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Thermal fatigue damage of the divertor plate
Fusion Engineering and Design 49 – 50 (2000) 343 – 348 www.elsevier.com/locate/fusengdes
Thermal fatigue damage of the divertor plate Satoshi Suzuki *, Koichiro Ezato, Kazuyoshi Sato, Kazuyuki Nakamura, Masato Akiba NBI Heating Laboratory, Naka Fusion Research Establishment, Japan Atomic Energy Research Institute (JAERI), 801 -1 Mukoyama, Naka-machi, Naka-gun, Ibaraki-ken 311 -01, Japan
Abstract Thermal fatigue of the divertor plate is one of the key issues which governs the lifetime of the divertor plate. A thermal cycling experiment of divertor mock-ups was carried out to investigate the thermal fatigue behavior of the divertor structure in a high heat flux test facility at Japan Atomic Energy Research Institute (JAERI). A cyclic heat flux of 5 MW/m2 was loaded onto the mock-ups to simulate the steady state thermal condition of ITER divertor plate. A pulse duration of 30 s was selected so that the mock-ups reach thermal steady state. The mock-ups showed water leakage due to thermal fatigue cracking of the cooling tube after 400 cycles. The fatigue crack was observed at the top of the cooling tube. According to a numerical analysis, the maximum strain amplitude was over 5% at the top of the cooling tube. The cracking of the cooling tube was caused by the large strain concentration due to structural discontinuity of the copper heat sink. It was found that the heat sink block should be separated into smaller blocks to reduce the strain concentration at the cooling tube. © 2000 Elsevier Science B.V. All rights reserved. Keywords: Thermal fatigue; Divertor plate; Copper
1. Introduction Plasma facing components of next generation fusion devices, such as international thermonuclear experimental reactor (ITER), are subjected to a high thermal/particle load and an electromagnetic load during operation. In particular, a divertor plate is requested to handle the highest thermal load among the plasma facing components in ITER. The thermal load is cyclically loaded to the divertor plate due to a repetitious * Corresponding author. Tel.: +81-29-2707552; fax: + 8129-2707558. E-mail address: [email protected] (S. Suzuki).
operation of ITER. Therefore, it is one of the most important issues for the divertor plate to investigate the thermal fatigue behavior against cyclic thermal loads, since thermal fatigue of the structural material strongly affects the integrity of the divertor plate. So far, thermal cycling experiments have been carried out in many institutions to investigate the thermal fatigue behavior of the divertor plates [1–3]. However, most of these experiments were focused on the integrity of the bond interface between the armor and the heat sink materials. The authors have reported the thermal fatigue on the cooling tube made of oxygen-free-high-conductivity copper (OFHC-Cu) [4], which have
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Fig. 1. Schematic drawing of particle beam engineering facility (PBEF).
pointed out the low cycle fatigue fracture of the OFHC-Cu cooling tube under the ITER divertor thermal conditions. In this study, aluminum-oxide-dispersion-strengthened copper (DSCu) was chosen as the structural material of the cooling tube to improve the durability against a cyclic thermal stress. In addition, a full-scale ITER divertor mock-up (vertical target) was fabricated to simulate the realistic mechanical condition of the ITER divertor plate, such as mechanical constraint from a rigid back plate and a stress concentration due to geometrical discontinuity of the divertor structure.
2. Experiment
2.1. Test facility The test facility used in this study is particle beam engineering facility (PBEF) in Japan
Atomic Energy Research Institute (JAERI). PBEF can produce hydrogen ion beam at an acceleration voltage up to 50 kV and at a beam current up to 30 A for 1000 s. One of the advantages of the usage of hydrogen ion beam as a heat source is low reflectivity for a metal target. In particular, it is suitable to simultaneously heat target surfaces, which are made of different kinds of materials, such as a combination of carbon and tungsten. The schematic drawing of PBEF is shown in Fig. 1. The dimension of vacuum chamber is 3.5 m (width) × 4 m (height) ×7 m (length). A large ion source has been developed and implemented to perform high heat flux experiments of large-scale mock-ups. The large ion source is capable of producing sheet-like hydrogen ion beams at an irradiation area of 10 cm × 1 m with a maximum heat flux of 10 MW/m2.
2.2. Di6ertor mock-up A divertor mock-up simulating an inboard vertical target was developed. Fig. 2 shows the schematic drawing of the divertor mock-up. The major dimension of the divertor mock-up is 35 mm (width) × 1.3 m (length), which corresponds to a single slice of the ITER inboard vertical target. The surface material of the mock-up is made of powder-sintered tungsten and a unidirectional carbon-fiber-reinforced-carbon composite (1D-CFC). This surface material combination is based on the ITER-FDR design [5]. These surface materials were directly brazed onto OFHC-Cu heat sink with a silver braze material (Ti–Cu– Ag). The braze process was conducted in a vacuum environment at a temperature of 850°C for a hold time of several minutes. The cooling tube has
Fig. 2. Schematic drawing of ITER divertor (vertical target) mock-up.
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the steady state thermal condition of the ITER divertor was conducted in PBEF. A heat flux of 5 MW/m2 was cyclically loaded onto the surface of the armor tiles. The heating duration of 30 s and the suspend duration of 120 s were selected so that the mock-up reached the thermal steady state. Since this experiment focuses on the thermal fatigue behavior of the structural material of the divertor plate, the heating duration of 30 s was chosen to avoid suffering from creep damage. In the thermal fatigue experiment, an infrared camera monitored surface temperature profile; the temperature evolution of the heat sink was monitored by thermocouples attached on the side. The coolant was purified water with an axial flow velocity of 10 m/s at a temperature of 25°C. Fig. 3. Surface temperature evolution of the mock-up (comparison between the experimental data and the analytical data).
Fig. 4. Fatigue crack appeared at the cooling tube (photo taken from the inside of the cooling tube).
a triplex layer. The outer and the inner layers were made of OFHC-Cu. The middle layer was made of DSCu. The role of the outer OFHC-Cu layer is to sustain good compatibility with the braze material; the role of the inner layer is to form a compliant layer for the mechanical fixation of the twisted tape. The OFHC-Cu heat sink, the DSCu cooling tube, and the stainless steel back plate were also brazed together at the same time. The upper heat sink in the tungsten-armored side, which was made of OFHC-Cu, consists of a single plate to simplify the fabrication process.
2.3. Thermal fatigue experiment Thermal fatigue experiment which simulated
2.4. Experimental results Fig. 3 shows the typical surface temperature evolution of the mock-up. According to the temperature evolution obtained from the infrared camera, the mock-up reached the thermal steady state within 30 s. The predicted temperature evolution from the finite element calculation showed good agreement with the experimental result. Thermal response of the mock-up showed no change up to 400 cycles. The thermal performance of the mock-up proved to be sound during the experiment. However, the water leakage from the mock-up was observed after 400 cycles. As a result of a microscopic observation after the experiment, it was found that the water leakage occurred at the cooling tube located in between the tungsten-armored region and the CFC-armored region. Fig. 4 shows that the crack penetrated the cooling tube wall. It is significant that location of the crack is independent of the position of the twisted tape attached inside of the cooling tube. It is expected that the plastic deformation of the inner wall of the cooling tube due to the twisted tape insertion does not cause the fatigue crack initiation. Fig. 5 shows the photo of the fracture surface of the cooling tube obtained from a scanning electron microscope (SEM) observation. The striation, which is typical in the fracture surface of fatigue cracking, was clearly
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observed in the outer and the inner OFHC-Cu layer of the cooling tube. On the other hand, no clear evidence of fatigue cracking was found in the middle layer made of DSCu. As shown in Fig. 6, the fracture surface of the middle layer is quite different from that obtained from the OFHC-Cu layer. This fracture surface implies brittle fracture that occurred in the DSCu layer. Therefore, the number of thermal cycles required so that the fatigue crack penetrates the DSCu layer is expected to be very small. Therefore, the experimental result in this study suggests that the fatigue lifetime of DSCu is not dominated by the number of crack growth but mainly by the number
Fig. 5. Fracture surface of the cooling tube (photo taken from the inner OFHC-Cu layer).
Fig. 6. Fracture surface of the cooling tube (photo taken from the middle DSCu layer).
of cycles for fatigue crack initiation. It is also supposed that DSCu is a crack-sensitive material and has high crack growth rate after crack initiation.
3. Discussion A thermal fatigue cracking of the copper-based cooling tube at this heat flux level (5 MW/m2) has never been observed in the high heat flux experiments formerly performed at JAERI [6]. In addition, the fact that the fatigue crack appeared at the cooling tube in between the tungsten-armored region and the CFC-armored region implies a specific reason of this fatigue damage. It is expected that the cyclic thermal deformation of the upper heat sink and the constraint due to the rigid back plate caused large strain concentration at the cooling tube. Three-dimensional finite element analyses were performed to clarify the reason of this fatigue fracture. The finite element code used was ABAQUS, version 5.7 [7]. Fig. 7 shows a finite element model and the boundary conditions of the analyses. The model includes six armor tiles with three tungsten tiles and three 1D-CFC tiles, which corresponds to the central part of the mock-up. Eight node-linear elements were used in the thermal and stress analyses. In the thermal analysis, uniform heat flux of 5 MW/m2 onto the surface of all armor tiles was assumed to simulate the experiment. Based on the temperature evolution obtained in this transient thermal analysis, subsequent elastoplastic stress analyses were performed. Prior to the elastoplastic stress analysis assuming the cyclic thermal loading, a residual stress analysis was conducted to obtain the initial stress field of the model. The residual stress behaved as mean stress during the cyclic thermal loading. The reference temperature of the residual stress analysis was 780°C, which corresponded to the solidified temperature of the Ti–Cu–Ag braze material. In the stress analysis, elastoplastic behavior of DSCu and OFHC-Cu was taken into account; the material properties of OFHC-Cu were assumed to be of a fully annealed material. Fig. 8 shows the equivalent mechanical strain
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Fig. 7. Finite element mesh and boundary condition (line ‘A’ corresponds to the top of the outer surface of the tube).
profile along the line ‘A’ corresponding to the top of the outer surface of the cooling tube. Large strain concentration appeared at the top of the tube locating in between the armor tiles. Particularly, the largest equivalent mechanical strain of 6.6% was obtained at the tube adjoining the tungsten-armored side and the CFC-armored side. The result of the residual stress analysis shows that the stress field of the OFHC-Cu layer is in an elastoplastic state after the braze process. Subsequently, the elastoplastic stress analysis simulating the experiment was carried out using this residual stress field as an initial condition. In the analysis, three thermal cycles were assumed to stabilize the stress–strain behavior of the model. Fig. 8 also shows the equivalent mechanical strain amplitude of the cooling tube obtained at the third thermal cycle. The maximum equivalent mechanical strain amplitude of 5.2% appeared at the tube adjoining the tungsten-armored side and the CFC-armored side. As a result of these stress analyses, it was found that the low cycle fatigue fracture was caused by the large strain concentration and by the large strain amplitude of the cooling tube. Moreover, the geometric discontinuity formed by the upper heat sink, which consists of a single OFHC-Cu plate, caused the strain concentration because the peak of the strain profile at the CFC-armored side along the line ‘A’ was
not as significant as that of the crack region. To avoid such fatigue cracking, separation of the upper heat sink is essential for this divertor structure.
Fig. 8. Strain distribution along line ‘A’ after braze process and mechanical strain amplitude at the end of the third cycle.
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4. Conclusion
Acknowledgements
Based on this study, we draw the following conclusions: 1. Thermal fatigue experiment of the ITER divertor (vertical target) mock-up was performed in PBEF. The DSCu cooling tube with triplex layer showed low cycle fatigue fracture at 400 thermal cycles under a cyclic heat flux of 5 MW/m2. 2. The fatigue crack was found at the cooling tube adjoining the tungsten-armored side and the CFC-armored side. According to the posttest microscopic observation, the DSCu layer showed a brittle fracture surface whereas the OFHC-Cu layer of the tube showed a clear evidence of the fatigue fracture. Since DSCu was proved to be a crack-sensitive material in this study, precise evaluation on the fatigue crack initiation/growth is necessary to use DSCu as a cooling tube material for the divertor application. 3. It was found that the geometric discontinuity formed by the single plate heat sink strongly affects the fatigue lifetime of the cooling tube. For the design of the divertor plate, separation of the heat sink for each armor tile is essential to avoid the large strain concentration at the cooling tube.
The authors wish to thank Y. Okumura and other members of NBI heating laboratory for their valuable discussions and comments. They would also like to acknowledge M. Ohta and S. Matsuda for their support and encouragement.
References [1] G. Vieider, V. Barabash, et al., Overview of the EU small scale mock-up tests for ITER high heat flux components, Fusion Eng. Des. 39 – 40 (1998) 211 – 218. [2] M. Akiba, S. Suzuki, Overview of the Japanese mock-up tests for ITER high eat flux components, Fusion Eng. Des. 39 – 40 (1998) 219 – 225. [3] I. Mazul, R. Giniatulin, et al., Manufacturing and testing of ITER divertor gas box liners, Proceedings of the 20th Symposium on Fusion Technology, vol. A, 1998, pp. 77–80. [4] S. Suzuki, T. Suzuki, et al., Development of divertor plate with CFCs bonded onto DSCu cooling tube for fusion reactor application, J. Nucl. Mater. 258 – 263 (1998) 318– 322. [5] ITER-EDA, Technical basis for the ITER Final Design Report, cost review and safety analysis, ITER EDA documentation series No. 16, 1998. [6] S. Suzuki, T. Suzuki, Development of divertor high heat flux components at JAERI, Proceedings of the 17th IEEE/NPSS Symposium on Fusion Engineering, vol. 1, 1977, pp. 385– 388. [7] ABAQUS/Standard user’s manual version 5.7, Hibbitt, Karlsson & Sorensen, 1999.
Thermal fatigue damage of the divertor plate Satoshi Suzuki *, Koichiro Ezato, Kazuyoshi Sato, Kazuyuki Nakamura, Masato Akiba NBI Heating Laboratory, Naka Fusion Research Establishment, Japan Atomic Energy Research Institute (JAERI), 801 -1 Mukoyama, Naka-machi, Naka-gun, Ibaraki-ken 311 -01, Japan
Abstract Thermal fatigue of the divertor plate is one of the key issues which governs the lifetime of the divertor plate. A thermal cycling experiment of divertor mock-ups was carried out to investigate the thermal fatigue behavior of the divertor structure in a high heat flux test facility at Japan Atomic Energy Research Institute (JAERI). A cyclic heat flux of 5 MW/m2 was loaded onto the mock-ups to simulate the steady state thermal condition of ITER divertor plate. A pulse duration of 30 s was selected so that the mock-ups reach thermal steady state. The mock-ups showed water leakage due to thermal fatigue cracking of the cooling tube after 400 cycles. The fatigue crack was observed at the top of the cooling tube. According to a numerical analysis, the maximum strain amplitude was over 5% at the top of the cooling tube. The cracking of the cooling tube was caused by the large strain concentration due to structural discontinuity of the copper heat sink. It was found that the heat sink block should be separated into smaller blocks to reduce the strain concentration at the cooling tube. © 2000 Elsevier Science B.V. All rights reserved. Keywords: Thermal fatigue; Divertor plate; Copper
1. Introduction Plasma facing components of next generation fusion devices, such as international thermonuclear experimental reactor (ITER), are subjected to a high thermal/particle load and an electromagnetic load during operation. In particular, a divertor plate is requested to handle the highest thermal load among the plasma facing components in ITER. The thermal load is cyclically loaded to the divertor plate due to a repetitious * Corresponding author. Tel.: +81-29-2707552; fax: + 8129-2707558. E-mail address: [email protected] (S. Suzuki).
operation of ITER. Therefore, it is one of the most important issues for the divertor plate to investigate the thermal fatigue behavior against cyclic thermal loads, since thermal fatigue of the structural material strongly affects the integrity of the divertor plate. So far, thermal cycling experiments have been carried out in many institutions to investigate the thermal fatigue behavior of the divertor plates [1–3]. However, most of these experiments were focused on the integrity of the bond interface between the armor and the heat sink materials. The authors have reported the thermal fatigue on the cooling tube made of oxygen-free-high-conductivity copper (OFHC-Cu) [4], which have
0920-3796/00/$ - see front matter © 2000 Elsevier Science B.V. All rights reserved. PII: S 0 9 2 0 - 3 7 9 6 ( 0 0 ) 0 0 3 9 4 - X
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Fig. 1. Schematic drawing of particle beam engineering facility (PBEF).
pointed out the low cycle fatigue fracture of the OFHC-Cu cooling tube under the ITER divertor thermal conditions. In this study, aluminum-oxide-dispersion-strengthened copper (DSCu) was chosen as the structural material of the cooling tube to improve the durability against a cyclic thermal stress. In addition, a full-scale ITER divertor mock-up (vertical target) was fabricated to simulate the realistic mechanical condition of the ITER divertor plate, such as mechanical constraint from a rigid back plate and a stress concentration due to geometrical discontinuity of the divertor structure.
2. Experiment
2.1. Test facility The test facility used in this study is particle beam engineering facility (PBEF) in Japan
Atomic Energy Research Institute (JAERI). PBEF can produce hydrogen ion beam at an acceleration voltage up to 50 kV and at a beam current up to 30 A for 1000 s. One of the advantages of the usage of hydrogen ion beam as a heat source is low reflectivity for a metal target. In particular, it is suitable to simultaneously heat target surfaces, which are made of different kinds of materials, such as a combination of carbon and tungsten. The schematic drawing of PBEF is shown in Fig. 1. The dimension of vacuum chamber is 3.5 m (width) × 4 m (height) ×7 m (length). A large ion source has been developed and implemented to perform high heat flux experiments of large-scale mock-ups. The large ion source is capable of producing sheet-like hydrogen ion beams at an irradiation area of 10 cm × 1 m with a maximum heat flux of 10 MW/m2.
2.2. Di6ertor mock-up A divertor mock-up simulating an inboard vertical target was developed. Fig. 2 shows the schematic drawing of the divertor mock-up. The major dimension of the divertor mock-up is 35 mm (width) × 1.3 m (length), which corresponds to a single slice of the ITER inboard vertical target. The surface material of the mock-up is made of powder-sintered tungsten and a unidirectional carbon-fiber-reinforced-carbon composite (1D-CFC). This surface material combination is based on the ITER-FDR design [5]. These surface materials were directly brazed onto OFHC-Cu heat sink with a silver braze material (Ti–Cu– Ag). The braze process was conducted in a vacuum environment at a temperature of 850°C for a hold time of several minutes. The cooling tube has
Fig. 2. Schematic drawing of ITER divertor (vertical target) mock-up.
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the steady state thermal condition of the ITER divertor was conducted in PBEF. A heat flux of 5 MW/m2 was cyclically loaded onto the surface of the armor tiles. The heating duration of 30 s and the suspend duration of 120 s were selected so that the mock-up reached the thermal steady state. Since this experiment focuses on the thermal fatigue behavior of the structural material of the divertor plate, the heating duration of 30 s was chosen to avoid suffering from creep damage. In the thermal fatigue experiment, an infrared camera monitored surface temperature profile; the temperature evolution of the heat sink was monitored by thermocouples attached on the side. The coolant was purified water with an axial flow velocity of 10 m/s at a temperature of 25°C. Fig. 3. Surface temperature evolution of the mock-up (comparison between the experimental data and the analytical data).
Fig. 4. Fatigue crack appeared at the cooling tube (photo taken from the inside of the cooling tube).
a triplex layer. The outer and the inner layers were made of OFHC-Cu. The middle layer was made of DSCu. The role of the outer OFHC-Cu layer is to sustain good compatibility with the braze material; the role of the inner layer is to form a compliant layer for the mechanical fixation of the twisted tape. The OFHC-Cu heat sink, the DSCu cooling tube, and the stainless steel back plate were also brazed together at the same time. The upper heat sink in the tungsten-armored side, which was made of OFHC-Cu, consists of a single plate to simplify the fabrication process.
2.3. Thermal fatigue experiment Thermal fatigue experiment which simulated
2.4. Experimental results Fig. 3 shows the typical surface temperature evolution of the mock-up. According to the temperature evolution obtained from the infrared camera, the mock-up reached the thermal steady state within 30 s. The predicted temperature evolution from the finite element calculation showed good agreement with the experimental result. Thermal response of the mock-up showed no change up to 400 cycles. The thermal performance of the mock-up proved to be sound during the experiment. However, the water leakage from the mock-up was observed after 400 cycles. As a result of a microscopic observation after the experiment, it was found that the water leakage occurred at the cooling tube located in between the tungsten-armored region and the CFC-armored region. Fig. 4 shows that the crack penetrated the cooling tube wall. It is significant that location of the crack is independent of the position of the twisted tape attached inside of the cooling tube. It is expected that the plastic deformation of the inner wall of the cooling tube due to the twisted tape insertion does not cause the fatigue crack initiation. Fig. 5 shows the photo of the fracture surface of the cooling tube obtained from a scanning electron microscope (SEM) observation. The striation, which is typical in the fracture surface of fatigue cracking, was clearly
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observed in the outer and the inner OFHC-Cu layer of the cooling tube. On the other hand, no clear evidence of fatigue cracking was found in the middle layer made of DSCu. As shown in Fig. 6, the fracture surface of the middle layer is quite different from that obtained from the OFHC-Cu layer. This fracture surface implies brittle fracture that occurred in the DSCu layer. Therefore, the number of thermal cycles required so that the fatigue crack penetrates the DSCu layer is expected to be very small. Therefore, the experimental result in this study suggests that the fatigue lifetime of DSCu is not dominated by the number of crack growth but mainly by the number
Fig. 5. Fracture surface of the cooling tube (photo taken from the inner OFHC-Cu layer).
Fig. 6. Fracture surface of the cooling tube (photo taken from the middle DSCu layer).
of cycles for fatigue crack initiation. It is also supposed that DSCu is a crack-sensitive material and has high crack growth rate after crack initiation.
3. Discussion A thermal fatigue cracking of the copper-based cooling tube at this heat flux level (5 MW/m2) has never been observed in the high heat flux experiments formerly performed at JAERI [6]. In addition, the fact that the fatigue crack appeared at the cooling tube in between the tungsten-armored region and the CFC-armored region implies a specific reason of this fatigue damage. It is expected that the cyclic thermal deformation of the upper heat sink and the constraint due to the rigid back plate caused large strain concentration at the cooling tube. Three-dimensional finite element analyses were performed to clarify the reason of this fatigue fracture. The finite element code used was ABAQUS, version 5.7 [7]. Fig. 7 shows a finite element model and the boundary conditions of the analyses. The model includes six armor tiles with three tungsten tiles and three 1D-CFC tiles, which corresponds to the central part of the mock-up. Eight node-linear elements were used in the thermal and stress analyses. In the thermal analysis, uniform heat flux of 5 MW/m2 onto the surface of all armor tiles was assumed to simulate the experiment. Based on the temperature evolution obtained in this transient thermal analysis, subsequent elastoplastic stress analyses were performed. Prior to the elastoplastic stress analysis assuming the cyclic thermal loading, a residual stress analysis was conducted to obtain the initial stress field of the model. The residual stress behaved as mean stress during the cyclic thermal loading. The reference temperature of the residual stress analysis was 780°C, which corresponded to the solidified temperature of the Ti–Cu–Ag braze material. In the stress analysis, elastoplastic behavior of DSCu and OFHC-Cu was taken into account; the material properties of OFHC-Cu were assumed to be of a fully annealed material. Fig. 8 shows the equivalent mechanical strain
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Fig. 7. Finite element mesh and boundary condition (line ‘A’ corresponds to the top of the outer surface of the tube).
profile along the line ‘A’ corresponding to the top of the outer surface of the cooling tube. Large strain concentration appeared at the top of the tube locating in between the armor tiles. Particularly, the largest equivalent mechanical strain of 6.6% was obtained at the tube adjoining the tungsten-armored side and the CFC-armored side. The result of the residual stress analysis shows that the stress field of the OFHC-Cu layer is in an elastoplastic state after the braze process. Subsequently, the elastoplastic stress analysis simulating the experiment was carried out using this residual stress field as an initial condition. In the analysis, three thermal cycles were assumed to stabilize the stress–strain behavior of the model. Fig. 8 also shows the equivalent mechanical strain amplitude of the cooling tube obtained at the third thermal cycle. The maximum equivalent mechanical strain amplitude of 5.2% appeared at the tube adjoining the tungsten-armored side and the CFC-armored side. As a result of these stress analyses, it was found that the low cycle fatigue fracture was caused by the large strain concentration and by the large strain amplitude of the cooling tube. Moreover, the geometric discontinuity formed by the upper heat sink, which consists of a single OFHC-Cu plate, caused the strain concentration because the peak of the strain profile at the CFC-armored side along the line ‘A’ was
not as significant as that of the crack region. To avoid such fatigue cracking, separation of the upper heat sink is essential for this divertor structure.
Fig. 8. Strain distribution along line ‘A’ after braze process and mechanical strain amplitude at the end of the third cycle.
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4. Conclusion
Acknowledgements
Based on this study, we draw the following conclusions: 1. Thermal fatigue experiment of the ITER divertor (vertical target) mock-up was performed in PBEF. The DSCu cooling tube with triplex layer showed low cycle fatigue fracture at 400 thermal cycles under a cyclic heat flux of 5 MW/m2. 2. The fatigue crack was found at the cooling tube adjoining the tungsten-armored side and the CFC-armored side. According to the posttest microscopic observation, the DSCu layer showed a brittle fracture surface whereas the OFHC-Cu layer of the tube showed a clear evidence of the fatigue fracture. Since DSCu was proved to be a crack-sensitive material in this study, precise evaluation on the fatigue crack initiation/growth is necessary to use DSCu as a cooling tube material for the divertor application. 3. It was found that the geometric discontinuity formed by the single plate heat sink strongly affects the fatigue lifetime of the cooling tube. For the design of the divertor plate, separation of the heat sink for each armor tile is essential to avoid the large strain concentration at the cooling tube.
The authors wish to thank Y. Okumura and other members of NBI heating laboratory for their valuable discussions and comments. They would also like to acknowledge M. Ohta and S. Matsuda for their support and encouragement.
References [1] G. Vieider, V. Barabash, et al., Overview of the EU small scale mock-up tests for ITER high heat flux components, Fusion Eng. Des. 39 – 40 (1998) 211 – 218. [2] M. Akiba, S. Suzuki, Overview of the Japanese mock-up tests for ITER high eat flux components, Fusion Eng. Des. 39 – 40 (1998) 219 – 225. [3] I. Mazul, R. Giniatulin, et al., Manufacturing and testing of ITER divertor gas box liners, Proceedings of the 20th Symposium on Fusion Technology, vol. A, 1998, pp. 77–80. [4] S. Suzuki, T. Suzuki, et al., Development of divertor plate with CFCs bonded onto DSCu cooling tube for fusion reactor application, J. Nucl. Mater. 258 – 263 (1998) 318– 322. [5] ITER-EDA, Technical basis for the ITER Final Design Report, cost review and safety analysis, ITER EDA documentation series No. 16, 1998. [6] S. Suzuki, T. Suzuki, Development of divertor high heat flux components at JAERI, Proceedings of the 17th IEEE/NPSS Symposium on Fusion Engineering, vol. 1, 1977, pp. 385– 388. [7] ABAQUS/Standard user’s manual version 5.7, Hibbitt, Karlsson & Sorensen, 1999.