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Using graphitic foam as the bonding material in metal fuel pins for sodium fast reactors
Journal of Nuclear Materials 441 (2013) 539–544
Contents lists available at SciVerse ScienceDirect
Journal of Nuclear Materials journal homepage: www.elsevier.com/locate/jnucmat
Using graphitic foam as the bonding material in metal fuel pins for sodium fast reactors Aydın Karahan ⇑, Mujid S. Kazimi Center for Advanced Nuclear Energy Systems, Nuclear Science and Engineering Department, Massachusetts Institute of Technology, 77 Massachusetts Avenue, Room: 24-215, Cambridge, MA 02139, United States
a r t i c l e
i n f o
Article history: Available online 28 March 2013
a b s t r a c t The study evaluates the possible use of graphite foam as the bonding material between U–Pu–Zr metallic fuel and steel clad for sodium fast reactor applications using FEAST-METAL fuel performance code. Furthermore, the applicability of FEAST-METAL to the advanced fuel designs is demonstrated. Replacing the sodium bond with a chemically stable foam material would eliminate fuel clad metallurgical interactions, and allow for fuel swelling under low external stress. Hence, a significant improvement is expected for the steady state and transient performance. FEAST-METAL was used to assess the thermo-mechanical behavior of the new fuel form and a reference metallic fuel pin. Nearly unity conversion ratio, 75% smear density U–15Pu–6Zr metallic fuel pin with sodium bond, and T91 cladding was selected as a reference case. It was found that operating the reference case at high clad temperatures (600–660 °C) results in (1) excessive clad wastage formation/clad thinning due to lanthanide migration and formation of brittle phases at clad inner surface, and (2) excessive clad hoop strain at the upper axial section due mainly to the occurrence of thermal creep. The combination of these two factors may lead to cladding breach. The work concludes that replacing the sodium bond with 80% porous graphite foam and reducing the fuel smear density to 70%, it is likely that the fuel clad metallurgical interaction would be eliminated while the fuel swelling is allowed without excessive fuel clad mechanical interaction. The suggested design appears as an alternative for a high performance metallic fuel design for sodium fast reactors. Ó 2013 Elsevier B.V. All rights reserved.
1. Introduction U–Pu–Zr alloy fuels are one of the candidates for sodium fast reactor applications [1]. Its higher heavy metal density, better neutronic feedback coefficients and reprocessing advantages are primary thrusts to pursue the metal fuel technology. On the other hand, the chemical instability of this material with steel cladding appears to be the primary performance limiting phenomena. Sodium has been adopted as bond material between the fuel and the cladding. Although it has excellent thermal properties, it does not provide a barrier against fuel–clad metallurgical interactions. Furthermore, it infiltrates into the fuel slug and brings the boiling risk in rapid heating transient scenarios, which may also cause a positive reactivity insertion. At power, metal fuel swells and comes to contact with the cladding after 1–2 at% burnup. Once the contact establishes, lanthanide migration into the clad and formation of brittle phases at the inner surface may occur at a significant level, especially above 620 °C clad inner temperature [2]. Moreover, iron migrates into the fuel slug and forms various phases such as (U, Pu)6Fe, which has a low melting point [3]. Above the fuel liquefac⇑ Corresponding author. Tel.: +90 530 774 63 63. E-mail address: [email protected] (A. Karahan). 0022-3115/$ - see front matter Ó 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.jnucmat.2013.03.049
tion temperature (between 650 °C and 725 °C), eutectic formation and penetration into the cladding occurs with a rate much more rapid compared to the steady state clad wastage rate [4]. Fuel clad metallurgical interactions limit the plant thermal efficiency, long term design basis and severe accident performance. Hence, a practical solution to this problem is essential. Barrier fuel concept for the metallic fuel has been examined by many researchers [5–8]. Typically, a suggested solution is to include a thin liner at the clad inner surface to eliminate the metallurgical interactions. Yang et al. has recently suggested adopting 30 lm of Cr liner based on high temperature experiments [8]. We note that in addition to fuel clad chemical interaction, in reactor conditions dictates (1) operation with a high dose due to fast neutron collisions with the material, (2) fuel clad mechanical interaction and (3) thermal stresses during transients. Karahan suspects if thermo-mechanical and micro-structural stability of the suggested thin liners under in-reactor conditions will yield a satisfactory performance. The present work suggests replacing sodium with a foam material as bonding between the metal fuel slug and cladding (Fig. 1). A high amount of porosity eases the flow of the material. It may prevent the stress build up due to irradiation induced dimensional changes and hardening under irradiation. This appears to be a un-
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Fig. 1. (a) Metallic fuel with graphitic foam, (b) foam micro-structure [10].
Fig. 2. Foam yield stress as a function foam density [9].
ique advantage of using foam material instead of using a thin solid barrier layer. The work considers using graphitic foam for demonstration purposes. It is also encouraged that various other foam materials should also be regarded as the future work. Graphitic foam materials possess low yield stress (Fig. 2) and high thermal conductivity [9]. The combination of a yield stress of on the order of few MPa and gap thermal conductivity on the order of 10 W/m/K appears to be promising for a bonding material to be used between the metal fuel and the cladding. Typically, thermal conductivity of the graphitic foam gets degraded with irradiation [10]. On the other hand, it is enhanced due to the reduction in foam porosity as the fuel swells. A conductivity of 10 W/m/K is conservatively assumed constant in this preliminary study. Compared to sodium thermal conductivity, the selected value, 10 W/m/K, for the foam is a much smaller number. Its effect on fuel temperature and overall fuel performance is given below. In addition to thermal and mechanical behavior, chemical compatibility of the foam material needs to be assessed. Based on Ref. [11] study, compatibility of the ferritic steels with graphite is satisfactory up to 700 °C. The authors assume that T91 [12], which is a 9Cr–1Mo ferritic/martensitic stainless steel, should also be compatible with graphite at temperatures up to 650–700 °C. The compatibility of the graphite with the U–Pu–Zr metal fuel alloys is an unexplored field. Operating the metallic fuel in contact with graphite above half of the melting temperature would result in the formation of different phases at metal fuel/graphite interface. Due to high free energy of formation of zirconium carbide as well as uranium plutonium carbide [13], it is expected that a stiff layer will form at the metal fuel/graphite interface which would hinder further diffusivity. Perhaps the biggest challenge of this design will be its fabrication. One way could be to braze the fuel and the foam first around
600 °C, then insert them together into the steel cladding. Another way would be to insert fuel into the steel cladding. Then, insert the foam into the gap region in small pieces as suggested by Ogata [14]. Another challenge appears to be the reprocessing of this fuel. Although the content of the carbon atoms is low due to high porosity, the inclusion of the carbide phases would complicate the usual reprocessing schemes. If once-through fuel cycle is preferred as suggested for breed-burn reactor concept, this should not cause any problem. Nevertheless, the fabrication and reprocessing aspects of the design problem are left as the future work. The first part of the paper introduces the reference metallic fuel pin description and an approximate neutronic analysis for the clad dose calculation. The second part will present the analyses of the thermo-mechanical behavior of the reference as well as the advanced metallic fuel performed with FEAST-METAL [15]. FEAST code is equipped with mechanistic models to simulate fission gas behavior, fuel clad chemical interaction, fuel constituent redistribution, and mechanical behavior of the fuel pin. The code is validated against EBR-II experimental database and in this work it is applied to an advanced design. The selected optimized configuration for the advanced fuel will be suggested using FEAST code results.
2. Description of the reference metallic fuel Selected properties of the metal fuel of near unity conversion ratio sodium fast reactor are given in Table 1. Typical EBR-II fuel pin geometry is adopted [1]. The zirconium content of the fuel is selected as 6 wt%. Most metallic fuels irradiated at EBR-II bear 10 wt% Zirconium; however, Refs. [16,17] showed that fuels containing 6 wt% and 10 wt% Zirconium acted fairly similar. T91 type
Table 1 Fuel composition, geometry and operating conditions for the reference metallic fuel. Property
Value
Fuel composition (wt%) Clad material Bond material Clad outer diameter (mm) Clad thickness (mm) Fuel smear density (%) Active fuel length (m) Pitch to diameter ratio Plenum to fuel volume ratio Peak burnup (at%) Peak dose (dpa) Coolant inlet temperature (°C) Peak clad temperature (°C) Peak linear heat rate (kW/m) Axial power profile Axial power peak
U–15Pu–6Zr T91 Sodium 5.84 0.38 75 1.0 1.2 2.0 15 200 390 660 35 Chopped cosine 1.25
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Fig. 3. Peak linear heat rate as a function of peak burnup (kW/m).
Fig. 5. Full core MCNP layout.
Fig. 4. Peak coolant outlet temperature as a function of peak burnup (°C).
steel [12] is selected as clad material over HT9 type steel in order to operate at clad temperatures as high as 660 °C without concerning about the thermal creep. The assumed clad dose limit is 200 dpa. Peak achievable burnup is 15 at% due to cladding dose limit. Finally, a large plenum to fuel volume ratio is adopted to accommodate for the fission gases. Operation with a plenum pressure above 10 MPa could deteriorate the high temperature transient performance (design basis accidents) significantly due to low thermal creep fracture margin of the ferritic–martensitic steels. Hence, a large plenum to fuel volume ratio is inevitable for this fuel pin design. Variation of the peak linear heat rate and peak coolant temperature within the assembly is given as a function of burnup in Figs. 3 and 4, respectively. Selected power and coolant temperature decreases linearly with burnup.
3. Neutronic analysis to assess the clad dose Neutronic analysis is a critical part of a fast reactor fuel design. Typically, the assessment of power and burnup distribution, neutronic feedback coefficients, and the dose on the cladding need to be performed. Note that the density of the foam is rather low and the effect of graphite bond on excess reactivity remains at a minor level. The present neutronic analyses were performed only to assess the clad dose for the selected fuels in an approximate manner. The dose equivalence of EBR-II core fast neutron fluence is known. Typically, 1E+22 n/cm2 (>0.1 MeV) fluence corresponds to 5 dpa clad dose [18]. The current study computed the neutron spectrum of EBR-II fuel and the reference fuel using MCNP-4C [19]. The MCNP full core model is given in Fig. 5. Inner assemblies consist of driver fuels and control assemblies. The peripheral regions are occupied by the blanket assemblies. By using EBR-II fuel properties, which is 66% enriched U–10Zr and the reference fuel properties given in Table 1, the driver fuel neutron flux spectrum was calculated for 35 kW/m linear heat rate conditions (Beginning
Fig. 6. Neutron flux spectrum for EBR-II fuel and U–15Pu–6Zr fuel at beginning of Life.
of Life). As depicted in Fig. 6, EBR-II operated with a much lower neutron flux level compared to a possible commercial fast reactor design. Although the fissile Pu isotopes are neutronically advantageous over U-235 at fast neutron spectrum, much larger fissile density of EBR-II lowers the neutron flux level. In addition, the neutron spectrum of EBR-II fuel is somewhat harder compared to the reference fuel again due to the large fissile density. Using energy deposition cross-sections given in Ref. [20], dose rate calculation of the reference fuel was performed in a relative manner. It was found that clad dose rate in typical unity conversion ratio sodium fast reactors is about 2.3 time higher compared to EBR-II conditions. 4. Thermo-mechanical analysis of the reference metallic fuel pin Steady state thermo-mechanical analysis of the reference U– 15Pu–6Zr metallic fuel with sodium bond was performed with FEAST-METAL code. Fig. 7 shows the axial variation of the clad inner temperature and fuel centerline temperature at the beginning of life. Upper axial part of the fuel operates at higher temperatures
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Fig. 11. Fission gas release for the reference fuel. Fig. 7. Fuel and clad temperatures at beginning of Life for the reference fuel.
Fig. 8. Axial variation of the clad inner wastage for the reference fuel.
Fig. 9. Clad hoop strain at end of Life for the reference fuel.
Fig. 10. Axial variation of the clad hoop stress as a function of peak burnup for the reference fuel.
and it is softer compared to lower axial sections. On the other hand, the operation of the clad at high temperatures leads to excessive lanthanide attack to the clad inner surface, formation of brittle
Fig. 12. Fuel swelling at the middle axial section for the reference fuel.
phases and the resulting clad thinning for more than 30% of the as-fabricated clad thickness, as shown in Fig. 8. Fig. 9 shows the axial variation of the clad hoop strain at end of life. Excessive thermal creep was observed towards the upper axial section mainly due to clad wastage formation and thinning due to formation of brittle phases at clad inner region. Lower axial sections were exposed to clad deformation via irradiation creep and a small amount of void swelling. Fig. 10 shows the axial variation of the clad hoop stress. Fuel clad mechanical interaction remains low at upper, hot, axial sections of the fuel pin. In these regions, the hoop stress is mainly due to plenum pressure. On the other hand, lower and middle sections of the fuel, which consists of stiff dual a + d phase at the outer radial region, have low compressibility, holding the contact pressure at a much higher level at high burnup. Figs. 11 and 12 shows the fission gas release and swelling behavior of the reference fuel, respectively. Operation of the fuel above half of the melting point leads to high diffusivity, large fission gas swelling and release starting at early irradiation. Solid fission product swelling linearly increases at the expense of pore shrinkage with the hot pressing mechanism under the external stress. Open pores are more compressible at the hotter section but they are not as compressible at colder sections of the fuel. Once the porosity concentration drops below 10%, fuel clad mechanical interaction rises significantly, as seen in Fig. 10. Predicted plenum pressure at the end of life is 8 MPa.
5. Thermo-mechanical analysis of the advanced metallic fuel pin The advanced fuel pin includes the graphite foam as the bonding material between fuel and clad. The operating conditions were kept the same as the reference fuel pin given in Table 1. The initial porosity of the foam is 80%. The foam is expected to yield under low external restraint, so that the fuel swelling and porosity forma-
A. Karahan, M.S. Kazimi / Journal of Nuclear Materials 441 (2013) 539–544
Fig. 13. Axial variation of the clad hoop strain at the end of Life for 65%, 70% and 75% smear density advanced fuels.
Fig. 14. Fuel and clad temperatures for 70% smear density advanced fuel.
Fig. 15. Swelling behavior of the middle region of the 70% smear density advanced fuel.
tion would be allowed without straining of the clad, excessively. The average behavior of the yield stress of the foam types given in Fig. 2 was used to calculate the compressibility and the stresses on the fuel/foam interface and foam/clad interface. Foam porosity is assumed to remain at plenum pressure. Irradiation creep and irradiation hardening effects of the foam are neglected for this preliminary study. FEAST-METAL was ran for 65%, 70% and 75% fuel smear density to simulate the advanced metallic fuel with 80% porous graphite foam. As can be seen in Fig. 13 and 75% smear density fuel caused a significant clad straining due to fuel swelling. For the new design, leaving a residue foam thickness at the end of life is inevitable; hence, keeping the fuel pin dimensions the same as the reference fuel, excessive straining of the clad is not surprising for the 75% smear density. This design does not allow enough space
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Fig. 16. Fission gas release behavior of 70% smear density fuel as a function of peak burnup for the advanced fuel.
Fig. 17. Axial variation of the foam porosity at beginning of Life and end of Life for 70% smear density advanced fuel.
Fig. 18. Fuel outer radius and clad inner radius axial variation at the end of Life for the advanced fuel.
for fuel to build the interconnected porosity especially at the lower sections at which the fuel is stiffer. The authors suggest to reduce the fuel smear density to 70% to eliminate/delay excessive fuel clad mechanical interaction (Fig. 13). Detailed results and discussion for 70% smear density, graphite bonded metallic fuel are given below. Fig. 14 shows the fuel and clad temperature variation at the beginning of life. Fuel temperature is higher compared to the reference case simulation (Fig. 7). It is due to (1) lower thermal conductivity of the graphite foam compared to sodium bond and (2) the absence of sodium logging into the fuel which improves effective fuel thermal conductivity. Note that the new design entirely eliminates the sodium infiltration into the fuel. Therefore, the sodium boiling should not cause any limitation to the fuel slug centerline temperature.
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ial parts and similar at the lower and middle axial sections compared with the reference case. 6. Conclusions and future work
Fig. 19. Axial variation of clad hoop stress as a function of peak burnup for 70% smear density advanced fuel.
The study suggests using a foam material as the novel barrier material between metallic fuel and steel clad to improve the chemical stability, to provide reasonable thermal conductance and compliancy to allow for fuel swelling without yielding the clad, excessively. Due to its favorable properties, the current study adopts a graphitic foam material as an example. Furthermore, using FEAST-METAL code for an advanced fuel design is accomplished. The fuel performance simulation results for a typical metallic fuel pin with sodium bond operating above 600–660 °C peak clad temperature demonstrate the necessity of a barrier material to hinder fuel cladding chemical interaction. Next, the simulation using the graphitic foam material between fuel and clad suggests lowering the fuel smear density to 70% to avoid excessive straining of the cladding. The new design is expected to improve the plant thermal efficiency and boost the long term transient and severe accident performance of the metal fueled sodium fast reactors. Further experimental studies as well as higher fidelity modeling and simulation efforts are needed to validate this behavior. Furthermore, the fabrication and reprocessing feasibility remains to be demonstrated. References
Fig. 20. Axial variation of the clad hoops strain and thermal creep strain for 70% smear density advanced fuel.
Higher available thermal activation resulted in an earlier start of fission gas release and higher rate of fuel swelling as seen in Figs. 15 and 16. Shortly after irradiation, interconnected open porosity forms and weakens the fuel slug. Fuel becomes both axially and radially restrained by the foam. Once the fuel open porosity drops below 10%, the fuel becomes less compressible and foam starts to shrink again under the applied compressive stress. The predicted plenum pressure at the end of life is 8.2 MPa. Figs. 17 and 18 show the foam porosity axial variation, and fuel slug outer radius and clad inner radius at the end of Life, respectively. The minimum residue foam thickness is on the order of 100 lm, which is expected by authors to be enough to prevent metallurgical interactions between the fuel and the clad. Figs. 19 and 20 shows the clad hoop stress and strain axial variations. At high burnup, the rate of change of clad hoop stress increased in parallel with the depletion of the fuel porosity. Compared to the reference case, the level of thermal creep is lower and acceptable. Hence, in this case, the clad hoop straining mostly occurs due to irradiation creep and a small amount of the void swelling. The total clad hoop strain is much lower at the upper ax-
[1] D.C. Crawford, D.L. Porter, S.L. Hayes, J. Nucl. Mater. 371 (2007) 202–231. [2] R.G. Pahl, C.E. Lahm, S.L. Hayes, J. Nucl. Mater. 204 (1993) 141–147. [3] K. Nakamura, T. Ogata, M. Kurata, T. Yokoo, M.A. Mignanelli, J. Nucl. Sci. Technol. 38 (2) (2001) 112–119. [4] A.B. Cohen, H. Tsai, L.A. Neimark, J. Nucl. Mater. 204 (1993) 244–251. [5] D.C. Crawford, C.E. Lahm, H. Tsai, R.G. Pahl, J. Nucl. Mater. 204 (1993) 157–164. [6] J.H. Kim, H.J. Ryu, J.H. Baek, S.J. Oh, B.O. Lee, C.B. Lee, J. Nucl. Mater. 394 (2009) 144–150. [7] H.J. Ryu, B.O. Lee, S.J. Oh, J.H. Kim, C.B. Lee, J. Nucl. Mater. 392 (2009) 206–212. [8] S.W. Yang, H.J. Ryu, J.H. Kim, B.O. Lee, C.B. Lee, J. Nucl. Mater. 401 (2010) 98– 103. [9] J. Klett, S. Jones, L. Klett, C. Walls, High Thermal Conductivity Graphitic Foam Reinforced Carbon–Carbon Composites, in: International SAMPE Symposium and Exhibition (Proceedings), vol. 48, 2003, pp-38-52. [10] N.C. Gallego, T.D. Burchell, J.W. Klett, Carbon 44 (2006) 618–628. [11] J.C. Bokros, J. Nucl. Mater. 3 (1) (1960) 89–100. [12] D.S. Gelles, J. Nucl. Mater. 239 (1996) 99–106. [13] R.G. Coltters, Mater. Sci. Eng. 76 (1985) 1–50. [14] Takanari Ogata, Central Research Institute of Electrical Power Industry, June, 2010, Personel Communication. [15] A. Karahan, J. Buongiorno, Modeling of Thermo-Mechanical and Irradiation Behavior of Metallic and Oxide Fuels for Sodium Fast Reactors, Report No: MITNFC-TR-110, Center for Advanced Nuclear Energy Systems, Massachusetts Institute of Technology, 2009. [16] R.G. Pahl, D.L. Porter, D.C. Crawford, L.C. Walters, J. Nucl. Mater. 188 (1992) 3– 9. [17] G.L. Hofman, A.M. Yacout, T.K. Kim, T.H. Bauer, N. Ueda, Assessment of Low Zr Metallic Fuel Alloys, in: Proceedings of Global 2009, 2009, Paper 9462. [18] F.A. Garner, D.L. Porter, J. Nucl. Mater. 1006–1013 (1988) 155–157. [19] J.F. Briesmeister (Ed.), MCNP–A General Monte Carlo N-Particle Transport Code, Los Alamos National Laboratory Report LA-12625-M, Los Alamos, NM, 1993. [20] D.R. Olander, Fundamental Aspects of Nuclear Reactor Fuel Elements, U.S. DOE, 1976.
Contents lists available at SciVerse ScienceDirect
Journal of Nuclear Materials journal homepage: www.elsevier.com/locate/jnucmat
Using graphitic foam as the bonding material in metal fuel pins for sodium fast reactors Aydın Karahan ⇑, Mujid S. Kazimi Center for Advanced Nuclear Energy Systems, Nuclear Science and Engineering Department, Massachusetts Institute of Technology, 77 Massachusetts Avenue, Room: 24-215, Cambridge, MA 02139, United States
a r t i c l e
i n f o
Article history: Available online 28 March 2013
a b s t r a c t The study evaluates the possible use of graphite foam as the bonding material between U–Pu–Zr metallic fuel and steel clad for sodium fast reactor applications using FEAST-METAL fuel performance code. Furthermore, the applicability of FEAST-METAL to the advanced fuel designs is demonstrated. Replacing the sodium bond with a chemically stable foam material would eliminate fuel clad metallurgical interactions, and allow for fuel swelling under low external stress. Hence, a significant improvement is expected for the steady state and transient performance. FEAST-METAL was used to assess the thermo-mechanical behavior of the new fuel form and a reference metallic fuel pin. Nearly unity conversion ratio, 75% smear density U–15Pu–6Zr metallic fuel pin with sodium bond, and T91 cladding was selected as a reference case. It was found that operating the reference case at high clad temperatures (600–660 °C) results in (1) excessive clad wastage formation/clad thinning due to lanthanide migration and formation of brittle phases at clad inner surface, and (2) excessive clad hoop strain at the upper axial section due mainly to the occurrence of thermal creep. The combination of these two factors may lead to cladding breach. The work concludes that replacing the sodium bond with 80% porous graphite foam and reducing the fuel smear density to 70%, it is likely that the fuel clad metallurgical interaction would be eliminated while the fuel swelling is allowed without excessive fuel clad mechanical interaction. The suggested design appears as an alternative for a high performance metallic fuel design for sodium fast reactors. Ó 2013 Elsevier B.V. All rights reserved.
1. Introduction U–Pu–Zr alloy fuels are one of the candidates for sodium fast reactor applications [1]. Its higher heavy metal density, better neutronic feedback coefficients and reprocessing advantages are primary thrusts to pursue the metal fuel technology. On the other hand, the chemical instability of this material with steel cladding appears to be the primary performance limiting phenomena. Sodium has been adopted as bond material between the fuel and the cladding. Although it has excellent thermal properties, it does not provide a barrier against fuel–clad metallurgical interactions. Furthermore, it infiltrates into the fuel slug and brings the boiling risk in rapid heating transient scenarios, which may also cause a positive reactivity insertion. At power, metal fuel swells and comes to contact with the cladding after 1–2 at% burnup. Once the contact establishes, lanthanide migration into the clad and formation of brittle phases at the inner surface may occur at a significant level, especially above 620 °C clad inner temperature [2]. Moreover, iron migrates into the fuel slug and forms various phases such as (U, Pu)6Fe, which has a low melting point [3]. Above the fuel liquefac⇑ Corresponding author. Tel.: +90 530 774 63 63. E-mail address: [email protected] (A. Karahan). 0022-3115/$ - see front matter Ó 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.jnucmat.2013.03.049
tion temperature (between 650 °C and 725 °C), eutectic formation and penetration into the cladding occurs with a rate much more rapid compared to the steady state clad wastage rate [4]. Fuel clad metallurgical interactions limit the plant thermal efficiency, long term design basis and severe accident performance. Hence, a practical solution to this problem is essential. Barrier fuel concept for the metallic fuel has been examined by many researchers [5–8]. Typically, a suggested solution is to include a thin liner at the clad inner surface to eliminate the metallurgical interactions. Yang et al. has recently suggested adopting 30 lm of Cr liner based on high temperature experiments [8]. We note that in addition to fuel clad chemical interaction, in reactor conditions dictates (1) operation with a high dose due to fast neutron collisions with the material, (2) fuel clad mechanical interaction and (3) thermal stresses during transients. Karahan suspects if thermo-mechanical and micro-structural stability of the suggested thin liners under in-reactor conditions will yield a satisfactory performance. The present work suggests replacing sodium with a foam material as bonding between the metal fuel slug and cladding (Fig. 1). A high amount of porosity eases the flow of the material. It may prevent the stress build up due to irradiation induced dimensional changes and hardening under irradiation. This appears to be a un-
540
A. Karahan, M.S. Kazimi / Journal of Nuclear Materials 441 (2013) 539–544
Fig. 1. (a) Metallic fuel with graphitic foam, (b) foam micro-structure [10].
Fig. 2. Foam yield stress as a function foam density [9].
ique advantage of using foam material instead of using a thin solid barrier layer. The work considers using graphitic foam for demonstration purposes. It is also encouraged that various other foam materials should also be regarded as the future work. Graphitic foam materials possess low yield stress (Fig. 2) and high thermal conductivity [9]. The combination of a yield stress of on the order of few MPa and gap thermal conductivity on the order of 10 W/m/K appears to be promising for a bonding material to be used between the metal fuel and the cladding. Typically, thermal conductivity of the graphitic foam gets degraded with irradiation [10]. On the other hand, it is enhanced due to the reduction in foam porosity as the fuel swells. A conductivity of 10 W/m/K is conservatively assumed constant in this preliminary study. Compared to sodium thermal conductivity, the selected value, 10 W/m/K, for the foam is a much smaller number. Its effect on fuel temperature and overall fuel performance is given below. In addition to thermal and mechanical behavior, chemical compatibility of the foam material needs to be assessed. Based on Ref. [11] study, compatibility of the ferritic steels with graphite is satisfactory up to 700 °C. The authors assume that T91 [12], which is a 9Cr–1Mo ferritic/martensitic stainless steel, should also be compatible with graphite at temperatures up to 650–700 °C. The compatibility of the graphite with the U–Pu–Zr metal fuel alloys is an unexplored field. Operating the metallic fuel in contact with graphite above half of the melting temperature would result in the formation of different phases at metal fuel/graphite interface. Due to high free energy of formation of zirconium carbide as well as uranium plutonium carbide [13], it is expected that a stiff layer will form at the metal fuel/graphite interface which would hinder further diffusivity. Perhaps the biggest challenge of this design will be its fabrication. One way could be to braze the fuel and the foam first around
600 °C, then insert them together into the steel cladding. Another way would be to insert fuel into the steel cladding. Then, insert the foam into the gap region in small pieces as suggested by Ogata [14]. Another challenge appears to be the reprocessing of this fuel. Although the content of the carbon atoms is low due to high porosity, the inclusion of the carbide phases would complicate the usual reprocessing schemes. If once-through fuel cycle is preferred as suggested for breed-burn reactor concept, this should not cause any problem. Nevertheless, the fabrication and reprocessing aspects of the design problem are left as the future work. The first part of the paper introduces the reference metallic fuel pin description and an approximate neutronic analysis for the clad dose calculation. The second part will present the analyses of the thermo-mechanical behavior of the reference as well as the advanced metallic fuel performed with FEAST-METAL [15]. FEAST code is equipped with mechanistic models to simulate fission gas behavior, fuel clad chemical interaction, fuel constituent redistribution, and mechanical behavior of the fuel pin. The code is validated against EBR-II experimental database and in this work it is applied to an advanced design. The selected optimized configuration for the advanced fuel will be suggested using FEAST code results.
2. Description of the reference metallic fuel Selected properties of the metal fuel of near unity conversion ratio sodium fast reactor are given in Table 1. Typical EBR-II fuel pin geometry is adopted [1]. The zirconium content of the fuel is selected as 6 wt%. Most metallic fuels irradiated at EBR-II bear 10 wt% Zirconium; however, Refs. [16,17] showed that fuels containing 6 wt% and 10 wt% Zirconium acted fairly similar. T91 type
Table 1 Fuel composition, geometry and operating conditions for the reference metallic fuel. Property
Value
Fuel composition (wt%) Clad material Bond material Clad outer diameter (mm) Clad thickness (mm) Fuel smear density (%) Active fuel length (m) Pitch to diameter ratio Plenum to fuel volume ratio Peak burnup (at%) Peak dose (dpa) Coolant inlet temperature (°C) Peak clad temperature (°C) Peak linear heat rate (kW/m) Axial power profile Axial power peak
U–15Pu–6Zr T91 Sodium 5.84 0.38 75 1.0 1.2 2.0 15 200 390 660 35 Chopped cosine 1.25
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Fig. 3. Peak linear heat rate as a function of peak burnup (kW/m).
Fig. 5. Full core MCNP layout.
Fig. 4. Peak coolant outlet temperature as a function of peak burnup (°C).
steel [12] is selected as clad material over HT9 type steel in order to operate at clad temperatures as high as 660 °C without concerning about the thermal creep. The assumed clad dose limit is 200 dpa. Peak achievable burnup is 15 at% due to cladding dose limit. Finally, a large plenum to fuel volume ratio is adopted to accommodate for the fission gases. Operation with a plenum pressure above 10 MPa could deteriorate the high temperature transient performance (design basis accidents) significantly due to low thermal creep fracture margin of the ferritic–martensitic steels. Hence, a large plenum to fuel volume ratio is inevitable for this fuel pin design. Variation of the peak linear heat rate and peak coolant temperature within the assembly is given as a function of burnup in Figs. 3 and 4, respectively. Selected power and coolant temperature decreases linearly with burnup.
3. Neutronic analysis to assess the clad dose Neutronic analysis is a critical part of a fast reactor fuel design. Typically, the assessment of power and burnup distribution, neutronic feedback coefficients, and the dose on the cladding need to be performed. Note that the density of the foam is rather low and the effect of graphite bond on excess reactivity remains at a minor level. The present neutronic analyses were performed only to assess the clad dose for the selected fuels in an approximate manner. The dose equivalence of EBR-II core fast neutron fluence is known. Typically, 1E+22 n/cm2 (>0.1 MeV) fluence corresponds to 5 dpa clad dose [18]. The current study computed the neutron spectrum of EBR-II fuel and the reference fuel using MCNP-4C [19]. The MCNP full core model is given in Fig. 5. Inner assemblies consist of driver fuels and control assemblies. The peripheral regions are occupied by the blanket assemblies. By using EBR-II fuel properties, which is 66% enriched U–10Zr and the reference fuel properties given in Table 1, the driver fuel neutron flux spectrum was calculated for 35 kW/m linear heat rate conditions (Beginning
Fig. 6. Neutron flux spectrum for EBR-II fuel and U–15Pu–6Zr fuel at beginning of Life.
of Life). As depicted in Fig. 6, EBR-II operated with a much lower neutron flux level compared to a possible commercial fast reactor design. Although the fissile Pu isotopes are neutronically advantageous over U-235 at fast neutron spectrum, much larger fissile density of EBR-II lowers the neutron flux level. In addition, the neutron spectrum of EBR-II fuel is somewhat harder compared to the reference fuel again due to the large fissile density. Using energy deposition cross-sections given in Ref. [20], dose rate calculation of the reference fuel was performed in a relative manner. It was found that clad dose rate in typical unity conversion ratio sodium fast reactors is about 2.3 time higher compared to EBR-II conditions. 4. Thermo-mechanical analysis of the reference metallic fuel pin Steady state thermo-mechanical analysis of the reference U– 15Pu–6Zr metallic fuel with sodium bond was performed with FEAST-METAL code. Fig. 7 shows the axial variation of the clad inner temperature and fuel centerline temperature at the beginning of life. Upper axial part of the fuel operates at higher temperatures
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Fig. 11. Fission gas release for the reference fuel. Fig. 7. Fuel and clad temperatures at beginning of Life for the reference fuel.
Fig. 8. Axial variation of the clad inner wastage for the reference fuel.
Fig. 9. Clad hoop strain at end of Life for the reference fuel.
Fig. 10. Axial variation of the clad hoop stress as a function of peak burnup for the reference fuel.
and it is softer compared to lower axial sections. On the other hand, the operation of the clad at high temperatures leads to excessive lanthanide attack to the clad inner surface, formation of brittle
Fig. 12. Fuel swelling at the middle axial section for the reference fuel.
phases and the resulting clad thinning for more than 30% of the as-fabricated clad thickness, as shown in Fig. 8. Fig. 9 shows the axial variation of the clad hoop strain at end of life. Excessive thermal creep was observed towards the upper axial section mainly due to clad wastage formation and thinning due to formation of brittle phases at clad inner region. Lower axial sections were exposed to clad deformation via irradiation creep and a small amount of void swelling. Fig. 10 shows the axial variation of the clad hoop stress. Fuel clad mechanical interaction remains low at upper, hot, axial sections of the fuel pin. In these regions, the hoop stress is mainly due to plenum pressure. On the other hand, lower and middle sections of the fuel, which consists of stiff dual a + d phase at the outer radial region, have low compressibility, holding the contact pressure at a much higher level at high burnup. Figs. 11 and 12 shows the fission gas release and swelling behavior of the reference fuel, respectively. Operation of the fuel above half of the melting point leads to high diffusivity, large fission gas swelling and release starting at early irradiation. Solid fission product swelling linearly increases at the expense of pore shrinkage with the hot pressing mechanism under the external stress. Open pores are more compressible at the hotter section but they are not as compressible at colder sections of the fuel. Once the porosity concentration drops below 10%, fuel clad mechanical interaction rises significantly, as seen in Fig. 10. Predicted plenum pressure at the end of life is 8 MPa.
5. Thermo-mechanical analysis of the advanced metallic fuel pin The advanced fuel pin includes the graphite foam as the bonding material between fuel and clad. The operating conditions were kept the same as the reference fuel pin given in Table 1. The initial porosity of the foam is 80%. The foam is expected to yield under low external restraint, so that the fuel swelling and porosity forma-
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Fig. 13. Axial variation of the clad hoop strain at the end of Life for 65%, 70% and 75% smear density advanced fuels.
Fig. 14. Fuel and clad temperatures for 70% smear density advanced fuel.
Fig. 15. Swelling behavior of the middle region of the 70% smear density advanced fuel.
tion would be allowed without straining of the clad, excessively. The average behavior of the yield stress of the foam types given in Fig. 2 was used to calculate the compressibility and the stresses on the fuel/foam interface and foam/clad interface. Foam porosity is assumed to remain at plenum pressure. Irradiation creep and irradiation hardening effects of the foam are neglected for this preliminary study. FEAST-METAL was ran for 65%, 70% and 75% fuel smear density to simulate the advanced metallic fuel with 80% porous graphite foam. As can be seen in Fig. 13 and 75% smear density fuel caused a significant clad straining due to fuel swelling. For the new design, leaving a residue foam thickness at the end of life is inevitable; hence, keeping the fuel pin dimensions the same as the reference fuel, excessive straining of the clad is not surprising for the 75% smear density. This design does not allow enough space
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Fig. 16. Fission gas release behavior of 70% smear density fuel as a function of peak burnup for the advanced fuel.
Fig. 17. Axial variation of the foam porosity at beginning of Life and end of Life for 70% smear density advanced fuel.
Fig. 18. Fuel outer radius and clad inner radius axial variation at the end of Life for the advanced fuel.
for fuel to build the interconnected porosity especially at the lower sections at which the fuel is stiffer. The authors suggest to reduce the fuel smear density to 70% to eliminate/delay excessive fuel clad mechanical interaction (Fig. 13). Detailed results and discussion for 70% smear density, graphite bonded metallic fuel are given below. Fig. 14 shows the fuel and clad temperature variation at the beginning of life. Fuel temperature is higher compared to the reference case simulation (Fig. 7). It is due to (1) lower thermal conductivity of the graphite foam compared to sodium bond and (2) the absence of sodium logging into the fuel which improves effective fuel thermal conductivity. Note that the new design entirely eliminates the sodium infiltration into the fuel. Therefore, the sodium boiling should not cause any limitation to the fuel slug centerline temperature.
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ial parts and similar at the lower and middle axial sections compared with the reference case. 6. Conclusions and future work
Fig. 19. Axial variation of clad hoop stress as a function of peak burnup for 70% smear density advanced fuel.
The study suggests using a foam material as the novel barrier material between metallic fuel and steel clad to improve the chemical stability, to provide reasonable thermal conductance and compliancy to allow for fuel swelling without yielding the clad, excessively. Due to its favorable properties, the current study adopts a graphitic foam material as an example. Furthermore, using FEAST-METAL code for an advanced fuel design is accomplished. The fuel performance simulation results for a typical metallic fuel pin with sodium bond operating above 600–660 °C peak clad temperature demonstrate the necessity of a barrier material to hinder fuel cladding chemical interaction. Next, the simulation using the graphitic foam material between fuel and clad suggests lowering the fuel smear density to 70% to avoid excessive straining of the cladding. The new design is expected to improve the plant thermal efficiency and boost the long term transient and severe accident performance of the metal fueled sodium fast reactors. Further experimental studies as well as higher fidelity modeling and simulation efforts are needed to validate this behavior. Furthermore, the fabrication and reprocessing feasibility remains to be demonstrated. References
Fig. 20. Axial variation of the clad hoops strain and thermal creep strain for 70% smear density advanced fuel.
Higher available thermal activation resulted in an earlier start of fission gas release and higher rate of fuel swelling as seen in Figs. 15 and 16. Shortly after irradiation, interconnected open porosity forms and weakens the fuel slug. Fuel becomes both axially and radially restrained by the foam. Once the fuel open porosity drops below 10%, the fuel becomes less compressible and foam starts to shrink again under the applied compressive stress. The predicted plenum pressure at the end of life is 8.2 MPa. Figs. 17 and 18 show the foam porosity axial variation, and fuel slug outer radius and clad inner radius at the end of Life, respectively. The minimum residue foam thickness is on the order of 100 lm, which is expected by authors to be enough to prevent metallurgical interactions between the fuel and the clad. Figs. 19 and 20 shows the clad hoop stress and strain axial variations. At high burnup, the rate of change of clad hoop stress increased in parallel with the depletion of the fuel porosity. Compared to the reference case, the level of thermal creep is lower and acceptable. Hence, in this case, the clad hoop straining mostly occurs due to irradiation creep and a small amount of the void swelling. The total clad hoop strain is much lower at the upper ax-
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