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Nuclear analyses of some key aspects of the ITER design with Monte Carlo codes
Fusion Engineering and Design 74 (2005) 133–139
Nuclear analyses of some key aspects of the ITER design with Monte Carlo codes H. Iida a,∗ , L. Petrizzi b , V. Khripunov c , G. Federici d , E. Polunovskiy d a ITER International Team, Naka Joint Work Site, Naka-machi, Naka-gun, Ibaraki-ken, Japan ENEA FUS-TEC Frascati, Sezione Tecnologie della Fusione, Via E. Fermi 45, 00044 Frascati, Italy c Nuclear Fusion Institute, Russian Research Center “Kurchatov Institute”, 123182 Moscow, Russia ITER International Team, Garching Joint Work Site, Max-Planck-Institut fuer Plasmaphysik, Boltzmannstr. 2, D-85748 Garching, Germany b
d
Available online 10 August 2005
Abstract The design of the ITER machine was presented in 2001 [1,2]. A nuclear analysis was performed at this time, using fairly detailed models and the best assessed nuclear data and codes that were available. As the construction phase of ITER is approaching, the design of the main components has been optimized/finalized and several minor design changes/optimizations have been made, some with the object to mitigate critical radiation shielding problems. These have required refined calculations to confirm that the nuclear design requirements are met. This paper reviews some of the most recent neutronic work with emphasis on critical nuclear responses in the TF coil inboard legs and vacuum vessel related to design modifications made to the blanket modules and vacuum vessel. © 2005 Elsevier B.V. All rights reserved. Keywords: ITER; Monte Carlo method; TF coil; Nuclear heating; Vacuum vessel; Helium production
1. Introduction Critical design issues related to the inboard radiation shielding capability are: the reduction of heat loads in the TF coils and the vacuum vessel to a sufficiently low level; the protection of the TF coil winding packs from radiation damage and the reduction of helium production at the location of field-joint between the ∗ Corresponding author. Tel.: +81 29 270 7730; fax: +81 29 270 7506. E-mail address: [email protected] (H. Iida).
0920-3796/$ – see front matter © 2005 Elsevier B.V. All rights reserved. doi:10.1016/j.fusengdes.2005.06.239
vacuum vessel segments. The main factors affecting the shielding capability where design optimisation is still possible are the toroidal and poloidal gaps between the blanket sectors. These are required to allow fitting of the sectors by remote handling but have a sensitive impact on the overall shielding capability. 1.1. Nuclear heat and damage of winding packs in the TF coils In the original design, the nuclear heat in TF coils was analysed and reported to be 12.4 kW in the
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H. Iida et al. / Fusion Engineering and Design 74 (2005) 133–139
Nuclear Analysis Report-2001(NAR-2001) [3,4] with 2 cm blanket gaps, while the cooling and refrigerator system of the TF coils were designed to remove up to ∼14 kW of nuclear heating at the reference pulse rate (2500 MW plasma pulses per hour). The calculation results of the nuclear heat have a certain degree of uncertainty, which is estimated to be of the order of ±70% on the inboard and ±80% on the outboard [3]. Consequently, there was a level of risk that the nuclear heat could exceed the design values for the refrigerator and coil cooling systems. Operating in this condition would have created a much higher cryogenic pumping work and would have been a very inefficient operating mode for the cryogenic heat extraction. Damage of the epoxy insulator in the winding packs was also at the limit, since the acceptable fast neutron fluence for some of the epoxy insulation systems being considered for the TF coils has turned out to be 5 × 1021 n/m2 , which is a half of the specified value.
A preliminary (1-D) calculation of the insulator damage documented in the NAR-2001 is 2.6 × 1021 n/m2 (ground insulation) and 2.3 × 1021 n/m2 (between first and second turns) with the assumption of 2 cm wide gaps between inboard blanket modules. These fluence values are close to the limit for the less radiation resistant epoxy insulator requiring a detailed 3-D analysis.
1.2. Helium production and heat loads at the surface of the vacuum vessel Due to radiation streaming through the gaps between blanket modules, locally peaked helium production and nuclear heat rates are expected on the vacuum vessel inner shell. The field joints require rewelding when replacement of the vessel segments is needed. The quality of such re-welding is guaranteed only when the helium production is less than 1 appm.
Fig. 1. Vertical cross section of the ITER 3-D model for the MCNP.
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When the heating density is larger than ∼0.3 W/cm3 , the thermal stress in the vacuum vessel can be a serious design issue. During the last few years, the design of some invessel components has undergone changes. Among the design changes, the reduction of the gap width between the blanket modules has a major impact on the radiation quantities above. This increases remote handling problem of the blanket modules but brings more margin, with more design flexibility, for the coils. Below, a summary of most recent nuclear analyses concerning these changes is presented.
2. Calculation tools and computing system The Monte Carlo transport code MCNP [5], nuclear data library FENDL1/MC and geometry input data of “ITER 3-D basic model” have been used for these analyses. After setting proper “weight window parameters”, most of calculation cases were finished by a 3 h run with 128 cpu (MPI) in the Origin3800 for obtaining acceptable statistics. Some cases for calculating very local peak values required ∼10 h to obtain sufficient statistics even with optimized weight windows.
3. TF coil nuclear heat and insulator damage 3.1. Nuclear heat The vertical cross section of the model is shown in Fig. 1. A TF coil is divided into 50 segments in the poloidal direction. The straight part of TF coils, which consists of segments from #1 to #14, will receive a major part (>70%) of the total heat and is defined here as “inboard leg of TF coil”. Shielding thickness for this part is minimal to reduce the size (hence cost) of the ITER machine.
Fig. 2. Nuclear heat poloidal distribution in the TF coil inboard leg. Case 1: vertical gaps 2.0 cm, all horizontal gaps on inboard 2.0 cm (2001 design), case 2: vertical gaps 1.4 cm, all horizontal gaps on inboard 1.0 cm (present design), case 3: vertical gaps1.4 cm, three horizontal gaps (#2/#3, #3/#4, #4/#5) near the equatorial plane 1.0 cm, other horizontal gaps 2.0 cm (back up option) and case 4: vertical gaps 1.4 cm, all horizontal gaps on inboard 2.0 cm.
Fig. 2 shows the poloidal distribution of the nuclear heat in the inboard TF coil legs. The nuclear heat is naturally high at the mid plane (segment #8). For the current design (case 2 in Table 1) the nuclear heat is lower by a factor of about 25% with respect to the 2001 design (case 1). Detailed assessment of the remote handing and operational displacements of the sectors confirms that the width can be reduced from 2 to 1.4 cm for the vertical gaps (the nuclear heat value for this step, with the original 2 cm horizontal gap, is shown as case 4). The reduction of the width of the horizontal gaps requires further work since the remote handling feasibility for the narrower gaps has to be confirmed by additional R&D. Narrower gaps also give some constraints on installation of diagnostic equipment. Case
Table 1 Nuclear heat in the TF coil inboard legs
Case 1 Case 2 Case 3 Case 4 a
Verical gap width (cm)
Horizontal gap width (cm)
Inboard leg (kW)
Total in super conducting coil (kW)a
2.0 1.4 1.4 1.4
2.0 1.0 Central 3 gaps 1.0, other gaps 2.0 2.0
10.0 7.51 7.78 9.14
12.4 9.91 10.2 11.5
Inboard leg + 2.4 kW.
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3, which has narrowed width only on the gaps near the mid plane and gives slightly higher heat than the case 2, is a back up option in case there are constraints in narrowing the gap width as a result of future R&D for remote handling technology. The total nuclear heat value of 9.91 kW (case 2) gives a much larger safety margin to cover some of the uncertainty in the calculation results than the 2001 design (case 1). A safety factor of 40% is now allowed on the calculated values in the design value of 14 kW assumed for nuclear heat in the coils, at which the design pulse rates can be maintained. Fig. 3 shows the uncertainty levels the present design will have [6]. This figure assumes uncertainties of 3-D modeling (30%) in the design calculation and of benchmark experiments (20%) at FNG (ENEA Frascati) [7] and FNS (JAERI) [8]. The average of the “most probable values” (µ1) is assumed to be larger than that of the “best estimates with calculation code” (µ2) by a factor of 1.15 [see ref. [6] in detail]. According to this figure, the probability of the nuclear heat being higher than 14 kW is ∼30% (R = 14 kW/9.91 kW = 1.4). This risk level has been accepted, since the coil design specification still allows for nuclear heating up to the full uncertainty range (a factor of 2, 18 kW) by permitting a reduction in plasma pulse rate, or an increase in the refrigerator capacity. The coil cooling design has been modified to allow a higher rate of heat extraction with a lower cost in cryogenic pumping power. 3.2. Fast neutron fluence on the epoxy insulator The maximum fast neutron fluence on the inboard legs occurs on the #8 segment, as easily expected from the nuclear heat distribution. The relation between
Fig. 3. Probability function of between calculated and probable 2 2 2 2 1 P(R) dR = · e−(Rµ1 −µ2 ) /2(σ2 +R σ1 ) 2πσ1 σ2 1 1 −β2 · 2 e + β{1 + ERF(β)} dR α 2
the ratio R nuclear heating
α=
, σ22 + Rσ12 σ 2 µ1 + Rσ 2 µ2 ,β= √ 2 1 , √ 2 2 2 2σ1 σ2 2σ1 σ2 σ2 + R σ1
β
2 t −2 dt ERF(β) = √ π 0 where fsd1 and fsd2 : fractional standard deviations of the calculated and probable values.
the maximum fluence and blanket module gap width is shown in Table 2. Considering the uncertainty of about a factor 2 in the calculated values, fluence values in the table should be less than 2.5 × 1021 n/m2 , which would allow the insulation requirement to be reduced to 5 × 1021 n/m2 when required. The fluence on the insulator between #1 and #2 turns is lower than
Table 2 Maximum fast neutron fluence on the epoxy insulator in the inboard TF coil leg
Case a Case b Case c Case d Case e Case f Case g
Vertical gap width (cm)
Horizontal gap width (cm)
Fast neutron fluence in #8 segment (1021 n/m2 ) at the end of life Ground insulator
Between first and second turns
2 1 1.4 1.4 1.4 1.4 1.4
2 1 2 1.5 1 0.5 0
5.02 3.02 4.50 3.81 3.25 2.88 2.71
2.73 1.64 2.45 2.07 1.76 1.56 1.46
End of life: 0.3 MW/m2 at the first wall (average).
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Fig. 4. Rear side view of the inboard blanket module with cooling water manifolds on the vacuum vessel surface.
2.5 × 1021 n/m2 in all cases with 1.4 cm of the vertical gap width, although those on the plasma side ground insulator can go up to 4.5 × 1021 n/m2 . The high fluence values on the plasma side ground insulator are judged to be acceptable since this part is not subject to significant operating stresses. Alternatively, the range of the TF insulation will be restricted so that the original specification can be achieved.
heating by 25%, compared to the case with no boron added. If the boron content is reduced to 1%, nuclear heat would increase by ∼4%.
4. Nuclear responses on the inboard vacuum vessel
3.3. Other related issues
4.1. Modeling of inboard blanket and vacuum vessel
• The effect of the gap width between the blanket module #1 and divertor cassette on the TF coil inboard leg heating has been analyzed. The gap width of 3.5 cm, which is more convenient than 2 cm for the divertor cassette handling, would increase TF coil inboard leg heating by ∼3%. • Nuclear heat distribution in the toroidal direction in the inboard leg winding pack is analyzed and found to be rather uniform. The peaking factor is found to be a few percent. The radiation distribution peaks due to the gaps between blanket modules will almost disappear as neutrons diffuse through the TF coil case. • Boron steel (2 wt% boron, natural enrichment) in the vacuum vessel (the present design) reduces nuclear
Distributions of nuclear responses on the vacuum vessel surface depend primarily on the presence of the gaps between blanket modules. They also depend, although to a lesser extent, on the fine structures of the blanket module and vacuum vessel. Among the fine structures, blanket cooling water manifolds, which are attached to the vacuum vessel surface and contain a large amount of water, have been implemented in the 3-D basic model for ITER. Figs. 4 and 5 show a rear side view of the inboard blanket module and a horizontal cross section of the inboard part of the calculation model “3-D basic model”, respectively. The following section gives the calculation results for the case with the gap width of the present design (1.0 cm of horizontal gap width).
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Fig. 5. Horizontal cross section of the 3-D basic model near the inboard mid plane.
4.2. Calculation results
Fig. 7. Nuclear heat distribution on the inboard vacuum vessel surface near #3/#4 module gap (gap width: 1.4 cm in vertical, 1.0 cm in horizontal directions).
Figs. 6–8 show distributions of the material damage (dPa), nuclear heat and helium production on the vacuum vessel surface around the #3/#4 module gap located near the mid plane. DPA, caused mainly by the high energy neutrons, is reduced by the shielding effect of the manifolds. The nuclear heat is instead higher due to the presence of the manifolds, suggesting that the water enhances slowing down of neutrons and their easier capture by the surrounding stainless steel, leading to generation of high energy photons. The effect of the presence of the manifolds on the helium production is in between of the above two, since it is caused by both high and low energy neutrons depending on Fig. 8. Helium production distribution on the inboard vacuum vessel surface near #3/#4 module gap (gap width: 1.4 cm in vertical, 1.0 cm in horizontal directions) (boron content in stainless steel: 10 appm).
the distance from the plasma and the boron content in the steels. Table 3 summarizes the nuclear response at the cross point of gaps (maximum point). The helium production is quite large (∼1.3 appm) at this point. Fortunately, field joint welding does not exist here, but near Table 3 Nuclear response values at the cross point of vertical and horizontal gaps between #3 and #4 blanket modules
Fig. 6. Material damage (DPA) distribution on the inboard vacuum vessel surface near #3/#4 module gap (gap width: 1.4 cm in vertical, 1.0 cm in horizontal directions).
Fast neutron fluence at the end of life (n/m2 ) Material damage at the end of life (dPa) Nuclear heat (W/cm3 ) He Pro. at the end of life (appm) End of life: 0.3 MW/m2 at the first wall (average).
8.42 × 1023 0.080 0.25 1.24
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the edge of the modules and ∼36 cm away from the vertical gap centre (see Figs. 7 and 8). The design limit for helium production is 1.0 appm for thick welding where re-welding is required. At the position of the field joint, the calculated helium production is 0.6–0.65 appm and lower than the limit but very close having only a minimal safety margin, which should cover the uncertainty (∼1.5) of the calculated value. Horizontal gaps wider than the present 1.0 cm width are very likely to cause a problem.
5. Conclusions Detailed analyses have been made on the inboard radiation shield capability of the present ITER machine with using the 3-D Monte Carlo code and sophisticated geometry input data. Thanks to the design changes of reducing gap widths between inboard blanket modules, the present inboard shield design satisfies design limits with an acceptable margin and risks.
Acknowledgements A special thanks to R.T. Santoro, who in the past years has supported and built the basis of this work. This report was prepared as an account of work undertaken within the framework of ITER Transitional Arrangements (ITA). These are conducted by the Participants: China, the European Atomic Energy Community, Japan, Korea, the Russian Federation and the
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United States of America under the auspices of the International Atomic Energy Agency. The views and opinions expressed herein do not necessarily reflect those of the Participants to the ITA, the IAEA or any agency thereof. Dissemination of the information in this paper is governed by the applicable terms of the former ITER EDA Agreement.
References [1] ITER Technical Basis, ITER EDA Documentation Series No. 24, IAEA, Vienna, 2002. [2] Y. Shimomura, et al., Nucl. Fusion 41 (2001) 309. [3] H. Iida, V. Khripunov, L. Petrizzi, “Nuclear Analysis Report”, G 73 DDD 2 01-06-06W 0.1, June 2001. [4] H. Iida, L. Petrizzi, V. Khripunov, et al., “Nuclear Analysis of ITER”, Proceedings of The 12th Biennial Topical Meeting of the Radiation Protection and Shielding Division of the American Nuclear Society (ANS), Santa Fe, New Mexico, April 14–18, 2002. [5] MCNP TM, A General Monte Carlo N-Particle Transport Code, in: J. Briesmeister (Ed.), Los Alamos National Laboratory, Los Alamos, New Mexico, LA-13709-M, March 2000. [6] H. Iida, V. Khripunov, L. Petrizzi “Nuclear Analysis Report”, G 73 DDD 2 04-04-17W 0.1, August, 2004. [7] P. Batistoni, M. Angelone, U. Fisher, H. Freiesleben, W. Hansen, M. Pillon, L. Petrizzi, D. Richter, K. Seidel, S. Unholzer, Neutronics experiment on a mock-up of the ITER shielding blanket at the Frascati neutron generator, Fusion Eng. Des. 47 (1999) 25–60. [8] F. Maekawa, C. Konno, Y. Oyama, Y. Ikeda, M. Wada, Y. Uno, Y.M. Verzilov, H. Maekawa, Validity assessment of shielding design tools for ITER through analysis of benchmark experiment on SS316/water shield conducted at FNS/JAERI, Fusion Technol. 30 (1996, Dec.) 1081–1087.
Nuclear analyses of some key aspects of the ITER design with Monte Carlo codes H. Iida a,∗ , L. Petrizzi b , V. Khripunov c , G. Federici d , E. Polunovskiy d a ITER International Team, Naka Joint Work Site, Naka-machi, Naka-gun, Ibaraki-ken, Japan ENEA FUS-TEC Frascati, Sezione Tecnologie della Fusione, Via E. Fermi 45, 00044 Frascati, Italy c Nuclear Fusion Institute, Russian Research Center “Kurchatov Institute”, 123182 Moscow, Russia ITER International Team, Garching Joint Work Site, Max-Planck-Institut fuer Plasmaphysik, Boltzmannstr. 2, D-85748 Garching, Germany b
d
Available online 10 August 2005
Abstract The design of the ITER machine was presented in 2001 [1,2]. A nuclear analysis was performed at this time, using fairly detailed models and the best assessed nuclear data and codes that were available. As the construction phase of ITER is approaching, the design of the main components has been optimized/finalized and several minor design changes/optimizations have been made, some with the object to mitigate critical radiation shielding problems. These have required refined calculations to confirm that the nuclear design requirements are met. This paper reviews some of the most recent neutronic work with emphasis on critical nuclear responses in the TF coil inboard legs and vacuum vessel related to design modifications made to the blanket modules and vacuum vessel. © 2005 Elsevier B.V. All rights reserved. Keywords: ITER; Monte Carlo method; TF coil; Nuclear heating; Vacuum vessel; Helium production
1. Introduction Critical design issues related to the inboard radiation shielding capability are: the reduction of heat loads in the TF coils and the vacuum vessel to a sufficiently low level; the protection of the TF coil winding packs from radiation damage and the reduction of helium production at the location of field-joint between the ∗ Corresponding author. Tel.: +81 29 270 7730; fax: +81 29 270 7506. E-mail address: [email protected] (H. Iida).
0920-3796/$ – see front matter © 2005 Elsevier B.V. All rights reserved. doi:10.1016/j.fusengdes.2005.06.239
vacuum vessel segments. The main factors affecting the shielding capability where design optimisation is still possible are the toroidal and poloidal gaps between the blanket sectors. These are required to allow fitting of the sectors by remote handling but have a sensitive impact on the overall shielding capability. 1.1. Nuclear heat and damage of winding packs in the TF coils In the original design, the nuclear heat in TF coils was analysed and reported to be 12.4 kW in the
134
H. Iida et al. / Fusion Engineering and Design 74 (2005) 133–139
Nuclear Analysis Report-2001(NAR-2001) [3,4] with 2 cm blanket gaps, while the cooling and refrigerator system of the TF coils were designed to remove up to ∼14 kW of nuclear heating at the reference pulse rate (2500 MW plasma pulses per hour). The calculation results of the nuclear heat have a certain degree of uncertainty, which is estimated to be of the order of ±70% on the inboard and ±80% on the outboard [3]. Consequently, there was a level of risk that the nuclear heat could exceed the design values for the refrigerator and coil cooling systems. Operating in this condition would have created a much higher cryogenic pumping work and would have been a very inefficient operating mode for the cryogenic heat extraction. Damage of the epoxy insulator in the winding packs was also at the limit, since the acceptable fast neutron fluence for some of the epoxy insulation systems being considered for the TF coils has turned out to be 5 × 1021 n/m2 , which is a half of the specified value.
A preliminary (1-D) calculation of the insulator damage documented in the NAR-2001 is 2.6 × 1021 n/m2 (ground insulation) and 2.3 × 1021 n/m2 (between first and second turns) with the assumption of 2 cm wide gaps between inboard blanket modules. These fluence values are close to the limit for the less radiation resistant epoxy insulator requiring a detailed 3-D analysis.
1.2. Helium production and heat loads at the surface of the vacuum vessel Due to radiation streaming through the gaps between blanket modules, locally peaked helium production and nuclear heat rates are expected on the vacuum vessel inner shell. The field joints require rewelding when replacement of the vessel segments is needed. The quality of such re-welding is guaranteed only when the helium production is less than 1 appm.
Fig. 1. Vertical cross section of the ITER 3-D model for the MCNP.
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When the heating density is larger than ∼0.3 W/cm3 , the thermal stress in the vacuum vessel can be a serious design issue. During the last few years, the design of some invessel components has undergone changes. Among the design changes, the reduction of the gap width between the blanket modules has a major impact on the radiation quantities above. This increases remote handling problem of the blanket modules but brings more margin, with more design flexibility, for the coils. Below, a summary of most recent nuclear analyses concerning these changes is presented.
2. Calculation tools and computing system The Monte Carlo transport code MCNP [5], nuclear data library FENDL1/MC and geometry input data of “ITER 3-D basic model” have been used for these analyses. After setting proper “weight window parameters”, most of calculation cases were finished by a 3 h run with 128 cpu (MPI) in the Origin3800 for obtaining acceptable statistics. Some cases for calculating very local peak values required ∼10 h to obtain sufficient statistics even with optimized weight windows.
3. TF coil nuclear heat and insulator damage 3.1. Nuclear heat The vertical cross section of the model is shown in Fig. 1. A TF coil is divided into 50 segments in the poloidal direction. The straight part of TF coils, which consists of segments from #1 to #14, will receive a major part (>70%) of the total heat and is defined here as “inboard leg of TF coil”. Shielding thickness for this part is minimal to reduce the size (hence cost) of the ITER machine.
Fig. 2. Nuclear heat poloidal distribution in the TF coil inboard leg. Case 1: vertical gaps 2.0 cm, all horizontal gaps on inboard 2.0 cm (2001 design), case 2: vertical gaps 1.4 cm, all horizontal gaps on inboard 1.0 cm (present design), case 3: vertical gaps1.4 cm, three horizontal gaps (#2/#3, #3/#4, #4/#5) near the equatorial plane 1.0 cm, other horizontal gaps 2.0 cm (back up option) and case 4: vertical gaps 1.4 cm, all horizontal gaps on inboard 2.0 cm.
Fig. 2 shows the poloidal distribution of the nuclear heat in the inboard TF coil legs. The nuclear heat is naturally high at the mid plane (segment #8). For the current design (case 2 in Table 1) the nuclear heat is lower by a factor of about 25% with respect to the 2001 design (case 1). Detailed assessment of the remote handing and operational displacements of the sectors confirms that the width can be reduced from 2 to 1.4 cm for the vertical gaps (the nuclear heat value for this step, with the original 2 cm horizontal gap, is shown as case 4). The reduction of the width of the horizontal gaps requires further work since the remote handling feasibility for the narrower gaps has to be confirmed by additional R&D. Narrower gaps also give some constraints on installation of diagnostic equipment. Case
Table 1 Nuclear heat in the TF coil inboard legs
Case 1 Case 2 Case 3 Case 4 a
Verical gap width (cm)
Horizontal gap width (cm)
Inboard leg (kW)
Total in super conducting coil (kW)a
2.0 1.4 1.4 1.4
2.0 1.0 Central 3 gaps 1.0, other gaps 2.0 2.0
10.0 7.51 7.78 9.14
12.4 9.91 10.2 11.5
Inboard leg + 2.4 kW.
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3, which has narrowed width only on the gaps near the mid plane and gives slightly higher heat than the case 2, is a back up option in case there are constraints in narrowing the gap width as a result of future R&D for remote handling technology. The total nuclear heat value of 9.91 kW (case 2) gives a much larger safety margin to cover some of the uncertainty in the calculation results than the 2001 design (case 1). A safety factor of 40% is now allowed on the calculated values in the design value of 14 kW assumed for nuclear heat in the coils, at which the design pulse rates can be maintained. Fig. 3 shows the uncertainty levels the present design will have [6]. This figure assumes uncertainties of 3-D modeling (30%) in the design calculation and of benchmark experiments (20%) at FNG (ENEA Frascati) [7] and FNS (JAERI) [8]. The average of the “most probable values” (µ1) is assumed to be larger than that of the “best estimates with calculation code” (µ2) by a factor of 1.15 [see ref. [6] in detail]. According to this figure, the probability of the nuclear heat being higher than 14 kW is ∼30% (R = 14 kW/9.91 kW = 1.4). This risk level has been accepted, since the coil design specification still allows for nuclear heating up to the full uncertainty range (a factor of 2, 18 kW) by permitting a reduction in plasma pulse rate, or an increase in the refrigerator capacity. The coil cooling design has been modified to allow a higher rate of heat extraction with a lower cost in cryogenic pumping power. 3.2. Fast neutron fluence on the epoxy insulator The maximum fast neutron fluence on the inboard legs occurs on the #8 segment, as easily expected from the nuclear heat distribution. The relation between
Fig. 3. Probability function of between calculated and probable 2 2 2 2 1 P(R) dR = · e−(Rµ1 −µ2 ) /2(σ2 +R σ1 ) 2πσ1 σ2 1 1 −β2 · 2 e + β{1 + ERF(β)} dR α 2
the ratio R nuclear heating
α=
, σ22 + Rσ12 σ 2 µ1 + Rσ 2 µ2 ,β= √ 2 1 , √ 2 2 2 2σ1 σ2 2σ1 σ2 σ2 + R σ1
β
2 t −2 dt ERF(β) = √ π 0 where fsd1 and fsd2 : fractional standard deviations of the calculated and probable values.
the maximum fluence and blanket module gap width is shown in Table 2. Considering the uncertainty of about a factor 2 in the calculated values, fluence values in the table should be less than 2.5 × 1021 n/m2 , which would allow the insulation requirement to be reduced to 5 × 1021 n/m2 when required. The fluence on the insulator between #1 and #2 turns is lower than
Table 2 Maximum fast neutron fluence on the epoxy insulator in the inboard TF coil leg
Case a Case b Case c Case d Case e Case f Case g
Vertical gap width (cm)
Horizontal gap width (cm)
Fast neutron fluence in #8 segment (1021 n/m2 ) at the end of life Ground insulator
Between first and second turns
2 1 1.4 1.4 1.4 1.4 1.4
2 1 2 1.5 1 0.5 0
5.02 3.02 4.50 3.81 3.25 2.88 2.71
2.73 1.64 2.45 2.07 1.76 1.56 1.46
End of life: 0.3 MW/m2 at the first wall (average).
H. Iida et al. / Fusion Engineering and Design 74 (2005) 133–139
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Fig. 4. Rear side view of the inboard blanket module with cooling water manifolds on the vacuum vessel surface.
2.5 × 1021 n/m2 in all cases with 1.4 cm of the vertical gap width, although those on the plasma side ground insulator can go up to 4.5 × 1021 n/m2 . The high fluence values on the plasma side ground insulator are judged to be acceptable since this part is not subject to significant operating stresses. Alternatively, the range of the TF insulation will be restricted so that the original specification can be achieved.
heating by 25%, compared to the case with no boron added. If the boron content is reduced to 1%, nuclear heat would increase by ∼4%.
4. Nuclear responses on the inboard vacuum vessel
3.3. Other related issues
4.1. Modeling of inboard blanket and vacuum vessel
• The effect of the gap width between the blanket module #1 and divertor cassette on the TF coil inboard leg heating has been analyzed. The gap width of 3.5 cm, which is more convenient than 2 cm for the divertor cassette handling, would increase TF coil inboard leg heating by ∼3%. • Nuclear heat distribution in the toroidal direction in the inboard leg winding pack is analyzed and found to be rather uniform. The peaking factor is found to be a few percent. The radiation distribution peaks due to the gaps between blanket modules will almost disappear as neutrons diffuse through the TF coil case. • Boron steel (2 wt% boron, natural enrichment) in the vacuum vessel (the present design) reduces nuclear
Distributions of nuclear responses on the vacuum vessel surface depend primarily on the presence of the gaps between blanket modules. They also depend, although to a lesser extent, on the fine structures of the blanket module and vacuum vessel. Among the fine structures, blanket cooling water manifolds, which are attached to the vacuum vessel surface and contain a large amount of water, have been implemented in the 3-D basic model for ITER. Figs. 4 and 5 show a rear side view of the inboard blanket module and a horizontal cross section of the inboard part of the calculation model “3-D basic model”, respectively. The following section gives the calculation results for the case with the gap width of the present design (1.0 cm of horizontal gap width).
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H. Iida et al. / Fusion Engineering and Design 74 (2005) 133–139
Fig. 5. Horizontal cross section of the 3-D basic model near the inboard mid plane.
4.2. Calculation results
Fig. 7. Nuclear heat distribution on the inboard vacuum vessel surface near #3/#4 module gap (gap width: 1.4 cm in vertical, 1.0 cm in horizontal directions).
Figs. 6–8 show distributions of the material damage (dPa), nuclear heat and helium production on the vacuum vessel surface around the #3/#4 module gap located near the mid plane. DPA, caused mainly by the high energy neutrons, is reduced by the shielding effect of the manifolds. The nuclear heat is instead higher due to the presence of the manifolds, suggesting that the water enhances slowing down of neutrons and their easier capture by the surrounding stainless steel, leading to generation of high energy photons. The effect of the presence of the manifolds on the helium production is in between of the above two, since it is caused by both high and low energy neutrons depending on Fig. 8. Helium production distribution on the inboard vacuum vessel surface near #3/#4 module gap (gap width: 1.4 cm in vertical, 1.0 cm in horizontal directions) (boron content in stainless steel: 10 appm).
the distance from the plasma and the boron content in the steels. Table 3 summarizes the nuclear response at the cross point of gaps (maximum point). The helium production is quite large (∼1.3 appm) at this point. Fortunately, field joint welding does not exist here, but near Table 3 Nuclear response values at the cross point of vertical and horizontal gaps between #3 and #4 blanket modules
Fig. 6. Material damage (DPA) distribution on the inboard vacuum vessel surface near #3/#4 module gap (gap width: 1.4 cm in vertical, 1.0 cm in horizontal directions).
Fast neutron fluence at the end of life (n/m2 ) Material damage at the end of life (dPa) Nuclear heat (W/cm3 ) He Pro. at the end of life (appm) End of life: 0.3 MW/m2 at the first wall (average).
8.42 × 1023 0.080 0.25 1.24
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the edge of the modules and ∼36 cm away from the vertical gap centre (see Figs. 7 and 8). The design limit for helium production is 1.0 appm for thick welding where re-welding is required. At the position of the field joint, the calculated helium production is 0.6–0.65 appm and lower than the limit but very close having only a minimal safety margin, which should cover the uncertainty (∼1.5) of the calculated value. Horizontal gaps wider than the present 1.0 cm width are very likely to cause a problem.
5. Conclusions Detailed analyses have been made on the inboard radiation shield capability of the present ITER machine with using the 3-D Monte Carlo code and sophisticated geometry input data. Thanks to the design changes of reducing gap widths between inboard blanket modules, the present inboard shield design satisfies design limits with an acceptable margin and risks.
Acknowledgements A special thanks to R.T. Santoro, who in the past years has supported and built the basis of this work. This report was prepared as an account of work undertaken within the framework of ITER Transitional Arrangements (ITA). These are conducted by the Participants: China, the European Atomic Energy Community, Japan, Korea, the Russian Federation and the
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United States of America under the auspices of the International Atomic Energy Agency. The views and opinions expressed herein do not necessarily reflect those of the Participants to the ITA, the IAEA or any agency thereof. Dissemination of the information in this paper is governed by the applicable terms of the former ITER EDA Agreement.
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