanket coolant supply connecting pipe feeds water simultaneously to the two extreme rear channels in the rear of the shield. The water in the rear channel flows down through the shield in the poloidal direction, then makes a transition into each of the three boxes at the bottom and flows back up through the blanket coolant panel. The water in the second to last channel also flows down through the shield then makes several transits through the shield ending up at the top.

VT.1.3 Copper Stabilizer Integration

The twin loop copper stabilizers are located on the upper and lower third of the side module, starting at z = 2.7 m extending to 4.3 m from the midplane at the first wall and to 5.0 m at the rear. The loops are in the form of 0.5-cm thick copper plates which enclose the blanket segment. Section B-B of Fig. VI.1-3 shows the copper plates on the outside of the side walls but on the

53 Fig. VI. 1-4 Poloidal cross sectional view of an inboard segment of the layered ceramic breeder blanket.

54 inside of the first wall. The plates are bonded to the blanket wall structure and thus do not require separate cooling.

VI.1.4 Penetration Accommodation

Radial ports have 1 to 1.3 m toroidal extent and 3.4 m poloidal extent. The central blanket module is split into an upper and a lower module, providing the space needed for the radial port at midplane. The upper module has service connections at the top and the lower module at the bottom. The difficult penetrations to accommodate are the three neutral beam ducts, which come in tangent to the circumferential centerline of the plasma. Each duct cuts across two side modules and one central segment as shown in Fig. VI. 1-5. The neutral beam duct is 0.8 m wide and extends 3.4 m in the poloidal direction. The layered concept with toroidal coolant flow can readily accommodate such difficult penetrations. The side segment modifications are straightforward as shown in Fig. VI.1-5. Two ways can be used to modify the central segment. (1) Tunnel through the segment at the midplane, with the upper and lower segments attached to each other by the two triangular segments, and duct coolant and purge gas through these triangular segments. In this way, the service connections remain at the top. (2) Split into an upper and lower segment with one sealed off triangular segment connected to one and the second to the other. In this case, the lower module will have service connections from the bottom. When the two halves are assembled in the reactor, *b?y appear continuous. In either case, the triangular connecting segments will have to be fabricated from PA' panels, since they would be exposed to surface heat loads from the plasma. Figure VI.1-6 shows the two ways to accomplish this.

VI. 1.5 Fabrication and Assembly

In the present design, the FW and the blanket panels are made of plates with flat sides which have coolant channels built in. Such panels can be fabricated as shown in Fig. VI. 1-7. Sheets of stainless steel are hot rolled or extruded with the imprint of one half of the coolant channels on one side. The sheets are continuously spot welded across the channel separations to form the complete channels. The panels are then bent into the proper shapes needed to form a segment. A segment can be made out of several segments and then E-beam welded together into a complete unit as long as the weld does not cut across any coolant channels. Such procedures are possibel based on the current technology (discussion at a meeting with technical representatives from Dean Products Inc. of Brooklyn, New York, U.S.A.). The panels will be shaped by drawing and bending operations, but care must be exercised to insure that the channels remain properly directed. Figure VII.1-7 shows a sequence of operations resulting in a completed side module.

55 '••/A

Layout of 0 Nfutrol B*o» lub* at ai

• Vrert "id'C

Fig. VI. 1-5 Blanket segments for neutral beam penetration

56 BLANKET COOLANT ANO Bo PURGE SUPPLY ANO RETURN SHUNT LINES

FW COOLANT ANO LI2O PURGE SUPPLY ANO RETURN SHUNT LINES

SECTION A - A

FRONT VIEW SIDE VIEW FRONT VIEW SIDE VIEW

CONTINUOUS MODULE SPLIT MODULE Fig. Vi.1-6 Central module modification for neutral beam penetration 00

O EXTRUDED SHEETS ARE SEAM WELDED © CUT AND FORM SECTIONS OF MODULE © EBEAM WELD SECTIONS TOGETHER TOGETHER TO FORM PANELS WITH BOX. DOTTED LINES INDICATE DIREC- BUILT IN CHANNELS TION OF CHANNELS

© CUT, BEND AND ASSEMBLE FW STIF- ASSEMBLE STIFFENERS. AND INNER SEAL WELD EDGES AND MACHINE FENERS. CUT AND SHAPE INNER AND OUTER COPPER PLATES OVEN SQUARE DRILL HOLES INTO FW COPPER PLATES BRAZE ASSEMBLY COOLING CHANNELS

(a) • 1 J •)•••<•

(7) STACK Be BLOCKS. BREEDER PLATES, COOLANT PANELS AND STEEL BLOCKS

(Tj) FORGE SECTIONS OF BLANKET REAR STRONG BOX. MACHINE (?) INSERT BLANKET COOLANT MANIFOLDS AND EBEAM WELD TOGETHER BRAZE COPPER PLATES WHERE ONE ATA TIME AND BRAZE TO PANELS NEEDED. MACHINE EDGES FOR A PERFECT FIT TO THE FRONT BREEDING BLANKET

@ ASSEMBLE SHIELD PLATES TOGETHER INTO THE STRONG BOX

©INSERT REAR STEEL PLATE FOR FW ^) OVERTURN FRONT BREEDING BLANKET AND PLACE ON THE COOLANT MANIFOLD AND SEAL WELD REAR STRONG BOX E-BEAM WELD TOGETHER TO MODULE SIDE S ALL AROUND

(b)

Fig. VI. 1-7 Blanket Fabrication and Assembly steps

59 Because the end segments of the IB blanket module get narrower toward the back, they must be assembled In steps as shown above.

The three segments are then welded together and the assembly bolted to the back steel structure. Finally the rest o1 the shield Is assembled and the last plate welded

(c)

The IB module is also made of the same kind of panels. Because of the need for segmenting the module into three electrically insulated parts and because the coolant channels are running poloidally, it was decided to make the individual boxes out of E-beam welded segments. The cross sections A-A and B- B of Fig. VI. 1-7 show how this is accomplished and the sequence needed for assembling an IB blanket module.

VI.2. PEBBLE CONCEPT Japanese design efforts of ceramic breeder blankets have been concentrated on the two types of the pebble bed blankets; (1) Layered breeder/multiplier pebble beds, (2) Mixed breeder/multiplier pebble bed. The incentives of using small pebbles are:

60 1) toughness against thermal cracking 2) good predictability of thermal performances of pebble bed 3) void space preparation for volumetric expansion of pebbles 4) easy packing process of pebbles into blanket box. The design of the two blankets is summarized below. (1) Layered pebble bed concept Cross-sectional views of the layered pebble bed blanket are shown in Fig.VI.2-1. The major design parameters are summarized in Table VI.2-1. The first wall integrated with the blanket has rectangular coolant channels which are oriented toroidally and poloidally for the outboard and the inboard blankets, respectively. A detailed parameter list is given in Sec. VI.5. In the blanket box, beryllium pebble beds and breeder (LLjO) pebble beds are alternately arranged. There exist six and two layers of the breeder in the outboard and inboard blankets, respectively. Packing ratio of 60 % is conservatively assumed for both the breeder and beryllium beds. The 5 % Lithium-6 enrichment is required for tritium breeding performance. Cooling panels with rectangular coolant channels are located in the beryllium bed and are oriented in the poloidal direction. Breeder pebble beds are clad by 1 mm thick stainless steel to avoid direct contact with beryllium pebbles. The beryllium layer works as a thermal resistant layer for the breeder. The minimum temperature of the breeder (450°C during normal operation in the technology phase) is maintained by the thickness of the beryllium layer, and the maximum temperature of the breeder (800°C during normal operation in the technology phase) is kept by the breeder layer thickness itself. Thicknesses of the breeder and beryllium layers increase in the radial direction in order to accommodate the attenuation of nuclear heating rate in the blanket. The thicknesses also vary in the poloidal direction in order to accommodate the poloidal variation of the neutron wall load. Thus the total thickness of the blanket module varies in the poloidal direction. The minimum thickness of the outboard blanket is 56.2 cm including the first wall and the back wall at the midplane, and the maximum thickness is 86.3 cm at the top/bottom ends of the blanket. Temperature of beryllium pebbles is designed to be kept below 500°C and temperature of breeder clad (316SS) to be around 500°C during the normal operation. Tritium generated in the pebble bed is recovered in the low pressure (0.1 MPa) helium purge gas with protium swamping, which flows poloidally in the ptbble beds. Coolant flow scheme in the outboard blanket segment is shown in Fig. VI.2-2 and VI.2-3. Water coolant connections are provided at the top except the lower central segment for which the coolant connections are provided at the bottom. In the breeding region, coolant flows poloidally with an inlet manifold behind the breeding region. In the outboard first wall, coolant flows toroidally wiht inlet and outlet manifolds behind the breeding region. Coolant flow in the inboard segment is oriented along with the coolant of the breeder region because of the limited space to provide manifolds and the shielding requirement of a large steel fraction behind the breeding region. In the fabrication process, poloidally segmented units of the first wall and the side wall are manufactured by three-dimensional joint technique using HIP bonding as shown in Fig.VI.2-4 and the segmented units are integrated into one module by EB welding. The breeder cans are mechanically attached to the blanket side wall, the cans are packed with pebbles in advance. This mechanical attachment allows thermal expansion of the cans. Cooling panels are also fixed to

61 INBOARD 1/32 FIRST WALL/BLANKET OUTBOARD LATERAL 1/48 FIRST WALL/ SEGMENT (SEC. A-AJ Scdle 1:8 £%«%„ BLANKET SEGMENT (SEC. B-Bl Sc3le 1:10 f: .1 7( •'

It '•".•; x 7 "«sa • fi';r *i' • •» • BREEDING Rl G-SN ' ———•••'y.y;7/V'//y',^ •/•.- • _Z^, ;.0(!:o\- M;VH,;,.; ., \ COOLANT MAN'> •?, wk//. •>. <-i i i I i ».

J—77 , -0Uf80

B?r^n5" BREEDING 7"F'T «='0N—.

£E M/^ C DETAIL D DETAIL E Scale 1: 5 Scale 1: BO Scale 1:1 Fig. VI.2-1 Layered Pebble Bed blanket conceDt TABLE VI.2-1 MAJOR DESIGN PARAMETERS OF LAYERED PEBBLE BED BLANKET Breeder :Li~O Form : petible ( < 1 mm diameter) Density : 85-95 T.D.* Li enrichment :50% Clad thickness : 1 mm Neutron multiplier :Be Form : pebble ( < 1 mm diameter) Density : 100 % T.D. Pebble packing fraction :60% Cooling channel : poloidal cooling panel with rectangular channel Panel thickness : 10 mm Channel size : 4 mm x 30 mm SS thickness : 3 mm front layer, 3 mm in rear layer Blanket thickness : 56.2 cm outboard midplane 86.3 cm outboard top/bottom 15 cm inboard midplane including first wall and back wall Temperature control of breeder : thicknesses of Be pebble bed and LL>O pebble bed layers Breeder max./min. temperature : 450-800 °C(Normal operation) Accommodation to power variation (Technology phase) : 25 % increase - Tritium breeding ratio (Technology phase) Outboard local : 1.35 at midplane sections 1.46 at upper and lower Inboard local : 0.54 at midplane Outboard net :0.72 Inboard net :0.08 Total net :0.80 Coolant : water Inlet pressure : 1.5 MPa Inlet/outlet temperature : 60/100 °C Velocity : < 3.5 m/s Pressure loss <0.3MPa Tritium recovery He purge gas Purge gas flow rate 250 NuT/h P ~ ure 0.1 MPa > ^ure loss lOkPa inydrogen addition : 1 % (H/T = 100) Tritium inventory in Breeder (Technology Phase) 117 g (Outboard), 9 g (Inboard)

* : Analyses of tritium breeding ratio, temperature distribution, tritium inventory etc. were performed for 85 % T.D..

63 BLMKil SHIELD

Fig. Vl.2-2 Coolant Flow scheme of outboard side segment IT'!

v i

I -

Fig. VI .2-3 Coolant flow scheme of outboard central segment Machining. Bending

First Wall Assembly

Fig. VI.2-4 Assembling Process of First wall/side wall unit

the side wall with welding. This process is performed iJternately. The beryllium pebbles are packed in the blanket box from the top with vibrating on a shaking table at the final stage before the coolant inlet header and the top wall are assembled. (2) Mixed pebble bed concept Cross-sectional view of the mixed pebble bed blanket at the outboard midplane (side module) is shown in Fig.VI.2-5. The major design parameters are summarized in Table VI .2-2. The first wall concept is the same as that of the layered pebble bed blanket. Homogeneously mixed beryllium and breeder pebbles are filled in the blanket box. The mixing ratio of LLO/Be was determined to be 1/3 for tritium breeding performance. Packing ratio of pebbles is expected to be 60%. The Lithium enrichment is unnecessary because of the good neutron economy to obtain the equivalent net TBR to that of the layered concept. Therefore the natural lithium is used in this concept. For the outboard side module, breeder region is poloidally separated into three zones by the intermediate coolant plenums as shown in Fig.VI.2-6 in order to accommodate the poloidal variation of the neutron wall load by changing number of cooling tubes in three zones. For the outboard center module, the blanket is originally divided into two parts: the top part and the bottom part, and no blanket structure exists at the midplane to accommodate large duct penetrations.

66 INBOARD 1/32 FIRST WALL/BLANKET OUTBOARD LATERAL 1/48 SEGMENT I SEC. A-A) Scale 1:8 i^V- BLANKET SEGMENT (SEC. B-Bl Scale 1:10

l ««v>; is ill \ -f 11/ !i i| ]tv:

:' .VLi S-if

"0"[ (No CHANNLL OKI I OING REGION *• (00/. (NO MANIFOLD-

Ml f

IN. I 1 DETAIL C DETAIL D DETAIL E Scale 1:5 Scale 1:20 Scale 2:1 Fig. VI .2-5 Mixed Pebble bed blanket concept TABLE VI.2-2 MAJOR DESIGN PARAMETERS OF MIXED PEBBLE BED BLANKET Breeder :LUO Form : pebble ( < 1 mm diameter) Density : 85-95 T.D.* Li enrichment : natural Neutron multiplier :Be Form : pebble (< 1 mm diameter) Density : 100 % T.D. Pebble mixing ratio : Li-O/Be = 1/3 Pebble packing fraction : ~ 60% Cooling channel : poloidal circular tube Outer/inner tube radius : 15/13 mm Blanket thickness : 60 cm outboard, 15 cm inboard including first wall and back wall

Temperature control of breeder : He gap and cooling tube arrangement Gap width : 1-3 mm around cooling tube Breeder max./min. temperature : 400-820 °C (Normal operation) Accommodation to power variation(Technology phase) : 20 % increase Tritium breeding ratio (Technology phase) Outboard local (poloidal) : 1.45 at midplane 1.49 at top and bottom ends Inboard local (poloidal) : 0.57 at midplane Outboard net :0.73 Inboard net : 0.08 Total net :0.81 Coolant : water Inlet pressure : 1.5 MPa Inlet/outlet temperature : 60/100 °C Velocity :_<3 m/s (outboard),< 1.5 m/s (inboard) Pressure loss :_<100kPa (outboard), < 20 kPa (inboard) Tritium recovery : He purge gas Purge gas flow rate : 240 NnP/h Pressure : 0.1 MPa Pressure loss : 0.8 kPa Hydrogen addition : 1 % (H/T = 100) Tritium inventory in Breeder (Technology phase) : 240 g (Outboard), 18 g (Inboard) * : Analyses of tritium breeding ratio, temperature distribution, tritium inventory etc. were performed for 85 % T.D..

68 |First/Side Waii] |Tube Bundle]

Outlet Tubesheet

Top Unit

Middle Unit

Bottom Unit

Inlet Tubesheet EBW Fig. VI.2-6 Poloidal subdivision of outboard side module and assembling process of blanket vessel and tube bundle (mixed pebble bed blanket)

To remove the heat generated in the breeder/multiplier pebble bed and keep the breeder temperature below the nominal maximum temperature, circular cooling tubes oriented poloidal:/ are arranged in the pebble bed region according to the nuclear heating rate. Thermal resistant layer (gas gap) is provided around the cooling tubes and on the blanket box wall by liner tubes or liner wall in order to keep the breeder temperature above the nominal minimum temperature. Thickness of the gas gap is changed poloidally and radially (e.g. 1 to 3 mm) to accommodate variations of nuclear heating rate and tube pitch. Temperature of the pebble bed zone during the normal operation of the technology phase was calculated and it was found that the temperature falls within the range of 400 to 820 °C. Tritium generated in the pebble bed is recovered in the low pressure (0.1 MPa) helium purge gas with protiuai swamping, which flows in the pebble bed poloidally. The first wall and the side wail of the blanket box are fabricated in the same manner of the layered pebble blanket. Liner walls are attached on all the inner surface of the blanket box by the spacer pins made of thermal insulating material. The blanket box and the tube bundle which includes tube sheets and intermediate plenums are assembled by EB welding as shown in Fig.VI.2-6. The breeder and multiplier pebbles are packed in the blanket box from the holes in the back wall with vibration on a shaking table in inert gas atmosphere.

69 VI.3. BIT CONCEPT

The BIT concept is based on a modular configuration of the blanket. Each blanket module is equipped with its own containment structure (not relying on the FW box as primary containment), so that there are no direct interactions between FW and Blanket and no additional loads are induced by the blanket on the FW. Two BIT design variants have been investigated, the first one based on poloidal modules, the second one on toroidal modules.

VI.3.1. Poloidal BIT concept

Design Features The main design features of the Poloidal BIT concept are summarized below: - LiAlC»2 breeder and Be multiplier, both in form of sintered pellets are contained in steel cladding, to allow for possible cracking. Be temperature is kept low (< 120°C) during operation to reduce the irradiation swelling. Use of other breeder materials is not excluded. - Poloidal water cooling (P = IMPa, Tin/out = 60/92°C) matches the general machine lay-out for vertical access ports. Further advantage of the poloidal cooling is the possibility to accommodate natural circulation of coolant.

- On line tritium recovery is performed by He purge stream with H2, O2, H2O or D2O. - He gap provides the thermal barrier to keep the breeder temperature at a proper temperature (> 450°C) for tritium recovery, despite the low coolant temperature. - Proper lay-out of the module internals produces a low surface flux (< 13W/cm ) which reduces the uncertainty in the breeder temperature control. - Neutron fluence of 1 MWa/m , possibility of extending the operating fluence to 3MWa/m is under investigation, depending on the uncertainty in the material damages under irradiation. - Allowable power variations are ± 20% the nominal value, without exceedinp the following temperatures: LiAlO^ Tmax/min = 900/430°C, 316SS Tmax 550/100°C (for cladding/structural steel, respectively), Be Tmax < 120°C - Tritium inventory in the breeder material is 5g., Tritium permeation rate into coolant is 0.1 Ci/d

Mechanical Configuration

The blanket modules consist of poloidal tubes (Fig VI.3-1) whose internals (Fig VI.3-2) are: - Central breeder rod (consisting of cylindrical pellets) is contained in a double steel cladding. The second cladding is a cooling pressure tube. - He gap between the first and second cladding provides a thermal

70 A L BLANKET MIDPUNE SECTION B-B

INBOARD 6UNKET MIOPLANE SECTION A-A

Fig. VI.3-1 Poloidal BIT blanket concept !••••( fO(K »0(.

*•',

BI.ANKFT MODULE - DETAIL A

4/ SECTION 0-0

outer clodc

berylliun

pressure l^i

wo tor

:'nd ciodd.r-

spocer

helium gop t its f.) • V: i 4- Fig. VI.3-2 Details of BIT Breeder modules configuration barrier for the breeder material and He purge flow for tritium recovery. He gap is charged in the poloidal direction by varying the first cladding thickness to accommodate the poloidal variation of the nuclear heating. - First steel cladding provides control of the breeder geometry in case of cracking during the operation. Second steel cladding provides containment of the He purge and it is directly cooled by water at low temperature ( < 100°C) to reduce the tritium permeation through the steel cladding. The first cladding (nominal operation temperature = 450-500°C) is made of CW 316SS to avoid swelling in case of extended operation up to 3 MWa/m . The second cladding and the other structural components operating at low temperature ( 100°C) are made ofSA316SS. - Direct contact between the He purge and the breeder material is ensured by proper channels in the breeder itself or - as an alternative option - by microholes in the first cladding. - Be annular pellets are located outside the cooling tube and protected by an additional steel cladding. - Elastic spacers (Inconel x-750-718) are located at discrete positions between the first and second cladding to maintain the He gap thickness. - Differential elongation of the first and second cladding (namely, 500 and 100°C, respectively) accommodated by proper springs and using a proper segmentation of the breeder rod in the poloidal direction.

The outboard tube bundles have a variable pitch in the poloidal direction (mid to top) to allow a constant number of tubes at both the mid and top sections, despite the poloidal variation of the FW cross section (Fig. VI.3-3). The coolant and He purge are supplied from the top, flowing through the front tube bundles and coming back through the rear ones. A two levels tube- sheets manifold (Fig.Vl.3-4) is used at the top and bottom of the segment. The He purg is immediately cooled by water at the outlet section in order to keep the He manifolds at low temperature to avoid tritium permeation. The tube bundles are supported from the top and equipped with 6 guiding giids to avoid vibration induced problems (Fig. VI.3- 5). Design solution for accommodating the horizontal ports and NBI penetration (Fig. VI.3-6) have been investigated.

Fabrication and Assembly Issues A preliminary assessment of the fabrication and assembly procedure of the blanket modules has been performed. Attention has been paid to avoid weldings in high fluence regions, in particular for the CW 316SS cladding. Since leak-tightness of the breeder rod is not required (on the contrary, tritium flow from breeder to He gap must be ensured), a simple mechanical joint can be used for assembling the breeder rod (i.e.: breeder pellet, first cladding, plug, spacer).

73 '//////// >

•''//// / / / / 4; /.••,,/ ,

1 '•

EQUATORIAL SECTION

Fig. VI.3-3 Toroidal BIT blanket concept

74 UPPER 1UBE 5HEFI OEM/I. A

HELIUM--, UAltK 3ERVLLIUM-

BERYLLIUM WA Ff R HELIUM -J BREEDER

LOWER TUBE 5HEE1 WAfER TANK DEJMi B HELIUM TANK

Fig. VI.3-4. Details of Header Design for Poloidal BIT Concept SPACfR

Ol'IfP WAIfR P/PE BRffDING PELlf! BfffHi /UN 1st CUOOING OUTfR ClAOOING He • GAP .J £US!IC 5PACEB 2nd ClADDING GRID SUPPORT GRID STRAP-. I

DETAIL OF 5UPPOR7ING GRID

Fig. VI.3.-5. Details of Grid Design for Poloidal BIT Concept 77 The following assembly procedure has been identified in case of poloidally curved modules: - Assemble the Be pellets in their steel cladding including the outside cooling tube. - Bend (only one curvature radius) the assembled tube. A proper machined shape of the Be pellets is required to allow bending of the cooling tube after assembling the Be pellets. - Asr .mble the breeder rods (breeder pellets, First cladding, plug, gap spacers) - Assemble the breeder rods in the second cladding - Assemble the pre-bent second cladding in the coolant

Investigations are in progress to use straight tubes instead of bent tubes in order to simplify the fabrication and assembly procedure.

VI.3.2 Toroidal BIT Concept

Design Features

The main differences of the Toroidal BIT concept with lespect to the Poloidal one consist of the additional segmentation of the FW/blanket in the poloidal direction (Fig. ^1.3-7), the toroidal orientation, and cooling of the blanket modules (Figs. VI .3-8,-9). The motivations for the differences are the following: - Facilitate the blanket fabrication procedure by reducing the size of the blanket segments. - Improve the reliability of the blanket module ( in particular, with respect to the fabrication and operation of the He gap) by reducing its length (i.e. 0.5 m) and keeping constant the He gap thickness. Further features of the Toroidal BIT concept are: - Use of Orthosilicate or Sital breeder materials, other breeder materials are not excluded. - Possibility of operating Be at high temperature, if it is required for tritium recovery by using an additional He gap to increase the Be temperature.

Mechanical configuration

- Internal configuration of the toroidal blanket module (Fig.IV.3-9) is similar to that of the poloidal BIT concept: - Central breeder rod is contained in a double steel cladding and located inside a toroidal pressure tube.

78 PD &. Ml i: v't'55f:

C:30API, 3LANK£' MODULE

: L n \i INBOARD

VACUUM

Fig. VI .3-7 Elevation view of the toroidal BIT blanket concept

79 SECTION A-A

PURGE GAS OUTLET

FIRST WALL COOLANT SUPPLY

BREEDING ELEMENT

MODULE INTERNAL ' SPACE CONTROL"

DETAIL C

BREEDER COOLANT INLET

BREEDER COOLANT OUTLET PURGE GAS INLET

Fig. VI.3-8 Toroidal cross section of the BIT blanket concept

- Coaxial coolant tubes allow inlet and outlet at the same side. - He gap between the first and second cladding provides thermal insulation for the breeder and allowing He purging flow for Tritium recovery. - He gap between Be and the steel cladding allows high operating temperature for Be material if it is required for tritium recovery. - Square Be pellets are located outside the cooling tube and contained in the outer steel cladding which also acts as primary vessel of the blanket module - Elastic spacers are located at discrete positions to control the He gap thickness.

80 M

[MINI Cl ADDING

1RI1WM BRFlOiNG CERAMICS

COOLANT CHANNEL

VACUUM GAP '/// j

Fig. VI .3-9 Poloidal cross section of the toroidal BIT blanket concept

Coolant and purge He supply are provided by poloidal manifolds located in the shielding region and connected with the toroidal blanket modules by welded joints. Accommodations of the horizontal ports is provided by omitting the equatorial blanket segments.

Fabrication and Assembly Issues

Large emphasis has been put on the supporting and assembly system of the blanket segments. Blanket segments are supported by the shielding plates (Fig. IV.3-9) and attached by a square beam at the center (260x260mm) which prohibits rotation in case of plasma disruption events. The shielding plates are in turn attached and supported by the Vacuum Vessel. A proper brazing technology based on the use of Pb-Cd is being investigated for the brazed joints in high fluence regions, for remote assembly and disassembly of the blanket modules.

VI.4. LITHIUM-LEAD EUTECTIC CONCEPT

DESIGN DESCRIPTION The blanket segment (Fig. VI.4-1) has a box-shaped design with plane side walls, a back wall following the vacuum vessel configuration and a first wall facing the plasma. Adjustments to the vacuum vessel are placed at the back of the segment. Blanket structural material is austenitic SS 04X16H11M3T. The blanket design provides for a double barrier between the plasma chamber, coolant and environment. Coolant and eutectic manifolds are located in the upper part of each segment (Fig. VII-2). The pipes outside the blanket segments have an additional steel jacket to act as a second barrier. Each segment also has a water- stainless steel radiation shield behind the breeding zone, header, and cooling pipes for the different components in the segment.

81 PIPE JGINJ

PLUG

SEEDING ELEMENT

Fig. VI.3-10 Cross section of the breeding module of the toroidal BIT blanket concept

82 ELEVATION

OUTBOARD 6 - B

Fig. VI.4-1 LiPb Eutectic blanket option

83 84 The outboard blanket segment contains eleven breeding channels (Fig.VII-3) divided into three rows and curved around the plasma chamber. The pitch between the channels is changing in the poloidal direction to match the change in the segment width. Three breeding channels are placed in each inboard segment. The tritium extraction is performed in a batch mode outside the reactor. The tritium extraction is necessary after about one week of full burn time. The eutectic is heated and melted before evacuation or replacement by hot gas through the coolant loop. Thermo-mechanical effects of the eutectic on the channel structure during transient conditions is a concern for the design. The proposed option of the cutectic driver blanket channel (Fig.VI.4-2,-3) reduces this concern due to reduction of eutectic/channel thermo-mechanical interaction by melting/solidifying. The main feature of the design is the segmentation of the eutectic channel into individual chambers with connection of free space of each chamber with the lower one by means of an overflow pipe. Free space above the eutectic surface in the chamber significantly reduces thermo-mechanical interaction of melting/solidifying eutectic with channel structure and acts as a mainfold for collecting helium and tritium released during melting. The overflow pipes are used for He flow through the molten eutectic from the top to the bottom to drain the eutectic from the overflow pipes or for in-situ tritium extraction. The channels schemes are shewn in Fig.VI.4-4. The difference between inboard and outboard channels is the water supply scheme: water passes through the inboard channel from the top and exits at the bottom to the collector; for outboard channel water is- supplied at the top to the first row channel, branches at the bottom, enters the two back rows channels in-parallel and exits at the top. The maximum height of the eutectic chambers is five times the diameters (0.9 m) to avoid thermo-mechanical interaction problem; the free space is greater than 5% of the chamber volume to collect the gases. The channel (Fig.VI.4-4) could be manufactured in the following sequence: chamber dividers with overflow pipes and sheath portions are assembled and welded to the central pipe and between each other. This assembly is inserted into the case then the upper and lower nozzles are welded to the whole assembly. Thr last step is to bend the channel to the required curvature. To prepare the channel for operation it must be cleaned by vacuum- thermal treatment and heated above eutectic melting temperature by gas (Fig.VI.4-4, upper scheme). Liquid eutectic is supplied to the bottom chamber, it fills the lower chambers pressurizing the gas and proceeds to the upper chambers through the overflow pipes. The pressure in the lower chamber at the end of filling will achieve 1-1.5 MPa. When the eutectics appear :n the upper pipe, it means that all the channels are full. Then the helium gas is pumped from the top to the bottom through the chambers to evacuate the eutectic from the small pipes. The channel is cooled to solidify eutectic and the heating gas is replaced with water coolant. During operation eutcctic could be solid, partly or completely melted, and it will not cause significant thermal stresses in the channel structure. The realesed gases from eutectic are kept in the free space of the chambers.

85 Fig. VI.4-3 Eutectic channel

86 EUJECJIC FILLING AND FORKING COOLING

•'URGE ..;

3-D ROW

EUTEC7/C AND WATER EyACUAJlON

tVlECHC

3-D ROW

SUPERSEOIW

Fig. VI.4-4 Outboard blanket channels connection scheme

87 To drain the water out of the channels (Fig.VI.4-4, lower scheme) the helium gas should be supplied in the upper coolant pipes, where the water coolant is drained from the bottom pipe. For the external tritium recovery or eutectic draining, molten eutectic can be evacuated from the channel by pushing the eutectic from bottom to the top through the overflow pipes by pressurized gas supplied at the bottom. The proposed channel design has the potential for in-situ tritium extraction. In this case the water coolant should be replaced with He gas, channel is heated until the eutectic is melted. The purge gas (helium) is supplied through the upper pipe, flows through the eutectic to the tritium extraction system until the final tritium concentration in the eutectic is achieved. Then the channel is gradually cooled and helium is replaced with water. To diminish the possibility of surface crust formation when eutectic is kept molten for a long time, some amount of fresh eutectic could be supplied through the upper pipe, forcing eutectic to flow slowly through the system . Polonium production from the eutectic is 60 Ci after two weeks of operation. Polonium will be separated from the eutectic during tritium extraction procedure. The accumulated tritium activity after 2 weeks of reactor operation is about 15 MCi. Since the permissible concentration of tritium is ~ 10 times more than that of Po-210, the biological hazard potential (BHP) from the generated polonium is ~ 40% of that of tritium.

VI.5 SUMMARY OF BLANKET DESIGN PARAMETERS

This section includes tables that contain the parameter lists for the ceramic breeder design options and the lead lithium design option. The design paramters of the ceramic breeder design options are given in table Vl.5-1 Table VI.5-2 lists the design parameters for the lithium-lead blanket.

88 TABLE VI.5-1 DESIGN PARAMETERS OF THE CERAMIC BREEDER DESIGN OPTIONS

PARAMETER LAYERD DESIGN LAYERO PEBBLE BED DESIGN BIT DESIGN INBOARD OUTBOARD INBOARD OUTBOARD INBOARD OUTBOARD BREEDER MATERIAL: Li?0 0 LiAlO, LiAlO., D 2 2 Li ENR1CHM.NT (*) 95 50 50 i FORM: sintered sintered sintered Pebbles sintered Pellets sintered Pellets GEOHETRY: blocks blocks Pebble (604PF) ylindrical Pellets Cylindrical Pel let -, MASS(HT): 2.07 10.53 -o -90 5.5 TEMPERATURE (C) WINDOW: 400-1000 400-1000 400-1000 400-1000 450-900 450 .,1) PHYSICS PHASE MIN.-HAX. TEMPERATURE (C) OURING NORHAL OPERATION: 509-603 502-609 120* NORHAL OPERATION: 611-737 610-744 150* NORHAL OPERATION: 778-961 764-964 TRITIUM INVENTORY (g): 0.42 2.18 TECHNOLOGY PHASE HIN.-MAX. TEMPERATURE (C) DURING NORMAL OPERATION: 452-526 451-537 450-700 450-800 500-7S; 500-/50 120* NORMAL OPERATION: 539-638 541-649 -500-800 "500-900 570-900 570-900 150* NORHAL OPERATION: 677-821 673-833 TRITIUM INVENTORYa (g): 1.94 11.6 ~ 9.2 ~ 117 MULTIPLIER HATERIAL: Be Be Be Be Be FORH: sintered sintered Fused Pebbles Sintered Pellets Sirtered Pellets GEOMETRY: blocks blocks Pebble (604PF) Annular Pellets Annular Pellets HASS(MT): 19 187 "8 "185 15 174.5 TEHPERATURE (C) WINDOW: <600 <600 <600 <600 <600 <:600 PHYSICS PHASE HIN.-MAX. TEHPERATURE (C) DURING NORMAL OPERATION: 172-455 130-471 120* NORHAL OPERATION: 190-558 143-576 150* NORHAL OPERATION: 216-712 164-752 TRITIUM INVENTORY (g): 4.6 22.4 TECHNOLOGY PHASE

MIN.-MAX. TEHPERATURE (C) OURING j A NORHAL OPERATION: 162-415 123-422 "100-500^ "100-700° 60-100 60-100 13 120* NORHAL OPERHION: 178-491 135-512 "100-600° "100-800 60-100 60-100 150* NORMAL OPERATION: 201-620 154-662 TRITIUM INVENTORYa (g): 230 1130 — BREEDER He PURGE GAS IN-OUT PRESSURE (HPa): 0.2 - 0.1 0.2 - 0.1 0.11-0.10 0.11-0.10 0.2 0.2 FLOW RATES (HOL/S): 0.53 2.71 0.31 2.79 2.6 INLET H2 (*): 0.2 0.2 0.817(lvpm) 0 .817 (lvpm) ~~ 1 1 HAX.-AVE. OUTLET HT(HTO) PRESSURE(Pa) 13.3-12.8 18.0-12.3 15.5(0.82) 15.5(0.82) 10-1 10-1 H/T RATIO: 29.7 30.6 100 100 100 100 oo PUMPING POWER (KH): 2.66 13.6 — — 10 TABLE VIi-1 DESIGN PARAMETERS OF THE CERAMIC BREEDER DESIGN OPTIONS (CONT.)

PARAMETER LAYERED DESIGN LAIRED PEBBLE BED DESIGN BIT DESIGN INBOARD OUTBOARD INBOARD OUTBOARD INBOARD OUTBOARD CLAD MATERIAL FOR BREEDER ANO MULTIPLIER SA 316SS 316SS SA 316SS SA 316SS CW316SS CW316SS HIN.-HAX. THICKNESS (im): 1-1 1-1 1 1 0.3-1.0 0.3-1.0 HASS(HT): 2.2 13.8 .4 -45 TEHPERATURE (C) WINDOW: <550 ^550 <500 <500 420^550 420^550 PHYSICS PHASE HIN.-HAX. TEMPERATURE (C) DURING NORMAL OPERATION: "S3-496 475-497 120* NORMAL OPERATION: 584-595 571-607 150* NORMAL OPERATION: 744-755 716-791 TECHNOLOGY PHASE

MIN.-MAX. TEHPERATURE (C) DURING J j NORMAL OPERATION: 431-440 424-446 400-600? 400-680? 480 480 120* NORMAL OPERATION: 514-523 iO6-541 450-700 400-780 550 550 150* NORMAL OPERATION: 647-658 631-693 STRUCTURE MATERIAL (COOLANT CHANNELS. SIDE HALLS. AND FIRST HALL) SA 316SS SA 316SS SA 316SS SA 316SS SA316SS SA316SS TEMPERATURE (C) WINDOW: -400 <400 <400 *400 <400 <400 HASS(HT): 19 187 "97 -500 PHYSICS PHASE MIH.-MAX. TEMPERATURE (C) DURING f NORMAL OPERATION: 95-239 85-278 60^IOOe 60^3"00T 120* NORMAL OPERATION: 98-250 90-293 l!0* NORMAL OPERATION: 103-266 95-326 TECHNOLOGY PHASE MIN.-nAX. TEHPERATURE (C) DURING NORMAL OPERATION: 92-187 77-225 60-170 60-170 60-100 60-100 120* NORMAL OPERATiON: 93-207 78-255 60-100 60-100 150* NORMAL OPERATION: 93-237 79-299 BLANKET/SHIELD COOLANT HATERIAL: H20 H20 H,0 H,0 H20 H-0 DIRECTION: POLOIDAL TOROIDAL PoloidSl Poioidai PoloidSl Poloidal IN-OUT TEHPERATURE (C): 60-100 60-100 60/100 60/100 60/92 60/92 IN-OUT PRESSURE (MPa): 1.50-1.32 1.50-1.45 1.5-1.3 1.5-1.2 1.2/1 1.2/1 MIN.-MAX. VELOCITY (H/S): 0.70-4.35 1.20-2.20 <3.5 <3.5 3-4 3-4 TRITIUM PERMEATION RATE (Ci/D): <0.1

PARAMETERS LAYERED DESIGN LAYERED PEBBLE BED DESIGN BIT DESIGN INBOARD OUTBOARD INBOARD OUTBOARD INBOARD OUTBOARO

GEOMETRICAL DETAILS FW-BLANKET RADIAL THICKNESS (m) MIDPLANE: 15-95 14-261 15-1509 15-5629 170 627 TOP/BOTTOM: 15-171 14-571 15- 15-8639 170 LAYER DIMENSION IN THE RADIAL DIRECTION OR TUBE DINEk-.CNS AND ATJIANGEMENT NIDPLANE: Table VIII.1-2 Table VIII.1-1 Fig.2.2-1 Fig.2.2-2 46 46/64 TOP/BOTTOM: Table VIP.1-2 Table VIII.1-1 Fig.2.2-3 46 46/64 PENETRATION ACCOHHOOATION FOfi NBI: Fig. Fig. PORTS: _ Fig. Fig. TRITIUM BPEEOING PERFORMANCE P"rLE;3 DATA BASE: ENDF/B-V JENDL3 EFF1 .rtANSPORT METHOD: ONEDANT-MCNP CODES . 1-DSN MCND-3B MIOPLANE-EXffiEHITY LOCAL POLOIDAL TBR: 0.755-0.895 1.375-1.310? 0.54- 1.35-1.46 MIDPLANE-EXTREHITY LOCAL 1OROIDAL TBR: 0.206-0.217 0.948-0.966? NET ESTIMATE TBR: 0.14 0.70° 0.05" "0.72 NET TBR BASED ON 30 CALCULATION: 0.12 0.69° 0.05 0.7T NUCLEAR HEATINGc (MW) TOTAL NUCLEAR POWER PHYSICS PHASE: 240 860 TECHNOLOGY PHASE: 179 640 -94 -750 BREEDER POWER DENSITY (HW/M3) MIN.-MAX. AT MIDPLAHE: -30.9 - 58.7 1.5-18 1-24 -25 HIN.-MAX. AT TOP/BOTTOM: 14.8 39.I 0.7-8 0.06-19 - 15 MULTIPLIER POWER DENSITY (MW/M3) MIN.-MAX. AT MIOPLANE: - 3.9 - 6.8 1.4-3 0.17-4 -4.5 MIN.-MAX. AT TOP/BOTTOM: - 1.5 - 3.3 0.62-1.3 0.02-1.8 -2.7 DECAY HEAT AT SHUT DOWN (MW): 14.8

a) Burn time of 2x10 and 1.2x10 s for physics and technology phases without any tritium release. b) Copper stabilizer included. c) Only nuclear heating d) Will be lowered by optimization of layer thickness e) With radiative-coiled tile . f) At local high flux region of 0.6 mW/m with radiative-cooled tile g) Including first wall and back wall TABLE VI.5-2 DESIGN PARAMETERS OF THE LEAD-LITHIUM BREEDER DESIGN OPTION

Parameter Value Parameter Value Parameter Value Li(17)-Pb(83) eutectic mass, T 1000 Li enrichment, % 90 Pressure losses - inboard blanket (IB) 125 Physics/Technology,MPa - outboard blanket (06) 875 Tritium breeding ratio 0.76-0.8 IB segment 0.159/0.10 IB channel 0.15/0.095 IB segment mass, T 40 Number of tubes, from Ob segment 0.25/0.073 IB/06 segments 7/9 First channel of OB 0.14/0.06 06 segment mass, T: Second channel of 06 0.10/0.099 - lateral 80 Maximum tube diameter,mm Third channel of 06 0.10/0.099 - central upper/lower 64/35 - IB segments 80 - 06 segments 120 Eutectic maximum temperature Breeding channels number in in breeding channel IB/OB segments 3/11 Thermal power from the Physics/Technology, °C breeding IB 241/211 Breeding channels diameter, mm Physics/Technology, HH 546/440 first channel of 0B 222/193 - IB/OB 180/190 IB 95/76 second channel of OB 181/169 OB 451/364 third channel of OB 131/124 Breeding zone thickness at Channel thermal power the midplane physics/technology, kW Maximum temperature of channel IB/OB, nm 180/510 IB 990/792 wall physics/technology, °C 0B first row 1500/1208 Ib 115/109 Blanket thickness 06 second row 688/556 first channel of 06 93/90 at the midplane 06 third row 333/271 second channel of 08 117/114 IB/OB, m 490/1100 third channel of 06 105/104 Coolant flow rate, kg/s: Eutectic mass, t - blanket 3266/2581 Total helium flow rate through - IB channel 1.6 - IB segment 18.2/14.4 blanket, kg/s 1-20 - 06 channel, 1-st row/2-nd - OB segment 57.4/45.8 and 3-d rows 2.0/1.65 - IB channel 6.08/4.81 Range of helium temperatures - first channel of 06 8.9/7.1 at the blanket inlet. °C 100/450 Water content in breeding - second channel of OB 4.5/3.6 channels, % - third channel of 06 2.0/1.7 - 1-st row/2-nd and 3-d rows 15.9/35.6 VH. DESIGN ANALYSIS

VII.l. NEUTRONICS ANALYSIS

VII.1.1 Introduction

The terms of reference for ITER require a tritium breeding blanket with a breeding ratio as close to unity as practical to produce the necessary tritium required for the ITER operation and the test program. The neutronics analysis was performed to define blanket configurations that maximizes the tritium breeding ratio and satisfies the other design guidelines. Detailed neutronics analyses were done to find such configurations for the four blanket options considered during the CDA phase. One/two-dimensional discrete ordinates codes ONE/TWODANT, ANISN and multi-dimensional Monte Carlo codes MCNP, BLANK were employed for the neutronics calculations. The cross section libraries used by these transport codes were generated from different nuclear data files. These files are ERF, ENDF/B-IV, -V, ENDL-83, and JENDL- 3. Multigroup libraries are used for all the transport codes except MCNP, it uses a continuous energy representation for the cross sections. Poloidai distribution of the neutron wall loading is calculated to help defining the poloidal change of the blanket dimensions. The total inboard blanket and shield thickness is defined based on the global system studies. This thickness does not permit the use of a full blanket module and an adequate shield thickness to protect the inboard section of the toroidal field coils. Therefore, the inboard blanket thickness is minimized to provide an adequate radial space for the shield. On the outboard region, an adequate radial space does exist to accommodate a full blanket module. Radioactivity analyses were carried out for the blanket to assess the material activation and decay heat source.

VH.1.2 Poloidal Neutron Wall Loading Distribution

The poloidal distribution of the neutron wall loading is calculated by ray tracing and Monte Carlo codes. The starting point for this calculation is the fusion power density distribution over the plasma volume. Accurate specifications of this distribution are not available at this point of time. Different approximations were used to perform the calculation where the fusion power is normalized to the reference values of 1100 and 860 MW for the physics and the technology phases. An example of the poloidal neutron wall loading distribution is shown in Fig. VII.1-1. The distributions resulted from the use of different approximations are in good agreement, the maximum difference in the neutron wall loading values at any poloidal angle is less than 5%. The neutron wall loading values at the blanket extremities are about 0.38 and 0.5 of the values at

93 1.8 _ Physics (1100 MW) ..Technology ( 860 MW)

i • • • i • ' ' i ' • • r 180 160 140 120 100 80 Poloidal Angle (Degree)

Fig. VII. 1-1 Neutron wall loading distribution

the midplane for the inboard and outboard regions, respectively. The blanket is designed to accommodate such change.

VII.1.3 Trititiun Breeding analysis

The tritium breeding analysis performed for the different blanket options are summarized in this section. Calculational methods and nuclear data used in the analyses are described.

1) Layered Blanket Concept

The first wall/blanket/shield design and optimization system (BSDOS) [1] was used to carry out the neutronics and thermal-hydraulics analyses in an integrated manner for the layered blanket concept [2]. BSDOS uses the one/two- dimensional discrete ordinates code ONE/TWODANT [3] to carry out the transport calculations. The analyses used a P^ approximation for the scattering cross sections and an Sg angular quadrature set. A 67-coupled group nuclear data library (46 neutron andzl gamma) based on ENDF/B-V is used for all the one- dimensional calculations. This library is based on VITAMIN-E [4] for the transport cross sections and KAOS/LIB [5] for the nuclear responses.

94 The radial build of the blanket is calculated at six poloidal locations to accommodate the change of the neutron wall loading. The radial build is defined at the midplane and the extremities of the blanket for the inboard and the outboard regions. The other two locations are at the starting point of the copper stabilizer in the poloidal direction. The analyses of two blanket versions based on the layered configuration are be presented. The first version has sintered (blocks) materials for the breeder and the multiplier. The second version uses lithium oxide pebbles and sintered beryllium blocks. The radial build of the first blanket version is defined based on neutronics and thermal considerations. In the inboard section, the beryllium material is used with a 0.65 density factor to reduce the thermal conductivity of the sintered block. This reduces the required beryllium thickness to get the temperature distribution of the solid breeder material within the temperature window. On the contrary, the beryllium material of the first outboard breeder zone has a density factor of 0.85 to get high thermal conductivity. This permits the use of a thick beryllium zone in the front section of the blanket where it is needed from the neutronics point of view. The beryllium material of the second breeder zone has a 0.65 density factor for the same reason similar to the inboard section. The range of the beryllium density factor of 0.65 to 0.85 is defined based on material considerations including swelling and mechanical properties. The calculated radial build of the outboard and inboard blankets are given in Table VH.1-1 and- 2. The change in the beryllium material thickness in the polciutti direction is similar to the poloidal change of the neutron wall loading on the first wall. The local tritium breeding ratio varies from 1.375 at Z = 0.0 to 1.461 at Z = 2.7 m where the copper stabilizer starts. The loss in the local tritium breeding ratio due to the copper stabilizer is 7.7% at Z — 2.7 m. The local tritium breeding ratio at the blanket extrimity is 1.310. The blanket thickness varies from 26.5 cm at the midplane to 58.5 cm at the end. The local tritium breeding ratio of the inboard blanket changes from 0.755 at Z = 0 to 0.895 at the blanket extrimity. The corresponding blanket thicknesses are 10.7 and 18.3 cm, respectively. In the second blanket version, the breeder zone has lithium oxide pebbles. First, the analysis was carried out for breeder zones with single size pebbles. The values of the thermal conductivity of the breeder zone and the gap conductance at the steel clad are low compared to the sintered breeder blocks. At the outboard midplane blanket, this change in the thermal characteristics produces the following effects relative to the blanket with the LLO blocks: a) the total thickness of the beryllium zones is 3.9 cm less, b) the tritium breeding capability is 8.6% less, and c) the breeder fi jiperature gradient is 222°C instead of 84°C for the sintered breeder blocks. To improve the thermal characteristics oi the breeder zones, the material fraction is increased by using two sizes for the pebbles. The resulting blanket is improved relative to the blanket with single size pebbles. The local tritium breeding ratio is increased from 1.257 to 1.331 and the breeder temperature gradient is decreased form 222°C to 143°C. However, the total thickness of the beryllium zones is increased by 2.2 cm.

95 TABLE Vn.1-1 OUTBOARD RADIAL BLANKET CONFIGURATION WITH TWO BREEDER ZONES AT THE MIDPLANE, THE BEGINNING OF THE COPPER STABILIZER, AN THE END OF THE BLANKET

ZONE MATERIAL THICKNESS (cm) (DF) 1.2 MW/tn2 0.958 MW/nr 0.6 MW/m2 Z - 0 1 " 1 2.7 m Z -,+ 4.3 m

First wall layers Tile(a > C 2.0 2.0 2.0 2.0 First wall steel 0.5 0.5 0.5 0.5

Coolant y,2o 0.4 0.4 0.4 0.4 Back wall steel 0.5 C.5 0.5 0 .5 Stabilizer Cu 0.0 0.0 0.5 0.5

Blanket Multiplier Be (0.85) 3.4 4.8 4.3 7.2 Clad steel 0.1 0.1 0.1 0.1 Breeder Li20 (0.80) 0.8 0.8 0.8 0.8 Clad steel 0.1 0.1 0.1 0.1 Multiplier Be (0.85) 5.9 7.9 7.9 12.7 Coolant channel steel 0.2 0.2 0.2 0.2 Coolant H20 0.2 0.2 0.2 0.2 Coolant channel steel 0.2 0.2 0.2 0.2 Multiplier Be (0.65) 5.7 8.4 8.4 19.C Clad steel 0.1 0.1 0.1 0 .1 Breeder Li20 (0.80) 0.8 0.8 0.8 0,.8 Clad steel 0.1 0.1 0.1 0,.1 Multiplier Be (0.65) 7.1 11.6 11.6 15, Coolant channel steel 0.2 0.2 0.2 7. 2(b) Coolant K20 0.2 0.2 0.2 0,.2

Total first wall/blanket thickness 26.5 37.1 37.1 58. 5 Local tritium breeding ratio 1.375 1.461 1.356 1.310

a - The carbon tile is used only in the physics phase analyses and it is not included in the total first wall/blanket thickness, b - A 7.0 cm of the steel in this zone is a part of the bulk shield and it is not included in the total first wall/blanket thickness.

96 TABLE VII.1-2 INBOARD RADIAL BLANKET CONFIGURATION WITH ONE BREEDER ZONE AT THE MIDPLANE AND THE END OF THE BLANKET

ZONE MATERIAL THICKNESS (an) (DF) 0.884 MW/m2 0.325 MW/rri Z = 0 Z = +, 3.4 m

First Wall Tile C 2.0 2.0 First wall steel 0.5 0.5 Coolant H20 0.4 0.4 Back wall steel 0.5 0.5

Blanket Multiplier Be (0.65) 3.3 9.0 Clad steel 0.1 0.1 Breeder Li20 (0.80) 1.0 1.0 Clad steel 0.1 0.1 Multiplier Be (0.65) 4.6 6.5 Coolant channel/shield steel 0.2 0.2 Coolant H20 0.2 0.2

Total first wall/blanket thickness 10.9 18.5 Local tritium breeding ratio 0.755 0.895

a - The carbon tile is used only in the physics phase analyses and it is not included in the total first wall/blanket thickness, b - A 2.6 cm of the steel in this zone is a part of the bulk shield and it is not included in the total first wall/blanket thickness.

Several calculations based on one-dimensional toroidal cylindrical geometry have been performed to determine tritium breeding in both the inboard and outboard blankets. Tlie net TBR has been estimated by coupling the one-dimensional results with the coverage fractions of the different blanket regions. The coverage fraction corresponds to the solid angle fraction subtended by the particular region as seen by the source neutrons in the plasma and represents the fraction of source neutrons going directly to this region. The coverage fractions for the mboard and outboard regions have been determined to be 16.4% and 68.1%, respectively. The actual coverage fraction of the outboard blanket was modified by subtracting the coverage fraction of the sixteen penetrations. This leads to a coverage fraction of 57.8% for the outboard blanket.

97 The effect of the 0.5 cm thick copper stabilizer loops has been considered by performing the neutronics calculations with the Cu in the outboard region. Also, the poloidal variation of the neutron wall loading and radial build were considered, the poloidally averaged TBR was determined to be 0.21 for the inboard region and 1.003 for the outboard region. In the one-dimensional toroidal cylindrical geometry model, 21.2% of the source neutrons go directly to the inboard region with the rest going lo the outboard region. The TBR results of the toroidal calculations were adjusted by the actual coverage of the inboard and outboard regions. In addition, space taken by the assembly gaps and side

IB PLASMA / \

TBR=0.21 ).212 0.78J TBR=1.003

Toroidal Calculation

DIV .155

TBR in IB with TBR in OB with coverage coverage fraction of 57.8% fraction of 16.4% (68.1-10.3) in ITER in ITER (excluding 16 TM, heating, CD, and maintenance ports) 0.16 0.74 I I TBR in IB excluding TBR in OB excluding assembly gaps and side assembly gaps and side walls (11%) walls (4.4%) 0.14 OVERALL 0.70 TBR

Fig. VI1.1-2 Estimation of net tritium breeding ratio for the layered blanket design with two outboard breeder zones walls amounting to 11% of the inboard region and 4.4% of the outboard region were considered. Figure VII. 1-2 illustrates the calculational procedure for the estimated net TBR for the first blanket version. The estimated net TBR is 0.84 with 0.14 contributed from the inboard blanket modules. For the second blanket

98 VERTICAL CRO8S-SECTION

TEST MODULE

IB HORIZONTAL CROSS-SECTION

Fig. VII. 1-3 Cross sections through the three-dimensional geometrical model for the net tritium breeding calculation.

99 TABLE VII.l-3 THREE DIMENSIONAL TRITIUM BREEDING RATIO FOR DIFFERENT BLANKET VERSIONS WITH DIFFERENT DATA BASE

Data Base ENDFB-IV ENDF/B-V Blanket Configurations 3-D Estimation 3-D Estimation 3-U Calculation Two breeding zones with n Li2 blocks 0.90 0.84 0.81 Three breeding zones with Li^O blocks 0.98 0.92 Two breeding zones with LipO pebbles 0.87 0.81 0.78 version with a two-size breeder pebble bed, the estimated net TBR is 0.81. Table VII. 1-3 gives the estimated net TBR for the different blanket variations. Three-dimensional neutronics calculations have been performed to determine the net TBR and the total nuclear heating. The continuous energy coupled neutron-gamma Monte Carlo Code MCNP, version 3B, has been used with cross section data based on the ENDF/B-V evaluation. Because of symmetry, only 1/32 of the reactor was modeled with two vertical reflecting surfaces at azimuthd angles of 0 and 11.25. The blanket segments were modeled in detail with the poioidaUy varying radial builds. The sidewalk and detailed layered configuration of the FW and blanket are included in the model. The sidewalk are 1.4 cm thick in the outboard segments and 1 cm thick in the inboard modules. The vertical extends of the inboard and outboard blankets are -3.4 to 3.4 m and -4.8 to 4.3 m, respectively. The 2-cm-thick assembly gaps between blanket segments and the copper stabilizer loops used in the outboard region were modeled. The first wall, divertor regions, and vacuum pumping ducts were modeled in detail. The divertor plate design consisting of tungsten, niobium, and water, was used. The sixteen standard 1.07 m x 3.4 m radial ports were used at the middle of the outboaid region. A typical Li/V blanket was used in the ports to represent a blanket test module. The test section used in the model is surrounded by a 25 cm thick steel buffer zone. Figures VII. 1-3 shows different cross section of the geometrical model used in the calculations. The neutron source was sampled from the D-shaped toroidal plasma zone whose boundary was determined from the reference plasma parameters. Thirty thousand histories were used in the calculation yielding statistical uncertainties less than 0.8% in the calculated net TBR and nuclear heating values. The tritium breeding results indicate that the net TBR k 0.81 with 15% of it contributed by the inboard blanket. The effect of using different materiak in the radial ports on tritium breeding has been investigated by performing the 3-D

100 calculation with 35% steel, 10% H2O, 5% Cu, and 50% void in the outboard radial port. This represents a LH or RF port. The net TBR calculated in this case is 0.802 implying that tritium breeding in the driver breeding blanket is not sensitive to the material used in the ports. It is interesting that the estimated net TBR given above is only 3.7% different from the value obtained from the detailed 3-D calculation. The total nuclear heating in the reactor excluding test modules is 996 MW, for the technology phase indicating that the total energy multiplication is 1.45. Adding the surface heating implies that the total reactor thermal power is 1150 MW. As mentioned before, all the analyses are based on ENDF/B-V, which results in a lower tritium breeding values compare to ENDF/B-IV. The calculations were repeated with ENDF/B-IV, the calculated difference is about 7% as shown in table VII.1-3. Several neutronics studies were carried out to improve the blanket performance including the tritium breeding capability. The dher objective is to find the blanket performance as a function of the different design parameters. The key parameters for these studies are the number of the breeder zones, the thickness of the breeder zone, the breeder clad thickness, the lithium-6 enrichment, the carbon tile thickness, and the thickness of the first wall materials. Tritium Breeding Enhancement: The first blanket version is reconfigured with three breeder zones and the same minimum breeder temperature. The third breeder zone extends from z = -2.7 to 2.7 m only in the outboard blanket. The local poloidal tritium breeding ratio at the midplane for the new configuration is 1.634 compared to 1.375 for the blanket with two breeder zones, which is about 19% increase. The estimated net TBR is 0.92 as shown in Table VII.1-3. Breeder Zone Thicknesses: The breeder zone thicknesses of the first blanket concept are varied simultaneously from 0.4 to 1.6 cm to study the impact on the local tritium breeding ratio. The results show that the tritium production from the second breeder zone is almost constant. This shows that the thickness of the second breeder zone should be sized based on other considerations. The tritium production from the first zone is increased slightly to reach a saturation value as the breeder zone thickness increases. The local tritium breeding ratio is increased only by 4% when the thickness of the breeder zone is doubled from 0.8 to 1.6 cm. From these results, a breeder zone thick of about 0.8 cm is adequate from the neutronics point of view. Breeder Clad Thickness: The breeder clad thickness is based on the beryllium steel compatibility consideration. The thickness of the reaction layer for ITER conditions is less than 0.01 cm at the end-of-life (3.8 fpy). Therefore, a clad thickness of 0.1 cm is used for the blanket design. The impact of the breeder clad thickness on the tritium breeding capability is calculated. The results show that the tritium breeding ratio is not sensitive to the clad thickness in this design. For example, doubling the clad thickness from 0.1 to 0.2 cm reduces the tritium breeding ratio by 2.6%. Lithium-6 Enrichment: The blanket design has a 95% lithium-6 enrichment for the solid breeder, which helps the blanket performance in several

101 ways. It reduces the solid breeder and tritium inventories in the blanket. It reduces the impact of the design parameters on the blanket performance such as the clad thickness and breeder zone thickness. Also, the use of a thin breeder zone thickness reduces its maximum temperature. This increases the blanket capability to accommodate a higher neutron wall loading. The analyses show that the triiium breeding in both breeding zones increases fasi and then slows down as the lithium-6 enrichment increases above 60%. First Wall Design: The impact of the first wall design on the tritium breeding is assessed. Up to 2.5 cm thick carbon tile, the decrease in the tritium breeding is about 7% per centimeter of the carbon tile and it is almost Linear. The cor responding value for the steel is about 14% loss per centimeter up to 1.6 cm of steel. The impact of the water-zone thickness in the first wall does account for 21% loss in the tritium breeding capability of the blanket per centimeter of water. These results encourage the use of a thin first wall with low water content. Beryllium Density Factor: The beryllium density factor is varied by 0.05 from the designed values of 0.85 and 0.65 for the beryllium zones. The corresponding change in the tritium breeding ratio is 0.023, which is only 0.11% per percent change in the beryllium density.

2) Pebble Blanket Concept

Neutronics analyses were performed for two blanket versions with pebble materials. The first version is layered configuration where the breeder and the multiplier materials are separated by steel clad. The second version consists of a mixture of beryllium and lithium oxide pebbles. One-dimensional calculations were performed for the two blanket versions to determine the tritium breeding ratio and the nuclear heating. A multigroup library based on JENDL-3 [9] was used for the AN1SN calculations. A Pr approximation for the scattering cross sections and an Sg angular quadrature set were used for the analyses. Two-dimensional calculations were performed to determine the impact of the penetrations, side walls, and the internal ribs inside the blanket box. The radial build of the blanket is calculated at three poloidal locations to accommodate the change of the neutron wall loading. The radial build is defined at the outboard midplane, the outboard extremity and the inboard midplane. The radial build of the first version is defined to operate the lithium oxide in the temperature range of 400 to 800°C. The outboard blanket has six lithium oxide zones, five coolant panels, and twelve beryllium zones The outboard blanket thickness varies from 56 cm at the midplane to 86 cm at the blanket extrimity. The inboard blanket has two lithium oxide zones, two coolant panels, and five beryllium zones. The inboard midplane thickness of the blanket is 15 cm to minimize the nuclear responses in the toroidal field coils. A 50% lithium-6 enrichment is used to enhance the tritium breeding capability. A pebble size of 1 mm is used for the lithium oxide and the beryllium materials. The packing fraction of the pebble bed is 0.6. The lithium oxide density is 0.85 the theoretical density.

102 COOLANT(HJO)

• ^ y SECTION a b A- 15 FIRST WALL B-B 30 SIDE SHELL

DETAIL C DETAIL D

Fig. VII.1-4 Cross-Sectional View of Outboard Blanket (side module at midplane, layered pebble bed type) 836

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Fig. VII.1-5 Cross-Sectional View of Outboard Blanket (side module at the top, layered pebble bed type) DETAIL "A 234

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Fig. VII. 1-6 Cross-Sectional View of Inboard Blanket Submodule (at midplane, layered pebble bed type)

The calculated radial build of the outboard and inboard blankets are shown in Fig. VII.1-4 through VII.1-6. The thicknesses of the breeder and multiplier zones increase in the poloidal direction as the neutron wall loading decreases. The local tritium breeding ratio varies from 1.35 at Z = 0.0 to 1.46 at the blanket extimity without the effect of the copper stabilizer. For the inboard blanket, the local tritium breeding ratio at the midplane is 0.54. The estimated net tritium breeding ratio is 0.80 based on one dimensional poloidal model and neutron coverage. This estimate is aiso account for the side walls, the major penetration ports, and the copper stabilizer. In the second version, homogeneously mixed beryllium and lithium oxide pebbles are used to fill the blanket box. The Be to LUO mixing ratio is 3 to 1. The packing fraction of the pebble is 0.6. Natural lithium is used for this configuration because of the large breeder volume. In the neutronics analyses, the blanket is approximated to three radial zones to model the change in the arrangement of the coolant tubes shown in Fig. VII.1-7 through VII. 1-9. Also, the

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DETAIL Fig. VII.1-8 Cooling Tube Arrangement at the Top (mixed pebble bed type O blanket) Fig. VII. 1-9 Cross-sectional View of Inboard Blanket at Midplane (mixed pebble bed type) blanket is divided to three section in the poloidal direction to accommodate the change in the neutron wall loading and the narrowing of the blanket in the toroidal direction above the midplane as shown in Fig. VII.2-6. The radial thickness of the blanket is 60 cm and it does not change in the poloidal direction. One-dimensional poloidal calculations were carried out to calculate the local tritium breeding ratio and the nuclear heating. The local tritium breeding ratio is about constant for the outboard blanket, varies from 1.45 at Z = 0.0 to 1.49 at the blanket extrimity without the effect of the copper stabilizer. The local tritium breeding ratio of the inboard blanket at the midplane is 0.57. The estimated net tritium breeding ratio is 0.81 based on one dimensional poloidal model and neutron coverage. This estimate is also account for the side walls, the major penetration ports, and the copper stabilizer. The loss in the tritium breeding ratio because of the four coolant plenum in the outboard blanket is not accounted for in this estimate. In the inboard section a constant local tritium breeding is assumed.

108 3) Breeder In Tube Concept

Neutronics analyses were performed for two blanket versions with breeder in tube (BIT). The first version uses LiAlO2 breeder while the second version has Li4SiO4 or Sital (0.6 S5O2, 0.2 Li2O, 0.2 Al2O3 + MgO) breeder. Both blanket versions are using beryllium multiplier. A steel clad is used for the breeder and the multiplier materials. Two- and three-dimensional calculations were performed for the first blanket version to find the tritium breeding ratio and the total nuclear heating in each element. A nuclear cross library based on the European Fusion File in ACE format was used for the MCNP calculations. The tritium breeding results form this library matches the results from the BMCCS library based on ENDF/B-IV. The two-dimensional calculations were performed to determine the local tritium breeding ratio while the net tritium breeding ratio was obtained from the three- dimensional calculation. First, the actual neutron source distribution was sampled for the three- dimensional calculations and at the same time to generate two plane surface sources facing the outboard and inboard first wall for the two-dimensional calculations. This procedure preserves the source characteristics in the two- dimensional calculations. Figure VII.1-10 shows a midplane cross section of the geometrical model used for the calculations. Table VII.1-4 gives the blanket box

TABLE VII.1-4 FIRST WALL/BLANKET BOX DIMENSIONS

MATERIALS INBOARD OUTBOARD

First wall thickness (cm)

graphite 2.00 ss 0.45 0.44 water 0.36 0.36 SS 1.00 0.80

SIDE WALLS (cm)

Cu --.-- 0.5 ss 0.45 water 0.36 0.36 ss 1.00 1.24

GAP BETWEEN SEGMENTS

void 2.00 2.00

109 INBOARD

B/T no ( concept) dimensions used in the calculations. A 2-cm assembly gap is included and a 0.5- cm copper layer is added only to the side walls. The dimensions of the blanket elements are given in section V1.3. However, the neutronics calculations assume a constant steel clad thickness of 0.5 mm compare to the actual clad thickness of 3 to 4 mm and a carbon tile of 2.0 cm for the inboard blanket only. The analyses of thj first blanket version used eighteen poloidal blanket modules for the outboard section, which uses 62.7 and 96.7 cm in the radial direction at the midplane and the extremities, respectively. The inboard section uses only four poloidal blanket modules occupying 16.8 cm in the radial direction at the midplane. A 50% lithium-6 enrichment is used to enhance the tritium breeding capability of the LiAlO2. The LiAlO2 density is 0.85 the theoretical density. Two-dimensional calculation was performed at the midplane to assess the local tritium breeding ratio with and without stiffening ribs. Table VII. 1-5 gives the calculated results for the two cases. The reduction in the tritium breeding is about 10% due to the ribs. Three-dimensional calculation was carried out to estimate the net tritium breeding ratio where the different components were modeled as much as possible. The results are also given in table 5 for two cases. The difference between the two cases is the modeling of the sixteen ports at the midplane. The net tritium breeding ratio is 0.80 form this model. This value does not include the effect of the copper stabilizer at the first wall, the manifolds, anrl the stiffening ribs. The total loss in the tritium breeding ratio due to the coppei stabilizer at the first wall and the manifolds is not calculated. In the second blanket version, One- and three-dimensional calculations based on toroidal model were performed. The results are modified based the neutron coverage of the different components. Discrete ordinates code ANISN and three-dimensional Monte Carlo code BLANK [10] were used to carry out the neutronics calculations. Multigroup cross section libraries based on BROND and ENDL were used for the calculations. Two configurations were calculated for the second version where one or five breeder pins are used per blanket module. The five pins configuration was introduced because it has higher tritium breeding capability relative to the single pin unit. The analysis is based on heterogeneous models because the homogenous models overestimate the tritium breeding results. The one dimensional analysis show that the use of 90% lithium-6 enrichment results in 12% increase in the tritium breeding ratio relative to the natural lithium. Therefore, the 90% lithium-6 enrichment is used for this blanket version. A three-dimensional calculation was performed based on midplane toroidal model where the inboard and outboard blankets are presented in details. In this model the neutron fractions for the inboard and the outboard blankets are 0.205 and 0.795, respectively. The results from this calculation are shown in table VII. 1-6. These results are modified by the actual neutron coverage to exclude nonbreeding sections. The results from this modification are also shown in Table VII.1-6. The net tritium breeding ratio given in Table VII.1-6 does not account for module side walls, side coolant channels, assembly gaps, and copper

111 TABLE VII. 1-5 TRITIUM BREEDING RESULTS FROM TWO-AND- THREE-DIMENSIONAL CALCULATIONS

Geometrical Model 2D 2-D with 3-D 3D with stiffness ribs portsa

inboard 0.146 0.145 0.090 0.091 outboard 0.861 0.758 0.788 0.707 total 1.007 0.903 0.878 0.79&

(a) 16 ports 100 x 340 cm

TABLE VII.1-6 TRITIUM BREEDING RESULTS FOR THE FIVE PINS CONHGURATION

Breeder Material Sital Li4Si04 Local Neutron Estimated" Loca 1 Neutron Estimated Blanket Section TBR coverage TBR TBR coverage TBR

inboard 1.0 0.19 0 19 1.09 0.19 0.21 outboard 1.33 0.57 0 76 1.47 0.57 0.84 TOTAL 0 95 1.05

a) It does not account for module side walls, side coolant channels, assembly gaps and copper stabilizers.

TABLE VII.1-7 NEUTRON BALANCE FOR THE LEAD LITHIUM BLANKET

0T neutron source 1.00 (n, 2n) - neutrons 0.42

Capture Tritium local by zones Breeding TBR

Inboard 0.27 0.13 0.69 Outboard Blanket: 0.97 0.70 1.04 Divertor Regions 0.18

TOTAL 1.42 0.83

112 stabiliser. Also, it does not account for the different toroidal manifolds and the support structure. The loss in neutron coverage due to side walls, assembly gaps, toroidal manifolds, attachment structure, and geometrical considerations is about 15 to 20% of the available first wall for the driver blanket. Also, it is expected that the used nuclear data libraries are over estimating the tritium breeding ratio by about 7% based on comparing results from ENDF/B-IV to -V. Several other studies were performed for this blanket version to define the impact of the carbon tile, the inboard blanket thickness, the copper stabilizer, and the use of a breeding blanket behind the divertor plate on the tritium breeding ratio. The carbon tile reduces the tritium breeding latio by 0.5% per mm. The use of a third blanket module in the inboard section increases the tritium breeding ratio by about 3.7% but it has a major impact on the shield performance. The integration of the cooper stabilizer behind the first wall in the outboard blanket reduces the net tritium breeding ratio by about 3%. Adding breeding modules behind the divertor plates increases the net tritium breeding by about 5.8%.

4) Lead Lithium blanket

Tritium breeding and nuclear heating, activation and decay heat analyses were performed for the lead lithium blanket. The net tritium breeding ratio is estimated based on three-dimensional calculation and modification to account for the sixteen ports and the copper stabilizer. The three-dimensional version of the BLANK Monte Carlo code [10] with a multigroup library based on ENDL data base was used to perform the transport calculation. A heterogenous geometrical model was used to calculate the tritium generation rate in each lead lithium channel. The (n,2n) reaction rate from lead is 0.42 per DT neutron. The lithium-6 enrichment is 90% to maximize the tritium breeding ratio. The net tritium ratio is 0.83 as shown in Table VII.1-7. This value does not include the effect of the sixteen ports at the midplane or the copper stabilizer. The midplane ports have a neutron coverage of 10.3% per DT source neutron. In the case of no tritium recovery from these ports, it is estimated that the loss in the net tritium breeding is only 0.05 for the lead Lithium blanket. The impact of the copper stabilizer is calculated based on one-dimensional analyses and it causes about 0.7% loss in the net tritium breeding ratio. Therefore the net tritium breeding ratio is about 0.77.

REFERENCES

[1] Y. (iohar et al., "First Wall/Blanket/Shield Design and Optimization System," International Symposium on Fusion Nuclear Technology, Tokyo, Japan, April 10-15, 1988. [2] Y. Gohar et ah, "Water-Cooled Solid-Breeder Blanket Concept for ITER," Fusion Technology, \15\, March 1989. [3] R.D. O'Dell, F.W. Brinkley, and D.R. Marr, "User's Manual for EDANT: A Code Package for One-Dimensional Diffusion-Accelerated,

113 Neulral-Particle Transport, Los Alamos National Laboratory Report LA-9184-M (February 1982). [4] Y. Gc-har "A Coupled 46-Neutron, 21 Gamma Ray Multigroup Nuclear Data Library for Fusion Analysis Based on ENDF/B-V," Argonne National Laboratory Report (to be published). [5] R.W. Roussin et al., "V1TAMIN-E: A Coupled 174 Neutron, 38 Gamma Ray Multigroup Cross-Section Library for Deriving Application Dependent Working Libraries for Radiation Transport Calculations," Oak Ridge National Laboratory Report ORNL/RSIC-XX (August 1984). [8] Y. Farawila, Y. Gohar, and C. Maynard, "KAOS/LIB-V: A Library of Nuclear Response Functions Generated by KAOS-V Code from ENDF/B-V, Argonne National Laboratory Report ANL/FPP/TM-241 (April 1989). [9] K. Shibata, ct al., "Japanese Evaluated Nuclear Data Library, Version-3 - JENDL-3," JAERI Report 1319 (June 1990). [10] S. V. Marin, D. V. Markovsky, G. E. Shatalov/'BLANK - The 1-D Neutron Transport Monte Carlo Code", Kurchatov IAE, Preprints IAE- 2832 (1977), IAE-3044 (1978).

VI1.1.4. Activation and Decay Heat Analyses

The radioactivity and the decay heat have been calculated, during operation and after shutdown for the two ITER phases. The physics phase has about 24 full power days (FPDs) at fusion power level of 1100 MW and the technology phase has 860 MW fusion power and operates for about 3.7 full power years.

1) Ceramic Blanket

A one-dimensional toroidal model at the midplane for the ceramic layered blanket given in section VI. 1 was used for the analyses. This model produces average vertical (poloidal) values representing part of the system rather than localized values. The inboard blanket receives 18.4% of the DT neutrons and the outboard blanket receives 73.8% of the DT neutrons. The corresponding average neutron wall loadings are .59 MW/m2 on the inboard and .895 MW/m2 on the outboard. These neutron wall loadings values should be scaled by 1.28 to obtain the physics phase values. In this model, the inboard/outboard effective vertical heights are 9.2 m and 10.9 m. These heights should be used to obtain the total inboard or the total outboard integral activation values. The neutron flux, used in the activation calculations, has been normalized to account for the average neutron wall loadings on the inboard and the outboard first walls. The activation responses are calculated with the RACC code [1] and its associated data libraries RACCDLIB and RACCXLIB. The average activity per cm height after 3.8 full power years of continuous operation (the two phases) is shown in Fig. VII.1-11 as a function of

114 100

io id ltf id1 io4 icf 10° io' ioT iof io10 Time [sec] Fig. VII. 1-11 Isotopic contributions to the average activity

time after shutdown. This calculation gives an upper bound, the actual value will depend on the operating scenario. The average radioactivity amounts to 1.5 MCi/cm at the shutdown of the physics phase, and at the end of the technology phase it reaches 2.2 MCi/cm. The corresponding values, after one pulse, are 0.81 and 0.78 MCi/cm. The higher value of the first pha~e (physics) after one pulse reflects the higher power and consequently, the high generation rate of the short- lived isotopes. After the full operations of the two phases, the second phase has higher activation level because of its higher fluence, which generates more long- lived isotopes. The contributions of the inboard zones and the outboard zoner to the average radioactivity are 30% to 35% and 70% to 65%, respectively. At shutdown, the Be/L^O zones produce about 30% of the average activity, which vanishes in few seconds. The 0.05 cm FW tungsten coating (inboard and outboard) produces more than 10% of the total activation for about one day after shutdown. At shutdown, the inboard FW lias 8 kCi/cc, and the outboard has 10 kCi/cc. The inboard and outboard FW coatings have 61 kCi/cc and 63 kCi/cc, respectively. Figure VI.1-11 shows the isotopic contributions to the average activity at the end of the technology phase. The dominant isotopes are: %e[.8s, Befoalpha)] at shutdown, 56Mn[2.6h, 55Mn(n,gamma)l up to 1 hour, 55Fe[2.68y, 54Fe(n,gamma)] from 1 Apur to 10 years^ 63Ni[100y, 62Ni(n,gamma)] from 12 years to 100 years, and 93Mo[3500 y, 92Mo(n,gamma)] beyond 500 years.

115 100 10'

ao i 10'

10° 101 lCf 10a 104 106 108 107 10° 10" 10'° Time fsec] Fig. VII.1-12 Isotopic contributions to the average decay heat

The time dependence of the decay heat generation rates (DHGR) after the shutdown of both phases and after the first DT pulse in the two phase is similar to that of the radioactivities. After the first DT pulse, the DHGRs are 8.63 and 8.58 kW/cm for the technology and the physics phases, respectively. At the end of the physics phase, the DHGR is 15.65 kW/cm compared to 15.57 kW/cm at the end of the technology phase. The isotopic contributions to the DHGR, are different from that in the radioactivity case. Figure VII.1-12 illustrates these contributions after the second phase operation. ^Mn with its 1.69 MeV gamma and .83 MeV Beta produces 40 to 70% of the total DHGR for about 4 hours, then, 58Co[71d;59Co(m2n); 1 MeV gamma,EC, Beta] takes the lead for about two months. After that, ^Mn[312 d; .84 MeV gamma] dominates for about 9 months, and then, Co[ 5.3 y; 2.5 MeV gamma] followed by 63Ni become the leading contributors to the decay heat. The integration of the DHGR over 1000 years results in 81 GJ/cm of nuclear heating where the release rate is shown in Fig. VII.1-12. In one week after shutdown, only 1 GJ/cm would have been released from Mn disintegration. The integration of the DHGR of each isotope provides more indicative measure of the importance of the isotope, since it combines the effects of its concentration, its disintegration energy, and its half-life. Figure VII.1-13 exhibits the integrated DHGRs of the isotopes, which contribute more than,J% of the total integrated DHGR at any time. ^Co, 54Mn, 58Co, 55Fe, and have the largest stored energy.

116 10*

io° id lcf id1 io' io6 io9 ioT lb8 io10 Time [sec] Fig. VII. 1-13 Isotopic contribution to the integrated decay heat 2) Lead lithium blanket

A one-dimension cylindrical toroidal model at the midplane for the lead lithium blanket given in secf«on VI.2 was used with ANISN for the transport calculations and FRINDA for the activation calculations. The calculations were performed for different operating scenarios to study the impact on the total activity and the decay heat after shutdown. Four operating scenarios were analyzed with different availability and total fluence. The total induced activity is shown in Fig. VH.1-14 as a function of time after shutdown. The maximum specific activity is 6-10 Ci/cc in the outboard first wall. The steel structure contributes 40-49% of the total. The total activity is about 260 MCi at shutdown after 1 MWa/m2 over 10 years of operation. Special attention was given to the polonium production from the lead- lithium eutectic. The total polonium inventory is 1080 Ci at the end of the technology phase with 3MWa/m average neutron fluence and 100% availability. Every two weeks of full power, polonium probably will be separated from the eutectic during the tritium recovery process. This will result fa 60 Ci as the maximum polonium activity at the end of two weeks of operation. The decay heat analyses were performed with the same geometrical model used for the activation analyses. Figure VII.1-15 give the decay heat as function of time after shutdown. The highest decay heat occurs in the outboard first wall. The decay heat value rt shutdown is about 2% the nominal operating power and it decrease fast after shutdown as shown in Fig. VII.1-15.

117 105| 00

10 4i

10 3i Technology phase:

10 2i o 10 i Physics phase

o °10-'ii Availability 0.016 1 0.25 0.1 "5 ZJ t° 10" Fluence, - MWa/m2 0.057 0.029 0.108 1.08 10 "3i

10 10M 1H 1D 30D 1Y 10Y 100Y 5 10" i i i mill]—i i 111in)—1n)i i i i i|i||in|)||i|ni|—i i|iinii|—i i 111ii||—i i niiH|—q[ i mni|—|||i |MIIII|—r—j-n \ | 10 10 102 103 104 10s 106 107 10" iO9 Time after shutdown, s Fig. VII.1-14 Total induced activity for different operating scenarios 10 4| ( Nominal Heat Power ) 10 3i

10 2i Technology phase:

10 i

1 i o a> x: 10 "'] Physics phase a> o "5 Availability 0.016 1 0.25 0.1

Fluence, - MWa/m7 0.057 0.029 0.108 1.08

10 "5 10M 1H 1D 30D 1Y 10Y -6 10 i i mni| I T 1 1111] 1 | | 11 III; 1 | 1 |||||| 1 l|HHH| 1 I I llll|| 1 I I 1NII| i \ i I 10 "' 1 10 10 z2 1in0 33 110Q 44 110Q 55 10 10 7 10 8 10 Time after shutdown, s

Fig. VII.1-15 Total afterheat for different operating scenarios REFERENCE

[1] J. Jung,"Theory and Use of The Radioactivity Code RACC,"ANL/FPP/TM-122, Argonne National Laboratory (1979).

VII.2. THERMAL/MECHANICAL ANALYSIS

VII.2.1. Breeder Temperature Control

The HÔ coolant in the ITER designs is at low temperature (60-100°C). Breeder ceramics have a minimum temperature (Tj^) above which tbs tritium release rate is high enough to keep the tritium inventory low. Generally, the ^min COTresPon(^nS to an inventory equal to one-day's generation rate is used as a design guideline. Based on very low burnup in-reactor tests with He + 0.1-1% H2, the T -j for LÔ ( ~ 80% dense with grain diameter < 20 fi,m) and LiAlO2 (~ 80% dense with grain diameter < 1 pim) are 320°C and 450°C, respectively. Because of other considerations (e.g., LiOT precipitation in LÔ), the recommended Tj^ for Li2O is 400°C to allow more flexibility in the purge flow design. In any case, the temperature rise from coolant to breeder interface must be OD the order of 300-400°C for adequate tritium retention/release performance. Two methods have been considered in ITER conceptual designs to provide a thermal barrier between breeder and coolant for breeder temperature control: Be layers (either sintered blocks or pebble beds); and He gas gap. Figure VI.1-1 shows the layered concept with four Be-block regions (30-150 mm in radial thickness and 15-35% porosity) [1]. An alternative to this concept is shown in Fig.VI.2-1 where the layers of blocks are replaced by Be pebble beds [2]. The second approach considered for ITER utilizes a He-filled gap in the annulus of a double-walled tube at the inner diameter of the coolant annulus (see Figs. VI.3-1 and VI.3-10) for the breeder-in-tube (BIT) concept [3] or at the outer diameter of the coolant tube (see Fig. VI.2-5) for the breeder-out-of-tube (BOT) concept [2]. Gas-gap thicknesses vary from ~ 1 mm to ~ 4 mm depending on radial location within a blanket segment and poloidal position. Several analyses [1,4] have been performed for the layered concept with Be-blocks. For an overly constrained blanket (which is good for interfacial heat transfer), unacceptably high stresses would be generated in the blanket layers and high loads would be transmitted to the first wall. For a relatively unconstrained blanket module which allows each layer to deform independently (good for minimizing stresses), increase of Be/steel interfacial heat flow resistance is a concern. Analyses have been performed to determine the unconstrained deformation and thermal feedback for the layered concept with Be-blocks. The first set of analyses focused on the beginning-of-life deformations due to temperature and temperature gradients within the continuous steel layers and the Be-block and the breeder-block layers. These analyses were used to determine fabrication clearances to allow unconstrained deformation of the blanket layers

120 TABLE VII.2.1-1 DEFORMATION DUE TO FREE THERMAL EXPANSION OF BLANKET COMPONENTS AT THE OUTBOARD CORE MIDPLANE POSITION FOR A 1 METER TOROIDAL SPAN OF THE BE BLOCK CONCEPT

Layer (DF) h, mm Min. Max. Thermal+Bending Local Gap, T,°C Expansion.jim Maximum, {Am Ah Al

SS first wall 5 77 191 0.007 1.39 -- A 4 __ SS back wall 5 82 225 0.009 1.80 Be (0.85) 34 280 401 0.175 6.27 17.7 SS clad 1 426 438 0.009 9.16 — Li.,0 (0.80) 8 453 537 0.109 13.63 -- ss'"clad 1 424 437 0.009 9.11 — Be (0.85) 59 192 401 0.235 5.56 17.1

SS coolant channel 2 77 138 0.002 0.87 — n i i. SS coolant channel 2 70 109 0.001 0.53 —

Be (0.65) 57 138 421 0.228 5.46 23.8 SS clad 1 435 446 0.009 9.45 — ,0 (0.80) 8 451 505 0.103 12.87 — Be (0.65) 71 123 427 0.261 5.45 20.5

SS coolant channel 2 68 99 0,.000 0.04 -- 1 I SS coolant channel 2 60 60 0 0 -- and to find acceptable Be-block dimensions (poloidal and toroidal) based on interface heat transfer characteristics. Based on analytical methods validated by comparing formula calculations to ANSYS 4.4 results, the total increase in the 253-mm radial thickness of blanket components (at the outboard core midplane position) between the shield and first wall is only 1.15 mm due to nominal thermal expansion (~ 1 mm relative to the 60°C coolant inlet temperature) ana to Be-block bending under the thermal gradient (0.16 mm). The bending calculation assumed poloidal/toroidal lengths of 34 mm for each block and a radial thickness of h (34, 59, 57, and 71 mm, respectively, for the four rows of blocks, starting from the first wall). The deformation results are summarized in Table VII.2.1-1. The secondary thermal stresses in the Be-blocks were negligibly small because of the nearly-linear temperature profile across each block. The breeder blocks have essentially a parabolic radial temperature profile, resulting in no bending, but in higher stresses. The maximum tensile

121 stress ( 80 MPa) may be large enough to crack the breeder blocks parallel to the heat flow direction. The dimensions of the Be blocks (34 mm by 34 mm by h) were chosen to minimize Be curvature at the Be/steel interface. In specific, the poloidal/toroidal length was reduced until the maximum local gap between the curved Be and the steel cladding was less than the estimated combined Be/steel surface roughnesses (~ 19 \Lta for 85% dense Be and ~ 41 ji,m for 65% dense Be). The resulting maximum gaps for the four Be zones at the outboard core midplane for 34-mm-by-34-mm-by-h blocks are 17-18 mm for the 85% dense blocks and 21-24 mm for the 65% dense blocks (see Table VII.2.1-1). The net effect of these small gas gaps is to raise the average temperatures of the two breeder layers at this position by ~ 20°C (first layer) and ~ 10°C (second layer). Thus to accommodate the differential thermal expansion of the blanket zones without inducing any stresses on the first wall, only 1.15 mm of radial clearance (relative to 60°C) needs to be fabricated out of the 253 mm of blanket thickness. Accommodating the toroidal expansions AL (1-14 mm) is no problem because this direction is not a heat transfer path and gas spaces are already specified between these layers and the side walls. However, two additional design issues which need to be addressed are the effects of Be/LLO swelling and first- wall pressure-induced deformation on thermal performance. The total radial increase in LL,O at the outboard midplane, using an upper bound estimate of 5 vol. % for T < 700°C, is only 0.27 mm. The total radiaJ increase in Be due to He generation, using ATR and EBR-II data extrapolation, is only 1.1Q mm at the outboard midplane for an end-of-life reactor fluence of 3 MWa/m . Thus, the end-of-life swelling is about the same as the beginning-of-life thermal expansion. The stresses induced in blanket and first-wall components due to this time- dependent swelling depend on the thermal and irradiation creep rates of the components, with first wall irradiation creep expected to be the dominant stress- relaxation mechanism. The resulting stresses on the first wall due to swelling of the blanket components requires at least a 2-D, time-dependent analysis of the multi-layered blanket module. If this analysis results in unacceptable loading of the first wall, then the option exists to replace only the first layer of Be blocks with a more compliant Be pebble bed. With regard to the impact of first-wall deformation, the problem is related to the overall design of the first waii to accommodate gas-pressure, thermal, and disruption loads. The allowable primary membrane stress intensity for the first wall with annealed type 316 stainless steel is 110 MPa. The maximum primary bending stress intensity is 125 MPa for the box-like first-wall structure. For a first-wall of 1-m toroidal span with no reinforcement, both the peak bending stress ( ~ 220 MPa) and the peak deformation (8 mm) are too high in response to an internal gas pressure of 0.1 MPa. As these values vary essentially linearly with gas pressure, reducing the gas pressure to 0.05 MPa reduces the stresses to within limits, but allows little margin for disruption loads and results in a maximum deformation at the Be/first-wall interface of 4 mm. It is clear that the first wall must be reinforced with beams (toroidal), ties (from first wall to shield) and/or a shorter toroidal span to satisfy stress

122 800.0-1 parameters max t 5 (celslus) mg.x...t 6..J.ce.lsljLi.sJ _ I0P.¥—k...Z..(cp.L.s.Ly.sj.. 700.0- max U 8 I Celsius) max t 9 ' (cei sIus)

600.0

a. S soo.o

400.0

300.0- 0.000 0.125 0.250 0.375 0.500 0.625 0.750 0.075 1.000 Gap conductance between • the first wall and the beryllium multiplier material (W/cm2-k) Fig. VII.2.1-1 Maximum temperatures of the breeder, clad, and multiplier materials of the first breeder zone of the blanket configuration at the midplane as a function of the gap conductance between the first wall and the beryllium multiplier material. intensity criteria. The first-wall deflection will decrease as more reinforcement is added. This subject is discussed in Section VII.3. In this section, a maximum displacement criterion is established for the first wall relative to the first Be layer from a heat transfer perspective. The maximum breeder temperature is relatively insensitive to Be-block/first-wall heat transfer coefficient for H > 2000 W /m -K (Fig. VII.2.1-1). This corresponds to a Be/first-wall gap of <82 mm. The reason for this is that as the Be/first-wall gap opens up, giving more resistance to heat flow, then a higher fraction of the heat from the first breeder plate flows away from the first-wall coolant towards the blanket coolant. Under the conditions that H > 2000 W/m2-K, Tmax < 600°C for the breeder (top line) and Tmax < 520°C (solid line) for the first Be layer. Thus, a reasonable criterion for first-wall displacement relative to the first Be surface is 6 _<. 80 \im. If the first Be-block layer was replaced by a Be pebble bed, the heat transfer would be less sensitive to first-wall displacement. The pebbles would redistribute somewhat to accommodate the first-wall displacement profile. Also the pebble bed would be more compliant than the block design, resulting in lower loads transmitted to the first-wall in regions (e.g., near the side walls) where the swelling displacement of the Be was greater than the deformation of the first wall. Thus, a first Be layer of pebbles offers the possibility of accommodating both Be-sweUing and first-wall pressure-induced deformation without

123 overloading the first wall and without seriously degrading the heat transfer. These potential benefits of the Be pebble bed need to be substantiated by more detailed analyses and experimentation. In particular, the issues of changing heat transfer characteristics with irradiation time (due to possible pebble sintering and/or redistribution) need to be addressed. In the second proposed means of thermal control, a He gap is used to increase the breeder temperature relative to the coolant temperature. In this approach, the Be is placed outside of the coolant/breeder interface and serves only the function of neutron multiplier. It is also operated at relatively low (< 150°C) temperature. Thus any time-dependent behavior of the Be does not impact the breeder temperatures. Also, with the BIT approach, the heat transfer within the tube is essentially decoupled from the behavior of the first wall. In order to realize these potential advantages, the He gap thicknesses must be fabricated to close tolerances (i.e. 0.1 mm) and these thicknesses be controlled during in-reactor operation. This necessitates the use of spacers between the first cladding (closest to the breeder) and the second cladding (in contact with the coolant). The proposed design foresees a solution (Fig. VI.3-2) in which the spacers are removed from the breeding/heat-transfer zone and plr ced at poloidal locations where there is no breeder pellet. In this design, the hot breeder and first cladding are discontinuous while the other three claddings are continuous. This allows for accommodation of differential thermal expansion in the poloidal direction between breeder/first-cladding and second cladding. In addition, a spring is used to maintain the pellet positions while allowing poloidal expansion. Temperature control is achieved (with respect to poloidal power variation) by varying discretely the wall thickness of the first cladding. For a tube closest to the first wall at the midplane, the first-cladding is 1.1 mm thick with a He gap of 1 mm. At the poloidal top and bottom of this tube, the first-cladding thickness is 0.5 mm with a 1.6 mm He gap. For the last radial row, the first wall cladding varies poloidally from 1.85 mm to 0.3 mm as the He gap varies from 1.55 mm to 3.1 mm. This results in an "average" Tmin of ~500°C at nominal power and a Tmin = 430°C for a 20% reduction in overall reactor power. With regard to maintaining gap thicknesses over the design life of the blanket, it remains to be demonstrated that the spacers and the auxiliary components (e.g., springs and insulators) are effective. Deformation could arise from bowing of the 10-m tubes. If maintaining He-gap tolerances either during tube assembly or in-reactor operation proves to be a problem, then it is possible that several poloidal segments of straight tubes could be used. The penalty from this approach would be in reducing the tritium breeding ratio. In summary, the two approaches proposed, Be-block thermal-resistance layer and He-gap thermal-resistance, have potentially attractive features and unresolved issues requiring further analyses and experimental validation. The primary issue with the Be-block concept is maintaining reasonable heat transfer at the Be/first-wall interface without over-stressing the first wall. Under beginning-of-life conditions, Be curvature due to the large AT across its radial thickness and pressure-induced deformation of the first wall will tend to create

124 local and global separations at this interface, with the net effect of raising both Be and breeder temperatures. A separation of ~ 0.1 mm can be tolerated in the design. The temperature profile across the blanket, however, is relatively insensitive to any increase in interface heat transfer due to Be/steel contact pressure. Thus, the major issue is the possible increase of Be and breeder temperatures due to deformation-induced separations at the interfaces. Further analysis, as well as out-of-reactor testing on a submodule and module level, is required to demonstrate satisfactory beginning-of-life behavior. With regard to lifetime thermal/mechanical behavior, it needs to be demonstrated that Be swelling can be accommodated. The swelling, which is on the order of ~0.1 mm across the radial thickness at the end-of-life, could be accommodated at the center of the module span by a first wall which deforms — 0.1 mm due to pressure loads. However, stresses will be induced by Be/steel contact for sections of the span near the side-walls. These stresses build up very slowly in time and are partially relaxed by the irradiation creep of the first-wall steel. Again, the magnitude of these stresses and their effect on first wall lifetime need to be evaluated. A back-up design alternative is to replace the first Be-block layer (only) with a pebble bed which may be more compliant and less sensitive with regard to heat-transfer in response to first-wall deformation. For the He-gap approach (in conjunction with concentric tubes), the blanket modules are decoupled, both thermally and mechanically, from the first wall. The primary issues have to do with the fabricability of the long (poloidal) concentric tubes, while maintaining close (0.1 mm) gap tolerances. Be swelling is not an issue, but thermal-mechanical feedback due to other deformation mechanisms (e.g., bowing) may be an issue. The critical area for the design is the satisfactory performance of the spacer region which functions both to maintain the He gap during operation to the desired tolerance and to allow differential thermal expansion between the hot (~400°C) first cladding and the cold ( ~ 100°C) second cladding. The consequences of higher- or lower-than-calculated heat transfer across the gas gap depends somewhat on the choice of breeder. In the current design with LiAlC^ breeder, lower-than-caiculated breeder temperatures would be a problem because of resulting higher tritium inventories. L^O and LLjZrO-j would be less sensitive to this. Higher-than-calculated temperatures would cause a problem with power variation capability due to the breedei maximum temperature limit.

REFERENCES

[1] Y. Gohar et al., "U.S. Technical Report for the ITER Blanket/Shield, A. Blanket," ITER-TN-BL-5-0-3, July-November 1990. [2] H. Yoshida et al., "Japanese Contribution to Blanket Design for ITER," ITER-IL-BL-5-0-10, July 1990. [3] W. Daenner et al., "EC Contribution to ITER Blanket Workshop: EC Blanket Design," 16-27 July 1990. [4] W. Daenner et al., "EC Short Presentations on Multilayer Configuration," ITER Blanket Workshop, Garching, 16-27 July 1990.

125 VII.2.2. Hydraulic analysis

In the blanket design, the water coolant is supplied and collected at the top of the reactor except for the central lower outboard segments where the supply and the return have to be from the bottom. The inlet temperature is 60°C, the outlet temperature is 100°C, the inlet pressure is less than 1.5 MPa. In the first wall, the coolant runs poloidally in the inboard section and toroidally in the outboard section where a separate coolant loop is used for the first wall. The analysis is summarized for the different blanket options.

1) Solid Breeder Blanket In the multilayer concept, the coolant runs poloidally in the inboard blanket/shield and outboard shield, while it runs toroidally in the outboard blanket. The first wall coolant is totally separate and independent from either the breeding zone or shield coolant. The main thermal hydraulics parameters are summarized in Table VII.2.2-1. The pebble bed concept have poloidal cooling for both blanket sections. The cooling loop of the first wall is separate form the blanket. The major thermal hydraulics parameters are summarized in Table VII.2.2-2.

TABLE VH.2.2-1. MAIN THERMAL HYDRAULIC PARAMETERS OF THE MULTILAYER BLANKET

PARAMETER INBOARD OUTBOARD Inlet coolant temperature,°C 60 60 Outlet coolant temperature 100 100 Temperature rise. °C FW coolant 31.7 Side Wall (SW) coolant 8.3 Shield (first /last channel) 36.5/3.5 0.6/0.2 Max. velocity, m/s FW 7.1 5.7 SW channel (Midplane) 4.0 Average velocity, m/s FW channel 6.7 1.6 SW channel 3.0 Blanket (lst/2nd channel) 2.1/1.2 Shield (first /last channel) 4.4/0.7 2.2/1.5 System pressure, MPa 1.5 1.5 Pressure drop, MPa FW/SW coolant system 0.56 0.03 Blanket 0.03 Shield (first /last channel) 0.18/0.0007 0.02 Coolant manifold (one side) 0.03 Volumetric flow rate per module, 1/s FW 21.5 90.6 Shield 24.4 92.0 Total volumetric flow rate, 1/s FW 688 2900 Shield 781 2944

126 TABLE Vli.2.2-2 MAIN HYDRAULICS PARAMETERS OF THE PEBBLE BED BLANKETS

Layered Mixed inboaru outboard inboard outboard

Inlet temperature,°C 60 60 60 60 Outlet tcmperature,°C 100 100 100 100 Inlet pressure, MPa 1.5 1.5 1.5 1.5 Outlet pressure, MPa 1.3 1.2 1.48 1.4 Maximum velocity, m/s 3.5 3.5 1.5 3.0

With respect to the Poloidal BIT concept, the coolant runs poloidally in the inboard and outboard first wall and blanket. The first wall cooling system is completely separate and independent from the blanket cooling system. The blanket coolant flows through the front bundle of 100 tubes and comes back through the back bundle of 85 tubes in the outboard blanket segment. Detailed hydraulic analysis has been performed for the outboard blanket. The main thermal hydraulics parameters are summarized in table VII.2.2-3. The toroidal BIT concept consists of short modules to.oidally cooled. The first wall coolant is completely independent, while the blanket and the shield have the same cooling system. The coolant flows through the shield (60/80°C) and comes back through the blanket (80/100°C). The main hydraulics parameters are listed in Table VII.2.2-3 From the hydraulics point of view, a toroidal cooling presents the advantages of lower pressure losses, lower flow velocity and a shorter coolant path in the irradiation zone. However, it requires flow control to insure the correct mass flow rate in each coolant path.

TABLE VII.2.2-3. MAIN THERMAL HYDRAULICS PARAMETERS OF THE BIT BLANKET

Poloidal BIT Toroidal BIT PARAMETER 1 st bundle 2nd bundle Inlet temperature °C 60 77 80 Outlet tenperature °C 77 92 100 Coolant velocity (1 st/last tube row).m/s 4.0/3.1 3.5/2.8 Mass flow rate (1 sVlast tube row), kg/s 1.6/0.9 2.0/1.2 Nominal pressure, MPa 1.0 1.0 Pressure drop, MPa 0.14 0.11 0.05

127 2) Lithium-Lead blanket

For the lithium-lead concept, the first wall coolant runs poloidally in the inboard blanket and toroidally for the main option of the outboard blanket. This cooling system is completely independent from the blanket cooling system which is poloidal for both inboard and outboard blankets. In the outboard the coolant runs from the top of the first row and comes back from the bottom through the two back rows. Thermal hydraulic analysis has been performed for the inboard and the outboard blanket for the physics and the technology phases. The main thermal hydraulics parameters are summarized in Table VII.2.2-4.

TABLE VII.2.2-4 MAIN HYDRAULIC PARAMETERS OF THE LITHIUM- LEAD BLANKET

PARAMETER INBOARD OUTBOARD

1 st row 3rd row Max coolant temperature °C 1.1 /0.8 1.1/0.9 1.0/0.8 Inlet coolant temperature °C 60/60 60/60 65.5/85.5 Outlet coolant temperature °C 100/100 85.5/85.5 95.8/95.8 Max coolant velocity, m/s 3.8/3.0 3.8/3.2 1.0/0.7 Coolant flow rate per channel, kg/s 6.14/4.8 14.3/11.5 8.0/6.2 Pressure losses, MPa <0.5 < 0.5 < 0.5 Channel thermal power, MW 1.0/0.8 1.5/1.2 0.3/0.3

VII.3. STRESS ANALYSIS

VII.3.1. Normal Operation

1) Cc mic Blanket concepts

Several stress analyses have been performed for the ceramic blanket concepts using advanced FEM computer programs (e.g.: ABAQUS, ANSYS, etc). For the multilayer concept based on a full integration of the blanket and first wall, attention has been paid to the interaction between the two components. For the BIT concept based on a modular configuration of the blanket, the interaction between the two components is not critical (at least during normal operation) so that their structural behaviour can be separately investigated.

Two sets of analysis have been carried out for the multilayer concept. The first one focused on the response of the FW box subjected to a pressure load (i.e.: 0.05/0.1 MPa, due to the He gas filling the box) and thermal load (i.e.: heat flux incident from the plasma, peak value up to 0.4 MW/m ). The second one

128 focused on the temperature and swelling induced stress/displacements in the blanket when sintered blocks of Be and LLO breeder layers are used.

A two-dimensional stress analysis of the FW box has been performed for generalized plane strain deformation taking into account both the pressure and thermal loads. The results show that: - The deformation and stress of the FW box are mainly due to the gas pressure load. - In the absence of any stiffening structures, the maximum displacement at the centre of the FW front panel (1.0 m surface area at the mid plane) is about 8 mm toward the plasma side. Based on the maximum displacement value, it is clear that the first wall must be reinforced or a shorter span (toroidal and poioidal) be used. The max. stress can be within the allowable stress limit of 316SS by using a gas pressure _<_0.055MPa. The radial stiffening ribs reduce substantially the stress and the deformation of the first wall.

Another analysL jias been performed by considering FW toroidal strip of a unit poloidal extent (located at *he mid plane) under the assumption that the 3D structural effects of the FW stresses are small because the toroidal span of the FW box is small compared to the poioidal length. Primary stress limits of 100 Mpa and 125 MPa for membrane and bending stress, respectively, have been used for the FW box which is assumed to be of a sandwich construction. The results show that: - The maximum permissible gas pressure in the FW is 0.05 MPa. - To avoid FW radial displacements larger than 100 |Xm, the blanket/FW segment must be divided into 3 subsegments ( ~ 0.33 m toroidal span) and FW must be stiffened with 3x3.3 cm toroidal beams at 40 cm poloidal spacing. A possible alternative is to stiffen the FW with a proper number of toroidal beams and radial ties connecting the beams to the shield, so that the distance between the ties is ~ 0.33 m. - The maximum effective thermal stress in the FW is 400-450 MPa at 0.25-0.4 MW/m2 heat fluxes, respectively. The 3 S design limit for thermal stress (330 MPa) is exceeded at 0.4 MW/m .Inelastic analysis will be needed to ensure that ratcheting does not occur at higher heat fluxes.

The analysis of the temperature and swelling induced stress/displacements on the blanket layers liis provided the following results: - minimum clearances of 1.53 mm and 15 mm are necessary to accommodate the radial and toroidal thermal strains, respectively. - Swelling (_<_2.6% in Be and <5% in LL0) causes an additional radial displacement of 1.37 mm which should oe accommodated in the FW box. The swelling-induced stress in the Be blocks should be partially relaxed by the ir'adiation/thermal creep. A multilayer swelling/creep analysis is required to quantify these effects and consequent loads

129 transmitted to the FW. However, only very limited data are presently available on Be irradiation creep. - Thermal performance is satisfactory for slight Be curvature at Be/steel interfaces. Hot gap of 100 mfi can be tolerated at Be/FW interface.

The results of a 3D thermal and stress analysis performed on *he poloidal BIT concept show that:

- If an assembly clearance of 50 flm is set at both breeder/cladding and Be/cladding interfaces, no contact pressure is generated between steel cladding and breeder or beryllium pellets during the operation. Consequently, negligible stress is generated (few MPa) in all the components of the blanket module except for breeder pellet. The stress level in the breeder is ~ 150 MPa which is expected to crack during the operation without consequences because of the steel cladding. The poloidal elongation of the breeder rods (~ 4 mm) does not lead to additional stress if proper poloidal segmentation of the breeder rod and springs are provided. Assuming a maximum Be swelling of about 2% (max. value expected at end of life), the assembly clearance between Be pellet and steel cladding must be set at 100 flm to avoid contact. In this case the stress in the Be is still negligible ( ~ 5 Mpa) even in case of signific ant differential swelling (_+_20%) in the Be pellet.

As the gap spacer is considered to be critical issue of the BIT design, a detailed 3D analysis has been devoted to it providing the following results:

- Thermal stresses in the first steel cladding (located in the breeder rod plug), the second steel cladding, and the gap spacer are far below the allowable limit. No significant cold spots in the breeder and negligible hot spots in the second steel cladding have been found due to the presence of the spacer. Operating temperature of the spacer (INCONEL X-750 or 718) is<_ 430°C.

Other structural operating analyses performed for the BIT concepts include:

- Vibrations of 'ube bundle were studied by menas of a model analysis. With a tube bundle supporting system consisting of 6 supporting grids spaced along the poloidal direction, natural vibration frequencies >32 Hz have been found. This value provides sufficient margin against seismic events. - Stresses due to temperature and dead weight of the blanket modules have been computed assuming the same supporting system used for the vibration analysis. This system allows poloidal elongations of the blanket modules and it maintains the same curvature in the poloidal

130 direction. Under these constraints the temperature induced stress is ~ 2 MPa and the weight stress is about 14 MPa.

A 3D transient thermal stress analysis of the FW box has shewn that the 3D effects due to the overall FW box structure strongly affect its mechanical behaviour, leading to different deformations from those obtained from 2D or partial 3D analyses. The AT between the hot FW front panel and the cold shielding region leads to a compressive stress of the front panel. The compressive forces result in displacements of the FW front panel towards the blanket volume, in contrast to the results obtained from 2D and partial 3D analysis. A maximum displacement of about 4 mm has been found (mid plane) due to the thermal loads. The max. thermal stress is about 450 Mpa in the FW front panel for 0.4 MW/m peak heat flux. It is due to the combined effects of the AT through the fronf panel and AT between the front panel and the shielding region.

2) Lithium-lead Blanket Concept

A stress and life-time analysis of the blanket module for the lithium lead concept has been performed taking into account the following cyclic loads:

- Hydrostatic pressure due to melted eutectic, i.e. 2 MPa. - Water coolant pressure, i.e. 2 MPa - Gas pressure due to the heating gas during the batch tritium recovery phase, i.e. 2 MPa.

A maximum stress level of 230 MPa has been found in the blanket channel with a stress intensity variation (due to the pulsed operation) of 115 Mpa, which corresponds to a fatigue life of 6x10 cycles.

VII.3.2. Disruption analysis

Disruption effects on the blanket box are closely connected to the eddy currents generated in the structure material. Major issues in the structural

TABLE VII.3.2-1 REFERENCE EM LOADS ON IB AND OB

IB OB

Quarter FR (MN) 1 10 FZ (MN) 2.55 20 Half FR (MN) 0.6 2.6 FZ (MN) 3.0 35 Peak pressure (MPa) 0.7 2.6

131 analysis of the blanket box are the values of electromagnetic (EM) forces, structural response, and internal support method. EM loads on the inboard and outboard blankets (IB and OB) have been evaluated by means of several numerical codes where some uncertainties exist in calculational model and the basic assumptions. Typical EM loads on IB and OB are shown in Table VII.3.2.-1. Stress analysis of the side module of the OB has been carried out under the disruption EM loads using a simplified rectangular prism model to examine the effect of the support conditions and reinforcement ribs on the stress and deformation. Distributed radial and uniform vertical EM forces are assumed, both of which are 10 MN in total. Four kinds of support (constraint) conditions (rear panel full, rear panel two-line, rear panel three-lines and side wall two-line) and three kinds of reinforcement ribs (lateral, toroidal and radial) have been examined. Results obtained by three-dimensional FEM analysis are summarized in Fig. VII.3.2.-1. Lateral and longitudinal reinforcement ribs are effective against the radial and vertical EM loads, respectively. For the three cases of the support conditions other than "side wall constraint" the effect of the support conditions is negligibly small. On the other hand, "side wall constraint" can minimize, the stress down to less than 60 MPa, and is lecommended to be applied for the blanket support design. Three-dimensional finite element analysis have been conducted on a typical poloidal blanket box segment consisting of four side walls and a first wall

0 WITHOUT RIBS 1~2 LATERAL AND RADI ^L RiB5 * LATERAL RIBS • LATERAL. RAOIAL ANDTOSOiDA

o o o O O o 80 n • • o o a: 40 [ft

FV=i10 MN FF MN FV • FR 0 1 1 1 1 | 1 1 i 1 | 1 1

1 1 • i 1 i CM => i i t u_ li- Li- 1 cr LU CC <

< ID E

CC EA R IDE - Q ar EA R CC EA R LLJ LU LLJ UJ CC LO •: in a: co cc LU LLJ rr CC cc rr CONSTRAINT CONDITIONS

Fig. VII.3-1 Summary of structural analysis of OB side machine under disruption EM loads

132 Toroidal length=0.8m, Pressure on SW=0.1MPa 5x4x5mm Walls, 3x3cm Toroidal Beam With Ties 1.2 T T T T

0.5 0.2 0.3 0.4 0.5 0.6 0.7 Pressure on First Wall (MPa) Fig. VII.3-2 Variation of the poloidal length of the blanket segment with the maximum pressure on the first wall.

stiffened by a toroidal beam with one or two radial ties attached rigidly to the shield. The results indicate that, in order to sustain a 1 MPa uniform disruption pressure at the mid plane, the poloidal length of the segment has to be ^. 40 cm. The poloidal length of the segments can be increased away from the midplane region depending on the disruption pressure (Fig. VII.3.2-2.). However, at the copper region, where there is a peak in the disruption pressure (0.6 MPa), the poloidal length of the segment should not exceed 50 cms.

VHATRITIUM RECOVERY AND INVENTORY

Several solid-breeder blanket options have been proposed for on-line tritium recovery and processing. In addition, a solid Pb-Li concept has been developed to be operated in the batch mode. These are described in Section VI. An important performance parameter for the on-line tritium recovery designs is the buildup and release of tritium during cyclic (physics and technology phase) operation. Table VII.4-1 summarizes the design information, operating conditions, and long-time, steady-state inventory for the solid-breeder blankets. These inventories are only reference numbers in that the pulsed nature of operation was not included in the calculations. Pulsing causes a lag in thermal response and a lag in tritium release response. The detailed input and calculational methods used to derive the results in Table VII.4-1 are presented in the remainder of this section, along with transient calculations for pulsed operation.

133 TABLE VII.4-1. SUMMARY OF SOLID-BREEDER PARAMETERS FOR ITER

Parameter Layered Layered^ BOTb^ BIT Blocks pebbles pebbles pellets

Material Li20 Li20 U20 LiA102

Form blocks pebbles pebbles sol.cyl.

Mass, MT IB 2.07 "3 OB 10.53 90 "40 ~50 Density/Packing Fraction, % 80 85-95/60 85-95/15 80 Li-6, % 95 50 7.4 Temperatures (C) at 100% Power

Physics Tmin/Tmax 502/609 Technology Tmin/Tmax 451/537 450/800 450/820 500/750 Tritium Generation Rate, g/day

Physics IB/OB 21.5/105.6 Technology IB/OB 18.5/92.7 10/119 13/119 106 He Purge

Flow rate (IB/OB), mol./s 0.53/2.71 0.31/2.79 0.31/2.79 0.33/2.45 Avg. temp., C 450 680 700 300 Pressure (in/out), MPa 0.2/0.1 0.11/0.10 0.101/0.100

H2, vol. % 0.2 0.817 0.817 '1

Avg. H/T ratio 30 100 100 100

Avg. outlet HT (HTO), Pa 12.4(0.69) 15.5(0.82) 15.5(0.82) 4(—) Tritium inventory, g

Physics (IB/OB) 0.42/2.18 Technology (IB/OB) 1.94/11.6 9.2/117b 18/240b

a) The generation rates assume a net tritium breeding ratio of unity. b) The breeder temperature ranges are assumed 500-850°C and 400-840°C for the layered and BOT pebble concepts.

134 TABLE VII.4-2. SUMMARY OF Be MULTIPLIER PARAMETERS FOR ITER

Parameter Layered Layered BOT BIT Blocks pebbles pebbles pellets

Form blocks pebbles pebbles hoi How pellets Mass, MT

IB 19 8 9 OB 187 185 130 — Density/Packing Fraction , % 65- 85 100/60 100/45 f90 Temperatures (C) at 100% Power

Physics Tmin/Tmax 130/471 ___ --- Technology Tmin/Tmax 123/422 450/800 <150 Tritium Generation Rate, g/day

Physics IB/OB 0.26/1.27 -_- Technology IB/OB 0.22/1.11 — — He + H2 Purge yes yes yes no Tritium Inventory, g

Physics (IB/OB) 5.5/26.9 --- ___ —

Technology (IB/OB) after 3 MW-y/m2 230/1130 --- —_ —

a) It will be reduced to ~ 500°C by a design optimization

Table VII.4-2 summarizes the information for the Be multiplier which also generates some tritium ( ~ 1% of the total). Detailed calculations have been performed for the layered design. The tritium generation rates were determined in the inboard and outboard Be during the physics and technology phases. These were multiplied times the number of full power days (24.3 days for physics phase and 1364 days for technology phase). For the purposes of tritium decay calculations, the physics phase was assumed to last 5 calendar years (giving a factor of 0.871 times the amount generated) and 10 calendar years for the technology phase (giving a factor of 0.764 times the amount generated). Based on recent data [1] for 100% dense Be in which < 2% of the tritium was released for T < 600C, it is assumed that all of the tritium is retained at these temperatures. Thus, the 3 MWa/m value of 1.36 kg is an upperbound in the sense that no tritium release was assumed. Data on porous Be is forthcoming which will shed some light on how conservative is the assumption of no release. Although no results have been reported for the other blanket concepts, it is assumed that the

135 low temperature Be retains all of its tritium, as well as the layered pebble design. The upper temperature of 700°C for the mixed pebble bed is a temporary number which will be lowered to <600°C by design optimization. The mixed pebble design involves a 3:1 mixture of Be: L^O pebbles. The higher operating temperatures may result in some release. However, a detailed calculation is required for this design which includes the effect of tritium recoil from the into the adjacent Be pebbles. In general, the end-of-life tritium inventory in the Be multiplier region for most designs should be in the range of 1.0-1.5 kg. This is far greater than the inventory estimates (Table VII.4-1) for the solid breeders. The main difference is that the solid-breeder maximum inventory is established relatively early and holds constant while in the Be it builds up slowly in time. In the following, a detailed description is given of the tritium inventory and release analyses for each of the blanket concepts.

1) Layered concept

The most detailed tritium inventory and release analyses were performed for the layered concept for both the physics and teclinology phases. The code TIARA [2] was used to perform a steady-state inventory analysis for the whole blanket. TIARA incorporates models for purge-flow thermal- hydraulics and chemistry, and steady-state inventory due to diffusion, desorption, solubility/ adsorption, and LiOT precipitation. Based on a 1-D (i.e., radial) temperature profile, it calculates a 2-D (i.e., radial and toroidal) partial pressure profile for He, H2, HT, H2O, HTO, and T2O. These are then input to the tritium analysis to give the tritium inventory in a breeder plate and purge outlet composition/flow-rate at a particular poloidal location. The analysis is repeated for other poloidal locations to obtain a whole-blanket analysis. For L^O properties which are well-known (e.g., diffusivity and moisture solubility) best-fit correlations are used in the models. For properties which have a high degree of uncertainty (e.g., desorption, hydrogen adsorption, and HT adsorption/solubility) conservative model parameters are used which are designed to produce an upperbound on tritium inventory. Sensitivity studies are also performed within the uncertainty range. Tables VII.4-3 and -4 show the results of the detailed inventory and purge-flow analyses for both phases at several poloidal locations. Table VII.4-5 is a summary of the whole breeder blanket inventory, including the uncertainty estimate. This uncertainty comes mainly from the factor of 4 spread in the data for hydrogen-isotope adsorption and solubility for reduced-form (e.g., H2, HT). The results presented in Tables VII.4-3 through VII.4-5 represent the behavior of the solid breeder under steady-state operating conditions. Because of the pulsed-nature of ITER operation, it is important to calculate excess inventory buildup because of thermal/tritium transport lags in the blanket and to calculate peak tritium release rates, which are important in the design of the tritium processing system. Two approaches have been used to obtain upperbound values

136 TABLE VII.4-3. LOCAL AND GLOBAL RESULTS FOR TRITIUM INVENTORY IN L^O UNDER STEADY-STATE PHYSICS AND TECHNOLOGY PHASE OPERATING TEMPERATURES AND GENERATION RATES FOR THE SINTERED-PRODUCT

Tritium Inventory, wppm

Phase Location(a) Desorp. Sol. Diff. Gas Phase Total

Physics IB Mi dplane 0.000 0.224 2.41E-4 7.66E-4 0.225 IB Average 0.000 0.202 2.61E-4 7.33E-4 0.203 IB Top 0.000 0.072 2.16E-4 3.35E-4 0.072

OBI Midplane 0.000 0.311 3.20E-4 1.05E-3 0.312 OBI Average 0.000 0.266 3.51E-4 9.78E-4 0.267 OBI Top 0.000 0.152 3.32E-4 6.53E-4 0.153

0B2 Midplane 0 .000 0.165 2.90E-4 6.60E-4 0.166 0B2 Average 0,.000 0.155 3.19E-4 6.52E-4 0.156 0B2 Top 0,.000 0.040 1.41E-4 1.93E-4 0.040

Tota 1 IB (grams) 0..00 0.42 0.00 0.00 0.42 Tota 1 OB (grains) 0..00 2 .18 0.00 0.01 2.19 Tota'1 IB+OB (grains) 0..00 2,.60 0.00 0.01 2.61

Technology IB Midplane 0.658 0..129 5.96E-4 6.•76E-4 0.788 IB Average 0.915 0.120 6.45E-4 6.60E-4 0.936 IB Top 0.202 0..045 4 .93E-4 3.03E-4 0.248

OBI Midplane 1.281 0. 183 7,.57E-4 9.33E-4 1.466 OBI Average 1.551 0. 161 8..34E-4 8.82E-4 1.714 OBI Top 1.114 0.097 7..52E-4 5.96E-4 1.212

0B2 Midplane 0.962 0.115 7.•24E-4 8.50E-4 1.079 0B2 Average 0.192 0.072 5.05E-4 4.27E-4 0.265 0B2 Top 0.000 0.026 3.20E-4 1.80E-4 0.027

Total IB (grains) 1.69 0.25 0.00 0.00 1.94 Total OB (grains) 10 .4 1.20 0.01 0.01 11.6 Total IB+OB (grams) 12.1 1.45 0.01 0.01 13.6 a) Inboard (IB), outboard plate #1 (OBI) closest to plasma, and outboard plate #2 (OB2) farthest from plasma.

137 TABLE VII.4-4. SUMMARY OF PURGE FLOW ANALYSIS FOR THE SINTERED- PRODUCT CONCEPT DURING THE ITER TECHNOLOGY PHASE WITH A He FLOW RATE OF 3.24 MOLES/S (2.80 x 105 MOLES/DAY) AN INLET PRESSURE OF 0.23 MPa, AN EXIT PRESSURE OF 0.1 MPa, AND 0.2% H2 AT PURGE INLET.

Exit Values

Parameter Location(a) H2 HT H20 HTO

Partial Pressure, Pa IB Midplane 190 13.3 6.29 0.720 0.0204 IB Average 190 12.9 6.11 0.678 0.0189 IB Top 196 5.96 2.91 0.146 0.0018 OBI Midplane 185 18.0 9.33 1.33 0.0527 OBI Average 186 17.0 7.91 1.17 0.0439 OBI Top 191 12.7 5.52 0.548 0.0136 0B2 Midplane 190 13.0 6.15 0.686 0.0192 0B2 Average 194 8.44 4.07 0.288 0.0051 0B2 Top 199 3.57 1.76 0.0514 0.0004

Total Flow IB 86.1 5.86 2.77 0.308 0.009 Rate moles/day OB 439 28.8 13.6 1.63 0.054 Total 525 34.7 16.4 1.94 0.063 a) Inboard (IB), outboard plate #1 (OBI) closest to the plasma, and outboard plate #2(OB2) farthest form the plasma) for pulsed inventory and peak release rate. The DISPL code has been used to calculate the transient inventory based on solubility and surface desorptior,. The calculation is consistent with the TIARA models and has been benchmarked to the TIARA steady-state results. Thus, it involves the slow- release models and properties used in TIARA to obtain upperbound inventories. Figures VII.4-la and -lb show the buildup of tritium as a function of the number of physics (600-s total cycle time) and technology (2490-s total cycle time) cycles, respectively, at the core-midplane, first-breeder-plate location. The transient results have been normalized to the steady-state (ss) results for constant- temperature, constant-generation. Thus, for the physics phase the transient inventory is only 1.25 times the values listed in Table VII.4-5, while for the technology phase the transient inventory is only 1.17 times the values in Table VII.4-5. Thus, the transient effects are within the uncertainties of the steady-state calculations and are not significant (at least for the corc-midplanc position). Additional calculations need to be performed for other poloidal locations to verify this.

138 100

Fig. VII.4-1. Tritium inventory and release rates la) Inventory II/Iss) and release (R/G) fractions (physics phase), lb) Inventory (I/Iss) and release (R/G) fractions (technology phase).

The MISTRAL code [3] was used to explore maximum values of tritium release rate during pulsed operation. This code has very detailed models for diffusion, desorption, and adsorption, along with models for gas-phase transport and compositional changes along the breeder interconnected porosity (radial direction). Solubility is not included, but low-temperature precipitation of LiOT is included. This results in excess tritium being retained during the start-up ramp when breeder temperatures are low and excess release of this tritium when breeder temperatures reach their full value. Within the range of model parameter uncertainties, the fastest tritium release rate parameters are used in the analysis to derive an upperbound on the transient release rate. The

139 TABLE VII.4-5. EFFECTS OF UNCERTAINTIES AND He PURGE PRESSURE ON THE TRITIUM INVENTORY RESULTS

Tritium Inventory, g Phase Case Inboard Outboard Total

Physics Baseline:

Upper Bound 1.54 8.30 9.48 Nominal 0.42 2.18 2.61 Lower bound 0.35 2.28 2.53

Low He Pressure/0.2% H2 (0.067 MPa in. 0.05 MPa out) 0.25 1.33 1.58

Technology Baseline:

Upper bujnd 6.87 37.8 44.7 Nominal 1.94 11.6 13.6 Lower bound 0.97 4. 70 5.67

Low He Pressure/0.2% H2 (0.067 MPa in. 0.05 MPa out) 3.68 20.3 23.9 calculation is performed for a radial slice of the breeder (from centerline to edge) with the appropriate purge partial pressures as input. Thus several calculations are needed in the toroidal purge flow direction to characterize one poloidal location. Figures VII.4-2a and -2b show the response of the solid breeder near the purge outlet (i.e., maximum tritium partial pressure) and the purge inlet (i.e.,minimum tritium partial pressure) for the technology-phase first-breeder plate at the outboard core-midplane position. Notice that at the inlet position the local, peak release rate is only ~ 1.5 times the generation rate, while at the outlet position it is ~ 5 times the generation rate (G). As the gas exiting this location reflects a spatial integration in the toroidal direction from Fig. VII.4-2b to Fig.VII.4-2a, the gas phase spike will be < 3 times the generation rate. The calculation would have to be repeated at each poloidal location to obtain an estimate of the peak blanket release rate seen by the processing unit. However, as the thermal, and hence tritium, transport lag increases in the poloidal direction, the blanket-integrated peak will be < 3 times the generation rate. An estimate of the peak tritium flow rate to the processing unit leads to < 2 times the steady-state value. While this is well within the range of the tritium processing unit design specifications, these conclusions need to be substantiated by additional calculations.

140 Figure VII.4-2c shows the release rate history at the purge-outlet breeder location (outboard core-midplane, first-breeder plate) during the physics phase. The peak release rate is ~5.5 G; the toroidally integrated peak value would again be ~ 3 G for rate of tritium exiting the purge at this location; and the overall release from the whole blanket would again have a peak of <2G, All of the TIARA and MISTRAL calculations are documented in detail in Ref.4. The DISPL calculations are documented in Ref. 2. The pebble concepts calculations were performed with the steady-state INVADER code [5]. For the layered and mixed pebble concepts, the breeder material (LLjO) is the same as for the layered concept and the purge composition and flow rates are similar (with the exception of ~3 times as much protium added). The main reason that the steady-state inventories are higher is because of the higher adsorption/solubility components at higher temperature and the larger breeder mass. In terms of average tritium concentration in wppm, the layered design has ~ \ wppm, while the layered and mixed pebble concepts have ~ 1.4 wppm and ~ 6 wppm. Thus, the two layered designs have about the same tritium concentrations based on a steady-state analysis. No transient analyses were performed for the pebble designs. However, it is anticipated that, at least for the layered concepts, the results will be similar. The BIT calculations vcre performed with the steady-state and transient TRIDYN coae [6]. Both temperatures and tritium retention/release are calculated. Repeated runs were conducted at various poloidal and radial locations to get a steady-state inventory for the outboard blanket of only 5 g. This translates into an average concentration of only 0.1 wppm in the LiAlC^. There are several differences between the layered and BIT calculation^ methods which result in such a low inventory calculation for LiAlO^ vs. ULO. In all in-reactor purge flow tests, Li^O exhibits a much faster tritium release rate and, hence, a much lower tritium inventory than LiAlC^. In the TRIDYN calculations, only diffusion and surface desorption mechanisms were considered as rate-limiting for release. The adsorption/solubility components were not included in the calculation as they were in the layered calculations. Also, the diffusion/desorption coefficients used were derived from the TEQUILA experiment and represent somewhat optimum properties for LiAlCK. Thus, while a broad data base was used to formulate the LUO properties and models, a very narrow data base was used for the L' tJO^ calculations. The uncertainties in the LiAlC>2 data base become very significant in regard to tritium inventory for T < 550°C. Future calculations should include the effect of these uncertainties. With regard to adsorption/solubility, these components are generally not significant for purge flow tests like TEQUILA because at the high purge flow rates (compared to tritium generation rates) the H/T ratio is quite high ( ~ 1000) and the partial pressures of HT and HTO are quite low. However, for ITER applications where typically H/T = 30-100 this might not be the case. Figure VII.4-3 shows the solubility of OH in LiAlOo vs. LL>O as a function of temperature at 10 Pa moisture pressure. This solubility in L1AIO2 is quite high at low temperature. For example, if all of the LiAlC^ were at 500^C with 10 Pa of T2O, the tritium inventory in 50 MT of material would be ~ 90 g. This number

141 Normalized Tritium Release (R/G) Normalized Tritium (ft/G) I 1 ^~j»»«- - 1 -

I i" mttmi"tti:it::n I 1 - f.'

.! I, '.V.V.V.V.'.V.V.V.'.V.V.*.1 u : : i i j i I •

5 - - c : m r. I 1 | s 4 ; - i : : I | * j i i 3 • : | : : | • j i : \ ; 1 ! '• ; 1 i • : j i : ; : i ; : : • j : i • 1 2 ! : ! | : - N • :

• j ; :

i : •••* • \ • •'" -f \ \ ] \. 1 1 I \ ! \| \ 1 Ni !: t o : i i j i ! 1! || il l : || : ; r i: ii : i n •; : • : i i ; • i : 0 0.5 1 1.5 2 Time, t (hr)

(c) Fig. VI1.4-2 Tritium transient analyses 2a) Tritium release history for the outboard LLO region 1 during the technology phas; with a 20 Pa tritium partial pressure in the purge, 2b) Tritium release history in the LUO outboard region 1 during the technology phase with zero tritium partial pressure in the purge, 2c) Normalized release history in the Li^O outboard region 1 during the physics phase with a 20 Pa tntium partial pressure in the purge.

£ Q. CL

O E x o

O CO

300 400 500 600 700 800 900 1000 TEMP CO

Fig. VII.4-3 Solubility of OH in Ll^O and LiAlO2 at 10 Pa moisture pressure.

143 would decrease with H or D addition to the purge and with purge flow rate. However, future calculations should take this component into account.

REFERENCES

[1] D.L. Baldwin, O.D. Slagle, and D.S. Gelles, "Tritium Release from Irradiated Beryllium at Elevated Temperatures," presented at ICFRM-4, Kyoto, Japan, December 4-8, 1989 (to be published in J. Nucl. Maters., 1990). [2] M.C. Billone, H. Attaya, Y. Gohar, and C.C. Lin, "Tritium Retention and Release Analysis foi the U.S.-ITER Driver Blanket," to be presented at the Ninth Topical Meeting on The Technology of Fusion Energy, Oak Brook, IL, October 7-11,1990. [3] G. Federici, A.R. Raffray, and M.A. Abdou, "MISTRAL: A Comprehensive Model for Tritium Transport in Lithium-base Ceramics - \Part I\: Theory and Descriptions of Model Capabilities," to be published in J. Nucl. Maters. (1990). [4] Y. Gohar et al., "U.S. Technical Report for the ITER Blanket/Shield: A. Blanket," ITER-TN-BL-5-0-3, July 1990. [5] INTOR Report, phase two A part II, p. 628-629, IAEA (1986) [6] A. Rossani and V. Violante, "Tritium Inventory and Permeation Analysis," presented at the ITER Blanket Workshop, Garching, FRG, 16-27 July 1990.

VII.5 POWER VARIATION CAPABILITY

1) Solid breeder concept Power variation capability of the solid breeder blanket concepts depends on the temperature limits and the normal operating temperature of the different blanket materials, and the method used for the breeder temperature control. Generally the breeder material has a temperature range for a satisfactory operating performance. Tritium inventory considerations or chemical compatibility with the steel clad define the lowest operating temperature for the breeder materials. The highest temperature is basically determined by material considerations and compatibility isssues. The temperature limits for the different materials are defined in section II. The extreme temperatures of the different materials in the outboard layered-blanket have been assessed at the midplane for a wide range of neutron wall loading. For the physics phase, the blanket can operate in the range of 0 to 150 % of the nominal ITER power level. That corresponds to about 0 to 2.3 MW/m . Low power operation results in higher tritium inventory. Full day operation at 50 % power level generates less than 68 g of tritium. Due to the very short period of operation, high power operation is only limited by the maximum breeder temperature. For continuous operation in the technology phase, the

144 lithium oxide material in this design has to operate within the temperature range of 350 C to 1000 C. The upper end of this range is reached at a 2.1 MW/m2 neutron wall loading (183 % of the nominal power level). The lower end of this temperature range insures a tritium inventory of less than 130 g for a full power day and corresponds to a 0.94 MW/m neutron wall loading. However, the neutron wall loading upper limit of 2.1 MW/m has to be decreased to 1.88 MW/m (154% of the nominal power leveJ) in order to get the maximum beryllium temperature less than 700cC which has been defined as design limit. Due to the helium-induced swelling, there can be a concern that this temperature is to kisure the beryllium integrity in case of using this material as a thermal barrier. Another consequence of the upper limit of the neutron wall loading is the high temperature of the stainless steel cladding ( ~ 700cC) which may be high due to the mechanical load induced by the differential tlu,iuial expansion of the materials in operation. The thermal performance of the multilayer concepts dependent on the gap conductance between the beryllium blocks and the other structures. The gap thickness considered in the thermal analysis is 100 (lm or less. Higher gap thicknesses could be produced due to the thermal deformations of the beryllium blocks and the first wall during operation. In this case, the used fraction of the temperature window is much than that considered in the power variation capability analysis, and consequently the margin for the power variation is lower. Therefore, future analyses and experiments are needed during the next phase. The thermal results of the layered pebble concept show that the low thermal conductivity values of the pebble materials require thinnner breeder and multiplier layers to reduce the maximum temperature of the breeder material. In the current configuration, the maximum breeder temperature of the last layer is 920°C. Design optimization can reduce this value to allow for power variation capability. In the mixed pebble concept, the breeder/multiplie temperature range is of 400 to 820°C. In this case, the maximum temperature increases to 950°C for 120% the nominal fusion power, from 1.2 to 1.44 MW/m . A fusion power decrease lowers the minium temperature below 400°C. The impact on the tritium inventory needs to be assessed where the lower operating temperature can be optimized for a satisfactory tritium inventory. Nevertheless, the mixed pebble concept might accommodate the power variation between the physics phase and the technology phase. With respect to the poloidal pin concept, the analysis has been performed for the first and last row of the first tube bundle and the first and last row of the second tube bundle at the equatorial plane and at the top/bottom of the blanket. The widest breeder temperature range at the nominal power level is 290°C (500 to 790°C) at the equatorial section of the First row of the second tube bundle. Due to the low beryllium temperature, the limiting factor to the power variation for this concept is the breeder temperature. For a power variation of + 20 %, the extreme breeder temperatures are 942°C (equatorial plane of the first row of die second tube bundle) and 433 C (equatorial plane of the first row of the first tube bundle) respectively. However, extreme capability for the power variation accommodation has not been assessed precisely and wider power

145 variations may be acceptable. A decrease of the fusion power will increase the tritium inventory which need to to be assessed.

2) LithL'in-lead concept

Thermal analysis has been performed for the inboard and the outboard blanket in the case of a power increase of 55 % for the physics phase and 86 % for the technology phase. Keeping the same cooling parameters as for the nominal power level, the maximum eutectic temperature are 338°C and 309°C at the inboard and the outboard blanket respectively for the physics phase and 341 C and 303 C at the inboard and the outboard blanket respectively for the technology phase. Also, a large volume (10 L) of the eutectic is melted in the inboard and outboard equatorial channels which is acceptable from the operational point of view. With the same cooling parameters, eutectic boiling is expected for a power equal to 2.1 times the nominal power level. The corresponding maximum temperature of the eutectic is 375°C, but a 4 mm-thick solid layer is still present between the melted zone and the channel case. Furthermore, a power decrease does not affect this concept more than for the nominal operation conditions since the main option for the tritium recovery is a batch mode. Therefore, the Lil7Pb83 concept has good capabilities for power variation accommodation.

VII.6 SAFETY ANALYSIS VII.6.1 LOCA Analyses of the BIT Concept 1. Assumed scenario: the blanket LOCA is caused by a break in the large pipes or the small pipes of the breeder region. The later case is the probable cause for LOCA. The affected segment experiences a transient which may be very severe, while the other segments are not affected. The change in the hydraulic paramters such as loop pressure may be negligible which does not generate a signal to shut down the plasma 2. Accident definition: The most critical case is a rear channel break at lower the manifolds, which implies: - direct flow increase in the broken channel; - direct flow increase in the intact rear channels, - flow reduction in the intact front channels; - discharge of fluid in the affected segment box. 3. Basic assumptions: - a small volume steam pressurizer is assumed; - double ended guillotine breaks are istantaneous; - plasma operates at nominal power during the accident; - the purge helium circuit remains intact. 4. Conclusions: - the break flow rate is too small to generate a signal to shutdown the plasma operation, however, the following two signals could be used: a) high flow unbalance between inlet and outlet legs in jae. of the 48 segments, b) high liquid level in one segment box;

146 - following a rear channel break, flow reduction in the intact front channels is about 40%; - even for the critical break, no void is generated before 6 seconds (see Fig.VII.6.1); - segment boxes are flooded and pressurized to one MPa in a short time; adequate protection is needed. 5. Special considerations for a LOCA generated by a break of the water-helium barrier. This accident implies flooding in the helium purge circuit, steam generation due to heat exchange with breeder tubes, helium circuit pressurization, cooling and eventually rewetting of the breeder tubes. The analyses show that the maximum pressure remains below the design limit, maximum temperature is 820 K, and rewet within 8 seconds.

VII.6.2. LOFA Analyses 1) The LAYERED concept 1. Assumed scenario: LOFA is caused by catastrofic mechanical failure or loss of power in all coolant pumps. 2. Methodology for catastrofic mechanical failure: - the flow is assumed to stop completely, - the effect of natural convection is small, due to the radial and toroidal direction of the coolant channels; - the heat transport mechanism is mostly conduction; - temperature histories are calculated for different plasma operating periods before shutdown at nominal power using the TOPAZ code. 3. Methodology for loss of power in all coolant pumps: - hydraulic analyses are used to characterize the tapering of the mass flow rate. - the water and solid breeder temperature histories are calculated as a function of the pump inertia and the plasma shutdown characteristic times. - the time was estimated for the water coolant to reach boiling and/or the breeder material to reach its maximum allowable temperatures 4. Conclusions - global natural convection has no effect following a LOFA; - the breedr temperature following a LOFA falls as soon the plasma is turned off. The coolant temperature limit is then binding constraint. - in the extreme case where the flow stops completely, the water coolant starts boiling after about 25 seconds (see Fig.VU.6-2). Corrective action, such as putting back-up pumps on line or starring an emergency heat removal system would need to be implemented within these 25 seconds; - in the case of loss of power accident, the pump inertia can be helpful (see Fig. VII.6-3), in retarding the coolant from reaching its boiling point. A pump half time of about 25 would provide several minutes for corrective action to be taken before the coolant reach its boiling point

2) The PIN concept 1. Assumed scenarios: LOFA can occur due to pump trip, loss of power, or pump seizure. Channels full of stagnant water can be developed due to break of large pipes. These accidents can be simulated as LOFA with an atmospheric loop pressure.

147 LOCA ANALYSIS: CASE OF THE PIN CONCEPT 00 386

350 3 4 3 TIME (S) Fig. VII.6-1 Front channel coolant temperature •Hit) 0-

400

CD •120 0 a ir. plasmn on a> 400.0 sec (h.ick) c ro 5 r.oc (back) o o O 380.0 10 sec (back) 0 sec (front) 360.0 0.0 10.0 20.0 Time following a LOFA (sec)

Fig. VII.6-2 Coolant Temperature History Following a LOFA with complete flow stoppage for different times during which the Plasma stays on.

a 2 o. E 0) 450 -

350 -f 20 40 60 Time Following a LOFA (s)

Fig. VII.6-3 Coolant Temperature at Exit of Breeder Region Following a LOFA for Ddifferent pump half-times (for a blanket thermal time constant of 600 s and with the plasma immediately turned off following the LOFA)

149 LOFA ANALYSIS: CASE OF THE PIN CONCEPT

.9 -

.8- Pump Seizure

,7 • (1.0 MPa) 55 o t-i .6 t- Large Break O a-» (0.1 MPa) Ex

Q ,4 O .3--

2 -•

. 1 • •

0. •Q- -B- -t— —I— 10 20 30 70 60 TIME (S) EC-Fig. VII.6-4 Front channel rod temperature 240- T, C Tmax

190-

140- Tmin

„ y 90-

40 • i i t i i i i i i i i i i i i 11 i i i i i i i i i i i i | i i i i i i i i i i i i i i i i i i i | i i i i 0 200 400 600 800 1000

Fig. VII.6-5 Temperature evolution of the first outboard row channel

2. Basic assumptions: - flow drops to zero in 2 seconds; - plasma operates at nominal power. 3. Main results: - Void generation takes place after 16 sceconds and breeder tubes dryout about 10 seconds later. This implies a breeder rod temperature increase of 150 K; - Breeder tube temperature remains below 1000 K (see Fig. VII.6-4); VIL6.3. Safety Analyses for the Lithium-Lead Concept 1) List of initial events for the blanket and its cooling and heating systems

LOCA SCENARIOS 1. Piping failure in the cooling system loop 2. Break of piping or headers inside the segment box 3. Failure of the breeding channels cladding 4. Break of piping or headers under gas heating mode. 5. Line failure in the tritium circuit.

LOFA SCENARIOS 1. Shutdown of one of the two circulating pumps 2. Failure of the reverse value of one cirulating pump 3. Decrease of water flow rate inside either blanket zone. 4. Loss of power

151 TEMPERATURE ISOUNE

0— 190 1 — 189 2— 187 3~ 185 M~ 183 5— 181 6~ 179 7— 177 8~ 175 9— 173 10— 172 11 — 170

TEMPERATURE FIELD IN THE CHANNEL OF IB (EQUATORIALCROSS-SECTION. OPERATION SCENARIO Bl) 1000 SECONDS PAST EMERGENCY

ThMPERATL'RE 1SOLINE

0— 187 1 — — 185 2— 183 3— 181 H— 179 5— 178 6— 176 7— 17H 8~ 172 9~ 170 10— 168 11 — 166

TEMPERATURE FIELD IN THE 1-ST ROW CHANNEL OF OB (EQUATORIALCROSS-SECTION, OPERATION SCENARIO Bl) 1000 SECONDS PAST EMERGENCY Fig. VII.6-6 Temperature field at 100 seconds

2) Break of one cooling system loop discharge header

The computation results show that the cooling systems of the breeding zone have enough safety margin, because temperatures in the channel do not reach the melting temperature of the eutectic, even in the case of the header rupture without the emergency pumps as shown in Fig. VII.6-5.

3) Water-eutectic interaction

1. Assumed scenario: The interaction between the cooling water and the eutectic may occur due to a leak or a complete mechanical failure of the cooling tubes inside the breeding region.

152 2. Experimental results: Water-eutectic interaction has been investigated at the BLAST facility (JRC, Ispra, Italy). On the basis of these investigations the following conclusions were made: - the pressure increase in the reaction zone does not exceed the pressure of the water being injected in the melted eutectic; - the formation of LLO and LiOH at the water-eutectic boundary moderates the reaction and reduces hydrogen yield. Vffl. SHIELD

VIII.l DESIGN LIMITS AND SAFETY FACTORS

The shield components of ITER have to satisfy two major requirements: - The magnet shield system has to adequately protect the Toroidal Field Coils (TFC) against excessive nuclear heating and against particular radiation loads to which the various components of the coils are sensitive. - The biological shield system has to ensure that during operation both the personnel and the public are adequately protected according to international regulations. It has also to allow personnel access to designated areas inside the reactor hall 24 hours after shutdown.

TABLE VIII.1-1: RADIATION LIMITS AND SAFETY FACTORS Safety Factors Toroidal Field Coils Limits ID Analysis*) 3D Analysis Total nuclear heating 55 kW 2 1.4 Peak nuclear heating in winding , pack 5 mW/cm 3 1,5 Peak dose to electrical q insulator 5.10 rad 3 1.5 Peak fast neutron fluence (E>0.1 MeV) .„ ~ to NbnSn superconductor 1.10 n/cm 3 1.5 Peak displacement danv.ge in o Cu stabilizer 6.10° dpa 3 1.5

Biological dose rate 0.5 mrem/h 10 10

*) The safety factors 2 and 3 are valid for all locations except the outboard! side, where they should be higher

The various radiation limits valid in the final stage of the ITER Conceptual Design Phase are summarized in Table VIII.1-1. This table lists also the safety factors which should be applied to calculated results depending on whether they have been obtained from one- or three-dimensional analysis. These safety factors account for uncertainties in the cross-section data, imperfections in modelling, and, in particular for ID calculated numbers, the influence of the assembly gaps in between adjacent blanket segments. They are based on a set of correction factors which is listed in Table Vm.1-2. These factors distinguish between local and integral responses, but do not differentiate between the various types of local responses.

155 TABLE VIII.1-2 RECOMMENDED CORRECTION AND SAFETY FACTORS FOR ITER SHIELDING ANALYSIS

ID Analysis 3D Analysis Responses Local Integral Local Integral

Correction factors for: Assembly gaps 1.7 1.2 __* * Modeling 1.3 1.3 1.1 1.1 Uncertainties in Xn data 1.4 1.3 1.4 1.3 Safety factors for 3 2 1.5 1.4 inboard and divertor regions

Safety factors for outboard regions >3# >2#

Safety factors for Biological shield 10 10

* Gaps included in 3-D models # Outboard blanket/shield design dependent

For a big part of the ITER magnet shield design, the recommended safety factors are 3 and 2 for the local and the integrated 1 D results, and 1.5 and 1.4 for the respective 3 D results. On the outboard side they are recommended to be higher because of the more significant gap effect. For the biological dose calculations a factor of 10 is suggested with respect to the relatively higher inaccuracy in the calculations. The specific dimensions of the assembly gaps (20 mm width, 520/1200 mm radial depth on the inboard/outboard mid plane) make this set of safety factors unique for the present ITER Reference Design. Therefore, they are not readily applicable to other machines. The safety factors listed in Table VIII.1-2 have been agreed among the shielding experts of the various ITER parties. They have also been approved by experts in shielding experiments [1] with the reservation that they need to be verified by an appropriate international R&D programme. This programme foresees experimental investigation of bulk shield configurations as well as of typical penetrations through a thick shield assembly. They are to be performed at the various existing and forthcoming 14 MeV neutron generators available to the four ITER parties.

156 VIII.2 NUCLEAR RESPONSES IN THE TF COILS

VIII.2.1 General Remarks

In the course of the ITER Conceptual Design Phase the poloidal shield coverage and the cross-sections of the vacuum vessel as one of the most significant shield components have been specified after frequent iteration, in order to balance shielding and engineering requirements. The poloidal distribution of the various shielding structures is shown irt Figure VIII.2-1, a cross-section through the inboard and the outboard section of the vacuum vessel at the level of the equatorial plane is given in Figure VIII.2-2. All contributors to the shielding analysis have essentially adopted this outer structure for their neutronics calculation models, but included different types of blanket/shield

UP PjE R VACUUM VESSEL D I V £ Z\ T OR PLATE p, HIE

OUTBOARD BLANKET TOROIDAL FIELD C qiL WINDIN PACK

INBO A R CD BLANKET

VACUUM PUMPING DUCT ATERAL SHIELD LOWER Dl V E R T OR PLATE SHIELD

FIG. VIII.2-1 Illustration of the Poloidal Shield Coverage

157 158 CROUN (1:5)

K»i«^tf mm

(MICE VACWft VESSEL 0 IOOO MO CMW SfCHO. ai , -n^ SOT VMI 101 \

FIG. VIII.2-2 Inboard and outboard Vacuum Vessel Cross Section on the Equatorial Plane segments according to the options described in Section VI. This is the main reason for the differences in the numbers for the various radiation responses reported in Section VIII.2.2. Further reasons are the differences in the calculational tools and in the parameters of their application, in the data base, and in a number of details of the blanket, shield, and coil configuration, composition, and representation. Most of the results were obtained from ID calculations. This method was extensively used for detailed layout and optimization studies. In general, codes like ANISN or ONEDANT have been used, mostly in approximation, the calculations have usually been done in cylindrical/toroi geometry, with the main machine axis as the axis of symmetry. This configuration allows to calculate inboard and outboard responses simultaneously and so to take into account the mutual interaction between the two sides of the machine. Careful source calibration has been done in each case to permit comparison both among various 1 D results and with those from more-dimensional calculations. In the later stage of the Conceptual Design Phase, also 2 D and 3D analyses were performed to assess the impact of design features not accessible to ID calculations (e.g. the assembly gaps) and to confirm the safety factors stated above. For 2D calculations, the DOT 3.5 and the TWODANT code were applied, which rely, like the 1 D codes mentioned above, on the discrete ordinates method. For 3 D calculations, the Monte Carlo code MCNP was used which uses a continuous energy coupled neutron-photon library. The data base for all the calculations was widely different. Depending on the availability to the users, data from ENDF/B-IV, ENDE/B-V, JENDL-3, and EFF-1 were used. Tool and data base effects on the resuls are commented on in Section Vm.5.

VIII.2.2. Comparative Analysis of all Relevant Results

In this Section, the most important results from all calculations are summarized. The various parts of the magnet shield system that have been considered in the analyses are: - the inboard region, - the recess region above and below the straight inboard section, - the outboard region, - the regions around the top and the bottom divertor.

VIII.2.2.1 Radiation loads on the inboard leg

The inboard shield configuration on the equatorial plane consists of a 300 mm thick vacuum vessel, including 50 mm of special shield material, and a 520 mm thick blanket segment, separated by a 20 mm "ide gap. The results for the radiation loads on the inboard leg as reported in [2,3,4,5] are not in complete agreement. The main reason is that different blanket designs (see Section VI) were included into the overall shield configuration, while

160 TABLE VIII.2-1: DIFFERENCES IN ASSUMPTIONS FOR INBOARD SIDE

Reference [2] [3] [4] [5] WP-Composition - SS 0.405 0.316 0.316 0.405 - Cu 0.220 0.355 0.260 0.220 - non-Cu 0.080 0.095 0.080 - He (liquid) 0.183 0.211 0.211 0.183 - Ins. 0.112 0.118 0.118 0.112 CC-Thickness (cm) 7 7 6 7 SM-Pb+B.C (+SS) 4 + 1 4 + 1 3 + 1(+1) 4 + 1 VV-Representation layered homogenized layered layered (SS:H90-70:30) RE-Thickness (cm) 4 5l 4 4 - SS fraction 0.75 0.2 0.85 1.0 SH-Thickness (cm) 30.8 37.0 38.4 32.3 BL-Type Pin Type Mixed pebble bed Multilayer Multilayer - thickness (cm) 16.7 10.5 10.1 15.2 FW-Structure Thickness (cm) 2.5 1.5 1.5 2.5 FW-Protection Thickness/Material Physics Phase 2.0/C 2.0/C 2.0/C 2.0/C Technology Phase 2.0/C 2.0/C 0.05/W 2.0/C

Abbreviations: WP - Winding Pack RE - Resistive Element CC - Coll Casing SH - Shield SM - Special Shield Material BL - Blanket VV - Vacuum Vessel FW - First Wall

the assumptions for the vacuum vessel and the TFC are virtually the same. There are a few minor differences in the rest of the assumptions which are summarized in Table VIII.2-1. These should, however, not have a significant impact on the calculated numbers. The results for the inboard leg are listed in Table Vm.2-2. The peak values given there refer to the midplane position. The fluence dependent quantities are given for an operation time of 3 full power years (FPY) . The numbers presented in the table do not include the safety factors. In this Section, only the peak values are discussed; the total nuclear heating is the subject of Section VIH.2.2.5. In the first four columns of Table VII.2.2 the results obtained from ID calculations are reported in which the 50 mm thick special shield material layer backing the inboard part of the vacuum vessel was assumed to be made of the reference materials, i.e. lead and boron-carbide. The differed ,"s in the local peak values amount to about a factor of 2 to 3. It is evident, that the higher numbers are characteristic for configurations which include blankets with a relatively high void volume fraction, while the lower numbers go along with highly dense blankets. It should be noted that the high void volume fraction is an inherent feature of a pebble bed blanket. In case of a pin type blanket, it can be avoided in principle, but it has been chosen here for engineering simplicity.

161 TABLE VIII.2-2: RADIATION LOADS ON INBOARD LEG (TECHNOLOGY PHASE OPERATION = 3 FPY)

ID 2D 3D ID Reference [2] [3] [4] [5] [5] [2] W W Special Shield Material Pb/B4C Pb/B4D Pb/B4C Pb/B4C Pb/B4C Pb/B4C W W Total nuclear heating [kW] - Physics Phase 9.3 7.3 5.5 4.9 8.3 16.0 11.1 5.0 - Technology Phase 7.1 5.4 5.0 3.9 6.5 1?.5 8.7 4.5 Peak nuclear heating [rrtrf/cm ] - Physics Phase 1.12 0.53 0.47 - - Technology Phase 0.86 0.39 0.57 0.36 0.9 0.6 0.43 Peak dose to insulator [10 rad] 2.3 2.5 1.7 1.2 2.2 2.7 2.2 1.3 Peak fast fluence to SC [1018 cm'2] 2.7 6.7 1.9 3.7 - - 3.3 1.5 Peak displacement damage [10 dpa] 1.7 1.7 1.0 1.0 - 1.5 - 0.7

Problems with respect to the radiation limits occur predominantly for the dose to the insulator, and this is more or less valid for all types of blankets considered. For this reason it is proposed to replace the Pb/BAZ layer by a tungsten layer in the region close to the midplane. For this specific case the ID results are reported in the last column of Table VIII.2-2 [4]. This improvement applied to the other types of blankets would bring the respective numbers also very close to the stated limit. The columns in between list the results from comparable 2D and 3D calculations. It can be noted that the discrepancies between those and ID calculated numbers are less than a factor of 2, except for the fast fluence to the superconductor. This is, however, one of the less critical responses. If the respective safety factors are applied tc the 2D and 3D results, radiation loads well within the limits are found.

VIII. 2.2.2 Radiation loads in recess regions

In the recess regions (see Fig. VIII.2-1) the total shield thickness is only 693 mm, as compared with 840 ma on the equatorial plane. There is no breeding blanket in this region; only the shielding components of the blanket segment extend into that part of the configuration. Inspite of the smaller thickness the radiation loads on the TFC are lower than on the equatorial plane. This is because the local wall loading in this position is significantly lower as a consequence of the general poloidal distribution of the wall loading, assisted by a screening effect caused by the protruding edges of the segment at the top and the bottom ends of the inboard breeding blanket. The radiation loads as calculated by ID and 3D methods are summarized in Table VTII.2-3. There is reasonable agreement between most results, even between ID and 3D values. Only the results listed in the first column show a major deviation. These value;: are based on the assumption that

162 TABLE VIII.2-3: RADIATION LOADS ON RECESS REGION (TECHNOLOGY PHASE OPERATION = 3 FPY)

ID 3D Reference [2] [4] [5] [2] [4] Total nuclear heating [kW] - Physics Phase 0.3 2.1 1.9 1.7 2.5 - Technology Phase 0.2 1.6 1.5 1.3 2.0 Peak nuclear haating [mW/cm ] - Physics Phase - Technology Phase 0.04 0.31 0.51 0.20 0.24 Q Peak dose to insulator [10 rad] 0.1 0.9 1.3 0.8 1.0 Peak fast fluence to SC [1018 cm"2] 0.13 1.0 1.7 - 1.4 Peak displacement damage [10 dpa] 0.08 0.6 1.3 0,4 - the breeding blanket ends with a sharp edge at the upper and lower end which increases the screening effect, while in all other models an inclined edge is assumed. None of the numbers exceeds any of the limits.

VIII.2.2.3 Radiation loads on the outboard leg

Because of the larger thickness of the outboard shield configuration as compared with the inboard one (1510 mm vs. 840 mm on the midplane) the radiation loads on the TFC on the outboard leg should be appreciably lower. ID calculations prove this statement as can be seen from the results listed in Table VIII.2-4. The numbers are well below the limits although it must be recognized

TABLE VIII.2-4: RADIATION LOADS ON OUTBOARD LEG (TECHNOLOGY PHASE OPERATION = 3 FPY)

LD 30 Reference [2] [4] [2] Total nuclear heating [kW] - Physics Phase 2.3 0.03 2.4 - Technology Phase 1.9 0.02 1.9 Peak nuclear heating [mW/cm ] - Physics Phase - Technology Phase 0.5 0.009 Peak dose to insulator [rad] 9.105 2.104 Peak fast fluence to SC [cm"2] 7.1014 2*1013 Peak displacement damage [dpa] 4.10"7 8*10"9

163 TABLE VIII.2-5: DIFFERENCES IN ASSUMPTIONS FOR THE CRITICAL DIVERTOR REGION

Reference [2] [4] Opt.l [5] [4] Opt.2 CC-Thickness (cm) 25.0-31 .0 27.0-30.0 25.0-29.0 27.0-30.0 Poloidal Width (cm) of high damage region 158 72 223 72 Srt-layer - material Pb=B.C W+SS Pb+B.C Pb+^.C+SS - thickness (cm) 44 4+1 4+1 4 3+1+1 SH+VV thickness (cm) 47.5-84.0 47.0-56.0 47.5-57.0 47.0-56.0 Divertor - plate material C+Mo/Re+H,0 W/Nb+H,0 C+Cu+H,0 W/Nb+H,0 - plate thickness [cm] 2.2 1.32 1.3 - structure thickness [cm] 3.3 -- 4.0 11.5.0? Wall loading (MW/nf) 0.34 0.1-0.46 0.17-0.43 0.1-0.46

that there are again significant differences originating from the different types of blanket assumed for the calculations. 3 D calculations for the BIT Pin Blanket Concept are in full agreement with the 1 D calculations. This is a surprising result and needs careful checking.lt was found that the major part of the TFC total nuclear heating is due to neutrons passing through the assembly gap.

VII1.2.2.4 Radiation loads in the divertor region

In the divertor region, the TFC is shielded by the vacuum vessel and a shield component of variable thickness in poloidal direction (see Fig VIII.2-1). The minimum total thickness is 470 mm at the outer edge of the 22° inclined divertor plate. The divertor plate and its supporting structure contribute additionally to the shielding effect, but it is questionable whether or not credit can be taken also of the collector tubes situated in between the plate structure and the shield component. Depending on the assumptions made, which are summarized in Table VIII.2-5, different results for the local responses have been calculated. They are reported in Table VIII.2-6 in the same way as for the other regions discussed above. In the first three columns the results from ID calculations are reported which refer to the case that the special shield material layer on the back of the vacuum vessel is extended up to the divertor region, and made of the reference materials lead and boron-carbide. With few exceptions, the numbers calculated are beyond the limitations. Therefore, it seems to be unavoidable to replace lead and boron-carbide by tungsten as was also suggested for the inboard side close to the equatorial plane. ID calculations for this case (i.e. column 6 in Table VIII.2- 6) show still too high radiation loadf in terms of insulator dose and fast fluence. 3D calculations for the same case (see column 7) yield in fact lower numbers which are within the limits. At this moment, it is, however, still questionable whether these results are accurate enough because it is difficult in this complex geometry to obtain precise results for a restricted volume. It should be noticed that all calculations discussed so far are based on the assumption that during the Technology Phase the same divertor is used as in

164 TABLE Vm.2-6: RADIATION LOADS IN DIVERTOR REGION (TECHNOLOGY PHASE OPERATION = 3 FPY)

10 20 3D ID 3D ID Reference [2] [4] [5] [5] [2] [4] [4] [4] Special Shield Material Pb/B4C Pb/B4C Pb/B4C Pb/B4C Pb/B4C W W New Div Total nuclear heating [kW] - Physics Phase * 9.3+9.3 19.0 13.0 8.5+6.8 13.0+9.0 - 14.0+10.0 14.8+15.7*J 9.0+9.0 - Technology Phase 7.3+7.3 15.0+10.0 6.7+5.4 10.0+7.0 11.0+8.0 11.6+12.3 ) 5.0+5.0 Peak nuclear heating [mW/cm ] - Physics Phase 1.7 1.6 - - Technology Phase 1.4 2.0 1.3 0.5 1.6 0.5 - 0.6 Peak dose to insulator [10 rad] 4.0 4.6 2.5 1.5 1.4 3.6 1.4 1.4 Peak fast fluence to SC [1018 cm"2] 4.7 5.4 3.2 3.0 - 4.0 2.3 1.7 Peak displacement damage [10 dpa] 2.5 2.2 1.8 1.5 1.9 1.6 - 0.7

N The two numbers refer to the upper and lower divertor region, respectively. ' Includes lateral shields on pumping duct ,-, 2.0

Pfus = 1000 MW

— 1.5 en

0.5

• Inboard- -Oiverfor- Outboord i -90 90 —- Pol. Angle [°] FIG. VIII.2-3 Poloidal Variation of the Neutron Wall Loading the Physics Phase. Recently, a specific divertor design for the Technology has been agreed, the characteristics of which are included in Table VII.2-5 ([4J, Opt.2). The last column of Table VIII.2-6 lists first ID results which indicate satisfactory shielding performance even in the case of lead and boron-carbide as special shield materials.

VIII.2.2.5 Total nuclear heating

The individual results for the total nuclear heating in the various parts of the TFC system have been included in the previous Tables VIII.2-2,-3,-4, and -6. They were obtained, in general, from multiple ID calculations and an appropriate integration procedure which takes into account the poloidal variation of both the neutron wall loading and the thickness of the various shield components. The poloidal variation of the wall loading which is shown in Fig. VIII.2- 3 was found by either ray-tracing or Monte Carlo methods. Reasonable agreement between the results from the two methods has been achieved. Discrepancies of the order of 5% for the midplane values occurred due to differences in the neutron source specification. The results based on this procedure have in a later stage of the Conceptual Design Phase been checked by 2D and 3D calculations. It has been found that for the inboard leg the 3D analysis yields about twice the total nuclear heating as compared to the ID based analysis. For the recess, the outboard, and the divertor regions reasonable agreement between ID and 3D analysis has been obtained.

166 TABLE VIII.2-7: TOTAL NUCLEAR HEATING IN TFC [kW] ID 3D Physics Phase - inboard 4.9 - 9.3 11.1 - 16.0 - recess 1.9 - 2.1 1.7 - 2.5 - outboard 0 - 2.3 0 - 2-4 - top divertor 8.5 - 14.0 14.8 - bottom divertor 6.8 - 10.0 14.5 - pumping duct - 1.3 - Total without safety factor 22.1 - 37.7 43.4 - 51.5 - Total with safety factor 44 - 75 61- 72 Technology Phase inboard 3.9 - 7.1 8.7 - 12.5 recess 1.5 - 1.6 1.3 - 2.0 outboard 0 -• 1.9 0-1.9 top divertor 6.7 - 11.0 11.6 bottom divertor 5.4 - 8.0 11.3 pumping duct - 1.0 Total without safety factor 17.5 - 29.6 33.9 - 40.3 Total with safety factor 35 - 59 47 - 56

In Table VIII.2-7 the final results for the total nuclear heatiog are summarised in terms of ranges which cover the various options for both the blanket concepts, the divertor designs, and the special shield material options. With the safety factors applied to the ID and 3D results, it turns out that for the Physics Phase a total nuclear heating of about 60 to 70 kW and for the Technology Phase about 45 to 55 kW have to be expected. From these numbers it seems that the Technology Phase conditions fit reasonably well to the stated limits of 55 kW. In the Physics Phase, however, the limit is significantly exceeded.

VIII.2.3 Three-Dimensional Analysis for a Specific Configuration

In the previous Section it was attempted to arrive at conclusions on the basis of all studies performed so far, which still involved different calculational methods and a number of different assumptions and shield configurations. In this Section, a selfconsistent set of 3D-results is reported instead which is based on the Layered Blanket Concept as outlined in Section VI.1. A similar analysis including the BIT Pin Concept (see Sect. VI.3) is in progress, but could not be completed in time. The analysis has been performed to determine the total nuclear heating in the TF coils and the end of life magnet radiation effects in areas with critical shielding space. The continuous energy coupled neutron-gamma Monte Carlo code MCNP has been used in the calculations with the ENDF/B-V cross section data evaluation. Two sets of calculations that complement each other have been performed. In the first calculation, the upper part of the reactor was modeled

167 with detailed magnet configuration in the inboard, outboard, and upper divertor regions. In the second calculation, the lower part of the reactor was modeled in detail to determine the radiation effects in parts of the TF coils adjacent to the vacuum pumping ducts and divertor coolant tubes. In the model of the upper part of the reactor, the detailed configuration of the layered blanket, the shield, vacuum vessel, coil case, winding pack and a 1.3 cm thick divertor plate was included, the poloidally varying blanket thickness, the layered inboard shield, and the 2 cm wide assembly gaps were modeled in detail. Due to symmetry, 1/64 of the reactor was modeled with reflecting boundaries. The magnet was divided into 6 segments in the inboard region, 6 segments in the divertor region and 3 segments in the outboard region in order to determine the poloidal profiles of the magnet radiation effects. The magnet segments are indicated in Fig. VIII.2-4. which shows a vertical cross section of the geometrical model used in the calculations. B4C/Pb shield is used in the 5 cm thick back layer outside the vacuum vessel in all regions except in regions II, D4 and D5 where W is used because of the limited shielding space. 100,000 source particles were sampled using the actual source profile in the D-shaped toroidal plasma. Several variance reduction techniques were utilized in the calculations. Surface flux tallies were used to determine the peak radiation effects at the irmer surface of the TF coils. The radiation effects at the TF magnets are given in Table VIII.2-8 for the different segments of the coils in the inboard and divertor regions. The D2 D3 D4 D5

w.p. BACK LAYER OF SHIELD FIG. VIII.2-4 Vertical Cross Section of the Geometrical Model Used in the 3D Analysis (Output from MCNP plotting routine)

168 TABLE VIII.2-8. RADIATION EFFLCTS IN THE TF COILS (SAFETY FACTORS INCLUDED)

Insulator Fast n Fluence Power Nuclear Heating in Technology Phase Region Dose (n/cm') Density .P. C.C Total (rads) (mW/cnr) (kWw ) (kW) (kW)

Inboard

11 4.2+9#(.O9)* 6.1+18 .08) 0.92 .10) 1 .80(.07 1.20 .07) 3 .00 .07 12 3.6+9 .06) 5.0+18 .07) 0.83 .08 3.00 .06 2.30 .06 5.30 .06 13 1.9+9 .20) 2.6+18 .17) 0.44 .20 1.54 .17 1.26 .16 2.80 .16 14 4.6+8 .14) 6.7+17 .15 0.12 .14 0.56 .12 0.50 .11 1 .06 .11 15 1.4+9 • 12) 2.1+18 .11 0.30 >14 0.78 .10 0.56 .10 1.34 .10 16 1.8+9 [.12) 2.7+18(.11 0.36 .13 .10 0.62 .10 .40 (.07 ) O4(.07 ;.jo

Upper Divertor

Dl 2 8+8( .12) 5.2+17 (.11) 0.06(.13 ) o ) 0.41 76(.11) D2 4 4+7 .19 7 5+16 .23 0 02(.20 0 06'.17 0.12 f.17 i 0 18 .17) D3 7 8+8 .09 1 3+18 .08 0 2O(.1O 0 29 .07) 1.27 :.o?) 1 56 .07) D4 2 6+9 .05 4 3+18 .05 0 66(.O6 0 87,.05 5.67 [.04 6 54 .04) D5 2 0+9 .06 3 4+18 .06 0 50f.07 0 921>.O5 5.38 [.05 | 6 30 .05) D6 5.6+7( .15 9 1+16 .13) 0 0Z(.15 0.05( 0.80 .08 85 .08) 7"§4 -05* 1335 04 lFT§ .04) Lower Divertor

Dl 2.8+8( • 12) 5.2+17( • 11) 0.06(13) 0.35(• X J. 0.41( .ir 0.76( 11) 02 4.4+7( .19 7.5+161 .23) 0.02 .20 0.06 17 0.12i 18i;i7) 03 6.5+8( .18 1.2+181 .10) 0.15 .19 0.381.09• i) 1.58i .07; o1.96. ( .06) D4 2.7+9( .15) 4.3+18 .08 0.72(.16 1.02 .07 5.80 .05 6.82 .05 D5 2.7+9( .15) 3.6+18 .12 54(.16 0.98 .09 5.12 .07 6.10 .06) 06 8.9+7( .22) 1.6+17 .22 0c.03(.23) 0.13 .22 1.26 .14 1.39 .12 .07 1O9- .05 lTTT .05

#Read as 4.2x10* W.P. = Winding Pack * Relative Error C.C. = Coil Case results reported herein include the safety factors.. Because of the thick outboard blanket and shield, only a few particles reached the TF coils. The 1 D neutronics analysis showed that the peak outboard magnet radiation effects are about 5 orders of magnitude less than those in the inboard region with negligible contribution (<0.1%) to the total magnet heating. Table VIII.2-8 indicates that the peak end of life insulator dose ana fast neutron fluence occur at the midplane in the inboard region. These values of 4.2 x 10 rads and 6.1 x 10 n/cm are below the design limits of 5x10 rads and 10 n/cm . The largest magnet radiation effects in the divertor region are in parts of the coil behind the outer end of the divertor plates (segments D4 and D5). The damage values in the lower half are considered to be the same as that in the upper half for all regions except for segments D3, D4, D5, and D6 where results of the 3-D calculation for the lower divertor region are used. The lower end of the blanket and shield protrudes farther into the plasma chamber than does the upper end resulting in lower neutron wall loading at the outer end of the lower divertor plate. However,

169 slightly higher magnet damage is obtained in the lower divertor region due to the presence of the vacuum pumping ducts. The end-of-life helium production at the inner surface of the vacuum vessel was also determined for the different poloidal regions. The peak in the inboard region at the midplane is 1.2 appm and the peak in the divertor region is 11 appm. This means that the helium production limit of O.I appm required for rewelding is exceeded everywhere except behind the middle part of the divertor plate and in the outboard region. The total nuclear heating n the 16 TF coils is listed in Table VIII.2-8 for the Technology Phase. While the coil case contribution is less than that of the winding pack in the inboard region, magnet nuclear heating in the divertor region is dominated by the thick coil case. Combining the results for the upper and lower halves of the coils, the total nuclear heating in the 16 TF coils is 48 kW in the Technology Phase and 59 kW in the Physics Phase. It should be mentioned that the recent divertor design calls for fairly thick divertor plates. In the Physics and Technology Phases, the divertor plate/structure is 5.5 and 11 cm thick, respectively. This design helps reduce the damage in the divertor region considerably and, therefore, B^C/Pb can be employed in the back layer instead of W. In segments D1-D5, factors of 1.2-1.7 reduction in damage are anticipated in both phases of operation. In this case, the heating in the divertor region is estimated to be ~38 and 20 kW in the Physics and Technology Phases, respectively. Including the inb.jard heating, the total heat load to the TF magnets is expected to be 55kW in the Physics Phase and 35 kW in the Technology Phase.

V1II.2.4 Conclusions

From the numbers reported and discussed above, the conclusion must be drawn that improvements to the Reference Design are advisable which cut down both the total nuclear heating, in particular during the Physics Phase, and some of the fluence dependent radiation loads. Some possibilities have occasionally been proposed in the course of the Conceptual Design Phase, but not accepted as Reference. The use of borated steel instead of low-boron steel, the use of borated instead of "pure" water are typical examples. The consequences of these choices may, however, interfere with other engineering requirements. The detailed results reported here suggest that especially for the inboard leg and for the divertor regions better solutions have to be found. On the inboard leg, the higher numbers are related to configurations which include blankets with a significant void volume fraction. Improvements of the respective blanket design concepts are possible which, together with the use of tungsten instead of lead/boron-carbide as special shield material layer in the region of the equatorial plane, will lead to an acceptable situation. In the divertor region the situation is not so clear. With the present geometrical features only an extensive use of tungsten and a redesign of the Physics Phase divertor plate configuration seem to solve the problem. A more straight forward solution would, however, be a slight increase of the vertical extension of the TFC which would help to avoid the application of unconventional materials.

170 VIII.3. DOSE AFTER SHUTDOWN

Several analyses were performed to calculate the dose equivalent outside the outboard shield, and around the neutral beam injector (NEI) duct.

VIII.3.1 Dose Equivalent from Bulk Shield

The dose equivalent after the full operations of the two ITER phases was calculated at different times after shutdown for the layered ceramic blanket design. The neutron flux calculated by the ONEDANT transport code was used by the radioactivity code RACC to generate the point-wise 21 groups decay gamma sources at the different times after shutdown. The gamma source, at

TABLE VIII 3-1 OUTBOARD GEOMETRICAL MODEL

Zone Thickness Zone material and composition cm Vol%

FW 5 C, steel, H2 Blanket 50 6% steel, 37% lead-lithium eutectic 11% water, 46 void Shield 64 79% steel, 21% water Vacuum vessel 54 63% steel, 37% water Cryostat 70 5% steel, 95% air Concrete 50 ordinary concrete, 2.3 g/cm

each time step after shutdown, is then used in the ONEDANT code to calculate the gamma flux ajud the dose equivalent. A 7 cm thick stainless steel cryostat and a 2 m thick steel reinforced concrete located at radii of 13 and 15.7 m, respectively, are included in the geometrical model. The results show that a total thickness of 175 cm for the first wall, blanket, shield, vacuum vessel, and cryostat is required to achieve a dose equivalent value of 0.5 mrem/h one day after shutdown outside the cryostat. Another detailed analysis was performed to study the effect of the different operating scenarios on the dose equivalent after shutdown. The calculations are based on a one-dimensional toroidal model where the specifications for the outboard section are shown in Table VIII.3-1. The transport calculations were performed with the ANISN code with P^Sg approximation. The activation analyses were carried out with SAM code in the pulse operating mode. A sample of the dose equivalent results is shown in Fig. VIII.3-1 as a function of the time after shutdown and the cobalt concentration in the steel. For the nominal cobalt concentration of 0.02 weight%, the dose equivalent is 0.4 mrem/h one day after shutdown with a total thickness of 177 cm after 1 MWa/m2 and 0.1 availability factor. The dose equivalent at the end of the Physics Phase is about a factor of ten less than after the Technology Phase.

171 T i i | i i i r i Curve 1 - 1700 pulses, Co-0.02 % 2-17000 pulses. Co-0.005% 3-17000 pulses. Co-0.02 % 4-17000 pulses.Co - 0.20 %

6h12h1d 1w 1m 6m 1y 2y TIME AFTER SHUTDOWN FIG. VIII.3-1 Dose Equivalent Outside Cryostat

VHI.3.2. Dose Equivalent from NBI Duct Shield

Two different sets of analyses were performed to calculate the dose equivalent along the outer surface of the NBI duct shield. The first analysis is based on a two dimensional RZ model where the two dimensional discrete ordinates code DOT3.5 was used for the transport calculations. The calculations used a P-i approximation for the scattering cross sections and a biased forward angular quadrature set with 100 divisions. The FUSION-J3[6] data library based on JENDL-3[7] was used for the calculations. The NBI duct shield has a steel type shield (80% steel and 20% water). A thin steel cryostat (ITER reference design) without a shielding function and a thick concrete cryostat with steel clad were used in the calculations. The neutron source has a uniform distribution and it is normalized to 1 MW/m2 at the first wall. The dose calculation was performed for one full power year. The second analysis calculates the angular distribution of the 14 MeV neutron current on the inside surface of the NBI duct. This surface source is used in one dimensional calculation with P-jS^ approximation. The large number of the quadrature set is required to describe the angular distribution of the neutron source. The activation code SAM was used to calculate the decay gamma source for the ANISN dose calculation with P3S0 approximation. The calculations were performed for two weeks of operation in the technology phase.

172 The results from the two sets of analyses show that a dose equivalent in the range of 0.5-2.5 mrem/h one day after shutdown requires a steel shield thickness of 110-100 cm for the NBI duct outside the cryostat. In case of thin steel cryostat, this duct shield thickness has to increase gradually to 125-115 cm at the back of the bulk shield. For a thick concrete cryostat (> 150cm), the required NBI steel duct shield thickness is 50 cm for a dose of 2.5 mrem/h one day after shutdown outside the cryostat. Both analyses did not account for neutron leakage between the different components. For example the neutron leakage between the toroidal field coils and the shield which has to be considered to assess the impact on the dose value.

VTII.4 RADIATION STREAMING ANALYSIS

The bulk magnet and biological shield of ITER are no continuous structures. They are locally interrupted by discontinuities and a number of penetrations, e.g. for heating a:id diagnostics purposes. The most important ones have been subjected to specific analyses in order to ensure that adequate local shielding be provided to avoid hot spots.

VIII.4.1 Assembly Gaps

The 20 mm wide radial assembly gaps in between adjacent blanket segments have been properly taken into account in 3D calculations. For application to the results from ID calculations a correction factor of 1.7 had been proposed on the basis of previous work, and was included in the respective safety factors (see Sect. VIIL1.). By means of 2D calculations with the codes DOT.3 and TWODANT it was attempted to verify this number. A typical geometrical model is shown in Fig.Vffl.4-1. Peaking factors of •> 2 and » 4 were found in the total and the 14 MeV neutron flux at the root of the gap as compared with a location behind the bulk of a blanket segment. The peaking tends to decrease and broaden with increasing depth into the vacuum vessel and the TFC. In addition to the gap streaming effect, the mechanical design of the inboard segment, and in particular of the vacuum vessel, calls for a wide variation in material arrangement within a single module. This toroidal variation in composition affects the damage level at the magnet. Furthermore, the thinning in the coil case of the inner legs creates hot spots at the corners of the winding pack. The variation in coil case thickness and the assembly gaps were included in the models. The calculations were performed in x-y and r-z geometries. The peaking in damage is given in Table VIII.4-1 at the middle and corner of the winding pack. The peaking factor is defined as the ratio of the 2-D to the 1-D values for the damage at the magnet and differs with the response function. The enhancement in damage at the corners of the winding pack stems from the relatively thin coil case (2.6 cm compared to 7 cm at the middle). A possible solution to alleviate this problem is to design a uniformly thick coil case for the inner legs. It is concluded that the increase in damage due

173 ©Amort (graph i te) (3)Blanket vessel

SS:H2O=0. 778:0. 222 (J/Breeder zone "5 - 0 3e:L;?.0:SS:H20 = 0. T:U:0. 1131 :0.0377:0.0904 ©Blanker, cooing header

—1 _/ SS:H8O=O.625:0. 375 ©Shield I 53 3 00 03 SS:R2O-0. 3:0.2 ©Shield tl 0.75 «.- 250 ®B4C=1.0 15 ® ©He-can SS=1. 0 ©Insulator (^finding pack 200 ©Extra Be-can SS»L0

I 1 I I II 1 1 lUllpllUlUHUlUIUIIIIIII II i I I II I 1 R 0 20 40 60 80 100 distance from TFC left side (cm)

FIG VUI.4-1 Horizontal Inboard Cross Section of the Geometrical Model Used in the 2D Analysis to the assembly gaps is ~ 1.7 for local responses, and thus confirms the previous assumption for the correction factor.

VIII.4.2. Vacuum Pumping Ducts

Detailed three-dimensional neutronics calculations have been performed to determine the radiation effects in the parts of the TF coils behind the lower divertor region and adjacent to the vacuum pumping ducts. The Monte Carlo code MCNP has been used in the calculations with the ENDF/B-V cross section data evaluation. Because of symmetry, only 1/64 of the reactor was modeled with surrounding reflecting boundaries. Detailed configurations of the blanket, shield, vacuum vessel, coil case, and winding pack in the divertor region were included in the model. In addition, the divertor plates, vacuum pumping ducts and divertor

174 TABLE VHI.4-1 PEAKING FACTORS IN THE WINDING PACKS OF THE INNER LEGS Middle of Corner of Winding Pack Winding Pack [5] [4] [3] [4]

Peak dose to insulator L64 L71 L5 Z09

Peak nuclear heating in winding pack 1.6 1.85 1.6 2.24

Peak fast neutron fluence 1.75 1.92

Peak dpa in Cu stabilizer 1.63 1.25 1.99 coolant tubes penetrating between TF coils were modeled in detail. At the side of the vacuum pumping duct the shield, vacuum vessel, and coil case are 30, 10 and 11 cm thick, respectively. 316 SS/HjO (at 20 v/o H2O) is used in the penetration shield. Seventy thousand source particles have been sampled in the MCNP calculation. The neutron source v/as sampled from the actual source profile in the D-shaped plasma. Several variance reduction techniques were utilized to improve the accuracy of the results. Surface flux tallies were used to determine the radiation effects at the front and side surfaces of the TF coils. Neutrons crossing surface detctors at the entrance and exit of the vacuum pumping ducts were tallied to quantify neutrons streaming into the divertor vacuum pumping ducts. The results indicate that 2.82 x 10 (±3%) neutions stream into each vacuum pumping duct per DT fusion. The number in parentheses corresponds to the statistical uncertainty in the calculation. 10.5 % of these neutrons are uncollided source neutons streaming directly iajo the duct. The leakage out of the vacuum pumping duct amounts to 7.75 x 10 Ci.4%) neutrons per DT fusion with a very soft spectrum. The end of life insulator dose and fast neutron fluence (E>0.1 MeV) averaged over the side surface of the winding pack are 6 x 10 C±26%) rad and 1.3 x 1017 (_+27%) n/cm at the end of 3.8 FPY's of operation.

VIII.4.3 Divertor Coolant Tube Penetrations

The ;j:bes for the divertor plate cooling penetrate through the shield and the vacuum vessel in vertical direction at the top and the bottom of the machine. The water in the tubes and the annular gaps around them provide a streaming path for thi radiation. Although the location of this penetration is in the middle between two adjacent toroidal field coils, the shield thickness is only - 20 cm and could, therefore, give rise to incrased radiation levels at the TFC. The problem was modelled in R,Z geometry far a 2D DOT analysis. This model is, of course, only a rough approximation of the real situation, it

175 should, however, yield results on the conservative side. It was found that the problem is governed by the empty gap rather than the water. Nevertheless, the values for all local responses calculated are within the limits applying the same safety factors to them as for the ID calculations. Additional results were obtained in the course of 3 D calculations with MCNP, as they are reported in Section VIII.4.2. It was found that the contribution to the tulal nuclear heating in the 16 TF coils is about 0.7 and 0.5 (+.15%) kW in the Physics and the Technology Phases, respectively.

VIII.4.4 Neutral Beam Injection Ports Three out of the 16 horizontal ports on the outboard equatorial plane are foreseen for neutral beam injection. They have a width of 80 cm and a hfight of 340 cm and penetrate obliquely through the shield (see Fig. VI. 1-6) thus allowing tangential injection into the plasma. In contrast to other ports, these ports have to be empty and enough shielding thickness has to be provided at the lateral sides of the duct to protect both the magnets and the environment from excessive radiation. Streaming calculations have been done [3,5,8] using both 2D (DOT) and 3D (MCNP) analysis to establish the neutron and gamma flux levels along tiie duct up to the far end which is 28 m from the duct entrance. As can be seen from Fig. VIII.4-2 the agreement between the results from the two methods is fairly good. The flux decrease over the entire length is more than 3 orders of magnitude.

> toral neulran flux tjy M[NP O gamma - ray flux Sy MCNP

- o

/totol neulron IIJ« DO 5 10 ce

gamma-ray flux by DOT 3S 1 10 o en

10' 0 3 6 9 12 15 18 21 2't 27 DISTANCE FROM NBI DUCT ENTRANCE Z[m] FIG. VIII.4-2 Total Neutron and Gamma Fluxes Along the Neutral Beam Duct as Calculated with DOT3.5 and MCNP.

176 Earlier 2D calculations [8] had uncovered a particularly high radiation load on the TFC where the duct passes the rear part of the coil. A design revision lead to a redistribution of the shield thickness for the two neighbouring coils (350 and 420 mm) which now assures acceptable local conditions. Further calculations have been done to specify the additional biological shielding on the duct walls outside the cryostat to meet the requirements for personnel access 24 hours after shutdown. In this context the transport of activation gammas from both the plasma chamber and the duct walls have been considered. The results have been discussed in Section VIII.3.

VIII.4.5 Diagnostic Channels

In the absence of well defined geometrical features of diagnostic channels to be used in ITER a parametric study was performed [9] vith the 2D DOT code in R,Z geometry, and a line-of-sigbt method. To improve the accuracy of the SQ calculations an asymmetric set of angular quadratures with a total of 166 directions was used. Channels of 5 to 20 cm in diameter and a length of 1.5 m have been considered which are surrounded by a water cooled stainless steel shielding structure. For these dimensions, the fast flux at the channel exit is higher by 4 to 6 orders of magnitude than behind the bulk shield. It decays to the background level within a radial distance from about 30 to 100 cm from the channel axis. There are indications that the exit spectrum is substantially harder than that on the channel entrance, because the high energy part experiences much lower scattering on the channel walls than the low energy one. For longer channels which even penetn•«. through a subsequent 2m thick concrete cryostat it was found that 2D calculations yield inadequate results. They underestimate the fast flux at the channel exit b> about 3 orders of magnitude as compared to the Une-of-sight method, while for the shorter channel the diviations were a factor of 2 for the 5 cm channel and below a factor of 10 for the 20 cm wide channel. The conclusions from these studies are, that the diagnostics equipment located inside the channel have to withstand an appreciable flux and hard spectrum, and that significant shielding is required outside to protect the environment from the leaking radiation.

VIII.5 SUMMARY OF BENCHMARK CALCULATIONS

Several shielding problems have been defined in detail andthe calculated results by the different participants wereanalyzed. The main function of this benchmark exercise is to establish a common methodology for the ITER shielding analyses. The other function is to compare the results from differenttransport codes and nuclear data files for ITER shielding problems. The main conclusions are summarized as follows:

177 - The maximum differences in the total nuclear heating values calculated by the different participants with common data base and transport code are 2 and 5% in the FW/Blanket/ shield and the inboard TF coils, respectively.

- The maximum differences in the nuclear response distributions calculated by the different participants with common data base and transport code are 3 and 9% in FW/Blanket/ shield and the inboard TF coils, respectively.

- The maximum differences in the spatial distribution and integrated nuclear responses calculated by ANSIN code (IFLU = 4) and ONEDANT are less than 0.5?

- The maximum difference observed in the nuclear response distributions obtained from ANISN code with IFLU= 0 and 4 is less than 3%.

- The maximum difference observed in the nuclear response distributions obtained from ANISN code with IFLU= 0 and 5 is less than 16%.

- The maximum differences in the integral and spatial distribution of nuclear responses due to the change in the mesh size from 2 to 10 mm are 2 and 10%, respectively.

- The neutron flux values in the inboard TF coils calculated with ENDF/B-IV are 20 to 25% less than the ENDF/B-V values.

- The nuclear heating in the carbon tile calculated with ENDF/B-IV is 35-40% grater than the ENDF/B-V value.

- The total nuclear heating in the inboard TF coils calculated with ENDF/B-IV is 20-25% lower than the ENDF/B-V value.

- The total nuclear heating in the inboard TF coils calculated with JENDL3 is 35 % lower than the ENDF/B-V value.

REFERENCES

[1] ITER Expert Meeting on Shielding Experiments and Analysis, Summary Report, ITER-IL-BL-5-0-5. [2] W. Daenner, F. Gervaise: ID Shielding Analysis on the basis of the EC Reference Ceramic Blanket; EC Contribution to the ITER Shielding Specialists' Meeting, Summer 1990. ITER-IL-BL-4-0-18 [3] K. Maki: Japanese Contribution to Shielding Neutronics, Japanese Contribution to the ITER Shielding Specialists' Meeting, Summer 1990, ITER-IL-BL-4-0-14.

178 [4] L. El-Guebaly, M. Sawan, E. Mogahed, I. Sviatnslawsky: US Contribution to the Shielding Session of ITER, ITER Shielding Speciahsts' Meeting, Summer 1990. [5] S. Zimin, V. Aborin, A. Svetchkopal, D. Markovskij: Bulk Shield Analysis Summary Report, USSR Contribution to the ITER Shielding Specialists' Meeting, Summer 1990. [6] Maki, K. et al., "Fusion Nuclear Group Constant Set FUSION-J3 and Shielding Properties for Fusion Experimetnal Reactor FUSION-J3." 1990 Annual Meeting of the Atomic Energy Society of Japan F8 (April 1990). [7] Shibata K., et al., "Japanese Evaluated Nuclear Data Library, Version-3 JENDL-3, "JAERI1319 (June 1990). [8] K. Maki, Japanese Contribution to Shielding Neutronic Design, Japanese Contribution to the ITER Shielding Specialists' Meeting, Winter 1990, ITER-IL-BL-4-0-1. [9] S.A. Zimin: An analysis of radiation streaming through ITER and ITER/OTR diagnostic channels and channels with water, stainless steels and entectic Li^Pbo-j, USSR Contribution to the ITER Shielding Specialists' Meeting, Summer 1989.

179 PARTB

ITER MATERIALS EVALUATION AND DATA BASE CONTRIBUTORS PART B (MATERIAL DATA BASE)

M. AKIBA, M. ARAKI, T. ANDO, D. BALDWIN, R. BASTASZ, R. BEHRISCH, U. BERGENLID, W. BOGAERTS, H. BOLT, M. BROSSA, A. BRUGGEMAN, M. BUDD, S. CASADIO, R. CAUSEY, F. CLINARD, S.COHEN, R.CONRAD, M. DEVRIES, W. DIENST, D. DUQUETTE, J. ELLIOT, M. ETO, A. FISCHER, T. FLAMENT, F. GARNER, D. GELLES, P. GIERSZEWSKI, Y. GOHAR, A.A. GRIGORIAN, A. HAASZ, R. HEIDINGER, A. HISHINUMA, E. HODGSON, G. HOLLENBERG, A. HULL, A. IBARRA, V.A. IGNATOV, H. ISE, C. JOHNSON, R. JONES, V.A. AZAKOV, T. KASSNER, D. KIRK, A. KOHYAMA, A. KRAUSS, T. KURODA, H. KWAST, T. LECHTENBERG, V. LEVY, G. LONGHURST, B. LOOMIS, G. MARTIN, D. MAZEY, L. MILLER, V. Z. NEPOMNYASCHCHIY, T. OKU, G. PELLS, Y.G.PROKOFIEV, A. S. POKORSKY, G.L. SAKSAGANSKY, V. ZEMLIANKIN, V.G. MARKOV, H. NICKEL, B.V. PETROY, J. ROTH, N. ROUX, A. ROWCLIFFE, G. SANTARINI, P. SCHILLER, M. SEKI, N. SEK1MURA, G.A. SERNYAEV, B. SINGH, D. SLAGLE, M. SMITH, YU.N. SOKURSKY, S.I. SOROKIN, J. STEPHENS, A. TAVASSOLI, V.N. TEBUS, YU.M. TRAPEZNIKOV, S.I. TURCHIN, V.P. KARKLINA, H. ULLMAIER, J. VAN DER LAAN, W. VANDERMEULEN, B. VAN DER SCHAAF, M. VICTORIA, E. VIETZKE, N.K. VINOGRADOVA, H. WATANABE, K. WILSON, H. WOLLENBERGER, A.P. ZAKHAROV, H. WERLE, J. WHITLEY, S.ZINKLE, I. INTRODUCTION

TABLE 1-1. CANDIDATE MATERIALS

Component Candidate Materials

Structural Materials First Wall/Blanket Type 316 austenitic steel (SA) Type 3/6 austenitic steel (CW) MN stabilized austenitic steel

Divertor Copper Alloys Molybdenium Alloys Niobium Alloys

Tritium Breeding Materials Ceramic Breeder Li2° Metal Breeder Materials 83Pb-17Li enteric Alloy Aqueous Lithium Salt Breeder LiOH.LLNiOo (solutions) Neutron Multipliers Beryllium Lead

Plasma Facing Materials Carbon based Materials Nuclear grade carbons pyrolytic graphite carbon fiber composites High-Z coatings Tungsten Low-Z coatings Beryllium

Ceramic Insulators Oxides AUO^, MgA^O^BeO Nitrides

183 needs and to recommend the required materials R&D needed to provide an adequate data base for the construction of ITER. The focus of this effort has been on the blanket, first wall, and plasma facing materials although a similar approach should be conducted for other components. Four types of materials have been included in this data base assessment: (1) structural materials for the first wall, blanket, and divertor, (2)Tritium breeding materials for the blanket, (3) plasma facing materials for the divertor and first wall armor, and (4) electrical insulators for use in the blanket and divertor. Based on guidance from the design activity the design requirements for the various materials applications were defined and candidate materials for each application were selected. Candidate materials for each application are listed in Table 1-1. A database for the candidate materials was compiled and presented at two Specialists Meetings GJ Materials Data base held at the ITER site in Garching in June 1988 and February 1990.

I.I. Structural materials

Type 316 austenitic steel in the solution annealed condition has been selected as the reference material for the first wall/blanket structure primarily on the basis of a more extensive data base and good fabricability. An important influence on this choice is the specification of low-temperature ( < 100°C) water as the first wall/blanket coolant. Therefore, the operating temperatures will generally be in the range of 50 - 200°C. A goal of the design is for the first wall blanket to last the entire reactor life of up to 3 MWa/m . The baseline physical and mechanical properties of Type 316 austenitic steel are well established. Key issues and concerns relate to the relatively low strength in the annealed condition, radiation induced and hydrogen embrittlement at low temperatures, and possible sensitivity to aqueous stress corrosion under the proposed operating conditions where significant radiolysis may occur. Of particular importance are these effects on weldrnents and ^raze joints. Type 316 austenitic steel in the cold- worked (cw) condition is the primary alternative. The cold-worked material provides a significant strength advantage but possibly increased concerns related to weldments, embrittlement, and stress corrosion. The manganese stabilized austenitic steel was originally considered as an alternative but it produces additional concerns regarding radioactivity afterheat, corrosion and embrittlement. Copper, Niobium and Molybdenum alloys have been selected as primary candidate materials for the divertor structure. As in the case of the first wall and blanket, low temperature water is selected as the reference coolant for the divertor. Therefore, much of the divertor structure will operate at temperatures below 200°C; however, temperatures of several hundred degrees Celsius will be reached in areas exposed to the peak heat fluxes of 10-15 Mw/m . Since it is anticipated that the divertor will be replaced several times during the reactor life because of high armor erosion rates, a radiation lifetime of approximately 5 dpa is probably acceptable. Because of the high coolant velocities required for heat removal corrosion is a concern.

184 The copper alloys arc favoured because of their high thermal conductivity which contributes to their high heat flux capability. The dispersioned strengthened alloys are currently recommended; however, age hardinable alloys have also been considered. Major issues with copper relate to their relatively high thermal expansion coefficients compared to carbon and tungsten armor, and the difficulty of welding the dispersioned strengthened alloys. Copper does provide a fairly good thermal expansion match for the beryllium armor option. Molybdenum alloys (particularly MoRe alloys) are candidates because of their relatively good thermal stress factor for accommodating high heat loads, high melting temperature which provides safety advantages, and a low thermat expansion coefficient which matches well with carbon and tungsten armor options. The dominant concerns for molybdenum alloys relate to fabrication and welding difficulties, radiation-induced embrittlement at the low operating temperatures, and safety issues associated with oxide volatility, particularly for the Mo-Re alloys. The Mo-5Re or Mo-40Re are the preferred compositions. Niobium alloys, such as Nb-lZr and Nb-V-Zr, provide the same advantages as the molybdenum alloys; however, they are more easily fabricated and welded and appear to be much less susceptible to radiation-induced embrittlement. Major concerns for the niobium alloys relate to the potential for hydrogen embrittlement and tritium permeation. Because of the more extensive data base Nb-lZr is considered as the reference niobium alloy; however, a Nb-V- Zr alloy such as Cb753 may provide superior properties.

1.2. Tritium breeding materials

This class of materials includes the neutron multiplier materials in addition to the lithium bearing tritium breeding materials such as the ceramic- breeders, Pb-Li alloys, and aqueous lithium salts. The primary considerations in the selection of the breeder materials include: tritium breeding capability; ease and reliability of tritium recovery; thermal transport properties; and thermal, chemical, and irradiation stability. Since the blanket is expected to last the entire reactor lifetime, these material will receive neutron fluences corresponding to a neutron wall load of up to 3 MWa/m . The operating temperature range will be set primarily by the tritium recovery scenario. Candidate ceramic breeder materials include LL^O, LiAlCL and LLZrOa and LLSiO^. The data base covers the baseline physical properties, baseline mechanical properties, chemical stability/compatibility, radiation effects, and tritium solubility/transport. Features of the ceramic breeder relate to relative safety associated with their chemical stability, potential for low tritium inventory, relatively good data base, and potential reactor relevance. Key issues are associated with reliable tritium recovery-transport- processing, thermal transport, and compatibility/mass transport, particularly for LLO. The 83Pb-17Li entectic alloy is a candidate breeder material which inherently provides neutron multiplication by the lead. This alloy melts at 235°C. The current design proposal provides for normal operation in the solid phase with subsequent melting during off-periods for tritium recovery. Key features of

185 the Pb-Li alloy include low tritium inventory and potential reactor relevance. Major concerns relate to tritium containment and recovery, compatibility with the structure and mechanical problems associated with the melting/solidification process. Considerable data base on the alloy has been developed in the last few years. The aqueous lithium salts ha*j& been proposed as a candidate breeder material. Solutions of LiOH and L1NO3 nave '3een evaluated in more detail. The primary feature of these materials are the flexibility associated with a liquid breeder, such as the ability to remove or replace the breeder easily. Primary concerns relate to corrosion/compatibility which is partly associated with radiolysis, safety in particular for the LiNOo, tritium recovery issues, and relatively low breeding performance. The data base for these materials is generally quke limited. Beryllium and Lead are considered as neutron multipliers to enhance the tritium breeding performance. Beryllium is the most effective neutron multiplier. Key features associated with the use of beryllium include its high thermal conductivity, low density, and low activation properties - both short term and long-term. Major concerns related to the use of beryllium include its chemical toxicity, tritium retention characteristics, radiation-induced swelling, and cost. A considerable data base exists for beryllium in various forms. Lead is currently proposed as a neutron multiplier only in the Pb-Li alloy and is discussed above. Key features are its low cost and ease of fabrication. Major concerns relate to its high density, low melting temperature and activation products. An extensive data base exist for lead.

1.3. Plasma Facing Materials

Viz leading candidates for use as plasma facing materials are carbon based materials including carbon-fiber-coniposites (CFC), tungsten and beryllium. The carbons have received the most attention because of their extensive use in current tokamaks. However, their use in the Technology phase is questioned because of radiation damage limitations. Nuclear grade-carbons, pyrolytic carbon and CFC's have all been considered. The 3-dimensional CFC's are currently the preferred form of carbon. Key features include their low-Z, high thermal conductivity, high vaporization (no melting) temperature, and low thermal expansion. Major concerns for the use of the carbons relate to tritium inventroy, sputtering erosion including chemical sputtering, limitations on component size for PG and CFC's in particular, radiation effects, chemical reactivity with air and water at high temperature, and joining difficulties. A considerable data base exists for the carbons; however, the properties vary widely depending on preparation methods. Tungsten is of interest primarily because of its very low sputtering rate. The major concern relates to the acceptability of high-Z materials in the plasma chamber. Other key features include its high melting temperatures, good thermal conductivity, low expansion coefficients, low tritium solubility and the ability to plasma spray. Other concerns relate to its low ductility, activation and nuclear

186 afterheat, and chemical reactivity with air at high temperature. The existing data base for tungsten is fairly extensive. Beryllium has been proposed «~j a plasma facing material primarily on the basis of its very low-Z. Other key features include high thermal conductivity, low tritium solubility, and low activation properties. Primary concerns relate to its high sputtering rate, relatively low melting temperature, radiation induced swelling and chemical toxicity.

1.4. Electrical insulators

Electrical insulators are required in both the blanket and the divertor to reduce disruption induced electromagnetic loads to acceptable levels. Ceramic insulators are proposed because of their superior radiation damage resistance compared to organic insulators. Even so, the radiation effects are a critical issue for the ceramics insulators. The oxides, e.g., AL^O, and MgAi^, are generally proposed as candidates. The oxides exhibit excellent insulating properties, are highly stable and readily available. However, radiation effects, particularly swelling and radiation induced conductivity are major concerns.

187 fl. STRUCTURAL MATERIALS

The data base for austenitic steel and copper-, niobium- and molybdenum-alloys are presented in this section. Information on braze materials is also presented. Baseline properties of the candidate structural alloys are summarized in Table II.0-1 and II.0-2 and Figs. II.0-1 to II.0-7.

II.1. Type 316 Austenitic stainless steel: FW and blanket structure

11.1.1. Selected materials

Solution-annealed (SA) 316 type austenitic steels have been selected as reference alloys on the basis of the large experience gained in the development and/or use of this type of steel for components or applications in the nuclear industry. Data are presented on both low-carbon Type 316 steel and a titantium modified composition. Typical compositions are given in Table H.l-1. Cold-worked 316 type austenitic stainless steels have been selected as alternative candidate alloys: 20% CW 316 and 25% CW 316Ti. These alloys could provide potential advantages in some applications due to their higher mechanical strength and dimensional stability under neutron irradiation. However then- fabrication flexibility is less than that of solution annealed material since high temperature brazing or welding will tend to anneal the cold-worked structure..

11.1.2. Status of existing data base

II. 1.2.1 Tensile properties

The effect of neutron irradiation on 316 type steel has been largely addressed in the various national R and D programmes for temperatures ranging from 250°C to 550°C and doses mainly up to 10 dpa with some data up to 20, 30 and 50 dpa. Conversely data for range 50 to 100°C are limited. Fig.II.1-1 gives the evolution of the yield strength of SA 316 and 316 Ti measured at the irradiation temperature after irradiation in mixed neutron spectrum reactors (R2, HFR and BR~ in EC; HFIR is the US) or fast breeder reactors (Rapsodie and Ph6nix in EC; Bor 60 in SU). Within the range cf variation investigated, the chemical composition and neutron spectrums appear to have marginal effect on the irradiation strengthening. The effect of temperature is marginal on the saturated value of yield strength. Conversely the saturation dose is strongly temperature dependent: ~ 0.1 dpa at 100°C, up to 10 to 15 dpa at ~ 400°C. Annealed material exhibits high ductability ( > 20% uniform elongation. However, as shown in Fig. II. 1-2, the uniform elongation is severely degraded to < 0.5 % at temperaturer of 200-350 °C after irradiation to fluences of 7-8dpa.

189 TABLE II.O-l

Property Alloy System 316 SS DS Copper TZH Mo-5Re Mo-40Re Nb-lZr Melting Pr, °C 1370-1400 1083 2610 2580-1600 2500-1550 2468 Density@2O°C.g/cm3 7.75-7.96 8.86 10.2 10.49 12.94 8.46 Poissons Ratio,20°C 0.27 0.33 0.33 - - 0.38 DBTT "C-5R NA NA 0 -150 -150 r Rx NA NA 50 -25 -150 -200 Elongation at 20°C total, * 69 (AU5)17, (A120)4 17 26 15 46 Uniform, * 52 (AU5)12 14 15 12 30 at 800°C total, 4 50 (A115)17. (A120)2 12 16 12 21 Uniform, V 12 (AU5)2 1.2 1 1.1 14

TABLE II.0-2 RECOMMEND EQUATIONS FOR USE IN DESIGN

Property Material Thermal Conductivity (w/m*k) »thermal Expansion (10 ) Modules of Elasticity (GP) 316 SS 14.28 + 1.42 x 10"* T 16.32=4.93xl0";)T-1.66xl0"bT;! 195.98-6.62xlO'2T-2.O2xl0"5 US copper 351.10-0.1241-3.llxlO"6T2 16.8=2.09xl0"3 T+2.81xlO"6T2 135.91-6.96xl0"2T-7.4xl0"6

2 5 2 TZH Molybdenum 119.24-4.96xl0' T-2.26xl0" T 4.94+2.72xlO"3T-1.43xlO"6T2 311.84-9.09x10-2 T-4.44xlO-6T Mo-5 Re 111.81-9.57 x 10"3 T 5.57=4.31xl0"4T-l.lxl0'7T2 Mo-40 Re 51.85 - 9.57xlO"3T 5.7+6.6xl(T9T Nb-lZr 43.15-1.41xlO"2T 6.51+1.59xlO"3T 89.25-33.9xlO"3T ^MATERIAL COMPARISON OF THERMAL EXPANSION | PROPERTY COEFFICIENT OF THERMAL EXPANSION _j Temperature (°F> 200 400 600 600 1000

700 800 Tomporaturo ( C)

| MATERIAL COA\PAglSON OF THERMAL EXPANSION | PROPERTY COEFFICIENT Of THERMAL EXPANSION | Temperature ( F) 2C0 400 600 800 1000 1200 MOO

! -o •o o I 331 —- -6 — o «• -fc I c o MB =a= mm •H c •HI — o _i •»• •••» ••— o •• =

—f— q LEGEND • - NIOBIUM-1X Zr. ANNEALED o - MOLYBDENUM-40% R«. S.R. v i •= MOLYBDENUM-5X R«. S.R. O o + = TZM MOLYBDENUM. S.R. O

100 200 300 400 500 600 700 Temperature ( C)

FIG. II.0-1 Coefficient of thermal expansion of ihe canaidate structural alloys

191 Electrical Resistivity, micro-ohms/cm Thermal Conductivity. W/in-K SO 100 150 200 50 100 150 200

P o

O o i ; ) m H

3" a q 3 E." cL IX a 5 (A o O C c £L 6:

0 20 40 60 80 100 Electrical Restivity. micro~ohms/cm Thermal Conductivity. BTU/hr-ft-F MATERIAL COMPARISON OF VARIOUS HEAT CAPACITIES PROPERTY SPECIFIC HEAT CAPACITY Temperature o 200 40CJ 600 80C 1000 1200 M00 o i 1

LEGEND *^ I a - TZM MOLYBDENUM. S.R. o = TYPE 316 SS, ANNEALED A = DS COPPER TYPE AL25 o I- . J/kg- K ) 0 100 + = NIOBIUM \X Zr. ANNEALED

O w MM =d Q.

60 0 =? Capa c ••a 3 O c i £§- MM = T

MB ifi c ea T HI 1 20 0 Spe c

I 100 200 300 400 500 600 700 800 Temperatura

FIG. II.0-4 Specific heat capacity of the candidate structural alloys

MATERIAL COMPARISON OF TENSILE STRENGTHS PROPERTY ULTIMATE TENSILE STRENGTH

Temperature 200 400 600 800 1000 1200 MOO I i 1 ! i i i 1 1 LEGEND 1 • - MOLYBDENUM-41% Re. S.R o o = TZM MOLYBDENUM. S.R. A = MOLYBDENUM-5% Re. S.R. i + - TYPE 316 SS. ANNEALED x - DS COPPER TYPE AL25 c •^ •^, | o - NIOBIUM IX Zr. ANNEALED t/> o —1 „ o •^ — — 2 ••• •— M » — c 3 mam m bi- ••>• i -— 2 « J) 3 E « =*= ^— • § -*> —^ =9•*• -4MM j ^i—1— 1 •—i —t-

FIG. II.0-5 Ultimate tensile strength of the candidate structural alloys

193 MATERIAL COMPARISON OF YIELD STRENGTHS I PROPERTY YIELD STOfNGTH Temperature ( F) 200 400 600 eco 1000 1200 1400

LEGEND MOLYBDENUM-41% Re. S.R TZM MOLYBDENUM. S.R. MOLYBDENUM 5% Re. S.R. DS COPPER TYPE AL25 TYPE 316 SS. ANNEALED NIOBIUM-lXZr. ANNEALED

100 200 300 400 500 600 700 800 Temperature (°O

FJG. II.0-6 Yield strength of the candidate structural alloys

MATERIAL COMPARISON OF MODULUS OF ELASTICITY PROPERTY STATIC YOUNG S MODULUS Temperature (°F) 200 -400 600 800 1000 1200 MOO

LEGEND TZM MOLYBDENUM. S.R. TYPE 316 SS. ANNEALED DS COPPER TYPE AL25 NIOBIUM 1% Zr. ANNEALED

200 300 400 500 600 700 800 Temperature ( C)

FIG. II.0-7 Static group's modulus of the candidate structural alloys TABLE II.l-l: CHEMICAL COMPOSITIONS OF REFERENCE SOLUTION-ANNEALED 316-TYPE AUSTENITIC STEELS Materials Cr Mn NI HO Coj

0.2 0.4 1.4 21 0 5 10 10 20 30 FLUENCE.1O2On/cm2(E>1MeV) DOSE (dpa) DCSE (dpa)

Fig. II. 1-1: Yield strength of Sa 316 and 316 Ti after irradiation at < 100, ___ 250

and ~ 400°C (Tt st=T-r) in mixed neutron spectrum reactor (R2, HFR and BR2 in EC; HFIR in the US) or fast breeders (Rapsodie and Phenix in EC; Bor 60 in SU) 60

UNIRRADIATED J316SA UNIRRADIATED JPCASA IRRADIATED BOTH ALLOYS

10 - 3 ••'v i'*r-t-^ • ' -'• .j i I i i 100 200 300 400 500 600 700 TEMPERATURE (°C) Fig. II. 1-2: Ductility of solution annealed austenitic steels irradiated to 7-8 dpa with fusion relevant He/dpa ratio of ~ 10 appm He/dpa.

UNIRRADIATED J316 20% CW 25 UNIRRADiATED US316 20% CW UNIRRADIATED JCPA 15% CW 20 UNIRRADIATED USPCA 25% CW

rio N UNIRRADIATED JCPA 25% CW o IRRADIATED ALL ALLOYS 15

1 ELO r I

FORf v 10

100 200 300 400 500 600 700 TEMPERATURE, C Fig. II. 1-3: Ductility of cold-worked austenitic steels irradiated to 7-8 dpa with fusion relevant He/dpa ratios.

The cold-worked material exhibits a much lower ductability than the annealed material in the unirradiated condition: however, after irradiation the uniform elongation of the cold-worked material is higher thar that of the annealed materW in the range 200-350oC (see Fig. n.1-3).

197 II. 1.2.2. Fracture toughness

The 300 series stainless steels exhibit good fracture toughness in the SA condition. After irradiation at 250 and 400°C up to 10, 15 and 50 dpa in fast breeder reactors or in test reactors, they show a decrease in Jrp by a factor in the range from 3 to 10. The fracture toughness expressed in term of Kr^ is as low as ~ 100 MPa m and 50 to 70 MPa m for material irradiated in solution-annealed and welded condition respectively. No data are available after irradiation at temperatures down to ~ 50°C. The rather good uniform and total elongation of SA 316 should nevertheless result in reasonable fracture toughness after irradiation. The fracture toughness of CW 316-type after irradiation up to ~ 10 and 75 dpa at ~ 400°C exhibit similar values as those of irradiated SA 316. After irradiation at ~ 50°C, a rather low fracture toughness is expected for CW 316 steels in accordance with their low ductility.

II. 1.2.3. Low cycle fatigue

Fig. II.1-4 gives the total strain range Aet versus the number of cycles to failure measured at ~ 400°C for SA316L, CW316 and CW316 Ti before and after irradiation at ~ 400°C up to doses in the range 10 to 50 dpa.

I i T

- 10°

ZZL C£ 5

I/O TEMPERATURE : TtK, = Tirr. = 430 °C MATERIAL: NON - IRRAO. IRRAO OOSEIdpa) SA316L-EC O • 10 CW316 -US 10 CW316 Ti-J 30 50 10" I I I I I I I I I 1111 103 104 10s 10' NUMBER OF CYCLES TO FAILURE Fig. 11.1-4: Total strain range versus the number of cycles to failure for SA 316L (EC) and CW 316 (US) and 316 Ti (Japan) before and after

irradiaiton at 43O°C (T =Tt .,) If icSl 198 Before irradiation the CW materials have a better fatigue lifetime than SA ones by about an order of magnitude at a given Ae*. However, the lifetimes of irradiated materials are very similar with a modest irradiation induced decrease for the SA 316L by a factor 2. There is either no significant or a small effect of hold time, as expected for test temperatures below the thermal creep range. The irradiated SA 316 shows no cyclic hardening or softening for Aet < 0.5% For Ae. > 0.5 % cyclic softening is observed. The procedure of post-irradiation testing is nevertheless a strong simplification as compared to reaJ FW conditions and its relevance is therefore being tested by additional in-pile fatigue testing performed within the EC program.

//. 1.2.4 Creep and rupture

Irradiation creep The data on irradiation creep 316 type steels above 250°C are quite abundant. The steady state irradiation creep rate of 316 type steel at temperatures above 250°C and doses below the onset of the steady state swelling rate regime is: (i) weakly dependent on temperature (ii) rather linear with dose rate and stress upto 150 MPa for CW 316 (iii) rather insensitive to the initial metallurgical condition and chemical composition. After the onset of the steady state swelling rate, the irradiation creep rate exhibits an enhancement by up to an order of magnitude. At lower irradiation temperature down to 60°C, 316Ti in CW and SA condition exhibits clear enhancement of the irradiation creep rate by a factor 5. CW and SA 316 show however significantly smaller enhancement. The irradiation creep rate is not significantly enhanced in experiments where the flux of light ions is cycled with and without temperature changes, the temperature range being from 160 to 400°C. Consequently data on stationary irradiation creep rates should apply to ITER pulsed operating conditions at least for temperature above 160°C. The existing evidence suggests that plastic deformattion controlled by irradiation creep is not accompanied by any cracking or ductility exhaustion phenomena. Therefore the interaction between irradiation creep and fatigue is expected to decrease the lifetime only slightly.

Thermal creep The rupture life and ductility of SA 316-type steel above 500°C are significantly reduced during or following irradiation, mainly due to He- embrittlement and premature intergranular fracture. This effect produces strong creep fatigue interaction. This He-embrittlement occurs even at very low He content. SA EC-316L steel containing 10 appm He produced by B(n,a) Li under thermal neutron irradiation at 50°C exhibits only 10% of the initial rupture life when creep tested at 625°C under 190 MPa.

199 II. 1.2.5 Swelling

The swelling behaviour of 316-type steels is characterized by an incubation dose followed by the steady state swelling regime with swelling rate of 0.5 to 1% per dpa. The incubation dose is strongly dependent on metallurgical condition and chemical composition conversely to the steady state swelling rate.

3 1 1 i I I i . FBR US DATA / RAPSOIE.PHENIX DATA : - • 377 -400°C 20% CW316 4 jj£L COO-435 °r SA 316 u ~o370-405°C SA 316 PH.tNIX DATA: — — - A 270 °C 20 % CW 316 / + 430°C EC316L - —- 3 _ORR DATA = / + •*430°C EC316LEB WLLD _ • 400'>[25%[W316Ti/ / x430°C 19-12-2 WELD " o400°CSA316Ti / / . I [J 2 - - / / 1. CO I ** 1 / i • < A 0 u _

* i * I 1 1 1 1 10 20 30 40 50 60 70 80 DOSE (dpa) Fig. II. 1-5: Swelling data on SA 316, 316 Ti, CW 316 and 316 Ti in fast breeders (EBRD and FFTF in the US, Rapsodie and Phenix in EC) or in mixed spectrum facility (ORP in the US)

In the temperature range of ITER FW, 60°C to 400°C, the swelling will be rather low but strongly temperature dependent. In fact below 300°C the recombination controlled low temperature regime will dominate with low or no swelling. For temperatures close to 400°C non-negligible swelling is however to be expected. Therefore the swelling issue for the FW will be related to swelling gradient rather than global deformation. The data on Fig.II.1-5 correspond to low He/dpa, 0.1 to 1, characteristic of fast breeders and should be handled with care for fusion environment. Results of swelling correlations for mixed spectrum reactor irradiations with more appropriate He/dpa ratios are shown in Fig. II.1-6.

II. 1.2.6 Stress corrosion cracking

It is generally well established that out of irradiation, SA 316-type steels are generally not prone to SCC in water containing Oo unless solution annealing is degraded by sensitization or some cold working is introduced. Under neutron

200 6.0 • FBR DATA FOR AISI316, 377-400'C OORR DATA FOR 316, PCA, 500'C # ORR DATA FOR 316, PCA, 400'C 4.0 * ORR DATA FOR 316, PCA^O'C + ORR DATA FOR '500'C LOW C PCA, 400'C

2.0

0 250

-0.5 I 0 10 20 30 40 50 DOSE (dpa) Fig. II. 1-6: Proposed fusion reactor swelling correlations for austenitic steels.

irradiation, grain boundary Cr-depletion has been reported at temperature 350°C for dose up to 1 dpa and shown to correlate reasonably well with occurrence of SCC in austenitic steels. Theoretical modelling and extrapolation of this behaviour down to 100°C show that very localized Cr depletion at grain boundaries could still occur at 100°C and for dose 1 dpa. In addition, segregation of minor elements such as Si or P and/or phase transformation such as the martensitic one could also promote sec. Preliminary slow strain rate tests on 316L type material have been performed on smooth and crevice specimens in SA and H-charged conditions. Under air or water with controlled oxygen content strain rates down to 3x10 s, and temperatures of up to 150°C do not result in SCC.

II. 1.2.7Hydrogen embrittlement

The solubilities, diffusivities and permeabilities of H-isotopes are reasonably well er.ablished in non-irradiated 316 type steel. At low temperature, 100°C, the diffusion mean path during a lifetime of ten years is < lem i.e. comparable to or smaller than the thickness of the plates involved in FW and blankets. In addition, there is some evidence that irradiation induces trapping centers which decrease the diffusivity and permeability. The austenite in Type 316 steels containing more than ~ 12% Ni is generally not prone to hydrogen cracking. Nevertheless, the low operating temperature could result in the

201 trapping of most of the H produced in the alloy and have higher impact on mechanical properties and fracture behaviour than observed on material irradiated in mixed spectrums where H/dpa ratio is only - 10 appm/dpa instead of > 50 in fusion environment. Furthermore, the presence of tritium and decay He seems to be rather effective in reducing the fracture toughness cf non- irradiated 304 or 316 type steels.

11.1,3 Main key issues

On this basis the following issues were recognized for SA 316 type reference material and welds: (i) Tensile strength and ductility after irradiation below ~ 250°C are significantly dependent on the initial metallurgical condition. The data are sparse. A special effort is to be devoted especially to welds, (ii) Low cycle fatigue data after irradiation below ~ 250°C are basically non- existent, (iii) A clear understanding of the enhancement of irradiation creep at low temperature has to be obtained, including possible coupling with pulse operating scenarios, (iv) Data on the effect of the presence of neutron flux on low cycle behaviour have to be obtained, (v) Data on fracture toughness due to irradiation below 300°C are basically missing, (vi) H-embrittlement and SCC are two issues that require clarification taking into account the particular fusion environment compared to that of fission reactors, (vii) He and H content effects on changes in properties during and after temperature excursions related to accident conditions, (viii) Data base on swelling behavior under fusion relevant He/dpa ratio should be extended.

REFERENCES

[1] ITER Specialists' Meeting on Materials Data Base, August 22-26, 1988, Garching (FRG).

[2] Structural Material Requirements and Related R and D For ITER Plasma Facing Components, J.L. Boutard, ICFRM-4, Dec. 4-8,1989, Kyoto (Japan). To be published in J. Nucl. Maier.

[3] ITER Specialist's Meeting on Materials Data Base, February 7-9,1990, Garching (FRG).

202 II.2.DISPff RSION STRENGTHENED (DS) COPPER: DIVERTOR STRUCTURAL MATERIAL

II.2.1 Basis for selection (Table H.2.1-1, [1,2])

DS Cu alloys of type Glidcop (US), CuAJO or CuZrO2 (EC) and MAGT-02 (SU) have been selected as reference materials. They show a better compromise between strength and thermal conductivity than OFHC Cu or age- hardenable alloys such as CuCrZr, CuCrZrMg or CuNiBe. Oxide dispersion strengthening results in a very high stability of the dislocation network up to temperatures close to the melting point. Heat treatments up to ~ 1000°C for ~ lh result in a marginal effect on the mechanical strength. DS copper alloys are therefore compatible with the high temperature brazing of the protective tiles of the divertor plate. The optimized metallurgical condition for CuCrZr alloys is obtained by cold-working and aging. The aging results in (i) high thermal conductivity by removing solute atoms to Cr rich precipitate and (ii) improvement of the stability of the cold worked structure compared to OFHC Cu. Nevertheless the high temperature brazing of the protective tiles call for a solution-annealed and aged condition. Under such conditions CuCrZr alloys will have lower mechanical strength and rather modest resistance to low cycle fatigue. The high thermal conductivity of CuNiBe alloys is obtained by ?ging to remove Be-atoms from solid solution. A high density of precipitates of y-phase is obtained. The age-hardening is very effective. Therefore the solution-annealed and aged condition, which is compatible with high temperature brazing of the protective tiles, can be used with excellent retention of tensile strength and thermal cond activity at room temperature.

II.2.2. Status of existing data base

The effect of neutron irradiation on the yield strength and electrical conductivity measured at room temperature is shown in Table II.2.2-2.

II.2.2.1 Tensile properties [1,3,4,5]

Glidcop A120 and 60 irradiated in the cold-worked and stress-relieved condition have remarkable stable yield strength after irradiation at 385°C up to 15 dpa. It should be noted also that Glidcop A125 irradiated in the cold-worked condition at 450°C up to 15 dpa exhibits significant, although tolerable softening. The ductility is marginally affected by irradiation at ~ 400°C. Significant loss of ductility is to be expected, however, for irradiation at low temperature ( < 250°C). CuNiBe alloy undergoes only marginal softening when irradiated in the annealed and aged condition at 450°C up to 15 dpa. In the cold-worked and aged condition, however, this alloy suffers significant softening.

203 TABLE II.2.1-1: YIELD STRENGTH, ELECTRICAL CONDUCTIVITY AND FIGURE OF MERIT OF HIGH CONDUCTIVITY, HIGH STRENGTH Cu-BASED ALLOYS (+ INTERNATIONAL ANNEALED COPPER STANDARD; ° : FIGURE OF MERIT: Y0 2%xIACSxl0"3'*:Cold"worked;**:Aged; Stress relieved

Glidcop Al 25 (Al-0,, 0.25 Al) 20 % CW 483 84 40.6 Glidcop Al 20 (Al,0^, 0.20 Al) 70 % CW-SR 850°C 337 94 31.7 Glidcop Al 60 (Al,0,. 0.60 Al) ditto 397 87 34,5 Trefimetaux Al 30tCdA10. Hot extruded 440 85 37,40 0.30 Al) [1] HAGT o.2 CW or hot extruded 430-540 91 39.1-49.1 ditto + Annealed 430-540 91 30.1-49.1 [2] 1000 °C 1 h Haterial Condition Yield Strength Electrical FOM° Conductivity (MPa) (% IACS) Cu (99.999%) MARZ Annealed 31.58 101,103 5.8,3.2 OFHC Cu [1] Annealed 26 101 2.6 Cu-0.9Cr-0.17Zr-0.05Hg 90 %CW-A 470°C.0.5h 450 83 37.4 MZC Cu-0.5Cr-).18Zr-0.04Mg 90%CW-A 425°C, lh -430 89 -38.3 HZC Cu-0.8Cr-0.15Zr-0.4Hg SA 950°C, lOOh 90 HZC SA 935°C.45mn-A 500°C 220 54 11.9 [1] lh Cu-0.75Cr-0.5Zr SA 1000°C lh-A 520°C 296 75 22.2 lh Cu-0.86Cr-0.07Zr SA 1000°C 0.5h-A480°C .600 2h-63% CW ditto + Annealed -200 [1] 700"C 15 nm Cu-1.8 Ni-0.3 Be Annealed-A480°C, 3h 561 61 34.2 [1] 20 % CW-A 480°C, 3 h 563 74 41. TABLE II.2.2-2: EFFECT OF FAST NEUTRON IRRADIATION ON YIELD STRENGTH, ELECTRICAL CONDUCTIVITY MEASURED AT ROOM TEMPERATURE FOR PURE COPPER, AGE HARDEN ABLE (CuZr, CuZr,CuNiBe) A1" ID ODS Cu- BASED ALLOYS

Material Irradiation upto 3 dpa irradiation upto 15 to 16 dpa 385°C 385° 450°C Y %IACS FOM Y % IACS Fom Y0.2% %IACS° FOM Q 2% Q 2% Glidcop A125 396 73 28.9 Glidcop A120 353 90 31.8 343 88 30.2 Glidcop A160 402 83 33.4 379 81 30.7 [3.4.5] CuNiBe - CW and aged 211 84 17.7 - annealed and 451 71 32.0 aged [1]

+ Yield strength: MPa;°: international Annealed Copper Standard; ,-3 * Figure of Merit: YQ 2% x IACSxlO " II.2.2.2. Electrical conductivity The effect of fast neutron irradiation is to decrease the electrical conductivity modestly (less than 10% IACS) except for CuNiBe, for which irradiation results in a small increase in conductivity. The decrease in conductivity of CuCrZr and Glidcop alloys is mainly due to the accumulation of Ni and Zn produced by transmutation of Cu-atoms [6]. A decrease of 10% IACS would imply an increase of ~ 50°C of the maximum temperature at the location of the heat flux peak of 15 MW/m Therefore the decrease in conductivity quoted in Table H.2.2-2 is most likely to be tolerable for DS Cu type alloys.

11.2.23 Swelling The resistance to swelling of DS Copper is extremely good. Upto ~ 100 dpa under fast neutron [1] or Ar ion [2], Glidcop and MAGT0.2 exhibit swelling below ~ 1%.

H.2.2.4 Microstructure DS Cu of type Glidcop or MAGT 0.2 show vciy high stability of dislocation density and y-AL^O^gg particles under fasst neutrone n irradiation [Ij or Ar ions [2J. Glidcop Al 60 irradiated with 28 MeV Si ions has shown however partial resolution of y-A^Oj particles after 10 dpa at 350°C [7], The y-phase m CuNiBe appears to be stable with some irradiation controlled coarsening for irradiation up to 10 dpa for temperatures and dpa rates in the range 225 to 400°C and 10 to 10"2 dpa/s respectively [8]. Conversely the fee Cr coherent phase obtained after aging solution-annealed CuCrZr alloy is dissolved during irradiation up to 0.2 or 1 dpa with temperature and dpa rate in the range 200 to 530°C and 10"5 to 10"2 dpa/s [9].

II.2.3 Key R&D needs On this basis the following main issues have to be investigated further for DS Cu alloys: (i) high temperature mechanical properties such as low cycle fatigue and thermal creep have to be obtained before and after irradiation with a special emphasis on He-embrittlement (ii) dynamic fracture toughness data are to be obtained in view of the expected loss of ductility under irradiation at low temperature (iii) the effect of radiolysed water on corrosion has to be assessed with emphasis on erosion corrosion and subsequent pitting corrosion on the 316 cooling loop (iv) a possibly detrimental effect of He and H production on the good swelling behaviour under fast neutron has to be checked. Finally among age-hardenable Cu alloys, only CuNiBe could be a back- up solution if operating temperature is maintained low enough: < 400°C.

206 RHFHRLNC1-S

[1] The Structural Materials for the First Wall and Divertor of NET. J.L. Boutard, Workshop on Radiation Damage Correlation for Fusion Condition, Sept. 28 to Oct. 3,1989, Silkeborg, Denmark. [2] Structural materials divertor Copper alloys properties. Dispersion Strenghthened copper alloys. V.R. Barabash, G.L. Saksagansky, Y.F. Shevakin, V.I. Solopov, S.V. Brabets, V.V. Rybin and S.A. Fabrisiev. ITER Specialist's meetings on materials data base. Feb. 7-9, 1990, Garching, JRG. [3] H.R. Brager, H.L. Heinisch and F.A. Garner, J. Nucl. Mater. 133-134 (1985) 676-673. [4] R.J. Livak, H.M. Frost, T.G. Zocco,J.C. Kennedy and L.W. Hobbs, J. Nucl. Mater. 141-143 (1986) 160-162. [5] O.K. harling, N.J. Grant, G. Kohse, M. Ames, T.S.Lee, and L.W. Hobbs, J. Mater. Res 3 (5), Sept. 1987, 568-579. [6] H.M. Frost and J.C. Kennedy, J. Nucl. Mater. 141-143 (1986), 169-173. [7] J.A. Spitznagel, NJ. Doyle, WJ. Coyke, J.G. Greggi Jr., J.N. Me Gruer and J.W. Davis, Nucl. Instr. and Methods in Physics Research B 16 (1986) 279-287. [8] R.P. Wahi, R. Koch, C. Abromeit and H. Wollenberger, J. Nucl. Mater. 127 (1985) 175-186. [9] C. Ramatchandra, N. Wanderka, R.P. Wahi and H. Wollenberger, ICFRM-4, Kyoto (Japan). Dec. 4-8, 1989. To be published in J. Nucl. Mater.

II.3. NIOBIUM ALLOYS

Pure niobium and low alloy forms of niobium (Nb-lZr) have been studied since the early 1930's; however, it was not until the mid 1950's that serious development of niobium as an engineering material began, concurrent with the development of niobium alloys was an exiv.nsion and modification of the manufacturing techniques used on niobium alloys. For example consolidation techniques such as electron beam melting and arc casting were introduced into the industry in the late 1950's and early 1960's. Prior to this time most of the alloys were made using powder metallurgy techniques. Therefore, there is a wide variation in data developed far many of these alloys because of the differences in manufacturing techniques a ad difficulties encountered in trying to scale-up laboratory size ingots to production size ingots. The data that is presented reflects the properties that are currently commercially achievable.

II.3.1 Basis for selection In selection a niobium alloy, none of the high strength commerical alloys availabel meet all of the requirements of fusion. For example some of the current

207 alloying elements (i.e. Ta and Hf) will present difficulties in a fusion environment. However, over the years extensive alloy development efforts have been conducted and if a low strength alloy such as Nb-lZr appears to offer potential advantage over other candidate materials then there is a good chance that a alloy can be developed that is higher strength than Nb-lZr but does not have the neutronic concerns that the Ta and Hf alloying elements have. One potential candidate, that is no longer commercially made, but has a modest unirradiated mechanical data base is an alloy that was known as Cb753 and contained Nb-V-Zr. In reviewing the niobium alloy data base, next to the high strength alloys the low strength alloy referred to as Nb-lZr has the largest unirradiated data base and can be easily formed and weleded. For purposes of evaluation, the properties of Nb-lZr are presented in this section and are recommended for use in this phase of ITER.

II.3.2. Status of Existing Database

Table II.0-1 compares several physical properties with the cnadidate structural materials, while Figures II .0-1 to II.0-4 show their temperature dependence. The equations used to generate the niobium curves are presented in table II.3-1. The niobium physical properties were obtained from references 1-5. Examination of the thermal expansion of the niobium alloy is about the same as molybdenum and less than stainless steel or copper. While not shown on this figure the expansion of niobium is about a factor of 2 higher than tungsten. The thermal conductivity of niobium is about half of that of TZM but significantly better than stainless steel. The combination of modest thermal conductivity and low thermal expansion coupled with a low modulus of elasticity indicates that Nb- lZr will have a low thermal stress, which is an advantage for the divertor. The mechanical properties are presented in Figures II.0-5 to II.0-7. These properties were obtained from references 6-7. Examination of the tensile curves indicates that the strength is essentially equivalent to annealed stainless steel. Relatively little data is available on the effects of irradiation on the mechanical properties. What data that exists was conducted at low fluence and temperatures. Since niobium is susceptible to contamination from the interstitial elements carbon, nitrogen, oxygen, and hydrogen, care must be taken in interpreting the irradiation data since contamination can occur from the encapsulating coolant. The unirradiated uniform and total elongation are shown in figure II.3-1 and figure II.3-2 shows the decrease in uniform elongation after irradiation to 20 and 34 dpa. Figure II.3-3 and II.3-4 show the effects of irradiation on both the ultimate and yield strengths after irradiation to 20 and 34 dpa. Figure II.3-5 shows the fatigue strength for 300k and 973 K, (reference 8) comparison of the data indicates that the fatigue strength is essentially independent of temperature. There is a substantial amount of information on the swelling behavior of niobium and niobium alloys under ion irradiation (9,12), but limited data is available for Nb-lZr under neutron irradiation (13,14). The most complete set of

208 TABLE II.3-1 RECOMMENDED Nb-lZr PROPERTY EQUATIONS

Property

Thermal conductivity, (w/m k) X = 47.2-1.47x10'^

Thermal Expansion, (10"6/c) a = 6.94 + 1.59x10"^

Specific Heat Capacity, (J/kgk) Cp = 268 + 5.65x10"^

ElectriciU Resisitivity (10"6Q/cm)e=21.6+4 x lO"2!

Densitiy, (g/cm3) d = 8.51-2.71 x 10'^

Ultimate Tensile Strength (MPa) Ou = 283-0.356T + 6.45 x lO4!^^ x lO'7!3

Yield Strength (MPa) Oy = 206-0.170T +1.23 x lO"4!2

Modulus of Elasticity (GPa) E = 80-33.9x10'^

Elongation (%)

Total 8t = 45.9-0.026^1.078x10'^ + 1.257x10 5 2 Uniform Eu = 31-4.501 x lO'^T + 3.144xlO" T 0 12 -0.6 Low cycle Fatigue Aef = 4.22 Ou (Nf)" - +0.281n (Eu) Nf" E 0.01/xEu when: T = temperature, °C Nf = cycles to failure

neutron data indicates a swelling peak at approximately 1070 k with essentially no swelling below 900k. The rate of swelling at the peak is approximately 0.13%dpa. Because of the low operating temperatures and low fluence in ITER, no swelling is expected. The predicted swelling curves for Nb-lZr are shown in figure II.3-6.

II.3.3 Compatibility with Water Coolant

In the current divertor studies, the coolant is water. Since niobium can be embrittled by oxygen and hydrogen some concern has been expressed about the suitability of niobium alloys is this environment. While there does not appear to be any information on the compatibility of niobium alloys in water, there was a study in the early 1960's on the effect of water vapour and hydrogen on the strength of pure niobium and a Nb-5Mo-5V. In this study specimens were exposed to H20/H2 mixtures of 1:1 and 3:1 over the temperature range of 200 to 816 C for times up to 1 hour. Tensile tests were conducted at the test

209 A'.ATGRIAL PPOPEKTY AVEKACfc ELONGATION dpo-0 Temperature ( F) 500 1000 1500 2000

• • I--,- - LEGEND \— r i = UNIFORM ELONGATION}— —-i — —\— • - TOTAL ELONGATION L "7

T — i—f— _ 1 —;— '7 — o ; — o • r / o T CD 1 C O - , _ - — r . ..— — > — ^4^1|—f~ f— 1 1—t I f+ —l 1 ] ' 1 i f i 1 1 200 400 600 800 1000 1200 Temperature (°O

FIG. II.3-1 Average elongation (dpa = 0) of Nb-l%Zr, annealed

Uniform Elongation of Irradiated Nb-1Zr (670-720K!rrT)

0 -r 2C3 400 T(k) 600 800 1000

FIG. II.3-2 Uniforma elongation of irradiated Nb-l%Zr (670-720 K irr T)

210 Ultimate Tensile Strength Nb-1Zr Irradiated 670-720K 700 UTS Odpa "JTS 20dpa 6C0 UTS 34dpa

5 500

CO 3 400

300

200

100 200 400 T(k) 600 eoo 1000

FIG. II.3-3 Ultimate tensile strength of irradiated Nb-]%Zr (670 - 720 K irr. T)

Yield Stress Nb-1Zr Irradiated 670-720K 700

( • YS Odpa —•— eoo • YS 20dpa < —.. —>. D • YS 34dpa "—*= , a • =——, 5 SOO ——-.

T 300

200

—a— 100 —a

i 200 400 600 800 1000 T(k)

FIG. II.3-4 Yield stress of irradiated Nb-l%Zr (670 - 720 K irr T)

211 Nb •1 a'igue 1 -

M i 11 o 300k u • 800k i o Oala 30Ok .1 - •rp-zr u » pala Q73k" IT ^- - itl 'I n T I; . _L . 1 • r" t j i _ . ^^ T 8 I I ll '! ! i Tf! ii 1I 1! 9 .01 - .>.. r" * -' '; • • i PTT" 44 - - TTT i 1 • —— Ti T i i I a * „ n Pur aNbD ati I ir -i f 1 (1 III 1 I I 'I'l 6 7 1 C 102 103 104 10 5 '10 io H Nf FIG. II.3-5 Nb-19*~ "Vquedata

Sweling of Nb-1Zr f \ 0.05- \ /

0.04- f- \ 0.03- \ a, / r 0.02- \ - 5dpa / \ \ - 1Sdpa i / 0.01 - -o—- 25dpa s X K. -o—• SOdpa A ,^ •Q-O-r 0.00- _ 800 900 1000 1100 1200 1300 T(K) FIG. II.3-6 Nb-l%Zr swelling data temperature. For the pure niobium specimens, no ductility loss due to exposure was observed until the specimens were exposed to 560 C. In this case the elongation was reduced form 36 % to 20 %. Above this temperature the ductility increased back to the unexposed values. Analysis of the specimen tested at 560°C revealed that this specimen picked up 1300 wppm oxygen and 100 wppm hydrogen. In the case of the alloy ductility loss did not occur until exposed above

212 650°C. At 760°C the ductihty loss was from 12% to 7%. At this point the oxygen absorbed was approximately 4000 wppm and the hydrogen was approximately 400 wppm. An interesting observation during these tests is that the niobium appeared to be more stable and experienced less oxidation in a water environment that in equivalent air exposure. Since neither of these materials had reactive metal additions to tie-up the interstitials, it is anticipated that the Nb-lZr material will preform equally or better than these specimens.

II.4. MOLYBDENUM ALLOYS-PFC STRUCTURE

Molybdenum and its alloys possess a combination of good thermal conductivity and elevated temperature strength, which makes this class of materials attractive for use in the high heat flux components of ITER. The inherent limitation in this class of materials is their difficulty in fabrication as a result of its ductile-to-brittle transition temperature (DBTT), which is at or above room temperature in the annealed condition. The alloys that are discussed in this section are those which either have been alloyed to reduce the transition temperature or have demonstrated to possess reasonable fabrication characteristics.

II.4.1 Basis for Selection

Fundamental studies, in the 1950's on the mechanisms controlling the DBTT of pure molybdenum found that cold working was effective in improving fabricability and that small additions of reactive elements increased the elevated temperature capability by raising the recrystallization temperature. The molybdenum alloy referred to as TZM (Mo-0.5 Ti-0.08 Zr) is representative of this class of materials. This material is supplied in the stress relieved condition and cannot be weided because the recrystaUized material in the weld zone has a DBTT that is well above room temperature and the stress concentrations between the low stress region of the recrystalized material and the high stress region of the stress relieved material frequently produces self-initiating cracks that propagate completely through the materials. Since this material was developed in the mid to late 1950's it possess an extensive manufacturing database as well as unirradiated material property database. In applications which require a welded component it was found that rhenium additions in the range of 40-50% were effective in suppressing the DBTT to well below room temperature and this material produced ductile welds in the annealed condition. Recent work in Japan, the USSR and the US indicates that rhenium additions below 20 % result in almost the same effect as the higher concentrations, the lower rhenium addition is an advantage because of its limited availability and high cost. Data developed by the USSR indicates that this concentration could be as low as 4-10%. Because the database of the Mo-Re alloys is not as extensive as that of TZM no recommendation is made on alloy selection but instead data is presented, when available, on all three compositions.

213 H.4.2. Status of Existing Database

Of the three candidate molybdenum alloys, TZM has the largest database, compiled over a number of years and covers a wide range of heats and sizes. The database for the molybdenum-rhenium is significantly smaller than that of TZM and in comparing the two rhenium alloys the database for Mo-40 Re is much larger than that of the Mo-5 Re alloy. For comparison purposes, when data is available the properties of the candidate molybdenum alloys are plotted on the same graph.

11.4.2.1. Physical Properties

Table 11.0-1 compares the physical properties of the various candidate structural materials, while figures II.0-1 to II.0-4 show the temperature dependence of several physical properties. The equations used to generate these curves are presented in Table II.4-1. The data used to develop these curves were obtained from references a-d. Examination of figure II.0-2 reveals that rhenium additions have a strong influence on the thermal conductivity of molybdenum, with the 40 % rhenium alloy possessing half of the conductivity of the 5 % rhenium alloy and the TZM alloy. In the case of thermal expansion (see figure II.0-1), all of the molybdenum alloys are low and about one third of that of copper or stainless steel. No data were presented on the modulus of elasticity for the moly-/henium alloys. However, in an overview presentation (ref.e) at the divertor workshop, values at 500°C of 260 and 290 GPA were cited for the alloys Mo-5 Re and Mo-41 Re respectively. Examination cf figure II.0-3 indicates that these values are within the range of the TZM molybdenum data and as a result it is recommended that the TZM curve be used for all three of the molybdenum alloys. No information was presented on either the electrical resistivity or the specific heat of the moly-rhenium alloys. Based on the thermal conductivity data, it is expected that the resisitivity of the Mo-5 Re alloy should be close to that of the TZM alloy (see figure II.0-3) and that of the Mo-40 alloy higher.

II. 4.2.2 Mechanical Properties

The ultimate tensile strength and the yield strengths of the molybdenum alloys are presented in figures II.0-5 and II.0-(> respectively, the data for these properties were obtained from references g-i. The ultimate tensile strengths of all three of the molybdenum alloys are higher than the other candidate meals. The 316 stainless steel is competitive with the Mo 5 % Re alloy up to about 500°C but beyond that point it begins to rapidly lose its strength. The higher strength of the molybdenum alloys is to be expected since they are in the cold worked and stress relieved condition. In the annealed or recrystalized condition the strengths would be substantially reduced to about the level of the 316 stainless steel. Depending upon the amount of cold work in the structure the temperature for recrystallization is in the range of 1200-1450°C with the Mo-5 Re having the lowest temperature and the TZM alloy having the highest. The yield strengths

214 TABLE II.4-1 RECOMMENDED MOLYBDENUM ALLOY PROPERTY EQUATIONS

Property Units TZM Mo-5Re Mo-40Re

Thermal conductivity (1) W/m k 119-4.46xl0"2T+2.26xl0'5T2 111-9.57xlO'3T 52-9.57xlO'3T Thermal Expansion (a) 10 /c 4.94+2.72xl0"3T-1.43xl0'6T2 5.57+4.37xl0~4T-l.lxl0~/T2 5.7+6.6xlO"4T Specific Heat Capacity(C ) J/kg k 268-3.51xl0"2T+7.27xl0"5T2 249+3.6xl0'ZT+2.71xl0"5T2 221+3.57xl0'2T+1.97xl0"5T2 Electrical Resistivity (e) 10'eQ/cm 5.31+2.85xl0'2T 18+3.71xlO'2T-5.7xlO'6T2 Ultimate tensile strength (Ou) MPa 936-1.04T+1.52xl0"3T2-7.94x10"'^ 712-0.756T+4.32xl0'V 1165-1.13T+7.26xlO'V Yield Strength (sy) MP? 772-C.478T+6.77xlO"4T2-4.17xlO'7T3 577-0.56T+3.5xl0"4T2 1094-0.991T+6.09xl0"4T2 Modulus of Elasticity (E) GPa 312-9.09xl0'2T-4.44xl0"6T2 Elongation, total (E.) % 18-4.46xl0"2T+3.2xl0"5T2

Note: T follow the same trend as the ultimate tensile strength. The elongations at room temperature and at 800°C are shown in table II.0-2. While the total elongations are all above 10 % for all of the candidate molybdenum alloys the uniform elongation is low and typically less than 1 %. The low uniform elongation and the strain rate sensitivity of the molybdenum alloys needs to be assessed for potential failures during disruptions.

REFERENCES

[1] Y.S. Toulo'iluam, ed., Thermophvsical Properties of High Temperature Materials. The MacMillian Co., New York (1967). [2] Metals handbook. Ninth Edition. Vol 2, Selection: Nonferrous Alloys and Pure Metals, American Society for Metals, Metals Park, Ohio, 1979. [3] C.A. English, The physical, Mechanical and Irradiation Behaviour of Niobium and Niobium-Base Alloys," Niobium, Proceedings of the International Symposium, H. Stuart, ed., The Metallurgical Society of the AIME, Warrendale, PA (1984). [4] L.J. Pionke and J.W. Davis, "Technical Assessment of Niobium Alloys Data Base for Fusion Reactor Applications," McDonnell Douglas Astronautics Co., COO-4247-2 (1979). [5] T.E. Tietz and J.W. Wilson, Behaviour and Properties of Refractory Metals. Stanford University Press, Stanford, CA (1965) [6] F.W. Wiffen, "The Tensile Properties of Fast Reactor Neutron Irradiated 3.C.C. Metals and Alloys," Defects and Defect Clusters in B.C.C. Metals and Their Alloys. Nuclear Meallurgy, Vol. 18, the Metallurgical Society, AIME, 176 (1973). [7] N. Nagata, K. Furuya, R. Watanabe, "Low Cycle Fatigue Behaviour of Blanket Structural Materials," J. Nucl. Mater.. 85&86.839 (1979). [8] N.S. Stoloff, "Cyclic Deformation of Refractory Metals for First Wall Application," J. Nucl. Mater., 85&86. 855 (1979). [9] B.A. Loomis, 6.B. Gerber, and D.E. Busch, "Reduction of void Number Density and Size in ion-irradiated Ti-Coated Nb," J. Nucl. Mater., 73, 58 (1978). [10] B.A. Loomis, S.B. Gerber, and D.E. Busch, "Reduction of void Number Density and Size in Ion-Irradiated Ti-Coated Nb", J. Nucl. Mater., 73, 58 (1978). [11] J.L. Brimhall and G.L. Kulcinski, "Void Formation in Ion Bombarded Vanadium,11 Radiation Effects, 20,25 (1973) [12] B A. Loomis et al., "Effect of Oxygen Impurity on void Formation in Ion bombarded Niobium," Defects and Defect Clusters in B.C.C. Metals and Their Alloys. Nuclear Metallurgy, Vol 17, the Metallurgical Society, AIME, 332 (1973) [13] H. Jang and J. Moteff/The Influence of Neutron Irradiation Temperature on the Void Characteristics of Niobium and Niobium-1% Zirconium Alloy," Proc. 1975 Int. Conf. on Radiation Effects and

216 Tritium Technology for fusion Reactors, J.L. Watson and F.W. Wiffen, eds., CONF-750989,1-106 (1976). [14] F.W. Wiffen, "Radiation Damage in CTR's," Proceedings of the International Working Sessions on fusion Reactor Technology, CONF- 710624,140 (1971).

II.4.2.3 Radiation Effects

There have been several investigations on the effects of neutron irradiations on the mechanical properties of molybdenum alloys. Examination of the data reveals that there is a wide variation in the radiation effects depending upon the initial condition of the starting material and the material composition. For TZM the data ranges from 0.7 to 4 dpa and irradiation temperatures of 370- 550 C (8,9). Analysis of this data revealed that for irradiations at 370 C, the strength of the annealed TZM after irradiation approached the strength level of the stress relieved material, which is the strength shown in figure n.0-5 and the total elongation was reduced to the level predicted by the equation shown in table II.4-1 for total elongation. Irradiation at higher temperatures (550 C) resulted in ductilities below 5%. While not presented in the data it is anticipated that there was a corresponding increase in the DBTT to well above room temperature. The limited radiation data on Mo-Re alloys which was presented at the ITER Workshop appears to indicate that this class of materials behaves in a similar manner to TZM and that the irradiation appears to suppress the beneficial effects of the Re doping and thermomechanical treatments (3,4). In a presentation at the July meeting (4) it was indicated that for irradiation at 300 C and a neutron dose of 10 n/cm the DBTT shift for Mo-Re class of alloys was 300-400C. This appears to be in conflict with a presentation in February (3) in which it was indicated that irradiations between 300-600 C and neutron doses of 10 to 10 n/cm that the rhenium additions were effective in suppressing the shift in the DBTT. Until more data are developed or presented the issue on the effects of rhenium on the DBTT shift after irradiation remains an open issue.

REFERENCES

[1] Anon., "Report on the Mechanical Properties of Tungsten and TZM Sheet Produced in the Refractory Metal Sheet Rolling Program," Southern Research Institute, SR-66H31 (AD-638631), 1966 [2] P. Falibriard and G. Nicolas, The Study of Refractory Metals for the NET Divertor," CIME BOCUZE, CDRMO-W-5-88,1988 [3] S.A. Fabritsiev, et al., "Refractory Alloys for Divertor," ITER-IL-NE-7- 0-4,presented at the ITER Structural Materials Database Meeting, February 1990, Garching, FRG. [4] V.R. Barabeash, et al., "Materials for Divertor", ITER-IL-PC-4-0-5, presented at the ITER Divertor Working group Meeting, July 1990, Garching, FRG.

217 [5J M.A. Merrigan and L.B. Lundberg, "An Initial Evaluation of Molybdenum Alloy for Reactor heat Pipes, "Los Alamos National Laboratory, LA-UR-83-1328, August 1983 [6] Anon., Technical Data Sheet-Rhenium and Rhenium Alloys, "Rhenium Alloys, Inc. (1978) [7] J.L. Boutard, "Assessment of Structural Materials for ITER Divertor Plates," presented at the Plasma Facing components Workshop, July 1990, Garching, FRG. [8] J.M. Steechen,"Tensile Properties of Neutron Irradiated TZM and...

II.2.5 BRAZING ISSUES

Brazing will be extensively used either to joint protective tiles to structure: TiCuSil (68.8% Ag; 26.7% Cu; 4.5% Ti), TiCuNi (70% Ti; 15% Cu; 15% Ni) or NiTi (90% Ni 10% Ti) or to have structural function in the FW: BNi6 (Ni; 11% P) or BNi7 (Ni; 13% Cr; 10% P). The behaviour and mode of failure of these joints have to be assessed under cyclic loading, neutron irradiation and H-isotope environment. The data on such behaviour are non-existing. From fundamental studies it is, however, clear that important effects due to microstructural changes are to be expected: most of the intermetallic compounds involved in the brazed joint being disordered or imorphized with important recoil effect expected for 14 MeV neutrons [1,2,3].

REFERENCES

[lj H. Mori, H. Fujita, M. Tendo and M. Fujita, Scripta Met. 18 (1984) 783 [2] C. Massobrio, V. Pontikis and G. Martin, Phys. Rev. Letters 62, March 1989,1142 [3] P.R. Okamoto, L.E. Rehn, J. Pearson, R. Bhadraand, M. Grimoditch, J. of Less Common Metals, 140 (1988) 231

218 m. BLANKET MATERIALS

The materials data base for candidate cermaic breeders, beryllium, Pb- Li eutectic alloy and aqueous salt breeders are presented in this section.

III.l Ceramic Breeder

III. 1.1 Basis for Selection

The functions of a ceramic breeder material are: 1) production of tritum from the neutron/lithium reaction and the release of tritium to the purge stream; 2) production of thermal energy and the conduction of this energy to the blanket coolant; and 3) shielding of neutrons. Thus, in evaluating ceramic breeding candidates for ITER, Li-atom density, tritium retention/release properties, and thermal transport properties were considered as primary parameters. However, because of special features of ITER, the tritium retention/release properties were rated highest of the three. In order to achieve a net tritium breeding ratio (TBR) close to 1, most blanket designs employ the use of enriched Li and of large quantities of Be multiplier material. This decreases the significance of the Li-atom density. Also, because ITER is not being optimized for power production, the thermal transport properties, while still very important, are not of primary importance to the main function of the blanket. For example, acceptable performance of the blanket is conditional upon satisfying many temperature- dependent constraints. As the breeder temperatures depend strongly on the ceramic thermal conductivity, it has an important indirect role in blanket performance. Beyond the tritium production/retention/release properties of breeder ceramics, the thermal, chemical, and irradiation stability properties of the ceramic breeder are important in lifetime and safety considerations. In particular, chemical and "thermomechanical" compatibility with stainless steel cladding (and with Be for some designs) is significant. Four ceramic breeding materials (LkO, Li^SiO^, L^ZrOg, and LiA]O2) were selected for detailed review based on their good tritium production/retention/release properties. The detailed review of the data base and recommendation of materials properties correlations were focused on baseline physical properties, baseline mechanical properties, chemical stability/compatibility and radiation effects, as well as tritium solubility/transport. While the choice of an optimum breeder ceramic, fabrication form and overall geometry is a design decision based on neutronics, tritium retention-transport-processing, thermal transport, compatibility/mass- transport, reliability, safety, and economics, the ceramic breeder properties are essential input to the design team for this optimization process.

219 TABLE III.l-l. SUMMARY OF RECOMMENDED CERAMIC BREEDERS

r O o 55

Li Atom Density o A. Large Scale Production Feasibility A. Activation o X A Thermal Conductivity o X X Melting Point/Phase Change o o 0 Thermal Expansion O o Mechanical Strength Jk> O o Qecp/Ductiliry J± X X X Chemical Stability Jk. 0 o Water Compatibility X •A. o Structural Compatibility A O 0 Be Compatibility 9 t •. o o Li Mass Transfer 9 • O o o Tritium Solubility Tritium Release (Transport) o o X Irradiation Effects on Physical Properties o o o o Irradiation Swelling X o o Eradiation Stability • X 0 Jk> Radiation Effects (T-Transport, Compatibility) 9 9 9 9 • • • •

220 HI. 1.2 Cr.'ical Issues

The critical issues for the four breeeder ceramics are summarized in Table III.l-l. LUO/Be compatibility is listed as a critical issue because of the high chemical potential for Be/L^O interaction. The kinetics of this reaction may be inhibited by the formation of a thin BeO layer. Further reaction may depend on the stability at this layer under irradiation. In most designs proposed for ITER, the Be and breeder are separated by a stainless steel cladding to avoid this potential problem. However, because of attractive neutronics and thermal- transport properties of a mixed Be-breeder bed, in-reactor experiments (e.g., SIBELIUS) are underway to explore this compatibility issue. The second critical issue for LuO is Li mass transfer. Based on small scale experiments and calculations, if the only source of oxygen comes from Li burnup, then mass transfer can be reduced to acceptable levels by proper control of the purge flow system. However, because of the potential of Li mass transfer from hot-to-cold regions of the blanket, the performance of the system needs to be confirmed on a component-scale level with prototypical temperature gradients The final critical issue for LUO is common to all of the breeder ceramics. Much of the data base for tritium transport and irradiation behavior comes from thermal and mixed-spectrum reactors at relatively low dpa and burnup. While no problems have surfaced to suggest serious degradation of properties with burnup and neutron exposure, it is important to demonstrate satisfactory performance for Li burnups up to 5 at. % and for anticipated end-of- life dpa values up to 50 dpa.

III. 1.3 Status of Existing Data Base

The data base for each breeder ceramic was reviewed in detail during the February 7-9, 1990 ITER Specialists' Meeting on Blanket Materials Data Base. Appendix I to the Summpary Report from this meeting contains recommended properties correlations for the materials whose data base is sufficiently developed. In some cases, where considerable uncertainty exists, the data base was merely summarized. The reader is referred to the Summary Report for the specific correlations. Graphs and tables are presented in this document for some of the more important properties. The assessment of the breeder ceramic data bases is contained in Table III.1-2. Ceramic breeder properties were grouped under baseline physical properties, baseline mechanical properties, chemical stability/compatibility, tritium solubility/transport, and radiation effects. With regard to baseline properties of ceramic breeders, the data bases for LLO and LiAlCK were judged to be reasonably complete. The open circles for Li^SiO^ and LuZrOo vapor pressure indicate that only a single set of limited data is available. This ranking suggests that more work is required in this area in order to develop a reliable data base for the vapor pressure of these ceramics. The solid circles for LUZrO-i melting temperature and LLSiO* thermal conductivity reflect the fact that new data sets have become available which are

221 TABLE III.1-2. ITER MATERIAL DATA £

Non-Structural Materials

Ceramic Breeder Materials 'Cn o Baseline physical properties Density_ Melting temperature Vapor pressure Thermal expansion o o Thermal conductivity Specific heat Baseline mechanical properties Elastic modulus A , •i Poisson's ratio o JL. Fracture strength Tensile Compressive O O Bending Strength o j± o J± Creep properties o

Chemical stability/compatibility Composition/purity Stability. Vapor pressure/transport. O O O Compatibility Water Beryllium. SS

Tritium solubility/transport Tritium solubility Tritium diifusivity O O • Adsorption/desorption properties •

Radiation effects Physical properties. O O O Swelling O o O o Creep. Tritium trapping/transport o O o Helium trapping/transport_ o o O o Fracture properties^ d (T

Adequate/good agreement limited/general agreement # Limited/important discrepancies O Single set of data Blank Very limited/non-existent/high uncertainties

222 significantly different from the results of previous studies. In such cases, a third data set would be useful in resolving the conflicting results. However, in the case of LiAlCU more than one recent data set suggest that the thermal conductivity is higher than previously measured values. Thus, it rates a closed rectangle meaning adequate data base and good agreement. Numerical values for the baseline physical properties are summarized in Table III.1-3. The thermal conductivity, thermal response time and linear thermal expansion strains are shown in Figs. III.l-l to III. 1-3, respectively, as functions of temperature. The behavior of other blanket materials (e.g., stainless steel and Be) are also shown for reference purposes. The data base for the mechanical properties is, in general, more limited than for the other properties. The important variables in establishing correlations are porosity and grain size, temperature, and fast-fluence/burnup. There is a reasonably complete data base for the mechanical properties of unirradiated LLAJO2. More data are needed for the compressive and bending strengths of L^O, the compressive strength of Li^SiO., and Poisson's ratio, compressive strength and creep properties of LLjZrO^. The mechanical properties are important in determining whether the breeder materials will crack under differential thermal and swelling strains and how much pressure they can exert on cladding materials if contact is established. Values based on current correlations are listed in Table III.1-3. The chemical stability/compatibility properties are important in establishing temperature limits and optimum purge gas composition and impurity control. LUO has received the most attention in the area of vapor pressure (LiOH/T)/transport/(Li). Calculations indicate that vapor transport should be considerably less for the ternaries. However, additional data are desirable to confirm these expectations for the ternary ceramics. With regard to breeder compatibility with Be, baseline data are needed for LUO/Be, particularly as one of the proposed ITER designs calls for mixing these two materials in a pebble bed. The data base on breeder/stainless-steel is reasonably complete. A recent data set for LUO/SS indicates that the reaction may be slower than indicated by previous studies and that the reaction rate may depend on the purity of the LuO. This observation should be confirmed by a third study. Figure III.1-4 shows the results of current correlations for stainless-steel/breeder and stainless-steel/Be wastage at the end of 3 MW-y/m average neutron wall loading. Tritium solubility and transport have been studied in significant detail for L^O, which is the only breeder material for which single-crystal tritium diffusion measurements have been made. The solubility data for LLO/THUO are quite good and consistent, but the solubility data for the LUO/^ system are quite scattered (factor of 15 spread in the data) and sensitive to the measurement technique. Thus, the closed circle in the table refers specifically to the solubility of gaseous hydrogen isotopes in reduced form in LUO. The desorption/decomposition data for moisture release from LiOH and LijO have a factor of 10 scatter, with the main cause appearing to be differences in measurement technique (e.g., tritium counting vs. mass spectrometer). Adsorption, as with solubility, needs to be determined more reliably for both the

223 TABLE III.1-3. PROPERTIES OF ITER CANDIDATE BREEDER MATERIALS AT 80% DENSITY, 10 MICRON GRAIN DIAMETER, AND 90% Li-6 ENRICHMENT. VALUES IN PARENTHESES ARE ESTIMATED OR EXTRAPOLATED WELL BEYOND THE DATABASE.

Property Li Si0 LiA10 Li20 4 4 2 3 2 RT Density, g/cm3 1.529 1.862 3.287 2.065 Li-6 Density, g/cm3 0.587 0.346 0.235 0.172 Tmelt, C 1432 1255 1695 1750 Thermal conductivity at 600C, W/m-K 3.54 0.82 1.42 2.83 Thermal diffusivity at 600C. mm2/s 0.857 0.300 0.357 0.715 Thermal expansion at 600C, % 1.50 1.41 0.57 0.62 Young's modulus, GPa RT 70.0 56.3 89.8 75.3 600C (60.7) (48.2) (77.3) 70.5 Poisson's ratio RT 0.19 0.24 0.2 0.22 600C (0.19) (0.24) (0.2) (0.22) Compressive strength, MPa RT (65.8) 253 145 86.2 600C (28.4) (-) (-) 39.6 Bending strength, MPa RT 50.1 36.8 41.1 16.7 600C (21.6) (-) 39.8 7.7 Creep rate, {A/m-s (800C, 20 HPa) 4.0 0.184 («2) (5.9E- Swelling (vol.%) at 1.E21captures/cm3 500C -3.0 1.5 0.8 0.4 Corresponding 700C 7.3 1.4 0.5 0.6 grain diameter (i,m 6 2 2

Tmin (C) for 1-day tritium residency time 320 390 320 450

Corresponding grain diameter (Am 16 21 1 0.4

224 500 400 300

200

100 Li4Si04(8mm,80%) a ^ 50 n 40 Li22r03(8n.m,80%) Ij

< CE 20 h- (80%) X LiAIO2(8mm,8O%TD) IOh

LiAIO2(8O%) 5 — 4 —

3 —

2 — 0.81

200 300 400 500 600 700 800 , ii T,°C 200 300 400 500 600 700 800 T,°C Fig. III. 1-1. Comparison of thermal conductivities of Fig. III.1-2. Comparison of thermal time constants of solid breeder ceramics, Be, and 316 SS. solid breeder ceramics, Be, and 316 SS. 9ZZ

THERMAL AL/Lo, % zr 2 3 Cu 13 0) en 01 • o 35' ^^ o W 5** a OJ

Compa r O 3 5 o

n 40 0 ence d i SO ^8 | O UOSI . •—^»

o th e 00 9 n _ n5' o 03 o cera r ther r 80 0

STEEL WASTAGE, mm. o o o b ro tn CD

OH (U20)

-JOG 500 600 700 800 900 I COO TEMP CO

FIG. III. 1-5 Solubility of H and OH in Li2O and LiAlO2 at 10 Fa pressure of or

Li2O/H2 and Li2O/H2O systems. On the other hand, solubility/ adsorption/ desorption are reasonably well characterized for the LiAlO2/H2O system, but not for the LiAlO2/H2 system. All of these mechanisms are poorly characterized for Li4SiO4 and Li2ZrO, ceramics. Figure III. 1-5 summarizes the solubility correlations for the LuO/f^O, U^O/Hj and LiAlO2/H2O systems. Radiation effects on the physical and mechanical properties, on swelling, and on tritium transport are, in general, poorly characterized in the anticipated range of ITER fast fluence and burnup. Most of the data come from postirradiation studies of the FUBR-1A samples. While this was a good experiment, it should be pointed out that it was a closed-capsule experiment with questionable moisture control which may have had an influence on the results. Several low-burnup experiments (e.g., MOZART, EXOTIC) have been conducted with breeder materials irradiated under similar conditions. These experiments are good for comparing tritium transport and retention performance of the breeders and for deriving performance parameters (e.g., tritium residency

227 10000

1000 LiAICL

CD 100- £ lDay CD 10- O U2Si03 C LLO CD T3 CO CD

or / Li2Zr03 0.1-

727 560 •4/ 352 283 C 0.01- —i— o.o 1.2 1.4- 1.6 1.0 Reciprocal Temperature, 1000/IC

FIG. III. 1-6 Tritium residence times comparison

time) for the particular experimental conditions (Fig. III. 1-6). However, the integral results by themselves do not give enough information as to how to extrapolate to ITER tritium generation rates and purge flow compositions and rates. It is important to have each separate mechanism characterized to be able to extrapolate. One exception to these generalizations is the case of L^O for which the single-crystal diffusivity for tritium has been determined as a function of neutron exposure.

HI. 1.4 Key R&D Needs

From the discussion in Section III. 1.3 on data base assessment, it should be clear which properties and performance parameters need to be determined to provide better baseline data, to resolve conflicts in the data base, and/or to confirm predictions based on a limited data set. However, it is instructive to re- organize these comments into specific recommendations within the framework of ITER design analysis and to indicate which R&D activities are in progress or planned. This is done in the following.

Thermal Performance of Breeder Ceramics

Two of the proposed ITER designs utilize the sphere-pac or pebble-bed form of solid breeder for Li^O and Ux^ZrOy The effective conductivity of these beds depends on many parameters (size and density of spheres, surface conditions, packing fraction, gas composition/pressure/flow-rate, etc.). Experimental data have been generated for the effective thermal conductivity of

228 Li-SiO, pebble beds. An active R&D program is on-going in Japan to generate the same kind of data for lÔ. Similar data are needed for Li^ZrO^. In addition, data are needed for the LiÔ/Be mixed pebble-bed design. While data and models are available for other metals and ceramics, the extrapolation of these models to now materials often has as much as a factor of 2 uncertainty associated with it. Some data have been generated on the effects of radiation on the thermal conductivity of sintered LiÔ and LiAlO^. Irradiation effects appear to be small and primarily at the lower temepratures. However, there is a great deal of scatter in what amounts to these "single data sets." Confirmational data are needed for these materials. In the case of the other two breeder ceramics, baseline data are needed on irradiation effects.

Tritium Performance of Breeder Ceramics

For Li^O, the scatter in solubility/adsorption/desorption data is too large to allow calculations to be made with confidence at this time. In particular, consistent data sets are needed for the LUO/H^ system under anticipated protium levels (100-1000 Pa) and moisture levels (1-50 Pa). The single-crystal diffusivity of tritium in LuO is well established at low burnup and dpa. Confirmation data are needed to lithium burnups at least as high as 5% and to much higher dpa. For the three ternary ceramics, the same information is needed as for Li^O. In addition, lattice diffusion coefficients are needed. It is very difficult to interpret, model, and extrapolate the results of in-reactor tritium- transport tests without a baseline knowledge of the lattice diffusion coefficient for these materials. Work is in progress (USA and France) to determine the lattice diffusion coefficient of LiAlO~.

Mechanical Performance of Breeder Cenmics

In general, mechanical properties data are needed for the breeder ceramics to determine deformation rates which may result in breeder/cladding contact and stresses on the cladding and to determine the likelihood of breeder cracking which may affect the heat transfer resistance of the ceramic. The properties needed are outlined in the following. The Poisson's ratio for l^ZrCX, needs to be determined. This parameter enters into the formula for calculating internal stresses in the breeder ceramic due to differential thermal expansion and swelling. The fracture toughness of Li^SiO^ has been determined for unirradiated material over a wide range of g;ain sizes and porosities at room temperature. Higher temperature data are needed for this material. The fracture toughnesses of the other three materials are needed as a function of porosity, grain size and temperature. Bending strength is another useful parameter. This needs to be determined for LLO as a function of porosity, grain size and temperature. More bending-strength data are available for the ternaries. However, the data base needs to be broadened. Both fracture toughness and bending strength are

229 responses of the material to applied mechanical loads. The data can be used to try to predict the response of the material to thermal loads. However, there is a great deal of uncertainty in this calculation. Usually it gives a lower bound on the temperature gradient which will cause fracture. It would be very useful to measure the thermal shock resistance of the ceramics directly in tests with temperature gradients or changes of up to 300C. This would cover the range of most of the ITER designs. As pulsed operation is considered for ITER, the tests should also be cyclic for pulses ranging form 200-1200 s. The data base for out-of-reactor thermal creep is reasonably well developed at high temperatures ( > 700C) for all of the ceramics except LUZrO-j. However, most of the ITER designs call for a significant fraction of the breeder to operate well below this temperature. Extrapolation of the creep correlations to in-reactor temperatures below 700C is complicated by the unknown effects of radiation on the deformation rates of these materials and by the changing of thermal creep mechanisms (e.g., matrix creep vs. grain boundary sliding). Lower temperature data are needed under in-reactor conditions in order to predict the mechanical response of the breeder/cladding system after contact has been established. The more the porous ceramic creeps and hot-presses in response to load, the less the stress will be on the cladding. The swelling of the breeder is important from two perspectives. Swelling is a driving force for breeder/cladding contact and stresses on the cladding. Also, differential swelling within the breeder can cause internal stresses which may crack the breeder. The FUBR-1A experiment has provided the primary estimate of breeder swelling rates. For the two materials which exhibited the highest swelling rates (i.e., LUO and LLSiO^) more data are needed in the temperature range of 400-600C under controlled moisture conditions. In FUBR-1A the LL^O actually experienced grain growth and sintering at 500C after long exposure. Such behavior is often associated with a formation of LiOH(T) due to possibly high moisture levels in the closed FUBR-1A capsules.

Chemical Stability/Compatibility of Breeder Ceramics

The chemical stability and compatibility of the breeder ceramics are important in establishing upper temperature limits for the bulk of the breeder and for breeder interfaces with other materials. A related performance parameter is Li mass transport. Based on calculations and data, it appears that mass transport may be an issue for LUO and for LLSiO^. The baseline data for Li^O comes from two data sources with consistent results. Basically, the data show that if the only source of oxygen in the system comes from Li burnup, then mass transport can be minimized by proper control of the purge flow system. However, these findings need to be substantiated on a component-scale level with prototypical temperature gradients. For LLSiO* material, baseline data on mass transport are still needed. With regard to compatibility, the key issue appears to be Li^O/Be compatibility both in- and out-of-reactor for the anticipated temperature range of 40O-7OOC. The thermodynamic driving force for the chemical reaction is high, but

230 there is some question about the kinetics of the reaction. The reaction between L12O and stainless steel, while apparently faster man the other ceramics, is still reasonably slow at anticipated 1TER interface temperatures. However, there are some recent results which suggest that for relatively pure Li^O the reaction is even slower than previously thought. Additional data are desirable to resolve this point. While the interaction between the other breeders and stainless steel and Be appears to be reasonably slow at the anticipated operating temperatures, in- reactor compatibility data are desirable. An in-reactor test (SIBELIUS) is in progress to study compatibility at 550C.

III. 2 BERYLLIUM

111.2.1. Basis for Selection

Most of the proposed ITER blanket designs use large quantities of Be as a neutron multiplier to increase the net TBR of the blanket. In addition to its excellent neutron-multiplication properties, it is also a low activation material with high thermal conductivity.

111.2.2. Critical Issues

Because of the use of Be in both the aerospace and fission reactor programs, there is a large data base on the physical, thermal, mechanical and irradiation behavior of Be. However, much of the data has been generated for relatively dense Be with 1-2 wt. % BeO. For proposed ITER blanket designs, Be with densities ranging form 65 to 98 % have been considered. Also, there is an interest in optimizing the mechanical, tritium transport and irradiation behavior of Be with regard to micorstructure and impurity content. Thus, while there are not any major "critical" issues with regard to the use of Be, there are a number of areas in which the data base needs to be expanded to minimize the uncertainties associated with the use of Be and to optimize its performance. These are summarized in Table III.2-1 and discussed in the next section.

III.2.3. Status of Existing Data Base

The proposed ITER blanket designs call for large quantities (> 100 metric tons) of Be to enhance the tritium breeding, and in some designs, to act as a thermal resistance between breeder and coolant. The baseline physical properties of Be are understood reasonably well. The decrease in thermal conductivity with increasing porosity is well characterized and consistent for other metals (e.g., stainless steel and copper). Some confirmational data are desirable to demonstrate that Be follows the same pattern, particularly for the range of porosities (0-35%) proposed for ITER applications. Thermal properties of Be are shown in Figs. III. 1-1 through III. 1-3 of the preceding section.

231 TABLE III.2-1. ITER MATERIAL DATA BASE ASSESSMENT FOR Be

Beryllium Baseline physical properties Density_ Melting temperature Vapor pressure Thermal expansion Thermal conductivIt^T Specific heat Electrical Resisuvity__ Baseline mechanical properties Elastic modulus Poisson's ratio Fracture properties. Tensile Bending. Compressive_ Creep properties

Chemical stability/compatibility Composition/purity Vapor pressure/transport Compatibility Water SS

Tritium solubility/transport Tritium solubility__ Tritium diffusiviry_ Adsorption/desorption properties.

Radiation effects Physical properties. TT Swelling Creep. Tritium trapping/transport Helium trapping/transport_ Fracture properties

•i—Adequate/good agreement Jk.—Limited/general agreement % Limited/important discrepancies O—Single set of data Blank—Very limited/non-cxisteni/high uncertainties I

232 Mechanical properties are highly sensitive to microstructural parameters (e.g., grain size and porosity) and impurity content (e.g., BeO) which vary with fabrication technique. There is a need to re-analyze existing data to isolate properties for anticipated FTER fabrication methods (e.g., isostatic-cold- pressing/sintering) from the broader properties data base which includes data from materials fabricated by methods which are not considered acceptable for ITER (e.g., arc-casting). More mechanical properties data are needed within the anticipated ITER fabrication methods. As with the breeder ceramics, the Be is not a structural member of the blanket. However, it is important from a heat- transfer perspective to know under what differential thermal expansion and swelling strains the Be will crack and what the Be/steel mechanical interaction will be. The compatibility of Be and stainless steel is well characterized for a nonradiation environment (see Fig. III. 1-4 of preceding section). The compatibility of Be and water is quite good below 600°C. The reaction is slowed considerably by the formation of a fine layer of BeO. More information is desirable for temperatures above 600°C where the interaction is more extensive. Tritium transport/solubility/trapping in Be is poorly characterized and understood. Based on one study with unirradiated Be, tritium retention was > 62% for T <600°C. Based on a more recent study of highly irradiated Be with __ 30,000 appm He, tritium retention was > 98 % for T < 600°C and only _ 5% for T = 611°C. Chemical and irradiation-induced traps, as well as He bubble formation and interconnection, may have a strong influence on tritium retetion and release in Be. More data are needed to characterize tritium behavior in Be as a function of fabricated porosity and impurities, temperature, and He content. In the ITER designs, 1-2 kg of tritium is generated in the Be at the end of 3 MW- y/m . It is important to be able to calculate how much of this tritium is retained under normal operating conditions and how much is released during overheating events. Very limited data are available for the effects of radiation on the Be physical properties. One study showed that neutron damage has no effect on the Be thermal conductivity for temperatues above room temperature. However, He- induced swelling will degrade the thermal conductivity in the same manner as fabricated porosity. It is sufficient to account for in-reactor effects on the thermal conductivity by decreasing the porosity correction factor as a function of He swelling. No degradation of the Be matrix conductivity is anticipated. He- induced swelling has been measured as a function of fast (E > 1 MeV) neutron fluence and in-reactor time at temperature for several low-temperature reactor (e.g., ATR and BR-2) positions and intermediate-temperature (427-487°C) reactor (EBR-II) positions. Postirradiation anneals from 1-24 hours have also been performed on these samples at intermediate and high temperatures. However, all correlations have been expressed in terms of fast neutron fluence. For ITER, which has a very different neutron spectrum than the test reactors, it is important to re-write the correlations in terms of He concentration. Unfortunately, this introduces a high degree of uncertainty as conversion factors range from 2200-6000 appm (He)/1022 n/cm2 fast fluence in the ATR and BR-2

233 reactors. Also, no data are available for porous Be. More work needs to be done in this area to provide reliable guidance for design analysis. Finally, fracture and creep properties need to be determined as a function of He content and temperature for the anticipated range of design parameters. In particular, irradiation creep, which relieves differential swelling stresses and Be/steel contact stresses, needs to be determined. Detailed correlations have been recommended for the properties listed in Table III.2-1. These are contained in Appendix I of the Summary Report on ITER Blanket Materials Data Base. Values for key parameters are listed in Table III.2-2 of this section.

III.2.4 Key R&D Needs

As with the breeder materials, the main issues are thermal and tritium transport, mechanical performance, and compatibility. Also, because the Be is not a structural material, the main mechanical performance issues are deformation, stress on structural components and possibly cracking due to differential expansion. Thus, the recommended R&D is organized in the same manner as its was for the breeder ceramics.

Thermal Performance of Be The effect of porosity on the thermal conductivity of Be has not been measured directly. While there is reasonable confidence in the model for this, it is desirable to confirm the model by measuring the thermal conductivity of Be as a function of porosity in the porosity range of 0-35%. More significant than the uncertainty in the bulk conductivity, however, is the uncertainty in the contact resistance at the Be/steel interfaces. Models for contact resistance have been formulated in terms of paramters such as surface roughness, thermal accommodation coefficients, hardness of the softer material, and gas composition and pressure. Either these parameters need to be determined in separate effects tests for the range of porosities considered or an integral heat transfer test needs to be conducted for Be/Steel. The latter is probably more productive in establishing nominal heat transfer coefficients and uncertainties than are the separate effects tests. In the case of proposed designs which call for Be pebbles (either separate or mixed with LLO pebbles), there is considerable uncertainty in the effective thermal conductivity. Experiments have been conducted in the U.S. on similiar metals (e.g., Cu and Al). However, extrapolation from these metals to Be is difficult because of uncertainty in model parameters. Thus direct measurement of the effective thermal conductivity of Be pebbles would be very useful in reducing the uncertainties.

Tritium Performance of Be Based on two data sets, it appears that tritium retention in dense Be is high for T < 500°C. For Be containing large quantites of He, the retention of

234 TABLE III.2.-2. PROPERTIES OF Be AT 80 % DENSITY AND 20 (im GRAIN DIAMETER. VALUES IN PARENTHESES ARE ESTIMATED OR EXTRAPOLATED WELL BEYOND THE DATA BASE.

Property Value Porosity (P) Extrapolation

RT density, g/cm3 1.85 1283 Thermal conductivity, W/m-K (1-PHl + llP2)"1 200°C 79.7 400°C 64.1 Thermal expansion, % 200°C 0.271 400°C 0*55 Young's modulus, GPa exp (-3.5 P) RT (147) 400°C (137) Poisson"s ratio RT 0.07 ± 0.06 400°C - Compressive strength, MPa exp (-5.0 P) RT 128 400°C 97 Bending strength, MPa exp (-5.0 P) RT (199) 400°C (151) Tensile strength, MPa exp (-5.0 P) RT 128 400°C 97 Irradiated ductility, % P > 0.1 RT 0.5 400°C 7.5 Thermal creep rate (jxm/m-s) at 20 MPa (1_p2/3r3.6

400°C 2.73E-12 600°C 1.90E-8 Irradiation creep rate Qj-m/m-s) at 20 MPa and 5 dpa/y (1.01E-5) Swelling (vol. %) at 1400° appm He RT 1.32 400°C 3.17

235 tritium is low for T > 610°C. Additional studies are needed on the tritium retention in Be as a function of initial porosity and generated He concentration, as well as time at temperature. In the U.S. ITER R&D program, samples with densities ranging from 80-100% are being irradiated in the ATR reactor at low temperature. These will be available for tritium transport testing (i.e., postirradiation annealing) during the summer, with results expected by September 1990.

Mechanical Performance of Be

There are a great deal of experimental results reported in the literature on the mechanical properties of unirradiated and irradiated Be. The mechanical properties, particularly the fracture and ductility properties, are highly dependent on the method of fabrication which affects the resulting oxygen impurity content and the grain size. Initial porosity and generated He also have strong influence on the mechanical properties. It is important to first summarize the Be properties for anticipated ITER fabrication methods and then to recommend additional tests to determine the ductility, fracture strength, and yield strength as functions of temperature, fabricated porosity, and He content. This work is in progress within the U.S. ITER R&D program. An important issue in the proposed ITER designs is the temperature limits on the Be to minimize He-induced swelling in the Be and to accommodate that minimum swelling. There are some interesting low-temperature data in the literature from experiments done in the ATR and BR-2 reactors. Unfortunately, the swelling data are correlated in terms of fast (E > lMeV) neutron fluence. The correlations should be expressed in terms of He content. The conversion from fast neutron fluence (in 10 n/cm ) to He content (in appm) ranges from 1500 to 6000 for test reactors, depending on the neutron spectrum at particular locations within the reactor and the type of reactor. While the swelling rate per 10 n/cm at low temperature has a reasonable uncertanty associated with it, the swelling rate per appm He has a very high uncertainty due to the conversion factor. It is very important to the design analysis to measure the He content directly for samples which have been tested and to correlate the observed swelling with this He content. This has been done in the U.S. ITER R&D program for ATR samples irradiated to a fast fluence of 5 x 10 n/cm . The results indicate 6000 appm (He) per 10 n/cm fast fluence. A number of postirradiation annealing studies have been performed on Be samples which have been irradiated at low temperature. Generally, the annealing times are 1-24 hours. However, in trying to correlate the data in terms of He content, the same problem arises as was noted in the previous paragraph. In only one study (EBR-II) was the Be irradiated at temperature (427-487°C) with both the He (1000-2000 appm) content and the swelling (0.1-0.9%) measured. However, proposed designs for ITER call for end-of-life He concentrations from 3000-15000 appm and temperatures over a broad range. There are insufficient data under these conditions to allow a reliable prediction of Be swelling in ITER. For example, the estimates of Be swelling at 10,000

236 appm He range form 1.5-5% for T < 300°C, 2-8% for T~400°C, and 3.5-14% for 500 < T < 700°C. These uncertainty ranges are much too large for design analysis. In addition to measuring the He content for samples which have already been irradiated, it is important to generate additional data on He swelling as a function of initial porosity, He generation, and temperature. Such experiments are planned for the U.S. ITER R&D program for T < 500°C. A related problem to He-induced swelling in Be is the time dependent stresses in Be due to differential swelling within the Be and between the Be and the stainless steeel cladding. These stresses tend to be relaxed by irradiation- induced creep, which is highly uncertain for Be. It is important to determine the irradiation creep rate of Be in order to do a design analysis on blanket lifetime.

III.3. LiPb BREEDER

111.3.1. Basis for selection The functions of LiPb breeder are combined requirements to multiply source neutrons and breed tritium. Shielding of neutrons has also to be provided. Being used as multiplier and shielding material LiPb eutectic could accommodate essential volume in the blanket design which permits to use moderately enriched or natural Li. Lithium-lead breeder has lowest cost in comparison with other breeder/multiplier compositions. 17Li83Pb eutectic was chosen among possible LiPb compositions because of lowest melting temperature and adequate stability. Low melting temperature provides with easier operation of the eutectic circuit and gives a possibility for out-of-blanket tritium extraction either in batch or continuous modes. Critical issues for 17Li83Pb are compatibility of structural materials and high tritium pressure in the eutectic. It was proposed to keep eutectic below or slightly above the melting point in the blanket during operation which can solve above mentioned problems. To extract tritium the eutectic has to be melted for limited period of time and transferred to the extractor or tritium can be removed in-situ in the batch processing mode. Continuous tritium extraction would require to keep eutectic at temperatures 300-350°C in the blanket which is still satisfactory for compatibility with stainless steel but requires protection to restrict tritium permeation to the water coolant. With a low cost, absence of radiation damage, easily solvable problem of lithium burn-up and wide power operation window, LiPb eutectic provides a good breeder/multiplier option for low and medium temperature blanket designs.

111.3.2. Critical issues Critical issues and status of their updated assessment are listed in Table HI.3-1. The main critical issues arc structural stability of 17Li83Pb under

237 TABLE III.3-1. LiPb MATERIAL DATA BASE ASSESSMENT

Liquid Breeder Materials Baseline physical properties Melting temperature • Vapor pressure •I Density • Specific heat •1 Viscosity IB 'I"hermal conductivity i hermal expansion •i Electrical conductivity •1 Chemical stability/compatibility Composition/purity Stability Impurity solubility. Compatibility Structure!sec structure sectionj Air/environment Bery Ilium Electrolytic decomposition.

Tritium solubility/transport Tritium solubility Tritium diffusivity

Radiation effects Radiolytic decomposition

•i Adequate/good agreement A—Limited/general agreement %—Limited/importDxit discrepancies O—Single set of data Blank—Very limitcd/non-existent/high uncertainties

preparation and operation conditions, compatibility with structural materials and tritium permeation due to high tritium pressure in the eutectic. Also thermal expansion while melting and solidifying requires specific designing of blanket channels.

III.3.3. Status of existing data base a) Thermophysical properties

A comprehensive set of measurements of the thermophysical properties of 17Li83Pb has been performed. Results in the form of equations are presented

238 TABLE III.3-2 THERMOPHYSICAL PROPERTIES OF Pbl7Li

Property Temperature Equation Accuracy range C 4986.7 4 Specific heat 25 - 235 cp=-0.0247+3.927xl0" (t-273)H 2 (Jg"1.^"1) 235-525 c -0.195-9.116xlO"6(t+273) 3 Latent heatjOf fusion (Jg~ ) 235 DHf=33.9±0.34 Dens i tw 25-235 d =10.25 (l-126xio;;55t) (g.cnT3) 235-350 dj=9.99 (1-168x10"°) Therma1 expansion (vol%) 235 3.5

Thermal 25-235 as=0.190-3. diffusity / — ~£. _ — 1 \ 235-350 aj-1.3xl0"3+13.0xl0'5(t+273)

4 Therma i 25-235 ms=0.177 +2.94xl0" (t+273) 10 conductivity (W.cm ."C'1) 235-350 m,»l.95xl0"2+19.6xl0"5(t+273) 10 8 + Electrical 25-235 ps=0.74xl0^+10.54xl0" (t 273) resistiuity (Q.cnf1) 235-660 PplO2.3 10"6+0.0426xl0"6 (t+273) Viscosity 250-700 11.64 T)=0.187 exp RT

in Table III.3-2. [1] Properties are determined sufficiently precise in operation temperature range of ~ 300-800 K.

b) Solubilities Solubilities of Ni, Mn, Cr and Fe in 17Li83Pb were studied in temperature range ~ 500-770 K. In general terms, the solubilities lie in the range between 1 and 10 appm, slightly higher for Fe. [3] Solubility of C, N, O were also measured. Oxygen solubility in 17Li83Pb was found extremely small. Solubility and diffusivity of hydrogen isotopes in the eutectic were measured. The tritium solubility is low but * discrepancy exists between old data and recent results [3]. The results on diffusivity fit rather well. Diffusivity coefficient lies on the level of 10 ms in the temperature rfinge 550-800 K. c) Compatibility with structural materials As for compatibility of lithium lead with austenitic stainless steels the compilation of data leads to an upper limit of temperature of ~ 400°C in flowing

239 550 500

§ § g O

O - 04X16H11M3T • - 316 L ( V - i cm/s) FIG. III.3-1 Corrosion resistance of 04X16H11M3T and 316 L steels in flowing

eutectic (velocity ~ 1.5 m/s see Fig. ni.3-1) and ~ 450°C in quasi-stagnant flow (1-lOcm/s). Welded joints were also investigated. It was shown that weld metal corrosion resistance is close to base metal resistance. There is an indication that corrosion rate will be increased approximately by 3 times in strained specimens. The compatibility issue for ITER lithium-lead blanket design is a minor concern because of small the time and low temperature of the eutectic operation in the liquid state.

d) Chemical reactivity The reactivity of 17Li83Pb with air, steam and water is low. The H^ production rate is about two orders of magnitude lower as compared to Li metal. The benign behaviour of 17Li83Pb in presence of steam is confirmed by the negligible temperature increase during testing. In order to simulate the eutectic water interaction a large-scale facility has been set-up in JRC-Ispra [3]. It was concluded from testing that in all cases the maximum pressure in the reactor vessel does not exceed the injection and the temperature of 17Li83Pb is below 600°C even for an injection of highly pressurized water (up to lOOMPa). Testing of solid eutectic (~ 150°C) with a low pressure water ( ~ 2MPa) shows a relatively rapid wash-out of lithium from the alloy. Thus during testing at

240 150°C for 50 hours lithium has been completely removed from subeutectic alloy and one third of it was removed from eutectic alloy. Interaction shows no temperature increase and is of no danger in the conditions clGse to ITER lithium-lead blanket conditions.

e) Eutectic preparation

Preparation of the eutectic alloy has been done on an industrial scale with a sufficient composition control. The impurity content is on the level of 0.005, 0.002 and 10" % for oxygen, nitrogen and hydrogen correspondingly. Structural study of the prepared eutectic were performed. The structure of the eutectic change after re-melting and solidifying [2] and the effect depends on the cooling rate. Experiments show the eutectic tendency to sedimentation in the vertical channels. This results in the structure nonhomogenity and may cause plug formation in the design. More detailed study is being conducted at the present time.

REFERENCES

[1]. U. Jauch, G. Haase, B. Schultz KfK report 1/11 (1986) [2]. A. Atanov et. Al. "Properties and behaviour of 17Li83Pb eutectic". To be published in Fusion Eng. & Design, 1990 [3] G. Casini, J. Sannier "Research and development of liquid PblTLi breeder in Europe". Fourth Int. Conf. on Fusion Reactor Materials, Kyoto, Japan, Dec. 1989

III.4 AQUEOUS LITHIUM SALT BREEDER

111.4.1 Introduction

Aqueous Lithium Salt solutions have been proposed as the breeder and coolant in the alternative driver blanket ALSB (Aqueous Lithium Salt Blanket). The main objectives of the corresponding R&D programme are the completion of the data base for lithium salt solution properties and the establishment of the experimental data base on corrosion and radiolysis effects by separate tests.

111.4.2 Critical issues

The main critical issues are the lithium salt data base assessmerl, corrosion, and radiolysis issues. Corrosion experiments are performed in a corrosion test loop and in standard cells to investigate more particularly the following issues: corrosion rate stress corrosion cracking effects of radiolysis induced solution chemistry changes

241 effects of crud formation corrosion risk under adverse operating conditions such as: salt/corrosion products precipitation, local boiling design related aspects such as welds, crevices, cold working, thermal loading, flowing environment corrosion behaviour of beryllium and materials for auxiliary equipments effects of long therm exposure. The radiolysis, i.e., the destruction and the chemical transformation of the water and the dissolved products by the ionizing radiation must be suppressed in the case of a closed cooling loop in order to prevent the formation of pockets of potentially explosive H2 and CK gas and a high O9 concentration in the solution which is detrimental for the corrosion behaviour 01 the 316 stainless steel. Computer simulation and capsule tests have been used for the assessment of the gas produced by radiolysis ia the ALS Blanket. Four lithium salts were selected from a previous overall survey on potentially useful lithium salts. They were: LiOH, LiNO-,, LiNCX, and LUSO4. For reasons of radiolysis, solubility, activation and corrosion, the salt ot most interest is LiOH, and the on-going R&D programme has been more particularly devoted to LiOH, with some less interest to LiNO-,.

III.4.3 Status of existing data base

I I 1.4.3.1 Property data base for lithium sails

A literature review of the thermophysical property data base of concentrated LiOH and LiNO-, solutions as well as new measurements for LiOH have been completed. Data on salt solubility, density, specific heat, thermal conductivity, viscosity, electrical conductivity, gas solubility, surface tension, heat of solution, vapor pressure, thermal expansion, compressibility and activity coefficient are presented in the reports [1-2]. Among the new results, there are in particular the measurements of the solubilities of hydrogen and oxygen in concentrated lithium salt solution. The solubility of hydrogen at one atmosphere partial pressure of gas for LiOH (2.2 mol/1), LiNOg (2 mol/1) and Li2SO4 (2 mol/l) at 70 248C is 3.1, 4.7 and 1.8. x 10"4 mol/1 respectively. In the same conditions, the solubility of oxygen is : 2.8, 5.2 and 2.2 x 10" mol/1 respectively. For ITER conditions, LiNO^ has about 10 % lower thermal conductivity and 30 % lower heat capacity than LiOH. The results will be soon published. Measurements of density and electrical conductivity are underway. The results are expected by September 1990.

IH.4.3.2 Corrosion investigation

With respect to corrosion experiments, potentiodynamic polarization test, corrosion rate assessment by polarization resistance measurements and weight loss measurements, exposure tests and Stress Corrosion Cracking investigation have been performed on AISI316 L, DIN 1.4914 stainless steels and

242 some conventional heat exchanger materials among which base nickel alloys, tungsten, and A106 carbon steel. In addition, some exposures tests in LiOH and LINOg solutions have been carried out on beryllium samples.

a) Exposures tests

An assessment of the corrosion rate of 316 L stainless steel by measuring the weight loss on specimens exposed in 5 M deaerated or hydrogenated LiOH for approximately 3300 hours has given results in the order of a few|Am/a at 95°C (less than 10 (Xm/a), and even lower for some nickel base alloy such as the Ni- 200, while the corrosion rate is approximately 400 \lmja and some millimetres per year for the A 106 carbon steel and beryllium [3-5]. Long-term exposure tests in oxygenated LiOH at 95°C are underway. Results are expected by end 1990. A slight Intergranular Attack of AISI 316 L has been observed in LiOH at 250°C (2/10 specimens), but no particular susceptibility was noted at 95°C.[6]

b) Stress Corrosion Cracking (SCC) investigation

The slow Strain Rate Test technique (SSRT) is used for the SCC investigation on 361 L specimens solution annealed and heat-treated at 650°C for 24 hours to simulate sensitization conditions, during most of these tests the electrochemical potential (ECP) of the specimens was controlled by means of a potentiostat. Relevant ECP values were selected on the basis of results form cyclic potentiodynamic polarization measurements. With respect to 4.2 M LiOH at 95°C, the results indicate that 316 L stainless steel may be immune to SCC as long as the corrosion potential (ECP) is in the primary passivation range of the metal i.e. below about -250 mV vs SCE. This low potential can be reached in a reducing solution such as for example a LiOH solution with dissolved hydrogen. Nevertheless, 316 L stainless steel is susceptible to SCC in LiOH solution when the ECP of the metal increases to higher values as it is shown in Fig. III.4-1 [6]. These values can easily be reached if the LiOH solution becomes oxygen-saturated or if high concentration of hydrogen peroxide (H9O2) are present. Exact critical concentrations for these species still need to be defined, but in most air-saturated solutions, the ECP seems to stay below -300mV, i.e. in the immunity zone of the metal. No significant differences were found in the SCC behaviour of the solution annealed and the 650°C heat-treated 316 L stainless steel [6,7]. A complementary series of similar experiments was also carried out in LiNOj sol itions. Until now, no sign of SCC has been found for the 316 J.SS [6]. These results are confirmed by U-bend exposure tests. No sign of SCC were seen in the U-bend exposed in 5 M LiOH solutions at 95°C for 2000 hours. Longer exposure tests are underway. However, SCC has been observed after 2000 hours on 316 L specimens exposed in 4.2 M LiOH at 250°C, while no sign of SCC has been seen in 2.9-5.8 M LiNOj solutions at 250°C, while no sign of SCC has been seen in 2.9-5.8 M LiNOg solutions at 250°C for the same duration. Experiments with longer exposure and on other materials are still underway.

243 Potential (mv vs SCE. eoo ^

-CO

400 ..

200 .. IGSCC (TGSCC) + IGSCC

0 . TGSCC* IGSCC TGSCC * (IGSCC)

-200.. IGA

DF -400 . _L

-600 .

-800 •H !0"-2 10*-1 10*+l l0*+3 10"+3 10"+4 Current density DF : Ductile fracture IGA : Intergranular attack IGSCC : Intergranular stress corrosion cracking TGSCC : Transgranular stress corrosion cracking FIG. III.4-1 Cyclic polarization curve of 316 L stainless steel in air saturated 4.2 M LiOH at 95°C

IIL4.3.3 Radiofysis investigation

The experimental programme aims at determining the main observed radiation chemical yields G(x) values ,(i.e., the final number of produced molecules or ions per 100 eV absorbed energy) for candidate ALSB solutions such as the LiOH and LiNO-j solutions hi particular, and at investigating the possible suppression of radiolytic gas production for the same solutions by the application of hydrogen overpressure. In-reactor capsule experiments on 4.3-4.5 M LiOH solutions indicate that the formation rates of H? (G(H2)) are about 0.6 (CRNL [8]) to 1.4 (Mol[9]) molecule per lOOeV while G(O2) are about 0.25 (CRNL) to 0.6 (mol). for a 4.35 M LiNO^ solution, a higher O2 formation rate is observed, together with NO2, while for a 2.18 M Li2SO4 solution, a higher H2 formation rate is observed [9]. The CRNL experiments are conducted at dose rates up to 4 10 Gy/s (0.4 MRad/s) and doses up to 9.6xlO5 Gy (96 MRad). The Mol experiments are

244 conducted at maximum dose rates of 50 Gy/s (0.005 MRad/s) and doses up to 106 Gy (100 MRad). Computer modelling [ 10,11] suggest that a hydrogen overpressure of 0.1- 0.2 MPa should be sufficient to suppress the radiolysis in the case of LiOH solution. With respect of LiNOo, the direct effect of the radiation on the nitrates causes a linear increase of the Oj concentration with dose. The high concentration of nitrate and its radiolytic products prevents completely radical induced recombination of nitrate and its radiolytic products prevents completely radiolytic induced recombination of water and there is no effect whatsoever of hydrogen addition. In-reactor irradiation (50°C) of closed-capsules filled by 4.6 M LiOH solution with 0.01 M methanol indicate complete suppression of O2 formation. Assuming that the mechanism is the formation of in-situ H2 from the breakdown of the methanol molecule, the corresponding hydrogen overpressure would be less than 0.2 MPa. However, direct experiments with hydrogen overpressure of 0.06-0.07 MPa (Mol, CRNL) and 0.2 MPa (CRNL) show no suppression of O formation. Tests at higher H9 overpressure up to 1 MPa are underway (M CRNL).

III.4.4. Key R&D needs

The main critical issues listed in the previous sections will be checked by the completion of this experimental programme. The corrosion experiments will be completed by December 1990 and October 1991 for the compatibility experiments performed at the University of Toronto and the long term exposure performed at the University of Leuven, respectively. The capsule irradiation experiments with higher gas overpressures and higher dose rates for radiolysis investigation will be completed by December 1990 and June 1991 for CRNL's and SCK/CEN Mol's contribution respectively, however, corrosion and radiolysis issues will be checked in separate tests. Possible synergism effects between corrosion, radiolysis and flowing conditions have to be studied, such as for example, the effect of transient corrosive species produced by radiolysis on the corrosion behaviour of the 316 stainless steel in lithium salt solutions. It was originally planned to build an in-pile loop to study these effects. After the decision taken in the summer 1989 session to keep the Aqueous Lithium Salt Blanket as an alternative candidate for ITER, the construction of this in-pile loop was cancelled.

RHFl£RLNCtS

[1] P. Gierszewski, P. Finn and D.W. Kirk. "Properties of LiOH and LiNO3 aqueous solutions". CFFTP-G-8932, July 1989. [2] A.J. Elliot, M.P. Chenier and D.C. Quellette. "Solubilities of hydrogen and oxygen in concentrated lithium salt solutions". CFFTP-G- 8913/AECL-10012, June 1989

245 [3] P. Gierszewski and al. "Preliminary measurements of materials compatibility with LiOH and LiNO-j aqueous salt solutions at under 100°C". CFFTP-G-8919, June 1989. [4] J.W. Graydon and D.W. Krik,"Corrosion of Carbon steel in Aqueous Lithium Hydroxide Under a Hydrogen Blanket." CFFTP-G-9045, June 1990. [5] J.W. Graydon and D.W. Krik, "Corrosion of Nickel and Stainless Steel in concentrated Lithium Hydroxide Solutions". CFFTP-G-9046, June 1990 [6] W.F. Bogaerts et al. "Hand-outs of the March '90 NET project meeting in Garching on March 13,1990." [7] J.H. Zheng and W.F. Bogaerts, "Causistic Stress Corrosion Cracking of stainless steel 316 L in concentrated LiOH". To be published in Journal of Nuclear Materials. [8] J. Elliot and M. Chenier/'Radiolysis of the ALSB". Proceedings of Canadian Nuclear Association/Canadian Nuclear Society. Toronto, June 1990. [9] A. Bruggeman et al. "Radiolysis experiments for the ALSB". Progress report of July 1990. [10] A.J. Elliot and D.R. Mccracken, "Computer Modelling of the Radiolysis in an ALSB-10013, June 1989 [11] P. Lorenzetto et al. "Radiolysis in EC ALSB". The proceedings of the 16 th SOFT, London, September 3-7,1990.

246 IV. PLASMA FACING MATERIALS

Plasma facing materials should be erosion resistant in order to minimize plasma contamination, they should retain their strength at elevated temperature, they should be thermal shock resistant, should be compatible with plasma (tritium inventory and recycling issues), and should be oxygen and water resistant at temperature > 1500 K (safety issue). A nominal heat flux of 15 MW/m on divertor plates is assessed for ITER double null configurations. In addition, the energy depositions by disruption will be as high as 20 MJ/m with a thermal quench time of 0.1-3 ms. The target temperature is expected as high as 1000-1500 K, and impinging particle flux is ~10 cm" s . From these, one can deduce, that the erosion behaviour and thermal shock resistance of material are critical. Candidate materials include carbon based materials tungsten and beryllium.

IV.l CARBON BASED MATERIALS

Carbon (CFC) is a candidate as a divertor armour material because of its low Z and excellent thermal mechanical properties. It has a low neutron absorption cross section and it is thermal shock resistant. However, graphite has the following principle limitations ': i) High specific surface area implying that tritium absorption, and, hence, inventory may a critical problem. ii) High affinity for hydrogen resulting in a high chemical erosion rate and enhanced sublimation during ion bombardment. iii) High chemical erosion by oxygen impurity. iv) Relatively high reactivity with oxygen and water (safety issue at h%h temperature, T > 1500 K).

IV.1.1 Data base

Erosion by plasma

Fig. IV. 1-1 shows the erosion yields of graphite by plasma as a function of temperature, for a particle energy of 1 KeV. Four mechanisms dominate the erosion process. The physical sputtering yields for hydrogen isotopes at relevant incident particle energies ;»re given in Table IV.1-1:

247 TABLE IV. 1-1 PHYSICAL SPUTTERING YIELD FOR HYDROGEN ISOTOPES

Energy Yield (eV) (Atom/ion)

H D 3 2 100 8.7 10r 2.6 10"2 3.110I- 200 1.1 10-2 3.8 10"2 5.210r2 300 1.2 10 4.2 10-2 5.8 10,-2

Phys Sublimation 10' .Sput

o o 10" CD UJ >~

£5 10" on CD cc UJ

10 J L S00 1000 1500 2000 TIKI FIG. IV.1-1 Graphite erosion yields by plasmas as a function of temperature for a particle energy of 1 keV

The physical sputtering yields are not expected to be flux dependent. The chemical erosion yield of graphite by hydrogen is temperature, flux and energy dependent 2'3'. At the divertor relevant flux of 10 -1019cm"2s , the peak chemical erosion yield is expected to decrease to ~10 . Although the erosion yield decreased with increasing flux density, the high flux density at the divertor plates is expected to result in high chemical erosion rates in the temperature range 500-1200 K. Chemical erosion of graphite is also affected by the presence of thermal neutrals leading to a synergistic effect. For first-wall and divertor-relevant fluxes this effect could lead to a factor of two increase in the erosion rate A For chemical erosion of graphite by oxygen, the yield is close to unity for temperatures > 500 K, and it is almost independent of flux density ' '. This implies that a very low oxygen impurity level must be maintained in a D-T tokamak device.

248 15 3 10 UJ Si 20

10 UJ 8h Slope =0.93 6 - 1000 eV H + 14 I I0 O cQrD o IO13 \0" FLUX DENSITY (ArVcm2s) Slope = 0.91

-J 1013 1 300 eV H+ a: o 10\2

IOOeVH+

10' 1012 10l3 I014 10l5 1016 FLUX DENSITY (H*/cm2s) FIG. IV-2 Carbon erosion via RES as a function of ion flux density, (a) Present results for H + -- > PyG at 1500 K + (b)for5keVAr --> PyGat2100K

Radiation-enhanced sublimation is dominating in the temperature range of 1200 K to 2000 K. The erosion yield is ~2xlO at 1200 K and increases with increasing temperature to a value of 2x10"1 at 1800 K. Although proposed models *' for the radiation-enhanced sub-limitation phenomena predict a substantial decrease in the erosion rate with increasing flux densitv. laboratory simulation experiments lead to the empirical relationshio,v~ 0 "v , where y is the carbon yield and 0 is the incident ion flux density^ (See Figs. IV.1-2 and IV.1-3).

249 -3 10 20 50 100 200 500 1000 ION ENERGY (eV) FIG. IV-3 RES as function of H + ion energy at 1500 K

Erosion by plasma impurities: oxygen and carbon

The impurities, such as oxygen and carbon will cause additional erosion, which depends on impurity flux density, energy and target temperature. The erosion yield of graphite by oxygen is unity, if the graphite temperature T > 500 K. Fig. IV.1-4 shows the sputtering yield of carbon self-sputtering as a function of particle energy and target temperature. It is seen, that runaway, self-sputtering occurs at 1800 K for carbon particle energy of 300eV, whilst at 2000 K for caibon particle energy of 100 eV. The runaway self-sputtering, y=l, as function of energy and angle of incidence is shown in Fig.IV.1-5 A simple particle balance for the carbon impurity from the surface ', is given by: V + *oYo

c If Y = 1, ij; tends to infinity, there is no steady state solution. Therefore, Y = 1 is the criteria for the uppermost operation temperature for steady state operation. Chemical erosion, radiation enhanced sublimation and high tritium retention are the serious limitations of using graphite, a number of carbon based ceramic materials have been tested with respect to erosion behaviour and tritium retention characteristics.

250 I 1 I I i 1 I

below 1200 K

103 10* C* ION ENERGY (eV) FIG. IV.1-4 Carbon self-sputtering as a function of temperature and energy

25 1 i 1 c* -— r / i 100eV 300eV 3keV 2 0 - i / O x Mayerhofer 1986 i • • ® Roth et ol 1989 ! 15 - // i r 1 7 i

1.0 - /O.ikeV ''" x'

0 5- /

~"'« 0 3keV\

i i i 500 1000 1500 2000 T(K) FIG. IV.1-5 Carbon self-sputtering yield Y = 1 as a function of temperature and energy

251 All metallic impurities in graphite will suppress the chemical erosion by H-isotopes, but medium Z metals are most effective to reduce radiation enhanced sublimation effects. In considering all the aspects: oxygen gettering, H- recycling, erosion, oxygen and water activity and low activation, the following combination should be seriously considered for further development: B/C, Be/C (low Z, O-gettering) Si/C, Ti/C (most effective to suppress RES and probably O-gettering). In conclusion, CFC composite with high thermal conductivity (a > 300 W. m , c ) doped with B, Be, Si and Ti should be seriously considered as potential candidates for divertor armour material.

IV.1.2 Thermo-mechanical properties

CFC's have good mechanical strength and superior thermal shock resistance. However, conventional CFC's have lower thermal conductivities and consequently poor erosion lifetimes. CFC's with high thermal conductivity, close to or over 400 W/mK at room temperature, have been developed during the last few years, and they are expected to have good thermal shock resistance. Tables IV. 1-2,3,4 summarize the thermo-mechanical properties of various grades of carbon-based materials recently developed in Japan. The first three columns, AC in the table, are isotropic graphites and the other columns, D-G, are CFC's. Experimental results from large tokamaks and laboratory test by means of heating devices such as JEBIS have shown CFC's are preferable to isotropic and pyrolytic graphites from the viewpoint of thermal shock resistance. Figure IV. 1-5 shows the temperature dependence of thermal conductivity of these CFS's. MFC-1, an unidirectional CFC, gives the highest thermal conductivity, followed by CX-20C2U and MCI-felt. Since MFC-1 has extremely anisotropic thermo-mechanical properties, the currently available grade has been proven to show relatively poor thermal shock resistance. Test samples of CX-2002U and MFC-1, heated by JEBIS undrr 6-9 MJ/m2 showed no failure. MCI-felt has lower thermal conductivity than CX-2002U at room temperature, and vice versa at elevated temperatures above 100°C. Thermal analysis shows MCI-felt has a larger allowable divertor armor thickness than CX- 2002U under the condition that the surface temperature remain below 1000°C. Of the CFC's developed in Japan so far, MCI-felt and CX-2002U show the most preferable features under high heat flux conditions as demonstrated by the results of thermal shock tests by electron beam heating facilities and JT-60 divertor operations. However, CFC's with still higher thermal conductivity, hopefully comparable to compression annealed pyroytic graphite, are required to assure greater armor thickness for protection against sputtering and disruption erosion, and further improvement of CFC's is indispensable. For example, thermal property of MFC-1 has been improved recently up to 300 W/mk at 800°C.

252 TABLE IV. 1-2

MATERIAL (A) (B) (C) (D) MCI-FELT (E) CX-2002U |(F) PCC-1-2S | (G) JC/C ETP-1O PD-330S IG-430U PROPERTY X Z « h X |, X Z DENSITY , 1760 1790 1810 1870 ~ 1700 — 1810 — 1780 — AT RT (KG/M3) YOUNG'S MODULUS 10.2 9.6 9.5 56.0 3.1 13.1 — 14.4 — 25.3 — AT RT (GPa) POISSON'S RATIO 0.14 0.14 0.18 ~ — 0.19 — 0.18 — 0.28 — AT RT THERMAL EXPANSION 4.4 4.6 4.7 -0.4 10.1 1.3 3.4 1.0 5.8 0.1 10.7 AT RT (10~b/K) TENSILE STRENGTH AT RT (MPa) 56.4 35.1 41.7 59.0 3.2 30.1 — 26.6 — 105.2 — at 1000°C 66.9 43.1 50.4 60.5 5.2 42.5 28.5 108.6 COMPRESSIVE STRENGTH 124.1 83.5 93.9 63.9 81.0 45.0 42.7 47.4 43.5 77.4 82.0 AT RT (MPa) BENDING STRENGTH 59.8 40.5 52.3 89.4 7.7 40.5 — 31.1 — 94.3 — AT RT (MPa)

< TABLE IV.1-3

MATERIAL (A) (B) C) | (D) MCI-FELT (E) CX-2002U (F) PCC-1-2S (G) JC/C ETP-10 PD-330S [0-430U 1 X Z X I Z X Z X Z PROPERTY | THERMAL CONDUCTIVITY / it /_i/ \ AT 300K 91. 2 180.8 155.0 355.0 62 0 381.1 215.0 312.9 188.3 299.4 126.7 AT 400K 89. 7 166.7 140.0 327.5 55 3 299.2 180.0 246.5 156.7 250.0 103.3 AT 500K 85. 0 146.7 130.0 289.2 48 0 254.2 150.0 203.5 130.0 213.4 87. 3 AT 7OOK 73. 0 116.7 100.0 228.4 36 7 192.5 116.7 153 5 93. 3 160.9 66. 0 AT 900K 62. 7 95. 0 78.3 175.0 30 3 152.5 93. 3 123 5 74.0 132.5 52.3 SPECIFIC HEAT (J/kgK) at 300K 740 735 718 746 - 718 — 717 — 712 — AT 400K 1030 1067 1000 1056 -- 1000 — 1000 _. 1017 — AT 500K 1300 1333 1300 1289 -- 1267 -- 1233 — 1289 — AT 700K 1533 1600 1500 1567 -- 1555 — 1511 — 1578 — at 900k 1700 1800 1700 1700 -- 1717 -- 1700 - 1717 — TABLE I V.I-4

MATERIA (H) MFC-1 MATERIAL (H) MFC-1 PROPERTY X Z PROPERTY X Z DENSITr 1960 — THERMAL CONDUCTIVITY AT RT (kg/mJ) (H/mK) AT 300 K 425.0 YOUNG'S MODULUS 100.0 -- AT RT (GPa) AT 700 K 270.0 __ POISSONS'RATIO -- — AT 900 K 250.0 -- AT RT AT 1100 K 235.0 THERMAL EXPANSION -0.9 12.0 AT RT (10~°/K) AT 1300 K 220.0 __ TENSILE STRENGTH SPECIFIC HEAT AT RT (MPa) 400.0 3.0 (J/kgk) AT 300 K 680 -- AT 1000°c AT 700 K 1305 COMPRESS IVE STRENGTH 216.0 16.0 AT 900 K 1510 AT RT (MPa) AT 1100 K 1695 -- BENDING STRENGTH 480.0 5.0 AT RT (MPa) AT 1300 K 1935

Graphites

For the divertor in the Physics Phase pyrolytic graphite has been proposed. Pyrolytic graphite is a unique form of carbon in that it is formed by the thermal decomposition of a hydrocarbon gas. The "as-deposited" material is not a true graphite in the crystallographic sense. This is because the deposition process occurs at temperatures below 2500°C. While the material is crystalline in nature, it must be heated above 2800°C to produce a true graphitic crystal structure. The deposition is formed in crystal layers which make the material highly anisotropic. The thermal conductivity in the x-y plane is roughly 150 times greater than through the "as-deposited" thickness, referred to as the c direction. Variations in the microstructure and in the thermal conductivity can be generated via irregularities in the substrate, or by the settling of soot particles into the growing pyrolyric graphite during deposition. To reduce these variations the pyrolytic graphite is usually given an annealing treatment at temperatures up to at least 3000°C. Figure IV.1-6 shows the effect of heat treatment on he thermal conductivity. It should be noted however that annealing at high temperatures significantly softens the material and makes it more susceptible to damage during machining and fabrication. The effects of irradiation is of primary concern in the use of pyrolytic graphite because it has been shown to reduce the thermal conductivity (see

255 500

200 600 BOO 1000 1200 1400 TEMPERATURE (K) FIG. IV.1-6 Temperature dependence of thermal conductivities for several grades of CFC composites. Symbols in the figure correspond to those given in Table IV.1-2,-3, and -4.

Figure IV.1-7). In relating this data to a fusion type spectrum it canbe assumed that the E>50 KeV fission spectrum conversion factor of 8 dpa/10 n/m can be used for the more energetic 14 MeV fusion neutrons. In addition to loss of thermal conductivity the dimensional changes pyrolytic graphite can be fairly large. The c axis growth and the a axis shrinkage have been shown to produce up to 800% dimensional changes. Kennedy (ref. 11) has developed a model which includes the effects of preferred orientation, density changes, and the influence of retained fabrication stresses. His analysis for the pyrolytic gra' -, are shown in Figure IV.1-8.

IV.1.3. Critical Issues

While the pyrolytic graphites offer advantage based on thermal conductivity, critical information needs to be developed before confidence can be gained in their survival in the ITER environment. The primary concern is the rate in the loss of thermal conductivity as a function of fluence. Experiments are

256 <;uuu i 1 i 0 I - ECPG (Carbone-Lorraine) — i • HOPS I " Compression Annealed @ 3000°C (Goodrich) 1500 • - Annealed @ 3000° C (PFIZER) \ \ •1MB • As Deposited (PFIZEH) >- \

> t

1000- uo

o 500- OS UJ

5 liuTuSnunHii ••••••• IM n- 500 1000 1500 2000 TEMPERATURE (°C) FIG. IV.1-7 Comparison of conductivity of different types of pyrolytic graphite with diffeient heat treatment

1050°[7]

12 3 4 Neutron fluence (1O20 n cm"2)

FIG. IV.1-8 Effect of neutron irradiation on the thermal conductivity of graphites

257 OOTTEO LINE CALCULATED FROM TSX DATA e - - a C o

c SMALL CRYSTALLITE SIZE /

< rr \ X / Wffi; HTAR^T CRYSTALLITE SIZE

200 400 600 800 1000 1200 1400 u TEMPERATURE (°C) FIG. IV.1-9 Irradiation induced growth rates for Pyro-graphite materials

currently in progress to resolve this issue. The second issue is delamination. It has been shown that under intense short pulse heating such as occurs in a disruption as deposited pyrolytic delaminates. It needs to be shown that for a series of disruptions, that the delaminating material does not spall release large pieces of the graphite into the plasma.

IV.2 TUNGSTEN

IV.2.1 Surface Properties

The high-Z material tungsten (W) is attractive for use as a divertor armor as well as for first wall protection because of its high melting point and its high energy threshold for physical sputtering by light ions. Figure IV.2-1 shows the hydrogen (and its isotope) sputtering yield of tungsten as a function of ion energy [20]. These values are quite low and do not yield any significant erosion rates. On the other hand, Figure FV.2-2 shows the tungsten self-sputtering yields as a function of ion energy for normal incidence [18]. It is seen, that the self- sputtering yield is about Y = 0.1 at ion energy of 200 eV. This yield increases with energy and exceeds unity at an energy of 1 KeV. However, at an oblique ion impact the yield Y = 1 is reached at about 400 eV at an incident angle of about 60. For the ITER devices, the incident angle of tungsten ion is expected to be r < 15. This implies, if plasma temperature, Te < 50 eV, the use of tungsten as divertor armor looks encouraging. By interaction of oxygen impurity with W, volatile oxide species will be formed [8] and it causes impurity productions. Figure IV.2-3 shows the sputtering yield of tungsten by oxygen as a function of particle energy and target

258 I I I I llll| 1 ! I Mlllj rTTi TTTTTTT W OH, AD, D4H» ROTH OXe HECHTL 10 PW ALMEN VHg WEHNER (0,4He, W) DSPUT J^ (H,D,T,*He, X«) IPP '^"•Hg

o z or N. UJ

Z> a

10"

10 10* 10s I04 ION ENERGY, tV FIG. IV.2-1 Energy dependent physical sputtering yields for tungsten

temperature [19]. It is seen that the sputtering yield is as high as 10"1 at temperature t_>_ 1100 k at particle energy E,*,_iy_>200 eV.

IV.2.2 Physical Properties

Tungsten has a body-center-cubic (bcc) crystal structure. Its room- temperature density is 19.25 g/cnr* and its melting temperature is about 3410°C. Several of the key physical properties of tungsten are given in Table IV.2-1. These properties are also given as a function of temperature.

REFERENCES

[1] C.H. Wu, "Material for Divertor Armour", EC contribution to ITER report, July 1990 [2] E. Zolti, NET/DM/89-015, Extended and revised version, 4.7.1989 [3] C.H. Wu, "Ceramic Materials for the Next European TORUS (NET) - Thermonuclear Fusion Reactor". Proc. of Seventh Cimtec World Ceramics congress, June 24-30,1990/Italy [4] J.W. Davis, AA Haasz and P.C. Stangeby. J. Nucl. Mat. 145-147 (1987), 417

259 s

3 10 •' 10* 10 to 10' ION ENERGY [eV] ION ENERGY [eV]

Fig. IV.2-2. Tungstein self-sputtering yield as a Fig. IV.2-3. Sputtering yield of tungsten by oxygen as function of ion energy for normal incidence a function of energy. TABLE IV.2-1 TUNGSTEN THERMOPHYSICAL PROPERTIES

Temperature Thermal Expansio1 n Thermal Conducting Heat Capacity Poisson's Ratio Elastic Modulus (K) (xlO'V ) a (w/mk) b(O/kgk) Ref. 21 (GPa) Ref. 22-25 aRef. 21 Ref. 26 DRef. 22 Ref. 21 300 4.1-4.4 163 177 131 0.28 406 500 4.13-4.55 145 150 138 0.285 398 1000 *.«-4.95 122 122 146 0.288 379 1500 4.85-5.30 110 109 157 0.295 355 2000 5.15-5.7 100 101 172 0.302 359 2500 5.54-5.95 92 -- 195 -- — 3000 -- 90 -- 225 -- -- a Recommended data b 2 3 9 3 Calculated from rCnp = 135.76 [1- 4805/T ] + 9.1159xlO" T + 2.3134xl0" T (TinfcC in J/kgK) [5] J. Roth, J. Nucl. Mat. 145-147 (1987), 87 [6] A.A. Haasz et al, J. Nucl. Mat. 145-147 (1987) 412. [7] E. Vietzke et al, J. Nucl. Mat. 145-147 (1987) 425 [8] C.H. Wu, J. Nucl. Mat. 145-147 (1987) 448 [9] J. Roth and W. Moeller, Nucl. Instr. Meth. B 7/8 (1985) 788 [10] H. Trinkaus, V. Philipps and E. Vietzke, to be published [11] A.A. Haasz and J.W. Davis, J. Nucl. Mater. 151 (1987) 77 [12] P.C. Stangeby, private communication [13] M. Hugon, P.P. Lallia and P.H. Rebut, JET-5 (89) 83 [14] M. Keilhacker and The JET Team, JET-P 989) 83 [15] C.H> Wu, "Operational Limitations of Plasma Facing Materials", 5th Workshop on Carbon Materials, 17-18 May, Juelich [16] J. Roth, W. Eckstein and J. Bohdansky, J. Nucl. Mater. 165 (1989) 199 [17] C.H. Wu, E. Hechtl, H.R. Yang and W. Eckstein, "Erosion of Beryllium by Oxygen", 1-roc. 9 th International conference on Plasma Surface Interactions, 21-25 May 1990, Bournemouth, U.K. [18] E. Hechtl, H.R. Yang, C.H. Wu and W. Eckstein, "Experimental Study on the Self-Sputtering of Tungsten", Proc. 9th International Conference on Plasma Surface Interactions, 21-25 May 1990, Bournemouth, U.K. [19] E. Hechtl, W. Eckstein, J. roth and J. Laszlo, "Sputtering of Tungsten by Oxygen at Temperature up to 1900 K, Proc. ICFRM-4, Dec. 3-8, 1989, Kyoto, Japan. [20] W. Eckstein and J.P. Biersack, Appl. Phys. A 37 (1985) 95. [21] Metals Handbook. Ninth Edition. "Volume 2, Properties and Selection: Nonferrous Alloys and Pure Metals", American Society for Metals, Metals Park, Ohio (1979) [22] Tungsten: Sources, Metallurgy, Properties and Applications. S.W.H. Yih and C.T. Wang, Plenum Press, New York (1979) [23] Petukhov, V.A. and Chekhovskoi, V. Ya., High Temp., High Pressure, 4, (1972), 671-677. [24] Properties of Refractory Metals. W.D. Wilkinson, Gordon and Breach Science Publishers, New York (1969). [25] Behavior and Properties of Refractory Metals. T.E. Tietz and J.W. Wilson, Stanford University Press, Stanford, CA (1965) [26] "Report on the Mechanical and Thermal Properties of Tungsten and TZM Sheet Produced in the Refractory Metal Sheet Rolling Program - Part 1", Southern Research Institute, Report No. AD-638631, August 31, 1966.

IV.3 BERYLLIUM Beryllium has been considered seriously as a plasma facing material, because of its low Z, favourable thermomechanical properties and expected low tritium retention. However, erosion by plasma and impurities is another important issue for judging the feasibility of this application. Bulk properties data base for beryllium is summarized under blanket materials, section JJI3.

262 10"

c — / o / ° / B^o 1 m / / i /^ I/) s a e aX a 2 o io- b f A '. fa: Q I 1 ha 1 - UJ : '•/ O - 10-3 • BeO ' : UJ i— • Be at RT i— a Be at ^00K Z5 Q_ ; > Be at 9Z0K CO

10'4 i I 1 I i • i i 1 i 1 1 1 i i : 1 101 10 2 104 ENERGY (eV) FIG. IV.3-1 Energy dependence of the sputtering yield of Be by D at different temperatures

IV.3.1 Data base

Erosion by plasma The sputtermg yield of Be by deuterium as a function of particle energy and target temperature was reported by J. Roth et al. 14). It is seen in Fig. IV.3-1, at lower particle energy, the sputtering yield is increased with increasing target temperature. However, at energy E > 1000 eV, the temperature effect on sputtering yield is almost negligible.

Erosion by plasma impurities: oxygen and beryllium Fig. IV.3-2 shows the exosion yields of beryllium bombarded by oxygen as a function of the energy at room temperature and at a target temperature of 600°C 5'. The erosion yields at 600°C seems to be slightly higher than those of room temperature, but more detailed experimental investigations are needed at a target temperature of > 600°C to establish a temperature dependence of the erosion yields. It is seen that the erosion yields of beryllium by oxygen first increase with increasing energy of the impinging particles to a maximum, and then decline. The mawwnim erosion yield of ~ 510 is reached at an energy of 3000 eV. By using the TRIM-SP code ' the erosion of beryllium by oxygen has been calculated and the results are also given b Fig. IV.3-2, The dashed line

263 1 ' ' i i i t i t i I 1 1 L

• //I

10" //

•'' / j 9

- V) o UJ ''I -2 0* —Be

10 Mi l 0 RT - a 600°C

10 * i i i t I J. J i i i 1111 10' 10* 0 30° 60" 90° ION ENERGY [eV] 6. ANGLE OF INCIDENCE

Fig. IV.3-2. Sputtering yield of Be by oxygen as a Fig. IV.3-3. Sputtering yield of BeO by oxygen as a function of energy. function of angle of incidence. 90 T 1 1 I I I

LLJ O Z UJ g 60'

LL o LU _J O 30'

Es = 3.38 eV Ec = 2.0 eV

j i i i i I i i i i i i 10' 10- 10 10'

Eo. INCIDENT ENERGY (eV)

FIG. IV.3-4 Be self-sputtering yield y = 0.5,1, and 2 in the energy versus angle of incidence plane. The shacked areas are limited by calculations using Es = 3.38 and 2.0 eV represents the oxygen ions impinging the beryllium, in which a Be sublimation energy of 3.38 eV has been used as binding energy, whilst the solid line represents the erosion characterisation of oxygen ions impinging on BeO, in which the heat of formation of BeO 6.33 eV has been used as binding energy. In the energy range of 10 to 10 eV, the calculated erosionvields for 0 on Be are higher than the measured values. At the lowest energy, 10 eV. the discrepancy is a factor of 6, continually decreasing to a factor of 1.3 at 10 eV. An excellent agreement has been found between calculated 0 - > BeO values and the experimental results. It implies that the beryllium is formed to oxide by impinging with oxygen ions. The fact that the measured erosion yields are very close to the calculated values of 0 - > Be suggests that the mechanism 0 - > BeO is dominant during the experiment. It is therefore assumed that once Beryllium contacts the oxygen impurity, it will immediately form a beryllium oxide layer by gettering of oxygen, Further calculations have been performed to investigate the erosion yields as a function of energy and incident angle, namely for energies of 200,300, 500 and 1000 eV. The results are given in Fig. IV.3-3. It is seen, that the erosion yields reach a maximum 200-1000 eV. The maximum erosion yield exceeds unity, at energy of already 300 eV and an incident angle of 60°, the maximum erosion yields is as high ad 2.5 for an energy of E ~ 1000 eV and 0 ~ 700°. The assessed Be self-sputtering yield of J. Roth et al. ' as a function of particle energy and angle of particle incident is shown in Fig.IV.3-4.

265 REFERENCES

[1] C.H. Wu, "Ceramic Materials for the Next European TORUS (NET) - Thermonuclear Fusion Reactor". Proc. of Seventh Cimtec World Ceramics congress, June 24-30,1990/Italy [2] J.W. Davis, A.A. Haasz and P.C. Stangeby. J. Nucl. Mat. 145-147 (1987), 417 [3] J. Roth, J. Nnci. Mat. 145-147 (1987), 87 [4] A.A. Haasz et al, J. Nucl. Mat. 145-147 (1987) 412. [5] E. Vietzke et al, J. Nucl. Mat. 145-147 (1987) 425 [6] C.H. Wu, J. Nucl. Mat. 145-147 (1987) 448 [7] J. Roth and W. Moeller, Nucl. Instr. Meth. B 7/8 (1985) 788 [8] H. Trinkaus, V. Philipps and E. Vietzke, to be published [9] A.A. Haasz and J.W. Davis, J. Nucl. Mater. 151 (1987) 77 [10] P.C. Stangeby, private communication [11] C.R. Kennedy and W.M. Woodruff, "The Irradiaton Dimensional Change:, of Grade TSX Graphite," Proceedings "CARBON-88," University of Newcastle upon Tyne, UK, 18-23 Sept. 1988 [12] E. Hechtl, H.R. Yang, C.H. Wu and W. Eckstein, "Experimental Study on the Self-Sputtering of Tungsten", Proc. 9th International Conference on Plasma Surface Interactions, 21-25 May 1990, Bournemouth, U.K. [13] E. Hechtl, W. Eckstein, J. roth and J. Laszlo, "Sputtering of Tungsten by Oxygen at Temperature up to 1900 K, Proc. ICFRM-4, Dec. 3-8,1989, Kyoto, Japan. [14] J. Roth, W. Eckstein and J. Bohdansky, J. Nucl. Mater. 165 (1989) 199 [15] C.H. Wu, E. Hechtl, H.R. Yang and W. Eckstein, "Erosion of Beryllium by Oxygen", Proc. 9 th International conference on Plasma Surface Interactions, 21-25 May 1990, Bournemouth, U.K.

266 V. ELECTRICAL INSULATORS

The use of ceramic insulators are proposed for the RF heating, blanket and magnet systems. The former application is the most demanding since the materials are located in a neutron and ionizing environment and have to be structurally and mechanically stable (high resistance to void swelling and microcracking) and retain good RF electrical properties both during and after irradiation. Extremely good dielectric properties together with high thermal conductivity and high tensile strength is required to allow for the transmission of the power needed with tolerable energy deposition in the window. For all frequencies considered (ICRH, LHRH or ECRH) the loss tangent must not exceed 10 . Therefore, there is only a limited choice of candidate materials fulfilling the stringed requirement. A number of works have reported about the characteristics of potential insulating materials. The main objective has been focused on the radiation effect on the change in electrical properties. Fast neutron irradiation investigation has shown that helium enhanced swelling well below those required for void swelling. Table V-l shows the transmutation product (appm/year) of several ceramic insulators with neutron wall loading of 1 MW/m . Ceramics differ widely in their ability to tolerate impurities. The more simple oxides such as BeO, MgO and A^o^ tend to have very low solubilities for impurities which would normally precipitate out as the appropriate compound.

TABLE V.I. TRANSMUTION PRODUCTS (APPM/YEAR AT 1 MW/m2)

Ceramic Transmutation BeO Carbon AljOj MgAlo4 AION S13N4 Product Hydrogen 103 - 292 298 370 630 Helium 2920 2098 505 352 356 658 Lithium 134 - - - - - Beryllium - 420 - - - - Carbon 569 - 624 600 506 328 Nitrogen 31 - 34 35 - - Sodium - - 69 60 60 - Magnesium - - 436 - 400 455 Aluminium - - - - - 71 Silicon - - 8 7 6 -

267 An attempting has been made to present the status of the relevant data for the applications of ceramic insulating materials in fusion devices. Some experimental results were obtained recently. Permitivity £ of single crystals of A^Oj for unirradiated and irradiated specimens is given in Table V-2. The temperature dependence of £ can be expressed by fol' owing equations:

4 Er'(T) = £r'(290K) = 5.10" (T-290K)

TABLE V.2. CHARACTERISTICS OF THE DIELECTRIC SPECIMENS AT 290 K DETERMINED AT 28-38 GHz (d: EFFECTIVE THICKNESS). SINGLE CRYSTALS ORIENTED E c.

unirradiated

Material HEMEX Sapphire AL23 Aluminia (C.Syst. ) (Criceram) (Friedr.) 99.9% (NGK)

d [inn] 5.015 5.038 6.224 5,.058 .'(±0.005) 9.400 9.400 9.695 10..065 irradiated to 3.5 * 1019 nlcm1

d [inn] 5.004 5.038 6.234 -

,'(±0.0005) 9.470 9.465 9.810 -

For resistive components in fusion reactors, a suitable insulator materials should keep 0 limited to smaller than 10 h. T is a critical time, in which the current increases by a chosen factor, limits for safe operation can be assessed, the "radiation-induced conductivity" (RIC) can be quantified by the relation:

where OQ the conductivity without irradiation, R the dose rate, d the dose rate exponent and K a materials constant. Several ceramics, both polycrystalline (p.c.) (MgAl2O4, 99,9 % A^Og) and single crystals (s.c.) (MgO) have been identified that comply with the conductivity limits up to dose rates of 10 Gy/h and temperatures as high as 800°C. Intrinsic breakdown occurs in ceramic insulators typically at 10-10 V/rnm and thermal breakdown at 10 V/mm for advanced oxide ceramics. Resistivity degradation in s.c. MgO was observed at 1200°C for 150 V/mm with t ~ 2 h. Extrapolations yield tc ~ 100 h at 900°C (in good agreement with the experiment and tc ~ 10 h for 600°C. In the same manner, it can be excluded that dielectric breakdown in s.c. MgA^O-j is a problem in absence of irradiation. When operating the resistive component in the presence of significant radiation

268 flux, resistivity degradation can become a serious life-time limiting factor. The increase of conductivity after a certain period of steady state was first demonstrated and quantified by the experiments of Hodgson for s.c. A^Oj (cf. table V-3) and MgO, and recently also in p.c. MgAl^O^. The reduction of iQ, called radiation-enhanced electrothermal breakdown, was found to follow R" .

TABLEV.3.RADIATION-ENHANCEDTHERMO-ELECTRICBEAKDOWN IN s.c. A12O3 IRRADIATED WITH 1.8 MeV ELECTRON AT 450°C AND 130 V/mm. Tc IS DEFINED BY | (tc) = 1.21 (STEADY STATE).

Dose rate [10B Gy/h] fa 0 >100 0.30 100 0.57 20 1.00 9

The dielectric parameters of advanced ceramic insulator materials are given in Table V-4. The thermophysical and mechanical properties of several insulating materials are given in Table V-5. Table V-6 shows the neutron irradiation effects on the strength at 400- 500 °C and 1-34 dpa fpr Al^ and ALN.

TABLE V.4. DIELECTRIC PARAMETERS OF CERAMIC INSULATOR MATERIALS AT MICROWAVE AND MM-WAVE FREQUENCIES.

S.C.A190, AL23 AD-995 VITOX s.c. SHAPAL Al,o, + ordinary (99.5 % (99.5% (99.9% MgAi-O. (99.7% 5%*ZfO, 24 l ray A12O3) A12O3) Al203) AIN) er' 9.4 9.8 9.7 10.8 8.4 8.3 10.3 tandd[10~4] at 19 GHz 1.5 1.7 35 GHz 0.4 0.8 9 2 90 GHz 1.5 3.8 4 135 GHz 1.7 4.5 145 GHz 2.0

269 TABLE V.5. THERMOPHYSICAL AND MECHANICAL PARAMETERS RELATED TO FAILURE RESISTANCE UNDER THERMAL LOADS. (VALUES AT ROOM TEMPERATURE, A FOR 20-1000°C). UBS: ULTIMATE BENDING STRENGTH.

thermal th. expansion Young's Poisson UBS R' Material conductivity coefficient a modulus number [W/m K] [10"7K] [GPa] [HPa] [103W/m] AIN 140 5.5 330 0.23 300±20 18 BeO 100 9 300 0.3 200 5

Al20399.9(llm) 30 8.2 400 0.22 410±40 3.1

Al20399.5991m) 32 8.2 375 0.22 200±20 1.9

MgAl2O2 17 7.2 270 0.26 -200 -1.2

TABLE V.6. STRENGTH CHANGE UNDER NEUTRON IRRADIATION AT 400 - 550°C.

Ultimate bending strength, MPa and Weibull modulus Material unirrad. - 1 dpa 10 dpa 20 dpa 31-41 dpa

A123, 99.5 % A103 208 210 128 8.4 7.7 4.6 BIO. 99.9 % / 322 214 174 182 d=O.< 9.8 6.1 6.2 3..0 AloOo single crystal 372 179 184 8.1 8.4 3.5

AIN.HIP 284 293 155 189 12.2 8.4 14.1 8.8

One of the main problems of ECRH systems for fusion devices is the development of radiofrequency windows to be used as vacuum and tritium barriers inside the wave guides. A very low radiofrequency absorption, very high thermal conductivity and mechanical strength are required for the candidate ceramic materials to be used in these windows, even under neutrons and gamma irradiation. By addition of zirconia particles to the ceramic matrix, the mechanical properties can be improved. However, in general, Zirconia induces a high increase in the dielectric properties.

270 Table V-7 shows the effects of Zirconia doping on the dielectric properties of insulting materials. It is seen, that ths permitivity and loss tangent increase with 'ncreaving concentration of Zirconia. Several inorganic materials, spinel MgAUCh, single- and poly-crystalline SUN4, AlON, and BeO are included in the R&D programme. The current and proposed investigations cover the effects of thermal reactor and charged particle irradiation on their structure and mechanical properties and in- reactor and posf-irradiation electrical properties. The radiation induced conductivity as ^ function of irradiation time and dose rate of 30, 57 and 100 Mrad/h is given in Table V-8. A progamme on the effects of ionizing radiation on the structural stabilities of at cryogenic temperature is also started.

TABLE V.7. DIELECTRIC PROPERTIES OF ZIRCONIA DOPED INSULATING MATERIALS

Sample permitivity loss tangent permitivity loss tangent (19 GHz) 1

REFERENCES [1] R. Heidinger, "Design Parameters of Ceramic Insulator Materials for Fusion Reactors", J. Nucl. Mater., in print. [2] R. Heidinger, "Dielectric Loss of Aluminia between 95 K and 330 K at ECRH Frequencies", J. Nucl. Mater., in print. [3] E.R. Hodgson, Cryst. Latt. Def and Amorph. Mat., 18 (1989) 169. [4] E.R. Hodgson, A. Ibarra, M. Jimenez de Castro, J. Molla, R. Vila, Report EUR-CIEMT 89/12. [5] G.P. Pells and S.N. Buckley, "Radiation Effects in Electrically Insulating Ceramics", Proc. of the Workshop on RAdiation Effects in Ceramic Insulators, 21-22 February 1989, Garching/Germany: [6] C.H. Wu, "Ceramic Materials for the Next European Torus (NET) - Thermonuclear Fusion Reactor. Invited Lecture, Proc. Seventh Cimtec World Ceramics Congress, June 24-30,1990, Montecatini Termo/Italy.

271 TABLE V.8: IRRADIATION INDUCED CONDUCTIVITY

SAMPLE DOSE RATE TIME BREAKDOWN ONSET a initial a final (Mrad h"1) 00 00 a 30 90 ioo ) 2 3

57 46 20 3 57

100 28 9 4.5 53

MgO 57 9 6 3 8 (~ 150ppmFe)

MgO 100 4 3 5.6 9 (~ 150ppmFe)

Mg 200 0.5 ~ 0.1 — ~ 700 (~ 500ppmLi)

4 4 ^3 0 ~ 100 No Breakdown 10- 10-

1 a) Estimated form base current (IQ') increase. All samples at 450°C and approx. 1300 V cm"

91-04789 INTERNATIONAL ATOMIC ENERGY AGENCY VIENNA, 1991