annual report of the association euratom/cea …annual report of the association euratom/cea 2000...

Annual Report of the Association EURATOM/CEA

2000

Compiled by : Ph. MAGAUD and F. Le VAGUERES

ASSOCIATION CEA/EURATOM DSM/DRFC

CEA/CADARACHE 13108 Saint-Paul-Lez-Durance (France)

- I -

CONTENTS

INTRODUCTION ................................................................................................................................. 1

EFDA TECHNOLOGY PROGRAMME ........................................................................ 3

Physics Integration

Diagnostics TW0-DIAG-DEV ITER diagnostic window development ........................................................................ 5

Heating and Current Drive TW0-ICRF/ANT ICRF Antenna and vacuum transmission line development - ICRH Antenna coupling: Near field computations....................................................... 7 CEFDA 00-519 Neutral beam heating design for EDA extension ......................................................... 11 TW0-NB.DEV.1 Neutral beam development for EDA extension - EU-JA collaborative experiment on KAMABOKO source.......................................... 15 TW0-NB.DEV.3 Neutral beam development for EDA extension - 1 MeV SINGAP accelerator ......................................................................................... 19 TW0-NB.DEV.4 Neutral beam development for EDA extension - Manufacture and testing of a prototype 1 MeV bushing for SINGAP ......................... 23

Vessel/In Vessel

Plasma Facing Components CEFDA 99-501 Critical heat flux and thermal fatigue testing of CFC monoblocks - 200 kW electron beam gun test .................................................................................... 27 CNET 98-480 Thermal fatigue testing of divertor full scale prototype - 200 kW electron beam gun test .................................................................................... 29 CNET 98-485 Thermal fatigue testing of baffle full scale prototype - 200 kW electron beam gun test .................................................................................... 35 DV4.3 Optimisation and manufacture of HHF components - Study of flat tile cascade failure possibility for high heat flux components ................. 37 TW0-DV4/01 Optimisation of manufacture of HHF components ...................................................... 41 TW0-DV4/02 Optimisation of Cu-SS-Tube Joints.............................................................................. 43 TW0-T438-01 Development and testing of time resolved erosion detecting techniques ..................... 45

- II -

Vessel/Mechanical structure T216-GB8 Small scale testing of FW/BS modules ........................................................................ 49 TW0-LASER/CUT VV intersector maintenance - Further development of high power Nd YAG - Improvement of YAG laser back-plate cutting using powder injection ....................... 51 TW0-LASER/HYD YAG laser process for cutting and welding of the hydraulic module blanket - CW YAG laser process for cutting and welding of the hydraulic connections ............ 53 TW0-LASER/REWELD VV intersector maintenance - Further development of high power Nd YAG - Laser rewelding after cutting........................................................................................ 55 TW0-LASER/WELD VV intersector joining - Further development of high power Nd YAG - Laser welding with multipass filler wire ...................................................................... 59 TW0-T301/01 Ultrasonic inspection for vessel inter-sector weld - Completion of ultrasonic inspection process development - Optimisation of the ultrasonic inspection method : effect of the wave polarity and frequency on the signal to noise ratio .................................................................... 61 TW0-T420/06 Fabrication of a first wall panel with HIP’ed beryllium armor .................................... 65 TW0-T420/08 Development of HIP fabrication technique - Faisability of using HIP for the manufacturing of Primary First Wall Panel (PFWP) ........................................ 69 TW0-T427/03 Development and characterization of one-step SS/SS and CuCrZr/SS HIP joints including mock-up fabrication............................................................................ 73 TW0-T508/04 Development of Be/CuCrZr HIPping technique .......................................................... 75 TW0-T508/05 Development of Be/CuCrZr brazing techniques .......................................................... 77 TW0-T519/01 Low temperature water radiolysis ................................................................................ 79 V63 Detailed investigation of CuAl25 (IG1) - Its joints with 316LN stainless steel and joints testing procedure.......................................................................................... 81

Assembly & Maintenance T252 Radiation tolerance assessment of standard components for remote handling and process instrumentation .......................................................................... 85 T329-5 In vessel RH dexterous operations - Task extension 2 ................................................. 91 TW0-DTP/1.1 Carrier and bore tools for 4" bend pipes....................................................................... 95 T329-4 TW0-DTP/1.2 Prototypical manipulator for access through IVVS...................................................... 99 TW0-DTP/1.4

Magnet Structure CEFDA 00-517 Design of cryogenic transfer lines and ring manifolds for ITER-FEAT ...................... 103 CEFDA 00-541 PF correction coils : conceptual design and analysis, CS cooling inlets, conductor analysis tools ............................................................................................... 105 M40 Design work on magnet R&D ...................................................................................... 109

- III -

M50 Conductor R&D - Development of NbTi conductors for ITER PF coils ..................... 111 TW0-T400-1 CSMC and TFMC installation and test ........................................................................ 113

Tritium Breeding and Materials

Breeding Blanket Water Cooled Lithium Lead (WCLL) Blanket TTBA-1.1 TBM Adaptation to Next Step...................................................................................... 117 TTBA-2.1 Blanket Manufacturing Techniques - Definition of specifications for demonstrators ............................................................. 121 TTBA-2.2 Blanket Manufacturing Techniques - Solid HIP demonstrator for fabrication and coating, fabrication of double wall tubes ................................................................................... 125 TTBA-2.4 Blanket Manufacturing Techniques - Feasibility of powder HIP fabrication for TBM or TBM components ......................... 129 TTBA-2.5 Blanket Manufacturing Techniques - Double-wall tube out-of-pile testing (DIADEMO experimental program) .................. 133 TTBA-3.1 Coating qualification and irradiation tests - Permeation barrier qualification - Fabrication and characterization of CVD specimens ................................................... 137 TTBA-3.5 Coating qualification and irradiation tests - Permeation out-of-pile testing ................. 141 WPA4.2.3 TTBA-4.1 Processes and components - Tritium extraction from Pb-17Li .................................... 143 TTBA-5.3 Safety and Licensing - Modelling of the lithium/lead water interaction ...................... 147 TTBA-5.4 Safety and Licensing - Development of minor components & instrumentation........... 149 TTBA-6.2 MHD Effects - Test and modeling of natural convection MHD .................................. 153 Helium Cooled Pebble Bed (HCPB) Blanket TTBB-2.3 Blanket Manufacturing Techniques - First wall manufacturing by HIP forming technique .................................................... 157 TTBB-5.D3 Development of ceramic breeder pebble beds - Characterization of Li2TiO3 pebble beds...................................................................... 159 TTBB-5.D4 Development of ceramic breeder pebble beds - Validation of Li2TiO3 fabrication with pre-industrial means of the lab fabrication steps - Mastering/Optimizing ..................................................... 163

Structural Materials development Advanced materials TTMA-1.2 SiC/SiC ceramic composites - SiC/SiC composite development and characterization................................................. 167 TTMA-1.4 SiC/SiC ceramic composites - Compatibility of SiC/SiC composite with liquid Pb-17Li alloy................................... 169

- IV -

TTMA-1.6 SiC/SiC ceramic composites - Brazing of SiC/SiC composites with thermo-mechanical testing................................. 173 Reduced Activation Ferritic Martensitic (RAFM) Steels SM-2-3-1 Mechanical properties of F82H weldments .................................................................. 177 SM-3-3 Corrosion in water conditions - Corrosion water with additives .................................. 181 SM-3-5.1 Compatibility with hydrogen and liquid - SCC behaviour of Eurofer in aqueous environments ................................................... 185 SM-4-3 F82H weldability with filler metal ............................................................................... 187 SM-5-7 Small size specimen technology................................................................................... 191 SM-6-1 Development of high dose high temperature - Low temperature, high dose irradiation test with ISTC................................................ 193 TTMS-1.1 RAFM steels - Irradiation performance - Neutron irradiation of RAFM steels to 30-35 dpa at 325°C......................................... 197 TTMS-2.1 RAFM steels - Metallurgical and mechanical characterization - SM-2-1 Metallurgical characterization of EUROFER 97 steel ................................................. 199 TTMS-2.2 RAFM steels - Metallurgical and mechanical characterization - Mechanical characterization of Eurofer 97 : tensile and impact properties.................. 203 TTMS-2.3.1 RAFM steels - Metallurgical and mechanical characterization - Mechanical properties of diffusion bonded welds (RAFM/RAFM HIP joint) ............. 207 TTMS-2.3.2 RAFM steels - Metallurgical and mechanical characterization - Mechanical properties of weldments ............................................................................ 211 TTMS-2.4 RAFM steels - Metallurgical and mechanical characterization - Mechanical properties of powder HIP Eurofer............................................................. 215 TTMS-4.1 RAFM steels - Qualification fabrication processes - Eurofer powder HIP development and characterization............................................... 217 TTMS-4.2 RAFM steels - Qualification fabrication processes - Mechanical alloying including ODS ............................................................................ 221 TTMS-4.3 RAFM steels - Qualification fabrication processes - RAFM weldability..................... 225 TTMS-4.4 RAFM steels - Qualification fabrication processes - Dissimilar welding RAFM/316LN............................................................................... 229 TTMS-4.5 RAFM steels - Qualification fabrication processes - Solid HIP process and qualification ............................................................................. 231 TTMS-5.1 RAFM steels - Rules for design and inspection - Design code assessment and development ................................................................... 235 TTMS-5.4 RAFM steels - Rules for design and inspection - Data collection and data base maintenance .................................................................. 239 TTMS-6.3.1 RAFM steels - Qualification of high performance steels - Development of RAFM-ODS steels............................................................................. 243 Neutron Source TTMI-001 IFMIF - Accelerator facility ......................................................................................... 245

- V -

Safety and Environment CEFDA 00-539 French contribution to ITER licensing working group................................................. 249 SEA1-14 Assist JCT on safety documentation ............................................................................ 251 SEA4-4 Plant safety assessment - ITER-FEAT design CAT V accident analysis : H2 production via delayed plasma shutdown................................................ 253 SEA5-2 Validation of computer codes and models - ISAS maintenance................................... 255 SEA5-3 Validation of computer codes and models - Thermohydraulical code validate experiments ............................................................. 257 TSW-2.1 Waste and Decommissioning Strategy - Detritiation of structural material and management of tritiated waste.......................... 259 TSW-2.6 Waste and Decommissioning Strategy - Development and assessment of overall strategies....................................................... 263 TW0-SEA3.5 Hydrogen deflagration/detonation analysis .................................................................. 265 TW0-SEA4 Plant safety assessment................................................................................................. 267

System Studies

Power Plant Conceptual Studies (PPCS) TW0-TRP-1.1 Safety and environmental requirements ....................................................................... 271 TW0-TRP-2.4 Economic and operational requirements ...................................................................... 273 TW0-TRP-2.5 Supply requirements..................................................................................................... 277 TW0-TRP-3D4 Reactor integration ....................................................................................................... 279 TW0-TRP-4D5 High heat flux components........................................................................................... 283 TW0-TRP-4D7 Economics and operational analyses - Maintenance schemes...................................... 285

Socio-Economics Studies SERF2 Externalities of fusion - Accident external costs and sensitivity analysis on the site location............................ 289

UNDERLYING TECHNOLOGY PROGRAMME ................................................ 293

Vessel/In Vessel

Plasma Facing Components UT-PFC&C-HFW Transparent polycrystalline ceramics ........................................................................... 295 UT-PFC&C-HIP Mechanical behaviour of HIP joints ............................................................................. 299

- VI -

UT-PFC&C-SiC C-SiC and boron doped graphite .................................................................................. 303 UT-PFC&C-SiC/MJ Development of SiC/metal joining technology ............................................................ 307 UT-SM&C-VVI Design and analysis of vacuum vessel and internals .................................................... 309

Vessel/ Mechanical Structure UT-SM&C-BIM Dissimilar diffusion-bonded joints mechanical testing ................................................ 313

Assembly and Maintenance UT-RH1 Remote Handling Techniques - Technology and control for remote handling systems.................................................. 317 UT-RH2 Remote Handling Techniques - Graphical programming for remote handling ............................................................... 319 UT-RH3 Remote Handling Techniques - Advanced technologies for high performances actuators ............................................. 323 UT-RH4 Remote Handling Techniques - Radiation tolerance assessment of electronic components from specific industrial technologies for remote handling and process instrumentation.................... 327

Tritium Breeding and Materials

Breeding Blanket UT-SM&C-BLK Development of breeding blankets ............................................................................... 333 UT-SM&C-COAT Development of an oxide layer on Fe-Al coatings ....................................................... 337 UT-SM&C-LM/DC Diffusion coefficient measurements in liquid metals ................................................... 341 UT-SM&C-LM/MAG Liquid metal corrosion under magnetic field................................................................ 345 UT-SM&C-LM/Refrac Compatibility of refractory materials with liquid alloys .............................................. 349 UT-SM&C-LM/WET Wetting of materials by liquid metals........................................................................... 353

Materials Development Structural Materials UT-SM&C-LAM/Mic Influence of the martensite morphology on the plasticity behaviour of the Eurofer steel ....................................................................................................... 357 UT-SM&C-LAM/Weld Laser weldability of LAM steels (Eurofer97) .............................................................. 361 UT-SM&C-LAM2 Irradiated behaviour of Reduced Activation (RA) martensitic steels after neutron irradiation at 325°C................................................................................. 363 UT-SM&C-LAM3 Microstructural investigation of Reduced Activation Ferritic-Martensitic (RAFM) steels by Small Angle Neutron Scattering (SANS) ....................................... 367 UT-SM&C-ODS Development of forming and joining technologies for ODS steels .............................. 371

- VII -

Nuclear Data UT-N-NDA Nuclear data assessment ............................................................................................... 375

Fuel Cycle UT-T1 Separation of the D/T mixture from helium in fusion reactors using superpermeable membranes - Development of superpermeable membranes suitable to operate under ion bombardment .............................................. 379

Safety and Environment UT-S-BLK Blanket Safety .............................................................................................................. 383 UT-S-MITIG Evaluation and mitigation of the risk connected with air or water ingress................... 385

System Studies UT-SM&C-REL Reliability modeling & assessment .............................................................................. 389 UT-SM&C-TDC Innovative Thermo-Dynamic Cycles studies (TDC) .................................................... 393 UT-SM&C-TMNC Thermal-hydraulic models for nuclear components ..................................................... 397

INERTIAL CONFINEMENT FUSION PROGRAMME .............................. 399

ICF01 Intense laser and particle beams dynamics for ICF applications .................................. 401 ICF02 Civil applications of inertial confinement fusion - Cryogenic targets production using magnetic levitation .............................................. 405 ICF03 Laser-matter interaction at relativistic intensities and fast igniter studies .................... 411 ICF-KiT-PRC Overview on power reactor concepts ........................................................................... 415

APPENDIX 1 : Directions contribution to the fusion programme ........................ 419

APPENDIX 2 : Allocations of tasks .......................................................................................... 423

APPENDIX 3 : Reports and publications .............................................................................. 429

APPENDIX 4 : CEA tasks in alphabetical order .............................................................. 441

APPENDIX 5 : CEA sites ................................................................................................................. 445

- 27 - EFDA technology / Vessel/In Vessel / Plasma facing components

CEFDA 99-501 Task Title : CRITICAL HEAT FLUX AND THERMAL FATIGUE TESTING OF

CFC MONOBLOCKS 200 kW electron beam gun test INTRODUCTION The vertical target (VT)of the ITER divertor consists in a assembly of a number of units in the toroidal direction (indicatively 1200). Each unit is composed with a straight CFC monoblock in the lower part and a bend tungsten armoured upper part. The peaked surface heat flux in the CFC monoblock part can reach very high values – up to 20 MW/m² and margins to critical heat flux (CHF) are a crucial point of the design. In a other hand, from an economical point of view, increasing the vertical target unit width offers substantial cost savings. Scope of this contract is to investigate the possibility to increase the VT unit width from the present 23 mm to the highest possible value in order to decrease the divertor cost keeping a margin to CHF occurrence. 2000 ACTIVITIES Initially critical heat flux testing of ten CFC monoblocks were foreseen in 2000 (in high heat flux facility FE200) but due to delays in the manufacturing of the components, the testing was delayed to year 2001. As the manufacturing of the components to be tested are not included in the contract activities, no activity in the frame of this contract was performed this year.

TASK LEADER F. ESCOURBIAC DSM/DRFC/SIPP CEA Cadarache 13108 St Paul Lez Durance Cedex Tél. : 33 4 42 25 44 00 Fax : 33 4 42 25 49 90 E-mail : [email protected]


CNET 98-480 Task Title : THERMAL FATIGUE TESTING OF DIVERTOR FULL SCALE

PROTOTYPE 200 kW electron beam gun test INTRODUCTION Thermal fatigue seems to be one of the most important damaging mechanism for the ITER divertor plasma facing components. Therefore, an assessment of the behaviour of the baffle component under cycling heat loads is essential to demonstrate the validity of the selected design solution. A serial of three straight and one curved samples was manufactured by Plansee in 1999 and delivered to CEA for infrared characterisation and high heat flux testing in FE200 facility. This report presents the main results obtained during the testing campaigns of the samples. 2000 ACTIVITIES STRAIGHT BAFFLE SAMPLES The 3 samples are monoblock-type, new generation : internal tube is made of CuCrZr and HIPped (High Isostatic Pressure) with OFHC layer previously AMC® on CFC NB31.

Monoblocks tiles are attached on a rear stainless steel part : BS1 and BS2 are “slotted” (CEA concept) whereas BS5 is “dovetail” attached (cf. Fig. 1). Each of them is equipped with a swirled insert into the cooling channel. Plasma facing surface dimensions are ~70 x 25 mm² (4 tiles of 18 x 25 mm²). Thermographic examination (SATIR facility) [1] Thermographic examinations were performed in Cadarache (France) with SATIR experimental device. Aim of this facility, mainly composed of two water circuits at different temperature levels and an infrared camera is to detect plasma facing components internal defects. Detection is based on the monitoring of the surface temperature during a transient period. A view of the temperature field before FE200 testing is shown Fig.2. A slower surface temperature response is interpreted as an higher thermal resistance, i.e. bad joining between the different layers of materials (defects = lack of brazing on HIP or AMC interfaces).

Figure 1 : View of slotted and “dovetail” monoblock types

Figure 2 : SATIR view of the samples during transient thermal shock

BS1 and BS2 BS5

BS2 BS1 BS5

0°C 100°C

T4 T3 T2 T1 T4 T3 T2 T1 T4 T3 T2 T1


The most important defect is located on BS1, Tile 4 lower part. Our usual numerical treatment gives for this one a Dtref value of 12°C [1;2]. Such a high value should correspond to a debonding of more than 50%. Edges of Tile 3/ Tile 2 of BS1 and Tile 3 of BS2, are characterised by values of Dtref from 7 to 10 °C. BS5 is more healthy, only the Tile 2 presents smaller symmetrical defect (Dtref < 3°C). Note that for Tore Supra toroidal pumped limiter fingers -CFC N11 EB welded on CuCrZr heat sink with AMC® interlayer, flat tile technology, more than 600 components - control quality acceptance criterion (DTref) is fixed at 3°C, corresponding to detection of a 4 mm width stripe defect at AMC® interface (with flat tile technology) [3]. As a conclusion, bad thermal behaviour of BS1 and BS2 -the two slotted monoblocks- are clearly detected whereas BS5 examination (dove-tail monoblock) is relatively satisfactory. High heat flux testing (FE200 facility) [4;5;6] Fatigue testing campaign was performed in the FE200 European facility situated in Le Creusot (France).

Screening at 11 MW/m² The first step of the campaign was devoted to the infrared monitoring of the component under an incident heat flux of 11 MW/m². During such a screening, EB gun shot on the sample up to thermal steady-state, attained after around ten seconds. At steady-state, an heterogeneous heat flux pattern was observed and interpreted as an image of thermal resistance from the surface to the cooling channel (see Fig.3). On Fig.3, temperature distribution on heated surfaces (view from the gun) are proposed after tentative correction of the very high measured temperature. Anyway, infrared device is not calibrated out of the temperature window 500-1500°C and one should consider higher measured values of 1500°C only as indicative data. Correlation with thermographic examination is remarkable : temperature on hot spot on BS1, Tile 4 is up to 2200°C (Dtref = 12°C), and from 1900 to 2000°C for edges of Tile 3/ Tile 2 of BS1 and Tile 3 of BS2 (Dtref from 7 to 10°C). Maximum temperature value for BS5 does not exceed 1700°C (Dtref < 3°C), to be compared with 1300°C in case of healthy tile.

(Different Infrared scale for the 3 samples due to differences on emissivity at such high temperatures)

Figure 3 : FE200 view of the samples under steady state incident heat flux of 11 MW/m² (Tin 100°C, 12 m/s, 33 bar)

Fatigue cycling A first step of 100 cycles at 5 MW/m² followed by a second one of 1000 cycles at 11 MW/m² (~20 kW removed in the water) did not point out strong evolution of the defect-free areas (cf. Table 1 : TC4x are located on Tile 4, sample x, 3mm deepness, upper part , TC1x idem on tile 1; Pyr2 is centred on Tile 1 for each sample). As far as hot spots are concerned, temperatures measured with infrared camera are out of calibration window (Tsurf > 2000°C for a calibration window of 500-1500°C).

We know that response of the camera is strongly non linear in this range of temperature, moreover, emissivities of the samples also changed during cycling and it is impossible to give absolute values. Nevertheless, global behaviour of the components did not change during this cycling phase : pattern of the temperature was kept unchanged. A third fatigue step of 1000 cycles was performed on BS5 at 15 MW/m² removed into the water but only on tiles 3 and 4 (up to 15 kW into the water). Other samples (and other tiles of BS5) were too overheated for cycling: during few tentative shots, electron beam gun was interrupted due to the vacuum level degradation (material vaporisation).

Table 1 : Thermocouples and pyrometer temperatures before and after fatigue cycling

BS1 BS2 BS5

T41 (°C) TC11(°C) Tpyr2 TC42 (°C) TC12 (°C) Tpyr2 TC45 (°C) TC15 (°C) Tpyr2

Before 880 1040 1300 970 1050 1240 1000 920 1250

After 960 1300 1000 1240 1000 1380

BS2 BS1 BS5 1400°C 2200°C 1300°C 2000°C 1300°C 1700°C

T4 T3 T2 T1 T4 T3 T2 T1 T4 T3 T2 T1


No evolution was observed during this step as it is shown in Fig.4.

Before After

Figure 4 : Zoom on Tile 3 and 4 of BS5 under 15 MW/m² before and after 1000 cycles at 15 MW/m²

LOFA (loss of flow accident) Objective of this last step was to determine critical heat flux (CHF) or maximum allowable heat flux (limitation due to facility limits: vacuum level, etc…) by decreasing the cooling flow rate under given incident heat flux (either by increasing the inlet water temperature), loop pressure being constant. 1) Heat flux was adjusted at 15 MW/m² on Tile 4 and 5 of

BS5 at 100°C of inlet temperature. Inlet velocity was decreased from 12 m/s down to 3 m/s without CHF occurrence. Nevertheless, it was observed an increasing of surface temperature measured by the pyrometer from 1740°C to 1820°C.

2) The same scenario was performed two times from

12 m/s down to 6 m/s respectively at 140°C and 176°C inlet temperature without CHF occurrence (+30°C increasing of surface temperature).

CURVED BAFFLE SAMPLE This curved sample (~200 mm length, cf. Fig. 5) was also produced by Plansee company.

The particularities of this sample are : 1) the design, allowing any shape for the Baffle, 2) the “cold” HIPping process with optimised parameters

avoiding high concentration of stresses during HIPping cycle (100 MPa; 550°C, 6h).

Figure 5 : View of Curved Baffle sample Both results on thermographic examination and fatigue testing are remarkable : the mock-up sustained 1000 cycles at maximum removed flux in the water of 9 MW/m² plus 1000 cycles at maximum removed flux in the water of 13 MW/m² without any noticeable evolution of surface temperature. Thermographic examination (SATIR facility) [7] The symmetry of the front view during the transient thermal shock (after fatigue testing) is remarkable and points out the quality of the bonding. Two lateral view are proposed and also demonstrate the good bonding. None of the DTref measured on the front view are higher than 3.6°C. High heat flux testing (FE200 facility) [8;9] 3 steps of fatigue cycling were performed in between initial and final screenings at 6 MW/m²: 1) 100 cycles at maximum removed heat flux evaluated at

6 MW/m² 2) 1000 cycles at maximum removed heat flux evaluated at

9 MW/m² 3) 1000 cycles at maximum removed heat flux evaluated at

13 MW/m²

Figure 6 : SATIR view of the sample during transient thermal shock (front view and lateral views)

T4 T3

1900°C 1000°C

T4 T3

0°C 100°C


Figure 7 : FE200 view of the sample under steady state max. removed heat flux of 6 MW/m² (Tin 40°C, 12 m/s, 34 bar)

Figure 8 : FE200 view of the sample under steady state max. incident heat flux of 13 MW/m² (Tin 40°C, 12 m/s, 34 bar) and surface temperature versus maximum removed heat flux into the water

The maximum surface temperature increased between the initial and final screenings from 780°C to 830°C which is quite negligible with regards to infrared measurements uncertainties. A view of the surface temperature at the end of third step of fatigue is proposed Fig. 8. The value of 13 MW/m² was chosen taking account the surface temperature (max allowable ~2000°C for cycling). The pyrometer surface temperature versus incident heat flux is shown on Fig. 8 and did not evolved during the cycling steps. The time response of the hottest point (1900°C) did not evolved during the steps and was kept constant at 3 s (normalised time response at 63%). CONCLUSION 4 samples of monoblock technology representative of the design of the ITER divertor baffle were infrared characterised and fatigue tested. The 3 straight ones presented large bonding defect detected both with the infrared non destructive method and during the fatigue testing. Nevertheless, two healthy tiles of the “dovetail” monoblock sustained 1000 cycles at 15 MW/m² without clear appearance of bonding defect. A LOFA test shown a good behaviour of the component even with an internal velocity as low as 3 m/s and an inlet temperature upper than 170°C, under a removed heat flux of 15 MW/m².

The curved one, manufactured after optimisation of “cold” HIPing cycle (100 MPa, 550°C, 6 hours), is remarkable and did not present any defect, even after 1000 cycles at 13 MW/m² (2000°C at the surface). This last design is promising, mock-ups with larger dimensions should be manufactured in the future. REFERENCES

[1] Tests SATIR sur 3 maquettes BS1, BS2 et BS5 CEA – V.Paulus, PCQ99.036, Nov.99. [2] A. Durocher et al. Interface quality control by infrared thermography

measurement Proc. On 15th World Conf. on NDT - Roma (Italy),

2000, to be published [3] A. Durocher- Proc. of 20th SOFT, Marseille, pp 121-

124, 1998 [4] Technical specifications of ITER BAFFLE SAMPLE

testing in FE200 facility CEA – P.Chappuis, CFP/NTT-1999.011, Nov. 99 [5] High Heat Flux Fatigue Tests of ITER MOCK-UP

BAFFLE SAMPLE (CEA71) FRAMATOME – I.Vastra-Bobin, MC/TS R 99.856

A, July 2000

540°C 830°C

780°C 830°C

1500°C 2000°C 0

500

1000

1500

2000

2500

0 5 10 15

Max. Removed Heat Flux (MW/m²)

Tsu

rf (

°C)


[6] Review of mock-up « ITER BAFFLE SAMPLE » testing (partial)

CEA – P.Chappuis – CFP/NTT-2000.003, January 2000

[7] Tests SATIR sur Curved CFC avant et après FE200 CEA – V.Paulus, RCQ2001-02, Febru ary 2001 [8] Technical specifications of Curved CFC testing in

FE200 facility CEA – F.Escourbiac CFP/NTT-2000.029, Dec. 2000 [9] Screening and Fatigue Tests of Curved CFC MOCK-

UP FRAMATOME – M.Febvre – MC/TS R 01.860 A,

January 2001

TASK LEADER

F. ESCOURBIAC DSM/DRFC/SIPP CEA Cadarache 13108 St Paul Lez Durance Cedex Tél. : 33 4 42 25 44 00 Fax. : 33 4 42 25 49 90 E-mail : [email protected]


CNET 98-485 Task Title : THERMAL FATIGUE TESTING OF A BAFFLE FULL SCALE

PROTOTYPE 200 kW electron beam gun test INTRODUCTION Thermal fatigue seems to be one of the most important damaging mechanism for the ITER divertor. Therefore, an assessment of the behaviour of the baffle component under cycling heat loads is essential to demonstrate the validity of the selected design solution. Scope of the contract is to perform thermal fatigue testing of a baffle full scale prototype in the high heat flux facility FE200. 2000 ACTIVITIES Initially thermal fatigue testing were foreseen in 2000 but due to delays in the manufacturing of the component, the testing was delayed to year 2001. As the manufacturing of the component to be tested is not included in the contract activities, no activity in the frame of this contract was performed this year.

TASK LEADER F. ESCOURBIAC DSM/DRFC/SIPP CEA Cadarache 13108 St Paul Lez Durance Cedex Tél. : 33 4 42 25 44 00 Fax. : 33 4 42 25 49 90 E-mail : [email protected]


DV4.3 Task Title : OPTIMISATION AND MANUFACTURE OF HHF COMPONENTS Study of flat tile cascade failure possibility for high heat flux

components INTRODUCTION The object of this task is to evaluate the possibility of failure in cascade of flat tiles under convective heat flux with a glancing incidence. Calculations and tests on the high heat flux (HHF) facility FE200 are foreseen. Preliminary works in 1999 consisted in the definition of the geometry, the study of the feasibility of the tests by finite elements calculations and the definition of foreseen tests and mock-ups to be prepared. In 2000 the fabrication of mock-ups were launched in Plansee company, used NS31 (2 mock-ups) and W (2 mock-ups) as armour material. The delivery of these mock-ups is foreseen by end of 2001. A meeting at Cadarache reminded the advantages of the concept (flat tiles with HV cooling) but also the issues to investigate. Review of the calculations showed clearly that a chamfering of the tile will be necessary for tests on the FE200. 2000 ACTIVITIES DEFINITION OF THE OBJECTIVES AT THE MAY 3rd, 2000 MEETING The flat tile with hypervapotron (HV) cooling has the following advantages: - It is cheaper than the corresponding monoblock

geometry by about 25% (excluding the armour cost). - The armour cost is reduced by about 50%. - The width of the VT unit can be increased appreciably. - The surface temperature is more uniform. - It has a higher critical heat flux limit and a lower

pumping power with respect to the “swirl tube” with a twisted tape having a twist ratio of 2.

- A curved component can be manufactured without

adding significant manufacturing issues. - The thickness of the cooling tubes is not intrinsically

limited to 1 mm as for the monoblock geometry. It is worth noting that the cooling tubes represent the only boundary between the pressurised water coolant and the plasma vacuum chambers.

- It is a more mature technology and is already foreseen in existing tokamaks.

However the following issues needs to be investigated: - It generates significantly higher cyclic thermal stresses

than a monoblock geometry; therefore the thermal fatigue lifetime is lower (19 MW/m2 x 1000 cycles vs. 24 MW/m2 x 1000 cycle for a flat tile and monoblock geometry, respectively). This is particularly important in the bottom part of the VT where the highest heat flux will occur.

- In the case of a convective heat flux, as for the lower

part of the VT, one should also take into account that if one tile falls off, the adjacent tile receives an extremely high heat flux localised on its edge as a consequence of the glancing incidence. As a result a rapid temperature rise occurs in the armour - heat sink joint which might seriously damage the joint. The heat flux to the coolant also increases significantly thus causing possible critical heat flux problems.

This last point is the objective of this task. REMIND OF TESTS DEFINITION 1. The following input parameters will be assumed for the

experimental tests: - Heat flux normal to the heated surface : 10 MW/m2. - Angle of incidence of the heat flux : 3°. - Coolant conditions : 100 °C, 3.5 MPa, 12 m/s.

2. Two mock-ups will be manufactured with a CFC (SEP

NS31) armour (tiles 20 x 27 x 6 mm) and two ones with a W armour (tiles 6 x 6 x 6 mm). The gap between each tile will be ~1 mm.

3. One mock-up with a CFC armour and one with a W

armour will be used for the thermal fatigue test (1000 cycles at 20 and 15 MW/m2 for the CFC and W armour, respectively).

4. One mock-up with a CFC armour and one with a W

armour will be used for the cascade failure test. Two CFC tiles, located 100 mm from each end of the

mock-up, will have a reduced thickness (5 mm instead of 6 mm) to simulate the loss of a tile. To prevent excessive CFC sublimation, the leading edge of one of the adjacent tiles will be chamfered.


A few W tiles, located 100 mm from each end of the mock-up, will have a reduced thickness (5 mm instead of 6 mm) to simulate the loss of a tile.

5. The study on the possible cascade failure will be

performed in two ways: (a) applying a heat flux normal to the surface which simulates the computed heat flux through the pure Cu interlayer, (b) applying a heat flux with a glancing incidence of 3°.

6. The mock-ups will be instrumented with 6

thermocouples each, 3 located in the armour and 3 in the heat sink.

Fig2a shows a possible heat flux deposition on the FE2000, 90MW/m2 are put on 1.6mm whereas 10Mw/m2 are put on the rest of the surface. This heat flux deposition gives about the same heat transfer at the CFC/Cu joint than a glancing incidence. However this heat flux deposition will be not possible on the FE200 facility, since the surface temperature would be too high. One can notice that the CFC/Cu joint reaches the temperature of 850°C, which is too high (maximum acceptable 550 to 650°C). Fig2b shows a heat flux deposition with a glancing incidence when using a shaping of 2 mm. The surface temperature is rather acceptable; the temperature at the CFC/Cu bond is still very high but under the limits on the edges of the tile.

FINITE ELEMENTS CALCULATION RESULTS

Figure 1 : Hypothesis of the calculation

Figure 2 : Possible heat load versus tile shaping and corresponding temperature isovalues MOCK-UP FABRICATION

Figure 3 : Hypervapotron mock-up with CFC tiles

NS31 tile

Cu layer (OFHC)

CuCrZr heat sink

Heat flux

27 mm2 mm

10MW/m2 10MW/m2

90MW/m2 38MW/m2

300°C

600°C

900°C

2100°C

390°C

855°C 1320 3180°C 2 mm

R =1.6mm


Figure 4 : Hypervapotron mock-up with tungsten tiles

Figure 5 : Hypervapotron mock-up with tungsten tiles Fig. 3 and 4 show the mock-up geometries with an armour length of about 500 mm and a width of 27 mm. The bond between AMC tiles and CuCrZr heat sink is made by electron beam weld. One extra mock-up with a bond obtained by high isopressure technique (HIP) is also expected. Fig. 5 shows the hypervapotron cooling system, about 700 mm long, with fins 3 mm thick, 4 mm high and a gap between fins of 3 mm. CONCLUSION In case of the falling of a tile, one can expect a plasma shaping of the adjacent tile so that the surface temperature is below 2200°C and a stopping of the erosion. But the probability of a failure of the second tile before the end of shaping is very high.

Even if it will be difficult to reproduce the exact condition on the FE200 this will be investigated with this task. The mock-ups are now in fabrication and the test are foreseen for the end of 2001 or beginning of 2002. REFERENCES [1] Task DV4.3 Intermediate report 2 (correction),

CFP/NTT – 1999.014, N. Curt, J. Schlosser, 11/02/2000.

[2] L-5 Progress Meeting Summary, Naka, Japan, 5-7

April 2000, Task 435, R. Tivey, 12/05/2000. [3] Summary report of the meeting held at CEA

Cadarache on 3 May 2000, M. Merola, EFDA CSU Garching (Plasma Facing Components), 9/05/2000.


TASK LEADER J. SCHLOSSER DSM/DRFC/SIPP CEA Cadarache 13108 St Paul Lez Durance Cedex Tél. : 33 4 42 25 25 44 Fax : 33 4 42 25 49 90 E-mail : [email protected]


TW0-DV4/01 Task Title : OPTIMISATION OF MANUFACTURE OF HHF COMPONENTS INTRODUCTION High heat flux components for the upper vertical target of the ITER divertor are made of tungsten tiles and copper alloy heat sink. The fabrication of these so-called tungsten monoblocks involves the insertion and joining of Cu-OFHC and CuCrZr tubes in tungsten alloy tiles. Among other processes, HIP diffusion welding is a promising way. In the past, mock-ups have been fabricated with limited success [1], [2]. The objectives of this work are to improve the process in terms of performance, to simplify its implementation and to test its applicability to curved monoblocks. 2000 ACTIVITIES MOCK UPS EXPERTISE Mock ups fabricated in early 1999 failed in thermal fatigue testing at rather moderate incident power density. Finite Element calculation and metallographic observations lead to the conclusion that the Cu-OFHC/CuCrZr interface was too weak to withstand the stresses due to thermal expansion mismatch between tungsten and copper. The location of hot spots during high heat flux testing correlate well to the FEM calculation : tube to tube decohesion occurs at the ends of the component, where the plastic deformation on cooling is the highest.

Figure 1 : Scheme of the tungsten monoblock design IMPROVEMENT OF Cu/CuCrZr JOINTS Cu/CuCrZr joining is made after Cu/W joining. The process temperature must be kept low enough to maintain acceptable mechanical properties of the CuCrZr. HIP experiments have been made at (500°C, 100MPa, 2h) using water quenched CuCrZr alloy as starting material (ageing of the material is made during joining). Various interlayer have been tried. The joints were tensile tested using axisymmetric notched specimens with the notch in the bonding area to force the rupture to occur in the joint (figure 2).

Figure 2 : Axisymmetric tensile specimen used in this study The results are shown on figure 3 and table 1. The best results were obtained with a Ni electroplated layer deposited on CuCrZr. Actually, with this specimen, rupture did not occur at the joint in spite of the notch. Deformation took place mainly on the Cu OFHC side (see figure 2, CuCrZr on the right hand side). All other specimens failed at the joint at less than about 200MPa, after only little deformation.

0

2

4

6

8

10

0 1 2 3displacement (mm)

forc

e (k

N)

Ni+Sn (EP)

Au (EP)

Ni-P

Ni foil

Ni (EP)

Figure 3 : Force-displacement curves for CuCrZr/Cu joints (EP = electroplated)

Table 1 : Summary of CuCrZr/Cu tensile strength

at room temperature. Monolithic CuCrZr is included for comparison

Sample Maximum strength

No interlayer Not measurable Ni foil 168 MPa Electroplated Ni 323 MPa Electroplated Ni + Sn 134 MPa Ni-P 202 MPa Au 170 MPa Monolithic CuCrZr 536 MPa

SIMPLIFICATION OF THE CANISTER DESIGN (BRAZING OF TUNGSTEN TILES) To get tube expansion during HIP, the tiles and tubes may be put in a tight canister. The canister must be removed after HIP, which is difficult and costly. In view of simplifying the process, it was tried to achieve tightness between tungsten tiles by brazing which would suppress the need for canister.

joint

Cu-OFHC tube

CuCrZr tube W-1%La2O3 tile

12 mm


The brazing process should allow maintaining a 0.5mm gap between the tiles (Ni or Cu spacers were used) and should provide end plugs compatible with CuCrZr welding (to achieve tightness at both ends). Unfortunately, common braze alloys for tungsten are usually silver-based alloys or high melting point pure metals and alloys. Work was thus mainly focused on the Cu-Ni-Ti system because is offers the possibility to carry out brazing at about 900-1050°C, which is compatible with the use of Cu spacers and potentially compatible with the subsequent HIP diffusion welding of the Cu tube at 950°C. Tiles were brazed with various materials and various conditions. Brazed joints possess a combination of defects (cracks and pores) as well as various microstructural features that depend of the braze composition : W dissolution, Ti penetration in W, intermetallic particles or layers are observed. Best results in terms of strength were obtained with Ni spacers associated with Ti braze. An example is given on figure 4. A crack in tungsten is visible.

Figure 4 : W/Ni/W joint brazed at 1050°C with 25 µm Ti + 10 µm Cu

Tightness measurement and shear testing were used to assess the brazed joints quality. Best results in terms of tightness were obtained with diffusion bonded and Ti-brazed copper spacers, but the results were not reproducible despite care was taken to get parallel joints by the application of a load during brazing. The use of a tight braze was thus given up. MOCK UP FABRICATION As tungsten brazing was not successful, a new canister system was designed. It includes a copper envelope and removable graphite tools. This technique has been estimated through the fabrication of a dummy mock up with molybdenum tiles. The HIP cycle was (950°C, 100MPa, 2h) for Mo/Cu joining and (500°C, 100MPa, 2h) for Cu/CuCrZr joining. The graphite tools were removed without damaging the mock up. A metallographic view of the CuCrZr/Cu joint is given on figure 5. Three tungsten straight mock-ups have been fabricated whereas three curved ones are under fabrication. Mock-ups will be controlled by NDT and tested under high heat flux.

Figure 5 : Metallographic view of the CuCrZr/Cu joint CONCLUSIONS The mechanical properties of the Cu/CuCrZr joints, which proved to be the weakest point of the high heat flux component mock-ups, were improved. Attempts to simplify the process thanks to brazing were not successful. However, the use of graphite tools and copper envelope has shown to be efficient. Straight mock-ups have been fabricated and curved ones are under fabrication. REFERENCES [1] E. Rigal, ITER Task DV4, NT DEM n°69/98. [2] M Rödig, interim report ITER task GB8-DV6, IWV2-

TN-5/99. REPORTS AND PUBLICATION [1] E. Rigal, interim report, task TW0-DV4/01, NT DEM

88/2000. TASK LEADER Emmanuel RIGAL DRT/DTEN/SMP/LS2M CEA Grenoble 17, rue des Martyrs 38054 Grenoble Cedex 9 Tél. : 33 4 76 88 97 22 Fax : 33 4 76 88 54 79 E-mail : [email protected]

400 µm

10 µm

CuCrZr

Cu OFHC

Ni


TW0-DV4/02 Task Title : OPTIMISATION OF Cu-SS-TUBE JOINTS INTRODUCTION During the ITER EDA there was considerable success in constructing full-scale vertical target mock-ups. However, driven by the requirements of RTO/RC-ITER, this task aims to investigate the possibility of developing cost effective alternatives to the reference design. The objective of this task is the further development and qualification of a bonding technique for CuCrZr to stainless steel tube to tube connection. Such junction is currently achieved using friction or welding techniques. In order to overcome drawbacks associated with Cu melting and to provide a reliable and tight joint, a diffusion bonding technique is proposed. A second possible route, namely brazing, has also been investigated. The properties of the junctions obtained with these two techniques have been addressed in terms of helium leak test, tensile test and metallographical examination. 2000 ACTIVITIES DIFFUSION BONDING ROUTE

Mo ring

SS tube

CuCrZr tube

SS ring

Figure 1 : Diffusion-bonding design process The proposed process for the diffusion bonding route is based on the assumption that the thermal expansion of the tubes provides the necessary load to join the surfaces to be bonded. The tubes can thus be simply introduced in a furnace, with no need of a container as in HIP assisted diffusion bonding. The diffusion-bonding process proposed is displayed in figure 1. According to the ITER data base, the thermal expansion coefficients of the two materials are very similar.

To reach a sufficient load in order to impose contact between the surfaces, a Molybdenum ring is added around the tubes. The low value of Mo thermal expansion coefficient induces a constraint of the tubes during heating. Moreover, in order to avoid diffusion bonding between the tubes and the Mo piece, a TiN layer has been deposited by CVD on the ring. BRAZE ROUTE Not to be disregarded in the ITER context, a braze should contain no Ag, Cd or P nor any elements that could corrode the only 1 mm thick CuCrZr tube. It must also be compatible with the water coolant (100-150°C, 4 MPa, 12m s-1). Thus, 2 commercial brazes have been identified NiCuMan37 and NiCuMan23. A preliminary study allowed to choose NiCuMan37 as a potential braze. EXPERIMENTAL RESULTS Tube/tube connections have been performed under several conditions as presented in table 1. Note that two trial tests without Mo ring at 1040°C and two other junctions with Mo ring, one at 1020°C and one at 1040°C, did not succeed the helium leak test. The results presented below only concern the junction that showed no helium leak.

Table 1 : Tensile properties of the CuCrZr/316 tube to tube connections

Specimen

n° Bonding Max load

(N) Max stress

(MPa) Rupture location

1 1020°C/2h/Argon + Mo ring

2970 86 joint


8467 245 gudgeon


7499 217 joint

3 1020°C/2h/Vacuum + Mo ring

6981 202 joint

4 NiCuMan37 8052 233 gudgeon

5 NiCuMan37 10748 311 head


7465 216 joint

The tensile properties at room temperature for these tube/tube junctions are given in figure 2. In some cases rupture occurred near the grips. Thus the results have to be taken with care and analysed mainly qualitatively. The main conclusions obtained from these tests are the following ones: - the influence of environment (vacuum versus argon) is

low, - the Mo ring is necessary since the coefficients of

thermal expansion of both materials are similar,


- scatter is observed in the tensile results at 1020°C and 1040°C : the process has to be improved to be reproducible,

- the 1040°C temperature does not improve the tensile

properties. The two brazed tube/tube connections show the best properties since it has not been possible to open the tube/tube joint.

Figure 2 : Tube-tube tensile results (overall load versus time)

Typical metallurgical observations of the junction after tensile test are displayed in figure 3 and 4. Although rupture occurs near the grips for specimen n°2, one can see that damage was growing through the joint interface.

Figure 3 : Specimen n°2 :316/CuCrZr interfaces after tensile tests – rupture near the gudgeons (HIP : 1020°C/2h) By contrast, the aspect of the joint of a brazed tube/tube connection does not appear damaged. The voids observed in figure 4 are due to the brazing process and not to the tensile loading.

Figure 4 : Specimen n°5 : NiCuMan37 braze – no rupture at 311 MPa

CONCLUSIONS Two routes to perform tube/tube CuCrZr/316LN junctions have been studied : a diffusion-bonding one and a brazed one. The performed junctions have been tested by means of helium leak test as well as tensile test at room temperature. The diffusion-bonding route is performed in a furnace. Since the two materials in concerned have similar coefficient of thermal expansion, it is necessary to add a Mo ring (low thermal expansion) surrounding the tubes during the heat treatment. Some joints have a good mechanical strength although lower than that of the weakest materials. Improvements can still be added to the process. However, the final grain size of the CuCrZr should be carefully analysed since a too low number of grains through the tube thickness could lead to leaks. The brazed junctions have the best properties. REPORTS AND PUBLICATIONS L. Briottet, E. Rigal, I. Chu, B. Riccetti, EFDA TW0 DV4-02 optimisation of Cu-SS tube joint, Note Technique DEM n°2000/113. TASK LEADER Laurent BRIOTTET DRT/DTEN/SMP/LS2M CEA Grenoble 17, rue des Martyrs 38054 Grenoble Cedex 9 Tél. : 33 4 76 88 33 15 Fax : 33 4 76 88 58 91 E-mail : [email protected]

CuCrZr

316L

voids

CuCrZr

316L

NiCuMan3

Brazed junctions:

No rupture in the


TW0-T438-01 Task Title : DEVELOPMENT AND TESTING OF TIME RESOLVED EROSION

DETECTING TECHNIQUES INTRODUCTION Carbon based material is widely used as plasma facing component in present fusion device due to its good thermophysical properties. This is also the material which has been retained for the ITER divertor. Nevertheless, physical and chemical sputtering yield of carbon are important and induce high erosion rate. The large carbon source reacts with the plasma and create a very complex Plasma Wall Interaction physic. One of this phenomena, known as redeposition occurs when carbon atoms or ion return to the wall : Due to the reactivity with hydrogen, carbon layers are built up with a large hydrogen content. In the case of ITER, the tritium retention in these carbon redeposited layers may limit the operation for safety reason. Although erosion and redeposition of carbon may reduce the life time of the divertor tiles and reduce the operation efficiency, only basic measurements have been undertaken in present tokamak and none of them can provide in situ in a fusion device a time resolved erosion/redeposition measurement. Thus, we propose to develop a technique aiming to measure the erosion and the redeposition process on the limitor in Tore Supra. 2000 ACTIVITIES In the framework of the CIEL program, it is planed to obtain in Tore Supra high performances long time discharges (up to 1000s). On such a duration, erosion of plasma facing components may become very significant. Taking into account the plasma parameter at the last closed surface (LCF) in a CIEL discharge (flux = 1019 cm-2s-1, Te = 100 eV) an average erosion rate of 10-2 has retained due to physical sputtering yield only. The limitor being actively cooled, the surface temperature should be kept below 1000°C and Radiation Enhanced Sublimation (RES) should not occur. It is anticipated that due to the high flux and high electron temperature, the chemical sputtering yield on the limitor, will be negligible. For a 1000 seconds discharge, the resulting gross erosion would be of about 1020C per cm2 or 10 µm. From previous measurements on actively cooled carbon limitor on inner wall in Tore Supra [1], the net erosion rate could be reduced by 2 orders of magnitude compared with the gross erosion rate. As a consequence, the erosion and redeposition process to measure should be in the range from 0.1 to 10 µm for a single discharge. With these requirements on the scale of the erosion/redeposition and with the constraint that measurements should be done without any contact, we start with a bibliographic study in order to determine the different techniques permitting to reach such objectives.

Other limitations came from the distance of observation which was in any case larger than 2 meters, the surface roughness of the carbon fibre material, the lack of absolute reference on the surface to analyse and the vibrations which cannot be avoided on a tokamak. From analysis of the relevant bibliography, it has been concluded that only optical methods using interferometric techniques could achieve such requirements and the Speckle interferometry would be the most promising. This work has been conducted by G .Roupillard and presented as thesis for an Engineering diploma in Optics and DEA in plasma physics. The principle of Speckle interferometry is presented on figure 1. A laser beam reflected on the surface to investigate interferes with a reference beam on a CCD camera. Displacements of the surface in the direction of the laser beam modify the interference fringes. Numerical analysis of the images allows to transform these modifications into displacements.

Figure 1 : Speckle interferometer set-up First experiments have been conducted in order to test the feasibility of such technique on a carbon fibre material. A CFC tile (32x20mm) brazed on a copper plate and a Yag laser at 532nm were used, three images were recorded with a CCD camera with different phases. The phase shifting is realised by changing the voltage on the piezoelectric mirror. Applying a light strength on the top part of the sample induces a very small displacement perpendicular to the laser beam. Then, 3 images were taken with different phases. We can see on figure 2, the image of the CFC sample as seen by the CDD camera and the displacement calculated from the analysis of the Speckle interferometers. This experiment shows that Speckle interferometry can be used on CFC material and provide qualitative and quantitative information on surface displacement.


Figure 2 : Image of the CFC sample as seen by the CDD camera (a) and the displacements

deduced from Speckle interferometer (b) In the second part of his report G. Roupillard described the possible implantation of such a diagnostic on Tore Supra. The main plasma facing component in Tore Supra is the Toroidal Pumped Limitor (TPL) on which the plasma is leaning. It consists of 576 fingers, with a length of 50 cm, made carbon fibre composite (N11) brazed on copper tubes. The only possible observation point of the PTL is a window localised on a upper port. Apart the leading edge, localised on the high field side, the totally of the flat surface of the fingers can be seen from this window. Then, the visible area allows a full radial coverage and toroidal extension of several fingers. The schematic view of the limitor with the accessible area from the window is presented on figure 3. Due to the vibrations of the limitor, the Speckle intererometry technique described on figure 1, is not applicable on the tokamak. In order to be insensible to the displacements induce by the vibrations, a Speckle interferometer using 2 wavelength has been proposed [2]. By using a second laser, it is possible not only to measure the relative displacement but also to perform a shape measurement of the surface. The choose of the second laser is such as the resulting wavelength called synthetic wavelength is in the range of the expected erosion to measure.

Figure 3 : Schematic view of the CIEL limitor and the diagnostic implantation

A design of a diagnostic based on 2 wavelength Speckle interferometer has been proposed and is described on

figure 4. The synthetic wavelength, θ⋅λ−λ

λ⋅λ=Λ

cos21

21

with a YAG laser at λ1 = 532 nm and an argon laser at λ2 = 514 nm is about 15µm. Therefore, the depth resolution with such a combination is in the order of 0.5 µm. Following the kick-off meeting in august 2000, a three parties meeting was held in September 2000 at Munich where this design has been presented to EFDA representative (Dr Wu) and Technic University Munchen (Pr Koch). Several points have been discussed but a general agreement has been given to this proposal. Later at the end of the year, a second experimental campaign has been done with this setting. The objective was to demonstrate the possibility to obtain a shape profile by using the 2 wavelength Speckle interferometer technique. Different tests have been performed on samples with calibrated roughness without been able to determine any shape measurement. Further experiments were conducted on metallic samples which showed that 2 wavelength Speckle interferometer gives qualitative measurements by using the synthetic wavelength whereas Speckle interferometer using only one of the wavelength did not give any information. Erosion on CFC tile induced by using a Ruby laser with energy in the range of 10 to 500 mJ focused on a 2mm2 spot has been performed ; although the laser spot was visible by eyes, no erosion measurement could be obtained. Even though these experiments did not provide reliable data on shape or erosion measurements, this 2 wavelength interferometer technique seems promising and need further development.


Figure 4 : Proposal design for a 2 wavelength Speckle interferometer diagnostic on ToreSupra CONCLUSIONS Bibliography study showed that erosion/redeposition measurements in the range of 0.1 to 10 µm could be obtained on carbon fibre composite material by using Speckle interferometer technique.. Preliminary experiments have been performed on CFC samples and allowed to measure displacements in the range of few micrometers. Due to the absence of absolute reference and the vibrations occurring on a tokamak, 2 wavelengths Speckle interferometer is required to measure erosion on CIEL limitor. Although the first experiments did not provide reliable data, this technique seems promising. REFERENCES [1] E.Gauthier et al, J. Nucl. Mater., 220-222, (1995) 506-

510 [2] A. W. Koch, M. Ruprecht, and R. Wilhelm, "Laser

Speckle Techniques for in situ-Monitoring of Erosion and Redeposition at Inner Walls in Large Experimental Fusion Devices," Max-Planck-Institut Für Plasmaphysik Garching bei München (1995)

REPORTS AND PUBLICATIONS - Etude et dimensionnement d’un diagnostic de mesure

d’érosion dans TORE-SUPRA par interférométrie de speckle.. Rapport de stage de DEA. G. Roupillard.

- Réunion de début de projet EFDA T438-01 du 2 août

2000 Kick-off meeting. CFP/CRR-2000.006 - G. Roupillard, E. Gauthier. - Etude de faisabilité (du 14.11.2000 au 16.11.2000).

Mesure d'ablation par interférométrie speckle à deux longueurs d'onde.

CFP/CRM-2001.006 - G. Roupillard, E. Gauthier. TASK LEADER Eric GAUTHIER DSM/DRFC/SIPP/ICI CEA Cadarache 13108 St Paul Lez Durance Cedex Tél. : 33 4 42 25 42 04 Fax : 33 4 42 25 49 90 E-mail : [email protected]

Nd YAG TPL

CCD

532 nm

Optic toward the PC Quartz

laser Ar+ 514,5 nm

interferential filter

- 299 - Underlying technology / Vessel/In Vessel / Plasma facing components

UT-PFC&C-HIP Task Title : MECHANICAL BEHAVIOUR OF HIP JOINTS INTRODUCTION The main objective of this task is to characterize the mechanical behaviour of joints fabricated using the Hot Isostatic Pressing (HIP) process, which is currently one promising joining technique [ref 1]. Both HIP 316LN to 316LN, and 316LN to a precipitation strengthened alloy (CuCrZr) joints will be tested. Previously, a programme was conducted on an oxide dispersion strengthened Cu alloy (Cu-Al-25) [ref 2]. Copper alloys are candidates for heat sink as first wall, divertor, limiter, baffle components. Validation of fabrication techniques requires performing mechanical properties such as impact test resistance (Charpy U), toughness, fatigue and thermal fatigue. The aim of the thermal fatigue programme is to get a better knowledge of component behaviour under cyclic severe peak temperature as for in-service conditions. All the HIP joints were manufactured by the Tecphy society. The proposed fabrication conditions were respectively 1100 °C, 140 MPa, 2 hours for 316LN to 316LN joints and 920 °C, 100 MPa for 316LN to CuCrZr. A first 316LN to 316LN HIP joint lot was rejected for unacceptable impact test resistance results. A second 316LN to 316LN HIP joint fabrication has been accepted. Both ultrasounds and dye penetrant tests do not show any major defect. All the tests on this lot are now completed. 2000 ACTIVITIES 316 LN TO 316 LN JOINT Tensile The new HIP joint was shown good by tensile test results with the final rupture occurring out of the diffusion-bounding zone. Necking at rupture is about 50 %. Additional tests performed on plate specimen using a video acquisition plainly confirm the good resistance of joint. In some cases, note that rupture starts far away from the HIP joint zone (Figure 1).

Figure 1 : Plate specimen tensile tests (HIP bounded-zone corresponds to the full line)

Impact test Resilience behaviour of 316 LN to 316 LN HIP joint is also very promising. Thus, absorbed energy is at least 45 Joules (Impact Charpy U with 358 Joules). In some cases, no rupture occurs. Toughness CT 20 specimen testing confirmed the good behaviour of the preceding tests. The notch is placed at the bounded zone. Jr-Da curve is presented on Figure 2. Measured crack initiation toughness (determined as a 0.2 mm crack growth) corresponds to about 500 KJ/m2 (KJ ~ 300 MPam1/2). Three tests have been performed, intermediate points correspond to compliance unloading sequences. Such values cannot be considered as intrinsic toughness characteristics referring to practice recommendations, since the tested material ductility is very large.

Figure 2 : Toughness properties of the 316LN to 316 LN HIP joint – Jr – Da curves obtained with CT 20 specimen

testing, full symbols correspond to post-mortem measurements

Four points bending tests Tests have been performed under two distinct conditions : Mixed mode (modes I + II) and pure shearing mode (mode II). A significant crack growth is obtained when KI = 15 MPam1/2 and KII = 45 MPam1/2 are applied simultaneously. As it is showed on Figure 3, a large plastic deformation is then clearly observed, confirming the good joint quality.

316 L 316 L HIP

0

500

1000

1500

0 1 2 3 4Da (mm)

J (

KJ/

m2 )

HIP joint

Before tested As tested


Figure 3 : Four points bending tests on 316 L 316 L HIP joint, a large plasticity occurs in the specimen

without breaking 316 LN TO COPPER JOINT Toughness Crack growth resistance has been determined using CT20 specimens. Jr-Da curve is presented on Figure 4. Crack initiation corresponds to 10 - 20 MPam0.5. Such values appear to be very low. But all are nevertheless significantly higher than one previously measured on Glidcop 316 L joint (ref 4). Three tests have been performed, intermediate points correspond to compliance unloading sequences.

Copper - 316 LN HIP

0

50

100

150

200

0 0,5 1 1,5 2 2,5Da (mm)

J (

KJ/

m2 )

Figure 4 : Jr - Da curve for the copper to 316 LN joint, full symbols correspond to post mortem measurements

Four points bending tests Tests have been performed for three conditions : Mixed mode (I + II), pure shearing mode (II), pure tension mode (mode I). Stress intensity factors KIC, KIIC attain 10 and 30 MPam1/2 respectively. A premature rupture can occur in the HIP zone at a maximum tension of about 90 MPa. Thermal fatigue tests Eight distinct thermal fatigue tests have been performed using the ‘’Cythia’’ thermal fatigue test facility. A ‘’Cythia’’ specimen is a thick tube with an external wall periodically heated by high frequency induction. The internal tube wall is permanently cooled by flowing water (Figure 5). Measured surface temperature range is 70 °C One test has been conducted up to 50 000 cycles (25 days). In every case, n any crack has been detected using ultrasound defect response, as previously defined by CEA/STA (ref. 5). Before each test, experimental parameters, such as water flow, electrical power…, are accurately determined using a reference specimen. The latter is identical to the test specimen, except that holds a larger number of thermocouples. To ensure an accurate evaluation of the testing conditions, temperatures will be registered on the heated surface, in the middle of the copper zone, at the interface zone, in the stainless steel, in the water (Figure 6). CONCLUSIONS The new 316LN to 316 LN HIP joint has a good mechanical behaviour. These results are plainly confirmed by all the mechanical tests such as tensile, toughness, four points bending tests. Thus, final rupture cannot occur in the diffusion-bounding zone. It was established that very low resistance of the previous fabrication is due to presence of chromium oxide in the interface. Such scattering behaviour on the joint shows that very accurate fabrication conditions are needed. As it was confirmed by the thermal fatigue test programme, 316 LN copper joint is acceptable. Besides, toughness is higher than previous fabrication (Glidcop – 316 LN). No any crack was detected using ultra-sound control up to 50 000 cycles.

Figure 5 : Thermal fatigue Figure 6 : Temperatures as a function of time determined ''Cythia'' specimen on reference specimen. Thermocouples are placed in different zones

CuCrZr

Water flow

316 LN

30

24

10

Water

Inner

Outer

HIP


REFERENCES [1] A.A. Tavassoli, Overview of advanced techniques of

fabrication and testing of ITER multiplayer plasma facing walls ISFNT, presented at Tokyo (Japan), April 6 - 11, 1997.

[2] B. Marini, Final Report ITER T214, Subtask CEA - 9,

CEA - 10, CEA - 11 Note Technique SRMA 98 - 2278.

[3] J.M. Gentzbittel, G. Nombalais, B. Ricetti, H. Burlet Note Technique DEM 97 – 66. [4] P. Forget, F. Ziolek Compte Rendu SRMA 98 – 1600. [5] M. Wojtowicz Mise au point de la détection ultra-sonore, intern

report 1999.

TASK LEADER A. FISSOLO DTA/CEREM/DECM/SRMA CEA Saclay 91191 Gif-sur-Yvette Cedex Tél. : 33 1 69 08 31 02 E-mail : [email protected]


UT-PFC&C-SiC Task Title : C-SiC AND BORON DOPED GRAPHITE INTRODUCTION Owing to its low atomic number, good mechanical and thermal properties, graphite is an attractive candidate as protection material for the first walls of fusion reactors. Addition of boron to graphite improves the erosion behaviour that is insufficient for pure graphite. Previous studies showed that the addition of boron to graphite improves the erosion behaviour. The objective of the present work is to investigate the effect of atomic boron addition on the relevant graphite properties. Thus, B-doped, highly dense materials are prepared using powder technology techniques, i.e. powders mixing, cold isostatic pressing (CIP), and hot isostatic pressing (HIP) that are well under control at SE2M/LECMA. 2000 ACTIVITIES

INSERTION OF BORON IN GRAPHITE In a first step the work was focused on the comparison of several procedures for the insertion of boron in graphite in order to select the most appropriate one. These trials were made with a boron concentration of 30 %. The graphite powder, reference TIM REX KS6 (0.55 µm) from LONZA, and the boron powder Grade 1 (0.8 µm) from H.C.STARCK were selected. Four different procedures were investigated : Turbulat procedure (liquid medium) Appropriate amounts of graphite and boron powders are weighed, mixed, and blended in the Turbulat device during 1 hour. The as-blended powder is sieved through a 400 µm sieve. Jar procedure (liquid medium) Powders are introduced in a jar together with ethyl alcohol and alumina balls which ease the blending combined with crushing effect. The jar itself is placed on rollers. After blending the liquid is removed and the blend is dried. The as-obtained powder is sieved through a 400 µm sieve. Ultra-turrax procedure (liquid medium) Here again, the graphite and boron powders are mixed in appropriate amounts, ethyl alcohol is added and the ultra-turrax treatment is operated during eight minutes. Ultra-turrax is used for its shearing effect on agglomerates. Same as in the previous procedure, the liquid is removed, the blend is dried, and the as-obtained powder is sieved through a 400 µm sieve.

Ultrasound/helix procedure (lquid medium) In this procedure the combination of ultrasonic waves and of helix stirring is used for blending the mixture of the graphite and boron powders. The presence or formation of agglomerates is therefore avoided. Liquid removal, baking and sieving are made, same as in the procedures described above. The four powders are shaped by CIP at 2500 bar during 1 minute and than pressed during heat treatment in HIP under 1900 bar at 1600°C. Results As a result of this comparison, the turbulat procedure was discarded for mixing the boron and graphite powders as it leads to heterogeneous materials. More promising results were obtained using the ultra-turrax procedure and the jar procedure. Ultra-turrax is quicker and using shearing forces may be more destructive for graphite platelets compared to crushing action of balls in jar. The jar procedure is retained for future work. DENSIFICATION OF BORON/GRAPHITE COMPOSITE In a second step we studied the densification of boron/graphite composite as a function of the chemical composition of the starting powder mix as well as the influence of the thermal treatment applied to the composite. Powder was prepared following the process selected on the basis of the results obtained during the first semester. After mixing powders in jars and ball milling in alcoholic medium, boron/graphite powder mix is dried then shaped using Cold Isostatic Pressing technique. Densification is performed using Hot Isostatic Pressing technique. Three compositions are prepared corresponding to 10, 20 and 30 vol% of boron. The main difference between the different cycles is the pressure applied during the isothermal stage at 1600°C. The isostatic pressure applied was 100, 127 and 150 MPa. Results Density Densities are measured by immersion in liquid (Archimedes’ principle) to detect open porosity, and by helium pycnometry for accurate closed porosity measurement. Both techniques give similar results for density. Highest difference represents 0.55% of the value that is smaller than accuracy expected for immersion technique. For the three different compositions, porosity is mainly closed. Figure 1 shows relation of closed porosity towards HIP pressure for the two cans pressed under 127 and 150 MPa.


Figure 1 : Evolution of the porosity of C (B) composites versus B content and HIP pressure

Photo 1 : Optical micrograph of 30 vol% boron/graphite composite densified 1600°C - 150 MPa

Porosity versus HIP pressure

0,00

1,00

2,00

3,00

4,00

5,00

6,00

120 130 140 150 160

HIP pressure (MPa)

clo

sed

po

rosi

ty (

%)

bore 10 vol%

bore 20 vol%

bore 30 vol%


On the graph, we see that at given pressure, the higher the boron content, the higher the density of the sample. The boron effect is more important between 10 and 20 vol% than it is between 20 and 30 vol%. increasing the HIP pressure above 127 MPa does not improve the densification of the material. Microscopic observation Samples are cut then enclosed in resin before to be polished. This type of composite is very difficult to polish due to the big difference in hardness between the two phases. Graphite that is very soft fills up the cavity and covers boron particles. Optical observations (photo 1) showed a good homogeneity of the structure for the three B Content in Graphite. X-Rays analysis Qualitative analysis was performed on three samples containing 10, 20 and 30 vol% boron, all of them were sintered at 1650°C, under 150 MPa isostatic pressure for 1.5 hour. Spectra showed only two types of signal, the main one corresponding to graphite and the other corresponding to boron carbide. There is no signal for free boron, even in the sample with highest boron content. We can consider that for HIP treatment at 1650°C for 1 hour, the reaction between boron and graphite is full. CONCLUSION Work performed during the second semester 2000 corroborated results obtained during the first semester, and showed that homogeneous structure can be obtained by mixing powders in jar with ball milling. Composites made of graphite with different boron contents, from 10 up to 30 vol% were elaborated. Density measurement performed on various samples with different boron contents showed that boron eases densification. After HIP treatment at 1650°C for 1 hour, porosity is mainly closed and averaged 3%. Increasing the HIP pressure above 130 MPa does not help to reduce the porosity. This work will be continued next semester and mainly focused on thermal and mechanical characterization of the boron/graphite composites

PUBLICATIONS [1] Boron doped graphites – interim report January/June

2000.N.LOCHET CE2M/LECMA 00-DT-152 [2] Boron doped graphites – June/November 2000

N.LOCHET CE2M/LECMA 00-DT-152 TASK LEADER Nicolas LOCHET DRT/LIST/DECS/SE2M/LECMA CEA Saclay 91191 Gif-sur-Yvette Cedex Tél. : 33 1 69 08 32 16 Fax : 33 1 69 08 57 54 E-mail : [email protected]


UT-PFC&C-SiC/MJ Task Title : DEVELOPMENT OF SiC/METAL JOINING TECHNOLOGY INTRODUCTION SiCf/SiC composites are evaluated for possible use in Fusion Power Reactor (FPR). Such materials need to be joined onto metal in plasma facing components. Last year [1], the joining of SiC base materials onto tungsten has been developed as divertor components will be fabricated with tungsten and this material is commonly used as interlayer in ceramic/metal joining in order to decrease residual stresses generated during the cooling of two joined materials with different Coefficient of Thermal Expansion (CTE). The experimental work was dedicated to the use of different BraSiC® brazing alloys: it has been demonstrated that BraSiC® H1 grade brazing alloy leads to sound SiC/W joints. From a mechanical point of view, two geometries have been proposed for shear tests: for an easier results interpretation, a ‘symmetric sandwich structure type’ such as W/SiC/W or SiC/W/SiC is better. Asymmetrical Four Point Bending has been proposed for the mechanical solicitation of the joint. This last remark is also supported in [2] which was focused on a literature review about mechanical tests on joined materials containing ceramic. The present work was focused on the design of a mechanical test on joined materials: this was based on the modelling of the joined sample after brazing and during the mechanical test in order to evaluate the exact geometry to be used for joint strength measurement. The modelling has been performed on SiC/Eurofer as the difference of thermal expansion behaviour is greater compared to the difference observed in SiC/W: it should then be possible to have a more discriminating results or at least comparable as the brazing of Eurofer is performed at lower temperature than with W. 2000 ACTIVITIES The residual stresses state at the SiC/metal interface has been modelled after brazing and also during the mechanical test in order to study the stress state change. MODELLING OF BRAZED SAMPLE DURING COOLING DOWN A sandwich type sample geometry SiC/metal/SiC has been used in order to have a symmetrical stress state (cf figure 1).

Figure 1 : sample geometry The typical Von Mises residual stresses in such configuration after brazing at 750°C is described in figure 2. The general Von Mises stresses variation orientation, given in figure 2, is reported in figure 3: it is oriented of about 45° from the joint interface. The mechanical test consists of applying a load in compression on top of the metal face as the SiC are blocked at the base of the sample.

Figure 2 : Von Mises stresses at 20°C

Figure 3 : Von Mises stresses Variation orientation

SiC SiC

Eurofer

5 mm

2.5 mm


MECHANICAL MODELLING DURING MECHANI-CAL TEST The typical Von Mises stresses at the end of loading is reported in figure 4.

Figure 4 : Von Mises stresses at 300°C at the end of the loading

Whatever the residual stresses state after the brazing process is taken into account, there is a rotation of the Von Mises stresses variation orientation between the beginning and the end of the test from 45° to almost 90° compared to the joint interface as drawn in figure 5.

Figure 5 : Rotation of the Von Mises stresses variation orientation from 45° to 90° during the loading

It should also be noted that there is large variation of the stress level on each side of the interface, so it is doubtful that the rupture will occur with the interface. CONCLUSIONS The main conclusions are: - even the stresses are not homogeneously and

symmetrically distributed within the sample on both side of the joint interface at the end of the brazing cycle, during the loading, the stresses shift to a more homogeneous and symmetrical distribution on both side of the joint interface,

- there is a difference in the stresses magnitude on each side of the interface: there are higher in the SiC than in the metal. Thus, it should break within the ceramic.

Further works are needed to generate the same kind of stresses level on each side of the interface in order to load the joint rather the material on each side of the joint. A difference in stresses level on each side of the interface will lead to a reinforcement of the adhesion properties on one of the two sides and a decrease ones on the other side depending on the Young modulus, the Poisson’s ration and the atom movement ability of the two materials joined as noted in [4]. REFERENCES [1] A. Gasse, P. Manicardi, R. Couturier, “Development

of SiC/metal joining techniques”, European Fusion Technology Programme, Task UT-PFC&C-SiC/MJ, Field PFC&C, 1st Progress Report, Note Technique DEM 93/99, 17/12/1999.

[2] R. Couturier, A. Gasse, “SiC-SiC ceramic composite:

definition of material grade and mechanical test”, Advancement Report M1-M2, Subtask n° ADV 1.1.2, Note Technique DEM 99/84, 9/12/1999.

[3] F. Saint-Antonin, P. Le Gallo, R. Couturier,

"Development of SiC/metal joining technology”, European Fusion Technology Programme, Field PFC&C, Task UT-PFC&C-SiC/MJ, 2nd Progress Report, Period 01/01/2000 to 31/11/2000, Note Technique DEM/SGM 123/2000, 13/12/2000.

[4] F. Saint-Antonin, “How to characterise heterogeneous

joint ?”, Note Technique DEM 04/2000, February 2000.

TASK LEADER François SAINT-ANTONIN DTA/DEM/SGM CEA Grenoble 17, rue des Martyrs 38054 Grenoble Cedex 9 Tél. : 33 4 76 88 54 77 Fax : 33 4 76 88 95 38 E-mail : [email protected]


UT-SM&C-VVI Task Title : DESIGN AND ANALYSIS OF VACUUM VESSEL AND INTERNALS INTRODUCTION The task UT-SM&C-VVI is a contribution to the European Fusion Underlying Technology Program. This task is intended to maintain/develop competence and analysis tools in the field of the design of tokamak vacuum-vessel and internals (divertor, limiters, baffle, first wall, shielding blanket and breeding blanket). Over the last years, the TAURO blanket concept has been developed at CEA. It is a self-cooled liquid metal breeder blanket using the SiCf/SiC as structural material and the lithium-lead eutectic both as coolant and breeder. Silicon carbide composites (SiCf/SiC) appear as one of the most promising low activation material for structural application in fusion power reactors. Furthermore, liquid metal cooling has the potential for high power density, low pressure and high coolant temperature suitable for high efficiency power conversion system. In the present study, the potentials of a liquid metal cooled divertor using SiCf/SiC as structural material, compatible with liquid metal-cooled blanket system is assessed. 2000 ACTIVITIES The 2000 activities covered analyses and more specific conceptual design work on liquid metal-cooled divertor using SiCf/SiC as structural material. A model describing the SiCf/SiC behaviour recently implemented in the CASTEM 2000 code was used for the mechanical analyses. ANALYSES AND RESULTS Geometry and materials Previous studies [1] on several liquid metal cooled divertor concepts resulted in a castellated square tube using W-alloy both as structural and armor material and SiCf/SiC as electrical insulator. Because of the unfavorable behavior of the W-alloy under irradiation, it would be preferable to avoid its use as structural material. Furthermore, improvements have been done in the performances of the SiCf/SiC composites and in its behavioral modeling, which could lead to consider the use of the SiCf/SiC as structural material for the square tube. The thickness of the W tiles and the thickness and the radial depth of the SiCf/SiC tube were assumed as parameters in the analyses. The length of the tube was fixed to 0.5 m and its toroidal width to 25 mm [2]. Figure 1 shows a three-dimensional view and a section of this divertor.

Figure 1 : Castellated SiCf/SiC divertor

Operating conditions It was assumed that an uniform surface heat flux is deposited on the entire length of the tube. A fixed volume heat flux per material corresponding to a neutron wall loading of 1.2 MW/m2 was included [1]. Parametric thermal analyses The sensitivity of the design to several parameters was evaluated with the aim of maximizing the allowable heat flux, while meeting the temperature limits (1300 °C for the SiCf/SiC and 2000 °C for the W) and maintaining the Pb-17Li velocity lower than 2.5 m/s (maximum value estimated to avoid significant MHD pressure drops). A simplified modeling was used to take into account boundary conditions with the liquid metal. A 2D section was analyzed, the liquid metal being included in the mesh. Thermal calculations were performed in transient, simulating the warming up of the liquid metal with the artifact described in [2]. The Pb-17Li inlet temperature was fixed to 300 °C. In Table 1, maximum, average and minimum temperature values at outlet section are summarized for the examined cases. It can be seen that the limiting factor is the SiCf/SiC maximum temperature, while the W tile temperature stays much lower than its maximum allowable value. In the table, we can outline also the influence of the SiCf/SiC conductivity and of the tube thickness on the maximum tube temperature and thus on its thermal gradient (being the minimum temperature quite constant). In Figure 2 is shown the temperature distribution at the outlet section for the case with Heat Flux = 5 MW/m2, v Pb-17Li = 2.5m/s, tube thickness= 1mm, tube depth = 15 mm, SiCf/SiC conductivity = 20 W/m/K, in bold on the table.


Table 1 : Temperature values of various divertor parts

SiCf/SiC T (°C) W tiles T(°C) HF (MW/m2)

v (m/s)

Tube thick (mm)

Tube depth (mm)

Tiles thick (mm)

SiCf/SiC Cond

(W/m/K) max ave min max ave min

6 2,5 1 15 5,5 20 1169 553 304 1507 1282 1053 6 2,5 1 15 5,5 15 1284 571 304 1625 1385 1144 5 2,5 1 15 6 20 1037 516 304 1338 1141 939 5 3 1 15 5,5 15 1094 514 304 1369 1176 980

5 2,5 1 15 5,5 20 1035 515 304 1310 1126 938 5 2,5 1 13 5,5 20 1036 527 304 1310 1127 939 5 2,5 1 10 5,5 20 1036 548 304 1311 1128 940 5 2,5 1 15 5,5 15 1131 530 304 1408 1212 1014 5 2,5 1 13 5,5 15 1132 543 304 1409 1213 1015 5 2,5 1 10 5,5 15 1133 565 304 1409 1214 105 5 2 1 15 5,5 20 1084 536 305 1361 1175 984 5 2 1 15 5,5 15 1181 551 305 1460 1261 1060 5 1,5 1 15 5,5 20 1154 567 306 1434 1244 1049 5 1,5 1 15 5,5 15 1253 582 306 1535 1331 1125 5 1 1 15 5,5 15 1368 636 309 1654 1445 1231 5 1 1 15 5,5 20 1268 621 309 1553 1357 1156 5 3 1,5 15 5,5 15 1293 555 305 1576 1358 1140 5 2,5 1,5 15 5,5 20 1187 548 305 1467 1265 1060 5 2,5 1,5 15 5,5 15 1332 572 305 1616 1395 1175

Figure 2 : Temperature distribution (°C) at the outlet section (Heat Flux = 5 MW/m2, v Pb-17Li = 2.5m/s,

tube thickness= 1mm, tube depth = 15 mm, SiCf/SiC cond. = 20 W/m/K)

Thermo-mechanical analyses In the mechanical calculations, only the structural material is meshed. A pressure of 0.3 MPa corresponding to the pressure of the liquid metal was applied on the internal walls. The 2D thermal fields obtained at different times were projected on the different planes of the 3D mesh, the number of the mesh elements in the third direction being adjusted to the number of thermal pitches [Erreur ! Signet non défini., Erreur ! Signet non défini.]. The SiCf/SiC behavioral model recently developed and implemented in the FEM CASTEM 2000 code was used for the calculations [3].

The "TAURO design criteria" were applied for the results check. Following these criteria, in all parts of the component, the called TAURO criterion must be lower than 1 and the module of the shear stresses trough the thickness must be lower than 44 MPa. In Table 2, the TAURO criterion value and the shear stress trough the thickness are reported for some of the examined cases. While the TAURO criterion value is everywhere fulfilled, the maximum value of the module of the shear stresses trough the thickness is, in all cases, higher than the limit. In Figure 3, the shear stresses distribution in various parts of the SiCf/SiC tube are shown for the case chosen previously (Heat Flux = 5 MW/m2, v Pb-17Li = 2.5/s, tube thickness = 1mm, tube depth = 15 mm, SiCf/SiC conductivity = 20 W/m/K). It can be seen that regions where the shear stresses trough the thickness exceed the corresponding limit are very restricted and limited to the region with discontinuity on geometry or on boundary condition. Furthermore, from a more accurate table check, it can be brought out that these values are quite insensitive to the different parameters, which tends to indicate that they probably depend more on the model than on physical reason. More detailed analyses are needed in order to verify the origin of these high stress values and try to reduce them.


Table 2 : TAURO criterion and maximum shear stress trough the thick for different parameters

HF (MW/m2)

v (m/s)

Tube thick (mm)

Tube depth (mm)

Tiles thick (mm)

SiCf/SiC Cond

(W/m/K)

TAURO criterion σxz MPa σyz MPa

6 2,5 1 15 5,5 15 0,9956 59 102

5 2.5 1 15 5.5 20 0.9103 49 92

5 2,5 1 15 5,5 15 0,95642 52 95

5 2,5 1 13 5,5 15 0,91768 51 94

5 1,5 1 15 5,5 20 0,94047 54 96

5 1,5 1 15 5,5 15 0,9764 58 100

5 1 1 15 5,5 15 0,99779 62 104

5 1 1 15 5,5 20 0,9645 58 101

Figure 3 : Shear stresses trough the thickness in the SiCf/SiC tube (Heat Flux = 5 MW/m2, Pb-17Li = 2.5m/s, tube thickness= 1mm, tube depth = 15 mm, SiCf/SiC cond. = 20 W/m/K)


CONCLUSIONS The limits and the potential of a castellated square divertor using SiCf/SiC as structural material and Pb-17Li as coolant, in line with the TAURO reactor characteristics were assessed in this study. A behavioral mechanical model suitably developed for the SiCf/SiC and the corresponding resistance criteria were used for the mechanical analyses. The calculations showed that this concept could be very performing in term of maximum allowable heat flux. However, high shear stresses trough the thickness resulted from the mechanical analyses. Considering that these stresses are limited to discontinuity region and quite insensitive to different parameters, more detailed mechanical analyses are needed in order to best assess their origin and the possible solutions. REFERENCES [1] L. Giancarli, J.-F. Salavy, Assessment of forced

convection cooled liquid metal divertors, CEA Report, SERMA/LCA/RT/99-2683/A.

[2] M. Fütterer, G. Aiello, F. Gabriel, J.-F. Salavy,

Thermal-hydraulics models for liquid-metal cooled vacuum vessel internals, CEA Report SERMA/LCA/RT/00-2755/A.

[3] H. Golfier, G. Aiello, L. Giancarli, Progress on the

TAURO blanket system, CEA Report, SERMA/LCA/RT/00-2837/A.

PUBLICATIONS A. Li Puma et al., Thermal - hydraulic and thermo-mechanical study of self cooled SiCf/SiC divertor concept, CEA Report SERMA/LCA/RT/01-2925

TASK LEADER Luciano GIANCARLI DRN/DMT/DIR CEA Saclay 91191 Gif-sur-Yvette Cedex Tél. : 33 1 69 08 21 37 Fax : 33 1 69 08 66 42 E-mail : [email protected]

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