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Non-destructive testing of divertor components
Fusion Engineering and Design 61 /62 (2002) 141 /146 www.elsevier.com/locate/fusengdes
Non-destructive testing of divertor components M. Merola a, , P. Chappuis b, F. Escourbiac b, M. Grattarola c, H. Jeskanen d, P. Kauppinen d, L. Plo¨chl e, B. Schedler e, J. Schlosser b, I. Smid f, S. Ta¨htinen d, R. Vesprini g, E. Visca h, A. Zabernig e a
EFDA Close Support Unit, Boltzmannstrasse 2, D-85748 Garching, Germany b CEA Cadarache, France c Ansaldo Ricerche, Genoa, Italy d VTT, Espoo, Finland e Plansee AG, Reutte, Austria f Seibersdorf Research Centre, Austria g ENEA Casaccia, Italy h ENEA Frascati, Italy
Abstract This task within the EU R&D for ITER had two main objectives: (1) qualification of inspection procedures for plasma facing components (PFC), (2) assessment of the behaviour of calibrated defects under high heat flux (HHF) cyclic loading. The ultimate goal of this work was to demonstrate that the reliable identification of fatal defects by the chosen non-destructive testing (NDT) methods can be achieved. This R&D was carried out according to the following steps: (1) manufacture of a divertor vertical target (VT) prototype with artificial calibrated defects; (2) blind nondestructive round robin test of the prototype; (3) HHF test in FE200 electron beam (EB) facility; (4) post-fatigue blind non-destructive round robin test; (5) destructive examination. The general final conclusion was that the NDT techniques can reliably detect and locate defects having dimensions well below those, which could impair the thermal fatigue lifetime. # 2002 Elsevier Science B.V. All rights reserved. Keywords: Non-destructive testing; Thermal fatigue; Electron beam facility
1. Introduction The overall goal of the non-destructive testing (NDT) development programme for ITER is to demonstrate that significant defects can be identi-
Corresponding author. Tel.: /49-89-3299-4220; fax: /4989-3299-4198 E-mail address: [email protected] (M. Merola).
fied without recourse to performing high heat flux (HHF) screening tests on the entire surface of each plasma facing component (PFC). In fact, to confidently satisfy lifetime, quality and safety requirements, it is necessary to have qualified NDT processes for all joints in the PFCs, namely: . carbon fibre reinforced carbon (CfC) to copper (Cu); . tungsten to Cu;
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M. Merola et al. / Fusion Engineering and Design 61 /62 (2002) 141 /146
. Cu to stainless steel; . stainless steel to stainless steel. This activity, within the EU R&D for ITER, had the following main aims: . Qualification of inspection procedures for PFCs. . Assessment of the behaviour of calibrated defects under HHF cyclic loading. . Demonstration that the smallest detectable defects do not impair the component performances appreciably. . Set up recommendations on the most suitable inspection method for each type of joint. The chosen joints for development and demonstration of NDT were: . CfC/Cu Active Metal Cast† for monoblocktype armour. . Cu/Cu-alloy brazing for monoblock-type armour with the cooling tube. . Cu/Cu-alloy electron beam (EB) for macrobrush-type tungsten armour with the heat sink. The main steps of this task were the following: . Manufacture of a vertical target (VT) prototype with artificial calibrated defects. . Blind NDT round robin tests of the prototype. . HHF fatigue test. . Post-fatigue blind NDT round robin test of the prototype. . Destructive analysis of the prototype.
2. Manufacture of the vertical target prototype A prototypical divertor VT, which included all the main features of the corresponding ITER divertor design, was manufactured with calibrated joint defects [1]. This VT prototype consisted in one unit with one cooling channel. The overall length and width was about 600 and 25 mm, respectively (Fig. 1). The HHF part and the 316L steel back plate were manufactured by the EU companies Plansee
Fig. 1. EU VT prototype with artificial calibrated defects.
AG and Ansaldo Ricerche, respectively. The integration between the two parts was carried out by Ansaldo Ricerche. The HHF part had a lower straight region made of CfC monoblocks and an upper bend region with tungsten macro-brush armour. The heat sink material was dispersion strengthened copper (DS /Cu) alloy by OMG Americas (GlidCop AL25, Cu/0.25% Al2O3). In the upper region, it consisted of a hollow bar (obtained by two half shells) where a DS /Cu tube (10/12 mm ID/OD) was inserted. The tube then continued in the lower monoblock region. The transition between the DS/Cu tube and the stainless steel tube was performed via a nickel sleeve EB welded at both ends. The steel tube was then TIG welded to the steel back plate on the top of the component and to the inlet and outlet flanges on the bottom. Thus, the steel back plate was actively cooled by the return water flow from the HHF part. A twisted tape with a twist ratio of 2 was inserted in the monoblock region to enhance the critical heat flux limit. The armour of the upper region was made of tungsten (1% La2O3). The thickness was 10 mm. A pure copper interlayer was inserted between the tungsten armour and the DS /Cu heat sink to alleviate the joint interface stress due to the thermal expansion mismatch. The lower region of the VT (CfC monoblock) was integrated onto the steel back plate via a dovetail mechanical attachment, which allowed the monoblocks to slide. The carbon material was 3D CfC (NB31TM manufactured by SNECMA) which was joined onto the DS /Cu
M. Merola et al. / Fusion Engineering and Design 61 /62 (2002) 141 /146
tube by Active Metal Cast† . Then the copper interlayer was brazed onto the DS /Cu tube at 880 8C by means of a Ti-based eutectic. The 11 CfC armour tiles had a thickness of 14 mm. Three different types of artificial defects were introduced: . Defects in the EB welded joint of the tungsten macro-brush. These defects had a rectangular cross section with the following dimensions: 3 /11, 6 /5, 6 /11, 3 /5 mm. They were obtained by machining from the side surface. . Defects in the CfC/Cu Active Metal Cast† joint and in the Cu/Cu-alloy brazed joint for monoblock-type armour with the cooling tube. These defects had a rectangular cross section (3 and 6 mm wide, 20 mm long) or a circular cross section (6 mm diameter). Since Cu does not wet CfC, the defects in the CfC/Cu joint were obtained by removing the activated surface of CfC (about 0.5 mm) prior to the Cu casting. The defects in the brazed Cu/Cu-alloy joint were obtained by applying a braze stopper onto the two surfaces to be joined prior to brazing.
3. Thermal fatigue testing After the first blind NDT round robin test, the VT prototype has been subjected to cyclic heat loads in order to assess the behaviour of the calibrated defects. Table 1 summarises the testing procedure; further details on the testing results can be found in Ref. [2]. After the HHF testing, a
Table 1 HHF testing procedure of the VT prototype with artificial defects CfC monoblock
Tungsten macro-brush
Initial screening up to 12 MW/m2 500 cycles at 12 MW/m2 500 cycles at 10 MW/m2 Intermediate screening at 12 MW/m2 500 cycles at 12 MW/m2 500 cycles at 15 MW/m2 2 Final screening at 12 MW/m
143
second blind NDT round robin test was carried out.
4. Non-destructive inspection Fig. 2 shows a comparison of the C-scans of the CfC to tube joint carried out by VTT, Seibersdorf and ENEA. This type of scan gives the extension of the defect on the joint surface. One can note that the C-scans are rather similar even if the chosen ultrasonic frequency varies from 10 to 25 MHz. The defects in the tube/Cu brazed joints can be identified and sized reliably, whereas those in the CfC/Cu joints can hardly be detected. This latter result can be easily explained by two reasons: (1) the significant difference in the acoustic impedance of carbon and Cu which leads to a high ultrasonic echo even if the joint is sound, (2) the corrugated joint surface which diffuses the ultrasonic echo in all the directions. In any case, all these C-scans remained practically unchanged before and after the HHF test. Fig. 3 shows a comparison between B- and Cscans. The B-scan gives information on the depth in which a defect is located. It is worth noting the undulation of the inner surface of the tube as a result of the manufacturing process (peak-tovalley distance about 0.1 mm). It is also possible to identify whether a defect is located in the CfC/ Cu or in the brazed joint. Fig. 4 shows a comparison between the infrared thermography and the surface temperature during the HHF test. One can notice that defects in the CfC/Cu joint have a more pronounced effect than those in the brazed joint. In fact the high thermal conductivity Cu material can better smooth the temperature perturbation due to the presence of embedded defects. Fig. 5 compares the C-scans of the EB welding of the Cu/Cu alloy joint in the tungsten macrobrush region. The inspections could only be carried out from the tungsten macro-brush side. It can be noticed that the sizing of the defects is rather poor since the ultrasonic examination is disturbed by the presence of the macro-brush. In
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Fig. 2. Comparison of different C-scans of the CfC/tube joint. From left to right; defects in the Cu/Cu-alloy tube brazed joint (dotted) and in the CfC/Cu joint (solid), C-scan by VTT (25 MHz), C-scan by Seibersdorf (15 MHz), C-scan by ENEA (10 MHz).
this case also the infrared examination did not give valuable results. On the other hand the surface temperature during the HHF test was not influenced appreciably by the presence of any of the artificial defects nor the thermal fatigue lifetime appeared to be impaired.
5. Conclusions All the details on the NDT procedures can be found in Refs [3 /5]. Here the main conclusions are summarised. . Ultrasonic inspections of the tube/Cu brazed joint proved to be able to locate and size defects of 2 /3 mm reliably. . B-scan proved to be useful to better evaluate the C-scan indications since it gave information on whether the defects are located in the Active Metal Cast† joint or in the brazed joint.
. Defects in the CfC material and/or in the Active Metal Cast have a stronger influence on the surface temperature than defects in the tube/Cu brazed joint. . The behaviour of a component under HHF loading is due to a combination of defects in the tube/Cu joint, in the Active Metal Cast and in the CfC armour. Only the first ones can be reliably detected by ultrasounds. . Infrared thermography proved to be an important and complementary inspection technique able to well predict the component behaviour under HHF loading. . However, the defect detection capability of the infrared thermography should be improved as far as tungsten armoured components are concerned. . Ultrasonic inspections of the EB welding of the tungsten macro-brush appeared more difficult but defects having all the dimensions greater than 3 mm could be identified reliably.
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Fig. 4. Comparison between infrared thermography (centre) and surface temperature during the heat flux test (right).
. The general final conclusion is that the NDT technique can reliably detect and locate defects having dimensions well below those, which could impair the thermal fatigue lifetime.
Fig. 3. Comparison between B- and C-scans of the CfC/tube joint. (ENEA, 10 MHz).
. The sizing of these defects seemed to require high frequencies but in this case the presence of W tiles may disturb the inspection. . However, the presence of large defects (up to 6 /11 mm) in the EB welding does not appear to impair appreciably the fatigue lifetime of the component nor its surface temperature. . All the investigated defects proved to be very forgiving with respect to the HHF loading in spite of their relatively large dimensions. This was confirmed by the final destructive examination of the component.
Fig. 5. Comparison of different C-scan of the EB Cu/Cu alloy joint. From left to right; defects in the EB welding between the Cu interlayer and the Cu alloy heat sink, C-scan by VTT (25 MHz), C-scan by Plansee (5 MHz), C-scan by ENEA (10 MHz).
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References [1] M. Merola, L. Plo¨chl, P. Chappuis, F. Escourbiac, M. Grattarola, I. Smid, R. Tivey, G. Vieider, Manufacturing and testing of a prototypical divertor vertical target for ITER, Journal of Nuclear Materials 283 /287 (2000) 1068 / 1072. [2] F. Escourbiac, P. Chappuis, J. Schlosser, M. Merola, I. Vastra, M. Febvre, High heat flux behaviour of damaged plasma facing components, Fusion Engineering and Design 56 /57 (2001) 285 /290.
[3] F. Escourbiac, J. Schlosser, Critical heat flux and thermal hydraulic of representative elements, non destructive testing, calibrated defects, heat load influence, CEA Report CFP/NTT-2000.030, December 2000. [4] E. Visca, Report on NDT of components with calibrated defects and destructive examination after high heat flux testing, ENEA Report FUS-TN-GE-VD-Q-001, May 2001. [5] S. Ta¨htinen, H. Jeskanen, P. Kauppinen, Ultrasonic and metallographic examination of ITER vertical target prototype, VTT Report BVAL62-011099, March 2001.
Non-destructive testing of divertor components M. Merola a, , P. Chappuis b, F. Escourbiac b, M. Grattarola c, H. Jeskanen d, P. Kauppinen d, L. Plo¨chl e, B. Schedler e, J. Schlosser b, I. Smid f, S. Ta¨htinen d, R. Vesprini g, E. Visca h, A. Zabernig e a
EFDA Close Support Unit, Boltzmannstrasse 2, D-85748 Garching, Germany b CEA Cadarache, France c Ansaldo Ricerche, Genoa, Italy d VTT, Espoo, Finland e Plansee AG, Reutte, Austria f Seibersdorf Research Centre, Austria g ENEA Casaccia, Italy h ENEA Frascati, Italy
Abstract This task within the EU R&D for ITER had two main objectives: (1) qualification of inspection procedures for plasma facing components (PFC), (2) assessment of the behaviour of calibrated defects under high heat flux (HHF) cyclic loading. The ultimate goal of this work was to demonstrate that the reliable identification of fatal defects by the chosen non-destructive testing (NDT) methods can be achieved. This R&D was carried out according to the following steps: (1) manufacture of a divertor vertical target (VT) prototype with artificial calibrated defects; (2) blind nondestructive round robin test of the prototype; (3) HHF test in FE200 electron beam (EB) facility; (4) post-fatigue blind non-destructive round robin test; (5) destructive examination. The general final conclusion was that the NDT techniques can reliably detect and locate defects having dimensions well below those, which could impair the thermal fatigue lifetime. # 2002 Elsevier Science B.V. All rights reserved. Keywords: Non-destructive testing; Thermal fatigue; Electron beam facility
1. Introduction The overall goal of the non-destructive testing (NDT) development programme for ITER is to demonstrate that significant defects can be identi-
Corresponding author. Tel.: /49-89-3299-4220; fax: /4989-3299-4198 E-mail address: [email protected] (M. Merola).
fied without recourse to performing high heat flux (HHF) screening tests on the entire surface of each plasma facing component (PFC). In fact, to confidently satisfy lifetime, quality and safety requirements, it is necessary to have qualified NDT processes for all joints in the PFCs, namely: . carbon fibre reinforced carbon (CfC) to copper (Cu); . tungsten to Cu;
0920-3796/02/$ - see front matter # 2002 Elsevier Science B.V. All rights reserved. PII: S 0 9 2 0 - 3 7 9 6 ( 0 2 ) 0 0 1 5 5 - 2
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M. Merola et al. / Fusion Engineering and Design 61 /62 (2002) 141 /146
. Cu to stainless steel; . stainless steel to stainless steel. This activity, within the EU R&D for ITER, had the following main aims: . Qualification of inspection procedures for PFCs. . Assessment of the behaviour of calibrated defects under HHF cyclic loading. . Demonstration that the smallest detectable defects do not impair the component performances appreciably. . Set up recommendations on the most suitable inspection method for each type of joint. The chosen joints for development and demonstration of NDT were: . CfC/Cu Active Metal Cast† for monoblocktype armour. . Cu/Cu-alloy brazing for monoblock-type armour with the cooling tube. . Cu/Cu-alloy electron beam (EB) for macrobrush-type tungsten armour with the heat sink. The main steps of this task were the following: . Manufacture of a vertical target (VT) prototype with artificial calibrated defects. . Blind NDT round robin tests of the prototype. . HHF fatigue test. . Post-fatigue blind NDT round robin test of the prototype. . Destructive analysis of the prototype.
2. Manufacture of the vertical target prototype A prototypical divertor VT, which included all the main features of the corresponding ITER divertor design, was manufactured with calibrated joint defects [1]. This VT prototype consisted in one unit with one cooling channel. The overall length and width was about 600 and 25 mm, respectively (Fig. 1). The HHF part and the 316L steel back plate were manufactured by the EU companies Plansee
Fig. 1. EU VT prototype with artificial calibrated defects.
AG and Ansaldo Ricerche, respectively. The integration between the two parts was carried out by Ansaldo Ricerche. The HHF part had a lower straight region made of CfC monoblocks and an upper bend region with tungsten macro-brush armour. The heat sink material was dispersion strengthened copper (DS /Cu) alloy by OMG Americas (GlidCop AL25, Cu/0.25% Al2O3). In the upper region, it consisted of a hollow bar (obtained by two half shells) where a DS /Cu tube (10/12 mm ID/OD) was inserted. The tube then continued in the lower monoblock region. The transition between the DS/Cu tube and the stainless steel tube was performed via a nickel sleeve EB welded at both ends. The steel tube was then TIG welded to the steel back plate on the top of the component and to the inlet and outlet flanges on the bottom. Thus, the steel back plate was actively cooled by the return water flow from the HHF part. A twisted tape with a twist ratio of 2 was inserted in the monoblock region to enhance the critical heat flux limit. The armour of the upper region was made of tungsten (1% La2O3). The thickness was 10 mm. A pure copper interlayer was inserted between the tungsten armour and the DS /Cu heat sink to alleviate the joint interface stress due to the thermal expansion mismatch. The lower region of the VT (CfC monoblock) was integrated onto the steel back plate via a dovetail mechanical attachment, which allowed the monoblocks to slide. The carbon material was 3D CfC (NB31TM manufactured by SNECMA) which was joined onto the DS /Cu
M. Merola et al. / Fusion Engineering and Design 61 /62 (2002) 141 /146
tube by Active Metal Cast† . Then the copper interlayer was brazed onto the DS /Cu tube at 880 8C by means of a Ti-based eutectic. The 11 CfC armour tiles had a thickness of 14 mm. Three different types of artificial defects were introduced: . Defects in the EB welded joint of the tungsten macro-brush. These defects had a rectangular cross section with the following dimensions: 3 /11, 6 /5, 6 /11, 3 /5 mm. They were obtained by machining from the side surface. . Defects in the CfC/Cu Active Metal Cast† joint and in the Cu/Cu-alloy brazed joint for monoblock-type armour with the cooling tube. These defects had a rectangular cross section (3 and 6 mm wide, 20 mm long) or a circular cross section (6 mm diameter). Since Cu does not wet CfC, the defects in the CfC/Cu joint were obtained by removing the activated surface of CfC (about 0.5 mm) prior to the Cu casting. The defects in the brazed Cu/Cu-alloy joint were obtained by applying a braze stopper onto the two surfaces to be joined prior to brazing.
3. Thermal fatigue testing After the first blind NDT round robin test, the VT prototype has been subjected to cyclic heat loads in order to assess the behaviour of the calibrated defects. Table 1 summarises the testing procedure; further details on the testing results can be found in Ref. [2]. After the HHF testing, a
Table 1 HHF testing procedure of the VT prototype with artificial defects CfC monoblock
Tungsten macro-brush
Initial screening up to 12 MW/m2 500 cycles at 12 MW/m2 500 cycles at 10 MW/m2 Intermediate screening at 12 MW/m2 500 cycles at 12 MW/m2 500 cycles at 15 MW/m2 2 Final screening at 12 MW/m
143
second blind NDT round robin test was carried out.
4. Non-destructive inspection Fig. 2 shows a comparison of the C-scans of the CfC to tube joint carried out by VTT, Seibersdorf and ENEA. This type of scan gives the extension of the defect on the joint surface. One can note that the C-scans are rather similar even if the chosen ultrasonic frequency varies from 10 to 25 MHz. The defects in the tube/Cu brazed joints can be identified and sized reliably, whereas those in the CfC/Cu joints can hardly be detected. This latter result can be easily explained by two reasons: (1) the significant difference in the acoustic impedance of carbon and Cu which leads to a high ultrasonic echo even if the joint is sound, (2) the corrugated joint surface which diffuses the ultrasonic echo in all the directions. In any case, all these C-scans remained practically unchanged before and after the HHF test. Fig. 3 shows a comparison between B- and Cscans. The B-scan gives information on the depth in which a defect is located. It is worth noting the undulation of the inner surface of the tube as a result of the manufacturing process (peak-tovalley distance about 0.1 mm). It is also possible to identify whether a defect is located in the CfC/ Cu or in the brazed joint. Fig. 4 shows a comparison between the infrared thermography and the surface temperature during the HHF test. One can notice that defects in the CfC/Cu joint have a more pronounced effect than those in the brazed joint. In fact the high thermal conductivity Cu material can better smooth the temperature perturbation due to the presence of embedded defects. Fig. 5 compares the C-scans of the EB welding of the Cu/Cu alloy joint in the tungsten macrobrush region. The inspections could only be carried out from the tungsten macro-brush side. It can be noticed that the sizing of the defects is rather poor since the ultrasonic examination is disturbed by the presence of the macro-brush. In
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Fig. 2. Comparison of different C-scans of the CfC/tube joint. From left to right; defects in the Cu/Cu-alloy tube brazed joint (dotted) and in the CfC/Cu joint (solid), C-scan by VTT (25 MHz), C-scan by Seibersdorf (15 MHz), C-scan by ENEA (10 MHz).
this case also the infrared examination did not give valuable results. On the other hand the surface temperature during the HHF test was not influenced appreciably by the presence of any of the artificial defects nor the thermal fatigue lifetime appeared to be impaired.
5. Conclusions All the details on the NDT procedures can be found in Refs [3 /5]. Here the main conclusions are summarised. . Ultrasonic inspections of the tube/Cu brazed joint proved to be able to locate and size defects of 2 /3 mm reliably. . B-scan proved to be useful to better evaluate the C-scan indications since it gave information on whether the defects are located in the Active Metal Cast† joint or in the brazed joint.
. Defects in the CfC material and/or in the Active Metal Cast have a stronger influence on the surface temperature than defects in the tube/Cu brazed joint. . The behaviour of a component under HHF loading is due to a combination of defects in the tube/Cu joint, in the Active Metal Cast and in the CfC armour. Only the first ones can be reliably detected by ultrasounds. . Infrared thermography proved to be an important and complementary inspection technique able to well predict the component behaviour under HHF loading. . However, the defect detection capability of the infrared thermography should be improved as far as tungsten armoured components are concerned. . Ultrasonic inspections of the EB welding of the tungsten macro-brush appeared more difficult but defects having all the dimensions greater than 3 mm could be identified reliably.
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Fig. 4. Comparison between infrared thermography (centre) and surface temperature during the heat flux test (right).
. The general final conclusion is that the NDT technique can reliably detect and locate defects having dimensions well below those, which could impair the thermal fatigue lifetime.
Fig. 3. Comparison between B- and C-scans of the CfC/tube joint. (ENEA, 10 MHz).
. The sizing of these defects seemed to require high frequencies but in this case the presence of W tiles may disturb the inspection. . However, the presence of large defects (up to 6 /11 mm) in the EB welding does not appear to impair appreciably the fatigue lifetime of the component nor its surface temperature. . All the investigated defects proved to be very forgiving with respect to the HHF loading in spite of their relatively large dimensions. This was confirmed by the final destructive examination of the component.
Fig. 5. Comparison of different C-scan of the EB Cu/Cu alloy joint. From left to right; defects in the EB welding between the Cu interlayer and the Cu alloy heat sink, C-scan by VTT (25 MHz), C-scan by Plansee (5 MHz), C-scan by ENEA (10 MHz).
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References [1] M. Merola, L. Plo¨chl, P. Chappuis, F. Escourbiac, M. Grattarola, I. Smid, R. Tivey, G. Vieider, Manufacturing and testing of a prototypical divertor vertical target for ITER, Journal of Nuclear Materials 283 /287 (2000) 1068 / 1072. [2] F. Escourbiac, P. Chappuis, J. Schlosser, M. Merola, I. Vastra, M. Febvre, High heat flux behaviour of damaged plasma facing components, Fusion Engineering and Design 56 /57 (2001) 285 /290.
[3] F. Escourbiac, J. Schlosser, Critical heat flux and thermal hydraulic of representative elements, non destructive testing, calibrated defects, heat load influence, CEA Report CFP/NTT-2000.030, December 2000. [4] E. Visca, Report on NDT of components with calibrated defects and destructive examination after high heat flux testing, ENEA Report FUS-TN-GE-VD-Q-001, May 2001. [5] S. Ta¨htinen, H. Jeskanen, P. Kauppinen, Ultrasonic and metallographic examination of ITER vertical target prototype, VTT Report BVAL62-011099, March 2001.