Czech Technical University in Prague. Faculty of Nuclear Sciences and Physical Engineering. Department of Nuclear Reactors

Transcription

1 Czech Technical University in Prague Faculty of Nuclear Sciences and Physical Engineering Department of Nuclear Reactors Feasibility study on high-conversion Th-U233 fuel cycle for current generation of PWRs Ph.D. thesis Author: Ing. Daniela Baldová Supervisor: Emil Fridman, Ph.D. Co-supervisor: Ing. Jan Frýbort, Ph.D. Prague, 2014

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3 Abstract This thesis explores a possibility of designing a high conversion (HC) Th-U233 core for current generation of Pressurized Water Reactors (PWRs). Increasing the conversion ratio in existing PWRs can potentially improve the utilization of natural resources, through the exploitation of vast thorium reserves and reduction in natural uranium demand. HC can be achieved through the use of heterogeneous seed-blanket (SB) Th- U233 fuel assembly design where the supercritical seed works as a neutron supplier, while the subcritical blanket acts as U233 breeder. One of the main challenges associated with the heterogeneous SB fuel assembly designs is a significant power imbalance between the seed and blanket regions caused by the concentration of fissile material primarily in the seed zone and consequently requiring a substantial reduction in the core average power density. The main objectives of the thesis are twofold: 1) to design a high conversion SB Th-U233 fuel assembly which is directly retrofittable into existing PWRs without introducing significant modifications into the core and plant design; 2) to estimate the reasonably achievable core power density level at which reactor safety is not compromised. The first objective is accomplished by performing assembly-level parametric study. Based on the results, a number of SB fuel assembly configurations are selected for the following whole-core analysis. The reasonable achievable core power density level at which reactor safety is not compromised is estimated by performing 3D full core coupled neutronic and thermalhydraulic (T-H) analysis of a typical PWR core fully loaded with HC Th-U233 SB fuel. The results demonstrate that in principle the Th-U233 PWR core can be operated without exceeding the main safety limits at reduced power density of 60 W/cc in three-batch annual fuel cycle. However, it was shown that the sustainable mode cannot be fully realized. Still the results indicate that potential saving of natural resources can be achieved with HC Th-U233 fuel cycle. Keywords: High Conversion Fuel Cycle, Th-U233 fuel, Seed-blanket fuel assembly 1

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5 Declaration I do solemnly declare that I have written the presented research thesis by myself under the direction of my supervisor Emil Fridman, Ph.D, and without the aid of unfair or unauthorized resources. In Prague, Daniela Baldová 3

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7 Acknowledgements I would like to express my great appreciation to Dr. Emil Fridman, my supervisor, for his patient guidance, encouragement, and support in all aspects of this work. I am also grateful for the assistance and useful critiques given by Dr. Jan Frýbort (Faculty of Nuclear Sciences and Physical Engineering, CTU in Prague) during the thesis preparation. Special thanks to Dr. Sören Kliem, head of the department of Reactor Safety, HZDR, for providing me with all the necessary facilities and financial assistance during my stay here. Advice given by Mr. Reuven Rachamin has been a great help in my research. I am deeply thankful to Dr. Yurii Bilodid and Dr. Silvio Baier for their help not only with the DYN3D code. Finally, I would like to thank my family and my friends for their support and encouragement throughout my study. 5

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9 Table of contents Abstract... 1 List of figures... 9 List of tables Abbreviations Chapter 1. Introduction Background Overview of advantages and drawbacks of Th-based fuels Historical Review of the Th Use in LWRs Recent Studies on the Potential Use of Th fuel in LWRs Open vs. closed fuel cycle Homogeneous vs. heterogeneous fuel design Once-through fuel cycle aimed at reducing the proliferation Closed fuel cycle aimed at Pu incineration HC fuel cycle aimed at improvement of fuel utilization Thesis objectives Chapter 2. Analysis tools D level analysis tools D level analysis tools Chapter 3. Assembly level analysis Calculation methodology Neutronic analysis T-H analysis Description of the reference HC Th-U233 SB PWR fuel assembly Parametric neutronic and T-H calculations Selection of SB configurations Chapter 4. Full core analysis

10 4.1 Description of the HC Th-U233 PWR core design Core loading pattern T-H constraints Analysis methodology Homogenized cross sections generation Specification of the limiting T-H criteria Results of the 3D full core calculations Selected SB configurations Modified configurations Core average reactivity coefficients Chapter 5. Fuel cycle consideration Analysis methodology Material flow in HC SB fuel cycle Natural uranium utilization Chapter 6. Summary and future work Thesis Summary Future work References Appendix A A.1 Power density distribution A.2 T CL distribution A.3 MDNBR distribution A.4 Void fraction distribution

11 List of figures Fig The η-factor of U-233, U-235 and Pu-239 depending on incident neutron energy (based on JEFF-3.1.1) Fig Th transmutation chain (IAEA, 2005) Fig Schematic view of SBU and WASB assembly designs (IAEA, 2005) Fig Heterogeneous Th-U233 SB fuel assembly designs Fig Comparison of k-inf and FIR, HELIOS vs. SERPENT Fig Comparison of actinide concentration, HELIOS vs. SERPENT Fig Pin-by-pin power distribution for 1/8th FA at 0 MWd/kg Fig DYN3D structure Fig FIR vs. burnup, variable blanket pin radius Fig U233 mass vs. burnup, variable blanket pin radius Fig Power share between seed and blanket vs. burnup Fig Reference SB fuel assembly: radial layout Fig MDNBR vs. power density and T in Fig Maximum T CL vs. power density T in Fig Maximum T out vs. power density and T in Fig Maximum outlet void fraction vs. power density and T in Fig Power density, W/cc Case Fig Power density, W/cc Case Fig Power density, W/cc Case Fig Power density, W/cc Case Fig Core loading pattern Fig Nodalization of SB fuel assembly in DYN3D code Fig DYN3D cross section calculation scheme Fig Differences in k-inf, HELIOS vs. DYN3D Fig Pumping power vs. mass flow rate Fig Pressure drop vs. mass flow rate Fig CBC vs. burnup Fig CBC vs. burnup, Cases 3.1 and Fig Power density, W/cc Case Fig Power density, W/cc Case Fig Microscopic capture XS of Th232, U234, and U236 (ENDF B-VI.8)

12 Fig Material flow diagram for U recycling Fig Fissile material content in discharged fuel after 7 years cooling Fig Main fuel components mass flow during recycling Fig Flow diagram of power generation Fig. A.1. Power density, W/cc Case Fig. A.2. Power density, W/cc Case Fig. A.3. Power density, W/cc Case Fig. A.4. Power density, W/cc Case Fig. A.5. T CL, C Case Fig. A.6. T CL, C Case Fig. A.7. T CL Case Fig. A.8. T CL Case Fig. A.9. MDNBR Case Fig. A.10. MDNBR Case Fig. A.11. MDNBR Case Fig. A.12. MDNBR Case Fig. A.13. Void fraction Case Fig. A.14. Void fraction Case Fig. A.15. Void fraction Case Fig. A.16. Void fraction Case

13 List of tables Table 2.1: Summary of SB fuel assembly parameters Table 3.1: Reference PWR core: summary of the main design parameters Table 3.2: Th232 in blanket: one-group σ a and absorption reaction rates Table 3.3:Reference SB fuel assembly: summary of the main design parameters Table 3.4: FIR and initial seed U233 content vs. average power density and T ave Table 3.5: Summary of the selected SB configurations for the full core analysis Table 4.1: Summary of the branch-off parameters Table 4.2: FIR and discharge burnup, Cases Table 4.3: Initial U233 mass load Table 4.4: MDNBR calculated for hot channels at BOC Table 4.5: Summary of the limiting T-H parameters, four selected SB configurations (Cases 1-4) Table 4.6: Operating parameters of the modified SB configurations Table 4.7: FIR and discharge burnup, modified configurations (Cases 3.1 and 4.1).. 70 Table 4.8:Initial U233 mass load Table 4.9: Summary of the limiting T-H parameters, modified SB configurations (Cases 3.1 and 4.1) Table 4.10: Summary of the core average reactivity feedback parameters Table 5.1: Initial Pu isotopic vector Table 5.2: U isotopic vector in discharged fuel after 7 years cooling * Table 5.3: U mass balance Table 5.4: Summary of material flow at BOL and EOL for SB Cycles Table 5.5: Total energy generated per core Table 5.6: Mass balance * Table 6.1: Comparison of the major design parameters, reference all-u PWR vs. HC Th-U233 PWR

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15 Abbreviations 2D 3D BOL BU CHF CR DC DNBR EFPD EOL FIR GIF HM IAEA LWBR LWR MDNBR NLRM NU ONB pcm ppm PWR R&D RTF SB SBU T CL T-H WASB XS Two dimensional Three dimensional Beginning of Life Burnup Critical Heat Flux Conversion Ratio Doppler Coefficient Departure from Nucleate Boiling Ratio Effective Full Power Days End of Life Fissile Inventory Ratio Generation IV International Forum Heavy Metal International Atomic Energy Agency Light Water Breeder Reactor Light Water Reactor Minimum Departure from Nucleate Boiling Ratio Non-Linear Reactivity Model Natural Uranium Onset of Nucleate Boiling percent mili rho (Reactivity change of 10 5 ρ) parts per million Pressurized Water Reactor Research and Development Radkowsky Thorium Fuel Seed Blanket Seed Blanket Unit Central Line Temperature Thermal-Hydraulic Whole Assembly Seed and Blanket Cross-section 13

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17 1.1 Background Chapter 1. Introduction World population growth has been one of the issues of the recent decades in terms of future energy needs and of the environmental conditions of the planet. In order to preserve the prosperity and peace on our planet we have to keep energy available and affordable. Nowadays, there are basically three energy options: renewable, fossil, and nuclear. The renewable energy sources - especially wind and solar - are unreliable by nature of their occurrence and have a limited potential to satisfy world energy demand. However fossil fuel is presently the most used energy resource which has the ability to satisfy the present energy demand, but cannot neglect environmental hazard due to the production of CO 2. By contrast, nuclear energy is the largest source of emission-free energy. The nuclear energy is costeffective and has potential to meet ever-increasing demands for word's electricity supply. The nuclear energy, dominated by Light Water Reactors (LWRs) comprising Pressurized Water Reactors (PWRs) and Boiling Water Reactors (BWRs), is currently based on burning uranium fuel in a once-through fuel cycle achieving relatively low natural uranium (NU) utilization of about 0.6% (Greenspan, 2012). The question of better utilization of natural resources, safety, reliability, and proliferation-resistance, gave rise to two main international projects, namely Innovative Nuclear Reactors and Fuel Cycles Programme (INPRO) initiated by the IAEA and the US-led Generation IV International Forum (GIF). Sustainable utilization of natural resources is one of the main goals set by the GIF (OECD, 2014). Current and future generations of LWRs can be modified or designed to achieve or approach the sustainability goals. Taking advantage of proven technologies and existing infrastructure would minimize the R&D expenditures. For instance, by making use of advanced LWR designs with innovative fuel cycles a more practical and cost effective pathway could be opened, at least in the near term before fast reactor technology becomes sufficiently mature. Recent studies performed in Japan by JAEA and Hitachi (EPRI, 2012) showed that self-sustainable operation of BWR with respect to fissile material is possible. This was achieved by modifying a BWR including reducing the fuel pins lattice pitch, increasing the core average void fraction and utilizing an axially heterogeneous core structure. 15

18 However, the Th-U233 fuel is the only combination practically capable of high conversion (HC) in thermal neutron spectrum, as U233 has a neutron yield (η-factor) per neutron absorbed greater than 2.0 over a wide range of the thermal neutron energies, unlike the other common fissile isotopes U235 and Pu239 (Fig. 1.1). Since Th is 2 to 4 times more abundant than U (IAEA, 2005), Th-based fuels used in HC fuel cycle in existing PWRs can potentially improve the utilization of natural resources through the exploitation of vast Th reserves and reduction in NU demand. Fig The η-factor of U-233, U-235 and Pu-239 depending on incident neutron energy (based on JEFF-3.1.1) 1.2 Overview of advantages and drawbacks of Th-based fuels Besides the aforementioned advantages, there are several additional benefits of using Th as the fertile material (IAEA, 2005): - Th232 has higher thermal capture cross section in comparison to U238; therefore, has higher fertile to fissile conversion rates. - ThO 2 is chemically more stable and has higher radiation resistance than UO 2, which may allow longer in core-residence times. - ThO 2 has favorable thermo-physical properties because of the higher thermal conductivity and lower coefficient of thermal expansion as compared to UO 2. - The melting point of ThO 2 is about 500 C higher than that of UO 2. 16

19 - ThO 2 is inert and does not oxidize unlike UO 2. - Being a lighter element than U238, Th232 produces virtually no transuranium isotopes. The disadvantages Th as a fuel (IAEA, 2005): - Th does not contain any fissile isotope. Therefore, to startup the Th fuel cycle, an initial fissile driver component must be provided. - For production of ThO 2, a higher sintering temperature (>2000 C) is required than that of UO 2, due to its high melting point (3350 C). - ThO 2 is relatively inert and hence does not dissolve easily in concentrated nitric acid. Therefore, the reprocessing of Th based fuel is more complicated as compared to U based fuel. The chemical method used for reprocessing of spent Th based fuel is called THOREX (THORium-uranium Extraction) process (Gresky, 1955). Till date, the THOREX process was used mostly on laboratory and pilot reactor scale. Therefore, the experience with THOREX process is limited as compared to that of commercial PUREX process (Anderson and Asprey, 1960). - By irradiation of Th-based fuel, significant amount of U232 is created. U232 has half-life of only 69.8 years and is associated with strong gamma emitting radionuclides in its chain, Bi212 and Tl208 with very short half-life (Fig. 1.2) - This significantly complicates spent Th fuel handling operations such as reprocessing, fuel fabrication, transport, disposal, etc. - U233 generated from Th232 is potentially weapons usable material. However, high energy γ emitters mentioned above provide a certain degree of proliferation self-protection (Laughter et al., 2002). - The database and experience with Th fuel cycle are limited in comparison with that of U fuel cycle. - Existence of so-called protactinium effect. Pa233 may not only decay to the fissile U233, buy may be also converted to Pa234. This leads to the significant loss of new fissile material U233. Moreover, it causes reactivity increase after shutdown, when Pa234 is no longer created and all Pa233 decays to U

20 Fig Th transmutation chain (IAEA, 2005) 1.3 Historical Review of the Th Use in LWRs The feasibility of using Th as a fertile material in LWRs has been explored since 1960s and the interest in it has lasted to the mid-1970s. During this period, several Thbased fuel design options were investigated and some even resulted in whole core demonstrations. - BORAX-IV (2.5 MWe) BWR (IAEA, 2005) experimental core located in Idaho and operated by Argon National Laboratory between 1956 and 1958 demonstrated the feasibility of stable operation with ThO 2 -UO 2 fuel. - Elk River (IAEA, 2005) was a 24 MWe BWR ( ) built by Allis- Chalmers in Minnesota. The reactor operated satisfactorily but its limited power and cooling system corrosion led to its shut down. Both reactors, BORAX-IV and Elk River were using high density ThO 2 -UO 2 fuel pellets containing 4-7% UO 2. - Indian Point 1 (IAEA, 2005), a 285 MWe PWR ( ) built by Consolidated Edison in New York led the way in the use of fuel mixture of Th and highly enriched (93%) U in the first U233 breeding core. - The Shippingport Light Water Breeder Reactor was developed by Pittsburgh Naval Reactors Office in framework of the Light Water Breeder (LWBR) program which aimed to confirm the feasibility of breeding of net U233 in the thermal spectrum (Connors et al., 1978). The reactor core consisted of seed- 18

21 blanket fuel assemblies surrounded by 15 reflector assemblies. Every fuel assembly was subdivided into two standalone sub-assemblies: a movable central seed region containing U233 and the peripheral stationary blanket region containing mostly Th. In LWBR movable seed rods were introduced as a replacement for control rods in order to further improve the neutron economy, and minimize neutron losses. The reactivity was controlled by moving up and down seed sub-assembly inside the blanket sub-assembly. The reactor was operated from 1977 until 1982 and the fuel achieved a maximum burnup of 60 GWd/t. The design lifetime for LWBR was 1200 Effective Full Power Days (EFPD) assuming three batch fuel management. The core was operated with the power output of 60MWe. The results of analyses on spent LWBR fuel confirmed that the ratio of the fissile content of the fuel at the end of operation to that at the beginning of operation was about It should be mentioned that the LWBR program provided the most extensive experience with Th based fuel cycle in LWRs. Despite of its potential advantages, the interest in Th based fuel cycle declined significantly, and by the 1980s most of the projects with Th fuel cycle had been terminated. One of the reasons of this declined interest was that Th could not compete with the relatively cheap and well-established U fuel cycle. 1.4 Recent Studies on the Potential Use of Th fuel in LWRs In recent years there has been a renewal of interest in Th-based fuel cycles, particularly for commercial PWRs, primarily motivated by their potential: - To improve the proliferation resistance of spent fuel not only by producing lower amount of Pu but also by reducing its quality. - To efficiently incinerate excess reactor-grade Pu (RG-Pu) and weapons-grade Pu (WG-Pu) and to reduce long-term radioactivity and toxicity of the spent fuel storage. - To improve the fuel utilization by reducing the NU requirements. Several R&D projects were aimed at investigation of Th-based fuel cycles for LWRs. The projects differ not only in the extent and depth of the R&D effort (from simplified study to a detailed assembly/core/cycle design), but also in fuel cycle strategies (once-through cycles vs. reprocessing), and fuel design options (homogeneous fuel vs. heterogeneous fuel). Noting that introducing Th component of 19

22 the fuel necessitates the use of a fissile driver component, several options were considered: enriched U, RG-Pu, WG-Pu, and a mixture of transuranium isotopes (TRU). The specified fissile driver, cycle strategy, and fuel design were selected based on a design objective, usually related to one of the three aforementioned reasons for Th utilization (proliferation, reduction of long-term radioactivity, reduction of NU requirements) or their combination Open vs. closed fuel cycle The main design constraint, which influences the design solutions, is the choice of closed (reprocessing) vs. open (once-through) fuel cycle. In the case of the open cycle in LWRs, the U233 is bred and fissioned in situ, without involving chemical separation of U233; therefore, there is always need for make-up fissile material. The open fuel cycle avoids the complex reprocessing processes and other complications associated with re-fabrication of highly radiotoxic U233 based fuels. Thus, the benefit from the produced U233 depends on burnup only, which is required to be as high as possible. In the case of the closed fuel cycle, the bred U233 and Th are recovered by chemical reprocessing of the spent fuel. As shown in (Shapiro et al., 1977), the use of Th in closed fuel cycle can improve the utilization of natural resources and reduce associated costs. Contrary to aforementioned closed fuel cycle, the open cycle avoids the handling of dirty U233 outside the core and therefore is considered to be safe from both, an environmental and non-proliferation point of view. Due to the mentioned advantages, most of the recent studies have focused on the open fuel cycle Homogeneous vs. heterogeneous fuel design There are two main fuel assembly design options that have been under discussion in recent studies for the Th fuel cycle in LWRs: homogeneous and heterogeneous. In the homogeneous approach, the U fuel rods are directly replaced by a fuel rods containing either ThO 2 -UO 2 or ThO 2 -PuO 2 mixture. The heterogeneous approach employs seed-blanket (SB) fuel configuration, where U and Th fuel parts are spatially separated either within a given assembly, or between assemblies. Although the homogeneous concept is attractive by its simplicity (without the need for any modifications of the core parameters), the results of several investigations, indicated 20

23 clearly that for the open cycle option and a homogeneous design, NU savings cannot be achieved. For instance, a comprehensive study (Galperin et al., 2002) on the use of the homogeneous ThO 2 -UO 2 mixture in current PWR was carried out with the aim of reducing NU requirement and proliferation potential of the fuel cycle. From the obtained results it is evident that (in case of a similar discharged burnup) the increase of the proliferation resistance of ThO 2 -UO 2 mixture as compared to all-u fuel is only moderated. In addition, it was shown that a homogeneous ThO 2 -UO 2 mixture does not provide any significant improvement in the fuel cycle; on the contrary, it will increase the NU requirement, thus resulting in a waste of this natural resource. The similar conclusion was also found in studies (Gungor, 1990; Clarno, 1999). In (Hyung-Kook Joo et al., 2004), alternative applications of homogeneous Th-U fuel in LWRs were investigated with the aim of enhancing the economic potential of Th-based fuels. The three alternative ThO 2 -UO 2 fuel cycles were considered: a Th-U fuel with a relatively low U235 enrichment (< 20 wt%), a mixed core of duplex ThO 2 -UO 2 and UO 2 fuels, and use of Th-U fuel in a small and medium LWR with a 5-yr cycle length. The results of core analysis indicated that there is no improvement in NU utilization and that the ThO 2 -UO 2 fuel cannot compete economically with UO 2 fuel. However, it was shown that the U utilization factor for the Th-U cores increases with fuel cycle length, while those for the U decrease. Hence, it can be assumed that the further improvement in economy will be achieved, when Th based fuel is used with longer fuel cycle schemes. It should be mentioned that the mixed core of duplex ThO 2 -UO 2 and UO 2 fuels which, in fact, does not belong to homogeneous fuel category provides slight increase in NU utilization by 3 to 6% compared to ThO 2 -UO 2 and UO 2 cores. The homogeneous design studies have also embodied micro-heterogeneous design where the U and Th fuel are spatially separated within a given fuel rod. In (Shwageraus et al., 2005) it was demonstrated that the spatial separation of the U and Th parts of the fuel on the order of a few millimetres to a few centimetres can improve the achievable burnup of the Th-U fuel designs through more effective converting of Th232 to U233. In contrast to homogeneously mixed Th-U fuel, the micro-heterogeneous fuel has a potential to be competitive economically with commercial UO 2 fuel Once-through fuel cycle aimed at reducing the proliferation The investigations of Th-U homogeneous mixture in once-through fuel cycle indicate that in order to avoid penalty on NU utilization and associated fuel cycle 21

24 costs, the U and Th parts of the fuel should be spatially separated. Two heterogeneous fuel design options for the implementation of Th fuel cycle in LWRs have been examined under the Nuclear Energy Research Initiative (NERI) funded by the United States Department of Energy. The two approaches are: 1) Seed-Blanket Unit (SBU, also known as Radkowsky Th Fuel (RTF) developed at Ben-Gurion University (Israel) (Radkowsky and Galperin, 1998)), which can replace the conventional PWR fuel assembly one-for-one (Fig. 1.3); 2) the Whole Assembly Seed and Blanket (Todosow et al., 2005) (WASB, the concept developed at MIT (USA)), where seed and blanket units each form one full-size PWR assembly (Fig. 1.3). Fig Schematic view of SBU and WASB assembly designs (IAEA, 2005) The both fuel assembly designs were considered to be directly retrofittable into existing PWR assemblies with minimum changes in the core and plant design. Both the RTF and WASB concepts exhibit significant deterioration of the proliferation potential due to a substantial reduction in the produced Pu quantity and quality (the critical mass of RTF Pu is greater than for PWR Pu). For example, in case of RTF, the amount of Pu discharged per year can be reduced to 20% of that of a PWR. Moreover the RTF concept, as well as WASB concept, is competitive economically. Several studies also investigated whether Th based fuel possess advantages in an intermediate neutron spectrum (tight lattice). For instance, the results of (Iwamura et al., 1999) 22

25 showed that introducing Th based fuels designed in tight-pitch lattice do have several attractive characteristics, such as a more negative void coefficient and a higher fuel conversion ratio than corresponding U based fuels Closed fuel cycle aimed at Pu incineration As aforesaid, Th based fuel can be also used for the purpose of incineration and or stabilization of excess RG-Pu and WG-Pu. The main objective of studies performed with focus on Th based cycles for reprocessing option was to maximize Pu destruction. The results of the studies reported in (Galperin et al., 2000; Shwageraus et al., 2004; Fridman and Kliem, 2011) demonstrated that both, homogeneous and heterogeneous fuel designs can achieve better Pu destruction rates and lower quality residual Pu vector in spent fuel than conventional U-Pu mixed oxide (MOX) fuel. For instance, in (Galperin et al., 2000) a heterogeneous SBU Pu-incinerator (also known as Radkowsky Th-Pu Incinerator - RTPI) was proposed for incineration of RG-Pu and WG-Pu and the results indicated that the Pu incineration rate is three times higher than that of the MOX core. In (Shwageraus et al., 2004) the Pu and transuranium (TRU) destruction efficiency of homogeneous Th-based fuels was investigated on a single fuel assembly level. The results of the study indicated that in case of Th-Pu fuel up to 75% of initially loaded Pu can be incinerated while in case of Th-TRU fuel the destruction efficiency of total TRU is limited by about 50%. The results of the full core analysis in (Fridman and Kliem, 2011) showed that in Th-Pu mixed oxide (TOX) PWR core, the Pu consumption is almost doubled as compared to that of MOX PWR core HC fuel cycle aimed at improvement of fuel utilization HC Th-U233 fuel cycle implemented in the existing reactors can potentially improve the utilization of natural resources through the exploitation of vast Th resources and the reduction in NU demand, since the Th-U233 fuel is theoretically capable of achieving a HC ratio in the thermal neutron spectrum. The possibility of thorium utilization in HC fuel cycle in PWRs has been investigated in a number of scoping neutronic studies ( Kotlyar and Shwageraus, 2012; Shwageraus et al., 2009; Yun et al., 2010). It was demonstrated through the assembly-level fuel depletion analysis that the HC ratio in PWRs can be achieved, in principle, via the use of heterogeneous Th-U233 SB fuel assembly designs, in which the fissile and fertile 23

26 materials are spatially separated from each other. In (Shwageraus et al., 2009) the SB fuel assembly design (Fig. 1.4.a) featuring a reduced assembly size, duplex seed pins, and a thorium hydride blanket fuel was developed. In the SB design investigated in (Yun et al., 2010), the SB fuel assemblies have typical PWR outer dimensions and the seed fuel pins do not form a distinct region in a fuel assembly but rather distributed among the blanket pins, as seen in Fig. 1.4.b. In (Kotlyar and Shwageraus, 2012), the arrangement of the seed and ThH 2 blankets fuel pins in (Fig. 1.4.c) and (Fig. 1.4.d) fuel assemblies was optimized using a simulated annealing approach. a (Shwageraus et al.,2009) b (Yun et al., 2010) c d (Kotlyar and Shwageraus, 2012) Fig Heterogeneous Th-U233 SB fuel assembly designs 24

27 1.5 Thesis objectives One of the main challenges associated with the aforementioned heterogeneous SB fuel assembly designs is a significant power imbalance between the seed and blanket regions caused by the concentration of fissile material mostly in the seed zone. A high power peaking in the seed region will most likely lead to exceeding the thermalhydraulic (T-H) safety limits and therefore will necessitate a substantial reduction in the average core power density. However, despite their importance, T-H design and safety related issues were typically not considered in the earlier works. An attempt to estimate the achievable power density levels was done in a previous T-H feasibility study (Volaski et al., 2009) where a simplified T-H analysis of a single SB fuel assembly was performed without taking the effect of the T-H feedback on neutronics into account. The main objectives of this study are: - To propose a design of HC SB Th-U233 fuel assembly which can be directly retrofit into existing PWRs without introducing significant modification into the core and plant design. - To estimate the reasonably achievable power density level of PWR core fully loaded with HC Th-U233 SB fuel at which reactor safety is not compromised. - To demonstrate feasible routes for generation of initial and makeup U To estimate potential savings of available resources that can be achieved with HC Th-U233 fuel cycle. The objectives will be pursued according to the following steps: - Assembly level depletion analysis of a number of SB configurations assuming several reduced power density levels and different inlet coolant temperatures. - 3D single assembly thermal-hydraulic (T-H) analysis aiming at the evaluation of major safety related parameters. - Based on the results of the parametric studies described by previous two steps, a number of SB configurations will be selected for the subsequent more realistic full core analyses. - Two-group cross sections (XS) generation for the selected SB configurations. 25

28 - Steady-state whole core coupled neutronic and T-H analysis of 100% Th- U233 fueled PWR aiming at the confirmation of the assembly level results with regards to the practically achievable power density. - Assembly level depletion and decay analysis including U recycling. The results of this study will be used to assess the basic operational feasibility of the HC Th-U233 PWR core from the neutronic and T-H point of view. 26

29 Chapter 2. Analysis tools This chapter presents the computational tools used in the dissertation D level analysis tools In this study, all 2D assembly level calculations including the generation of homogenized two-group XS for a 3D full core analysis were performed using HELIOS HELIOS is a well-known commercial deterministic neutron transport lattice code developed by Studsvik Scandpower (Studsvik Scandpower, 2008). The HELIOS transport solver is based on current coupling and collision probabilities (CCCP) method. The HELIOS code employs the 190-group cross section library based on ENDF/B-VI evaluated data files. The system to be calculated in the HELIOS code (e.g. a fuel assembly) is built up by a cumulative coupling of the space elements (structures such as pin cells) by means of the interface currents, and the properties of the space elements are obtained from collision probabilities. These probabilities are evaluated by assuming the flat-flux approximation inside the sub-regions of the space elements. These sub-regions, which are in fact closed material areas, may be further subdivided in order to obtain a satisfactory spatial mesh for the flat-flux approximation. The angular dependence of the aforementioned interface currents is taken into account by subdividing the structures into sectors, where the cosine currents are assumed. This angular representation between two structures is specified by choosing one of the coupling options (k) available in HELIOS. Presently, k ranges between -5 and 18, where k equals to 0 represents the most accurate but expensive option of direct collision probabilities over the entire geometry. The HELIOS code capability for analyzing homogeneous Th-based fuels was already demonstrated in the past with the help of the code-to-code benchmark (Mittag and Kliem, 2011) and comparison with the experimental data (Björk et al., 2013) in the framework of the European LWR-DEPUTY project (Verwerft, 2007). However, the HC fuel assembly designs investigated in this dissertation are expected to be highly heterogeneous. For that reason, further verification of the HELIOS code regarding this matter is necessary. In order to demonstrate the applicability of the HELIOS code for the analysis of heterogeneous Th-U233 fuel assembly designs, HELIOS was benchmarked against 27

30 the SERPENT code (Leppänen, 2013). SERPENT is a 3D continuous-energy Monte Carlo (MC) reactor physics code, recently developed at VTT Technical Research centre of Finland. The SERPENT code was especially developed for the reactor physics application. SERPENT runs significantly faster than other MC codes. The reasons for that are twofold: first, the use of the Woodcock delta-tracking (Leppänen, 2010) in a combination with a typical surface-to-surface ray-tracing in a geometry routine, and second, the use of the unionized energy grid (Leppänen, 2009) for all point-wise reaction cross sections. The SERPENT code has built-in decay and burnup routine which is based on the solving Bateman equations using the Chebyshev Rational Approximation method (CRAM) (Pusa, 2011). Each SERPENT update is benchmarked against MCNP by running a standard set of assembly calculation problems. The SERPENT used in this study was validated by comparision to MCNP5 (Jeremy, 2003). The benchmark was carried out for six different geometry types (Leppänen, 2009a). The number of neutron histories was same for both codes. It was shown that the difference in effective multiplication factor is in the order of statistical accuracy from the reference results. Likewise, the discrepency in 4-group constants does not exceed significantly the statistical accuracy from the reference results. In general, the results show that the differences between the codes are well within the range of statistical accuracy and that the SERPENT code performes faster - in some of the investigated fuel types 50 times faster - than the MCNP code. As mentioned, the SERPENT code has a burnup solver, and for that matter it can be used to verify the HELIOS code. A series of burnup calculations were performed for heterogeneous Th-U233 fuel assembly design adopted from the Volaski (2009). The considered fuel assembly has an pin lattice, and is subdivided into the two regions, namely, seed and blanket. The geometry and operating conditions of the fuel assembly are summarized in Table 2.1. The assembly layout is shown in Fig All the fuel lattice calculations were carried out with the reflection condition. The SERPENT calculations were performed employing ENDF/B-VI.8 based XS library and using 100,000 neutron histories, 20 skipped and 250 active cycles. The HELIOS calculations were executed with a fixed resonance category parameter RES = 6 and coupling parameter k = 4. The RES chosen value is suggested in the HELIOS manual for the treatment of Th232 resonances (Studsvik Scandpower, 2008). In the both codes, the seed and blanket fuel pins were subdivide into two radial regions with equal volume (Fig. 2.2). In HELIOS pin cell model, the coolant zone was 28

31 subdivided into eight azimuthal regions, as illustrated in Fig. 2.2.a, in order to cope with flat-flux approximation and to account for spatial variation in neutron flux in the coolant. Fig SB assembly layout a. HELIOS b. SERPENT Fig Seed and blanket pin cell nodalization The compared parameters included burnup dependent k-inf, concentration of important actinides, and ratio of the instantaneous to initial fissile nuclide inventory (fissile inventory ratio - FIR). Pin-by-pin power distribution at beginning of life (BOL) predicted by the codes was also compared. 29

32 Table 2.1: Summary of SB fuel assembly parameters Assembly array Array geometry square Square Number of fuel rods per assembly 121 Fuel pin pitch, cm 1.26 Power density, W/cc 70 Tfuel, K 900 Tclad and Tcoolant, K 600 Inner Seed Radius, cm Inner Seed Fuel (Th,U233)O 2 Outer Seed Radius, cm Outer Seed Fuel ThO 2 Blanket Radius, cm 0.55 Blanket Fuel ThH 2 The results of the SB fuel assembly benchmark calculations are presented in Fig. 2.3 Fig.2.5. The discrepancy in k-inf is about 0.35% at BOL, and drops to less than 0.15% at end of life (EOL) as shown in Fig. 2.3.b. The maximum difference in the FIR is about 0.2% and was observed at the EOL (Fig. 2.3.b). The differences in isotopic concentration presented for Th232 and U233 in Fig 2.4.c is lower than 0.01% and 0.25%, respectively, over the whole burnup. In case of Pa233, the maximum discrepancy is notably higher and reaches 2% (Fig. 2.4.c). This difference can be partially explained by its relatively low concentration in the fuel, as shown in Fig. 2.4.a. The BOL pin power distribution calculated by Serpent and HELIOS is presented in Fig 2.5. In the seed region hosting most of the initial fissile material the maximum deviation is about 0.8%. Somewhat higher maximum discrepancy of about 3% was observed in the low power blanket region. The HELIOS results show generally good agreement when compared against the SERPENT code. Therefore, it can be concluded that HELIOS is a suitable tool for the analysis of heterogeneous Th-U233 fuel assembly configurations. Based on the results, the HELIOS model was adjusted for the next studies as follows: resonance category RES = 6, coupling k = 4, and pin-cell nodalization according to Fig. 2.2.a, which means eight azimuthal regions in the moderator zone, one radial region in cladding, and two radial regions in fuel. 30

33 a. K-inf and FIR as a function of burnup, HELIOS b. Relative difference in k-inf and FIR as a function of burnup Fig Comparison of k-inf and FIR, HELIOS vs. SERPENT 31

34 a. Concentration of U233 and Pa233 as a function of burnup b. Concentration of Th232 as a function of burnup c. Relative difference in actinide concentration as a function of burnup Fig Comparison of actinide concentration, HELIOS vs. SERPENT 32

35 SERPENT HELIOS 2.86% 2.86% Absolute error % 2.95% 3.01% % % 2.69% 2.18% % 1.83% 1.43% 0.26% % 1.15% 0.38% 0.52% 0.76% % 1.22% 0.20% 0.65% 0.09% 0.12% Max. error in Blanket Max. error in Seed Fig Pin-by-pin power distribution for 1/8 th FA at 0 MWd/kg 2.2 3D level analysis tools The 3D full core analysis was performed using the nodal diffusion code DYN3D (Grundmann et al., 2000) developed for 3D steady state and transient core calculations in square and hexagonal fuel element geometry with thermal-hydraulic (T-H) feedback. The DYN3D code is also capable of performing burnup calculation. The neutronic kinetic model is based on the solution of time dependent 3D multigroup diffusion equation using nodal expansion method. The two-phase T-H module is based on the solution of four differential balance equations for mass, energy and momentum of the two-phase mixture and the mass balance for the vapor phase in parallel coolant channels. A scheme of the DYN3D is shown in Fig DYN3D is a best estimate code and is also one of the reference 3D core simulators within a European simulation platform for nuclear reactor safety NURESIM (Chauliac et al., 2011). The DYN3D code was extensively verified and validated via numerous numerical and experimental benchmark problems. For example the DYN3D code was applied in NEA CRP benchmarks for PWR and BWR (Grundmann and Rohde, 1996), the Three-Miles-Island-1 main steam line break and the Peach Bottom Turbine Trip benchmarks (Grundmann and Kliem, 2003; Grundmann al., 2004), and others (Rohde et al., 2009). Results of verification and 33

36 validation activities demonstrated that DYN3D is an effective tool for steady-state and transient core analyses. Fig DYN3D structure 34

37 Chapter 3. Assembly level analysis This chapter presents the results of the assembly-level parametric study aiming at the selection of a number of SB fuel assembly configurations for the following wholecore analysis. The assembly configurations are selected according to their potential to satisfy the specified fuel cycle requirements and comply with the T-H safety limits. 3.1 Calculation methodology This sub-section refers to the calculation methodology of the neutronic parameters that were used during the neutronic optimization, and of the T-H parameters used for the final nomination of optimized SB fuel assembly configurations Neutronic analysis As previously mentioned, all neutronic and burnup calculations on 2D assembly level were performed with the HELIOS code. The results of single assembly burnup calculations were used to estimate the behavior of a full core by applying Non-Linear Reactivity Model (NLRM) (Driscoll et al., 1990; Fridman et al., 2007). The NLRM is typically used for approximate calculation of discharged burnup, reactivity, and fuel cycle length of cores loaded with advanced fuels having non-linear dependence of reactivity on burnup. In such cases, the calculated reactivity data versus burnup are interpolated by a polynomial. NLRM assumes equal power share between the fuel batches in the core. Therefore, the core reactivity is calculated as an average of the reactivity values of each fuel batch corrected for leakage. The leakage reactivity worth of 2% was used in the NLRM. This value was estimated from a simplified full core analysis. In this study, the 3 rd order polynomial was used to fit the reactivity versus burnup data, and the reactivity was calculated according to following expression: 2 3 (BU) = A0 + A1 BU + A2 BU + A3 BU (3.1) where represents the reactivity and BU the burnup. The expansion coefficients A i ( i=0,..,3 ) were obtained with a help of the MATLAB function polyfit. We assumed the equal power sharing between all fuel batches within the core, and we set the average core reactivity corrected for leakage (ρ leakage ) to zero at the end of cycle (EOC). The 35

38 burnup accumulated by each batch in one cycle (BU C ) can be calculated from the following equation: (BUC ) + (2 BUC) + (3 BUC) core (EOC) = - leakage 0 (3.2) 3 where the discharged burnup is equal to EOC burnup of the last batch in our case 3 rd batch, accordingly; DISCH BU = 3 BUC (3.3) The conversion performance of SB fuel assemblies was assessed by evaluating the FIR at EOL. The FIR, calculated as a ratio of instantaneous to initial fissile nuclide inventory, estimates the Th232 to U233 conversion performance and therefore serves as the main feasibility criterion of the HC fuel cycle. In FIR calculations, the instantaneous fissile inventory also included Pa233 due to its relatively fast conversion into U233 through the β-decay with the half-life of about 27 days. Fissile Nuclides Inventory at time (t) FIR = (3.4) Initial Fissile Nuclides Inventory The neutronic calculations for the reference SB fuel assembly were performed assuming reduced core average power densities of 70 W/cc. This value was adopted as an initial guess from the previous T-H study (Volaski et al., 2009) performed on a single fuel assembly channel level. However, it must be mentioned that the study analyzed somewhat different assembly with reduced radial dimensions, duplex seed pins and ThH 2 blanket fuel. Moreover, the core radial power profile was not taken into account in the analysis. Therefore, further reduction in the power density level down to 55 W/cc was considered in the parametric neutronic and T-H calculations T-H analysis The T-H behavior of the SB fuel assemblies was analyzed using T-H module of DYN3D. At this stage, the calculations were performed with a given power for a single 3D fuel assembly. Every fuel pin was modeled as a separate T-H channel. The 3D power distribution was calculated as a product of the pin-wise radial power profile obtained from HELIOS simulations, core axial power distribution, core radial power 36

39 peak, and finally, the average power density in question. Chopped cosine shape axial power profile was used in the analysis. The maximum core radial power peak was assumed to be equal to 1.3. The objective of this simplified T-H analysis was to select the SB fuel assembly configurations for the full core analysis through an evaluation of following T-H parameters: center line fuel temperature (T CL ), Departure from Nucleate Boiling Ratio (DNBR), and void fraction. The T CL refers to the highest temperature occurring at the central line of the solid fuel and it is an important indicator of fuel melting which should be avoided at all normal operating conditions and during design base accidents. The void fraction is defined as fraction of channel volume that is occupied by the gas phase. Although there is no any exact limit for the void fraction, the only subcooled boiling can be accepted. The sub-cooled boiling occurs when the surface temperature is higher than the temperature of the saturated fluid. The location where the vapour can exist in a stable state on the heated surface without condensing is designated as the Onset of Nucleate boiling (ONB). Departure from Nucleate Boiling (DNB) describes the phenomenon of transition from nucleate to film boiling which occurs when the heat flux reaches some critical value. Critical heat flux (CHF) is the heat flux at which the vapour film around a heated surface is formed and, subsequently, boiling crisis takes place. The vapour film, reducing the surface-to-coolant heat transfer, causes prompt increase in the fuel rod surface temperature and, consequently, a failure of the cladding material. The DYN3D code calculates directly DNBR as a ratio of the CHF and a local heat flux by using one of the correlations IAE-4 (Osmachkin, 1978), OKB-2 (Bezrukov et al., 1976), Pernica (Pernica and Čižek, 1992) or BIASI (Biasi et al., 1967). In this work, DNBR values were obtained using the Pernica CHF model. It is based on approximately 18,000 experimental data points in about 350 test geometries with both relatively wide and very tight rod lattices. The range of validity of Pernica CHF model with respect to pressure (p), coolant mass flux (G), steam quality (x), and pitch-to-diameter (P/D) ratio is 0.26 < p < 18.7 MPa, 34 < G < 7500 kg/m 2 s, <x < 0.99, and 1.02 < s < The P/D ratios for seed and blanket lattices are equal to and respectively. Since these values are within the validity range of Pernica model, it appears to be applicable to CHF prediction in SB fuel assemblies. 37

40 The thermal conductivity of ThO 2 base fuel is higher than that of UO 2 base fuel over most of the temperature range of interest (Belle and Berman, 1984); however, the ThO 2 -UO 2 mixture has the thermal conductivity comparable with that of the pure UO 2. Therefore, in this study, the thermal conductivity of ThO 2 -UO 2 mixture was conservatively presumed to be equal to that of UO Description of the reference HC Th-U233 SB PWR fuel assembly One of the main design goals in the development of HC SB fuel assembly is to maximize FIR, while meeting fuel cycle and safety requirements. The optimization of SB assembly design is a complex multi-objective problem, which involves many independent parameters (e.g. fuel pin cell and fuel assembly lattice pitch dimensions, number of seed and blanket fuel rods, their radial dimensions, fissile content in seed and blanket, average power density, etc.). Performing the full scale optimization is very computationally expensive, iterative procedure, requiring a large number of 2D fuel lattice and 3D full core calculations. It should be stressed that the selected reference SB design is not the optimized one, but rather one of the several possible configurations. It was developed by fixing some of the independent parameters as well as by taking the previous experience into account. This section presents the reference SB fuel assembly design and explains rationale behind the selection of the main operating parameters. In this study, a large Westinghouse PWR core with available operating data (Galperin et al., 1995) was selected as a reference for investigation into the operational performance of HC Th-U233 fuel. The summary of the main operating parameters of the reference PWR is presented in Table 3.1. The SB fuel assembly is required to be directly retrofittable into the reference PWR without introducing any major changes into the core and plant design. For that reason, the proposed SB fuel assembly has standard outer dimensions, a typical pin lattice and 25 guide tubes at the usual locations. A target fuel cycle length was fixed to twelve months assuming 3-batch reloading scheme and a capacity factor of 90%. This is equivalent to the fuel cycle length of 330 effective full power days (EFPD). The relatively short fuel cycle was considered because of the degradation of FIR with burnup as observed in the previous study (Shwageraus et al., 2009). 38

41 Table 3.1: Reference PWR core: summary of the main design parameters Total thermal output, MW 3358 Number of fuel assemblies 193 Average core power density, W/cc System pressure, bar 155 Total core flow rate, kg/s Core inlet temperature, ºC Active fuel height, cm 366 Assembly array Number of fuel rods per assembly 264 Assembly pitch, cm 21.5 Fuel rod pitch, cm 1.26 Fuel pellet radius, cm Cladding outer radius, cm Cladding material Zircaloy Number of guide tubes 25 Guide tube inner radius, cm Guide tube outer radius, cm Earlier studies suggested using Th hydride as a blanket fuel form because the presence of hydrogen in the fuel promotes neutron captures in Th. However, softer spectrum also increases the burnup rate of generated U233, resulting, on the overall balance, in lower FIR at the EOL. Therefore, in this study, an oxide fuel form was selected also because of the substantial experience accumulated with thoria in the past. As mentioned earlier, the transition from homogeneous to heterogeneous fuel assembly layout is a crucial condition for achieving high conversion. Therefore, the fuel assembly was subdivided into the two rectangular seed and blanket regions. Most of the U233 is concentrated in the central 9 9 seed, while remaining 191 fuel pins, with significantly lower than in seed U233 content, form the blanket region. The number of the pins in the regions was selected to approximately preserve the seed to blanket volumetric ratio close to that of the optimized SB design reported in (Shwageraus et al., 2009). It is worth mentioning that the square seed region configuration may be not the optimal one from a neutronics point of view, as was shown in (Kotlyar and Shwageraus, 2012). Nevertheless, the square layout was selected to enable subsequent 3D full core analysis with the nodal diffusion DYN3D 39

42 code, which is only capable of modeling reactor cores with regular square or rectangular mesh. While the seed fuel pins have typical PWR dimensions, the blanket fuel pin radius was enlarged in order to enhance the conversion performance of the SB fuel assembly. In order to demonstrate the effect of the blanket pin radius on FIR, two burnup calculations were performed with the blanket pin radius increased from cm (i.e. typical PWR value) to cm. In both cases, the U233 content in the seed was adjusted to achieve the target fuel cycle length of 330 EFPD. Fig. 3.1 shows that an increase in blanket radius results in higher FIR during the entire irradiation period. The burnup depended concentration of U233 for both cases is presented in Fig 3.2 showing somewhat steeper U233 generation rate in the blanket with enlarged radius. This increase in FIR is a trade-off between two effects. On the one hand, the increase in blanket dimensions reduces the moderation, hardens the neutron spectrum, and decreases effective σ a of Th232 in the blanket (Table 3.2). On the other hand, the total absorption rate in the blanket Th232 increases with the radius (Table 3.2). This is because larger blanket pins have significantly higher volume and, consequently, higher Th232 mass. Fig FIR vs. burnup, variable blanket pin radius 40

43 a. Blanket radius = cm b. Blanket radius = cm Fig U233 mass vs. burnup, variable blanket pin radius 41

44 Table 3.2: Th232 in blanket: one-group σ a and absorption reaction rates Blanket pin radius, cm Blanket pin volume (V), cm σ a, barn Σ a φ, 1/cm 3 s 3.91E E+14 Σ a φ V, 1/s 7.54E E+16 An additional series of burnup calculations was performed with the blanket U233 content varying from 2.33 to 0.25 w/o. In every case, the seed U233 content was modified to preserve the target fuel cycle length. Reducing the blanket U233 content decreases the reactivity swing (Fig. 3.3) and improves the FIR (Fig. 3.4). However, reducing the blanket U233 content below 0.5 w/o may lead to sub-critical core (Fig. 3.3). Therefore, starting from this point, the blanket U233 content was fixed at 0.5 w/o. Fig. 3.5 shows the seed-blanket power share as a function of burnup for the case with 9.55 and 0.5 w/o U233 content in seed and blanket, respectively. The power imbalance at BOL can be clearly observed while the seed is generating about 75% of the power in the assembly. The power imbalance between seed and blanket reduces with burnup and diminishes entirely towards the EOL. Fig Core reactivity vs. burnup, variable seed and blanket U233 content 42

45 Fig Fuel Assembly FIR vs. burnup, variable seed and blanket U233 content Fig Power share between seed and blanket vs. burnup 43

46 The reference SB fuel assembly is presented schematically in Fig 3.6. Summary of the main assembly design parameters is given in Table 3.3. Fig Reference SB fuel assembly: radial layout Table 3.3: Reference SB fuel assembly: summary of the main design parameters Pin pitch, cm 1.26 Seed fuel pellet radius, cm Seed cladding outer radius, cm Blanket fuel pellet radius, cm Blanket cladding outer radius, cm Guide tube inner radius, cm Guide tube outer radius, cm Seed/Blanket fuel material U233O 2 -ThO 2 U233 content in blanket, w/o 0.5 Number of seed pins 72 Number of blanket pins 192 Number of guide tubes 25 44

47 3.3 Parametric neutronic and T-H calculations This section reports on the selection of SB configurations for the following 3D full core analysis. The task was accomplished using the following two-step procedure. At the first stage, depletion analysis of the SB assembly was performed assuming different core average power densities of 55, 60, 65, and 70 W/cc. The reference PWR core has somewhat lower than typical coolant inlet temperature (T in ) of C (Table 3.1). Therefore, two additional inlet coolant temperature values of 275.0, and C were considered in the analysis and calculations were performed using average coolant temperatures (T ave ) of 280.0, 290.0, and C approximately corresponding to the inlet coolant temperatures mentioned above. For every combination of average power density and T ave, the fissile content was adjusted to achieve the target fuel cycle length of 330 EFPD. At the next stage, 3D single assembly T-H analysis of all aforementioned cases was performed in order to determine the most promising combinations of power density level and T in at which the safety limits are not exceeded. The following limiting T-H criteria were specified: - Maximum acceptable T CL was set to 3100 C which is 200 C below the melting point of the Th-U oxide (Belle and Berman, 1984). - Minimum DNBR (MDNBR) must be at least 1.3, which is a limit typically used in the PWR thermal analysis. - The maximum temperature of the coolant bulk should be below saturation temperature allowing for a limited sub-cooled boiling only. As described in Section 3.2, blanket fuel pins are enlarged as compared to seed fuel pins. Consequently, blanket lattice has somewhat higher flow resistance as compared to the standard UO2 and seed fuel pins resulting in a higher overall pressure drop over the active core, Δp. Since pumping power is proportional to Δp and the volumetric flow rate ( ), preserving the reference flow rate will lead to increase in pumping power and will require the uprate of the primary coolant pumps. In order to avoid any significant modifications of the plant infrastructure, it was decided to preserve the pumping power of the Th-U233 core equal to that of the reference PWR core by adjusting the total coolant mass flow rate ( ). The results of depletion calculations are summarized in Table 3.4. In all cases, FIR values close to unity were observed. Increasing the average coolant temperature 45

48 reduces the moderation, increases the resonance absorption, and slightly improves FIR. However, the initial U233 content required to achieve the desired fuel cycle length also increases with the coolant temperature. The results presented in Table 4 suggest that neutronic design of HC Th-U233 fuel cycle is feasible. Table 3.4: FIR and initial seed U233 content vs. average power density and T ave Case number Power density, W/cc T ave, C U233 content in seed, w/o FIR at EOL The safety related T-H parameters including MDNBR, maximum T CL, outlet coolant temperature (T out ), and void fraction are plotted as functions of T in and power density in Fig. 3.7, Fig. 3.8, Fig. 3.9, and Fig. 3.10, respectively, at both BOL and EOL. Typical PWR values are given for comparison purposes. The results of T-H analyses can be summarized as follows: a) BOL - 70 W/cc o DNB occurs regardless of the coolant T in (Fig. 3.7a). o For T in values of 275 and 289 C the coolant at the outlet reaches its saturation temperature (Fig. 3.9.a) and the corresponding maximum void fraction is as high as 29 and 61% respectively W/cc o DNB is avoided only for T in = 265 C case (Fig. 3.7.a). 46

49 o The average outlet coolant temperature reaches saturation for T in = 289 C (Fig. 3.9a). In this case, the maximum outlet void fraction is about 47% (Fig a) W/cc o DNB and the coolant bulk boiling at the outlet are observed only for T in = 289 C case. o In this case, MDNBR is equal to 1.23 and the maximum outlet void fraction is about 27% W/cc o All T-H parameters are within the specified limits for all considered T in values. - For all power densities, there are sufficient margins to fuel melt (Fig. 3.8.a). - For SB cases that comply with the specified T-H limits, the margin to DNB becomes tighter as compared to a typical UO2 case. b) EOL - The T-H limits were not exceeded for any considered T in and power density level. 47

50 a. BOL b. EOL Fig MDNBR vs. power density and T in 48

51 a. BOL b. EOL Fig Maximum T CL vs. power density T in 49

52 a. BOL b. EOL Fig Maximum T out vs. power density and T in 50

53 a. BOL b. EOL Fig Maximum outlet void fraction vs. power density and T in 51

54 3.4 Selection of SB configurations Based on the results of the neutronic and T-H parametric studies, four SB configurations were selected for the following 3D full core analysis. The selected cases are presented in Table 3.5. First, three cases (one for 55, 60, and 65 W/cc power density levels) were chosen as a trade-off between the potential to meet the specified safety criteria and the capability to achieve maximum FIR at discharge. The T-H analysis was performed assuming, somewhat arbitrary, core power peaking factor of 1.3. However, it may be possible to optimize the core loading pattern in a way that the power peaking factors in the core regions hosting fresh SB fuel assemblies will be sufficiently lower than 1.3 and, in this way, to assure satisfactory DNB margin. Therefore, an additional configuration (70 W/cc, T in = 265 C) was considered despite the slight violation of MDNBR criterion. The pin by pin power density distribution of the selected cases is presented in Fig Fig The maximum power density for Case 1 reaches 230 and 135 W/cc at BOL and EOL (Fig. 3.11), respectively, whereas the maximum power density for Case 4 is around 179 and 110 W/cc (Fig. 3.14). It can be concluded that, for all the cases, the maximum power density is approximately equal to three times and two times the average power density of the corresponding Case at BOL and EOL, respectively, and the maximum values have been observed in the seed region. It is worth mentioning that the current T-H analysis was performed without taking the T-H feedback effects on neutronics as well as a real core radial power distribution into account. Therefore 3D full core coupled neutronic T-H analysis is required in order to assess the actual performance of SB fuel. Case number Table 3.5: Summary of the selected SB configurations for the full core analysis U233 content in blanket, w/o Power density, W/cc T in, C FIR at EOL MDNBR Max. void fraction, % Max. T CL, C Max. T out, C

55 G.T. MAX G.T G.T G.T. 95 G.T G.T MIN 27 a. BOL G.T. MAX G.T G.T G.T. 95 G.T G.T MIN 27 b. EOL Fig Power density, W/cc Case 1 53

56 G.T. MAX G.T G.T G.T. 88 G.T G.T MIN 25 a. BOL G.T. MAX G.T G.T G.T. 88 G.T G.T MIN 25 b. EOL Fig Power density, W/cc Case 2 54

57 G.T. MAX G.T G.T G.T. 81 G.T G.T MIN 24 a. BOL G.T. MAX G.T G.T G.T. 81 G.T G.T MIN 24 b. EOL Fig Power density, W/cc Case 3 55

58 G.T. MAX G.T G.T G.T. 74 G.T G.T MIN 22 a. BOL G.T. MAX G.T G.T G.T. 74 G.T G.T MIN 22 b. EOL Fig Power density, W/cc Case 4 56

59 Chapter 4. Full core analysis This chapter is focused on the 3D coupled neutronic and T-H analysis of 100% Th-U233 fueled PWR core with the aim to confirm the assumptions of the assembly level calculations (Chapter 3) with regards to the practically achievable power density levels and assess the operational feasibility of the design from neutronic and T-H standpoint. 4.1 Description of the HC Th-U233 PWR core design Core loading pattern The SB fuel assemblies were arranged in a core according to the three-batch loading pattern (Fig. 4.1). Since the SB fuel assembly has very low reactivity swing (Fig. 3.3), varying the fuel assembly locations has a relatively minor effect on the core power distribution, and the maximum power peak is likely to occur in the center of the core. However, the high power peaking in the seed region can be somewhat mitigated by concentrating the twice burned fuel in the core center, since the power share between the seed and blanket for the twice burned fuel is more balanced as compared to the fresh and once burnt fuels (Fig. 3.5). On the other hand, the core patterns with the fresh fuel located in the core periphery suffer from higher neutron leakage. In the case of HC core, the sub-critical blanket of the peripheral fuel assemblies will absorb a certain amount of neutrons leaving the seed, which will in turn reduce the neutron losses T-H constraints As compared to the reference all-u fuel assembly, the SB fuel assembly has a higher flow resistance due to the enlarged radius of the blanket fuel pins (Table 3.3). Therefore, introduction of the HC SB fuel assemblies into a PWR core will increase the overall pressure drop across the active core, Δp. Preserving the reference coolant flow rate will lead to higher pumping power requirements because pumping power is proportional to Δp and the volumetric coolant flow rate,. In order to avoid uprating the primary coolant pumps, it was decided to preserve the pumping power of the HC PWR core and set it to be equal to that of the reference PWR core by reducing the total coolant mass flow rate,. 57

60 4.2 Analysis methodology Fig Core loading pattern As mentioned earlier, the 3D full core analysis was performed using the DYN3D code. Since the seed and blanket parts of the SB fuel assembly have significantly different neutronic and T-H characteristics, every SB fuel assembly in full core analysis was modeled using nine separate neutronic and T-H nodes as depicted in Fig As a result of such subdivision, some of the nodes within the fuel assembly can have rectangular rather than square shape (Fig. 4.2). Therefore, in this study, the DYN3D code was upgraded with the capability of modeling rectangular nodes. Fig Nodalization of SB fuel assembly in DYN3D code 58

61 4.2.1 Homogenized cross sections generation In this work, two-group cross sections for the four considered SB fuel assembly configurations (Table 3.5) were generated using single assembly level HELIOS calculations. In order to solve the neutron diffusion equations in DYN3D, the reactor core is divided into the elements called nodes, for each of which a set of homogenized nuclear data is needed. The homogenized few-group neutronic data including fluxweighted, node-averaged neutron diffusion coefficients and XS are generated using lattice transport codes (e.g. HELIOS) and compiled in data libraries that can be used as DYN3D input. The few-group XS were generated according to the following procedure. First, single assembly depletion calculations were performed using average core operational parameters. Subsequently, the branch-off calculations were performed for certain BU points with perturbed state parameters (i.e. coolant density (ρ M ), fuel temperature (T F ), and soluble boron concentration (C B )). The range of every state parameter can be found in Table 4.1. The XS sets including the dependence on burnup and core operational parameters were compiled in data libraries. The libraries are actually multi-dimensional tables, in which all possible combinations of parameter variation are considered. Using such a data basis, the DYN3D calculates the actual nuclear data for every node by multidimensional interpolation. The actual values of core operational parameters are calculated by DYN3D itself (see Fig. 4.3). In this study, every SB fuel assembly was subdivided into the nine nodes (as shown in Fig. 4.2) and described by three separate XS sets generated for seed, blanket-side and blanket-corner regions within the same HELIOS run. Table 4.1: Summary of the branch-off parameters Parameter # of branch-off points Value Fuel temperature, K 5 538, 900, 1300, 1500, 2500 Boron concentration, ppm 3 0, 200, 400 Moderator density, kg/m , 746, 705, 661,

62 Fig DYN3D cross section calculation scheme In order to account for the high non-uniformity of neutron flux between the seed and blanket regions and to reduce the homogenization errors, the superhomogenization (SPH) method (Hebert, 1993) was used to correct the flux-volume weighted two-group cross sections. The correction SPH factors can be obtained iteratively using the following procedure. First, the diffusion solution is obtained for the nine region single assembly problem (Fig. 4.2) with DYN3D employing twogroup cross sections generated by HELIOS. Then, the SPH factors for every region r (r=1...9) and energy group g (g=1,2) are calculated using the following relation: Het r,g r,g (4.1) where Hom r,g Het r,g and Hom r,g are the average heterogeneous and homogeneous neutron fluxes in region r and group g obtained from heterogeneous HELIOS transport solution and homogeneous DYN3D diffusion solution respectively. After that, modified cross sections, Mod r,g, are calculated for every region and energy group using the SPH factors generated according to Eq. 4.1: (4.2) Mod r,g r,g r,g Then, the nine-region single assembly diffusion problem is solved again by the DYN3D code using the modified cross sections. The obtained homogeneous neutron fluxes are used for the calculation of a new set of the SPH factors. This iterative 60