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Department of Nuclear Engineering And Radiation Health Physics
DESIGN AND ANALYSIS OF A NUCLEAR REACTOR CORE FOR INNOVATIVE SMALL LIGHT WATER REACTORS.
By Alexey I. Soldatov
A DISSERTATION
Submitted to Oregon State University
March 9, 2009
1
AN ABSTRACT OF THE DISSERTATION OF
Alexey I. Soldatov for the degree of Doctor of Philosophy in Nuclear Engineering presented on March 9, 2009. Title: Design and Analysis of a Nuclear Reactor Core for Innovative Small Light Water Reactors. Abstract approved:
Todd S. Palmer
In order to address the energy needs of developing countries and remote
communities, Oregon State University has proposed the Multi-Application Small
Light Water Reactor (MASLWR) design. In order to achieve five years of operation
without refueling, use of 8% enriched fuel is necessary.
This dissertation is focused on core design issues related with increased fuel
enrichment (8.0%) and specific MASLWR operational conditions (such as lower
operational pressure and temperature, and increased leakage due to small core).
Neutron physics calculations are performed with the commercial nuclear industry
tools CASMO-4 and SIMULATE-3, developed by Studsvik Scandpower Inc.
The first set of results are generated from infinite lattice level calculations with
CASMO-4, and focus on evaluation of the principal differences between standard
PWR fuel and MASLWR fuel. Chapter 4-1 covers aspects of fuel isotopic
composition changes with burnup, evaluation of kinetic parameters and reactivity
coefficients. Chapter 4-2 discusses gadolinium self-shielding and shadowing effects,
and subsequent impacts on power generation peaking and Reactor Control System
shadowing.
2
The second aspect of the research is dedicated to core design issues, such as
reflector design (chapter 4-3), burnable absorber distribution and programmed fuel
burnup and fuel use strategy (chapter 4-4). This section also includes discussion of the
parameters important for safety and evaluation of Reactor Control System options for
the proposed core design.
An evaluation of the sensitivity of the proposed design to uncertainty in
calculated parameters is presented in chapter 4-5.
The results presented in this dissertation cover a new area of reactor design and
operational parameters, and may be applicable to other small and large pressurized
water reactor designs.
3
© Copyright by Alexey I. Soldatov March 9, 2009
All Rights Reserved
4
Design and Analysis of a Nuclear Reactor Core
for Innovative Small Light Water Reactors.
By
Alexey I. Soldatov
A DISSERTATION
Submitted to
Oregon State University
in partial fulfillment of
the requirements for the
degree of
Doctor of Philosophy
Presented March 9, 2009
Commencement June 2009
5
Doctor of Philosophy dissertation of Alexey I. Soldatov presented on March 9, 2009.
APPROVED
Major professor, representing Nuclear Engineering
Head of the Department of Nuclear Engineering and Radiation Health Physics
Dean of the Graduate School
I understand that my dissertation will become part of the permanent collection of
Oregon State University libraries. My signature below authorizes release of my
dissertation to any reader upon request.
Alexey I. Soldatov, Author
6
ACKNOWLEDGEMENTS
I would like to take this opportunity to thank Dr. Todd S. Palmer for all of his
support, ideas and positive input through the research and writing phases of this
dissertation.
I would like to thank Dr. Kord S. Smith of Studsvik Scandpower, and
personnel of this company for guidance and support of this project.
I would like to thank the faculty members of Oregon State University Nuclear
Engineering and Radiation Health Physics department, and particularly Dr. Jose N.
Reyes Jr., Dr. Qiao Wu, Dr. Brian Woods, Dr. Steve Reese Dr. Michael Hartman and
dissertation committee members Dr. Goran N. Jovanovic and Dr. Jamie J Kruzic, who
have helped me with development of the research concepts and ideas, and writing this
dissertation.
Also I would like to acknowledge the valuable input into the MASLWR core
design from OSU undergraduate and graduate students Mark Galvin, Jeff Dahl, Jeff
Magedanz, Wade Marcum, Alex Misner, Sander Marshall and Benjamin Nelson.
Their input, comments and ideas were important for the development of the
MASLWR concept and for further feasibility studies.
I would like to thank my instructors and professors from Moscow Engineering
Physics Institute, who guide me through my education and help with information
collection and analysis for this dissertation.
Finally I would like to acknowledge my family for the inspiration and support
in my education, and particularly my father, Igor Mikhailovich Soldatov P.E., who has
encouraged me to give my heart and soul into Nuclear Engineering and Science.
7
TABLE OF CONTENTS
Page
Chapter 1 – Introduction ................................................................................................ 1
1.1. Introduction......................................................................................................... 1
Chapter 2 Literature review ........................................................................................... 3
2.1. Small reactor designs on market (Competitors for MASLWR) ......................... 3
2.2. Innovative LWR research programs and Research Areas................................. 26
2.3. National and International requirements and regulations ................................. 35
Chapter 3 – Methodology ............................................................................................ 43
3.1. Initial Data – Definition of research goals .................................................. 43
3.2. Initial Data – Geometry and materials ........................................................ 44
3.3. Tools Available for Neutronic design ......................................................... 47
3.4. Studsvik Scandpower Tools.............................................................................. 48
3.5. Physical Models implemented in the tools ....................................................... 49
3.6. Algorithms of the tools used for the standard calculations ............................... 56
3.7. Organization of the feasibility study ................................................................. 59
3.8 Goals of Feasibility Study, Criteria of completion ............................................ 69
Chapter 4 Results ......................................................................................................... 70
Chapter 4-1 – Study of the Physical Effects in MASLWR 8% Enriched Fuel (General
Analysis) ...................................................................................................................... 71
4-1.1 Calculation model ........................................................................................... 71
4-1.2 Multiplication factor ....................................................................................... 73
8
TABLE OF CONTENTS (Continued)
Page
4-1.3 Fuel Depletion, Plutonium Production..................................................... 78
4-1.4 Neutron flux and energy spectra for PWR and MASLWR .......................... 100
4-1.5 Prompt neutron lifetime and Effective fraction of delayed Neutrons........... 110
4-1.6. Reactivity Effects......................................................................................... 116
4-1.7. Control Rod Worth, Shielding and Shadowing Effects............................... 133
4-1.8 Conclusions................................................................................................... 157
Chapter 4-2 – Results Gadolinium Burnable Absorbers for MASLWR Fuel ........... 159
4-2.1 Burnable Absorbers (Functions, Design Options)........................................ 159
4-2.2. Calculation Model and Assumptions........................................................... 168
4-2.3. Fuel Assembly with Gadolinium BP (standard geometries). ..................... 170
4-2.4. Self-shielding effects in fuel with burnable absorbers................................. 175
4-2.5. Shadowing effects of Gd burnable absorbers in V and W fuel assembly
modifications.......................................................................................................... 187
4-2.6. Control rod shadowing effect ...................................................................... 202
4-2.7. Multiple burnable absorbers for programming of multiplication factor...... 209
4-2.8. Conclusions.................................................................................................. 211
Chapter 4-3 Results – Reflector Studies .................................................................... 213
4-3.1. Basic requirements for a neutron reflector .................................................. 214
4-3.2. Neutron Reflector Materials ........................................................................ 216
4-3.3. Neutron Reflector Geometry ....................................................................... 219
9
TABLE OF CONTENTS (Continued)
Page
4-3.4. Reflector Studies with SIMULATE-3 ......................................................... 222
4-3.5. Reflector Studies with CASMO-4E............................................................. 227
4-3.6 Conclusions for Reflector Studies ................................................................ 235
Chapter 4-4 Results – Core Design............................................................................ 236
4-4.1. MASLWR core model in SIMULATE-3 .................................................... 237
4-4.2. MASLWR Core without burnable poisons................................................. 240
4-4.3. MASLWR Core with radial BP profiling .................................................... 257
4-4.4. MASLWR Core with 3D BP profiling ........................................................ 276
4-4.5. Prototypical core: Thermal Hydraulic parameters....................................... 284
4-4.6. Prototypical core: Kinetic and Safety parameters........................................ 294
4-4.7. Conclusions.................................................................................................. 307
Chapter 4-5 Results – Core Design Sensitivity Study ............................................... 308
4-5.1. Effects of fuel thermal conductivity variations............................................ 309
4-5.2. Effect of coolant flow rate variations .......................................................... 312
4-5.3. Effects of fission cross section uncertainty for 8.0% fuel ........................... 316
4-5.4. Effects of gadolinium worth uncertainty at BOC-MOC.............................. 318
4-5.5. Effect of variations in residual gadolinium worth at EOC .......................... 321
10
TABLE OF CONTENTS (Continued)
Page
Chapter 5 Discussion, Conclusions, Recommendations............................................ 323
5.1. Conclusion ...................................................................................................... 323
5.2. Recommendations for a future work............................................................... 328
List of References ...................................................................................................... 331
11
LIST OF FIGURES
Figure Page
2.1.2.A. Evolution of IRIS design characteristics…………………………... 8
2.1.2.B. IRIS reactor vessel and internals…………………………………... 10
2.1.2.C. IRIS core configuration……………………………………………. 11
2.1.4.A. SMART (KAERI) reactor vessel and internals……………………. 15
2.1.4.C. SMART core geometry……………………………………………. 16
2.1.4.F. SMART development program……………………………………. 18
2.1.5. SMART (USA) reactor vessel and reactor building………………. 21
2.1.6.A. CAREM Reactor internals and core……………………………… 23
2.1.6.D. CAREM Fuel element and core design (1/3 Core)……………….. 26
3.2.1. Westinghouse 17x17 PWR fuel assembly design………………… 45
3.2.2. MASLWR reactor vessel geometry……………………………….. 45
3.5.2. CASMO-4 calculations flow diagram……………………………... 52
3.6.1. The flow path for steady-state calculations with Studsvik tools…... 56
3.6.2 Standard flow-chart for steady-state calculation with flow correction and control rod, boron concentration, power lever iterations, and reactivity coefficients calculations………………… 57
3.6.3. Flow charts for refueling optimization (left – for manual optimization, right – for automatic optimization using the X-Image code)……………………………………………………………….. 58
3.6.4. Flow chart for the comparison analysis between CASMO 4E and SIMULATE 3 results (used for uncertainty evaluation)………….. 58
3.7.1. General research organization……………………………………... 60
12
Figure
LIST OF FIGURES (Continued)
Page
3.7.2. Step 1 Details ……………………………………………………… 62
3.7.3. Step 2 Details ……………………………………………………… 63
3.7.4. Step 3 Details ……………………………………………………… 65
3.7.5. Step 4 Details ……………………………………………………… 67
4-1.2.1. Multiplication Factor for different fuel enrichment in MASLWR and PWR operation conditions…………………………………….. 73
4-1.2.2. Initial Multiplication Factor for MASLWR and PWR conditions… 75
4-1.2.3. Fuel assembly M1 - burnable absorbers map……………………… 76
4-1.2.4. Multiplication factor in fuel with burnable absorbers……………... 77
4-1.3.1.A. Contribution of fertile fission for PWR fuel………………………. 78
4-1.3.1.B. Contribution of fertile fission for MASLWR fuel ………………… 79
4-1.3.1.C. Comparison of the fertile fission contribution for MASLWR and PWR fuel burnup …………………………………………………..
79
4-1.3.2. Uranium 235 depletion nuclear reactions………………………….. 81
4-1.3.3.A. Uranium 234 depletion for MASLWR and PWR fuel…………….. 82
4-1.3.3.B. Uranium 235 depletion for MASLWR and PWR fuel…………….. 83
4-1.3.3.C. Uranium 236 depletion for MASLWR and PWR fuel…………….. 84
4-1.3.3.D. Uranium 238 depletion for MASLWR and PWR fuel…………….. 86
4-1.3.3.E. Total Uranium depletion for MASLWR and PWR fuel…………… 86
4-1.3.4. Uranium 238 reactions, plutonium production…………………….. 87
4-1.3.5. Production of fissile plutonium in the fuel……………………….. 89
13
Figure
LIST OF FIGURES (Continued)
Page
4-1.3.6. Fissile plutonium production rates at the small burnup levels…….. 90
4-1.3.7. Fissile Plutonium concentrations and fuel burnup limits………….. 91
4-1.3.8. Total plutonium productions in MASLWR and PWR fuel………... 92
4-1.3.9. Total Plutonium weight fraction for MASLWR and PWR (details). 93
4-1.3.10. Plutonium qualities in MASLWR and PWR fuel…………………. 94
4-1.3.11. Plutonium quality comparison for MASLWR and PWR………….. 95
4-1.3.12. Plutonium weight fractions for 4.5 % enriched PWR fuel………… 96
4-1.3.13. Plutonium weight fractions for 8.0 % enriched MASLWR fuel…... 97
4-1.4.1.A. Total neutron flux in PWR fuel……………………………………. 101
4-1.4.1.B. Total neutron flux in MASLWR fuel……………………………… 101
4-1.4.1.C. Comparison of total neutron flux in MASLWR and PWR………... 102
4-1.4.2.A. Comparison of Fast Neutron Flux in MASLWR and PWR……….. 103
4-1.4.2.B. Comparison of Thermal Neutron Flux in MASLWR and PWR…... 103
4-1.4.3.A. Fraction of the thermal flux in PWR fuel…………………………. 105
4-1.4.3.B. Fraction of the thermal flux in MASLWR fuel…………………… 106
4-1.4.3.C. Fraction of the thermal flux in MASLWR fuel…………………… 107
4-1.4.3.D. Fraction of the thermal flux in PWR fuel…………………………. 108
4-1.5.1.A. Prompt neutron lifetime (PWR)…………………………………… 111
4-1.5.1.B. Prompt neutron lifetime (MASLWR)……………………………... 112
4-1.5.1.C. Prompt neutron lifetime (PWR vs. MASLWR)…………………… 113
4-1.5.2.A. Effective Fraction of Delayed Neutrons (MASLWR)……………. 114
14
Figure
LIST OF FIGURES (Continued)
Page
4-1.5.2.B. Effective Fraction of Delayed Neutrons (MASLWR)…………….. 114
4-1.6.1.A. Moderator Temperature Coefficient (HFP, BOR=0.1 ppm)………. 119
4-1.6.1.B. Moderator Temperature Coefficient (HFP, BOR=800 ppm)……… 119
4-1.6.1.C. Moderator Temperature Coefficient Comparison…………………. 120
4-1.6.2.A. Fuel Temperature Coefficient (HFP, BOR=800 ppm)…………….. 122
4-1.6.2.B. Fuel Temperature Coefficient (TMO=493K, BOR=800 ppm)……. 122
4-1.6.3.A. Void Reactivity Coefficient (HFP, BOR=0.1 ppm)……………….. 124
4-1.6.3.B. Void Reactivity Coefficient (HFP, BOR=800 ppm)………………. 124
4-1.6.3.C. Void Reactivity Coefficient Comparison………………………….. 125
4-1.6.4.A. Zero Power Heating Reactivity Defect……………………………. 127
4-1.6.4.B. Power reactivity Defect During Power Increase Period …………... 129
4-1.6.4.C. Relative power reactivity defect (linear approximation)………….. 130
4-1.6.4.D. Total reactivity defect……………………………………………… 131
4-1.7.1.A. Soluble Boron Worth, cents/ppm (HFP, BOR: 0.1-800 ppm)…….. 133
4-1.7.1.B. Soluble Boron Worth, cents/ppm (HFP, BOR: 800-1600 ppm)…... 133
4-1.7.1.C. Soluble Boron Worth Comparison………………………………… 134
4-1.7.1.D. Soluble boron worth loss ratio for higher concentration…………... 135
4-1.7.2.A. Soluble Boron Worth, cents/ppm (CZP, BOR: 0.1-800 ppm)…….. 136
4-1.7.2.B. Soluble Boron Worth, cents/ppm (CZP, BOR: 800-1600 ppm)…... 137
4-1.7.2.C. Soluble boron worth loss ratio for higher concentration…………... 137
4-1.7.3.A. Soluble Boron Worth Comparison (CZP, HZP, HFP)…………….. 139
15
Figure
LIST OF FIGURES (Continued)
Page
4-1.7.3.B. Soluble Boron Worth Comparison (CZP, HZP, HFP)…………….. 140
4-1.7.4.A. Soluble Boron Worth, cents/ppm (Boiling, Void=0%)……………. 141
4-1.7.4.B. Soluble Boron Worth, cents/ppm (Boiling, Void=40%)…………... 142
4-1.7.4.C. Soluble Boron Worth Comparison at Boiling Conditions………… 142
4-1.7.4.D. Soluble boron worth loss ratio with boiling……………………….. 143
4-1.7.5.A. B4C Control Rod Worth (HFP, BOR=800 ppm)………………….. 145
4-1.7.5.B. B4C Control Rod Worth (CZP, BOR=800 ppm)………………….. 146
4-1.7.5.C. Control rod worth comparison (CZP, HZP, HFP conditions)……... 147
4-1.7.6.A. Ag-In-Cd Control Rod Worth (HFP, BOR=800 ppm)…………….. 150
4-1.7.6.B. Hafnium Control Rod Worth (HFP, BOR=800 ppm)……………... 150
4-1.7.6.C. Control Rod Worth Comparison (B4C, AIC, HAF)………………. 151
4-1.7.6.D. Control Rod Worth Comparison (B4C, AIC, HAF)………………. 152
4-1.7.7.A. B4C Control Rod Worth (HFP, Void 0 %, BOR)…………………. 154
4-1.7.7.B. B4C Control Rod Worth (HFP, Void 40 %, BOR)………………... 154
4-1.7.7.C. Control rod worth gain ration with boiling………………………... 155
4-2.1.1. The total neutron absorption cross section of boron………………. 161
4-2.1.2.A. Gadolinium Reaction tracked in CASMO-4………………………. 161
4-2.1.2.B. Total neutron absorption cross sections of gadolinium……………. 162
4-2.1.3.A. Erbium Reaction tracked in CASMO-4…………………………… 163
4-2.1.3.B. Total neutron absorption cross sections of erbium………………… 163
16
Figure
LIST OF FIGURES (Continued)
Page
4-2.1.4 Comparison Gd-155, Gd-157 and Er-167 absorption cross sections with U-235 fission cross section…………………………………... 164
4-2.3.1 The geometry specifications of some standard fuel assemblies…… 170
4-2.3.2.A. Multiplication factor in a fuel with standard arrangement of burnable poisons (8.0% enriched fuel, standard BA conc.)……….. 171
4-2.3.2.B. Multiplication factor in a fuel with standard arrangement of burnable poisons (8.0% enriched fuel, 9.0% BA conc.)…………... 172
4-2.3.2.C. Comparison of Multiplication factor in a fuel with standard arrangement of burnable poisons………………………………….. 172
4-2.4.1. Fuel Assembly geometry specifications for Gd self-shielding survey……………………………………………………………… 175
4-2.4.2.A. The results of Gadolinia Self-Shielding Study (BOR=800 ppm)….. 176
4-2.4.2.B. The results of Gadolinia Self-Shielding Study (BOR=0.1 ppm)…... 177
4-2.4.2.A. The results of Gadolinia Self-Shielding Study (BOC-MOC)……... 178
4-2.4.3.A. Atomic concentrations of Gadolinium isotopes in BA-pins in Fuel Assembly FA_8TG02……………………………………………… 179
4-2.4.3.B. Atomic concentrations of Gadolinium isotopes in BA-pins in Fuel Assembly FA_8TG04……………………………………………… 179
4-2.4.3.C. Atomic concentrations of Gadolinium isotopes in BA-pins in Fuel Assembly FA_8TG06……………………………………………… 180
4-2.4.3.D. Atomic concentrations of Gadolinium isotopes in BA-pins in Fuel Assembly FA_8TG08……………………………………………… 180
4-2.4.4.A. Linear atomic concentrations of gadolinium isotopes per unit length of fuel assembly FA_8TG02……………………………………….. 181
4-2.4.4.B. Linear atomic concentrations of gadolinium isotopes per unit length of fuel assembly FA_8TG08………………………………………... 181
17
Figure
LIST OF FIGURES (Continued)
Page
4-2.4.5.A. Linear atomic concentrations of Gd-152…………………………... 182
4-2.4.5.B. Linear atomic concentrations of Gd-155 and Gd-156……………... 183
4-2.4.5.C. Linear atomic concentrations of Gd-157 and Gd-158……………... 184
4-2.5.1. Fuel Assembly geometry specifications for Gd-shadowing survey.. 187
4-2.5.2.A. U-Type fuel assembly multiplication factor versus depletion……... 188
4-2.5.2.B. V-Type fuel assembly multiplication factor versus depletion……... 189
4-2.5.2.C. W-Type fuel assembly multiplication factor versus depletion ……. 190
4-2.5.3. U, V and W type fuel assembly multiplication factor comparison... 191
4-2.5.4. The initial multiplication factor as a function of BA concentration.. 192
4-2.5.5.A. The depletion coordinate for the maximum of multiplication factor, as a function of burnable absorber concentration…………………… 193
4-2.5.5.B. The maximum multiplication factor, as a function of BA concentration………………………………………………………. 194
4-2.5.6. Fuel assembly pin peaking factor versus depletion………………... 195
4-2.5.7. Initial fuel assembly pin peaking factor…………………………… 196
4-2.5.7.A. U-type Fuel Assembly Pin-Power Distribution …………………… 197
4-2.5.7.B. V-type Fuel Assembly Pin-Power Distribution …………………… 198
4-2.5.7.C. W-type Fuel Assembly Pin-Power Distribution …………………... 199
4-2.6.1 A. U-type fuel assembly control rod worth vs. depletion ……………. 202
4-2.6.1 B. V-type fuel assembly control rod worth vs. depletion ……………. 203
4-2.6.1 C. W-type fuel assembly control rod worth vs. depletion …………… 204
4-2.6.2 A. U-type fuel assembly soluble boron worth vs. depletion …………. 205
18
Figure
LIST OF FIGURES (Continued)
Page
4-2.6.2 B. V-type fuel assembly soluble boron worth vs. depletion …………. 206
4-2.6.2 C. W-type fuel assembly soluble boron worth vs. depletion ………… 207
4-2.7.1. Multiplication Factor as a function of exposure for fuel assemblies with 4.0% and 9.0% Gadolinium concentration rods …………….. 209
4-2.7.2. Multiplication Factor as a function of exposure for fuel assemblies with 5.0% and 9.0% Gadolinium concentration rods …………….. 209
4-3.3.1. Upper-core structures, and CASMO model for axial reflector …… 218
4-3.3.2. Reactor internals in a mid-core cross-section plane ………………. 220
4-3.3.3. Radial reflector options …………………………………………… 220
4-3.4.1. Simulate-3 segments assignment …………………………………. 221
4-3.4.2.A. Burnup distribution in the MOC (Burnup 20 MWD/kgHM) ……... 222
4-3.4.2.B. Burnup distribution in the MOC (Burnup 40 MWD/kgHM) ……... 223
4-3.4.2.C. Burnup distribution in the EOC (EOC Criterion Keff=1) ………… 223
4-3.4.2.D. Burnup distribution in the EOC (EOC Criterion burnup limit)…… 223
4-3.4.3.A. Burnup distribution at MOC………………………………………. 224
4-3.4.3.B. Burnup distribution at EOC……………………………………….. 225
4-3.5.1. CASMO-4E input geometries …………………………………….. 227
4-3.5.2. CASMO-4E Segment assignment and boundary conditions ……... 228
4-3.5.3.A. Energy deposited in full-height reflector by neutron thermalization 231
4-3.5.3.B. Energy deposited in full-height reflector by gamma attenuation….. 232
4-3.5.3.C. The contributors to total energy deposited in full-height reflector... 232
4-3.5.3.D. The contributors to total energy deposited in reflector material ….. 233
19
Figure
LIST OF FIGURES (Continued)
Page
4-4.1.1. Simulate-3 model nodalization (fuel assembly and core)…………. 236
4-4.2.1. Multiplication Factor (HFP, BOR=0 ppm) ……………………….. 240
4-4.2.2 Power Peaking Factor (HFP, BOR=0 ppm) ……………………… 241
4-4.2.3 Axial offset (HFP, BOR=0 ppm) ………………………………… 241
4-4.2.4.A. Fuel assembly names assignment ………………………………… 243
4-4.2.4.B. 3-D Relative Power Fraction (HFP, BOC, 0 MWD/kgHM) ……. 243
4-4.2.4.C. 3-D Relative Power Fraction (HFP, MOC, 25 MWD/kgHM) …… 243
4-4.2.4.D. 3-D Relative Power Fraction (HFP, EOC, 50 MWD/kgHM)…….. 244
4-4.2.5. 3-D Fuel Burnup (HFP, EOC, 50 MWD/kgHM)…………………. 244
4-4.2.6. 3-D Coolant Temperature (HFP, BOC, 0 MWD/kgHM)………… 245
4-4.2.7. 3-D Coolant Density (HFP, BOC, 0 MWD/kgHM)……………… 246
4-4.2.8. Effective fraction of delayed neutrons ……………………………. 247
4-4.2.9. Reactivity coefficients associated with moderator temperature…… 249
4-4.2.10. Uniform Doppler coefficient ……………………………………… 250
4-4.2.11. Power reactivity coefficient ……………………………………….. 251
4-4.2.12. Flow reactivity coefficient ………………………………………… 252
4-4.2.13. Pressure reactivity coefficient …………………………………….. 253
4-4.2.14. Power defect (HZP-HFP, 100% Flow) ……………………………. 254
4-4.2.15. Soluble boron worth ………………………………………………. 255
4-4.3.1 MASLWR Fuel – Radial Arrangement of Burnable Absorbers ….. 257
4-4.3.2.A. Multiplication Factor (HFP, BOR=0 ppm) ……………………….. 258
20
Figure
LIST OF FIGURES (Continued)
Page
4-4.3.2.B. Multiplication Factor as a function of depletion and soluble boron concentration……………………………………………………….. 258
4-4.3.3 Power Peaking Factor (HFP, BOR=0 ppm) ………………………. 259
4-4.3.4 Axial offset (HFP, BOR=0 ppm) ………………………………….. 260
4-4.3.5.A. Core 3-D thermal hydraulic parameters (BOC, 0 MWD/kgHM) …. 262
4-4.3.5.B. Core 3-D thermal hydraulic parameters (MOC, 13 MWD/kgHM) .. 264
4-4.3.5.C. Core 3-D thermal hydraulic parameters (MOC, 25 MWD/kgHM) .. 265
4-4.3.5.D. Core 3-D thermal hydraulic parameters (MOC, 30 MWD/kgHM) .. 266
4-4.3.5.E. Core 3-D thermal hydraulic parameters (EOC, 50 MWD/kgHM) .. 267
4-4.3.6 Reactivity coefficients associated with moderator temperature…… 269
4-4.3.7 Uniform Doppler coefficient ……………………………………… 270
4-4.3.8 Power reactivity coefficient ……………………………………….. 271
4-4.3.9 Flow reactivity coefficient ………………………………………… 271
4-4.3.10 Pressure reactivity coefficient ……………………………………... 272
4-4.3.11 Power defect (HZP-HFP, 100% Flow) ……………………………. 272
4-4.3.12 Soluble boron worth ………………………………………………. 273
4-4.4.1 MASLWR Fuel – 3-D Arrangement of Burnable Absorbers ……... 276
4-4.4.2.A. Multiplication Factor (HFP, BOR=0 ppm) ……………………….. 278
4-4.4.2.B. Multiplication Factor as a function of depletion and soluble boron concentration ……………………………………………………… 279
4-4.4.3.A. Power Peaking Factors (HFP, BOR=0 ppm) ……………………… 279
4-4.4.3.B. Power Peaking Factors comparison (HFP, BOR=0 ppm) ………… 280
21
Figure
LIST OF FIGURES (Continued)
Page
4-4.4.4 Axial offset (HFP, BOR=0 ppm)…………………………………... 281
4-4.4.5. Multiplication factor as a function of operation strategy…………... 282
4-4.5.1. The fuel assembly coolant exit temperature as a function of burnup 283
4-4.5.2.A. Core 3-D thermal hydraulic parameters (BOC, 0 MWD/kgHM)….. 284
4-4.5.2.B. Core 3-D thermal hydraulic parameters (MOC, 7 MWD/kgHM)…. 285
4-4.5.2.C. Core 3-D thermal hydraulic parameters (MOC, 25 MWD/kgHM)... 286
4-4.5.2.D. Core 3-D thermal hydraulic parameters (MOC, 45 MWD/kgHM)... 287
4-4.5.2.E. Core 3-D thermal hydraulic parameters (EOC, 50 MWD/kgHM)… 288
4-4.5.3. Coolant density at different depletion steps………………………... 289
4-4.5.4. Fuel thermal conductivity …………………………………………. 291
4-4.5.5. Fuel Temperature at different depletion steps……………………... 292
4-4.6.1. Critical soluble boron concentration ……………………………… 293
4-4.6.2. Effective fraction of delayed neutrons vs. burnup…………………. 294
4-4.6.3. Reactivity coefficients associated with moderator temperature…… 295
4-4.6.4. Uniform Doppler coefficient………………………………………. 297
4-4.6.5. Power reactivity coefficient……………………………………….. 297
4-4.6.6. Pressure reactivity coefficient……………………………………… 298
4-4.6.7. Flow reactivity coefficient…………………………………………. 299
4-4.6.8.A. Soluble boron worth……………………………………………….. 299
4-4.6.8.B. Soluble boron worth (details)……………………………………… 300
4-4.6.10. Reactor control system options…………………………………….. 302
22
Figure
LIST OF FIGURES (Continued)
Page
4-4.6.11. Control rod worth………………………………………………….. 303
4-4.6.12. Multiplication factor at different operational conditions…………... 305
4-5.1.1. Average fuel temperature in the most heated pellet as a function of thermal conductivity variation and core average burnup………….. 309
4-5.1.2. Variation in the multiplication factor with fuel thermal conductivity variation ± 15%……………………………………………………... 310
4-5.1.3. Variation in the axial offset with fuel thermal conductivity variation ± 15%……………………………………………………………….. 310
4-5.2.1. The axial location of coolant bulk boiling in the most heated channel as a function of flow rate variation and burnup…………... 312
4-5.2.2. The core multiplication factor as a function of flow rate variation and burnup…………………………………………………………. 312
4-5.2.3. The 3PIN peaking factor as a function of flow rate variations and burnup……………………………………………………………… 313
4-5.2.4.A. Radial peaking factor variations as a function of flow rate and burnup ……………………………………………………………... 314
4-5.2.4.B. Axial peaking factor variations as a function of flow rate and burnup………………………………………………………………. 314
4-5.3.1. The core average multiplication factor as a function of changes in νΣf and fuel burnup…………………………………………………. 316
4-5.3.2. The 3PIN peaking factor as a function of changes in νΣf and fuel burnup………………………………………………………………. 316
4-5.4.1. Calculation of keff correction coefficients, assuming a gadolinium worth variation 5%………………………………………………….. 317
4-5.4.2. The core average multiplication factor as a function of depletion and gadolinium worth variation…………………………………….. 318
23
Figure
LIST OF FIGURES (Continued)
Page
4-5.4.3. 3PIN Peaking factor as function of variation in gadolinium worth and fuel burnup……………………………………………………… 319
4-5.4.4. Axial offset as function of variation in gadolinium worth and fuel burnup………………………………………………………………. 319
4-5.5.1. The multiplication factor as a function of depletion and residual gadolinium worth…………………………………………………… 320
4-5.5.2. The relative difference in multiplication factor as a function of depletion and residual gadolinium worth…………………………… 321
24
LIST OF TABLES
Table Page
2.1.1. Potential Design Challenges and Options for MASLWR core………... 7
2.1.2.D. The main parameters of the IRIS Reactor…………………………….. 11
2.1.2.E. The main neutron-physics characteristics of the IRIS core…………… 12
2.1.4.B. The main parameters of the SMART plant design……………………. 16
2.1.4.D. SMART neutron-physics characteristics ……………………………… 17
2.1.4.E. Thermal hydraulic parameters of KAERI SMART Reactor ………….. 18
2.1.6.B. The main parameters of the CAREM …………………………………. 24
2.1.6.C. CAREM neutron-physics characteristics……………………………… 26
3.2.1. Main parameters of the MASLWR reactor……………………………. 44
4-1.1.1 Fuel Assembly parameters, used in research model…………………... 71
4-1.1.2 Operational parameters for MASLWR and PWR comparison ………. 72
4-1.2.1 The Multiplication factor graphs characterizing parameters …….…… 74
4-1.3.1. The isotopic composition of PWR and MASLWR spent nuclear fuel at 60 MWD/kgHM, atomic number densities (Minor Actinides)……………………………………………………… 98
4-1.3.2. The Isotopic Composition of PWR and MASLWR spent nuclear fuel at 60 MWD/kgHM, atomic number densities (Fission products and other isotopes)………………………………….. 99
4-1.4.1. MASLWR and PWR Flux comparison summary……………………... 104
4-1.6.1. Fuel segments used in a feedback evaluation study…………………… 117
4-2.1.1. Long-term reactivity excess (CASMO-4 Calculations)………….……. 158
4-2.1.2. Functions of Burnable absorbers with regards to the Reactivity control…………………………………………………………………. 159
25
Table
LIST OF TABLES (Continued)
Page
4-2.3.1. The fuel pin specifications of some standard fuel assemblies………… 169
4-2.4.1 Fuel Assembly specifications for Gd self-shielding survey…………… 174
4-2.4.5. Self-shielding survey summary ……………………………………….. 185
4-2.5.1. The specifications of fuel assemblies used in a shadowing survey …... 186
4-3.4.2. Summary of Simulate 3 reflector studies …………………………….. 222
4-3.5.2. Multiplication factor as a function of reflector and fuel enrichment….. 228
4-3.5.3. Energy deposition in full-height core reflector by gammas ………….. 229
4-3.5.4. Energy deposition in full-height core reflector by neutrons………….. 229
4-3.5.5. Total energy deposition in full-height core reflector………………….. 230
4-4.1.2 MASLWR operational conditions……………………………………. 238
4-4.6.9. Prototypical MASLWR core reactivity coefficients…………………. 301
4-5.1.1. Core cycle length variation……………………………………………. 309
26
LIST OF ACRONYMS
Reactor Types
MASLWR Modular Advanced Small Light Water Reactor
LWR Light Water Reactor
PWR Pressurized Water Reactor (Generic name for all PWRs)
BWR Boiling Water Reactor (Generic name for all BWRs)
ABWR Advanced Boiling Water Reactor (Designed by General Electric)
EPR European Pressurized Water Reactor (Designed by AREVA) Could also mean: Evolutionary Pressurized Water Reactor
IRIS International Reactor Innovative and Secure
STAR LW Safe Transportable Advanced Light Water Reactor
SMART System-integrated Modular Advanced Reactor (KAERI)
CAREM Central ARgentina de Elementos Modulares (Argentinean small modular LWR)
Systems and Components
NPP Nuclear Power Plant
FA Fuel Assembly
BA Burnable Absorber, same as BP (Burnable Poison)
BP Burnable Poison, same as BA (Burnable Absorber)
FA NBA Fuel Assembly without Burnable Absorbers
CR Control Rod
SNF Spent Nuclear Fuel (Nuclear Fuel After Irradiation)
NSSS Nuclear Steam Supply System
RCS Reactor Control System
27
LIST OF ACRONYMS (Continued)
Systems and Components
CVCS Chemical and Volume Control System
ECCS Emergency Core Cooling System
Operation and Accident modes and conditions
ATWR Anticipated Transient Without Scram
RIA Reactivity Initiated Accident
PCI Pellet Cladding Interaction
LOCA Loss of Coolant Accident
LOFA Loss of Flow Accident
BOC Beginning of Cycle
MOC Middle of cycle
EOC End of cycle
CZP Cold, zero power conditions (operation mode)
HZP Hot, zero power conditions
HFP Hot, full power conditions
RPF Relative power fraction
Organizations
OSU Oregon State University
US NRC United Stated Nuclear Regulatory Comission
US DOE United Stated Department of Energy
IAEA International Atomic Energy Agency
NEA Nuclear Energy Agency
28
LIST OF ACRONYMS (Continued)
Organizations
OECD Organization for Economic Cooperation and Development
WNA World Nuclear Association
WNU World Nuclear University
ORNL Oak Ridge National Laboratory
INL Idaho National Laboratory
AREVA AREVA – The biggest NPP vendor and nuclear fuel supplier
WHS Westinghouse Electric Co.
GE General Electric
EUR Club European Utility Requirements Club (the club which include most of the European utilities in order to develop unified requirements for LWR in Europe)
EC European Commission (European Government)
JRC Joint Research Center of European Comission
CEA Commissariat de l’Energie Atomique – French Nuclear Ministry
Documents
NPT Non-Proliferation Treaty
INFCIRC IAEA Information Circular
TOR Terms of References
TS Technical Specifications
1
DESIGN AND ANALYSIS OF A NUCLEAR REACTOR CORE FOR INNOVATIVE SMALL LIGHT WATER REACTORS.
Chapter 1 – Introduction
1.1. Introduction
The sustainable development of the world’s energy sector cannot be achieved
without extensive use of nuclear energy and the advantages of nuclear related
technologies. In order to meet national and global (international) requirements,
achieve maximum safety and efficiency levels, and address worldwide proliferation
concerns, a variety of innovative nuclear reactors will need to be developed.
Oregon State University has proposed a design of the Multi-Application Small
Light Water Reactor (MASLWR). The passive safety systems, natural circulation,
long core lifetime, off-site refueling and use of standard equipment are the basic
design features of the MASLWR reactor design.
The design challenge of MASLWR is to find a suitable and sustainable reactor
core configuration that will fit with these complex requirements. Despite the use of
standard equipment (standard fuel geometry, in-core internals), the combination of
features listed above will require fuel characteristics that are different from traditional
LWR fuel assemblies. The design of small, natural circulation LWR cores requires an
understanding of the physics, an accurate and efficient set of computational tools and a
well-developed design methodology.
The purpose of this research is to evaluate the effects of increased enrichment,
deeper burn-up rates, increased leakage, and burnable poison distribution on reactor
performance, and to design a prototypical core for use in a small light water reactor.
State of the art tools used in the modeling of traditional LWRs was evaluated
for the design of small, modular LWRs. The data generated through these simulations
cover a new range of fuel enrichments, temperatures, pressure, power, and flow rates.
2
An evaluation of the reactor control system options was also performed to meet the
stringent design criteria of this reactor.
The research work is primary focused on the increased enrichment fuel, and
analysis of phenomena related with use of this fuel at a specific operational conditions
of a MASLWR reactor. The first level of analysis is performed with a neutron
transport code CASMO-4 for a fuel assembly with reflected boundary conditions. This
analysis allowing evaluating characteristics of the MASLWR fuel and compare them
with characteristics of conventional PWR fuel. The chapter 4-1 provides the results of
effects of the increased fuel enrichment, and operational conditions onto:
1) Infinite lattice multiplication factor,
2) The fuel isotopic content,
3) Neutron energy spectrum
4) Reactor kinetic parameters
5) Reactivity feedbacks
6) Worth of the control rods and soluble boron.
The use of the increased enrichment fuel in a MASLWR core leads to the high
reserve of the reactivity for a fuel burnup. This reactivity reserve is compensated with
a gadolinium burnable absorbers. The self-shielding of the gadolinia in a MASLWR
fuel, fuel shadowing and control rod shadowing effects was evaluated for MASLWR
fuel, and discussed in a chapter 4-2.
The small reactor cores have a high neutron leakage. The evaluation of the
neutron leakage on the multiplication factor and fuel utilization and discussion of the
core reflector issues are performed in a chapter 4-3.
The neutron leakage, use of the burnable absorbers and coupled neutronic and
thermal hydraulic should be taken into account during a core design. The prototypical
MASLWR core design is evaluated in a chapter 4-4 of the dissertation.
3
Chapter 2 Literature review
2.1. Small reactor designs on market (Competitors for MASLWR)
Today, the world’s major reactor vendors have targeted markets in developed
countries and currently offer designs that have large power outputs (1000–1700
MWe). [3] However, these large reactors are unsuitable for many developing countries
for several reasons [79].
Many developing countries have limited electric grid capacity that cannot
accommodate a single power plant with output approaching or exceeding 1000 MWe.
Also, the grid in some countries is localized in a few isolated population centers with
minimal interconnections. This situation favors the use of smaller power plants sited at
geographically separated locations.
Nuclear power plants traditionally have a large capital cost relative to fossil
power plants, which creates an additional barrier to choosing nuclear power. By virtue
of their reduced size and complexity, smaller-sized nuclear plants will have a lower
capital cost per plant and shorter construction time. Thus the initial power unit can be
generating revenue before the second and third units are constructed, reducing the
maximum capital outlay for the combined generating capacity. This is especially
important for developing economies, which typically have limited availability of
capital funds.
Because of the lower power levels of small or medium-sized nuclear plants,
countries have more flexibility to install generating capacity in smaller increments that
better match their rate of power demand and economic growth. The reduced power
levels allow greater use of passive safety systems and plant simplifications, such as
natural circulation of the primary coolant. These features enhance the safety and
4
reliability of the nuclear power facility, which is especially advantageous in countries
that have limited nuclear experience and trained workforces.
Current, worldwide nuclear energy is based on medium and high capacity level
reactors [489] (IAEA classification of capacity levels 300 - 900 MWe and 900 – 1600
MWe respectively [54]). The primary technology for operational nuclear reactors
worldwide is Light Water Reactor (LWR) technology [48, 268, 489]. The world
“technology” here does not specifically refer to the nuclear power plant design, but
also the LWR-oriented fuel cycle facilities and comprehensive supply chain. LWR
technology is well studied and well reflected in regulatory practice [68-71, 288]. This
makes LWR technology most attractive for innovative reactor design in the short-term
and mid-term time frames.
Several countries are currently working on the design of innovative nuclear
power plants in the low-power range (10 – 300 MWe). There are a variety of design
concepts being considered worldwide (such as High Temperature Gas Cooled
Reactors and Liquid Metal Fast Breeder Reactors) [27, 52, 54]. However, LWRs are
considered the most economically feasible and deployable technology for innovative
reactors in the near future [21].
In order do design a competitive reactor, and successfully promote in into
international markets, the overview of a close competitors is necessary. In order to
perform this analysis, first of all necessary to recognize and list all major reactor
designs concepts (projects, designs) being proposed during the last decade, identify
major advantages and disadvantages of the proposed designs, identify designs which
has similarities with MASLWR, and try to understand the ideas and motivation behind
these designs. Learning the difficulties and unsuccessful designs of small reactors
(particularly LWRs) is also important in order to avoid the repetition of the others
mistakes and potential challenges in the early design stage. The general overview of
the competing designs is given for large scale innovative LWRS [3, 61, 62, 63, 157]
and for different designs of small reactors [21, 24, 27, 52, 53, 54, 55, 79, 85, 158]
including innovative LWR. Detailed discussions of the innovative reactor designs are
given in the sources [101-156].
5
2.1.1. MASLWR (USA)
In order to give an overview of the competing design, the presentation of the
basic concepts and ideas of Modular Advanced Small Light Water Reactor
(MASLWR) is necessary.
The MASLWR Project was inspired by the ideas of the small, modular
transportable LWR with long-lived core and off-site refueling, implemented in the
STAR-LW concept of the late 1990-s early 2000 [126]. The development of the
STAR-LW lead to the creation of IRIS design, and future evolution of IRIS leads to
the forced circulation, non-transportable reactor. In 2000 INL and OSU developed the
STAR-LW concept and proposed the MASLWR design. The scaled thermal hydraulic
test facility was commissioned in OSU to study the transportable, modular LWR with
natural circulation of coolant. Analysis of natural circulation and safety systems
thermal hydraulics were performed at this facility [101, 102, 103]. The synthesis of
this analysis extended with plant layout and main systems stretches and associated
studies performed during the period 2000-2004 are presented in the NERI Report
[104]. The research of natural circulation continues with flow stability tests [106] and
modeling of passive safety systems [105].
The results of the OSU work created an interest to the project from the
commercial entities [495]. Commercialization of the MASLWR technology [109,
110], requires design and analysis of the prototypical core and associated reactor
control system for the demonstration of the feasibility of the declared design
parameters and further correction of the reactor design. Reactor core parameters were
analyzed with MCNP (initial core studies) and CASMO-3 (initial fuel assembly level
studies) [107, 108]. These calculations allow identify major tasks for neutronic
calculations, however it was clear that for further design, mode detailed and accurate
calculations with utilization of the state-of-art tools necessary for modeling of
burnable absorbers depletion and reactor control system with feedbacks.
6
The design objective of the MASLWR project were to provide a nuclear
energy option for small energy generation markets [104, 107, 109]. The market of
“small generators” refers to small communities (small networks), remote locations
(away from fuel supply and energy supply networks), and regions with high potential
risks of natural disasters (flooding, earthquakes, tornados, etc.) – in other words,
regions where the construction of large nuclear power plants (NPPs) is impossible
[109, 110]. The MASLWR concept is also oriented toward industrial users of energy
and low-potential heat [110]. The customization of MASWLR for foreign markets
should also address Nuclear Non-Proliferation issues [107].
MASLWR must incorporate the following design requirements:
• Operational flexibility (maneuverability) of the MASLWR-based NPP
to accommodate load following operation.
• Enhanced safety of the NPP in a wide range of operational conditions
and a robust response to internal and external phenomena
• Low construction, decommissioning, O&M, and fuel costs.
• Long core lifetime
• Design consideration of end-of-life issues, such as decommissioning
and spent nuclear fuel handling
These key features of the MASLWR design lead to the design requirements:
• Passive safety systems
• Defense in depth philosophy
• Negative reactivity feedbacks, especially at full power
• Natural circulation reactor cooling
• Transportable reactor unit modules
• Short construction schedule (~1.5 years for first unit)
• A five year core lifetime, without refueling
• Components manufactured at conventional steel works, leading to
simplified manufacturing compared to traditional PWR heavy
components
7
• Lower temperature and pressure parameters
• Standard fuel assembly design
• Implementation of an automatic plant operational system for balance of
plant-equipment and safety systems
The MASLWR NERI Report [104] does not refer to any particular fuel
enrichment and gave the range of the potential fuel compositions that could be utilized
in the reactor (covering range from 4% up to 19.5% of enrichment, and assuming that
MOX and Thorium fuel could be used). The discussions and estimations of the
potential design options in 2006 resulted in two basic options for enrichment: 4.95%
and 8.0%. Additionally consideration of the potential markets for MASLWR resulted
two approaches: On-site refueling option for the domestic market, and transportable
option with off-site refueling for the foreign customers. The matrix of the potential
options and associated design challenges is given in the Table 2.1.1.
Table 2.1.1. Potential Design Challenges and Options for MASLWR core
On-site refueling (Partial core)
On-site refueling (Full core)
Off-site refueling (Full core)
4.95
% • Fuel Use optimization
• Refueling Strategy,
• Fuel Use optimization
• Transportation with fresh and spent fuel. (prove subcritical)
8.00
%
• Fuel Use optimization • Refueling Strategy, • New Fuel/cladding
materials necessary for burnup above 60
MWD/kgHM
• Prove 5 year core • Great absorption. • Spectrum shifting • Max burn-up limit
(60 MWD/kgHM)
• Prove 5 year core • Great absorption. • Spectrum shifting • Max burn-up limit
(60 MWD/kgHM) • Transportation with
fresh and spent fuel. (prove subcritical)
8
2.1.2. IRIS (USA, Italy, International Consortium)
The IRIS (International Reactor Innovative and Secure) originated from
STAR-LW reactor concept developed under the US Department of Energy Nuclear
Energy Research Initiative (NERI) program [126]. Evolution of the IRIS is presented
in the figure 2.1.2.A. The evolution of the IRIS design and rationale for modifications
are discussed in a progress reports [116, 117, 118, 119, 123].
Figure 2.1.2.A.
The basic evaluation of the STAR LW – IRIS could be characterized in
following logical steps:
• Elimination of the transportability (STAR – IRIS)
• Change from Natural to Forced Circulation
• Decrease of the fuel enrichment down to 4.95%
• Use of the soluble boron for the long-term reactivity compensation
The basic considerations, which make IRIS deployable, licensable and
competitive reactor design are:
9
• IRIS is based on proven LWR technology, newly engineered
• IRIS is specifically designed to meet or be significantly within current
licensing regulations
• If current licensing requirements are relaxed, IRIS will still meet top-
level safety goals
• Enhanced safety through safety-by-design and simplicity – probabilistic
risk assessment (PRA) to guide final design and safety analysis
• Westinghouse will obtain early NRC input on testing and licensing
issues
• Westinghouse will establish a continuing interaction with and feedback
from NRC and ACRS as design progresses
The fuel assemblies in the IRIS core are similar to those of a loop type
Westinghouse PWR design, specifically, Westinghouse 17x17 XL Robust Fuel
Assembly and AP1000 fuel assembly designs. An IRIS fuel assembly consists of 264
fuel rods with a standard 0.374” OD in a 17x17 square array. The central position is
reserved for in-core instrumentation, and 24 positions have guide thimbles for the
control rods. The core configuration consists of 89 fuel assemblies. This configuration
has a relatively high fill-factor (i.e., it closely approximates a cylinder), to minimize
the vessel diameter (see Figure 6). The IRIS 1000 MWt core has a low power density;
the active fuel height is 14 ft. (4.267m) and the resulting average linear power density
is about 75 percent of the AP600 value. The improved thermal margin provides
increased operational flexibility, while enabling longer fuel cycles and increased
overall plant capacity factors [21, 123].
The reactor vessel, and arrangement of the reactor internals for the integral
reactor design are given in the figure 2.1.2.B.
10
Figure 2.1.2.B. IRIS Reactor vessel and internals
Reactivity control is accomplished through Integral Fuel Burnable Absorbers
(IFBA), control rods, and the use of a limited amount of soluble boron in the reactor
coolant. The reduced use of soluble boron ensure the moderator temperature
coefficient negative, thus increasing inherent safety, and lessening boric acid induced
corrosion concerns. In addition to using IFBAs, erbium in form of Er2O3 mixed in the
fuel is another standard Westinghouse integral burnable absorber.
11
Figure 2.1.2.C. – IRIS Core Configuration
The core cross-section and location of the control rods is shown at the figure
2.1.2.C. The main parameters of the IRIS reactor are given in Table 2.1.2.D [27].
.
Table 2.1.2.D. The main parameters of the IRIS Reactor.
12
Table 2.1.2.E. The main neutron-Physics characteristics of the IRIS core:
Several options for the reactor core are still under consideration [122,124].
Some publications also discuss the necessity of fuel modifications to accommodate
higher fuel enrichment and long core lives. The main focus of the IRIS design group is
13
thermal-hydraulics [123,125], reactor safety studies and the design of reactor internals
[127]. The neutronic models currently used in the reactor design are simplified, and
based on point reactor kinetics [120,121] coupled with thermal hydraulics through
correlations.
2.1.3. STAR LW (USA)
In the late 1990-s concept of the small, transportable nuclear reactors, that
could be used as a “nuclear batteries” was created [24].
The STAR (Safe, Transportable Advanced Reactor) reactor and fuel cycle
concept is devised to attain Gen-IV goals by responding to foresee mid century needs
and market conditions. It is targeted to fill energy and potable water needs for urban
centers in developing countries and is designed to fit within a hierarchical hub-spoke
energy architecture based on regional fuel cycle centers, using nuclear fuel as the long
distance energy carrier – with distributed electricity generation as the local carrier to
mesh with existing urban energy distribution infrastructures using grid delivery of
electricity, potable water, and communications (and sewage return) through a common
grid of easements.
STAR is also intended for independent power producers in industrialized
countries seeking to service emerging markets for hydrogen and water production.
STAR concept development is being conducted for a portfolio of specific reactor and
balance of plant designs to enable an incremental market penetration that is time-
phased according to the degree of R&D required.
The STAR family has two major design STAR-LM (Liquid Metal) and STAR-
LW (Light Water). The light water reactor design later evolve to IRIS and similar
Ideas were implemented in MASLWR.
14
2.1.4. SMART (KAERI)
System-integrated Modular Advanced Reactor (SMART), developed by
KAERI is forced circulation modular PWR [24, 27, 156]. Another reactor design,
Small Modular Advanced Reactor Technology [155], which is also refer to SMART
abbreviation, is a US designed modular BWR with natural circulation. The US
SMART design will be discussed in the section of this chapter.
Since 1997, KAERI has been developing the system-integrated modular
advanced reactor (SMART). The SMART is a promising, advanced integral PWR-
type reactor with a rated thermal power of 330 MW. All major primary components,
such as the reactor core, steam generator (SG), main coolant pump (MCP) and
pressurizer (PZR), are installed in a single reactor vessel assembly (RVA). The
conceptual and basic designs of SMART and a coupled desalination system were
completed in March of 1999 and March of 2002, respectively. SMART development
has been conducted under the nuclear research and development program supported by
the Ministry of Science and Technology (MOST) of the Republic of Korea and thus
KAERI and MOST are the principal stakeholders.
The SMART design is an advanced reactor for dual purposes. It can be used to
supply electricity and fresh water to isolated areas where the main grid is not
interconnected. The SMART has a daily load following capacity can be finely
controlled by combining with the amount of seawater desalination. Safety functions
are performed by passive safety systems, so safety and radiation protection could be
achieved without offsite power.
Figure 2.1.4.A. shows the structural configuration of the SMART reactor. Four
main coolant pumps are installed vertically at the top of the reactor pressure vessel
(RPV). The reactor coolant flows upward through the core and enters into the shell
side of the steam generator (SG) from the top. The SGs are located at the
circumferential periphery between the core support barrel and the RPV above the core.
This design excludes the possibility of the large-break loss of coolant accident
(LBLOCA) by the elimination of coolant loops. Additional innovations include the
15
canned motor pumps, which remove the necessity of pump seals and the possibility of
the small-break LOCA (SBLOCA) associated with pump seal failure, and the passive
pressurizer that does not have an active spray and heater. This pressurizer design
eliminates complicated control and maintenance requirements and reduces the
possibility of malfunction [27].
Figure 2.1.4.A. SMART (KAERI) Reactor vessel and internals
The SMART design adopts a three-year refueling cycle, and soluble boron-free
operation. These two design features can reduce the amount of liquid waste
dramatically compared with a conventional PWR.
16
Table 2.1.4.B. The Main Parameters of the SMART plant design
Reactor thermal output: 330 MW(th) Power plant output: Electricity 90 MW(e);
40,000 tons of fresh water /day Availability factor: more than 90% Fuel material / Enrichment Sintered UO2, 4.95 wt % U-235 Fuel Assembly / Core Load 57 square FA, 17×17 WHS Design Type of coolant / moderator Light water Core characteristics Soluble boron free, Low power density Maximum peaking factor (HFP) 3.29Core dimension: Active core height 2.0 m
Equivalent core diameter 1.832 m Reactor vessel: Cylindrical shell inner diameter 4072 mm
Wall thickness of cylindrical shell 264 mm
The SMART core consists of fifty-seven fuel assemblies based on the 17×17
KOFA designed by KAERI/Siemens-KWU and used in the 900 MW(e)
Westinghouse-type Korean PWRs. The active fuel height of the SMART is 200 cm.
using a 4.95 weight % enrichment of U-235; the core can be operated for three years
without refueling. The core design is characterized by a long cycle operation with a
single or modified single batch reload scheme, low core power density, soluble boron-
free operation, enhanced safety with a large negative moderator temperature
coefficient (MTC) at any time during the fuel cycle, a large thermal margin,
inherently-free from xenon oscillation instability and minimum rod motion for the
load follow with coolant temperature controlled load following.
Figure 2.1.4.C. The SMART core geometry
17
Table 2.1.4.D. SMART neutron-physics characteristics
Reactivity Effects/Deffects Excess reactivity at cold (20°C) condition, BOC 14.8 % ∆ρReactivity defects, Xenon worth 1.9 % ∆ρPower defect (HFP to HZP1 1.4 % ∆ρTemperature defect (HZP to CZP) 8.1 % ∆ρReactivity coefficients at HFP: Moderator temperature coefficient (MTC) -72 < MTC < -42 pcm/°CFuel temperature coefficient 4.52 < FTC < -2.54 pcm/°CReactivity control mechanism: Type of control rod drive mechanism (CRDM): Linear pulse motorNumber of CRDMs: 49Absorber rods per control assembly 24Absorber material Ag-In-CdControl Rod Worth: HFP Scram rod worth (from critical) 30 %∆ρCZP2 total bank worth 25 %∆ρCZP bank worth at stuck rod condition 20 %∆ρBurnable absorber material Al2O3-B4C and Gd2O3-UO2
Soluble neutron absorber Number of independent active reactor controls: Power maneuvering: CRDM + negative MTCEmergency shutdown CRDM + emergency boron injectionFuel campaign parameters Cycle length: 990 Effective full power daysAverage discharge burn-up (nominal) 26.2 MWD/kgUMaximum discharge burn-up (nominal) 31.0 MWD/kgUFuel inventory 12.47 metric ton UAnnual consumption of uranium 13,9303 kg/GW(th) yearAnnual consumption of natural uranium 143,8504 kg/GW(th) year
The data presented in tables 2.1.4.B – E. will be important for the comparison
to the MASLWR neutronic design parameters with another designs. The core design
and reactor control system design of KAERI SMART has a lot of similarities with
IRIS reactor, however the major differences between SMART and MASLWR are:
• Higher thermal parameters and forced circulation
• Bigger core, with more fuel assemblies
• Smaller Power Density
18
Table 2.1.4.E. Thermal hydraulic parameters of KAERI SMART reactor
Thermal-hydraulic characteristics: Circulation type Forced circulationNumber of coolant loops Integral typeReactor operating pressure 15 MPaCoolant inlet temperature, at RPV inlet 270°CCoolant outlet temperature, at RPV outlet 310°CMean temperature rise across core 40°CPrimary circuit volume, including pressurizer 56.27 m3
Core average heat flux 402 kW/m2Core average linear heat generation rate 12.0 kW/mSteam flow rate at normal conditions 152.5 kg/sFeedwater flow rate at nominal conditions 152.5 kg/sSteam temperature/pressure ≥ 274/3.0°C / MPaFeedwater temperature/pressure 180.0/5.2°C / MPaMaximum peaking factor (HFP) 3.29The design limit DNBR 1.41Minimum operating thermal margin 15%Design basis lifetime for vessel and structures 60 yearsDesign basis lifetime for steam generators 15 years
The development and implementation of SMART Design panned to be done in
a 3 major steps, presented in the figure 2.1.4.F.
Figure 2.1.4.F. SMART development program
19
The design approach, design and operating experience of SMART reactor, as
well as the marketing experience, and evolution of the design driving by these factors
should be considered and evaluated with application to MALWR design.
2.1.5. SMART (USA)
Another design under SMART abbreviation is a US DOE design, presented in
2003 [155].
The US SMART concept utilizes an innovative BWR design, that reduces the
overall system complexity and eliminates the need for steam generators, thereby
reducing the overall cost, which should offset the additional cost for construction of a
stronger containment. The US-SMART core design is based on a concept that uses
either low-enrichment uranium or a mixture of low-enrichment uranium and thorium.
Both fuels are relatively proliferation-resistant, and in conjunction with advanced fuel
pin and core materials, the current design would allow continued operation (with
provisions for on-line maintenance) for periods exceeding 10 years, without refueling.
The US SMART core design is a long, cylindrical core, using standard 8x8-
assembly BWR design. Two cores were studied: 244 cm diameter and 183 cm
diameter. The design was chosen to keep the “local” power density at a desired level,
that of typical BWRs (56 kW/litre).
The US SMART reactor control concept is derived from the INEEL’s
Advanced Test Reactor (ATR), where local fluxes and power densities can be much
higher than the average core flux or power density. For this concept, an increase in
local power density will reduce the effective full power days of operation, but will also
allow for better sectional burnup of the fuel due to the shorter reflector height needed.
The base design control element is a drum surrounding the periphery of the
core. A second design was also studied to verify the affect of having the control drum
as close to every outside element as possible. The thickness of the drum was varied to
check the reactivity worth. Both Cadmium and B4C were used and compared. A
20
central “safety” rod is also present to allow safe shutdown in the case of control drum
failure, which also used Cadmium or B4C. The drum can be raised and lowered as
needed to allow for “local” power generation. The idea of local power generation can
be accomplished by either using a reflective material (e.g., Be) at the bottom of the
drum, or by simply allowing water to replace the absorber material. Both a reflective
material (Be) and non-reflective design alternatives are used to determine the affect on
reactivity.
The number of fuel assemblies for the larger-size reactor core, was selected in
such a way as to allow the local power density (i.e., the density in the reflector region)
to be limited to 56 kW/l, and yet have a reflector height that was small enough to
allow for localized power generation (i.e., this is achievable, because in the SMART
concept, only one axial section of the core is burning at-a-time in an upside candle-like
fashion). Higher power densities were used for the smaller core in order to keep the
non-absorber height to approximately 54.6 cm while operating at steady-state power
of 150 MWt. The density at this reflector height is 50% higher than the typical density,
and calculations were performed to adjust the height for a 20% and 30% increase in
power density. The reflector heights necessary to achieve these power densities have
been calculated. The burnup with reactivity calculations were performed using the
MOCUP (MCNP-ORIGEN2.1)
Coupled Utility Program code to analyze the reactivity characteristics and
isotopic concentrations of unit fuel pins/cells, with 17 actinides and 41 fission
products being tracked through the MCNP portion of the analysis. In order to increase
the power level of the reactor, a change in the height of the reflector/absorber drum is
necessary, or the power density needs to increase for similar drum heights.
21
Figure 2.1.5. SMART (USA) reactor vessel and reactor building Reactor Building Reactor Vessel
22
lower power density and bigger (longer core) where natural
circula
The design solutions and some ideas implemented in the US-SMART reactor
concept is similar to early MASLWR concepts and ideas. It is not surprising,
considering that INL/INEEL was involved in both designs. Particularly the use of the
natural circulation and reflector-absorber drums correlate with some considerations of
the MASLWR report [104]. The differences are: boiling water reactor concept (no
steam generators),
tion flow is improved by boiling in the upper core.
2.1.6. CAREM (Argentina)
Another small modular reactor design is the CAREM (Central ARgentina de
Elementos Modulares) design from Argentina. The information on the CAREM design
was ac
AREM is an Argentine project to develop, design and construct an
innovative, simple and small nuclear power plant (NPP). This plant has an indirect
cycle reactor wi d contribute to
a high safety level. Some of the high-level d ant are: an
integrated primary cooling system; self-pressu system and safety systems
relying on passive features. CAREM is a ional de Energía
Atómica) project, which has been jointly developed with INVAP, an Argentine
company [21, 27].
The CAREM concept was first presen arch 1984, in Lima, Peru, during
the IAEA’s conference on small and medium cally CAREM
was one of the first of the present new generation of reactor designs [21]. The first step
of this
rospective [27].
cumulated and summarized from IAEA Reports [21, 24, 27] and other
publications [79, 137, 157, 163].
C
th distinctive features that greatly simplify the design an
esign characteristics of the pl
rized primary
CNEA (Comisión Nac
ted in M
sized reactors. Chronologi
project is the construction of the prototype of about 27 MW(e) (CAREM-25).
This project allows Argentina to sustain activities in nuclear power plant design,
assuring the availability of updated technology in the mid-term p
23
The CAREM design is intended for the base load operation, however some
modes allow to perform load following. The availability factor of CAREM is 90% or
greater.
Figure 2.1.6.A. CAREM Reactor internals and core
1 Control rod drive
2 Control rod drive structure
Feed water inlet
t ha
as nuclear desalination, where it co l
energy supply for a Reverse-Osmosis
Integral primary system react but
a new approach
final step in the R&D and system ed
unless other strategies are possible mic
reasons. R&D costs for safety accep ed. A
be co e cheaper strategy in the
next step o of a prototype of
3 Water level am generator 4 ste
5 SG6 SG Steam outlet 7 CR Absorbing element 8 Fuel Element 9 Core 10 Core Support Stucture
The CAREM power plan s a potential use for non-electric applications such
uld be used either as a heat source or electrica
-based desalination plant.
ors used classical PWR or BWR technologies
this configuration is that needs demonstration. Demonstration is the
verification strategy and it should be perform
and convenient due to, for example, econo
tance of different options should be compar
CAREM reactor prototype will nstructed because it is th
context of CAREM. The f this project is the construction
27 MW(e) (CAREM-25).
24
Table 2.1.6.B. The main parameters of the CAREM
Reactor Vessel Design Basis 40 Years Fuel type / Enrichment PWR Hexagonal, 3.5 wt% U-235 Coolant / Moderator Light Water Core design The core of CAREM-300: The core of CAREM-25:
199 fuel assemblies / 2.85 m active length. 61 fuel assemblies / 1.40 m active length
Fuel Assembly Design: Fuel assemblies of hexagonal cross section. Number Fuel Pins: 108 Number of guide tubes: 18 control Rods and 1 instrumentation thimble. Fuel Pins outer diameter 9 mm Reactor Vessel (CAREM-25) Vessel material / Lining material: SA508 Grade 3 Class 1 / SS-304L Height / Inner diameter / Wall thickness: 11 m / 3.16 m / 0.135 m Operation Parameters (CAREM – 25) Primary System Configuration Integrated Circulation Type Natural circulation for normal operation as well
as hot shutdown for low power modules (be150 MWe).
low
natural circulation for hot shutdown for high power modules (over 150 MW(e)).
Forced circulation for full power operation and
Coolan
ds to
t conditions Self-pressurization of the primary system in the steam dome is the result of the liquid-vapour equilibrium. Due to self-pressurization, bulk temperature at core outlet corresponsaturation temperature at primary pressure.
Core inlet / outlet temperature 284°C / 326°C Primary pressure / mass flow 12.25 MPa. / 410 kg/s Steam P, T / mass flow 290°C / 4.7 MPa / 175.32 t/hr Max. fuel centerline temperature: 950°C. MDNBR > 1.7 Maximum Peaking Factor 2.7
The CAREM design utilizes several reactivity control mechanism based on
different principles and are physically independent.
Long-term core reactivity reserves is controlled by the use of Gd2O3 as
burnable poison in specific fuel rods and movable absorbing elements belonging to the
25
adjust
t-term reactivity corrections (adjustments) and emergency
scram f
commonly used Ag-In-Cd alloy. Absorbing elements are used for reactivity control
during normal operation (adjust and control system) and to produce a sudden
interruption o stem).
is diversified to
: this n
CAREM-25 l
6880 pcm
y for the fast shutdown system and
u designed to
obtain a minimal dropping time so that it takes only a few seconds to completely insert
them inside the core. In CAREM-25 design,
pcm, with one rod unavailable [21, 27].
The second shutdown system (SSS) is a gravity-driven injection device of
borated water at high pressure. In CAREM-25, this system provides a total negative
reactivity at cold shutdown of 5980 pcm, assuming single failure.
he fuel cycle can be tailored to customer requirements, with a reference
design of 330 full-power days. In CAREM-300, 1/3 of the core is refuelled and 1/2 is
refuelled in CAREM-25.
n
a 3.5%
and control system. Liquid chemical compounds are not used for reactivity
control during normal operation.
Transients or shor
unctions are performed by control rods. Each absorbing element consists of a
cluster of rods linked by a structural element (namely, “spider”), so the cluster moves
as a single unit. Absorber rods fit into the guide tubes. The absorber material is the
f the nuclear chain reaction when required (fast shutdown sy
The shutdown system fulfill Argentine regulatory-body
requirements [27].
The first shutdown system (FSS) system consists of gravity drive
neutron-absorbing elements. In the design, this system provides a tota
negative reactivity at cold shutdown of , with all rods inserted.
Many of the elements are intended onl
during normal operation they are kept in the pper position. They are
this system has a minimum worth of 3500
T
CAREM-300 has about 200 tones of natural U (feed)/GW(e) per year based o
initial enrichment, 35,000 MW·d/Mt U of average discharge burn-up, 33%
thermodynamic efficiency, 0.25% of enrichment tail and 90% of load factor [21, 27].
26
.0 mm / 7.6 mm Table 2.1.6.C. CAREM neutron-physics characteristics
Fuel pellet length / Diameter 8Fuel Density 93 – 95 % TD Fuel Temperature reactivity coefficient < - 2.1 pcm/°C Coolant Temperature reactivity coefficient < - 40 pcm/°C (N
< - 4 pcm/°C (Cold shutdown) ormal operation)
Coolant Vo r /% (Normal operation) < - 43 pcm/% (Cold shutdown)
id eactivity coefficient < - 147 pcm
Burn-up rea ti 3600 pcm c vity swing Maximum aPe king Factor 2.7 Maximum / Average burn-up (CAREM-300): 45.0 / 35.0 MWD/kgU
Fi rgu e 2.1.6.D. CAREM Fuel element and core design (1/3 Core)
f the
design
2.2. Innovative LWR research programs and Research Areas
In addition to the new reactor designs, a significant amount of the research
work is focused on the design and improvement of a specific reactor and fuel
components and elements of LWR Technology. The purposes of such studies are
always involve of performance improvement, increasing of safety or resolution o
issues and discrepancies of the past design.
In order to perform a feasibility study for the MASLWR core design, the
following fields of studies was considered in the literature review:
27
ameters
• Spent fuel handling, storage and transportation
• Coupled Codes and Coupled Neutronic-Thermal hydraulic calculations
• Fuel improvement and behavior of LWR fuel
• Increased burn-up of LWR fuel
• Increased performance and reliability of LWR fuel
• New materials for LWR fuel
• Materials and design of the CR, burnable poisons (integrated,
removable) and relative issues.
• Studies on optimization of the fuel use
• Benchmark of the tools, libraries and new area of par
• Natural Circulation Systems [11,29]
2.2.1. LWR fuel behavior
Vendors, national labs, international research consortiums, universities and
other research organizations constantly perform the studies and publications on the
fuel im
ts [248, 249,254, 258, 271, 272], NRC NUREG
Publication [3
The m r studies is evaluation of the basic
phenomena
migration and
and potential et cladding interaction, fuel behavior in a
different a d
parameters of
generation pro rn-up.
provement and fuel behavior. The knowledge base op the fuel behavior and
fuel failure criteria is presented in the IAEA Publications [2,7,10,14,19,20,25], NEA
Technical and research repor
s 43, 360, 361] and other publications [76,201,211,476, 477, 483].
ain focus of the fuel behavio
in the irradiated fuel (i.e. fuel radiation growth, swelling, fission products
fission gas pressure inside fuel pin, thermal extension and oxidation)
fuel failure modes (i.e. pell
cci ental modes). This documents are also contains an important thermal
the fuel such as thermal-conductivity, thermal capacity and heat
file as a function of the power level and bu
28
temperature,
greater temperature gradients axial and radial for the fuel assembly
These documents are important for the MALWR fuel design and analysis,
because the “standard” fuel technology, geometry and materials of the MASLWR fuel
should work in a “non-standard” conditions of lower coolant and fuel
2.2.2. Increase burnup of LWR Fuel and Burnup effecrts
In order to provide 5 years of operation with off-site refueling, MASLWR fuel
should be exposed to the 50 MWD/kgU average burn-up, and fit current LWR design
burn-up limits of 60 MWD/kgU (fuel Assembly Average).
could be spitted
into sev
a depletion range of 60 – 90 MWD/kgU, with
mation on this research area are presented in the IAEA Publications
[3, 13, 15, 60], NEA Technical Reports [254, 255, 258, 267, 271] NRC NUREG
Series [382,384, 385, 386, 397] Other Publications [203, 204, 464, 210].
These limits are settled for standard LWR (PWR and BWR) fuel, however the
knowledge base of the fuel behavior around and above 60 MWD/kgU has very small
amount of experimental data, comparing to the knowledge base in the range of 30-40
MWD/kgU.
The studies of the increased burn-up is an actual an on-going process for all
nuclear energy industry, utilities and vendors. Currently these studies
eral main directions:
• The studies, modeling and experiments for highly burnt the LWR fuel
with depletion < 60 MWD/kgU
• Experimental and Theoretical Studies of the fuel pins, pellets and fuel
assembly behavior in
standard enrichment (<5%)
• Feasibility studies and fuel behavior prediction with enrichment of 5-
9% and burn-up levels around 90-100 MWD/kgU, structural behavior
of the fuel pellet and cladding materials.
The infor
29
hese publications are mostly focused on the structural and mechanical
behavior of the irradiated fuel (pellet cracking, fission product release, cladding
oxidation and thickness) and isotopic composition of the spent fuel. The core design
for the highly burnt fuel rarely considered [464, 477], however even in those papers
yed neutrons,
reactiv
igh bur-up levels. The sources [15,
19] are
e fuel
parame
matized evaluation of the control
rod his ry, burnable absorber history, operational history impact on the multiplication
is information will be useful for the
s in the MASLWR fuel over the complete
operati
T
information on the fuel kinetic parameters (effective fraction of the dela
ity coefficients) is given partially.
For the MASLWR core design mentioned documents would be useful
primarily as a source of the knowledge of the mechanical data of the highly burnt fuel
and source of the potential design concerns at the h
also given some ideas and figures related with limitation on transients and low
temperature load of the fuel. Other sources will be used as a cross reference cases at
for the comparison of the calculated results with a measured figures.
Another big area of study is on the influence of the burn-up history on th
ters, such as use of control rods, burnable poison and other factors on the
criticality properties of the burnt fuel. Oak Ridge National Laboratory has provided a
number of NUREG Publications [376-380, 388, 390-393, 396, 410] and ORNL
Publications [418-456] on the topic.
These publications mainly discussing burn-up credit approach to the criticality
safety evaluation, and at the same time gives a syste
to
factor of the fuel and isotopic composition. Th
evaluation of the control rod and other effect
on cycle.
30
2.2.3. Increased performance and reliability of LWR fuel
In order to increase performance reliability of the LWR Fuel vendors
constantly performing studies on improvement of the mechanical characteristics of the
fuel, as well as modification and optimization of the fuel composition and geometry in
order to
NBR margins trough avoiding sub-cooled boiling.
emblies was learned from [483,485]. Evolution of Westinghouse
fuel ass
signs and approaches were considered.
improve neutronic physics characteristics.
The main areas for the mechanical improvement of the fuel is modification of
the fuel assembly design, in order to provide better debris protection, maintain
guarantee fuel assembly geometry, decrease hydraulic friction and improve mixing of
the coolant and so increase D
In this literature review, the basic modification of the western LWR are
considered in [343] and compared with VVER fuel assembly design in [19]. Evolution
of the VVER fuel ass
emblies was learned from the publications [174-179,] for PWR fuel assembly
and [172, 173, 184, 185] for other reactor types. Also Areva [138] and Mitsubishi
[133] fuel de
2.2.4. New materials for LWR fuel
Analysis of new materials for the LWR fuel are divided into four main groups:
• New Fuel Pellet composition (or manufacturing technology)
• New Clading and Structural Materials Composition
• New Control rod absorbing Materials
• New Burnable Absorber Materials
Current fuel pellets technology – is centered ceramic fuel pellet with real
density equal to 0.95 – 0.98 of theoretical density. The discussion of use UO2 fuel with
increased density is given in the paper [55, 22, 209]. Also several publications are
available on use of the micro spheres [225] or dispersed fuel for LWR reactors [18, 55,
31
mented in a future.
ssia [483] and use
of dupl
release current burn-up limitation from 60 MWD/kgU up to
80-100 MWD/kgU.
Also use of new cladding techniques could decrease cladding failure (leakage)
e transportable version of
MASLWR.
194]. Currently these technologies are not considered for the MASLWR fuel, however
they could be imple
Cladding and structural materials – is another field of research. All vendors are
working on the improvement of the cladding reliability and extension of the cladding
lifetime. Some examples are the use of the annotated cladding in Ru
ex ZIRLO [182] cladding technology by Westinghouse were considered during
the literature review stage.
New cladding could increase the burn-up limits, so it could open new horizons
for the core design, and
probabilities, which could have a positive effect for th
2.2.5. Materials and design of the CR, burnable poisons (integrated, removable) and relative issues.
Traditionally 3 basic composition of the control rods were used in the NPP
with LWR: Hafnium, Boron Carbide (B4C) and Silver-Indium-Cadmium (AIC)
compos
he motivation for the use of the new materials in one
case is incr s tion if
the control l
ATWS and CR staking. The other publications on the CR phenomena [256]
In
now available nuclear industry: Er2O3, ZrB,
Borated G s
concentration w in the
literature [2 ]
itions. However, now several new compositions suggested, such as enriched
boron and Dysprosium [5, 483]. T
ea e of the control rod efficiency (worth), in the other case minimiza
pe let swelling and radiation grow effects, and improvement resistance to
addition to traditional B4C and Gd2O3, new burnable poisons materials
and tested for the use in the commercial
las , and compositions with Yt, Eu. [5, 477]. Also use of the high
of gadolinium BP (with a self shielding) is widely discussed no
56 .
32
he manufacturing technology for the burnable poisons also gives a lot of
options, including BP uniformly mixed with the fuel, BP coated on the sides of the
fuel pe
The e able poisons will be
investigate o
Oth p
burn-up [390,
T
llet of on internal side of the cladding, removable BP roads.
b havior, advantages and disadvantages of the burn
d f r the forming of the set of the burnable poisons for MASLWR.
er ublications related with control rod efficiency and effects on the fuel
391, 392]
2.2.6. Studies on optimization of the fuel use
Optimization of the fuel use is another aspects for the modern development of
the nuclear power plants. Current dissertation does not intent to optimize MASLWR
parameters or fuel utilization, however the understanding of the optimization
techniques, parameters of the reactor optimization and tools available will be useful.
The several publications we considered [212-223] as a different examples of the
reactor optimization.
idered publications several levels of optimization are
he
eactor [274]
According to the cons
applicable to the reactor design:
• Fuel Pin Pattern in Fuel Assembly Optimization (2D, 3D) [215]
• Fuel Assembly Loading and Refueling optimization (2D-3D) [186,
191, 216, 217]
• Burnable Poisons Optimisation (in order to increase fuel burn-up or
decrease pin peaking factors) [213, 214]
• Optimization of the neutronic parameters trough variation of t
Fuel/Moderator ratio (spacing, pith-diameter ratio) [218, 221]
• Optimization of the thermal-dynamic and Thermal-hydraulic
parameters of the r
33
ses:
the current dissertation, and feasibility studies, we do not
need to
to consider various solutions and
various is analysis, systematize
and narrow e
of the importa ll as development of analysis and
search algo perform
our researc e
2.2.7. Be c
Separately we could split optimization studies on a several clas
• Development of optimization algorithm (method)
• Optimization of the calculation algorithm or search algorithms
• Creation of the optimization tools for a given task (like fuel pin pattern
optimization in FA, or FA load pattern optimization in the reactor)
For the purposes of
find an optimal solution in order to prove feasibility of the design, however,
the iteration process will be necessary, in order
configuration of the reactor core. In order to perform th
th scope of the studies analysis of the important parameters and creation
nce function will be necessary as we
rithm. The analysis of the examples listed above will help us to
h b tter.
n hmark of the tools, libraries in a new area of parameters
The proposed MASLWR design param
PWR or BWR param
and natural circulation, resul
should be done at the edge of the proved and verified code applicability limits, and
sometimes even beyond them. Several publications and bench-mark databases were
considered and will be u
eters are different comparing to the
eters. The fuel temperatures, enrichment, power density level,
ting the fact those calculations of the rector parameters
sed as a references in a further chapters of the dissertation.
Particularly IAEA / NEA Benchmark Data-base for the code verification and
related publications [57, 58], NEA technical Reports [248, 250, 261-267] and other
publications [224, 232,360,388].
34
rtation2.2.8. Spent fuel handling, storage and transpo
odes (when the fuel assemblies could not be moved or removed from the
core, or any additional (new, external) absorbing elements inserted into the core,
besides designed elements.
this topic
provide
own phase) [373, 377,396, 428]
at
reactor will remain subcritical in a potential transportation ascidians, and well as prove
nd
transportation. This problem even more complicated if we assume that reactor could
have fu
Spent fuel transportation is usually a separate part of the NPP design. However
for the MASLWR option with off-site refueling creates necessity to consider spent
fuel wet storage, dry storage and transportation condition as a part of the core
operation m
The good knowledge base and collection of the documents on
a NRC NUREGs and ORNL documents.
Particularly we could specify several important issues:
• Reactivity criticality issues during the transportation of the fresh fuel
(or reactor modure with a fresh fuel)
• Reactivity/criticality issues during on-site or off-site refueling
• Spent fuel criticality issues in a wet storage (spent fuel pull, reactor
core during a cool d
• Dry Storage criticality and fuel behavior issues (also for a dry storage
of the reactor fuel in reactor module) [387, 402]
• Reactivity/criticality issues during the transportation of the MASLWR
module with a spent fuel [379, 383, 394, 400, 404, 411, 427, 433-435,
440-445]
For the off-site refuel able MASLWR option will be necessary to prove th
reasonable heat removal from the spent nuclear fuel during dry storage phase a
el pin leak acceding, and should be removed in with partially burnt fuel and the
middle of the cycle
35
2.2.9. Natural Circulation Systems [11, 26, 29]
2.2.10
The natural circulation and passive safety systems of the MASLWR is a
unique feature of the current reactor design. In order to learn about natural circulation,
natural circulation systems and phenomena, several IAEA publications were studied
[11,26,29].
. Coupled Neutronic-Thermal hydraulic calculations
Coupled neutronic and thermal hydraulic has a greater importance for the
natural circulation driven system, such as MASLWR. The Power Flow correlation
could result oscillations and instabilities in the reactor. Also having temperature
gradien
2.3. N
t across the core equal to 60K, we should understand that reactor will be very
sensitive to the fuel and coolant temperatures and distribution. Smaller flows rate
(compare to PWR reactor) increase importance on the cross-flow between fuel
assemblies in the reactor analysis.
The set of tools, developed by Studsvik Scandpower allow to model coupled
system with a constant flow trough the fuel assemblies [236] for the steady state
depletion analysis in SIMULATE 3. However SIMULATE-3K Coupled with RELAP
model of the primary look and 3D model of reactor core could allow to model
transients with natural circulation with a cross flow [236].
Other publications were also considered in order to investigate coupling
phenomena in LWR and capabilities of the other codes systems [251-253, 457-460]
ational and International requirements and regulations
The purpose of this, literature review section, is to present the current structure
of the regulations and design requirements on national and international level,
applicable to MASLWR core design, and demonstrate a knowledge basis (requirement
basis) for a MASLWR core design.
36
2.3.1. International Requirements, IAEA
Traditionally the top and a most general level of any norms and regulations are
nventions. For the nuclear
te and
safe
agr of
, peer-reviews,
m
ns programs (India,
Pakista
ion in nuclear energy use.
ommittee, Nuclear Suppliers
accountancy and control, and organization and functioning of export control system
h
the international multilateral and bilateral agreements, co
energy this top level represented by Non-proliferation treaty [32], IAEA statu
agreements that each country conduct with IAEA, and set of conventions on nuclear
ty and liabilities.
Particularly, IAEA Statute covers the system of other documents and
eements that create a basis for the work of IAEA with the country in terms
safety assessment and experience exchange, as well as inspections
se inars, educational programs and development activities.
The Nonproliferation Treaty – is a fundamental document, and a basis for a
Non-proliferation regime, where IAEA represents international community and report
to the Security Council of the United Nations. Non-Proliferation Treaty (NPT) is a
basis for a next level of the international treaties and organizations, responsible for the
control on export of nuclear material and technologies. NPT postulate the existence of
five nuclear weapon states (USA, Russia, UK, France, China) – the countries that
internationally allowed to have nuclear weapons, and military nuclear programs, non-
nuclear weapons states – the countries which voluntarily agree to not have a nuclear
weapons programs, and receive a benefits of the international cooperation in peaceful
use of nuclear energy use. However there is still several countries who are not
members of NPT and have nuclear weapons or nuclear weapo
n, Israel, North Korea). These countries could not fully participate and have
benefits from international cooperat
Institutionally NPT supported by IAEA, Zanger C
Group, Euroatom and a number of the less important organizations. From a norms and
regulation point of view NPT cover such aspects as Nuclear materials security,
(involving Zanger Committee and Nuclear Suppliers Group). The documents, whic
37
h
NPT, a
FCIRC-225 “The Physical Protection of Nuclear Material” [35]
• INF
rt
control
• INFCIRC-335 “Convention on Early Notification of a Nuclear Accident” [36]
• INFCIRC-336 “Convention on Assistance in the Case of Nuclear Accident or
Radiological Emergency” [37]
• INFCIRC-500 “Vienna Convention on Civil Liability for Nuclear Damage” [39]
are important for a MASLWR design ad related with functioning and compliance wit
re:
• INFCIRC-66 “The Agency’s Safeguard System” [31]
• INFCIRC-140 “Treaty on The Non-Proliferation of Nuclear Weapons” [32]
• INFCIRC-153 “The structure and content of arrangements between the Agency
and States required in connection with the Treaty on the non-proliferation of
Nuclear Weapons” [33]
• INFCIRC-254 “Communication Received from Certain Member States Regarding
Guidelines for the Export of Nuclear Material, Equipment or Technology” [34]
• IN
CIRC-274 “Convention on the Physical Protection of Nuclear Material” [40]
These documents are the basis for the nonproliferation system and provide a
practical guidance and reference figures and definitions for the legal issues related
with functioning of NPT. However these document could also be useful for the
designers, because the identify the quantities and qualities for the terms such as “Low
Enriched Uranium”, “Significant Quality” and other figures important for the reactor
design. Also this document gives a classification of the nuclear materials and describes
how accountancy and security systems should be organized for this kind of material
(Nuclear Fuel) during manufacturing, transportation, use in reactor and final disposal,
which finally affect economic parameters and design requirements. Also Expo
guides give us understanding of how the international cooperation and trade
could be organized for MASLWR for a different design options.
The Third Set of Basic International Documents and requirements are the
international conventions, such as:
• INFCIRC-449 “Convention on Nuclear Safety” [38]
38
tion of Nuclear Material” [40]
material or technology. Unfortunately not
all cou
o these conventions IAEA provides a publications which explain
how th
andards [42], Safety Issues for a transport of
Nuclea
• INFCIRC-274 “Convention on the Physical Protec
These conventions are mandatory requirements for all countries, which ratified
them and ratified them, and effective for vendor country (or countries), end-user
country and countries of transit of nuclear
ntries yet ratified all this conventions (for example convention on Civil
Liability for Nuclear Damage is not enter into force in some CIS countries), so than
bileteral or multileteral agreements will be necessary for the supply of nuclear
facilities or services. This conventions creates a basis for the national legislations in
the IAEA memberstates, and befines a basic principles and figures for the legislation
acts.
In addtiton t
is conventions should be implemented and reflected in the country legislation
and international contracts [47]. In order to facilitate the implementation of this
conventions and assist memberstate legislators and operators of nuclear facilities,
IAEA also publishing a series of Safety Standards and Guides, Quality Standards and
Evaluation Guides. The documents importaind to the MASLWR Core Design are
given in a references [41-51]. This documents covers such aspectrs as Radionuclide
Source Term Evaluation [41], Quality St
r Materials [43], Core design standards [44], Fundamental Safety Principles
[46], as well as NPP economical and live-time management guides [45, 49, 50, 51].
Most of these principles and requerment are reflected in the US National
Rulles and Regulations, however for the international marketing of MASLWR will be
importaint to demonstrate complience with IAEA safety standatds, as a basis for many
national safety standards and requerments (even while some national requerments
could be harder and more restictive than IAEA recommendations).
39
2.3.2. International Requirements (Other)
In additions to IAEA Safety and Security Standards and requirements, some
other type requirements could be applicable for the MASLWR design, and promotion
of the MASLWR on international markets.
ISO – International Standards on quality assurance, business and industry
organization, are well known, and applicable not only for the Nuclear Industry, but
also for the relative supply chain for the plant design, construction, operation and
decommissioning. Certificate of compliance with ISO is becoming a mandatory
require
ype of international or multinational requirements, is requirements
and saf
te a Utility Clubs, in order to
generate combined general requirements. The most famous Utility Club – is European
EUR-Club club was started as a research project focused on harmonization of
the requirem
e
experience of large LWR reactors in Europe, and covering only PWR and BWR
e 1 covers
“Generic
ment for most of the international tenders for the supply of equipment or
services for the nuclear facilities. The comparison of ISO and IAEA standard is given
in publication [42].
Another t
ety standards developed by multinational groups or organizations. For example
Vendors Creates a so-called owners groups (For Example Westinghouse Owners
Group) and within the owners group, the additional or specific safety and quality
standards could be implemented. Utilities also could crea
Utility Requirements Club (EUR-Club).
ents in the European Union Countries for the nuclear power sector, and
summarizing the general utilities demands and requirements for the potential vendors
[74]. However later EUR-Club become a strong player in the European Market [72,
73] by creating of well-structured set of utility requirements [68-71] and performing
evaluation of the basic potential designs for the European Market (including
W stinghouse AP/EP-1000 [78] and Russian VVER-1000/AES-2006).
The EUR requirements currently created according to the operational
reactors [77]. The requirements are structured in four basic volumes. Volum
requirements related with main policies and objectives [68], Volume 2 –
40
pplication of EUR
lly
4 are available for a general public, Volume 3 contains some specific design
nd available only for the EUR-Club
set of
require
the utilities and vendors. (for
exampl
h EUR requirements is also important marketing fact for
even fo
ure marketing.
Nuclear Island Requirements” [69], Volume – 3 demonstrate the “A
requirements to the Specific Projects” evaluated by EUR-Club [70], and fina
Volume 4 gives a “Generic Conventional Island Requirements” [71]. Volume 1,2 and
information, and design evaluation results a
members. For the MASLWR core design Volume 2 is the most valuable
ments, however volumes 1 and 4 also should be considered as long as they
could affect a core and reactor design.
The creation of EUR also related with an opening of the electricity market in
Europe [75] and increased internationalization of
e Italian utility ENEL own NPPs in Slovakia, even while nuclear energy
banned in Italy).
The Compliance wit
r Non-EUR utilities in neighboring countries (such as CIS-Countries, and
Middle East). The demonstration of the MASLWR compliance with EUR-
requirements will be important for the fut
2.3.3. US National Requirements
The main nuclear regulatory body in the USA for the operation of the nuclear
power
88]
• 10
plants is United Stated Nuclear Regulatory Commission. The basic principles
and requirements directly related with design, licensing, construction, operation and
decommissioning of the NPP is given in the Codes of Federal and Regulations 10
CFR. For the MASLWR Design following sets of documents is important:
• 10 CFR.50 “Domestic Licensing of Production and Utilization Facilities” [287]
• 10 CFR.52 “Domestic Licenses, Certifications, and Approvals for Nuclear Power
Plants” [2
CFR.20 “Domestic Standards for protection against radiation” [289]
• 10 CFR.71 “Packaging and transportation of radioactive material” [290]
41
reement”, [294]
• 10
Plant protection [318-324] – The materials
protection and accounting will be necessary or both transportable and non-
transpo le configuration of the MASLWR Reactor, However the difference in the
safeguard requirements for transportable reactor module could be different compare to
the complex of requirements for the design with on-site refueling. Evaluation of both
will be necessary to estimate the best choice from non-proliferation and other
prospective.
Section 7 – Transportation [325-331] – The standards for the transportation
packages for the nuclear materials should be learned for the case of transportable
• 10 CFR. 73 “Physical protection of plants and materials” [292]
• 10 CFR.74 “Material control and accounting of special nuclear material” [293]
• 10 CFR.75 “Safeguards on nuclear material-implementation of US/IAEA
ag
CFR.100 “Reactor Site Criteria”, [295]
• 10 CFR.110 “Export and import of nuclear equipment and material” [296]
These documents give a general, top-level of the design and safety
requirements, and has a similar ideas as IAEA safety and non-proliferation guides, but
applicable for the US entities and US jurisdiction, with a specific requirements
representing USA regulatory and legislation system.
The more detailed and specific requirements are given in NRC regulatory
guides, which are spitted into 10 sections. For the MASLWR Core Design following
sections will be applicable:
Section 1 – Power reactors [297-317] – Regulatory guides regarding operation
of the commercial power reactor, licensing, Material acquainting and Security. This
guides applicable for all NPP in the US.
Section 3 – Fuel and Material Facilities [314-317] – This guides are important
to learn in order to understand the limitation of the fuel facrication facilities (for the
fuel fabrication with 8% enrichment) and for the on-site and off-site refueling and
storage facilities, which could be necessary in some design scenario.
Section 5 – Materials and
rtab
42
reactor with off-site refueling. Than Reactor in transport configuration should comply
with the most of the tra ss NRC would issue a
odule loaded with a
fuel.
asibility study of the reactor concept
and do
number of requirements and publications, which should be taken ito
conside
nsport package requirements, unle
special requirement concerning the transportation of the reactor m
In addition to the NRC requirements, the specific non-nuclear requirements
could also be implemented and affect core design of the MASLWR. The plant
designers should consider particularly ASME, ANSI and other requirements in some
design stages. Current dissertation covers only fe
n’t intend to go tot that level of requirements, unless they has a strong effect of
the core design, and mentioned in NRC requirements or regulatory guides.
NRS also issuing a NUREG series of the publications, which covers the
research works and conferences and contain information important for the safe
operation of the NPP and related fuel and radiation source management and
transportation. These publications are given in the list of reference [332-411]. There is
no space and necessity to characterize each of these publications here. It will be done
in corresponding design chapters of the dissertation. Here important to mention only
that the transportable option of the MASLWR design with off-site refueling creates
necessity for the core designer to take into account a transport configuration of the
core, and address potential transportation accidents and issues. This fact significantly
increases the
ration during the core design.
43
e safety power system. A
rce in the OSU thermal
hydraulic test facility as a model for MASLWR.
estinghouse PWR fuel assemblies, with a shorter
U fuel with
-Pu or U-Pu-Th fuel. The power density proposed in
l to the power de R. The reactor
ed to be cluster control rods in 16 (of the total 24) fuel
our spider driving . The refueling was
proposed to be on-site, but would take place in ent,
the stationar ign,
m design sections of the 2003 report consists of
, the design has passed through several evolutions [106-109]. In
ly interact with po ze
MASLWR. As a result, a feasibility study of the reactor core was necessary.
Chapter 3 – Methodology
3.1. Initial Data – Definition of research goals
The design of the MASLWR reactor was initiated as a U.S. Department of
Energy NERI research project in 2000. The research associated with MASLWR was
primarily focused on the study of a natural circulation passiv
generic, simplified core model was used as the heat sou
The 2003 NERI report [104] and subsequent publications on MASLWR [101-
106] were mainly focused on thermal-hydraulics. The core in these publications was
assumed to use standard 17x17 W
active length. The 2003 report also discusses th use of LEe optional
enrichment up to 19%, or MOX U
that report was 100 kW/liter, equa
control system was propos
nsity of the large PW
assemblies, combined into f mechanisms
a separate refueling compartm
where reactor should be moved from
te
y position. In total, the fuel des
core design and reactor control sys
seven pages.
Since 2003
2006, OSU started to active tential investors to commerciali
The primary goal of the current research is to investigate the feasibility of core
design for a small, natural circulation light water reactor (MASLWR), for 5 years of
non-refuelable operation and maximum fuel enrichment of 8% of U-235. In addition,
we wish to identify and evaluate potential design issues that could appear during
MASLWR commercial design.
44
LWR operational conditions, which is different from
convention er power
ensity, large adients across th he
magnitude of the control rod shadowing and self-shielding effects for the fuel with
higher concentration of burnable poisons (like gadolinium).
3.2. Initial Data – Geometry and materials
Many of the MASLWR design characteristics are not fixed; however the
design specifications of the MASLWR reactor for the purposes of the current research
are given in table 3.2.1 [107-109]:
Table 3.2.1. Main parameters of the MASLWR reactor Thermal power o
Another goal of the current research is to evaluate the neutronic characteristics
of 8% enriched fuel for MAS
al PWR reactors (lower operating temperature and pressure, low
d r temperature gr e fuel assembly), and evaluate t
utput per unit 150 MWt Electrical power output per unit 35 MWe Reactor coolant pressure 8.6Mpa Coolant inlet / outlet temperature (Full power, core average)
491.8 K / 544.4 K
Coolant flow rate trough core 424 kg/s Fuel Type Ceramic UO2,
Standard PWR 17X17 design Fuel Enrichment Not greater than 8% U-235 in Reactor Campaign Length (required) 5 Effective full power years Maximum allowable Fuel Burnup 60 MWD/kgHM (FA Average) Average Fuel Burnup (at 5 years) 49 MWD/kgHM (Core Average) Fuel Load (mass) 5500 kg of fuel Fuel Density 10.1 g/cc Core Average Power Density 85 kW / litre (for Hcore=160 cm)
The core contains 24 assemblies of Westinghouse 17x17 fuel. Each fuel rod is
197.10
ign margins was
conside
4 cm long and contains 160.0 cm of fuel. (An increase in the core height to
176.8 cm to decrease the power density and extend the fuel campa
red; however, for this dissertation, a core height of 160.0 cm was used as the
main reference figure for calculation). Twenty five rods per assembly are guide tubes,
where burnable absorbers, control rods, or instrumentation can be placed, leaving a
45
total of 6336 fueled rods in the core, for a total of 5500 kg of fuel (assuming fuel
density 10.1 g/cc [248, 254]).
Figure 3.2.1. Westinghouse 17x17 PWR fuel assembly design
Fuel Assembly CASMO 3 1/8 Assembly Model
Figure 3.2.2. MASLWR reactor vessel geometry
46
produce
more sp
4) Potential reactivity related accidents should be addressed for the
dule and reactor control system.
The core is located inside a barrel with an inner diameter of 1.5 m and
thickness of 10 cm. At full power, the coolant enters at 491.8 K and exits at 544.4 K,
flowing at 424 kg/s (core average). These figures may be strongly dependent on the
reactor operating mode, secondary side parameters and power generation profile.
MASLWR will produce 150 MWt for approximately 1825 days, leading to a
core average burn-up of 49 MW-days/kgU. Initial calculations have used 8% enriched
fuel, although it is likely that with an optimal layout of burnable poisons to
atially uniform fuel burn-up, a somewhat lower enrichment can be used.
There were no limitations or constraints assumed on power peaking factors and
DNBR limits for the MASLWR core in this current feasibility study.
The originally proposed reactor control system has 16 control rod clusters,
joined into 4 spider drives [104] However, this option is not considered as a final
design solution [107] and other options [113] for the reactor control system will be
evaluated within the current feasibility study. The following recommendations were
made as limiting design constraints:
1) The reactor control system should comply with current US regulations and
should reflect global best practices (i.e. IAEA safety standards, EU safety
requirements)
2) The reactor control system should be able to handle operational transients at
full power, without anticipated reactivity insertions (accidents) (i.e. the worth of the
control unit of the control group should be less than 1$ of reactivity).
3) The reactor core should be designed with negative reactivity feedbacks and
the reactor control system (reactor shutdown) should be able to keep the reactor
subcritical at cold, zero-power conditions.
transportation configuration of the reactor mo
47
ther reactor design codes are currently
availab
ld save resources for the designer and avoids
mistake
e core simulation. Core
calcula
also sh
idates.
43, 244] and the
CASM
asons. Calculations performed with the SSP tools cover all stages of
the reac
3.3. Tools Available for Neutronic design
A large variety of neutronic and o
le on the market [233-244]. Some codes are freeware and supported by national
and international research centers; others are designed for the commercial use by
reactor vendors, utilities, fuel manufacturers, engineering firms and regulatory
authorities. The right choice of tools cou
s and miscalculations.
In the current feasibility study we will need codes for lattice calculations to the
study effects in the fuel assemblies caused by increased enrichment and different
operational parameters. Also we will need a code that can perform depletion of the
nuclear fuel and creates neutronic data libraries for th
tions also should be performed using three-dimensional core models with
consideration of the fuel and moderator temperature changes throughout the core, and
complex power, reactivity and other feedbacks typical for the reactor system.
The results of the calculations should be accurate and reproducible. The code
ould be recognized by the industry as an industry standard, and verified and
licensed for the reactor calculations. Availability of the code for the research team and
future availability for the reactor designers, vendors and utilities is another criteria of
choice. Finally, the input and output formats and interfaces between code components
and interfaces with other codes should be clear and formulated in the units used in
industry. This set of requirements narrows the list of the potential cand
Three set of codes available to OSU were considered: MCNP-5 / MCNP-X
(created and maintained by Los Alamos National Laboratory) [241, 242]; SCALE
(created and maintained by Oak Ridge National Laboratory) [2
O/SIMULATE CMS family of tools by Studsvik Scandpower Inc (SSP) [233,
240]. Each of the packages was used at certain stages of the core design calculations;
however this dissertation will be dedicated to the calculations performed with the SSP
tools for several re
tor design: from pin-cell calculations to core design and RCS simulations. SSP
48
tools h
MULATE 3K with RELAP
or other thermal hydraulic codes for the detailed analysis of the effects of the primary
and secondary circuits on reactor transients, as well as 3D cross-flow analysis inside
dely used by utilities
and ve
3.4. Studsvik Scandpower Tools
s the use of qualified computational tools.
Studsv
exposure and power level). [237]
ed to create the fuel assembly’s cross-
section
comp
It wi for cross-verification of the codes, pin-power reconstruction, time
dependence of burnable poison concentration, and evaluation of reflector geometries.
ave the built-in capability of 3D coupled thermal hydraulic and neutronic
calculations, which may be extended by the coupling of SI
reactor core. The SSP tools are recognized by industry, and wi
ndors in the US and worldwide for core analysis and refueling design.
Consultancy and training provided by Studsvik Scandpower, and the interest of the
company in participating in the future commercialization of the MASLWR is another
driving factor for doing calculations with SSP tools.
Commercial reactor design require
ik Scandpower Inc, develops and markets one such set of industry standard
simulation software. This set includes codes for almost all stages of reactor neutronic
calculations, from the pin (pellet) up to plant simulators. The research tools to be used,
will include the following codes:
INTERPIN – Performs pin-power and pin-temperature calculations. The
results of these calculations will be used as input to CASMO (fuel and moderator
temperature and power density) and SIMULATE (the matrix of the fuel temperature
versus
CASMO-4 – Performs a deterministic solution of the multi-group neutron
transport equation in 2D geometry for LWR fuel. This code simulates fuel assembly
behavior during the whole exposure cycle (and even during the after-use cooling
period). The files generated by CASMO are us
libraries for further use by SIMULATE. [233, 234]
CASMO 4E – A modification of CASMO that allows calculations on more
licated geometries, including the modeling of a 2D slice of the whole core. [235]
ll be useful
49
[236]
reacto portant for safety and operation of the power plant. Code has
loadi
neces
analy can be coupled with RELAP5-3D to model
mp
l
and
ntrol
and S
CASM
3.5.
in the SSP tools. The two principal codes used for the calculations are CASMO – 4
and SIMULATE – 3. The methodology description presented here is summarized and
based on the CASMO – 4 m
other publications [203,204, 212-223].
3.5.1
SIMULATE – Performs a nodal solution of the multi-group diffusion equation
. Used for 3D modeling of the core, fuel optimization and calculation of basic
r parameters im
capabilities for the modeling of different operational modes and complicated fuel
ng patterns. This code is also used for the creation of “restart files”, which are
sary for refueling optimization using SIMULATE or X-IMAGE
SIMULATE 3K – A version of the SIMULATE 3 code with built-in transient
sis capabilities. This code
co licated transients where thermal hydraulics is very important.
X-IMAGE – An extension of SIMULATE used for the optimization of fue
loading patterns and refueling. This code is equipped with a user-friendly interface
built-in optimization scripts to address traditional optimization tasks (co
functions). [240]
CMS-View – Code used to view and builds plots and tables from the CASMO
IMULATE output files. [238]
CMS-Link – Code used for the creation of the fuel assembly libraries from
O output files for further use in SIMULATE. [239]
Physical Models implemented in the tools
In this section, we provide a brief description of the physics methodology used
anual [233, 234], the SIMULATE – 3 Manual [236] and
. CASMO – Code Overview
CASMO-4 [234] and CASMO-4E [235] are multi-group two-dimensional
nsp
simple pin cells. The code handles a geometry consisting of cylindrical fuel rods of
tra ort theory codes for burnup calculations on BWR and PWR assemblies or
50
gadol
instru
fuel
4/4E.
absor
such as MICBURN-3 is no longer required) and a fully heterogeneous model is used
• Nuclear data for CASMO-4 are collected in a library containing microscopic
cross sections in 70 energy groups. Neutron energies cover the range from 0 to
10 MeV.
fuel bundles. However, since most
b
lculations in half, quadrant, or octant symmetry.
• Absorber rods or water holes covering 1x1, 2x2, 3x3 or 4x4 pin cell positions are
allowed within the assembly.
• Thermal expansion of dimensions and densities is performed automatically.
• Effective resonance cross sections are calculated individually for each fuel pin.
• A fundamental mode calculation is performed to account for leakage effects.
• Microscopic depletion is calculated in each fuel and burnable absorber pin.
• In the depletion calculation a predictor-corrector approach is used which greatly
reduces the number of burnup steps necessary for a given accuracy. This is
particularly important when burnable poison rods are involved.
varying composition in a square pitch array with allowance for fuel rods loaded with
inium, erbium, IFBA, burnable absorber rods, cluster control rods, in-core
ment channels, water gaps, and cruciform control rods in the regions separating
assemblies. Reflector/baffle calculations can also be performed with CASMO-
CASMO-4/4E incorporates the direct microscopic depletion of burnable
bers such as gadolinium and erbium into the main calculation (an auxiliary code
for the two-dimensional transport calculation.
Some characteristics of CASMO-4 are listed below:
• The two-dimensional transport solution is based upon the Method of
Characteristics and can be carried out in a number of different energy group
structures.
• CASMO-4 can accommodate non-symmetric
undles are symmetric, it is possible to take advantage of this symmetry and
perform the ca
51
• The CASM s edits of few-
group cross sections (macro and micro) and reaction rates for any region of the
assembly. An ASCII card image file (CI-file) is created for linking to various
diffusion theory core analysis programs, e.g., CMSLINK/SIMULATE.
• The user may specify group structures and regions for which CASMO output is
desired. The standard output is brief, but options are provided to generate
extremely detailed output if desired.
• Reflector calculations are easily performed and discontinuity factors are
calculated at the assembly boundaries and for reflector regions. (However there
are limitations on the modeling of strong absorbers in the reflector region and
creation of the necessary segments for SIMULATE 3, so CASMO 4E is
necessary for the evaluation of reflector/absorber configurations)
• CASMO-4 can also perform an 18 group gamma transport calculation if
requested.
• CASMO-4 is a “user’s” code and as such has a simple user oriented input.
Default values are available for many input quantities and nuclear data are
automatically read from the library (which is an integral part of the CASMO-4
code package). Complicated assemblies can be modeled with just a few basic
input cards.
• Input and number densities may be saved on a restart file at each burnup step to
be used in subsequent calculations, e.g., coefficient calculations at different
exposures.
• CASMO-4 and CASMO-4E are written entirely in FORTRAN 77 and will run
on most modern UNIX® and Linux workstations, and P.C.’s.
• Multi-assembly, MxN calculations in CASMO-4E for both BWRs and PWRs
use the same transport theory methodology (Method of Characteristics) as the
si
O-4 output is designed to be flexible and generate
ngle assembly calculations.
52
3.5.2 CASMO – Methodology Overview
The simplified flow of calculations in the CASMO-4 program is shown in
Figure 3.5.2. [234]
Figure 3.5.2. CASMO-4 calculations flow diagram
53
tc., provided in the user’s input.
elates tabulated
effectiv
self-collision probability. The resonance integrals
obtaine
defined to lie between 4 eV and 9118 eV. Absorption
above 9
n energy spectra in 70 energy groups
to be u
so that individual spectra are obtained for each pin type, e.g., pins
contain
ro-group spectra for water
holes within the assembly.
A 2D macro group calculation in 40 energy groups (default) is then performed
for the entire bundle using homogenized pin cells. The solution from this calculation
Macroscopic group cross sections are prepared for the micro-group
calculations. The nuclear data library in 70 energy groups is an integral part of the
code system and macroscopic cross sections are directly calculated from the densities,
geometries, e
The effective cross sections in the resonance energy region for resonance
absorbers are calculated using an equivalence theorem, which r
e resonance integrals for each resonance absorber in each resonance group to
the particular heterogeneous problem. The equivalence expression is derived from
rational approximations for the fuel
d from the equivalence theorem are used to calculate effective absorption and
fission cross sections. The “shadowing” effect between different pins is taken into
account through the use of Dancoff factors that are calculated internally
by CASMO-4.
The resonance region is
118 eV is assumed to be unshielded. The 1 eV resonance in Pu-240 and other
low energy resonances in plutonium and other nuclides are adequately covered by the
concentration of thermal groups around these resonances and are consequently
excluded from the special resonance treatment.
The cross sections thus prepared are used in a series of collision probability
micro-group calculations to obtain detailed neutro
sed for the energy condensation of the cross sections.
The micro-group calculation is quite fast and is repeated for each type of pin in
the assembly,
ing fuel of different enrichment. To provide micro-group spectra for
condensation of an absorber pin cell, a micro-group calculation is carried out for the
absorber rod surrounded by coolant and a buffer region representing the surrounding
fuel pins. The same procedure is used to determine mic
54
ensation of cross sections for use in 2D
transport calculations.
ckling calculation is used to modify the results
obtained from the transport calculati
for each region containing a burnable absorber. 10 radial rings are typically used for
the depletion of Gd within a pellet.
ed using a predictor-corrector approach. For
each bu e, first using the spectra at the start of
the ep ing the spectra at the end of the
step ese two calculations are then used as starting
val
rs for use in core
calc a for assembly
inte c s can be used by
SIMULATE in two group diffusion theory in order to preserve net currents calculated
by
cluding data for homogenized baffle/water are accurately
generated by a two-dimensional calculation modeling one segment plus the reflector
on
ated case matrix capability (S3C) for generating
data suitable for downstream 3D nodal codes, e.g. SIMULATE. This capability will be
used also for evaluation of the reactivity coefficients and control rod worth from
assemb level calculations for various sets of parameters.
gives neutron spectra for the final energy cond
The data generated in the previous steps constitute the input to the
heterogeneous, two-dimensional characteristics based transport calculation, normally
performed in 8 energy groups, which gives the eigenvalue and the associated flux
distribution.
A fundamental mode bu
ons to include effects of leakage.
Isotopic depletion as a function of burnup is calculated for each fuel pin and
The burnup calculation is perform
rnup step, the depletion is calculated twic
st , and then after a new spectrum calculation, us
. Average number densities from th
ues for the next burnup step.
CASMO-4 has a flexible output format and print options, e.g., the eigenvalue,
the power distribution, reaction rates and few-group paramete
ul tions. The output also contains flux discontinuity factors
rfa es and reflector regions. These discontinuity factor
the CASMO-4 multigroup transport solution.
Reflector data, in
one side.
CASMO-4 contains an autom
ly
55
3.5.3 SIMULATE – Code Overview
SIMULATE-3 [236] is an advanced two-group nodal code for the analysis of
both PWRs and BWRs. The code is based on the QPANDA neutronics model which
em n
both the fast and thermal groups. Key features of SIMULATE are:
• Pin power reconstruction
• No normalization required against higher order calculations
• No adjustable parameters in the nodal model
• Explicit representation of the reflector region
• Easy-to-use free format input
• Easy-to-use binary cross-section library
• Expanded cross-section modeling capabilities
• Automatic geometry expansion
• ritten entirely in FORTRAN 77
rt for plant operations
•
ploys fourth-order polynomial representations of the intranodal flux distributions i
W
SIMULATE can be used for fuel management, core follow, and reload physics
calculations. The types of calculations performed by the code are:
• Depletion in two or three dimensions, 1/8, 1/4, 1/2 or full core
• Reload shuffling including reinsertion of discharged fuel
• Reactivity coefficient calculations
• Rod worth, including shutdown margin, dropped and ejected rod in two or
three dimensions
• Xenon transients
• Core follow suppo
Start-up predictions
• Criticality searches
56
following logic for use of the tools is
implem nted:
Figure 3.6.1. The fl with Studsvik tools
3.6. Algorithms of the tools used for the standard calculations
For the standard core analysis, the
e
ow path for steady-state calculations
CMS-VIEW
INTERPIN
SIMULATE 3
CMS-LINK
CASMO 4
CMS-VIEW
binary library of the fuel and reflector se
for viewing SIMULATE output files, and
correction of the inputs.
calculation of control rod worth, power-iterations, and flow
correction (which is important for modeling of natural circulation system) several
Figure 3.6.1. Shows the standard flow path for core calculations, which
includes the basic elements of the SSP code package. INTERPIN is used for
calculations of the fuel temperatures, and fuel temperature table at different power and
exposure levels. CASMO is used for the preparation of the fuel assembly and reflector
libraries. CMS-View can be used at this stage for the graphical representation of
CASMO CAX files or for analysis and comparison of different fuel assemblies. When
all potential fuel assembly files are ready, CMS-Link is used for the preparation of the
gments for use with SIMULATE – 3 in core
modeling. CMS View could be also used
The steady state analysis of the core requires more than one run of
SIMULATE. For the
57
iterati dard
algo
Figure 3.6.2. Standard flow-chart for steady-state calculation with flow correction and control rod, boron concentration, power lever iterations, and
reactivity coefficients calculations.
ons could be required. Figure 3.6.2 shows the modification of the stan
rithm for the steady state core analysis.
CMS-VIEW
INTERPIN
SIMULATE 3
CMS-LINK
CASMO 4
Analysis, Iterations
tion could be performed manually in
SIMULATE – 3 or automatically using the X-IMAGE optimization code. The flow
p
should be cre ement of the
fuel assembly in the core or the replacement of the fuel assembly with fresh fuel, or
the fuel assembly from the separate batch, which contains in restart file. The
optimization in this case is performed manually, with restarting of many cases.
X-IMAGE automates this process, with optimization of the fuel-reloading
pattern using multiple criteria.
Refueling is out of the scope of the current feasibility study; however, the
output and restart files created during the current research could be adopted for the
refueling studies.
Fuel, Power, FlowControl Rods, Bor
The refueling studies and optimiza
ath for the refueling studies is shown in Figure 3.6.3. In both cases, restart files
ated for the refueling period. SIMULATE 3 allows the mov
58
tion (left – for manual de).
Figure 3.6.3. Flow charts for refueling optimizaoptimization, right – for automatic optimization using the X-Image co
INTERPIN
CMS-VIEW
SIMULATE 3
CMS-LINK
CASMO 4
INTERPIN
CMS-LINK
CASMO 4 CMS-VIEW
Creation of RESTART FILES
SIMULATE 3Optimisation
Refueling (Manual)
SIMULATE 3Creation of
RESTART FILES
X - IMAGEOptimisation
Refueling (Auto)
Two tools can be used for the evaluation of reactor control system (RCS)
options. CASMO-4E is useful for the quarter-core analysis of complicated RCS
geometries in 2D core with buckling parameters. SIMULATE–3 can be used only for
the standard geometries of the RCS, but provides a more detailed analysis because of
its 3D core geometry and thermal feedbacks.
Figure 3.6.4. Flow chart for the comparison analysis between CASMO 4E and SIMULATE 3 results (used for uncertainty evaluation)
INTERPIN
CMS-LINK
CASMO 4
SIMULATE 3 CASMO 4E
Comparison Analysis
59
3.7. Organization of the feasibility study
3.7.1. General research organization
The organization of the design of the MASLWR reactor is different from the
traditional LWR (PWR or BWR) with forced circulation. This difference is caused 1)
by the physics of t orical order of the
research work.
The natural circulation of the coolant means that the designer has a limited set
of controls on the coolant flow rate [26,29]. In MASLWR, the flow rate is not defined
by the capability of coolant pumps, but by the combination of reactor parameters such
as reactor power, coolant inlet temperature, pressure, density difference on core inlet
and outlet, length of the riser and thermal hydraulic friction in the reactor core and
steam generator [104]. For the designed reactor, the only means of the flow rate
control is 1) change of the core power, 2) changing of the coolant pressure and 3)
changing of the core inlet temperature (through the change of the steam generator feed
water temperature and flow rate).
Historically MASLWR began as a research project focused on thermal
hydraulics of the natural circulation driven system [101, 102]. So the parameters of the
coolant were chosen based on the natural circulation requirements and experiments
with a simplified core model involving a uniformly distributed heat source without
reactivity feedbacks [104]. As a result, the thermal hydraulic parameters of the reactor
core considered as defined for the full power, and the purpose of the this feasibility
study is to investigate feasibility of a core design that fits these parameters, safety and
operation requirements.
The design approach and logic chosen for the feasibility study reflect the
current situation. The general flow chart presented in Figure 3.7.1 covers five major
steps of feasibility studies structured in order of increased complexity and required
knowledge about the new fuel and core. The first step covers the study of the new fuel
type, and effects in the reactor caused by increased enrichment and different
operational parameters. The second step is dedicated to the prototypical design of the
he natural circulation system, and 2) the hist
60
neutron reflector, assuming a uniformly loaded core. The third step is focused on
design of the burnable absorber layout for the prototypical core. The fourth step is
of prototypical core.
The fin
focused on the evaluation reactor control system options for the
al step of the current study is dedicated to the discussion of the operational
issues and uncertainty analysis.
Figure 3.7.1. General research organization
1. Analysis of initial Data, Design Criteria, Phenomena important for design
OutputBasic knowledge about effects of increased enrichment and lower
operational parameters on fuel and reactor behavior (New area of knowledge)
Knowledge base for future licensing of the increased enrichment fuel
OutputSelection of the preferred reflector design
Discussion and Justification of preferable reflector optionAll further studies will be performed for the selected reflector geometry
3. Burnable Absorber Studies
OutputPrototypic core with 3D profiling of burnable absorbers,
programable burn and adjustConclusion of feasible cycle len
ed peaking factors.gth, and fuel burn-ups
Prototypic Core ConceptEvaluation of Thermal hydraulic parameters for prototypic core
4. Reactor Control System (RCS) Studies
OutputConclusions on RCS design options and evaluation of RCS parame
Recommendations for further design of the Reactor Control Syters
stem
Initial Data, (NERI Report, TOR, Technical Specification)
5. Relevant studies regarding operation modes, strategiesSensitivity and uncertainty analysis for the parameters deviation
Support studies for basic conclusions and recommendations
61
3.7.2. Step 1 Initial Fuel and Core Analysis
The first step of the feasibility study is analysis of the initial data, design
criteria and phenomena important for the MASLWR core design. The main goal of the
first step is to learn the basic differences between the MASLWR core design and a
traditional PWR; identify and evaluate phenomena caused by increased enrichment
and lower operational parameters; and evaluate leakage effects for the small core.
At this step, fuel assembly studies are performed for the core average
parameters assuming fuel depletion in a core with chemical shim for the compensation
of the excess reactivity (soluble boron in coolant) and for the non-borated coolant.
Standard geometry of the cluster control rods and standard control rod materials (B4C,
Ag-In-Cd, Hafnium) and semi-standard arrangement and concentrations of burnable
absorbers will be modeled.
he core will be uniformly loaded with fresh fuel without burnable absorber
The coolant flow rate will be assumed equal to the core average in all fuel assemblies,
without mixing, cross-flow and bypass flow. The standard PWR baffle will be
considered. Eigenvalue calculations will be performed for the core with and without
ds. F latio ditions will be considered (unless
otherw
um calculations will be made for the fuel assembly average spectrum,
and cha
T .
control ro or all calcu ns Hot-Full Power con
ise specified; for example, for reactivity coefficient calculations).
The comparison will be done for 8% and 4.0 – 4.95 % fuel for both the
MASLWR and PWR operational conditions.
Spectr
racterized by the fraction of the thermal neutron flux versus total neutron flux
at the full power level.
The effective fraction of the delayed neutrons will be calculated, and used for
the evaluation of the reactivity effects and control rod worth in $.
62
Figure 3.7.2. Step 1 Details
1. Analysis of initial Data, Design Criteria, Phenomena important for design
Comparison of4.0%, 4.95% and 8
Enriched FuelAnalysis, Phenomena Identification
.0%
Modeliwith 4.0
ng of Reactor Core%, 4.95% and 8.0%
Enriched Fuel (Uniform Load, no BA)Analysis, Phenomena Identification
Fuel Assembly Level Reactor Core Level
Multiplication Factor vs. DepletionNeutron Flux and SpectrumEffective fraction of delayed neutronsPlutonium ProductionCR WorthReactivity Effects
Multiplication Factor vs. DepletionTemperature and Power ProfilesPeaking FactorsEffective fraction of delayed neutronsCR WorthReactivity Effects
OutputBasic knowledge about effects of increased enrichment and lower operational parameters
on fuel and reactor behavior (New area of knowledge) Knowledge base for future licensing of the increased enrichment fuel for MASLWR
3.7.3. Step 2 Reflector Studies
The second step of the feasibility study is a study of various reflector options
The goal of step two is to investigate the effect of the reflector material and design on
neutron leakage and core design, and to select the reflector design for further
ns.
Several possible designs of the reflector will be simulated with CASMO and
SIMUL
pass flow, and Hot-
Full po
ycle length will be subjects
for investigation.
.
calculatio
ATE. The core model for the reflector studies involves uniformly loaded fuel,
no-soluble boron for the depletion calculations, uniform flow rate in all fuel
assemblies equal to the core average flow rate, no cross-flow or by
wer operational conditions.
The eigenvalue versus burnup, burnup maps and c
63
The selection of the reflector option will be based on the following criteria: the
longest cycle achievable with the reflector configuration, and the burnup distribution
in the core. The selected option will be used for all further simulations and
calculations.
Figure 3.7.3. – Step 2 Details
2. Reflector Studies
Creation of the reflector Data Libraries with CASMO 4
Modeling of Reactor Corewith different
Fuel Assembly Level Reactor Core Level
Axial and Radial reflector segments Multiplication reflector Options
No adFactor vs. Depletion
Cycle limits B=60 MWD/kgHM)
ditional analysis at this step Burn-up Profiles (assembly average)Peaking FactorsEnd of (K=1 or
Reflector Concept Options (Geometry, Materials)List of Reflector materialsDrawings of potential reflector geometriesCalculation models for reflector studies
OutputSelection of the preferred reflector design
Discussion and Justification of preferable reflector optionAll further studies will be performed for the selected reflector geometry
3.7.4. Step 3 Burnable Absorbers
The third step of the feasibility study involves burnable absorbers. The goal of
step three is to determine the burnable absorber material, design, geometry,
technology and loading pattern for the prototypic MASLWR core.
The core design will evolve from the radial profiling of the burnable poisons to
the three-dimensional profiling of the burnable absorbers.
64
The effort will be ini r the configuration with an
acceptable level of excess reactivity over the reactor campaign, and will evolved to
achieving acceptable fuel assembly burn-up (below burnup limits) with increasing
utilization of the outer fuel assemblies. The next series of iterations will be focused on
the decreasing the peaking factor and implementation of the programmed burnup
strategy. Finally, thermal hydraulic parameters for the prototypic core will be
evaluated.
The core model for the calculation will involve the following features and
assumptions. The nodalization of the problem will include 4 radial and 40 axial nodes
per fuel assembly and up to 8 loading segments per fuel assembly. The neutron
reflector will be chosen as that selected in step two. A uniform flow through all fuel
assemblies equal to the core average flow rate without cross-flow and bypass-flow is
assumed. Hot full power conditions are used and control rod (CR) drives are assumed
to have double positioning precision (200 positions, with 0.8 cm insertion step)
compared to standard PWR CR drives (1.59 cm insertion step).
The eigenvalue iterations, soluble boron iteration, control rod iterations and
power iterations are performed during the search process.
The burnable absorber material considered in this stage is gadolinium oxide
uniformly mixed with the fuel, with concentrations from 0.5% up to 10%. The
maximum number of gadolinium concentrations per fuel assembly segment is three.
The number of different Gd concentrations in the fuel assembly is no more than ten
with a step of 1.0% (in order to fit to manufacturing requirements).
Additional studies of shielding effects and shadowing effects will be
performed in fuel assembly and core level calculations to develop an increased
knowledge base and better support the calculations in the evaluation of the reactor
control system.
tially focused on the search fo
65
Figure 3.7.4. Step 3 Details 3. Burnable Absorber Studies
Fuel depletion with BANon-standard geometry
Fuel Assembly Level Reactor Core Level
OutputPrototypic core with 3D profiling of burnable absorbers,
programable burn and adjusted peaking factors.Conclusion of feasible cycle length, and fuel burn-ups
Discussion and Selection of BA Material, Concept and TechnologyMaterials:Gadolinium (Gd O ) Erbium
2 3
(Er O )Boron Carbide
2 3
ium Boride (ZrB )
2
(B C) ZirconOther
4
TechnologyUniformly mixed with fuel Coated on the fuel pellet or cladding
Combination of options
Separately placed, non-removableRemovable
Burnable Absorber Studies
Fuel depletion with BAStandard geom
Different concentrationsetry
Self SCR Sh
Different concentrationsSelf Shielding effectSelf Shadowing effectCR Shadowing effectNeutron flux transformation in FA
hielding effectadowing effect
Radial BA Profiling
Single BA concentrations per FAMultiplication factor vs. DepletionStudy of under-burn or the fuel
Radial and Axial BA Profiling(3D Profiling)
Single BA concentrations per FAMultiplication factor vs. DepletionBoron concentration vs. DepletionPower peaking, offsetsBottom centered core
Segments with multiple BA concentrations
2 and more concentrationsPrograming of multiplication factor curveSmall BA concentrations for BOC
Radial and Axial BA ProfilingVarious Segments
Multiplication factor vs. DepletionBoron concentration vsPower peaking, offsets
. Depletion
Design IterationsProgramable fuel burn (burn-in-wave)
Prototypic Core ConceptEvaluation of Thermal hydraulic parameters for prototypic core
66
3.7.5. Step 4 RCS Evaluation
The fourth step of the feasibility study is the evaluation of reactor control
system (RCS) options. The goals of step four are to evaluate reactor kinetics
parameters (for steady state operation modes), evaluate different options of the reactor
control system design and RCS functions, and propose a framework for further
development.
The initial evaluation of the RCS begins with a discussion of the RCS
functions, mechanisms and design criteria. Various geometries of the primary RCS
will be considered, such as cluster and cross-blade control rods, and reflector absorber
drums. The evaluation of these geometries will be performed with CASMO 4E in two-
dimensional geometry.
The remainder of the RCS evaluations for the prototypic core will be
performed with SIMULATE 3 assuming cluster control rods. The evaluation of the
RCS with control rods will be performed for the core with and without soluble boron
at Hot Full Power, Hot Zero Power (Zero flow) and Cold Zero Power operational
conditions. Control rod worth will be calculated for transportation accident conditions
(core flooded with fresh water at 4 °C and atmospheric pressure, with failure of the
most effective single CR cluster).
The flow model for all evaluations assumes a uniform flow rate through all
fuel assemblies equal to the core average flow rate, without bypass flow and cross
flow. The core reflector is chosen to be the result of step two of our study, and the
burnable absorber loading pattern is chosen to be the result of step three.
Reactivity coefficients are calculated for the whole core. Eigenvalue
calculations, boron iterations, and control rod iterations will be performed for each
configuration of the RCS. The comparison of the “CR+Boron” and “Boron Free”
cores is performed in terms of excess reactivity control, peaking factors, and burnup
distortions.
67
Figure 3.7.5. Step 4 details 4. Reactor Control System (RCS) Studies
Fuel Assembly Level Reactor Core Level
OutputConclusions on RCS design options and evaluation of RCS parameters
Recommendations for further design of the Reactor Control System
Class cation of RCS functions, control mechanisms and design requiermentsFunctions:
ontrol cy rea
Long term excess reactivity compensation NeutroOther
ifiMechanisms:
Reflector / Absorber RCS mechanisms
Power c and transientsEmergen ctor scram
Control rods (Cluster or X-blades)Chemical Shim(Soluble Boron)
Burnable AbsorbersPower, Temperature, pressure, Flow and other reactivity control mechanism
n flux correction and compensation
Evaluation of RCS Compliance with Rules and Regulations
Correction and creationof Necessary FA Segments
RCS Model for Prototypic core
CZP, HZP and HFP conditionsControl rods + BoronControl rods onlyTransportation Conditions RequirementsModifications of fuel and CR designintegral CR worth, CR Group assignment and worth
Evaluation of the RCS design options and materialsNeutron Absorbing Materials:Silver Indium Cadmium (AgInCd)Hafnium (Hf)Boron Carbide (natural and enriched boron)
Boric Acid Other
Design Options Evaluation WithCASMOCASMO 4EMCNPSIMULATE 3
Analysis of the reactivity effects and reactor kinetics parameters for different modes of operation
Modeling and Evaluation of the RCS options for different modes of operation
68
he modification of the core design, or design basis may be required as a result
of this set of calculations. The recommendations will be formulated in the conclusions
T
section of this dissertation.
3.7.6 Step 5 Other relevant studies of Prototypic core
Finally, the fifth step of the feasibility study addresses the issues of the
operation strategy and uncertainty and sensitivity analysis of the prototypical core
design.
decreased power
operation at the end of cycle could give some extra days / month of operation. This
effect will be simulated and studied for the simulated low power conditions.
Another concern of the natural circulation system – cross-flow between fuel
assemblies, caused by natural (density) driven flow, will be modeled as a corrected
flow case for the prototypic core. The corrected flow assumption means that flow
through fuel assemblies will corrected in order to preserve the core average flow rate,
and achieve the same exit density for all fuel assemblies. The difference in neutronic
characteristics between the uniform flow and corrected flow cases will be also
discussed in later in this dissertation.
The operational strategy of the reactor could yield an additional economical
benefit if power and temperature effects are used. For example,
69
3.8 Goals of Feasibility Study, Criteria of completion
he main products of the current feasibility study will be a determination of
the feasibility of the proposed reactor design with the given thermal hydraulic
parame rs, a prototypic core design (with associated assumptions) and
recomm necessary design considerations and
modifications.
he structure of this dissertation differs slightly from the presented research
methodology, and will include two chapters dedicated to fuel assembly-level studies
and two chapters for the prototypic core design and RCS evaluation.
T
te
endations for further design and
T
70
lts
Chapter 4 Resu
71
Cha
helf” equipment would be used in order to simplify the licensing
pro e
standard y g l be use increased
enrichm he effe n ent is
necessary for the MASLWR core design. It also icensing
of the M ntially applicable for conventional LWR in the event
of future d cycle leng
The issues related with potential use of enrichm in PWRs
were also discussed in publications [53, 60, 201, 203-205, 209, 210, 271, 273, 342,
343, 3
table 4-1.1 below:
pter 4-1 – Study of the Physical Effects in MASLWR 8% Enriched Fuel (General Analysis)
4-1.1 Calculation model
The two major differences between MASLWR and traditional LWRs are use
of 8% enriched fuel and different core parameters (geometry and operational
conditions). [104, 111] However, the design philosophy of MASLWR states that
standard, “of-the-s
cess. [109, 110] The application of this statement to the fuel design means that th
bl Westinghouse 17x17 fuel assem eometry wil d, with
ent fuel. The evaluation of t cts caused by i creased enrichm
w mportanill be i t e lfor futur
ASLWR fuel and is pote
increases in enrichment an th.
increased ent fuel
79, 382], however detailed studies of the neutronic characteristics of 8%
enriched fuel were never performed for the MASLWR parameter range, until this
research. The preliminary results of the current study were presented in 2008 ANS
Annual meeting, and could be found in Transactions [111-112]. The key design
geometry parameters of the 17x17 fuel, used for the modeling of the fuel assemblies,
are given in
Table 4-1.1.1. Fuel Assembly parameters, used in research model Fuel pin lattice arrangement 17 x 17 Number of guide tubes (no fuel) 25 Number of the control rods 24 Fuel pellet OD 0.8260 cm Fuel cladding OD 0.9522 cm Fuel pins spacing 1.26 cm Fuel pellet density 10.1 g/cm3 Guide tube ID/OD 1.14 / 1.22 cm Control rod ID/OD 1.018 / 1.116 cm Control pellet OD 1.010 cm
72
eparately. A comparison of neutronic characteristics of all considered cases can then
y
lower operational parame g operational parameters
were considered:
Table 4-1.1.2. Operational parameters for MASLWR and PWR comparison
A comparison of MASLWR and PWR fuel was performed for two sets of
parameters: MASLWR and a typical, generic PWR. This approach facilitates an
evaluation of effects caused by increased enrichment operational parameters
s
ield general conclusions on the combined effect of the increased enrichment and
ters of MASLWR. The followin
Reactor Type MASLWR PWR Soluble Boron Concentration 0 and 800 ppm 0 and 800 ppm Power Density 30 W/g(U) 40 W/g(U) Fuel Temperature 785 K 1000 K Moderator Temperature 525 K 583 K Primary Coolant Pressure 86 atm 150 atm Coolant Density 0.801 g/cm3 0.704 g/cm3 Saturation Temperature 574.1 K 615.8 K
The above table shows that MASLWR fuel has a lower power density, and a
lower moderator temperature. Both lead to a lower fuel temperature. Lower
tem nd
better modera owever, incre c
e fuel, and as a hardening of the neutron energy
semblies were per e s
, 4.0%
U-235. Two s p
ct c
the fuel with enrichme .5 8 h
e poisons contains 12 fu ns
8% Gd2O3. The layout of the burnable poisons is similar to the M1 modification of
17X17 fuel assembly mentioned in an ORNL report [430].
perature and pressure of the coolant leads to increased density of the coolant, a
tion capabilities. H ased enrichment will lead to in reased
absorption of thermal neutrons in th
spectrum.
Simulations of the fuel as formed for th both ets of
parameters for fuel enrichments of 3.5% , 4.5%, 4.95% (some times referred as
5.0%) and 8.0% weight percent of fuel assemblie with burnable oisons
were also simulated in order to evaluate spe ra el eff ts , and ot ffher e e ucts ca se d by
use of the burnable absorber for nt of 4 % and .0%. T e fuel
assembly with burnabl el pins with 4% Gd2O3 and 16 pi with
73
4-1.2
n conditions
Multiplication factor
The multiplication factor calculations were performed with CASMO-4 in two-
dimensional geometry with reflective boundary conditions. All results presented here
are for an infinite two-dimensional lattice of fuel assemblies (infinite reactor).
Figure 4-1.2.1. Multiplication Factor for different fuel enrichment in MASLWR and PWR operatio
74
depletion under MASLWR and PWR operating conditions are presented in figure 4-
1.2.1. The calculations for fuel with MASLWR parameters are labeled “FM” and
represented by thick solid lines. The calculations for fuel with PWR parameters are
labeled “FP” and represented by thin solid lines. The figure 4-1.2.1 illustrates the
effects of burnup increase with increase of the fuel enrichment, and increase of the
multiplication factor due to lower operational parameters in MASLWR. The summary
of the principal points characterizing the curves is given in the table below:
Table 4-1.2.1. The Multiplication factor graphs characterizing parameters
The results of the multiplication factor calculations as a function of fuel
Fuel Enrichment 3.50% 4.00% 4.50% 4.95% 8.00%EOC (K=1) MASLWR, MWD/kgU 32.9 37.7 42.3 46.6 71.6 EOC (K=1) PWR , MWD/kgU 31.4 35.9 40.3 44.2 68.1 Burnup Extension 1.5 1.8 2 2.4 3.5 Cycle Extension (Relative to PWR) 4.8% 5.0% 5.0% 5.4% 5.1% Initial K (B=0 MWD/kgU) MASLWR 1.3674 1.3939 1.4153 1.4313 1.4963Initial K (B=0 MWD/kgU) PWR 1.3411 1.3664 1.3868 1.4021 1.4650Initial increase of K in MASLWR 0.0262 0.0275 0.0285 0.0292 0.0313Initial reactivity increase in MASLWR, Abs 0.0143 0.0144 0.0145 0.0145 0.0143Initial reactivity increase in MASLWR, $ 1.99 2.01 2.02 2.03 2.00 Burnup K(MALWR)=K(PWR) 42.75 48.20 53.75 58.50 N/A Mulpiplication Factor 0.9295 0.9240 0.9235 0.9215 N/A
An increase in enrichment increases the cycle length and initial excess
reactivity. An enrichment of 8.0% leads to a maximum burnup of 68-71 MWD/kgU;
this corresponds to the point at which the reactor becomes subcritical. An enrichment
of 4.95% corresponds to a maximum burnup of 44-46 MWD/kgU. This fuel does not
meet the design goals for the transportable version of MASLWR with full core
refueling and five effective full power years of operation, which requires a maximum
fuel burnup of 49 MWD/kgU (core average).
Figu ASLWR
fuel reactor is greater than that for a PWR el reactor with the same enrichment, and
that the difference increases with increasing enrichment. This means that the initial
re 4-1.2.2 shows that the initial multiplication factor for the M
fu
75
excess reactivity in MASLWR could be greater than in a traditional PWR, especially
if incr
f the multiplication
properties with the increase of enrichment and decrease of operational parameters in
MASL
Figure 4-1.2.2. Initial Multiplication Factor for MASLWR and PWR conditions
eased enrichment fuel (8.0%) is used. This may require more burnable
absorbers, or other mechanisms for the compensation of initial excess reactivity. On
the other hand, the power defect in MASLWR may be comparable with that of the
PWR (for the same enrichment). Another contributor to the multiplication factor –
neutron leakage - has a greater importance for the small MASLWR core. So, the
necessity of more burnable absorber is obvious; however, current calculations
demonstrate that designers should be aware of the increase o
WR.
Fuel burned at MASLWR’s operational conditions may allow longer cycles
(assuming k∞ =1 as a criteria of the cycle length). However at high burn-ups, the lines
representing the PWR and MASLWR fuel multiplication factors cross. This is caused
76
Figure 4-1.2.3. Fuel assembly M1 - burnable absorbers map
by plutonium accumulation in the fuel. Harder parameters of PWR fuel, greater power
density, lower moderation, harder spectrum (for the same level of enrichment) leads to
a greater level of plutonium accumulation at the end of cycle. This is not important for
the core design of the “transportable” version of MASLWR with full-core off-site
refueling, where k∞ =1 is an EOC criteria. It could be important for core designs with
partial on-site refueling, when local k∞ <1 is possible. The mechanisms of plutonium
accumulation and relative results will be discussed in the next section of this chapter.
Two pairs of fuel assemblies with burnable absorbers were simulated for PWR
and MASLWR operational conditions for fuel enrichments of 4.50% and 8.0%. In
each simul used: 12
fuel pins contain 4.0% Gd2O3 and 16 fuel pins contain 8.0% Gd2O3 uniformly
distributed in fuel pellets with same uraniu enrichment as non-poisoned fuel pellets.
The burnable absorber map is shown in figure 4-1.2.3. The results for the
multiplication factor calculations are presented in figure 4-1.2.4.
ation, the same geometry and types of burnable absorbers were
m
77
Figure 4-1.2.4. Multiplication factor in fuel with burnable absorbers
fuel with bu
n
The behavior of the MASLWR and PWR rnable absorber is
similar for given the same initial enrichme t, However, an increase in the initial fuel
enrichment up to 8.0 % results in a reduction of the effectiveness of the burnable
absorber. Also the shape of the m
richment, the competition between uranium and gadolinium for the
absorption of ther
ultiplication factor curve for 8.0% fuel with burnable
absorber is different from the 4.5% enriched fuel. For 8.0% enriched fuel, the peak of
the multiplication factor curve is not as clear as that for the 4.5% enriched fuel.
Furthermore, the difference between initial multiplication factor and multiplication
factor at the peak is smaller for 8.0% enriched fuel, and the burnup coordinate of the
peak is shifted toward greater burnup levels. The main reason for these changes is the
self-shielding of the burnable absorbers, and fuel-absorber shadowing. In the fuel with
increased en
mal neutrons leads to increased absorption in the fuel, so greater
concentrations of gadolinium will be necessary in order to compensate the initial
excess reactivity.
78
4-1.3 Fuel Depletion, Plutonium Production
4-1.3.1. Contribution of fission of fertile isotopes with burnup
In light water reactors initially fueled with enriched uranium fuel, there are two
processes that primarily contribute to the energy production and cycle length: fission
of U-235 and production and fission of Pu-239 and Pu-241. Although the fast fission
of U-238 is still possible, the contribution of the fast fission to overall energy
production in LWRs is relatively low, about 5-8 %.
Calculations of the inventory of fertile isotopes (U-238, Pu-238, Pu-240) were
performed in the analysis, in order to evaluate the contribution of fast fissionto overall
fuel burnup as a function of the initial fuel enrichment, presence of the burnable
absorbers and operational conditions. The results of these calculations are presented in
figures 4-1.3.1 A, B and C
Figure 4-1.3.1.A. Contribution of fertile fission for PWR fuel
79
ribution of fertile fission for MASLWR fuel
Figure 4-1.3.1.B. Cont
Figure 4-1.3.1.C. Comparison of the fertile fission contribution for MASLWR and PWR fuel burnup
80
sing
enrichment in both the PWR and MASLWR. However, lower operational parameters
and greater density of the moderator in MA LWR lead to the improved moderation of
fast neu
he main contribution to the burnup comes from the fission of fissile isotopes.
Although fast fission is possible for fissile isotopes, these isotopes (U-235, Pu-239 and
Pu-241) have large neutron absorption and fission cross-sections in the thermal energy
range
evaluation of uranium depletion and plutonium production and further depletion rates
is crucial to understanding the behavior of these light water reactors.
These figures show that the fraction of fertile fission decreases with increa
S
trons, so the contribution of the fertile isotope fission in MASLWR is smaller
than in PWR for the same initial fuel enrichment. The increase in enrichment (and
related spectrum hardening) does not increase the contribution of the fertile isotope
fast fission. This implies that fissile isotopes play a greater role in the energy
generation.
The use of gadolinium burnable absorbers may increase the contribution of the
fertile isotope fission in the beginning of the cycle. Gadolinium absorbs thermal
neutrons, increasing fast fission in the fuel pellets with BA compared to non-poisoned
fuel pellets. As a result, the curve for poisoned fuel is above the curve for non-
poisoned fuel of the same enrichment.
T
, so thermal and epithermal neutrons cause the majority of fissions. The
81
4-1.3.2. Uranium Depletion and Nuclear reactions
Figures 4-1.3.1 and 4-1.3.2 show the basic nuclear reactions of interest for
fissile U-235 and fertile U-238 under neutron irradiation.
Figure 4-1.3.2. Uranium 235 depletion nuclear reactions.
Some U-235 may be produced from fertile U-234. A small fraction of this
isotope is present in natural uranium. During enrichment, the concentration of U-234
may increase slightly. However the impact of the U-235 production from U-234 is
negligible, and we may assume that there
initially
is no path to produce U-235 in the LWR
fueled with uranium fuel, so the concentration of U-235 will decrease with
time and fluence. The simplified formula for U-235 depletion is:
( )∫∞
Φ−=0
235235
235 )()( dEEEtNdt
dNaσ
An important component of the U-235 chain is the production and
accumulation of U-236, which is strong resonance absorber of the neutrons. Further
captures and beta-decays will lead to the production of Np-237 and Pu-238, which
accumulate in the spent fuel.
82
Figure 4-1.3.3.A. Uranium 234 depletion for MASLWR and PWR fuel
The above figure shows the initial concentration of U-234 growth with
enrichment. The depletion rates for the same initial enrichment remains similar for a
standard 4.5% enriched fuel, and have minor deviations for the 8.0% enriched fuel.
Also as we can see the de
U-234 compared to standard 4.5% enriched
PWR fuel (
pletion rate of U-234 for MASLWR operational conditions
(dark red bold line) is slightly slower than the depletion rate for PWR operational
conditions (light red solid line). The presence of burnable absorbers may also have an
effect on the U-234 depletion rate. At the end of the irradiation cycle, 8.0% MASLWR
fuel will have twice the concentration of
U-234 concentration at the EOC of 8.0% enriched fuel almost equal to U-
234 concentration at the BOC of 4.5% enriched fuel).
83
Figure 4-1.3.3.B. Uranium 235 depletion for MASLWR and PWR fuel
At the beginning of the cycle, the depletion rates of Uranium 235 under
MASLWR and PWR operational conditions are almost identical for the same initial
enrichment. However the depletion rate of U-235 in 8.0% fuel is greater than the
depletion rate for 4.5% enriched fuel. This is caused by the larger neutron capture and
fission macroscopic cross section, and results in a greater contribution of U-235 to the
burn the
difference in depletion rates become obvi s. Because of the different operational
parame rs, and greater moderator density, the MASLWR consumes U-235 more
actively than the PWR, so the curve representing the weight fraction of U-235 for
up in higher enrichment fuels. However in the middle and at the end of cycle
ou
te
84
R operational
conditi
and could be recycled in PWR, BWR, CANDU or RBMK
reactors using a simple DUPIC [54, 65] process (no water solutions, no separation of
uranium, ty of the
reprocessing of MASLWR fuel may also be greater than that of standard PWR fuel.
Figure 4-1.3.3.C. Uranium 236 depletion for MASLWR and PWR fuel
MASLWR operational conditions lies below the curve for PW
ons, for the same initial enrichment.
8.0% enriched MASLWR fuel will have about 2.5-2.7% weight percent of U-
235 at the end of cycle
plutonium and solid fission products). The economic feasibili
U-236 production ra nder MASLWR and PWR
operational conditions are very similar; even the addition of the burnable poisons does
not significantly increase U-236 production rates (the differences are in the fourth
tes for the 4.5% enriched fuel u
85
signific
centration of U-236 equal to 0.63% for a
burnup
echanisms of U-
238 de
ant figure). For 8.0% enriched fuel irradiated under MASLWR and PWR
operational conditions, the greatest difference between the U-236 weight fraction
curves occurs at the middle of cycle for the burnup of 20-40 MWD/kgHM; however
this difference is also not significant. The fuel irradiated under PWR conditions has a
slightly higher concentration of U-236 than MASLWR fuel in the same burnup range.
The relative difference between the two results is about 0.1%.
Comparing the curves for 4.5% and 8.0% initial enrichment, we see that the
curve for 4.5% has a local maximum con
around 60-70 MWD/kgHM.
The curve for the 8.0% enriched fuel does not have a local maximum in the
current burnup range (0-80 MWD/kgHM), so the balance concentration is not
achievable at this range for MASLWR and PWR operational conditions. The use of
the balance concentration approach is not appropriate for the whole depletion cycle of
8.0% fuel at either MOC or EOC.
At the EOC (60 MWD/kgHM) the concentration of the U-236 in 8.0%
enriched MASLWR fuel will be 1.5 times greater than in standard PWR fuel irradiated
in a typical PWR operational conditions.
In order to have a complete picture of the depletion of uranium isotopes, we
need to consider U-238 depletion (see figure 4-1.3.3. D). The major m
pletion are fast fission and neutron capture with a further production of
plutonium isotopes. The U-238 depletion rates under MASLWR operational
conditions are smaller than those under PWR conditions for fuels with the same initial
enrichment. Also for 8.0% enriched fuel, the initial concentration of U-238 is smaller
than for 4.5% enriched fuel, and depletion rates are smaller as well.
Figure 4-1.3.3. E indicates that the overall consumption of uranium in
MASLWR is slightly lower than in a PWR. MASLWR’s different spectrum is
affecting all reaction rates.
86
Figure fuel 4-1.3.3.D. Uranium 238 depletion for MASLWR and PWR
Figure 4-1.3.3.E. Total Uranium depletion for MASLWR and PWR Fuel
87
4-1.3.3. Plutonium production and accumulation in fuel
a reactor with fresh fuel, the plutonium concentration is initially zero;
however plutonium could be produced from fertile U-238 through the chain of nuclear
reacti
In
ons shown below.
Figure 4-1.3.4. Uranium 238 reactions, plutonium production
U23892 U239
92
Pu24094 Pu241
94Pu23994 Pu242
94 Pu24394
Np23993
Am24195 Am243
95
β 23.5 m
2.70 b
β 2.35 d
, Fissionn10742.5 b
268.8 b 289.5 b 368 b 18.5 b
, Fissionn101009 b
β 13.2 y β 4.98 h
The system of equations describing these reactions are written as:
( )
( )
( )
( )
⎪⎪⎪⎪⎪⎪⎪⎪
⎩
⎪
−Φ+=
+Φ−=
∫
∫∞
∞
...
)()()(
)()()(
)(
2402400
239239
240
2392390
239239
239
239
239239
tNdEEEtNdt
dN
tNdEEEtNdt
dN
dN
tN
PuPuPucPu
Pu
NpNpPuaPu
Pu
Np
UU
λσ
λσ
λ
and so on down to Americium
⎪⎪⎪
⎨−= )()( 239239239239
0
tNtNdt
dt
NpNpUU λλ
⎪⎪⎪⎪⎧
−Φ=
Φ−=
∫
∫∞
∞
)()(
)()(
238238
239
0
238238
238
dEEEtNdN
dEEEtNdt
dN
UcU
U
UaU
U
σ
σ
88
The plutonium concentration depends on the total neutron flux and neutron
spectrum. For Pu-239, Pu-240 and Pu-241 some equilibrium
(production=consumption) concentrations may be achieved at different points in time,
and values of total fluence. The isotope of plutonium with lower atomic weights may
achieve equilibrium earlier and the isotopes with greater atomic weights may not
achieve these concentrations during LWR cycle length (in particular Pu-242, Am-241
and Am-243 have very small absorption cross sections, and during the LWR cycle
they will be accumulating in the fuel).
The Pu-239, Pu-240, Pu-241 and total plutonium weight fractions were
calculated for MASLWR and PWR operational conditions, and the enrichment levels
specified above. The three series of graphs are presented in current dissertation: 1)
fissile plutonium (Pu-239 and Pu-241) weight fraction versus burnup, 2) total
pluto atio
of the fissile isotopes weight fraction to t e total plutonium weight fraction) versus
burn-up
hat of a PWR core with the same enrichment. It is
caused
where we can see that the plutonium
production rate for 4.5% enriched fuel is greater than for the fuel with 8.0%
enrichment.
nium (all isotopes) weight fraction versus burnup and 3) plutonium quality (r
h
. The first curve illustrates an inventory of the fissile isotopes available in the
fuel for fission. The first and second curve illustrates the accumulation of the
plutonium isotopes in the fuel (important quantity for the fuel reprocessing). The third
curves illustrate the level of attractiveness of the fuel for the potential diversion and
use for the creation of nuclear weapons (less quality, less attractive).
Figure 4-1.3.5 shows that the production of the fissile plutonium (239 and 241)
in the MASLWR core is less than t
by the reduced power density and reduced neutron flux. The balance
concentrations of the fissile plutonium for the fuel with greater enrichment are higher
in both the MASLWR and PWR. This is caused by neutron energy spectrum
hardening, and the increased fission of U-235 atoms, which means that in terms of the
burn-up units (MWD/kgHM), the fuel with higher initial enrichment will have more
fissions of U-235 at the same burnup, than the fuel with lower enrichment. These
effects are illustrated in Figure 4-1.3.6,
89
Figure 4-1.3.5. Production of fissile Plutonium in the Fuel
90
Figure 4-1.3.6. Fissile Plutonium production rates at the small burnup levels
Another interesting period of the fuel campaign is the end of cycle, which may
be caused either by sub criticality or exceeding burnup limits. Figure 4-1.3.7 shows
that at the end of cycle, at the moment when k∞ = 1.0 (see also table 4-1.2.1 for the
burnup at k∞ = 1.0 conditions) both MASLWR and PWR fuel reach balanced fissile
plutonium concentrations. At this point, the production and depletion rates of Pu-239
are balanced, and Pu-241 is close to this balance. Small deviations caused by changes
of spectrum (softening) and neutron flux (increase) at the end of cycle, require that a
constant power density be maintained.
91
Figure 4-1.3.7. Fissile Plutonium concentrations and fuel burnup limits
The principal difference between conventional enriched fuel and increased
enrichment fuel is burn-up limits. The burnup limit of 60 MWD/kgHM allows 4.5%
enriched fuel extended operation at 50% power, which can be achieved with a partial
refueling approach. During this extension period, the fissile plutonium weight fraction
will almost remain the same. The 8.0% enriched fuel will reach burnup limits before
the balance concentration of plutonium will be established, and before a k∞ = 1.0 will
be achieved. This principal difference should be taken into account for the design of
the MASLWR reactor with off-site refueling option, especially for the burnup credit
and shutdown and cool down margin calculations.
Finally, we conclude that in the burnup region of 45-55 MWD/kgHM, which is
a reasonable burnup region for an operational reactor (corresponding to the MASLWR
core average burn-up requirement), fissile plutonium concentration in 8.0% enriched
MASLWR fuel is almost equivalent to the fissile plutonium concentration of 4.5%
enriched fuel depleted in PWR operational conditions.
92
Figure 4-1.3.8. Total Plutonium productions in MASLWR and PWR fuel
A: Weight fraction of the all (total) Pu isotopres accumulated in (MASLWR)
B: Weight fraction of the all (total) Pu isotopres accumulated in (PWR)
The total plutonium production is also important for the evaluation of nuclear
fuel. Because Pu-242 requires a whole chain of nuclear reactions for its production,
the balance concentration of this isotope is not achievable during the fuel campaign.
As a result, the total plutonium concentration curve will be an increasing function of
93
time (see Fig 4-1.3.8). The total plutonium concentration is an important parameter for
ides to be
produc
Figure 4 details)
the fuel reprocessing facility. It indicates the potential amount of minor actin
ed during fuel reprocessing.
PWR operational conditions lead to greater total plutonium concentrations at
the same level of burnup for the same fuel enrichment. At the middle of the operating
cycle, the total concentration of plutonium in 8% enriched MASLWR fuel is always
lower than in 4.5% enriched fuel depleted in PWR conditions. Another important
observation is that the total amount of plutonium in 8.0% enriched MASLWR fuel at a
burnup of 60 MWD/kgHM is equal to the total amount of plutonium in 4.5% enriched
PWR fuel at a burnup of 52 MWD/kgHM.
-1.3.9. Total Plutonium weight fraction for MASLWR and PWR (
94
A: Plutonium quality during the depletion cycle (MASLWR)
Figure 4-1.3.10. Plutonium qualities in MASLWR and PWR fuel
B: Plutonium quality during the depletion cycle (PWR)
95
Knowing the total plutonium and fissile plutonium concentration, we can
h
increas
fuels with greater enrichment is higher; however Figure 4-
1.3.11 shows that for the same enrichment levels, the difference between PWR and
MASLWR fuels become y v lity of
the plutonium produced in MASLWR is sligh than the lutonium
produced in the fuel with the same initial enrichment in a PWR core.
calculate plutonium quality (see Fig 4-1.3.10). Plutonium quality decreases wit
ing burnup, due to production of non-fissile plutonium isotopes. This graph
also shows that gadolinium burnable absorbers do not have a significant impact on
plutonium quality (in standard arrangements and concentrations of Gd). The quality of
plutonium produced in
s significant onl at high burnup le els. Also, the qua
tly smaller q puality of
Figure 4-1.3.11. Plutonium quality comparison for MASLWR and PWR
96
In a comp th partial on-site
refueling, we observe that 4.5% enriched PWR fuel, irradiated above level of kinf=1 up
to burnup of 60 MWD/kgHM has a smaller plutonium quality than 8.0% enriched fuel
irradiated in MASLWR operational conditions. However if we consider EOC criteria
as kinf=1, than 8.0% enriched fuel and 4.5% enriched fuel in MASLWR and PWR
operational conditions will have similar plutonium qualities. That means that
MASLWR with full off-site core refueling or MASLWR with on-site full core
refueling options will have the same quality of plutonium in the range of 0.71-0.73 for
8.0% and 4.5% fuel enrichment (almost insensitive to the enrichment). This is
important for the transportation criticality evaluation and for further fuel reprocessing
conditions.
A summary of the plutonium weight fractions (plutonium vector) for 4.5%
PWR and 8.0% MASLWR fuel is given in Figures 4-1.3.12 and 4-1.3.13.
Figure 4- WR fuel
arison of off-site refuelable MASLWR and PWR wi
1.3.12. Plutonium weight fractions for 4.5 % enriched P
97
Figure 4-1.3.13. Plutonium weight fractions for 8.0 % enriched MASLWR Fuel
These two previous figures show that MASLWR fuel at the EOC has a slightly
greater weight fraction of Pu-239. Concentrations of Pu-240, Pu-241 and Pu-242 are
greater in the PWR fuel. The concentration of Pu-238 is almost the same in both fuel
types.
98
4-1.3.4 Spent Fuel Isotopic composition at 60 MWD/kgHM
performed for each burnup step. However, the detailed discussion of reaction rates and
specific of production and accumulation of each isotope is out of the scope of this
dissertation. Knowledge of the isotopic composition of the spent nuclear fuel is
important for the evaluation of the nuclear fuel cycle back-end [436-444]. Here we
provide a comparison of the isotopic compositions for MASLWR and PWR fuel with
initial enrichments of 4.5% and 8.0%. A comparison table is presented for fuel burnup
of 60 MWD/kgHM at the full power conditions. Cool down effects and effects of the
radioactive decay of the short-lived isotopes after irradiation are not covered.
Table 4-1.3.1. The isotopic composition of PWR and MASLWR spent nuclear fuel at 60 MWD/kgHM, atomic number densities
(Minor Actinides) FA averaged number densities of burnable nuclides (atoms / cm3)
A detailed study of the isotopic composition of the irradiated fuel was
Minor Actinides MASLWR Fuel PWR Fuel
Isotope ID 4.50% 8.00% 4.50% 8.00% U-234 3.43E+18 7.90E+18 3.34E+18 7.60E+18U-235 1.18E+20 5.68E+20 1.43E+20 5.92E+20U-236 1.40E+20 2.22E+20 1.38E+20 2.21E+20U-238 2.02E+22 1.97E+22 2.00E+22 1.95E+22U-239 1.20E+16 8.30E+15 1.62E+16 1.18E+16Np-237 1.92E+19 2.18E+19 2.05E+19 2.36E+19Np-239 1.74E+18 1.20E+18 2.34E+18 1.71E+18Pu-238 1.02E+19 8.57E+18 1.11E+19 9.51E+18Pu 1.84E+20-239 1.26E+20 1.64E+20 1.43E+20 Pu 5.56E+19-240 6.15E+19 5.23E+19 6.49E+19 Pu-241 3.88E+19 3.90E+1 E+19 4.46E+199 4.52Pu-242 2.42E+19 1.15E+1 E+19 1.20E+199 2.42Am-241 1.63E+18 2.15E+18 .50E+18 1.85E+181Am-242 2.38E+16 3.62E+16 E+16 3.44E+162.44Am-243 6.71E+18 2.58E+18 E+18 2.91E+187.12Cm-242 7.17E+17 4.96E+17 E+17 5.12E+177.42Cm-243 2.21E+16 1.22E+16 2.39E+16 1.34E+16Cm-244 3.21E+18 7.89E+17 3.51E+18 9.46E+17Cm-245 1.87E+17 4.56E+16 2.44E+17 6.23E+16Cm-246 3.19E+16 3.36E+15 3.45E+16 4.09E+15
99
MASLWR spent nuclear fuel at 60 MWD/kgHM, atomic number densities Table 4-1.3.2 The Isotopic Composition of PWR and
(Fission products and other isotopes)
FA averaged number densities of burnable nuclides (atoms / cm3) Fission products and other Isotopes
MASLWR Fuel PWR Fuel Isotope ID 4.50% 8.00% 4.50% 8.00%
Kr-83 3.71E+18 4.99E+18 3.77E+18 4.92E+18 Rh-103 2.97E+19 3.22E+19 3.02E+19 3.21E+19 Rh-105 5.69E+16 4.21E+16 7.38E+16 5.68E+16 Ag-109 5.03E+18 3.42E+18 5.06E+18 3.54E+18 I-135 1.77E+16 1.76E+16 2.35E+16 2.34E+16 Xe-131 2.40E+19 2.78E+19 2.36E+19 2.70E+19 Xe-135 6.55E+15 1.13E+16 8.59E+15 1.43E+16 Cs-133 7.10E+19 7.71E+19 7.00E+19 7.56E+19 Cs-134 1.10E+19 8.58E+18 1.24E+19 9.94E+18 Cs-135 2.99E+19 4.72E+19 2.73E+19 4.36E+19 Cs-137 8.32E+19 8.26E+19 8.40E+19 8.34E+19 Ba-140 7.38E+17 7.64E+17 9.77E+17 1.01E+18 La-140 9.69E+16 1.00E+17 1.28E+17 1.32E+17 Nd-143 4.38E+19 5.93E+19 4.63E+19 5.98E+19 Nd-145 3.86E+19 4.35E+19 3.80E+19 4.26E+19 Pm-147 6.45E+18 8.32E+18 6.91E+18 8.83E+18 Pm-148 3.52E+16 3.23E+16 4.65E+16 4.48E+16 Pm-149 4.45E+16 3.84E+16 6.24E+16 5.51E+16 Pm148M 5.77E+16 7.77E+16 7.15E+16 9.55E+16 Sm-147 5.44E+18 7.13E+18 4.19E+18 5.47E+18 Sm-149 8.19E+16 1.61E+17 1.10E+17 2.08E+17 Sm-150 1.76E+19 1.71E+19 1.87E+19 1.80E+19 Sm-151 5.88E+17 9.14E+17 7.55E+17 1.12E+18 Sm-152 5.62E+18 5.59E+18 5.52E+18 5.46E+18 Eu-153 7.18E+18 6.65E+18 7.37E+18 6.74E+18 Eu-154 2.26E+18 2.09E+18 2.50E+18 2.30E+18 Eu-155 1.28E+18 1.13E+18 1.47E+18 1.27E+18 Eu-156 2.47E+17 1.04E+17 3.06E+17 1.39E+17 Gd155F 7.63E+15 1.59E+16 8.93E+15 1.70E+16 LFP1 1.84E+21 1.81E+21 1.83E+21 1.80E+21 LFP2 3.81E+20 3.93E+20 3.78E+20 3.89E+20* LFP – Lumped Fission Products
100
4-1.4 Neutron flux and energy spectra for PWR and MASLWR
The differences in isotopic compositions discussed above can be explained
through the analysis of the neutron flux and neutron spectra. For the purposes of this
study, we consider a two-group model, implemented in MASLWR. The group
boundaries are follows:
Group Number Referenced as: Lower Boundary Upper Boundary Croup 2 “Thermal” 0.0 eV 0.625 eV Group 1 “Fast” 0.625 eV 10 MeV
In order to analyze the difference between MASLWR and PWR fuel we first
consider values of the total, fast and thermal neutron fluxes, required to maintain a
constant power density specified as 40 W/g(HM) for the PWR and 30 W/g(HM) for
the MASLWR. The neutron fluxes determine reaction rates, and they are the key to
understanding the accumulation and depletion mechanisms.
We also consider the change over time in the relative neutron energy spectrum.
For a two-group model, we will discuss the fraction of the thermal flux as a measure
of spectrum hardening. As long as fissile nuclides are the primary absorbers of thermal
neutrons, the spectrum plots will be increasing functions of the fissile isotope weight
fraction, with iso-burnup lines.
4-1.4.1 Neutron Flux
For all test cases of MASLWR and PWR fuel, we have calculated the fast,
thermal and total neutron. The results for the total flux calculation are presented in
Figures 4-1.4.1. A,B and C
101
Figure 4-1.4.1.A. Total neutron flux in PWR fuel
Figure 4-1.4.1.B. Total neutron flux in MASLWR fuel
102
Figure 4-1.4.1.C. Comparison of total neutron flux in MASLWR and PWR
Because MASLWR has a lower power density, lower values of the total
neutron flux are required to maintain the required power density. For fuels with higher
initial enrichment, a lower flux is required for the same specified power density. The
addition of burnable absorbers requires an increased neutron flux at BOC in order to
maintain the power density and compensate for the increased parasitic absorption.
The fast and thermal flux comparison for MASLWR and PWR is presented in
Figure 4-1.4.2 A and B. The overall flux decrease in MASLWR due to lower power
density is still present. The solid lines represent fuel without burnable poisons and the
dashed lines represent fuel with burnable poisons.
103
Figure 4-1.4.2.A. Comparison of Fast Neutron Flux in MASLWR and PWR
Figure 4-1.4.2.B. Comparison of Thermal Neutron Flux in MASLWR and PWR
104
en PWR and MASLWR is ¾, the
flux ra
The comparison of the fast, thermal and total neutron fluxes is summarized in
Table 4-1.4.1. While the power density ratio betwe
tios for the same initial fuel enrichment are different. This is caused by the
different neutron energy spectrum in PWR and in MASLWR.
Table 4-1.4.1. MASLWR and PWR Flux comparison summary
Parameter 4.5% Fuel 8.0% Fuel Power Density Ratio MASLWR / PWR 0.75 0.75 Total Flux Ratio MASLWR/ PWR 0.68-0.72 0.68-0.70 Fast Flux Ratio MASLWR/ PWR 0.67-0.69 0.66-0.67 Thermal Flux Ratio MASLWR/ PWR 0.75-0.93 0.76-0.88
4-1.4.2 Neutron energy spectrum transformations
Two competing effects cause the differences between the MASLWR and
conventional PWR neutron energy spectrum. The spectrum is softened due to lower
operational parameters, and resulting increased moderator density and enhanced
moderation. The spectrum is hardened because of the increased fuel enrichment and
the resu increased absorption of the thermal neutrons.
Figures 4-1.4.3 A and B show the spectral change with increasing initial fuel
enrichment under PWR and MASLWR operational conditions. The fuel used in these
calculations does not have burnable absorbers, control rods are out and there is no
soluble boron in the coolant. These figures show that the neutron energy spectrum is
hardening with increased initial enrichment. The initial fraction of thermal neutrons
for the same initial enrichment is higher for MASLWR operational conditions. The
shapes of the curves are most strikingly different in that the fraction of thermal
neutrons has a maximum value in the EOC for PWR operational conditions, and does
not exceed this value for standard enrichments (3.5-4.5%). The MASLWR spectrum
curves continue to increase with increasing burnup.
lting
105
Figure 4-1.4.3.A. Fraction of the thermal flux in PWR fuel
106
Figure 4-1.4.3.B. Fraction of the thermal flux in MASLWR fuel
107
Figure 4-1.4.3.C. Fraction of the thermal flux in MASLWR fuel
108
Figure 4-1.4.3.D. Fraction of the thermal flux in PWR fuel
109
In order to understand the operational spectrum of the MASLWR and PWR
were produced.
ntaining burnable absorbers and soluble boron in the coolant. The
PWR w
range than a conventional PWR.
Fur r
should
harden
reactors, several additional curves
The effect of burnable absorbers and soluble boron is represented by three
curves in Figure 4-1.4.3 C and D. The blue rectangular area represents the expected
spectral values for a conventional PWR loaded with conventional fuel containing
burnable absorbers and soluble boron in the coolant. The red rectangular region
represents the expected spectral values for a MASLWR core loaded with increased
enrichment fuel co
as built according to data shown in Figure 4-1.4.3.C, and copied to Figure 4-
1.4.3.D, and the MASLWR area was built according to MASLWR operational data
shown in Figure 4-1.4.3.D, and then copied back to Figure 4-1.4.3.C.
From these plots, we conclude that the MASLWR core with 8.0% enriched
fuel will have a harder spectrum in epithermal energy
the more, if some PWR utilities decide to switch to increased enrichment fuel, they
expect a significant decrease of the thermal flux fraction, and spectrum
ing.
110
layed
The prompt neutron lifetime and the effective fraction of delayed neutrons are
important reactor kinetics parameters, characterizing the sensitivity of the reactor
These pa a fun fuel
y spect nt fi rtile
t yields (em yed n the
fraction of delayed neutrons will depend on the fission of each of the isotopes. Also
delayed
rons best characterizes the behavior of the
reactor
4-1.5 Prompt neutron lifetime and Effective fraction of deNeutrons
power to reactivity insertions [482, 487]. rameters are ction of the
isotopic composition and neutron energ rum. Differe ssile and fe
isotopes have different fission produc itters of dela eutrons), so
neutrons are born with a lower energy than prompt fission neutrons, so they
have a lower probability to cause fast fission or leak in or out of the problem. [234]
The effective fraction of the delayed neut
. Prompt neutron lifetime is also important and depends on the neutron spectra
and the absorption and fission reactions rates and cross sections.
The results of the prompt neutron lifetime calculation are given in
figure 4-1.5.1 A and B, and summarized in Figure 4-1.5.1 C
111
Figure 4-1.5.1.A Prompt neutron lifetime (PWR)
112
Figure 4-1.5.1.B. P time (MASLWR) rompt neutron life
113
Figure 4-1.5.1.C. Promp (PWR vs. MASLWR) t neutron lifetime
These last three figures show that 8.0% enriched fuel leads to a smaller prompt
neutron lifetime in both the MASLWR and the conventional PWR. For the same
initial enrichment, MASLWR operational conditions lead to a greater prompt neutron
lifetime at the EOC than in a PWR. A comparison of a conventional PWR and the
proposed MASLWR fuel with increased enrichment shows that the MASLWR core is
expected to prompt neutron lifetime that is smaller than that of a conventional PWR
core by a factor of 1.5 to 2.
The results for the calculations of the effective fraction of the delayed neutrons
are given in Figures 4-1.5.2. A and B for the PWR and MASLWR respectively.
114
Figure 4-1.5.2.A. Effective Fraction of Delayed Neutrons (MASLWR)
Figure 4-1.5.2.B. Effective Fraction of Delayed Neutrons (PWR)
115
These last two graphs sho e fraction of delayed neutrons is
greater in the MASLWR fuel than in PWR ame enrichment. The most
important difference occurs when ith different initial enrichments.
The 8.0% enriched fuel has delayed neutron fraction that decreases more slowly with
burnup than conventional PWR fuel; at the MASLWR EOC (Burnup = 60
MWD/kgHM) the effective fraction of delayed neutrons is equivalent to a
conventional PWR core at burnup 30-40 MWD/kgHM. This means that “deep
burnup” of 8.0% MASLWR fuel does not cause a significant decrease of delayed
neutrons, which is a concern for deep burnup for conventional PWR fuel.
w that the effectiv
fuel with the s
we consider fuels w
116
tures. This requirement is reflected
in US
or pressure changes with different concentrations of soluble
boron (if any).
• The reactivity effect of fuel temperature, assuming constant power density and
isotopic composition.
• Combined or integral reactivity effects caused by power changes (power
defect) with relative changes of the isotopic composition and fuel and
moderator temperature, start-up and transient conditions.
In order to evaluate the reactivity effects and reactivity coefficient changes in
MASLWR fuel caused by increased enrichment and different operational conditions
(lower power density, lower fuel and moderator temperature and lower pressure
4-1.6. Reactivity Effects
In the previous section we demonstrated that differences in initial fuel
enrichment and reactor operational conditions lead to significant changes in the fuel
and reactor properties during depletion. The different isotopic composition and
neutron energy spectrum hardening will also affect the reactor kinetics and dynamics.
In particular, reactivity feedback has the greatest importance in the reactor design and
safety evaluation.
The current approach to reactor safety requires having negative reactivity
feedback mechanisms to provide inherent safety fea
regulatory guides [287, 288, 309, 312], IAEA safety standards [44, 46] and
evaluation criteria developed by international and national user groups [69].
Traditionally, reactivity feedback evaluations cover:
• The reactivity effect of moderator temperature, at full power, intermediate
power levels and at shutdown conditions for a constant fuel temperature,
pressure without boiling and with different concentrations of soluble boron (if
any)
• The reactivity effect of moderator density, including void reactivity effects
caused by boiling
117
compared to a PWR), we will c ssemblies out of the set of fuel
assemblies mentioned above at PWR and MASLWR operational conditions.
Table 4-1.6.1. Fuel segments used in a feedback evaluation study
Operational Conditions
onsider four fuel a
Fuel Type MASLWR PWR
4.5% Enriched Fuel Without Gadolinium FM_45_00 FP_45_00 4.5% Enriched Fuel With Gadolinium FM_45_M1 FP_45_M1 8.0% Enriched Fuel Without Gadolinium FM_80_00 FP_80_00 8.0% Enriched Fuel With Gadolinium FM_80_M1 FP_80_M1
The cross comparison of all eight test cases provides the basis for an evaluation
of the effect of the initial enrichment and operational conditions on the reactivity
coefficients for PWR and MASLWR fuel.
It is also important to mention that all test cases were done with CASMO-4 for
fuel
the multiplication
factor
a assembly, in two-dimensional geometry with reflective boundary conditions.
The reactivity coefficient calculations were performed in the following manner: first,
the initial depletion of the fuel assembly at constant operational conditions and
constant power density (see the table above for the operational conditions description)
and constant soluble boron concentration of 800 ppm were performed, in order to get
information about fuel isotopic composition and effective fraction of delayed neutrons
at each burn-up step. Next, a series of branch-case calculations of
were performed for each burn-up step. Finally, calculations of the reactivity
effects were performed as a difference between the multiplication factor at two state
points.
118
4-1.6.
1. Moderator temperature reactivity coefficient:
The moderator temperature reactivity coefficients (MTC) for hot full power
conditions (HFP) were calculated for two soluble boron concentrations of 0.1 ppm and
800 ppm
and MASLWR at the boron concentration of es slightly
positive for 4.5% enrich
or boron free, we may conclude that it
is unlik
rences in the behavior of
the MTC for MASLW
and are presented in figure 4-1.6.1.A.
This figure shows that the MTC at HFP conditions is negative for both PWR
0.1 ppm; however it becom
ed fuel under MASLWR operational conditions after a burnup
level of 55 MWD/kgHM. Assuming that at this burn up level the multiplication factor
of the fuel will be below 0.95, and that at this burn-up level the reactor will probably
be operated with smaller boron concentrations
ley that the reactor will be operated under these operating conditions, but they
should be discussed and addressed in safety analysis report.
A more detailed examination of these plots shows that increasing the initial
fuel enrichment decreases the magnitude of the MTC for fuel under PWR operational
conditions at a boron concentration of 0.1 ppm, and has no significant effect for PWR
fuel at a boron concentration of 800 ppm. The increase of initial enrichment from
4.5% up to 8.0% for fuel under MASLWR operational conditions does not cause a
significant difference in the MTC for a boron concentration of 0.1 ppm up to a burnup
level of approximately 50-55 MWD/kgHM. However, for a boron concentration of
800 ppm, the absolute value of the MTC for 8.0% enriched fuel is initially twice as
large as that for 4.5% fuel, and the curve for 8.0 % enriched fuel lies below the curve
for 4.5% fuel at all burn up levels. These fundamental diffe
R and PWR are caused by the different HFP conditions of both
reactors and the resulting increased moderation in the MALSWR reactor compared to
a PWR.
119
Figure 4-1.6.1. A
Figure 4-1.6.1. B
120
ture and reactor power oscillations requires additional study and evaluation.
The comparison of MASLWR and PWR fuel at different boron concentrations
is shown in figure 4-1.6.1 B. The HFP MTC for MASLWR is generally 2-3 times
lower than HFP MTC for a PWR reactor. This fact creates a potential challenge for the
safety justification at Hot Full Power conditions, but gives additional margins against
secondary criticality during reactor cool-down. The influence of the lower magnitude
of the MTC at the HFP conditions in a reactor with natural circulation on coolant flow,
tempera
Figure 4-1.6.1.C.
*Solid lines represent boron concentration of 0.1 ppm
Dashed lines represent boron concentration of 800 ppm
121
4-1.6.2 Fuel temperature reactivity coefficient
he fuel temperature reactivity coefficient generally represents the importance
of reso
e fixed at 493ºK, which is close to
the operational temperature (of which system?). These calculations were necessary to
elim nce
caused by the different irradiation history of the fuel and the resulting different
isotopic compositions.
The graphs below show that an increase in initial fuel enrichment decreases the
magnitude of the FTC for both MASLWR and PWR. This is caused by the decrease
in resonance absorption in U-238 and a slight increase in U-235 fissions. The
magnitude of the FTC grows with depletion.
The difference between the MASLWR and PWR FTC at HFP conditions is
caused by operational conditions and the amount of moderator available. In the
MASLWR, the probability of avoiding resonance absorption is greater than in the
PWR (for the same initial enrichment). The second graph illustrates that at the same
moderator conditions, the difference in FTC is caused by fuel isotopic composition
and irradiation history. We may see the increase of FTC at the BOC caused by the
presence of burnable absorber (gadolinium). At the EOC, the difference between the
curves for fuel with the same initial enrichment is caused by the different isotopic
compositions of fuel irradiated in MASLWR and PWR operational conditions.
T
nance absorption of the neutrons in uranium 238 and the Doppler-broadening of
the resonances with the increase of the temperature. The evaluation of fuel
temperature reactivity coefficients was performed in two series. Both series were done
for fuel temperatures in the range between 500 ºK and 1500 ºK.
The first series of calculations (figure 4-1.6.2 A) were performed for HFP
conditions for the MASLWR and PWR respectfully. These conditions are different
and the PWR has higher fuel and moderator temperatures. It allows us to compare the
magnitude of the FTC at real operational conditions for the MASLWR and PWR.
The second series of calculations (figure 4-1.6.2 B) were performed for full
power conditions, but for the moderator temperatur
inate the difference in operational conditions and evaluate the FTC differe
122
Figure 4-1.6.2. A
Figure 4-1.6.2. B
123
-1.6.3. Void Reactivity Coefficient4
oid
reactiv
the moderator temperature is equal
to the
maller reactivity insertion from moderator density changes at the MOC and
EOC. The depletion history and concentration of the fissile isotopes and fission
products in the irradiated fuel are also important. For example, both MASLWR and
and
MASL
ation of 800 ppm is 20-30% smaller than for a
boron-free moderator. The importance of the neutron absorption moderator in this case
is grea r for the fuels with lower initial enrichment, so the difference of the void
The moderator density reactivity coefficient, and in particular, the v
ity coefficient at boiling conditions, is one of the most important reactor safety
parameters. The evaluation of the void reactivity coefficient is required according to
the NRC regulatory guides [287, 288, 300, 309] and IAEA safety standards [44, 46].
The methodology for the evaluation of the void reactivity coefficient is given in [1,
236, 490]. The standard set of branch cases in CASMO 4 specified by card “S3C”
provides data to calculate the void reactivity coefficient for the coolant boiling at the
saturation temperature.
The void reactivity coefficient calculations were performed for two soluble
boron concentrations: 0.1 ppm and 800 ppm. The fuel temperature was chosen to be
the average fuel temperature at hot full power, but
saturation condition at the specified pressure (86 atm for MASLWR and 150
atm for PWR): 571ºK and 615ºK respectively. The void coefficient was calculated
through the branch cases for 0% and 40% void fraction.
An increase in initial fuel enrichment up to 8.0% leads to a greater U-235
weight fraction than fuel with an initial 4.5% enrichment at the same burnup level. So
the absorption in the fuel becomes more important, and the density of the moderator
leads to a s
PWR 4.5% enriched fuel has smaller differences at the EOC, than PWR
WR fuel with an initial enrichment of 8.0%.
The use of soluble boron affects the void reactivity coefficient. Boiling in the
moderator decreases moderator density (negative feedback), but also decreases the
amount of soluble absorber (positive feedback). As a result, the void reactivity
coefficient for a soluble boron concentr
te
124
124
reactivity coefficients for the enrichment 4.5% irradiated in
MASLWR and PWR operation conditions, is greater than for the fuel with initial
enrichment of 8.0%.
Figure 4-1.6.3. A
fuel with initial
125
Figure 4-1.6.3. B
A comparison of the graphs of the void reactivity coefficient for a boron-free
moderator and a moderator with 800 ppm of boron shows that, for 4.5% enriched fuel,
the voi
core with 800 ppm of boron. For 8.0%
enriched fuel, the void reactivity coefficient for MASLWR is lower than for PWR for
boron free conditions, however f ith 800 ppm of boron, the void
reactivity coefficients for MASLWR and P R is almost the same.
MASLWR fuel is shown below.
d reactivity coefficient for MALSWR is slightly lower than for the PWR in a
boron free-core, and significantly higher for the
or the moderator w
W
Finally, if we compare void reactivity coefficients at different burnup levels
relative to the criteria of K∞=1, we find that at this point the void reactivity
coefficients for MASLWR and for PWR operational conditions are almost the same
for both initial fuel enrichments (see fig. 4-1.6.3. C).
A comparison of the void reactivity coefficients for standard 4.5% PWR fuel,
and for 8.0%
126
Figure 4-1.6.3. C
4-1.6.4. Power Defect
The integral effect of the different reactivity mechanisms can be evaluated
through the power reactivity defect: the reactivity insertion during a transient from
state A to state B. In particular, for the purposes of the current dissertation, we
consider three important states of reactor operation:
1) CZP – Cold, Zero Power or, Cold shutdown case
2) HZP – Hot, Zero Power or, Reactor Startup conditions, Hot standby
3) HFP – Hot, Full Power, or Normal operational conditions.
We also consider two different HFP conditions: when the moderator
temperature is equal to the core inlet coolant temperature (marked as HFP(S)) and
when the moderator temperature is equal to the core average coolant temperature
(marked as HFP (A)).
raditionally, startup of a PWR is split into several stages. First, the reactor is
heated from Cold Shutdown conditions up to Host Standby conditions with the
T
127
external source, while the reacti pensated by the reactor scram
system. The second stage is the removal of e reactor scram rods, and removal of the
soluble
power range. The third step is po removal of the soluble boron or
control rod group at constant temperature and pressure parameters. Calculations of all
reactivity defects were performed for a soluble boron concentration of 800 ppm, and
“control rod out” conditions, in order to evaluate these defects without biases caused
by the presence of control rods. The purpose of this study is to provide comparisons of
the different fuel options for MASLWR and PWR. The data for the complete
MASLWR core will be provided below in chapter 4-4.
The reactivity difference between CZP and HZP conditions for each depletion
step is the Zero Power Heating Reactivity Defect (ZPHRD). This parameter combines
the effects of fuel temperature, moderator temperature and moderator density on
reactivity. The values of the Zero Power Heating Reactivity Defect are presented in
figure 4-1.6.4.
The lower operational parameters of MASLWR (pressure, temperature) lead to
lower absolute values of the ZPHRD compared to conventional PWR operational
conditions. During this operational stage, the moderator temperature reactivity effects
’s are
negativ
vity excess is com
th
boron down to the startup concentration to achieve criticality in the source
wer increase, with
and soluble boron density effects dominate. At the beginning of cycle, ZPHRD
e for all cases; however at the EOC, positive values occur for the fuel with an
initial enrichment of 4.5%. On first examination, this appears to be significantly
problematic; however the EOC concentration of boron is usually smaller than 800
ppm (so the MTC is more negative), and the multiplication factor K∞=1 at 40
MWD/kgHM for PWR operation conditions. For the MASLWR operation conditions
K∞=1 at 42 MWD/kgHM, and the Zero Power Heating Reactivity Defect is slightly
positive. This issue should be addressed in the safety analysis if the reactor operation
with multiple batch refueling and 4.5% enriched fuel is considered for MASLWR.
128
Figure 4-1.6.4. A
For the fuel with initia .0%, the Zero Power Heating
Reactivity Defect is always negative. The absolute values of ZPHRD under
MASL
n a PWR, however, could lead to greater zero power heating
reactivity defects, which could cause potential reactivity issues during reactor cool
down.
The use of burnable poisons has a significant effect on the ZPHRD at the
beginning of the cycle, however after burnup of 10-15 MWD/kgHM the deviations
caused by burnable absorbers are almost negligible. At the BOC, the same amount of
burnable absorber can cause greater deviations for the fuels with lower initial
enrichment.
The second set of reactivity defect calculations was performed for the HZP and
HFP(S) conditions. For both cases, the moderator temperature and pressure and
l enrichment of 8
WR operational conditions are always 1.5–2.5 times smaller than those for
PWR operational conditions, and similar to 4.5% fuel in the BOC and MOC. The use
of 8.0% enriched fuel i
129
soluble boron concentration were fuel temperature is increased to
the full power fuel temperature. The difference in the multiplication factor is caused
by the different power density and relative fuel temperature increase, and changes in
concentrations of the fission products and their daughters (particularly Xe and Sm).
This change in reactivity is called the power reactivity defect.
The result of the calculations is shown in figure 4-1.6.4. B. The MASWLR
operational conditions involved a lower power density than a standard PWR, so for the
same initial enrichment, the absolute value of the power reactivity defect is smaller in
MASLWR. Also, with an increase in initial enrichment, the contribution of the fission
products to overall neutron absorption is relatively small, compared to the fraction of
neutrons absorbed by U-235. This means that the magnitude of the power defect is
smaller for fuels with higher enrichment. The quantity of fissile isotopes is decreasing
with depletion, so the magnitude of the power reactivity defect is increasing towards
the EOC.
Figure 4-1.6.4. B
the same, and the
130
and PWR fuel,
we may
e 4-1.6.4. C Relative power reactivity defect (linear approximation)
In order to make another fair comparison between MASLWR
calculate the power reactivity defect per unit of power density. It is roughly a
linear function of the power reactivity coefficient in the range between zero power and
full power. The results of such calculations are presented in the figure 4-1-.6.4. C.
According to these calculations the power reactivity effect per unit of the
power density is stronger in MASLWR operational conditions for both 4.5% and 8.0%
enriched fuel. This is important information, however for the detailed comparison and
licensing purposes more detailed comparison will be necessary, for a different power
levels, and detailed investigation of the fuel temperature profiles and distributions for
intermediate power levels in MASLWR and PWR.
Figur
The total power reactivity defect was calculated as the difference in reactivity
between cold zero power conditions (CZP) and hot full power conditions, with a
131
moderator temperature equal to th erator temperature. The results of
these calculations are presented at the figure 4-1.6.1. D.
For all test cases, the total reactivity defect is negative. For the MASLWR
operational conditions, the magnitude of the reactivity defect is usually smaller than
for PWR conditions with the same initial fuel enrichment. For the 4.5% initially
enriched fuel, the magnitude of the total reactivity defect increases from the beginning
of cycle up to a burnup of 35-40 MWD/kgHM and then decreases. For the 8.0%
initially enriched fuel, the magnitude of the total reactivity defect increases at all
points. The lower magnitude of the total reactivity defect for the MASLWR
operational conditions means a smaller excess reactivity during reactor cool-down,
which is good, because the secondary cold criticality margin is larger.
Figure 4-1.6.4. D
e core average mod
The reactivity defects and reactivity coefficients calculated here is a first step
in understanding the effect of operational parameters on MASLWR fuel; these are
132
very important considerations dur The core characteristics might be
slightly different than those considered in this analysis because of multiple fuel types
loaded into the core, three dimensional burnup irregularities, neutron leakage and
reflector effects.
ing reactor design.
133
4-1.7. Control Rod Worth, Shielding and Shadowing Effects
lude control rods with solid neutron absorbers (such as boron
carbide
uel enrichment and isotopic composition.
We wi
For the safe and efficient operation of the nuclear reactor over long periods, a
reliable and diversified reactor control system should be designed capable of providing
reactor scram and shutdown margins at any time during reactor operation and to
compensate burnup reserves of the reactivity at the BOC. Traditionally, reactivity
control mechanisms inc
, hafnium, silver-indium-cadmium or other neutron absorbing materials or
compositions), a neutron absorber dissolved in the coolant (such as boric acid) and
burnable absorbers (such as gadolinium, erbium or boron compositions uniformly
distributed in the fuel, coated on the surfaces of the fuel pellet, or other designs of
removable or non-removable burnable absorbers).
The relative and absolute worth of the reactor control mechanisms and
components are strongly dependent on the f
ll consider the worth of the soluble boron and reactor control rods for the
MASLWR and PWR operational conditions and for fuel with initial enrichments of
4.5% and 8.0%. The results presented in this chapter are from assembly level
calculations, performed with CASMO-4, for a standard Westinghouse 17x17
geometry.
4-1.7.1. Soluble boron worth for the Hot Full Power Conditions
The use of soluble boron distributes a neutron absorber uniformly throughout
the core and enables a fine-tuning of the boron concentration during the reactor
campai
A and B.
gn. The evaluation of the soluble boron was performed for two concentration
intervals: 0.1 - 800 ppm and 800 – 1600 ppm. The results of the calculations are
presented in figures 4-1.7.1.
134
Figure 4-1.7.1. A
Figure 4-1.7.1. B
135
These figures show that he larger initial enrichment, the
relative absorption in the fuel is greater, so the worth of the soluble boron is almost
half as large for 8.0% enriched fuel than for the 4.5% enriched fuel.
The lower operational parameters of MASLWR lead to a larger coolant
density, so for the same relative concentration, the absolute amount of boron is
greater, and the worth of the boron in higher, than in a PWR fuel assembly with the
same initial enrichment.
The comparison of the soluble boron worth in the concentration intervals 0.1 –
800 ppm and 800 – 1600 ppm is presented in figure 4-1.7.1. C. The solid lines
represent the soluble boron worth concentration interval 0.1-800 ppm, and the dashed
lines represents the soluble boron worth for the interval 800 – 1600 ppm. This figure
shows that the boron concentration slightly decreases the relative soluble boron worth
due to a shielding effect.
Figure 4-1.7.1. C
for the fuel with t
136
The evaluation of the solu ding effect for the hot full power
conditions is presented in the figure 4-1.7.1. D
ble boron self-shiel
Figure 4-1.7.1. D
This figure demonstrates that the self-shielding effect is stronger in the fuel
with a
.
lower initial enrichment. The magnitude of the self-shielding effect is also
greater in this softer spectrum. Also, the self-shielding effect magnitude grows with
depletion. This non-linearity in the soluble boron worth curve should be taken into
account when calculating startup concentrations after reactor scram or standby
operation
137
s 4-1.7.2. Soluble boron worth for the Cold zero Power Condition
Soluble boron worth was also calculated for cold zero power conditions for
fuel initial enrichments of 4.5% and 8.0% irradiated at MASLWR and PWR
operational conditions.
The cold zero power conditions are similar for the PWR and MASLWR and
assume room temperature for the fuel and moderator so the density of the moderator
and the amount of boron may be considered identical for the MASLWR and PWR
fuel. The differences in the curves for the soluble boron worth are caused only by the
history of the fuel irradiation and corresponding isotopic fuel composition.
The results of these calculations are presented in figures 4-1.7.2 A (for the
concentration interval 0.1 - 800 ppm) and 4-1.7.2 B (for the concentration interval
800-1600 ppm). The density of the moderator at CZP conditions is greater than for
HFP conditions, and the worth of soluble boron is greater as well.
Figure 4-1.7.2. A
138
Figure 4-1.7.2. B
The evaluation of the self-shielding effect is presented in the figure 4-1.7.2 C.
Figure 4-1.7.2. C
139
4-1.7.3 Soluble boron wort erational conditionsh at different op
Previous calculations showed that the boron worth at cold zero power
conditions is greater than for hot full power conditions. An understanding of the
influence of power and heating on soluble boron worth is important for further reactor
design. We perform a comparison of the soluble boron worth for MASLWR and PWR
fuel with initial enrichments of 4.5% and 8.0% at CZP, HZP and HFP operation
conditions. The results of these simulations are shown in figure 4-1.7.3. A. The bold
lines represent MASLWR operational conditions. The solid lines correspond to hot
full power conditions, the dashed lines represent hot zero power conditions, and the
dotted lines represent cold zero power conditions.
The figure shows that moderator density has a greater effect on the soluble
boron worth than fuel temperature and power associated effects. Moderator density
effects are also associated with operational parameters, such as temperature and
pressure. These parameters are smaller in MASLWR, so the greater coolant density
leads to a greater worth of soluble boron at the hot conditions for the MASLWR fuel
than for PWR fuel with the same initial enrichment.
n increase of MASLWR fuel enrichment up to 8.0%, compared to traditional
PWR fuel, will lead to spectrum hardening and a decrease of the ratio of neutrons
absorbed by boron atoms to the neutrons absorbed in fuel. As a result, the worth of
the soluble boron in MASLWR fuel with initial enrichment of 8.0% is 1.5-1.7 times
smaller than in PWR fuel with an initial enrichment of 4.5%. The difference between
hot and cold conditions is also smaller for the MASLWR fuel. At the MOC (~30
MWD/kgHM) the worth of the soluble boron in the MASLWR fuel with an initial
enrichment of 8.0% is the same as that in the PWR fuel with an initial enrichment of
4.5% at the BOC.
A comparison of the 8.0% MASLWR fuel and 4.5% PWR fuel is given in
figure 4-1.7.3 B.
A
140
Figure 4-1.7.3. A
The current figure is too complicated, so for the better comparison of the
MALSWR fuel with initial enrichment of 8.0 % and standard PWR fuel figure 4-
1.7.3.B is more useful.
141
Figure 4-1.7.3. B
Figure 4-1.7.3. B shows that the soluble boron worth curves of 8.0%
MASLWR fuel and 4.5% PWR fuel overlap; however for the same enrichment levels,
the boron worth in 8.0% MALSWR fuel will always be smaller than in PWR fuel.
Also the lower boron worth at the cold conditions will require greater concentrations
of boron to compensate for the excess reactivity during refueling.
142
4-1.7.4. The soluble boron worth at boiling conditions.
The worth of the boron uniformly distributed in the moderator with a given
concentration is strongly dependent on the moderator density. At boiling, the amount
of boron available for neutron absorption will decrease with a decrease of the
moderator density and increase of the void fraction. For both the PWR and MASLWR,
bulk boiling or sub-cooled boiling are out of the range of normal operational
parameters. Nevertheless, boiling could occur under accident conditions, due to an
uncontrolled power increase, loss of feedwater or depressurization. The soluble boron
worth was calculated for a moderator temperature equal to the saturation temperature
for the given pressure with void fractions equal to 0% and 40%. The results are
presented in figures 4-1.7.4 A and B and, figure 4-1.7.4.C provides a comparison of
the cases.
Figure 4-1.7.4. A
143
Figure 4-1.7.4. B
Figure 4-1.7.4. C
144
hese figures show that the soluble boron worth decreases under boiling
conditions. The relative loss of soluble boron worth due to moderator boiling is shown
relative loss is a function of the operational parameters
(densit
itivity of the soluble
boron
T
in figure 4-1.7.4.D. The
y at saturation temperature for the given pressure) and fuel enrichment. For
MASLWR operational conditions the moderator density is greater for the zero void
fraction, so the small changes in the density result in a larger impact than in the PWR
case. An increase of initial enrichment also increases the sens
worth to the boiling conditions. This effect is also complicated by spectrum
changes due to the decrease of the moderator density, so a more detailed study of the
boron worth at boiling conditions should be performed for the safety analysis of
MASLWR.
Figure 4-1.7.4. D
145
4-1.7.5. The Control Rod Worth
Modern LWRs use control rods to provide active reactor control and shutdown
margins. The design of the control rod components and materials composition is
usually
ine different material combinations were considered: hafnium;
silver-i
feasibility and manufacturability of the chosen enriched boron compositions is out of
the scope of this dissertation, so here we will consider only the influence of the boron
enrichment on the neutronics of the reactor.
We consider 24 control rods per fuel assembly. The control rod tubes are made
of stainless steel and the guide tubes are made of Zircalloy. For all materials, except
enriched boron mixtures, standard CASMO-4 cards were used for the specification of
material properties and thermal expansion coefficients for all cases. Detailed
descriptions of the materials can be found in the CASMO-4 manual [234].
The control rod worth is calculated as the reactivity difference between the
case with fully inserted control rods and the case with fully removed control rods with
the same power density and operational parameters.
All evaluations were performed for the control rods with boron carbide and
natural boron as the neutron absorber composition unless otherwise specified.
The first series of evaluations was performed for the B4C control rods at hot
full power conditions (normal operation conditions). The results of the evaluation are
shown in figure 4-1.7.5.A.
proprietary information; however, some general characteristics available in
publications [5] can be used for the computer simulation of control rods, and
evaluation of the impact of operational parameters and increased enrichment on the
control rod properties.
In this study, n
ndium-cadmium, boron carbide with natural boron, and 6 compositions
simulating boron carbide with enriched boron. The boron enrichments for those
compositions were: 40%, 50%, 60%, 70%, 80% and 90% correspondently. The
146
Figure 4-1.7.5. A
This figure shows that the worth of the control rod increases with depletion.
The increase of initial fuel enrichment leads to a control rod shielding effect due to
increased absorption of thermal neutrons in the fuel. The worth of the control rods is
smaller for both MASLWR and PWR operational conditions.
Comparing the results for fuel with the same initial enrichment but different
operational conditions, we find that MASLWR operational conditions (at full power)
lead to a slight decrease in control rod efficiency. This can be explained by comparing
the resu eutron
ut th rmal neutrons in the
or u e eso
ation is greater due to hi od
uch as in PW era con .
lts for the thermal neutron flux for MASLWR and PWR. The thermal n
flux for the MASLWR is smaller at full power, b e role of the
fission process is greater and the probability f the ne t ron to scape r nance
absorption during moder gher m erator density, so a local
perturbation (local decrease of the thermal neutron flux) around the control does not
affect the overall control rod worth as m R op tional ditions
147
A second set of evaluations was performed for the B4C control rods at cold
zero power conditions (or cold shutdown conditions). The results of the evaluation are
shown in figure 4-1.7.5.B.
Figure 4-1.7.5. B
* Solid lines: FA without burnable poisons,
* Dashed lines: FA wi
he initial inc
th st load p
rease of enrichment up to 8.0% leads to a decrease in the control
rod wo
od worth at the beginning of cycle. The
difference between the two curves at the
ent of 4.5%.
andard BP attern,
T
rth, but the operational conditions (which are similar for the MASLWR and
PWR at CZP) do not affect the control r
end of cycle is caused by the history of
irradiation and isotopic composition of the fuel. We also conclude that for the fuel
with an initial enrichment of 8.0% the effect of the irradiation history is not as
significant as for the fuel with an initial enrichm
A third set of evaluations was performed for the B4C control rods at hot zero
power conditions (or intermediate conditions during startup). The results of this
148
evaluation were combined with the results of the two previous evaluations and are
shown in figures 4-1.7.5.C.
Figure 4-1.7.5. C
149
During reactor startup, as we progress through CZP, HZP and HFP, the control
rod worth increases during reactor heating, and then slightly decreases during the
power increase and rise of the fuel temperature. The PWR has a greater difference
between hot and cold parameters than MASLWR, so the change in control rod worth
between CZP and HZP is greater for the PWR than for MASLWR.
In order to describe separate the effect of different parameters on the control
rod worth, we define two terms. The heating correction of the control rod worth is
equal to the difference between the control rod worth at HZP and CZP. For both the
MASLWR and PWR, the heating correction is less than 0 (increase of the control rod
worth). The power correction of the control rod worth is equal to the difference
between control rod worth at HFP and HZP. For both the MASLWR and PWR, the
power correction is greater than 0 (loss of the control rod worth).
The figures show that the power correction is slightly higher for the MASLWR
operational conditions than for the PWR, while the heating correction is greater for the
PWR.
he correction of the control rod w rth is characterizes the application of the
reactivity effect for the reactor with control rods. This means that during the startup
hould consider the correction of the control rod worth in
combin
hoose boron-free
operation and reactor startup with control rod motion at intermediate power and
heating lev ffects will
be necessary for the intermediate power, eating and poisoning levels in order to
predict reactor response to the control rod movement with acceptable accuracy.
oT
and power transients we s
ation with temperature and power reactivity defects of the fuel described
above.
These effects are important for the calculation of startup or shutdown boron
concentration in order to avoid repeated criticality (or unexpected criticality during the
transient), and is especially important if MASLWR designers c
els. Additional calculations of control rod worth and reactivity e
h
150
ifferent control rod compositions4-1.7.6 Comparison of d
ower
plants. Common control rod absorber materials include boron carbide (B4C; B4C),
silver-indium-cadmium (Ag-In-Cd, AIC), and hafnium (HAF) [234]. Some
publications [5] also mention dysprosium compositions and different composites of
boron, hafnium and dysprosium. However, these materials are available from single
vendors, and detailed specifications are currently unavailable for a variety of reasons.
In this study, we evaluate only the three most common control rod compositions
mentioned above.
The results of the control rod worth calculations for boron carbide control rods
are shown in figure 4-1.7.5.A (above), for silver indium-cadmium in figure 4-1.7.6. A,
and for hafnium in figure 4-1.7.6. B. The combined results are shown in figures 4-
1.7.6 C and D.
These figures show that the worth of the Ag-In-Cd and hafnium control rods
are smaller than worth of B4C control rods. The worth of the hafnium control rods is
slightly greater than the worth of Ag-In-Cd. All three control rod compositions exhibit
a decrease of the worth for the MASLWR fuel in comparison with PWR fuel of the
same initial enrichment.
ds will be
limited
There are a variety of control rod compositions available for nuclear p
Increasing the enrichment to 8.0% in MASLSR leads to a decrease of the
control rod worth, and current control rod compositions cannot provide sufficient
worth to match that applied in generic PWR fuel.
These evaluations show that boron carbide control rods have the greatest
worth, and are the most preferable option for reactor scram control rods, and for the
reactor power control rods, assuming that the number of the control ro
by the requirement of a transportable reactor design.
151
Figure 4-1.7.6. A
Figure 4-1.7.6. B
152
Figure 4-1.7.6. C
153
Figure 4-1.7.6. D
154
4-1.7.7 Control rod worth under boiling conditions
The boiling of the coolant could occur during accident conditions. The
reliability and the efficiency (worth) of the reactivity control systems under these
conditions will be extremely important to reactor safety evaluations. In a previous
section
moderator temperature equal to the
saturati
trol rod
worth is calculated as the reactivity difference between these state-points. The same
n. The
1.7
the worth of the control rods increases with
conditions; the MASLWR and PWR curves lie close to each other. The increase of the
trol
neu
, we investigated the loss of the soluble boron worth caused by boiling of the
coolant. This series of calculations was performed in order to evaluate the change in
control rod worth under boiling conditions.
Boron carbide was taken as a reference control rod material. The evaluations
were performed for full power conditions, with a
on temperature for the given pressure. The soluble boron concentration was
800 ppm.
The worth calculations were performed ffor zero void-fraction with the control
rods completely removed (CR-Out) completely inserted (CR-In). The con
procedure was used for the calculation of control rod worth for 40% void fractio
results of the calculations are presented in figures 4-1.7.7.A (void fraction 0%) and 4-
.7.B (void fraction 40%).
Figure 4-1.7.7.C shows that
increasing void fraction. At boiling, control rod worth is indifferent to operational
con rod worth is greater for the MASLWR fuel, which is likely related to the
tron energy spectrum at the non-boiling conditions.
155
Figure 4-1.7.7. A
Figure 4-1.7.7. B
156
Figure 4-1.7.7. C
157
tween MASLWR and PWR fuel is increased initial
enrichm
challenges for reactor
refueling and transportation of spent nuclear fuel (or the fueled reactor in case of off-
site
e isotopic comp the ir ASLW be
di rm topic osition t PWR ecific
1) gr ater weight fractions of U-235 and U-236 at the EOC; 2) lower
total plutonium weight fraction, but better quality of the plutonium (greater fraction of
Pu-239 and Pu-241; 3) slightly different composition of the fission products caused by
iation period at a lower power density. These differences should be
considered during the evaluation of the critica ety and radiatio or the
agement, storage nsportation and reprocessing. The extension of the
MASLWR burnup beyond 60 MWD/kgHM wil a larger differe the spent
position comp o PWR spent nuclear fuel.
ed initial enrichment and di e in the operati rameters
lea ill
cause increased absorption of thermal neutro
moderat PWR) will l he
4-1.8 Conclusions
The major difference be
ent (up to 8.0%) and lower operational parameters (lower power density, fuel
temperature, coolant temperature, coolant density). The combined effect of these
factors leads to differences in the core neutronic characteristics.
The increased enrichment of the MASLWR fuel allows greater fuel burnup,
and a longer reactor campaign. The lower operational parameters will also give some
advantages for an extended reactor campaign. The main limitation to the fuel burnup
will be caused by mechanical properties of the fuel cladding and fuel pellet radiation
growth and swelling. Currently, burnup is limited to 60 MWD/kgHM average per fuel
assembly. This means that the spent nuclear fuel of MASLWR with initial enrichment
of 8.0% still could be critical at this burnup level. This creates
refueling).
Th
fferent fo
osition of
comp
radiated M
of the spen
R fuel will also
fuel. The spthe iso
differences are e
a longer irrad
lity saf n safety f
EOC fuel man , tra
l lead to nce in
nuclear fuel com
The increas
ared t
fferenc onal pa
ds to a neutron energy spectrum shift. The increase of the initial enrichment w
ns, and spectrum hardening. Increased
ead to better thermalization of tor density (compared to a
158
neutrons an than in a PWR with the same fuel enrichment. The
ev e c com ss s
will have a slightly harder neutron energy spectrum than a conventional PWR, and this
fact should be taken into account.
The difference in neutron energy spectrum and the difference in isotopic
composition of the irradiated fu l will also lead to differences in the prompt neutron
lifetime and effective delayed neutron fraction between MASLWR and a conventional
PWR. Some assumptions and models used for the PWR should be corrected in
modeling MASLWR.
tor kinetics and dynamics para o be different for a PWR
and for MASLWR. This means that reactivity coefficients, control rod worth and
soluble boron worth should be accurately evaluated for the MASLWR safety analysis.
d a softer spectrum
aluation of th ombined effect of this peting proce hows that MASLWR
e
The reac meters will als
159
ersion ratio of light water
reactors is less than 1, the amount of fissile nuclides will decrease with fuel burnup,
and the multiplication factor will n order to operate MASLWR for
five years without refueling, reactivity reserves are necessary to offset burnup.
According to preliminary calculations [108], if the core average fuel burnup is
approximately 50 MWD/kgHM, then 8.0 % enriched fuel should provide a target of
five years of operation without refueling. Fuel assembly-level calculations performed
for MASLWR and PWR operating conditions demonstrate that MASLWR fuel with
8% enriched fuel will have initial reactivity equal to 33.2% (see table 4-2.1.1). The
compensation of this initial reactivity is not possible using only control rods and
soluble boron because a proper shutdown margin cannot be maintained. Burnable
poisons are therefore essential for the compensation of the initial reactivity excess.
Table 4-2.1.1 Long-term reactivity excess (CASMO-4 Calculations)
Chapter 4-2 – Results Gadolinium Burnable Absorbers for MASLWR Fuel
4-2.1 Burnable Absorbers (Functions, Design Options)
The fission of uranium and plutonium fissile isotopes, as well as neutron
capture and transmutation of nuclides during the reactor operation leads to changes in
the multiplication properties of the core. Because the conv
decrease as well. I
Parameter / Reactor Type (operation conditions) MASLWR PWR Fuel Enrichment, wt % U-235 4.5 % 8.0% 4.5 % 8.0% Initial reactivity excess (infinite reactor) 29.3% 33.2% 27.9 % 31.7 Initial control rod worth -42 % -30 % -53 % -39 % Soluble boron Worth (pcm / ppm) -6.5 -4.0 -5.8 -3.5 Reactivity change Rate (%/per MWD/kgHM) -0.61 -0.40 -0.65 -0.45 Reactivity change Rate (%/per Month) N/A 0.55 0.5 * N/A * Figure for an conventional PWR with partial refueling was taken from [500]
160
The general approach for in the MASLWR is presented in
table 4-2.1.2. The primary function of the burnable absorbers is long-term reactivity
excess compensation; soluble bo ds will be used as a secondary
mechanism for compensation of reactivity changes caused by burnup and a primary
mechanism for operational reactivity control.
Table 4-2.1.2 Functions of Burnable absorbers with regards to the Reactivity control
Control
rods Soluble boron
Burnable absorbers
reactivity control
ron and control ro
Load following operation (Power transients)
Primary, Hour and Dailycompensation
Secondary, Weekly
compensation
N/A (Passive)
Reactor start control, Reactivity defects compensation
Depends on start procedure
Depends on start procedure
N/A (Passive)
Ope
ratio
nal
Reactivity excess compensation (long term, reactor campaign)
Secondary Requires a large amount of CR
elements
Secondary, Active
compensation
Primary, Passive
compensation
Reactor scram, Safety functions
Primary, Fast Scram
Secondary, Ensure Sub criticality
N/A (Passive)
Non
-ope
ratio
nal
Refueling, Long term sub criticality, while control rods may be removed with SNF Assembly
Secondary Primary N/A (Passive)
The burnable absorber material must satisfy several general requirements
[500]:
The capture cross section of the poison material must be significantly higher
than that of the fissile material so that poison residue at the end of the depletion cycle
is low. It must also be low enough so that it does not burn out before a significant
reactivity loss due to depletion takes place.
The reaction products of the burnable poison should have sufficiently low
cross sections so that their buildup does not cause a significant reduction in cycle
length.
161
s currently used as burnable absorbers: boron,
gadolinium
burnable absorbe
specifications for order to understand the nuclear
reactions and imp
is presented in the
4-2.1.1. Boron
The poison material or its reaction products should not compromise the
mechanical integrity of the reactor. If a design calls for use of separate burnable
poison elements, the radiation and corrosion resistance of these elements should be
verified. If the poison material is mixed with the fuel, its behavior with irradiation
should be evaluated to assure that the fuel failure mechanisms are not enhanced.
There is several material
and erbium [391]. Other elements are currently being studied for use as
rs, but are not used in industry and the detailed technical
these elements are not available. In
ortant properties of the burnable absorbers, some basic information
next section.
two naturally occurring isotopes of boron: B-10 (19.9%) and B-11
ron absorption cross section of B-10 (see figure 4-2.1.1.A) is several
ude greater than that of B-11. Some vendors can provide boron
The technical details and specifications for the enriched boron are
There are
(80.1%). The neut
orders of magnit
enriched in B-10.
confidential commercial information; we make some general assumptions about
enriched boron in this study.
or
burnab
boride ZrB2, borosilicate glass, steel and aluminum
alloys o
The principal nuclear reaction of the boron absorber element (control rod
le absorber) is (n, alpha) reaction:
HeLinB n42
73,
10
105 +⎯→⎯+ α
The typical chemical compositions of boron used in the nuclear industry are
boron carbide (B4C), zirconium
f boron and boric acid HBO3.
162
sorption cross section of boron Figure 4-2.1.1. The total neutron ab
4-2.1.2. Gadolinium
There are seven naturally occurring isotopes of gadolinium: Gd-152 (0.20%),
Gd-154 (2.18% ), Gd-155 (14.80%), Gd-156 (20.47%), Gd-157 ( 15.65% ), Gd-158
(24.84%) and Gd-160 (21.86%). The major nuclear reaction mechanism with incident
neutrons is neutron capture. The chain of the gadolinium isotopes and nuclear
reactions tracked in CASMO-4 [234] are shown in figure 4-2.1.2 A:
Figure 4-2.1.2 A Gadolinium Reaction tracked in CASMO-4
163
gadolinium in terms of neutron absorption; owever, the natural occurrence of Gd-152
is very low. The total neutron absorption cross-section of gadolinium isotopes are
ented in figure 4-2.1.2 B.
Figure 4-2.1.2 B Total neutron absorption cross sections of gadolinium
This figure shows that Gd-152, Gd-155 and Gd-157 are the most important isotopes of
h
pres
Gadolinium isotopes are used in LWRs primarily because their absorption
cross sections have a significant slope in the thermal energy range and there are
epithermal resonances in Gd-155 and Gd-157, which will have an effect on the
xellian spectrum of the thermal neutrons. Also, large neutron absorption cross
ions in the thermal and epithermal energy ranges lead to significant self-shielding,
shielding and control rod shadowing effects in gadolinium-doped fuel.
The typical chemical composition of gadolinium used in the nuclear industry is
gadolinia powder (Gd2O3), uniformly mixed with UO2 powder during fuel pellet
pressing.
Ma
sect
fuel
164
4-2.1.3. Erbium
There are six naturally occurring isotopes of erbium: Er-162 (0.14%), Er-164
(1.61%), Er-166 (33.6%), Er-167 (22.95%), Er-168 (26.8%), Er-170 (14.9%). The
major nuclear reaction with incident neutrons is neutron capture. The chain of the
erbium isotopes and nuclear reactions tracked in CASMO-4 are shown in figure
4-2.1.3 A [234]:
Figure 4-2.1.3.A. Erbium Reaction tracked in CASMO-4
Figure 4-2.1.3 B Total neutron absorption cross sections of erbium
165
Er-167, which has a highes isotopes,
is the m st important isotope of erbium from actor design. Er-167
also ha
of erbium used in the nuclear industry is
erbia p
oss sections
t cross section among other natural erbium
the perspective of reo
s an epithermal resonance with a tail covering the entire thermal energy range.
This allows erbium isotopes to be used as a reactivity feedback corrector as well as a
burnable absorber. The absorption cross section of Er-167 is comparable with the
fission cross section of U-235, (see fig. 4-2.1.4) which means that at the end of the
cycle some erbium will remain in the fuel. More information about the use of erbium
in a 8.0% MASLWR enriched fuel can be found in [108].
The typical chemical composition
owder (Er2O3), uniformly mixed with UO2 powder during fuel pellet pressing.
However in research reactors, erbium is often used in a matrix of dispersed fuel (like
TRIGA reactor fuel). There is ongoing research investigating the possibility of erbium
coating of fuel.
Figure 4-2.1.4 comparison Gd-155, Gd-157 and Er-167 absorption crwith U-235 fission cross section
166
Burnable poisons designs, s and manufacturing technologies
are very dependent on the specific manufacturer [397]. Table 4-2.1.3 presents a
summary of the options available on the market, according to the sources [48, 397].
Table 4-2.1.4 Burnable absorbers technology and material options
absorbing material Technology
Abs
orbe
r El
emen
t
Chemical composition
B B4C
B Aluminum alloy with B
B ZrB2 Rem
ovab
le
Cluster Assembly (Similar to PWR
control rod assembly)
B Boron-silicate glass
Gd Gd2O3 Mixed with fuel
Er Er2O3
B ZrB2 Coated
Er Coated Erbium Composition
B Aluminum alloy with B
Bur
nabl
e A
bsor
bers
Non
rem
ovab
le
Fixed pins with burnable absorber B B4C
Removable burnable absorber rods are widely used in conventional PWRs with
partial refueling. The use of removable burnable absorber clusters in control rod guide
tubes of standard fuel assembly when the assembly is placed in a position where
control rod drives are unavailable, allows a certain degree of freedom for fuel
management. In particular, the standard fuel assembly design can be used for the
reactor, and burnable absorbers could be loaded into the first or second cycle fuel
assemblies (i.e. with 3 or 4 batch refueling pattern). However application of this
167
design option for MASLWR with is not practical, as the refueling
will be done for the whole core and positions for conventional clusters of removable
burnable absorbers are limited.
Non-removable burnable absorbers require detailed specification for each fuel
assembly modification, with a precise description of the burnable absorber location,
material, concentration and other manufacturing specifications. Different
manufacturers may have their own technical limitations, like material type and
properties, and number of the different absorber rods or pellets per fuel assembly.
The marketing and manufacturing of the fuel for MASLWR from an
independent fuel vendor creates additional requirements (recommendations) for the
burnable absorber design: burnable absorbers should be designed such that the
MASLWR operating utility can order fuel with a standard specification from more
than one vendor (non-monopolistic obligations).
Several different compositions of burnable absorbers (BA), including erbium
and gadolinium were evaluated in previous preliminary research [107, 108, 111]. The
burnable absorber with gadolinium uniformly mixed in the fuel pellet was chosen for
the performance of the current research (feasibility study) for several reasons:
1) Gd burnable absorber technology is well known, licensed and certified.
2) A wide variety of manufacturers can provide fuel with Gd burnable absorbers
(some removable BA or integrated BA have specific features, which are
manufactured by unique vendors)
3) A wide variety of codes and other data libraries are available for core design
with Gd burnable absorbers.
4) The self-shielding effect of the gadolinium allows the designer to program fuel
burnup with respect to the multiplication factor curve maxima and shape using
multiple Gd concentrations in the fuel assembly.
off-site refueling
168
4-2.2. Calculation Model and Assumptions
In order to design a core for MASLWR some preliminary study of the burnable
absorber behavior should be performed. The previous chapter discusses the results of
research focused on differences between PWR and MASLWR fuel caused by
increased enrichment and operational conditions. In this chapter we present a study of
the behavior of burnable absorbers in 8.0% enriched MASLWR.
First, a study of the standard PWR fuel burnable absorber pattern will be
investigated with 8.0% enriched fuel at MASLWR operational conditions. We then
investigate self-shielding, self-shadowing and control rod-shielding effects. Finally,
the option of multiple burnable absorbers will be discussed as it may yield a more
optimal fuel burnup pattern. These studies form a set of fuel assembly segments for
the feasibility study of the MASLWR core.
Several assumptions will be made for the fuel assembly design within the
scope of the feasibility study:
Assumption 1: The densit adolinium will be assumed equal
to the density of the fuel without the bur ble absorber, and equal to 96% of the
theoretical density of uranium oxide at cold conditions. This is a reasonable
assumption because the actual density of the fuel pellets vary from one manufacturer
to another in the range between 95%-97% of theoretical density, and the technology
for producing fuel pellets with gadolinium also varies, making the correlation for the
density correction vendor-dependent. Also such corrections would not have a
significant effect at the current stage of the feasibility study, but the complications
caused by these correlations may effect the understanding of some effects in the fuel.
(For some comparison studies it is better to have similar initial conditions and
properties of the fuel in order to specifically investigate the effect of the burnable
absorber).
Assumption 2: Traditionally, uranium with reduced enrichment is used in fuel
pellets with burnable absorbers, compared to the enrichment of the fuel pellets without
burnable absorbers. This minimizes the pin peaking factors in the fuel assemblies, and
y of the fuel with g
na
169
burnup limits for the pins with burnable poisons. However, current correlations
available in the literature are designed for PWR fuel with standard enrichment. The
operational conditions and initial enrichment of the MASLWR fuel is different from
those of a conventional PWR, and allowable peaking factors will not be available until
the feasibly study of the core will be completed, analyzed and discussed with other
MASLWR designers and technical specification for the preliminary core design will
be issued. In this dissertation, it is assumed that fuel pellets with burnable absorbers do
not have reduced enrichment, and all fuel in the fuel assembly have the same initial
enrichment of 8.0%.
Assumption 3: Traditionally, fuel pellets have a very specific form at cold
conditions (thyroidal sides of the cylinder and spherical dishes of the top and bottom
of the fuel pellet). This shape is necessary to accomodate thermal and radiation
expansion of the fuel pellet, which is a function of the operational conditions and
irradiation history. For a detailed core design the actual shape of the fuel pellet should
be take
fuel, and discussion of the potential core design issues that should be
addressed in the future design.
n into account. However, for the current feasibility study it was considered that
the fuel pellets would have the default, quasi-cylindrical shape, and default thermal
expansion coefficients of CASMO-4 will be used for the calculations.
In addition to these assumptions, it is necessary to recall that the purpose of the
current feasibility study is an investigation of the possibility of the proposed
MASLWR design to reach five effective years of safe operation with 8.0% initially
enriched
170
4-2.3. Fuel Assembly with Gadolinium BP (standard geometries).
Gadolinium burnable absorbers are widely used in modern PWR fuel. The
MASLWR design requires the use of standardized, off-the shelf equipment and
technologies as much as it reasonable and possible. Several standard fuel assembly
types, similar to Areva/Siemens PWR fuel assemblies [397] were simulated with
CASMO-4. The technical specifications of the considered fuel assemblies are given in
the table 4-2.3.1, and the geometric layout of the burnable absorber pins are presented
in figure 4-2.3.1. The considered fuel assemblies all have geometries similar to
standard PWR fuel assemblies, but the fuel enrichment of all pins is 8.0%. Simulations
were performed for the standard arrangement of burnable absorber pins, but the
gadolinium concentration in all absorber pins is equal to 9.0%.
Table 4-2.3.1 The fuel pin specifications of some standard fuel assemblies
Fuel assembly Marker Number of
Pins Uranium
Enrichment Weight Percent of
Gadolinium
236 8.0 % No Gadolinium
16 8.0 % 8.0 % FM_8GSM1
12 8.0 % 4.0 %
240 8.0 % No Gadolinium
16 8.0 % 8.0% FM_8GSM1
8 8.0 % 4.0%
244 8.0 % No Gadolinium
16 8.0 % 6.0 % FM_8GSM1
4 8.0 % 2.0 %
260 8.0% No Gadolinium FM_8GSM1
4 8.0% 2.0%
171
Figure 4-2.3.1 The geometry s me standard fuel assemblies
Sem Modified standard fuel assembly layout
pecifications of so
i-standard fuel assembly layout. Gadolinium concentration
as in table 4-2.3.1 Gadolinium concentration 9.0%
CASMO 4 Model (1/8) Fuel Assembly Layout CASMO 4
Model (1/8) Fuel Assembly Layout
FM_8GSM1
FM_8G9M1
FM_8GSM2
FM_8G9M2
FM_8GSM3
FM_8G9M3
FM_8GSM4
FM_8G9M4
172
emblies with burnable absorbers, and make
conclus
cation
factor
The purpose of the CASMO simulations is to investigate the decrease of the
initial reactivity excess in the fuel ass
ions on the possibility of using standard burnable absorbers layouts.
The multiplication factor as function of time is presented in figure 4-2.3.2 A
for fuel assemblies with semi-standard arrangements of burnable absorber pins, and in
figure 4-2.3.2 B for fuel assemblies with a gadolinium concentration of 9.0%. These
graphs show that the proposed burnable absorber arrangement and design does
decrease the initial reactivity excess. However, the decrease in the multipli
is not significant and a high concentration of soluble boron and increased
control rod use will be necessary during operation. It is clear that the standard
burnable poison map was tuned for the use in a conventional PWR with lower fuel
enrichment and with a partially refueled core, so the irradiated fuel after the first and
second campaign compensates the initial reactivity excess.
Figure 4-2.3.2 A
173
Figure 4-2.3.2 B
Figure 4-2.3.2 C
174
Figure 4-2.3.2 C shows multiplication factor changes in
standard and modified fuel assemblies in soluble boron-free operation at the beginning
and middle of cycle. In the case of a single concentration of burnable absorber equal to
9.0% of gadolinium, the principal difference in the curves occurs at the beginning of
the cycle. After 35-40 MWD/kgHM of burnup, the curves are parallel and the
difference is negligible. This indicates that all the burnable absorber is gone, and after
this period the change in multiplication is due primarily to uranium fuel depletion and
the products of the uranium reaction chains.
Figures 4-2.3.1 A and C show that lower concentrations of gadolinium could
lead to changes in fuel behavior. For example, consider fuel assemblies FA_8GSM1
and FA_8GSM2. Both assemblies have 16 pins with a gadolinium concentration of
8.0%, but fuel assembly FA_8GSM1 has 12 fuel pins with gadolinium concentration
2.0% and fuel assembly FA_8GSM2 has 8 fuel pins with gadolinium concentration
2.0%. The curves for these fuel assemblies are significantly different up to a burnup of
10 MWD/kgHM. After this burnup, the assemblies behave similarly. This means that
the effective burnup period for distributed pins with a gadolinium concentration of
2.0% is smaller than for distributed fuel pins with a gadolinium concentration 8.0%,
primarily due to different amount of absorbing material and self-shielding effects. So
from this study we may conclude:
1) Standard PWR burnable absorber layouts are not sufficient for
compensation of the initial excess reactivity of the MASLWR fuel
with an initial enrichment of 8.0%
2) Fuel pins with higher concentrations of gadolinium hold down
reactivity longer (in terms of burnup time units) due to a greater
inventory of absorbing material and self-shielding effects.
3) A new burnable absorber layout should be designed for the
MASLWR fuel assembly so the specific requirements for the initia
excess reactivity compensation will be fulfilled, and self-shielding
and shadowing effects of the burnable absorbers should be
investigated for MASLWR fuel with an initial enrichment of 8.0%.
a comparison of
l
175
4-2.4. Self-shielding effects in fu l with burnable absorbers.
he high neutron absorption cross-section of gadolinium leads to self-shielding
effects in fuel pins with large gadolinium concentrations. MASLWR fuel has greater
enrichment and different operatin conventional PWR fuel, so the
study of the self-shielding effect should be performed for MASLWR fuel irradiated at
MASLWR operating conditions.
In order to evaluate the self-shielding effect, four fuel assemblies were
considered with Gd2O3 concentrations of 2.0%, 4.0%, 6.0% and 8.0%. The number of
fuel pins with burnable absorbers were chosen to be 96, 48, 32 and 24, respectively, to
have equal initial burnable absorber material in each assembly. The technical
specifications and loading patterns are presented in table 4-2.4.1 and figure 4-2.4.1
Table 4-2.4.1 Fuel Assembly specifications for Gd self-shielding survey
e
T
g conditions than
Fuel assembly name FA_8TG02 FA_8TG04 FA_8TG06 FA_8TG08Number of pins without burnable poisons
168
216
232
240
Uranium enrichment, wt% 8.0% Number of pins with burnable poisons (Gd2O3)
96
48
32
24
Uranium enrichment, wt% 8.0% 8.0% 8.0% 8.0% Concentration of Gd2O3 2.0% 4.0% 6.0% 8.0% Volumetric fuel density 10.10 g/cm3 (CZP) and 9.937 g/cm3 (HFP) Linear fuel density (HFP) 5.383 g/cm Linear Gd2O3 load per fuel pin (HFP) 0.108 g/cm 0.215 g/cm 0.323 g/cm 0.431g/cm
Linear Gd2O3 load per fuel assembly (HFP) 10.34 g/cm
Gadolinium wt% in pin 1.74% 3.47% 5.21% 6.94% Linear Gadolinium load per fuel pin (HFP) 0.093 g/cm 0.187 g/cm 0.280 g/cm 0.374g/cm
Linear Gadolinium load per fuel assembly (HFP) 8.967 g/cm
176
Figure 4-2.4.1 Fuel Assembly geometry specifications for Gd self-shielding survey
FA_8
TG
02
FA_8
TG
04
FA_8
TG
06
FA_8
TG
08
177
le absorbers is smaller with respect to the weight fraction of the
burnable absorber – or, the greate of burnable absorber, the smaller
amount of uranium oxide in the fuel pi . However, the load of uranium (and
particularly U-235) was the same for all fuel assemblies, due to the smaller amount of
the abs
ase depletion case with soluble
boron c
The design of the fuel assemblies used in this study was made in a way that
fuel assemblies has a similar mass of the burnable absorber (gadolinium), and fuel
pins with burnable absorber were distributed evenly. The amount of uranium in the
fuel pins with burnab
r concentration
n
orber pins in the fuel assembly with greater concentration of the burnable
poison. So we may conclude that the difference in initial reactivity, and changes in
reactivity with depletion pattern are caused by the self-shielding effect of gadolinium
in the burnable absorber pins.
The multiplication factor behavior for the fuel assemblies considered in self-
shielding survey is presented in figure 4-2.4.2 A for a b
oncentration 800 ppm; and in figure 4-2.4.2 B for a boron-free core (soluble
boron concentration 0.1 ppm).
Figure 4-2.4.2 A
178
Figure 4-2.4.2. B
These two figures show that the major differences between fuel assemblies
occur in the beginning of the cycle. The initial multiplication factor is smaller in the
fuel assembly where the gadolinium is distributed in a greater number of the fuel pins.
However, with depletion, the multiplication factor in the fuel assembly with 96 2.0%
gadolinia pins grows faster than in the fuel assembly with 48 4.0% gadolinia pins, and
fuel as
. The self-shielding effect in fuel assemblies
with greater gadolinium concentration results in different shapes for the multiplication
factor curves; the curve for FA_8TG08 does not have a local maximum.
sembly FA_8TG02 (96 2.0% Gd pins) reaches the maximum multiplication
factor earlier than other fuel assemblies.
The self-shielding effect in fuel assembly FA_8TG08 is strong, so the same
amount of gadolinium provides only half of the initial decrease of multiplication factor
compared to fuel assembly FA_8TG02
179
Figure 4-2.4.2 C shows the or in the middle of the cycle. The
black curve represents the standard PWR fuel assembly with 8.0% enriched fuel and
without burnable poisons.
Figure 4-2.4.2 C
multiplication fact
fter a certain burnup level, the multiplication factor curves for poisoned fuel
are par
resented in figure 4-2.4.3 A-D). These curves were built
A
allel to the non-poisoned fuel curve. This represents the fact that at some point,
the concentration of burnable absorber atoms becomes negligible. Consider the atomic
concentration curves for all gadolinium isotopes in fuel pins with 2%, 4%, 6% and 8%
concentration of gadolinia (p
for single pin atomic concentrations, so for a fair comparison the number of pins with
burnable absorbers should be taken into account through calculation of the linear atom
density per pin (multiplying by fuel pellet cross section) per fuel assembly segment
(multiplying by number of pins with burnable absorbers, See figures 4-2.4.4 A, B.)
180
Figure 4-2.4.3 A
Figure 4-2.4.3 B
181
Figure 4-2.4.3 C
Figure 4-2.4.3 D
182
Figure 4-2.4.4 A
Figure 4-2.4.4.B
183
These figures show that th gadolinium isotopes Gd-155 and
Gd-157 are decreasing with depletion a
produc
Figure 4-2.4.5 A
e concentrations of
nd the concentrations of their daughter
ts Gd-156 and Gd-158 are increasing correspondingly. The change in the
concentration of the other gadolinium isotopes is negligible. The initial concentration
of Gd-152 is relatively low and its impact is minor; however this isotope has a large
absorption cross-section, and its concentration is decreasing with depletion (figure 4-
2.4.5.A).
The depletion of Gd-155 and Gd-157 and production of the daughter isotopes
are different in the different fuel assembly segments due to self-shielding. The
concentrations of Gd-155, Gd-156, Gd-157 and Gd-158 are presented in pairs in in
figures 4-2.4.5 B and C.
184
Figure 4-2.4.5 B
185
Figure 4-2.4.5 C
186
These graphs demonstrate ding results in different burnout
rates for the major burnable absorber isotopes Gd-155 and Gd-157. Also we can see
that effective burnout of Gd occurs faster in a fuel assembly with a smaller
concentration of the burnable absorbers. Finally, we conclude that the presence of two
major burnable absorber isotopes leads to an early effective burnout of Gd-155, and
then effective burn-out of Gd-157. A summary this data is presented in table 4-2.4.5
Table 4-2.4.5 Self-shielding survey summary
Fuel assembly segment name FA_80_00 FA_8TG02 FA_8TG04 FA_8TG06 FA_8TG08
that the self-shiel
Initial K∞ (base case BOR=800ppm) 1.4282 0.9892 1.1245 1.2081 1.2390
Initial K∞ at 0.1 MWD/kgHM (base case BOR=800ppm) 1.3970 0.9794 1.1072 1.1861 1.2157
Maximum value of K∞ 1.3970 1.2571 1.2300 1.1967 1.2157 Fuel depletion at K∞ = Max(K∞) 0 13.5 18.0 22.0 0 Gd-155 Effective depletion end* N/A 12 18 24 28 Gd-157 Effective depletion end* N/A 22 28 35 40
* Approximate end effective depletion in MWD/kgHM determined from graph
The results of the present survey aid the designer in programming reactivity
excess compensation by mixing different concentrations of gadolinia burnable
absorber pins in a fuel assembly. Two pins types with different concentrations will
allow getting four principal burnout points, which means that it is possible to
accurately program the change in multiplication factor with depletion.
However, in order to start feasibility study for a MASLWR design of the core
we sho
uld note, that the presented survey was performed for evenly distributed
burnable absorber pins. Knowing that the gadolinium self-shielding effect has a great
importance, we may assume that the shadowing effect of neighboring pins with
burnable absorbers may also have important consequences for the fuel behavior.
187
4-2.5. Shadowing effects e absorbers in V and W fuel assembly modifications
To evaluate the importance of the burnable absorber shadowing effect on
regular fuel pins, three series
etry type (FA_8UGXX, or U-type) assumes an even distribution
of 96 burnable absorber pins in the fuel assembly, similar to the pattern of fuel
assembly FA_8TG02 used in the previous survey. The second geometry type
(FA_8VGXX, or V-type) assumes positioning of the 96 burnable absorber pins in the
center of the fuel assembly, causing the center burnable absorber pins to be shadowed
by neighboring burnable absorber pins.
The third geometry type (FA_8WGXX, or W-type) assumes positioning of the
96 burnable absorber pins in the corners of the fuel assembly. Assuming reflective
boundary conditions, this geometry has some similarities with V-type, except for the
fact that now water pins (or control rods) are located in the non-poisoned fuel zone.
Table 4-2.5.1. The specifications of fuel assemblies used in a shadowing survey
of Gd burnabl
of fuel assembly calculations were considered. The fuel
enrichment for all calculations was 8.0%, and each fuel assembly series has the same
number of burnable absorber pins (96). Three geometries were considered (see the
figure 4-2.5.1). For each series of experiments, ten simulations were performed with
concentrations of gadolinia 1.0%, 2.0%, 3.0% … 9.0% and 10.0% respectively.
The first geom
Fuel assembly Marker
Number of Pins
Uranium Enrichment Weight Percent of Gadolinium
168 8.0 % No Gadolinium FA_8UGXX 96 8.0 % XX% = 1.0%; 2.0%; … ; 10.0% 168 8.0 % No Gadolinium FA_8VGXX 96 8.0 % XX% = 1.0%; 2.0%; … ; 10.0% 168 8.0% No Gadolinium FA_8WGXX 96 8.0% XX% = 1.0%; 2.0%; … ; 10.0%
188
Figure 4-2.5.1 Fuel Assembly g ions for Gd-shadowing survey
FA_8
UG
XX
eometry specificat
FA_8
VG
XX
FA_8
TG
XX
* The same geometry U, V and W will be used for a control rod shadowing survey
The complete S3C matrix of branch cases was performed for each combination
of geom
pletion state
points)
n
figure 4-2.5.2 A – for U-type series of fuel assembly segments, B – For V-type series
of fuel assembly segments and C – For W-type series of fuel assembly segments.
etry type and gadolinium concentration, assuming a base case of depletion
under MASLWR operation conditions, with soluble boron concentration of 800 ppm.
We generate multiplication factor curves for the base case (BOR=800ppm, 100
depletion state points) and a boron-free branch case (BOR=0.1 ppm, 15 de
.
The multiplication factor as a function of depletion is presented i
189
Figure 4-2.5.2 A
190
Figure 4-2.5.2 B
191
Figure 4-2.5.2 C
Each of these curves is important, because the selection of the assembly
segment geometry and burnable absorber concentrations for prototypic core design
will be done with implementation of the curves, and “burnup-time-scale” approach.
192
Figure 4-2.5.3
For the evaluation of the self-shielding effects, some comparison of the
segments is necessary. Figure 4-2.5.3 represents comparison of U, V and W assembly
geometries for three gadolinia concentrations: low, medium and high self-shielding
(2%, 5% and 10%).
193
ication factor compared to the V-type and W-type fuel assemblies. The
shadow
The shapes of the multiplication factor curves are depend sensitively on the
arrangement of the BA pins in the fuel assembly, even while the number of the BA
pins and the concentration of burnable absorber are the same. The differences are
caused mainly by 1) the shadowing effect of neighboring pins, and 2) the spectrum
changes (shifting) in combinations of “BA-pin + Fuel Pin”, “BA-pin + water rod
(guide tube)” and “Fuel Pin + water rod (guide tube)”.
There are several ways to evaluate and characterize the shadowing effect on
the fuel assembly multiplication factor. Shadowing decreases the effect of the
poisoned pins on the overall decrease of the multiplication factor (see figure 4-2.5.4).
The U-type fuel assembly has distributed gadolinia-doped pins, so it has a lower initial
multipl
ing also that initial multiplication factor in a U-type fuel assembly is more
strongly dependent on and decreases faster with the increasing gadolinia
concentration, than in the V-type and W-type fuel assemblies.
Figure 4-2.5.4
194
tions of gadolinia in a range of 0-10%.
On the
Figure 4-2.5.5 A
The initial multiplication factors for V-type and W-type fuel assemblies are
above 1 for boron-free operation and concentra
contrary, the initial multiplication factor for the U-type assembly with
distributed BA pins, is above 1 for gadolinia concentrations up to 2.5%, and then
strongly drops below 1 with growth of gadolinia concentration.
Another important characteristic is the coordinates of multiplication factor
maximum with respect to the fuel assembly burnup, and the value of the maximum
multiplication factor (figure 4-2.5.5 A and B). Shadowing moves the maximum of the
multiplication factor curve toward the end of the cycle, which is very important for the
five-year cycle of the MASLWR core.
195
Figure 4-2.5.5 B
Moving the multiplication factor towards the end of the cycle also leads to the
decrease of the multiplication factor value at the maximum. The use of V-Type and
W-type fuel assemblies with shadowing effect causes an overall flattening of the
multiplication factor curve.
The shadowing effect also has some disadvantages. The shadowing in V-type
and W-type fuel assemblies will create areas with depressed flux and decreased power
generation, so the non-poisoned fuel pins would have a greater load in order to
maintain fuel assembly average power generation. The pin power reconstruction and
pin peaking factor were studied for all series of fuel assembly segments. The pin-
peaking factor as a function of depletion is presented in figure 4-5.2.6.
196
Figure 4-2.5.6
197
ct does not necessarily lead to a
dramat
f the peaking factor requirements will be discussed in the core
design
el assembly could either
increas
These plots show that the shadowing effe
ic increase of the pin peaking factor. The pin-peaking in the range of 1.4-1.6 is
higher than the industry average for PWR fuel assemblies [198], however the lower
power density of the MASLWR reactor could permit greater peaking factors [104,
112]. The fulfillment o
chapter. Here, our purpose is to discuss the behavior of pin-peaking related
with self-shielding and shadowing effects.
For all fuel assembly segments, the initial pin-peaking factor is increases with
gadolinia concentration (see figure 4-2.5.7). The growth rate is not constant, and
decreases with increasing gadolinia concentration due to self-shielding. The
shadowing affects the pin-peaking factor indirectly, through the redistribution of the
power generation profiles. The shadowing effect in the fu
e pin peaking factor (as for W-type fuel assembly) or decrease it (as for V-type
fuel assembly) relative to the fuel assembly with distributed poisoned pins (U-type).
Figure 4-2.5.7
198
generat
In order to investigate the shadowing mechanisms, and demonstrate the energy
ion profile evolution in the fuel assembly with shadowing, we consider pin-
power reconstruction for U-type, V-type and W-type fuel assemblies with gadolinia
concentrations 2% and 10% at the beginning, middle and end of the cycle (burnup
equal to 0, 30 and 60 MWD/kgHM respectively). The results are shown in figure 4-
2.5.8 A, B and C
Figure 4-2.5.8 A
199
Figure 4-2.5.8 B
assembly with fuel pins mixed with burnable absorber are
filled w
Locations in the
ith gray; non-poisoned pins are in white. The figures in boxes are representing
pin power generation relative to average fuel assembly power generation. The red font
is used for relative pin-powers above 1.1, and green font is used for relative pin
powers below 0.95. Bold font is used for the highest and lowest relative pin powers.
200
C Figure 4-2.5.8
The figures show that at the beginning of the cycle, evenly distributed absorber
pins ha
ower distribution.
s reduced pin-power peaking. The relative power drop in the BA pins is
decreasing increasing gadolinium concentration. By the middle of the cycle, the
difference in power generation between regular fuel pins and poisoned pins is much
smaller, and at the end of the cycle, both BA pins and regular pins have the same
relative pin p
201
pins are concentrated in the center of
lot of water holes or guide tubes in this region, which
leads to better moderation of fast neutrons. In both cases (10% and 2% gadolinia
concen
n, gadolinia concentrations in
the fuel pins decrease. However there is some spare uranium in the poisoned pins that
ha
of gadolinia) at the middle of cycle, all gadolinium is gone, and the saved uranium
leads to more energy production in the poisoned pins than in regular fuel pins. In fuel
assembly FA_8GW10 (10% of gadolinia), self shielded and saved gadolinia still
causes a depression of the power generation at the middle of cycle. At the end of the
cycle, the major power generation happens in the pins that were initially poisoned
and/or shadowed by other burnable absorber pins.
In a W-type fuel assembly, fuel pins with burnable absorbers are located on the
periphery, and non-poisoned pins are located in the center where more water holes are
available and the thermal flux is greater due to better moderation and lower
absorption. These phenomena cause increased pin-peaking factors in the central area
of the fuel assembly. However, at the middle of the cycle for fuel assembly
FA_8WG02, and at the end of cycle for fuel assembly FA_8WG10, uranium saved in
the BA pins leads to greater energy generation in these pins.
This study shows that burnable absorber pin shadowing effects provide
additional capabilities for the extension of the core cycle and flattening of the
multiplication factor curve. The shadowing effect may lead to an increase of the pin
peaking factor; however this is not dramatically high, and in some cases peaking
factors could be even decreased. This study provides a basis for a core feasibility
study.
In the V-type fuel assembly, all poisoned
the fuel assembly. There are a
trations) there is depression of the power generation in the beginning of cycle,
with local increases around water holes. With depletio
s been saved due to shielding from gadolinium. In fuel assembly FA_8GW02 (2%
The assembly segments used in this survey will be included in the fuel library
for the core design.
202
sections of this chapter we demonstrated that the use of
burnab
tion and initial gadolinia concentration demonstrates that the location
of the
he worth of the control rods strongly depends on the unperturbed flux in the
region where the control rod is inserted. We can say that the unperturbed flux
In the case of the V-type fuel assembly, burnable
absorbe
the burnable absorber pins are distributed evenly,
some c
ent (see figure 4-2.6.1 A, B and C).
4-2.6. Control rod shadowing effect
In the previous
le absorbers leads to strong self-shielding and shadowing effects. The use of
this shadowing effect in fuel assembly design (V-type, W-type) leads to the
redistribution of the neutron flux in the fuel assembly, which changes with time and
depletion.
The discussion of the relative pin power distribution in the fuel assembly as a
function of deple
“water holes” or guide tubes is very important. Increased moderation effects
around “water holes” leads to flux suppression in the area around it. Burnable absorber
pins cause a flux depression in their immediate neighborhood. In order to maintain the
assigned fuel assembly power, the power generation (and neutron flux) must increase
in the non-poisoned areas, effectively pushing neutron flux out of the burnable
absorber region.
T
represents neutron importance.
r pins will surround most of the control rods, so the control rods will be
inserted into the area with a depressed neutron flux. In the W-type fuel assembly,
control rods will be inserted into the non-poisoned region of the fuel assembly, and
will see increased neutron flux (compared to the fuel assembly no burnable poisons).
In the U-type fuel assembly where
ontrol rods will be inserted in increased flux areas, others will be inserted into
depressed flux areas, and the overall effect is more difficult to characterize.
To study the shadowing effect of the burnable absorbers on the control rods,
branch case simulations were performed with insertion of the B4C control rods. The
multiplication factor and control rod worth were calculated as a function of depletion
for each fuel assembly segm
203
Figure 4-2.6.1 A.
In the U-type fuel assembly, the importanc
in the fuel assembly without burnable poisons (FA_80_00). W
e of the water holes is greater than
e conclude that the
presence of the burnable absorber in the U-type fuel assembly increases the efficiency
(worth) of the control rods. The control rod worth in U-type fuel assembly grows with
increasing burnable absorber concentration. After the burnup of the gadolinium in the
BA pins, the control rod worth behaves similar to the non-poisoned fuel assembly
(FA_80_00).
204
Figure 4-2.6.1 B.
V-type fuel assem rnable absorber pins shadow most of the control
sing a decrease control rod worth (absolute value of negative
ntrol r fter gadolin
(FA_ al
concentration of gadolinia leads to a longer burnout time (effective burnup).
In the bly bu
rods. It is cau of the
reactivity inserted by co ods). A ium burnout, the control rod worth
curve behaves similar to the non-poisoned fuel assembly 80_00). The high initi
205
Figure 4-2.6.1 C.
“Pushing” of the neutron flux towards the center of the fuel assembly leads to
increased control rod worth in the W-type fuel assembly. This effect is stronger with
increasing initial concentration of the gadolinia. For gadolinia concentrations of 8%-
9%, the control rod worth is almost constant in the W-type fuel assembly.
segments, for boron concentrations in the range 0-800 ppm. (Fig. 4-2.6.2 A,B and C)
The evaluation of soluble boron worth was performed for all fuel assembly
206
Figure 4-2.6.2 A.
In the U-type fuel assembly, burnable poison pins are distributed evenly and
the overall thermal neutron absorption is greater than the fuel without burnable
poisons
increase in areas with non-poisoned pins. This causes an increase
in the initial soluble boron worth with increasing gadolinia concentration.
(FA_80_00). This leads to a decrease of the soluble boron worth in the U-type
fuel assembly. However, the increase of the burnable poison concentration leads to
self-shielding of the burnable poison pins, and to a flux decrease around them, and a
corresponding flux
207
Figure 4-2.6.2 B.
In the V-type fuel assembly, flux is “pushed” out from the center of the fuel
assembly, so the relative power generation and relative neutron flux is greater in the
non-poisoned region. This also means that the importance of the thermal neutrons in
that re
hment is greater, so the boron
worth is slightly smaller than for the same burnup in non-poisoned fuel.
gion is greater, and the fuel multiplication factor is sensitive to the boron
concentration in that region. The magnitude of this “push-out” effect increases with
increasing initial burnable poison concentration, so the soluble boron worth is greater
at the BOC. At the MOC and EOC, “saved” uranium leads to the fact that in shadowed
regions, where gadolinium is burned out, uranium enric
208
Figure 4-2.6.2 C.
The shadowing effect of the poisoned pins and “pushing” of the neutron flux
towards non-poisoned pins discussed above is even stronger in the W-type fuel
assembly. The water holes in the W-type fuel increase the thermal flux in the center of
the fuel assembly, where non-poisoned pins are located. This increases the importance
of the soluble boron in the central area.
It is clear from this study that burnable poisons can either decrease or increase
the efficiency of the control rods and required soluble boron concentration. These
effects must be taken into account during the core design.
209
4-2.7. Multiple burnable absorbers for programming of multiplication factor.
The previous sections of this chapter demonstrate that the self-shielding effect
in burnable absorber pins extends the effective gadolinia burnout time. The use of
multiple burnable absorber concentrations in a fuel assembly segment allows a
programming of the multiplication-depletion characteristics of the fuel assembly. In
particular, the combination of the small concentration, “fast-burning” gadolinia pins
and high concentration, “slow-burning” gadolinia pins create a superposition of the
effect.
Several series of fuel a mbly segments with multiple gadolinia
concentrations are used in the MASLWR core feasibility study. All fuel pins have an
enrichment of 8.0%. This name of the fuel assembly segment gives some parameters
of segment design. For example, the segment name: “GXY_NxNy” indicates a
segment with gadolinia burnable poisons, with two concentrations of gadolinia X%
and Y%. This fuel segment contains Nx fuel pins with gadolin oncentration X%
and Ny fuel pins with gadolinia concentration Y%. Different combinations of X,Y, Nx
and Ny were considered during this research.
Two series of assemblies were chosen for the MASLWR design feasibility
study: G49_N4N9 (G49_0808; G49_0812; G49_0816; G49_0820; G49_0828) and
G59_N5N9 (G59_0808; G59_0812; G59_0816; G59_0820; G59_0828). The
arrangement of the fuel pins was designed thro h tions of the
“FA_8GSM1” and “FA_8GSM2” geometries discussed above. Graphs of the
multiplication factor as a function of depletion
2.7.2.
These segments were created because of
have a nearly constant multiplication factor in a range of k∞ = 1.20-1.25 during the
depletion period 0-30 MWD/kgHM. The graphs demonstrate that G59_0812 satisfies
this requirement.
sse
ia c
ug some minor modifica
are presented in figures 4-2.7.1 and 4-
the need for fuel segments that could
210
Figure 4-2.7.1
Figure 4-2.7.2
211
4-2.
The implementation of burnable absorbers is essential in order to compensate
initia and
technologies are currently available on the market. MASLWR does not belong to one
of the world’s biggest fuel and reactor vendors (like AREVA, GE, Westinghouse); it
is, however, necessary to use burnable absorber technology that can be provided by
several of these suppliers. Burnable absorbers based on gadolinium oxide mixed
uniformly with fuel in a fuel pellet can be produced by a variety of vendors, so this
technology was chosen as a baseline for the MASLWR core design feasibility study.
Two gadolinium isotopes, Gd-155 and Gd-157, have absorption cross-sections
several orders of magnitude greater than that of uranium-235, but their daughter
products have relatively low absorption cross-sections. This is why 1) gadolinium
burns out faster than uranium and 2) gadolinium-doped fuel has a great self-shielding
effect.
Gadolinium self-shielding increases the gadolinium effective burnout time
when larger gadolinium concentrations are used. The use of fuel pins with differe
gadolinium concentrations in the same fuel assembly allows programming of the fuel
burnup p during
burnup.
he great neutron absorption of gadolinium-doped pins depresses the neutron
flux ar
emblies with built-in shadowing features (called here V-type and W-
he fuel assembly, which may
cause an increase in pin-peaking factor compared to non-poisoned fuel, or fuel with
evenly distributed burnable absorbers. At
8. Conclusions
l reactivity reserves. Different types of burnable absorbers materials
nt
attern and programming of the multiplication capabilities of the fuel
T
ound the pin. Combinations of gadolinium-doped pins may lead to shadowing
of the other fuel pins and control rods. Use of this shadowing effect can extend the
effective burnout time for gadolinium-doped pins in a fuel assembly, and flatten the
multiplication factor curve.
Fuel ass
type) also transform the neutron flux distribution in t
the same time, this shadowing saves some
uranium in poisoned pins as reserves for use at the middle and end of cycle.
212
n a decrease of the
reactivity control mechanism’s efficiency. The understanding of these effects is
necessary for high-quality reactor core design.
The increased enrichment fuel, operational conditions and
spec an
envelope of the design parameters that is different from traditional PWR or BWR
designs. This motivates the study of the fuel behavior and behavior of the burnable
absorbers in fuel up to 8.0% enrichment in a MASLWR operational conditions. The
data generated in this study creates a basis for a core design feasibility study and may
be useful for other designs of small LWRs with extended core life and increased
enrichment, such as modifications of IRIS, CAREM, and SMART [27].
The transformation of the neutron flux in a fuel assembly with burnable
absorbers affects the efficiency of the reactivity control mechanisms such as control
rods or soluble boron. These effects do not necessarily mea
of the MASLWR
ific requirements, such as off-site refueling and 5-year core lifetime creates
213
– Reflector Studies
on a
Fuel Assembly level with neutron transport code CASMO-4E. These studies
considered a fuel assembly in a reflective boundary conditions,
which age,
or effects on the boundary of the different fuel assemblies and core surrounding
structures were not yet considered. The next chapters will be dedicated to the core
design using 3D diffusion code SIMULATE-3 and 2D transport code CASMO-4E.
The logic of the study will be a following: First we will discuss the choice of
the proper reflector. These studies could be done for a core loaded with a non-
poisoned fuel. At that moment, the great details of the fuel loading pattern and
campaign strategy and fuel load optimization do not have a principal impact on the
decision-making. Than we will consider a several steps for improvement of core
parameters with through 3D burnable poison load profiling, and will discuss the issues
related with reactor safety in control associated with 3D core design.
Chapter 4-3 Results
The previous two chapters were dedicated to the design studies performed
2D geometry with
is analog of the infinite lattice reactor. The effects related with neutron leak
214
4-3.1. Basic requirements for a neutron reflector
There are many requireme e core reflector design. Some of
them are more general, others are sp ogies,
manufacturing, construction and maintenance requirements. Here, we discuss some
basic requirements for the neutron reflector design, which are important for the
preliminary design phase and choice of reflector.
4-3.1.1. Neutron scattering and slowing down
nts applicable to th
ecific to current reactor technol
order to be a good reflector, the reflector material should have a large
neutron elastic scattering cross-section. It clear that the scattering cross-section is a
eflector material should have a
large sc
urn neutrons with a lower energy, and it is
possibl
4-3.1.2. Neutron absorpt
In
function of energy, so it is preferable that the neutron r
attering cross-section in a wide range of energies.
During a scattering event, the neutron will likely lose a fraction of its energy.
The fraction of energy that a neutron loses per scattering is greater for lighter atoms.
Reflectors based on light atoms will ret
e to have an increase of the epithermal neutron flux near reflector. Reflectors
with heavy atoms will not slow down neutrons as efficiently as light atoms. However,
the heavy-atom reflector could flatten the fast neutron flux in the reactor core.
ion in reflector
While the main purpose of the neutron reflector is neutron economy, the
reflector atoms should have a relatively low neutron absorption cross-section. In some
cases, the reflector may contain a combination of atoms, like multi-atomic molecules
or alloys, and macroscopic scattering and absorption cross-sections should be taken
into account. Some reflector materials require cladding in order to protect them from
contact with coolant or other structural materials. In this case, the shielding effect of
the cladding material should also be considered.
215
4-3.1.3. Reflector weight and geometry
Engineering considerations impose limits on the size and weight of the
reflector. The technology of the reactor may require channels for coolant and control
rod drives through the reflector, as well as cavities and channels for the detectors and
other I&C equipment.
The lower axial reflector should provide enough flow area for the coolant. Also
it shoul
a reasonable weight of upper reflector and avoid weight loading of the fuel
assemblies during normal operation and in case of accident.
The radial reflector should also have a reasonable size and weight. It also
should fit the reactor technology and arrangement of reactor internals. In the
MASLWR reactor, possible space for an axial reflector is the space between the core
baffle and core barrel. Additional reflector pads are possible in the down chamber as
long as they do not affect the natural circulation flow rate dramatically. The reflector
design criteria, technical specifications and acceptance criteria should be discussed
and coordinated with all designers involved in the reactor core and vessel design.
4-3.1.4. Reflector radiation heating
d not increase friction or create disturbances or oscillations of the coolant flow.
The upper reflector should provide enough flow area for coolant flow and additional
area for control rods and instrumentation channels. It also should not increase friction
or create disturbances or oscillations of the coolant flow. Also it is important to
maintain
The interaction of neutrons and absorption of other forms of radiation may lead
to the heating of the reflector material. There are several mechanisms which play a
principal role in the reflector radiation heating: neutron and gamma absorption,
neutron and gamma scattering and deposition of the absorption reaction products’
energy (like absorption of secondary gamma, produced in n,γ-reaction).The radiation
heating of the reflector requires a proper reflector cooling mechanism. Cooling
channels or other heat exchange systems will be needed.
216
4-3.1.5. Radiation effects on a materials and reflector maintenance
The interact ns, gammas) may
result in changes in reflector properties, structural integrity, chemical reactions or
other effects. Some well-known effects are stainless steel embrittlement, swelling, and
radiation g
4-3.2. Neutron Reflector Materials
Several materials are traditionally used as neutron reflectors in a reactor:
4-3.2.1. Water
ion of the reflector material with radiation (neutro
rowth. Another effect kind of effect is related to the accumulation and
redistribution of “Wigner’s energy” in graphite [487].
Proper reflector maintenance is also needed during the reactor cycle and at the
end of plant life, due to accumulation of radiation defects, activation of the reflector
atoms and environmental effects (such as interaction with the coolant).
Water is a universal material that could be used as moderator, coolant and as a
reflector. The scattering cross-section of hydrogen, and its ability to slow neutrons
rapidly down to thermal energy makes water a cheap and obvious choice for the
reflector. Disadvantages of water include its relatively low density, and as a liquid, it
cannot perform structural functions.
4-3.2.2. Graphite
Graphite is widely used in research and power reactors (RBMK, CGR, AGR)
as a neutron reflector. Graphite has a smaller absorption cross-section than water. Its
average neutron energy loss per interaction is smaller than for hydrogen, but is still
very large, which makes graphite a good moderator. A graphite reflector also converts
fast neutrons to thermal, which increases the thermal flux inside reflector and near to
it.
217
Disadvantages of graphite include: 1) the return of fast neutrons from the
graphite reflector to the reactor is relatively small, 2) accumulation of the “Wigner’s
later be released in an
which
of graphite over the long operation period. Proper cladding
or insu
energy” under certain irradiation conditions, which could
unpredictable manner, and 3) chemical interaction of graphite with reactor coolant,
could lead to a chemical reactions (due to water chemistry and chemical shim)
or mechanical degradation
lation may be necessary for graphite blocks. Graphite is not a good structural
material that can sustain heavy loads.
4-3.2.3. Stainless Steel
Stainless steel is widely used in nuclear reactors for manufacturing of the
reactor
absorption cross
section than graphite or hydrogen; however the average neutron energy loss per
intera rons.
That is why stainless steel may be used to flatten the neutron flux in a small
MASLWR core. Stainless steel can also be used as a structural support material.
Disadvantages of stainless steel include its much greater weight, and increased
thermal neutron absorption and gamma absorption, which may lead to significant
reflector heating. Other disadvantages of stainless steel include embrittlement,
swelling, and radiation growth and accumulation of radiation cause
activation.
4-3.2.4. Beryllium
internals and structural elements. In PWRs, it is used for the core baffle and
core barrel. Some vendors also use stainless steel blocks as a reflector or a monolithic
baffle-reflector block. Stainless steel has a larger thermal neutron
ction is not as high, so stainless steel could be a good reflector for fast neut
d defects and
Beryllium is well known in the nuclear industry as a good reflector for fast
neutrons. It was widely used in early research reactors and critical assemblies.
However, this material is very expensive and not used in the light water reactor
industry. While one of the main design requirements is use off-the-shelf equipment,
beryllium is not likely a good choice for MASLWR.
218
4-3.2.5. Uranium (Depleted Uranium)
As a heavy atom, uranium may be a good reflector for fast neutrons. However,
use of depleted uranium in axial blankets (reflectors) will lead to the production of
plutonium, which could create a threat to a non-proliferation regime. Depleted
uranium could also be used as an axial reflector in a form of fuel pellets placed in the
bottom nd the top of the fuel rod. a
4-3.2.6. Thorium
The use of thorium is a similar to use of depleted uranium, as a heavy atom,
thorium may be a good reflector for fast neutrons. However, use of thorium in axial
blankets (reflectors) will lead to the production of U-233, which could create a threat
to a non-proliferation regime. Thorium could also be used as an axial reflector in a
form of fuel pellets placed in the bottom and the top of the fuel rod.
The use of thorium is also related with a radiation safety risk. The by-product f
the neutron capture reaction (U-232) has a decay mechanism associated with emission
of high-energy gamma. Currently thorium is not used in commercial industry as a
blanket or reflecting material.
4-3.2.7. Other Materials
Other materials may be used as a neutron reflector. However, the
implementation of these materials may create licensing difficulties due to the necessity
of additional testing and studies.
4-3.2.8. Choice of reflector material
For the current design, it would be reasonable to consider a limited number of
reflecto
tor
and structures, and uranium pellets can be used as built-in reflectors for the fuel
assemblies.
r candidates and geometries: standard PWR water reflector, graphite reflector
and stainless steel reflector. Stainless steel can be considered as a part of the reflec
219
4-3.3. Neutro
xial Reflector
n Reflector Geometry
4-3.3.1. A
lector could have several li tions related with
natural circulation driven flow.
re structures, and CASMO model for axial reflector
The geometry of the axial ref mita
reactor technology and
Figure 4-3.3.1. Upper-co
At the current design stage, the following two-component approach may be
implemented for the axial reflector:
Component 1 – Axial fuel blankets with depleted uranium
Component 2 – Axial stainless steel structures, like lower and upper core
support plates
220
or uranium blankets it is reasonable to consider reflector thickness, or the
height of the additional depleted fuel pin in the range of 8-16 cm above and below the
be formed by a combination
of the
pleted uranium blankets)
and 18
F
active part of fuel rod.
The second component of the axial reflector will
fuel assembly upper and lower heads and upper and lower spacer plates. The
thickness of the upper plates and coolant flow area should be calculated with
involvement of mechanical designers (for weight distribution and other related issues)
and thermal hydraulic specialists (for optimization of the flow area and coolant
friction issues).
In the current research, a thickness of the upper core plate and lower plate with
fuel assembly heads was taken equal to 10 cm, and the coolant flow area is equal to
the flow area of the fuel assemblies. The non-fueled part of the fuel pins were also
considered equal to 10 cm (for a fuel assemblies with axial de
cm (for a fuel assemblies without axial depleted uranium blankets), unless fuel
design specifications are not finished.
4-3.3.2. Radial Reflector
Space for the radial reflector is limited by the MASLWR modular approach.
Becaus r dimensions is
unacce le. Cu
baffle and core b
steel blocks of m lector.
steel pads could be used in the MASLWR core, outside of the core barrel. This type of
reflecto pansio use graphite
pads will be inef
by gamma or ne re design of reflector cooling channels or
other cooling mechanisms.
e the reactor module is movable, any increase of the reacto
ptab rrently the radial reflector is located in the space between the core
arrel, which could be filled with water, graphite blocks or stainless
onolith ref
In order to increase the reflector efficiency for corner fuel assemblies, stainless
r ex n is reasonable only for the stainless steel reflector, beca
fective because of shielding by the core barrel. The reflector heating
utron radiation may requi
221
r internals in a mid-core cross-section plane
Figure 4-3.3.2 Reacto
Figure 4-3.3.3 Radial reflector options
Option 1 Option 3 Option 4
tion 2
cooling channels and reflector pads
Op
List of options
1. Standard PWR baffle 2. Graphite Reflector 3. Steel Reflector 4. Steel Reflector with cooling channels 5. Steel Reflector with
Option 5
222
4-3.4. Reflector Studies with SIMULATE-3
4-3.4.1. The SIMULATE-3 model.
The first set of reactor studies were focused on an investigation of the best
choice of the reflector from a neutron economy prospective. These calculations were
performed with CASMO-4 and SIM SMO-4 was used for a creation of
the fuel assembly segments and reflector segments in a 2D geometry, using the
solution of the neutron transport equation. SIMULATE-3 was then used for 3D core
simulat
rom 0 to 50 MWD/kgU (core average). Three basic configurations were
considered for the reflector studies, modeling the geometries shown in figure 4-3.3.3
as Option 1, Option 2 and Option 3.
Figure 4-3.4.1 Simulate-3 segments assignment
ULATE-3. CA
ions, and calculation of the core multiplication factor and fuel depletion for
every node using a 3D nodal diffusion model.
Reactor calculations were performed with the SIMULATE-3 [236] computer
code, assuming quarter core symmetry with reflector regions. Each fuel assembly was
represented as a composition of 40 axial and 4 radial nodes. Calculations were
performed at “100% power”, “100% coolant flow”, “control rod-out” and “no boron”
conditions. In total, 99 depletion steps were performed for each case, covering fuel
depletion f
223
4-3.4.2. The SIMULATE-3 Simulation Results.
The duration of the core campaign was determined according to two simple
criteria: 1) The multiplication factor equal to 1; and 2) One or more fuel assemblies
reach the burnup limitation of 60 MWD/kgHM.
Table 4-3.4.2. Summary of Simulate 3 reflector studies
Campaign Length Reflector Case MWD/kgU* Years**
Limiting Factor
A (Water) 41.50 4.2 keff = 1 B (Graphite) 47.00 4.8 keff = 1 B (Graphite) 46.00 4.7 Burn-up limit reachedC (S. Steel) 49.00 5.0 keff = 1 C (S. Steel) 47.50 4.8 Burn-up limit reached* Core average exposure, ** Effective full power years
he fuel burnup represents the efficiency of the fuel use, so the burnup
distribution calculations were performed for three types of reflectors at core average
/kgHM. The fuel burnup distribution
was als
T
burnup equal to 20 MWD/kgHM and 40 MWD
o calculated for the end of the cycle (according to Table 4.1).
The burnup distribution at MOC and EOC is presented in figure 4-3.4.2 (A-D)
Figure 4-3.4.2 A Burnup distribution in the MOC (Burnup 20 MWD/kgHM)
224
Figure 4-3.4. WD/kgHM)
2 B. Burnup distribution in the MOC (Burnup 40 M
Figure 4-3.4.2 C. Burnup distribution in the EOC (EOC Criterion K=1)
Figure 4-3.4.2 D. Burnup distribution in the EOC (EOC Criterion burnup limit)
* This EOC criterion is not applicable for standard PWR baffle case
225
Another representation of this data is given on the figure 4-3.4.3 (A-B).
Figure 4-3.4.3 A: Burnup distribution at MOC
As we can see from this figure, the use of graphite and stainless steel reflector
allow a better (flatter) distribution of the burnup, so the fuel in the outer fuel
assemblies has a greater burnup in a case of solid reflector, than in case of a standard
PWR baffle case.
226
Figure 4-3.4.3 B bution at EOC
: Burnup distri
These simulations demonstrate that use of a stainless steel reflector provides
the maximum reactor campaign length. All subsequent studies were performed
assuming a stainless steel reflector.
227
4-3.5. Reflector Studies with CASMO-4E
4-3.5.1. The CASMO-4E model.
ctor cooling. The solution of
the gam
he simulation was performed with the neutron and photon transport code
CASMO-4E. Reactor internals, such as core baffle, core barrel, reflector, and reflector
mbination of individual segments. The segment breaking
and me
(see figure 4-3.5.1):
Up to 5% of all energy produced in a conventional PWR reactor is generated in
the coolant and structural materials due to neutron absorption and moderation, and
gamma attenuation. Use of the massive stainless steel reflector in a reactor with
natural circulation could create issues related with refle
ma transport equation in addition to neutron transport is necessary in order to
evaluate the importance of the reflector radiation heating. The evaluation of radiation
heating was performed for two fuel enrichments 4.95% and 8.0%, and for five options
of the reflector geometry.
T
pads were simulated as a co
shing inside the segments are presented in the drawings below. According to
these drawings, CASMO-4E input files were prepared.
The following CASMO-generated inputs were used to approximate the quarter
core geometry
228
Figure 4-3.5.1 CASMO-4E input geometries
1 2
4 3
List of options
1. Standard PWR baffle 2. Graphite Reflector
5
channels and reflector pads
3. Steel Reflector
Reflector with cooling
4. Steel Reflector with cooling channels
5. Steel
229
4-3.5.2. The CASMO-4E Multiplication Factor Calculations.
The results of these simulations provide multiplication factors as a function of
reflector geometry and fuel initial enrichment (see Table 4-3.5.2). The maximum
value f
Additio
or the multiplication factor occurs for the stainless steel reflector. However,
adding the cooling channels in the reflector geometry decreases its efficiency.
nal reflector pads outside of the core barrel add some more efficiency to
stainless steel reflector.
Table 4-3.5.2 Multiplication factor as a function of reflector and fuel enrichment
Fuel Enrichment Reflector Option 4.95% 8.00% 1. Standard PWR baffle 1.35155 1.41347 2. Graphite Reflector 1.36681 1.43039 3. Steel Reflector 1.36892 1.43262 4. Steel Reflector with cooling channels 1.36766 1.43126 5. Steel Reflector with cooling channels and reflector pads 1.36794 1.43158
The following segment assignments and boundary conditions were used in the
CASMO 4E model for the energy deposition study:
r BC
Mirr
Figure 4-3.5.2 CASMO-4E Segment assignment and boundary conditions
or BC
Mirr
o
13 14 15 16
9 10 11 12
5 6 7 8
1 2 3 4
Black BC
Bla
ck B
C
The fuel segments are marked with green and they are not currently of interest.
The gray segments are reflector segments, and the deposited energy in these segments
is the goal of these calculations. Due to symmetry, we consider only segments
1,2,3,6,7, and while segments 8,11,12 and 16 are symmetrical to 3,6,2 and 1
respectfully.
230
Table 4-3.5.3 Energy deposition in full-height core reflector by gammas
Table 4-3.5.4 Energy deposition in full-height core reflector by neutrons
231
Table 4-3.5.5 Total energy deposition in full-height core reflector
Due to the current geometry and segment assignment, reactor vessel walls and
water in down chamber and other reactor internals, which are not part of the reflector,
are not included in the survey. The contribution of those components to the reflector
heating is low and their effect on multiplication factor is almost negligible.
4-3.5.3. Discussion of CASMO-4E Results
The total energy deposited in the reflector segments is nearly the same for al
options; however, the distribution of the energy absorbed by different core structures
is diffe
l
rent.
Figure 4-3.5.3 A shows that the energy deposited in reflector by neutron
thermalization is greater for the water reflector; however the use of the water reflector
leads to a smaller energy deposited by neutrons in a core barrel.
232
Figure 4-3.5.3 A: Energy deposited in full-height reflector by neutron thermalization, kW,
reactor vessel.
he second great contributor to the reflector heating is attenuation of the
gamma
The use of the cooling channels increases amount of energy deposited by
neutrons in the steel reflector area. The slight decrease of the energy deposited in
outside water in Option 5 caused by the fact that reflector pads replace some volume
of that water. Another conclusion that we could make from this graph could be made
regarding absorbed dose outside of the reactor vessel, and associated shielding. The
water reflector is allowing reduction of the neutron share into the dose rate outside
T
radiation. Figure 4-3.5.3.B shows that the use of a stainless steel reflector
leads to better absorption of photons energy, and a resulting smaller photon energy
flux on the core barrel and reactor vessel.
233
Figure 4-3.5.3 B: Energy deposited in full-height reflector by gamma attenuation, kW,
Figure 4-3.5.3 C: The contributors to total energy deposited in full-height reflector, kW,
Figure 4-3.5.3 C shows that the attenuation of prompt gamma makes is a major
contributor into a reflector and reactor structures heating. While the energy deposited
by neutrons in water reflector is significantly higher than for other cases, the overall
role of neutrons into energy deposition is less than 20%.
234
radiation types (see figure 4-3.5.3 D). As
we can
nto graphite reflector heating is almost similar.
The heating of the stainless steel re caused by attenuation of photons.
The addition of the cooling channels increases the amount of energy deposited into
stainless steel reflector by prompt gamma.
Finally, in order to evaluate reflector options, we need to evaluate the energy
deposited in a reflector material by different
see from the figure the total energy absorbed in a reflector material that occupy
space between core baffle and barrel, is smallest for the graphite reflector. The major
contributor for the reflector heating of water reflector is neutrons thermalisation. The
contribution of neutrons and gamma i
flector is mainly
Figure 4-3.5.3. D: The contributors to total energy deposited in reflector material, kW
235
4-3.6
core, fuel
burnup in the outer fuel assemblies could increase by up to 150% compared to the
standar
core barrel, but design of the reflector cooling system is
out of the scope of the current report.
For the further calculations the stainless steel reflector with reflector pads will
be used as a base line preferred option for neutron economy and fuel utilization
prospective. The same argument for use of stainless steel reflector was also made in
other reactor designs [3, 27, 133, 138]. The use of the reflector pads is necessary for
the MASLWR core in order to improve fuel utilization of the outer ring of fuel
assemblies (especially fuel utilization in the pins on the corners of the assemblies).
In order to simplify the neutronic modeling of the MASLWR core at the stage
of feasibility study, the bypass flow and reflector heating contribution into energy
generation would be considered as insignificant (only at current stage), however it
should be considered at the stage of the detailed core design (the next stages of
design).
Conclusions for Reflector Studies
A variety of reflector materials are considered for the MASLWR reactor:
water, graphite and stainless steel. Different geometries of the reflector are possible.
These reflector options have different advantages and disadvantages, so the proper
choice of reflector is a complicated task.
The use of a stainless steel reflector may provide improved fuel usage in a
small reactor core, where neutron leakage is significant. For the MASLWR
d PWR water reflector.
Heating rates in the stainless steel reflector and structures could be as high as
580 kW, requiring proper heat removal mechanisms. Also different reflector materials
provide different levels of protection of the reactor vessel against neutrons and
photons, which could affect the reactor vessel lifetime and plant lifetime management.
A heat removal mechanism should be designed for the reflector in order to avoid
reflector melting and degradation. The current calculations estimate the heat
generation in the reflector and
236
Chapter 4-4 Results – Core Design
I In previous chapters we have considered specific issues related to the use of
8.0% enriched fuel in MASLWR operational conditions, the effect of using
gadolinium burnable absorbers and the impact of the neutron reflector on the fuel
economy and reactor behavior. These studies permit the creation of a library of fuel
assembly and reflector segments. These segments will be used by the neutron
diffusion code SIMULATE-3 for the 3-D core calculations.
237
4-4.1. MASLWR core model in SIMULATE-3
4-4.1.1. Geometry and nodalization
Core calculations with SIM rformed for two geometry types:
full core and quarter core.
The active core height was chosen to be 160 cm. The core was divided into 40
axial segments (4 cm per segment). A simplified model with 20 axial segments (8 cm
per segment) was also used for some intermediate calculations. Axial reflector
segments were identified for each fuel assembly to formulate boundary conditions for
the top and bottom of the reactor core.
The large radial power and burnup gradients determine the necessity of four
radial segments per fuel assembly for final calculations and pin-power reconstruction.
The model of fuel assemblies and full core is shown in figure 4-4.1.1
Figure 4-4.1.1 Simulate-3 model nodalization (fuel assembly and core)
ULATE-3 were pe
238
e full core model.
The full core model was also used for a control rod worth calculations with
SIMULATE-3.
4-4.1.
Six fuel assemblies comprise the quarter core model, with reflective boundary
conditions on axes of symmetry. This model was used for burnable absorber
placement studies. The final detailed study was performed with th
2. Thermal hydraulic conditions
The specific feature of the natural circulation system is dependence of the core
flow rate on a core power. This relation is non-linear, and depends on many factors,
such as geometry of the reactor core, upper core space (raiser), and heat exchangers
and down chamber. It also depend on operational conditions for the primary and
secondary circuit. The detailed for a power flow correlations are usually unavailable at
the stag
t channel were not considered
in our
analysis. While the
real power/flow correlations were unavailable at the current design stage, the
evaluation of the parameters was performed either for “full-power” and “maximum
core average flow” or “zero-power” and “zero flow” conditions.
The operational conditions of the MASLWR reactor, were chosen according to
a MASLWR report [104] with a correction [108] and are given in the table 4-4.1.2:
e feasibility study. It leads to a certain assumptions and limitation for a flow
model used in a core design calculations.
The “full-power” thermal-hydraulic conditions were used in the SIMULATE-3
model. The core flow was selected equal to the core average flow rate. The cross flow
between fuel assemblies and flow acceleration in the ho
studies. A discussion of the appropriate flow model for a natural circulation
system will be presented in later section dedicated to uncertainty
239
Table 4-4.1.2 MASLWR operational conditions
Parameter Value Units Core thermal power 150 MW(t) Core operating pressure 8.6 MPa Core inlet temperature 220 ºC Core average flow rate 1526.4
137.6 MT/hr kg/cm2 per hr
Core bypass flow fraction 0 % Core average power density 84.5 kW/liter Core volume (cold) 1775 Liters Core fueled area 11094 cm2 Number of fuel assemblies in core 24 Number of the fuel pins in core 6336 Control rod position All control rods out Soluble boron concentration for keff search No Soluble boron Criteria for boron search keff = 1 Control rod history / boron history Not accumulated SIMULATE-3 thermal-hydraulic model PWR
he PWR thermal hydraulic model was used for the MASLWR core, with 0%
core by
T
pass flow. A multiplication factor search was performed for a boron free core
with all control rods out of the core. Neither boron nor control rod history were taken
into account at this stage of the design.
240
4-4.2. M ble poisons
4-4.2.
ASLWR Core without burna
1. Calculation model and assumptions
The first attempts to model the MASLWR core were performed with MCNP
and SCALE [107, 108]. These simulations were performed for a fresh loaded core
with different initial enrichments, without depletion and without coupling of neutronic
and thermal hydraulic parameters of the core. The results allowed an initial estimate of
the multiplication factor keff and the importance of the neutron leakage.
Depletion calculations for the MASLWR core were performed and
documented in the 2003 MASLWR report [104], however these calculations were
performed for different core geometry and fuel compositions. Neutronic and thermal
hydrau
ly level) was described in previous chapters. The same major design steps
were used for the SIMULATE-3 core studies. The first set of calculations was
performed for 4.5% and 8.0% fuel in the MASLWR core without burnable poisons.
The basic assumptions in the model for these calculations are:
a) Semi-forced coolant circulation –a constant flow rate, independent
from reactor power for minor power fluctuations around nominal
power.
b) Each fuel assembly is considered as an isolated channel, ignoring
coolant cross flow.
c) The coolant flow rate is the same in all fuel assemblies.
d) Depletion is performed for full power, soluble boron free conditions
with all control rods out.
lic parameters were not coupled in these calculations.
In order to investigate the feasibility of the MASLWR core with 8.0% enriched
fuel and five effective full power years of operation, some studies were necessary in
order to build a basis for future design work. Some work performed with CASMO-4
(fuel assemb
241
4-4.2.2. Core overview
The primary focus of the current simulations was estimation of multiplication
factor and burnup pattern as a function of the time (depletion). The the thermal
hydraulic and kinetic parameters were evaluated for these case studies, in order to
understand magnitude and nature of reactivity effects in a MASLWR core and
perform comparison with fuel assembly level estimation described above.
Figure 4-4.2.1
The multiplication factor as a function of time is presented in figure 4-4.2.1.
The use of 8.0% fuel in the MASLWR core with a stainless steel radial reflector
allows a doubling of the reactor campaign length compared to 4.5% enriched fuel. The
initial r
f 8.0% fuel yields a slightly smaller radial and
eactivity reserve for burnup is significantly higher in 8.0% fuel.
Figure 4-4.2.2 shows the axial, radial and nodal peaking factors for 4.5% and
8.0% fuel. The increase of the fuel enrichment does not lead to any significant changes
in power peaking factors. The use o
242
nodal peaking factor at the BOC, but slightly greater peaking factors at the MOC. At
the EOC the peaking factors are comparable for 8.0% and 4.5% fuel. The values of the
power peaking factors decrease with burnup, so the BOC is an area of interest for
safety analysis.
Figure 4-4.2.2
The power generation profile changes with core depletion and fuel burnup. The
simplest way to describe the axial power generation profile is to consider the axial
power offset. A positive value indicates an upper peaked core and a negative value
indicates a lower peaked core. The power generation axial offset is presented in figure
4-4.2.3. At the BOC, the power generation profile is shifted towards the bottom of the
core, where the lower moderator temperature provides better neutron moderation. The
value of the offset is approaches zero with depletion, and becomes slightly positive at
the EOC.
243
Figure 4-4.2.3
4-4.2.3. 8.0% Core thermal hydraulic
SIMULATE-3 allows the reconstruction of the three-dimensional heat
generation profile, using the relative power distribution in the fuel assemblies. The
relative power generation profiles are a function of burnup. During the research work
on prototypical core design, th ower generation profiles were
considered for every depletion step of all considered cases. Here, we present only the
relative power profiles for BOC, MOC and EOC conditions in figures 4-4.2.4 A-D.
At BOC, the relative power of the center fuel assembly (FA 1-1) is significantly higher
than other fuel assemblies and the curve has a typical bottom-centered cosine shape.
At the MOC the shape of the curve has a plateau near the core center. At the EOC the
curve has two maxima at the lower core and upper core, caused by significant fuel
burnup in a center of the core.
e 3-D relative p
244
Figure 4-4.2.4 A: Fuel assembly names assignment
Figure 4-4.2.4 B:
Figure 4-4.2.4 C:
245
Figure 4-4.2.4 D:
These graphs show that the burnup of the fuel in the center of the core has a
great importance for the relative power fraction (RPF) distribution. It is reasonable to
expect greater burnup in a center of the core for a core without burnable absorbers.
The burnup distribution at the EOC is shown in figure 4-4.2.5.
Figure 4-4.2.5
246
While the average burnup for the center fuel assembly is 62 MWD/kgHM, the
center part of this fuel assembly has a burnup around 70 MWD/kgHM. The outer fuel
assemblies and fuel assembly edge pins have a relatively small burnup.
At the next design step, it would be rational to choose the burnable absorber
loading pattern to increase fuel usage (or burnup) of the outer fuel assemblies.
The large values of the RPF in the center fuel assembly raises safety issues
related to heat removal from this fuel assembly with a relatively small natural
circulation coolant flow rate. In order to investigate the potential safety issues, we
consider coolant temperature changes over the length of the fuel assemblies (modeled
as insulated channels with the same coolant flow rate in each fuel a bly). The
coolant temperature distribution is presented in figure 4-4.2.6.
Figure 4-4.2.6
ssem
The calculations show that the MASLWR core will have coolant bulk boiling
in the center fuel assembly at BOC at a height equal to 120 cm, or ¾ of the core
height. Usually bulk boiling means zero safety margins and leads to a resetting of the
core design parameters. However, this is feasibility study and the problem could be
247
reformu
ssemblies?” or “Are there any reserves of the coolant
flow th
lated in the following way: “Is it possible to design a burnable absorber
loading pattern that will minimize power peaking, and redistribute heat generation
more evenly towards other fuel a
at could allow better heat removal without boiling?”
Figure 4-4.2.6 shows that the coolant temperature of neighboring fuel
assemblies is lower, and the coolant temperature of the outer fuel assemblies is
significantly lower than the central assemblies. At constant pressure, the coolant
density is strictly related to its temperature. The coolant density is the main driving
force in natural circulation driven systems. The 3-D coolant density distribution is
shown in figure 4-4.2.7.
Figure 4-4.2.7
This figure shows there is a large radial density gradient between the center
and ou
de n
assume that cross-flow between fuel assem lies may significantly increase the mass
flow rate in the upper part of the center fuel assemblies.
ter fuel assemblies. The value of this gradient is about ½ of the axial core
nsity gradient, which creates the driving force for the natural circulation. We ca
b
248
The discussion of natural circulation mechanisms is out of scope of this
dissertation, and we remark that the availability of appropriate coupled neutronic and
thermal hydraulic calculations for this natural circulation system with cross-flow
would significantly simplify further calculations, and help to resolve many safety
related issues.
4-4.2.4. Core kinetics
The influence if the reactor thermal hydraulics and operational parameters onto
reactor kinetic, reactivity feedbacks and RCS worth was discussed in chapter 4-1 for
infinite lattice geometry. In order to be consistent, we consider kinetic parameters and
reactivity coefficients for the finite MASLWR geometry for 4.5% and 8.0% initially
enriched fuel.
Figure 4-4.2.8 Effective fraction of delayed neutrons
The curves of effective fraction of delayed neutrons for the finite core (see
figure 4-4.2.8) look similar to the curves presented in chapter 4-1. However a slight
decrease of effective fraction of delayed neutrons occurs for the finite geometry case.
This difference is caused by several facts. First, Non-uniform power distribution and
249
non-uniform burnup leading to th er fuel assemblies have a greater
burnup, greater plutonium concentration and lower effective fraction of the delayed
neutrons, comparing to the fuel assemblies at the edge of the core. Second, the leakage
from the outer fuel assemblies is greater, which is applicable for both prompt and
delayed
The reactivity feedback oderator temperature could be
characterized through 3 different reactivity coefficients:
1) The isothermal temperature coefficient (ITC) is the reactivity change
associated with a uniform change in the fuel and moderator inlet temperatures divided
by the change in the averaged moderator temperature. [236]
2) The moderator temperature coefficient (MTC) is the reactivity change
associated with a change in the moderator inlet temperature divided by the change in
the averaged moderator temperature. [236]
3) The inlet temperature coefficient is the reactivity associated with a uniform
change in the inlet temperature divided by the requested change in temperature. [236]
All these three coefficients are useful for the reactor design and safety analysis,
and represent the different mechanism or the reactivity insertion associated with
moderator temperature changes. The results of the moderator reactivity coefficients for
MASL
e fact that cent
neutrons. Finally due to different burnup levels of the fuel assemblies and
neutron leakage, a spectrum in a center of the core is different from the spectrum on a
core periphery.
related with m
WR core at HFP conditions, loaded with 4.5% and 8.0 enriched fuels are shown
at figure 4-4.2.9. As we can see from this figure, MTC is always negative for a HFP
conditions, for boron concentrations 0 ppm and 800 ppm. The magnitude of the MTC
for MASLWR core is higher for finite size core comparing to for infinite lattice
(chapter 4-1).
250
Figure 4-4.2.9 Reactivity coeffi with moderator temperature.
cients associated
251
for a MASLWR core, than the results of uniform lattice calculation. It
happen
The fuel temperature feedback could be characterized by the uniform Doppler
coefficient, which is the reactivity change associated with a uniform change in the fuel
temperature divided by the change in the averaged fuel temperature. [1, 236] The
results for uniform Doppler coefficient are shown in figure 4-4.2.10. The magnitude of
UDC is higher
ed because the power generation and maximum fuel temperature occurring in a
center fuel assemblies, which has a burnup above core average. So we can mention
that value of UDC at 40 MWD/kgHM core average is equivalent to the value of fuel
temperature coefficient at 55 MWD/kgHM for uniform infinite lattice. This figures are
equivalent to the average burnup of the center fuel assembly for a given core average
burnup.
Figure 4-4.2.10
The power coefficient is the reactivity change associated with a uniform
change in the power level divided by the percent change in power. The power
distribution used to evaluate cross sections is unchanged [236]. The power reactivity
coefficient was evaluated for full power conditions, with assumption that there are no
subsequent changes of the coolant flow rate, and flow rate remain constant and equalt
to the nominal flow rate (see figure 4-4.2.11).
252
Figure 4-4.2.11
As we may see from the figure, the core loaded with 8.0% enriched fuel has a
lower power reactivity coefficient, than core loaded with 4.5% enriched fuel. This is
caused bur a cumulative effect of the moderator and fuel temperature reactivity
coefficients, which are also smaller for 8.0% core than for 4.5% core. The use of the
soluble boron decreases the magnitude of the power reactivity coefficient, however it
remains negative, which is good for reactor safety. Aslo we can conclude that the
magnit
[104], t
ure 4-4.2.12. the
evaluat
ude of the power reactivity coefficient at the EOC (Keff=1) for 4.5% and 8.0%
core is equal.
While the flow rate of the reactor with natural circulation of a coolant is
smaller than for PWR with a forced circulation, and oscillation of the flow is possible
he evaluation of the flow reactivity feedback is extremely important. The flow
coefficient is the reactivity change associated with a uniform perturbation of the flow
density divided by the percent change in flow [236]. The result of the flow reactivity
coefficient evaluation for 4.5% and 8.0% core are given at the fig
ion were performed for a constant power conditions, for 2 soluble boron
concentrations: 0.0 ppm and 800 ppm.
253
Figure 4-4.2.12
The values of the flow reactivity coefficient are positive for 4.5% and 8.0%
core. It is mean, that increase of the coolant flow will lead to a positive reactivity
insertion, while the decrease of the flow could lead to the negative insertion of the
reactivity. The main reason for such feedback is a coolant temperature rise along the
core height and relative decrease of the moderator density with a temperature increase.
This phenomenon is extremely important for moderation properties in the upper core
where sub cooled boiling may occur. The increase of the coolant flow decreases the
coolant temperature, and a chance of the coolant sub cooled boiling in the upper core.
The use of the soluble boron, leads to the decrease of the flow reactivity
coefficient. At the soluble boron concentration of 800 ppm, the increase of the
moderator density at the upper core will also increase the neutron absorption in the
upper core coolant, so the overall effect of floe increase will be smaller than for a
these experiments the flow oscillation could be ±4% of a nominal flow at HFP
boron-free core.
The safety importance if the flow oscillations could be evaluated according to
the results of the flow experiment at OSU MASLWR Test facility [104]. According to
254
conditions. These oscillations could lead to the reactivity insertions ±7.5 cents at BOC
and ±1
, so the reactivity insertions related with this fluctuations should
be eva
7.5 cents at the EOC. In terms of the total reactivity uncertainty window these
figures will be 15 cents at BOC and 35 cents at EOC.
The coolant density and saturation boiling temperature are also depend on the
primary circuit pressure. The fluctuations of the coolant pressure at the MASLWR
core are also possible
luated. The pressure coefficient is the reactivity change associated with a
perturbation of the primary system pressure divided by the pressure change. The
results for the pressure reactivity coefficients calculations are show at figure 4-4.2.13.
Figure 4-4.2.13
The pressure reactivity coefficient is positive for both 4.5% and 8.0% core,
operated boron-free, so the increase of the pressure lead to increase of the moderator
density and a positive reactivity insertion. For a soluble boron concentration of 800
ppm, the magnitude of pressure reactivity coefficient is smaller than for a boron free
core, and could be even negative at the BOC for 4.5% core. The phenomena that
defines a pressure reactivity feedback is similar to a flow reactivity coefficients, and
defined by a moderator density in the core and relative amount of boron in the coolant.
255
ese
assump ons, with a remark about calculation conditions.
The power defect is the reactivity defect associated with a change in power
level from reference to the perturbed power. Both moderator and fuel temperatures
were changed based on the perturbed power. For a MASLWR core we consider a
power defect from HFP conditions to HZP conditions, assuming a nominal flow rate at
HZP conditions. These choice is caused by the fact the PWR perturbation model used
in SIMULATE is created for PWR reactor with a forced circulation, so while the
power-flow correlation for a MASLWR reactor is not developed, we may use th
ti
Figure 4-4.2.14
The power defect for the 8.0% core is smaller than for 4.5% core at the same
depletion step, and same soluble boron concentration. The presence of the soluble
boron decreases the power defect. At the MOC (depletion between 10 and 40
MWD/kgHM) the power defect of the 8.0% boron free core is equal to a 4.5% core
with soluble boron concentration of 800 ppm. The values of the power defect
evaluated for a 3D MASLWR core is ~20% lower then the figures of the power defect
evaluated for the infinite lattice in chapter 4-1. The main causes for such difference is
the 3D effect of the fuel an coolant temperature distribution and different depletion
rates for a center and outer fuel assemblies.
256
The boron coefficient is the reactivity change associated with a uniform
perturbation of the boron concentration divided by the boron change. The soluble
boron worth was calculated for hot full power conditions, for a 2 soluble boron
concentrations 0.0 ppm and 800 ppm.
Figure 4-4.2.15
The values for the soluble boron worth calculated for a 3D MASLWR core
average depletion range of 0 –50 MWD/kgHM are comparable to the results of the
infinite lattice calculation for the depletion range of 0 –60 MWD/kgHM. The slight
desertion of the solkuble boron worth in a 3D model is caused by neutron leakage and
related burnup and RPF distribution. Also we should remember that neutron
importance at the core edges is smaller that in a center of the core due to leakage.
4-4.2.5. Conclusion
The next stage of the neutronic calculations are focused on the design of a
burnable poison loading pattern allowing five operational year, controllable core in
operational and transport conditions, and a power generation distribution that allows
operation without coolant boiling or heat exchange crisis.
257
4-4.3. MASLWR Core with radial BP profiling
4-4.3.1. Core overview
The previous section of the feasibility study shows that high values of the RPF
in the center fuel assemblies could lead to coolant boiling and under utilization of the
outer fuel assemblies. The intuitive solution for this problem is to load the core with a
burnable absorber, with more burnable absorber in the center. However, the practical
design is more complicated. In order to choose the burnable absorber pattern we need
to take into account economic and manufacturing issues, and physical phenomena
related to use of the burnable absorbers that have been discussed previously.
rom the economic and manufacturing prospective, it is better to use standard
fuel as
described in literature [387] and discussed previously in chapter 4-2
(FA_8GSM1, FA_8GSM2, FA_8GSM3 and FA_8GSM4). The fuel assembly design
is presented in figure 4-4.3.1
F
sembly loading maps. Most of the current manufacturers can produce fuel
assemblies with two types of the burnable absorber pins, and three axial zones [176,
181, 343] (i.e., a center axial zone two edges of the fuel assembly containing burnable
absorber). For fuel assembly segments, we may use the semi-standard fuel assembly
designs
258
Figure 4-4.3.1
The main idea behind this design was to depress the neutron flux and power
generation profile in the center of the core, to flatten the space and time (burnup)
distribution of power generation.
The multiplication factor as a function of depletion and soluble boron
concentration for this core configuration is presented in figures 4-4.3.2 A and B.
259
Figure 4-4.3.2 A
Figure 4-4.3.2 B
These figures show that the use of the fuel assemblies with burnable absorber
material decrease the initial multiplication factor of the MASLWR core. The effect of
260
the burnable e at a core
average burnup level of 25 MWD/kg M, and after ~ 30 MWD/kgHM the
multiplication factor curves of the poisoned and non-poisoned core are parallel and the
difference in the values is so small that it canot not be resolved in the graphs.
Figure 4-4.3.2 B shows that use of soluble boron could decrease the value of
the multiplication factor. However, even a concentration of 2400 ppm of soluble boron
does not bring the HFP MASLWR core to critical conditions at BOC and MOC. The
use of a “control group” of control rods will be necessary for the operation of this
core.
The dependence of the core power peaking factors with depletion is shown in
figure 4-4.3.3. This figure also gives a comparison of the peaking factor in the
poisoned core (dashed line) and the non-poisoned core (solid line).
Figure 4-4.3.3
absorbers on the multiplication factor becomes negligibl
H
261
We c th burnable
poisons is smaller than the non-poisoned core, and remains smaller until a burnup of
approximately 10 MWD/kgHM. The nodal peaking factor curve also has two maxima
at burnups equal to 0 and 13 MWD/kgHM. The behavior of the relative power
fraction, fuel burnup and moderator temperature should be examined at these points,
where a heat exchange crisis potentially may occur.
If we also consider the axial peaking factor, we may conclude that initially it is
smaller in the poisoned core, and after a burnup of 13 MWD/kgHM it is slightly
greater than that of the non-poisoned core. The local maximum of the axial peaking
factor occurs at a burnup ~ 14.5 MWD/kgHM. The radial peaking factor behaves
similarly. Initially it is lower for the poisoned core, but after ~ 10 MWD/kgHM the
non-poisoned core has a smaller radial peaking factor.
The axial offset curve is also helpful to evaluate the behavior of the power
generation in the core with burnable absorbers (see figure 4-4.3.4)
Figure 4-4.3.4
an see that the initial nodal peaking factor for the core wi
262
Initia wer profile,
which is represented by negative offset. The axial offset grows with burnup and
reaches 0 at a burnup of approximately 24 MWD/kgHM. The axial offset continues to
increase slightly and reaches a maximum at 30 MWD/kgHM. After that point a slight
correction of the axial offset occurs, and its value stabilizes at +0.02 at the EOC.
4-4.3.2. Core thermal hydraulics
lly the core with burnable absorbers has a bottom-peaked po
The peaking factors give us an overview of the core behavior with respect to
core average depletion. To evaluate core safety and the influence of the burnable
absorbers on the spatial distribution of the power generation and fuel burnup, we need
to consider the 3-D reconstruction of the relative power fraction (RPF), 3-D fuel
burnup profiles and moderator heating in each fuel assembly, at least in the most
important points of core cycle. These points are:
A) Beginning of cycle (B = 0 MWD/kgHM);
B) Local maximum of the node-peaking factor (B = 13 MWD/kgHM);
C) Middle of the cycle and axial offset =0 (B = 25 MWD/kgHM);
D) Maximum of the axial offset (B = 30 MWD/kgHM);
E) End of cycle (B = 50 MWD/kgHM);
The 3D burnup, 3D relative power fraction and 3D coolant temperature are
shown in figures 4-4.3.5 A-E respectfully. The order of the curves in each figure
represents the cause-consequence chain for each state point. The burnup distribution
defines the availability of the fuel and remaining concentration of the burnable poison
in a fuel assembly, which in turn defines the local neutron energy spectrum and
multiplication properties of the fuel and power generation in the node. The power
generation defines the coolant temperature in the fuel assembly channel, which
provides a feedback to a power generation through the moderator density (or
moderating ability of the coolant).
At the BOC, the greater concentration of burnable absorbers in the center of
the core depresses neutron flux and power generation in this area. As a result the RPF
in the center fuel assemblies is close to the RPF of the outer fuel assemblies. The
263
magnitude o leakage (fig
4-4.3.5 A).
Figure 4-4.3.5 A
f the RPF is defined by the coolant temperature and neutron
The spikes in the RPF curve at the non-poisoned edges of the fuel assemblies
do not affect coolant temperature and heat removal conditions. None of the fuel
assemblies reach saturation temperature, but the exit coolant temperature is very close
to the saturated temperature and some sub-cooled boiling may occur.
264
Near f the center
fuel assemblies is very similar to the shape of the power generation profile in the
BOC. The lower core experiences a greater burnup. Recalling Chapter 4-2, figure 4-
2.3.2 A, we find that the multiplication factor for fuel assembly segments FA_8GSM1
and FA_8GSM2 do not change dramatically in the range between 0 and 25
MWD/kgHM. But the multiplication factor of the non-poisoned edges of the fuel
assembly (segment FA_80_00) decreases significantly. As a result, there are no spikes
near the edges of the center fuel assemblies at this burnup.
The power generation profile at this stage leads to significant heating of the
center fuel assembly and bulk boiling may occur after 140 cm of height. The relative
power density at this spot varies from 1 to 0.6, which is large enough to make a heat
exchange crisis possible in this area.
At the MOC (fig.4-4.3.5 C), the burnup of lower part of the center fuel
assemblies is greater than 30 MWD/kgHM, making the effect of the burnable
absorbers in this fuel assembly negligible, and multiplication factor behavior is similar
to non-poisoned fuel. At the same time the upper part of the fuel assembly has a
burnup less than 30 MWD/kgHM, and some gadolinium remains here. The isotopic
content of fissile U-235 is also greater here. The core average burnup of 25
MWD/kgHM characterizes the balance between upper core where the uranium excess
is compensated by remained gadolinium and by effect of the lower moderator density.
The coolant heating at this depletion step is relatively high in the center fuel
assembly, and bulk boiling occurs above the core level of 144 cm.
The core average depletion of 30 MWD/kgHM is the location of the maximum
axial offset. There is no gadolinium effect in the center fuel assemblies (1-1 and 1-2)
so the multiplication factor decreases linearly with increasing burnup. However the
fissile uranium “stored” in this part of the core during the previous depletion allows
better multiplication properties, and a greater RPF in the upper core. The advantages
of the stored uranium outweigh the disadvantages of the lower moderator density in
terms of neutron balance. The shapes of the RPF curves are similar.
the beginning of the MOC (fig 4-4.3.5 B), the burnup curve o
265
Figure 4-4.3.5 B
266
Figure 4-4.3.5 C
267
Figure 4-4.3.5 D
268
Figure 4-4.3.5 E
269
FA_2-2
of the core by
increasing the power generation in outer fuel assemblies. Using standard burnable
absorbers does n nditions, and a
compensation group of the control rods wil e necessary, which decreases the number
of control rods available for reactor scram and decreasing the cold shutdown margin.
As a r
nels.
4-4.3.3. Core kinetics
At the end of the cycle, the RPF curves are very similar for pairs of fuel
assemblies FA_1-1 and FA_1-2 (the innermost fuel assemblies) and FA_1-3 and
(outer fuel assemblies).
The burnup curves show that use of the burnable absorbers allows better
utilization of the fuel in the outer fuel assemblies and at the periphery of the core. The
use of burnable poisons also decreases boiling in the upper part
ot allow control rod-free operation at hot-full power co
l b
esult, an increase of the concentration of the burnable poison and a more
complicated loading pattern is necessary to meet the core design goals.
The calculations show that use of burnable poisons decreases the rate of
depletion for poisoned fuel assemblies. This different time-scale, or “depletion
dependent” time scale can be used for programming the fuel burnup pattern. This
programming is necessary to 1) flatten the multiplication factor curve, 2) decrease
peaking factors and 3) tune up coolant heating in the fuel chan
The evaluation of the core kinetic parameters and reactivity coefficients were
performed for a boron concentration of 0.0 ppm, 800 ppm, 1600 ppm and 2400 ppm.
The effective fraction of the core with standard burnable absorbers is slightly different
from it for a non-poisoned core, however this difference is negligible and almost
invisible in graph. The same set of the reactivity coefficients was calculate for 8.0%
core with standard burnable absorbers as for non-poisoned case.
The moderator temperature reactivity coefficients are shown at figure 4-4.3.6.
270
Figure 4-4.3.6 Reactivity coefficients associated with moderator temperature.
271
f positive
moderator temperature coefficient at the high soluble boron concentration levels,
which,
As we can see from these figure, the ITC, MTC and Coolant inlet temperature
reactivity coefficients at HFP conditions are negative during all cycle even for high
soluble boron concentration of 2400 ppm. This fact resolves concerns o
was occurred early in some VVER and PWR designs [482, 483, 484]. However
we should notice that increase of the soluble boron concentration up to 2400 ppm
leads to a 4 times decrease of moderator temperature reactivity coefficients at BOC
and 3 times decrease at the EOC.
The results for the uniform Doppler reactivity coefficient (UDC) are shown at
the figure 4-4.3.7. The value of UDC is weakly depending on the boron concentration.
The dependence is caused by the spectrum hardening and smaller resonance escape
probability at higher boron concentration.
Figure 4-4.3.7
he power reactivity coefficient (see figure 4-4.3.8) was calculated for the
same b
T
runch case conditions, as described in a previous section, i.e. power increase
does not associate with simultaneous increase of the flow, so the increase of the
reactor power lead to the increase of the fuel and moderator temperature at a nominal
flow conditions.
272
Figure 4-4.3.8
The flow reactivity coefficient calculation results are presented at the figure
4-4.3.9. The increase of the soluble boron concentration decreases the flow reactivity
coefficient, due to increase of the neutron absorption by atoms of boron.
Figure 4-4.3.9
273
cient increases too.
The pressure reactivity coefficient graphs are presented in figure 4-4.3.10. As
we can see from these figure, the increase of the soluble boron concentration leads to
the decrease of the pressure reactivity coefficient down to ~ 0 for a 2400 ppm of
soluble boron. At the EOC, when the axial offset moves towards the upper core, the
sensitivity of the multiplication factor from the moderator density increases (assuming
that sub cooled boiling may occur), so the pressure reactivity coeffi
Figure 4-4.3.10
Figure 4-4.3.11
274
er defect at
the BOC, assuming that at flow rate at the HZP conditions is equal to a nominal, so a
the heat exchange conditions is al MASLWR, where flow rate
depends on power.
r ppm) with increase of
the solu
The use of the burnable poisons leads to a slight increase of the pow
better that in a re
Finally the increase of the soluble boron concentration lead to the hardening of
the epithermal neutrons energy spectrum and self-shielding of the boron in a coolant,
which resulting a slight decrease of the soluble boron worth (pe
ble boron concentration.
Figure 4-4.3.12
4-4.3.4. Conclusion
The use of the standard burnable poisons and radial profiling of the burnable
absorber load in the core, requires the use of the control rods for the compensation of
the excess reactivity, which is decreasing the amount on the control rods available for
a reactor scram and shutdown sub criticality. The core with standard burnable poisons
will have a sub cooled boiling and some fraction of the bulk boiling in the upper core,
275
which could rise a issues related w ge and issues related with use of
soluble boron.
The high concentration of the soluble boron does not lead to the positive
reactivity feedbacks in HFP conditions, which was reported for some PWR cores
[482,483], However, the increase of the boron concentration decreases the magnitude
of these effects, and may lead to the greater amplitude of the core flow oscillations, or
to the longer relaxation periods in a case of the power of flow transients. The overall
ity coefficient decrease on the stability of the nat lation
with multiple feedback mechanisms should be evaluated.
ith heat exchan
effect of the reactiv
system
ural circu
276
4-4.4.
MASLWR Core with 3D BP profiling
4-4.4.1. Core Overview
This se a feasible core configuration, which
meets the requ cycle length, 2) a controllable
core over the w as maximum
burnup, proper heat exchange and 4) a subcritical, transportable configuration under
normal and accide
The search for a feasible core configuration was performed first for a radial
profiling of the burnable poisons load. A total of 100 different core configurations
with different segments used for the radial profiling were considered. While the results
of the radial profiling search did not meet all requirements, the search was continued
for a 3D burnable poison loading. The primary core designs were symmetric relative
to the core midplane. These core configurations have a fundamental disadvantage. Due
to the large coolant temperature gradient, the lower core always has a greater
modera
n
criteria
ction is dedicated to the search for
irements for 1) the declared effective core
hole cycle, 3) fuel safety and performance criteria, such
nt conditions.
tor density and better moderation, and symmetrical core designs have
underburned fuel in the upper core, or possibly coolant boiling.
The final configuration we have developed is a prototypical core for the
MASLWR transportable configuration. There is no claim that this is an optimal core;
in fact, the results of our design calculations suggest that the some of the desig
should be reconsidered. This core meets the design criteria, and with some
consideration, could be evaluated for further studies of MASLWR.
The core was designed with the same assumptions in terms of natural
circulation (semi-forced flow, no cross-flow, same flow rate for all fuel assemblies).
The issues related with peaking factors, heat exchange crisis and coolant boiling were
taken into consideration, and addressed in the proposed core design.
The design of the fuel assemblies used in the core model, the fuel assembly
segments and nodalization map are shown in figure 4-4.4.1.
277
Figure 4-4.4.1.
The notation of the segment names used in this diagram is the same as that
used in Chapter 4-2. This figure shows that W-type fuel segments were used for the
center fuel assemblies, and multiple BA concentration segments were used for burnup
programming in the outer fuel assemblies and edges of the center fuel assemblies.
278
t the end of the fuel campaign, when all gadolinium will be burned out, it is
difficul
ively low power generation in the upper core,
where s ooled
Several operational strategies were proposed in order to extend the effective
operati ore ca
improve safety m
We now d ith the
multipl n fac uss a
proposed operational strategy that allows the core to operate for five effective years
(or rea
control system efficiency.
The core has a greater concentration of burnable absorbers below the middle
plane. The original intent for such a BA loading was to allow a greater burnup of the
upper core, while the fuel pellets does not have a lot of fission products and have a
better thermal conductivity. Also, in the BOC, the impact of the irradiation history is
minimal, so the behavior of the thermal hydraulic parameters could be predicted with
greater confidence.
A
t to have much control over burnup profiling, so there definitely will be a
greater heat generation in a center fuel assembly. The only adjustment that can be
made here, as a core designer, is to design core such that it will have a bottom-peaked
power profile at the EOC, and relat
ub-c boiling may occur.
ng c mpaign and reduce the power load on the high-burnup fuel, and to
argins.
iscuss the behavior of the proposed core in detail, beginning w
icatio tor, power peaking factors and axial offset. We will then disc
ch burnup levels of 50 MWD/kgHM core average, within the limits of 60
MWD/kgHM fuel assembly average). We then consider thermal hydraulic parameters
of the proposed core design and compliance of the core design with the thermal
requirements. Finally we will discuss core kinetic parameters, reactivity coefficients
and reactor
279
he multiplication factor for the proposed core design is presented in figure 4-
4.4.2 A. The 3D profiling of the burnable poisons flattens the effective multiplication
multiplication factor for the HFP conditions varies between
1.02 up to 1.04 during the most of the cycle, allowing some reserve of multiplication
factor
T
factor curve. The value of
for a maneuverability and possible poisoning issues during transients. The
multiplication factor for HFP conditions reaches 1 at a burnup 47.5 MWD/kgHM core
average. However use of the power reactivity effect allows keff=1 at a burnup of 51.0
MWD/kgHM core average with 50% of nominal power and 50% of nominal flow.
Figure 4-4.4.2. A.
Soluble boron can be used as a primary means of compensating for the
reactivity reserves, with control rods used only for reactor scram and transportable
conditions. The branch cases were calculated for soluble boron concentrations of 800
ppm, 1600ppm and 2400ppm. The results for the multiplication factor at HFP
conditions with these boron concentrations are given in figure 4-4.4.2 B.
280
Figure 4-4.4.2. B.
The axial, radial, nodal and 3PIN power peaking factors were calculated for all
state points of the depletion cycle and are shown in figure 4-4.4.3 A.
Figure 4-4.4.3. A.
281
A comparison of the peaking factors of the final design with other preliminary
the proposed
core de
core design options is given in figure 4-4.4.3 B. This figure shows that
sign does not have significant advantages or disadvantages in terms of the
power peaking, compared to the cases considered previously. At the BOC power
peaking factors of the proposed core are smaller than for other options; however at the
MOC and EOC the peaking factors of the proposed core are greater than for other
options.
Figure 4-4.4.3. B.
The axial offset of the proposed core design is presented in figure 4-4.4.4. The
behavior of the axial offset for the proposed core is different from the core options
considered previously. This figure clearly shows that the proposed core design has a
top-centered power profile in the BOC and bottom centered power generation at the
EOC.
282
Figure 4-4.4.4.
These graphs also show that there are several depletion points where peaking
factors
and axial offset have local maximum. These points may be problem areas with
respect to heat exchange. The behavior of the thermal hydraulic parameters at these
points should be examined carefully.
283
4-4.4.2 Operation Extension based on “Power Effect”
The MAS equire
transportation of the loaded reactor module to the manufacturer’s site [104, 110] for
the refueling. At the end of the full power cycle (B=47.5 MWD/kgHM), the reactor
core will reach HFP subcritical conditions. However, during the power decrease,
positive reactivity will be released into the core. The nature of this power reactivity
defect is a cumulative effect of the fuel poisoning, and negative reactivity feedbacks
from sources including fuel and moderator temperature .
The power defect could be used to increase the length of the reactor effective
full power campaign. (See figure 4-4.4.5). The decrease of the reactor power at the
EOC allows the reactor to reach a burnup of 51.0 MWD/kgHM core average, which
corresponds to 5.2 effective years.
Figure 4-4.4.5.
LWR option with full core off-site refueling will r
284
posed design, there is a possibility of heat exchange problems caused
by coo
exit coolant temperature,
which is useful in diagnosing potential boiling events where the core does not remain
covered. The center fuel assembly coolant exit temperature is close to the saturation
temperature at the EOC, and its maximum occurs around 45 MWD/kgHM.
Figure 4-4.5.1
4-4.5. Prototypical core: Thermal Hydraulic parameters
In the pro
lant bulk or sub cooled boiling. For future designs of the MASLWR reactor,
some details on the prototypical core thermal hydraulics might be useful, so we
present several graphs of neutronic and thermal hydraulic parameters. 3D core relative
power fraction, 3D fuel burnup, 3D coolant temperature and density, and 3D fuel
temperature are presented for a following depletion steps:
A) Beginning of cycle (freshly loaded core, B = 0 MWD/kgHM)
B) 1st local maximum of Node and 3PIN peaking factor (B = 7 MWD/kgHM)
C) Middle of cycle (B = 25 MWD/kgHM)
D) 2nd local maximum of Node and 3PIN peaking factor and depletion of the
transient to lower power operation stategy (B = 45 MWD/kgHM)
E) Formal End of cycle (B = 50 MWD/kgHM)
Figure 4-4.5.1 contains a plot of the fuel assembly
285
The 3D burnup, 3D RPF and 3D coolant temperature curves for all fuel
assemblies are presented in figures 4-4.5.2 A-E. The figures for moderator density and
fuel temperature are presented in figures 4-4.5.3 and 4-4.5.4, respectively.
Figure 4-4.5.2 A Thermal Hydraulic Parameters at BOC
As we see from these figures, the burnable poisons profiling allows to increase
the fuel use and power generation in the outer fuel assembly and at the edges of the
core at BOC.
286
Figure 4-4.5.2 B Thermal Hydraulic Parameters at MOC
287
Figure 4-4.5.2 C Thermal Hydraulic Parameters at MOC
288
Figure 4-4.5.2 D Thermal Hydraulic Parameters at MOC
289
Figure 4-4.5.2 E Thermal Hydraulic Parameters at MOC
290
Figure 4-4.5.3 Coolant density at different depletion steps
291
As we may see from the previous calculations, the use of 3D burnable poisons
profiling allows C-MOC, and
distributing power gene
However a chance of the sub cooled boiling and a slight chance of the bulk
boiling occurring at the center fuel assembly at the end of the cycle. Fortunately due to
design strategy, relative power generation in the upper core at the EOC is relatively
low, so decreasing of the reactor power level or increase of the operational pressure at
the EOC may provide some reserves for preventing of a heat exchange crisis.
It is important to mention here that flow model for the current research does
not allow modeling the accurate power-flow dependency and coolant cross-flow
between fuel assemblies. Also we have considered that initially no flow profiling
devises were used, such as flow restrictors for the outer fuel assemblies’ inlet.
However we may perform evaluation of the potential cross-flow rates for a
current core design. If we consider a coolant density graphs (figure 4-4.6.3) we may
see that at the EOC axial change of the flow density is 0.15 g/cm3 per 160 cm of core
length, or ~ 20% of the density change comparing to core average density. Axial
density gradient for the hot channel could be evaluated as 0.094 (g/cm3)/meter.
0.73 g/
compar
coolant heating at the center fuel assembly will facilitate a flow suction from the outer
fuel assemblies.
decreasing the load of the center fuel assembly in BO
ration in a fuel more evenly.
At the BOC the coolant density varies from 0.76 g/cm3 (periphery) to
cm3 (center) at the core exit plane. This change could be evaluated as 4.0%
radial density gradient over core radius, or 0.069 (g/cm3)/meter. At the EOC the
coolant density varies from 0.765 g/cm3 (periphery) to 0.71 g/cm3 (center) at the core
exit plane. This change could be evaluated as 7.0% radial density gradient over core
radius, or 0.127 (g/cm3)/meter. The radial gradients are twice lower at the middle of
the core, however they’re at the same order of magnitude as axial gradients. So we
may conclude, in for the current flow model, axial and radial flow gradients are
able, and radial (planar) component of the flow velocity vector could be
significant. In the other words, there is a driving force for the cross-flow between the
fuel assembly, and the value of the cross flow is greater at the EOC. The greater
292
re depends on the heat generation
level, thermal conductivity of the fuel and heat transfer coefficient from the fuel pin
s
irradiation history [249]. The fu amatically decreases with
burnup. The effect is stronger if the fuel is irradiated at low temperature. Most of this
damage ef thermal
annealing [2 burnup and
irradiation is shown at figure 4-4.5.4.
Figure 4-4.5.4 Fuel thermal conductivity
The fuel temperature is important safety parameter. The average fuel
temperature and maximum centerline fuel temperatu
urface to the coolant. Thermal conductivity of the fuel is a function of the burnup and
el thermal conductivity dr
fect is permanent and only a small part of it is recovered by
49]. The fuel thermal conductivity as a function of the
he average fuel temperature of the most heated pin for MASLWR core as a
function of the depletion is shown at the figure 4-4.5.5. As we can see from the graphs
at the fuel temperature at the BOC is relatively low (maximum ~ 850 ºK). At the EOC,
maximum value for hot pin average fuel temperature is ~ 1100 ºK. The increase of the
fuel temperature at the EOC is caused by RPF redistribution and thermal conductivity
decrease.
T
293
Figure 4-4.5.5 Fuel Temperature at different depletion steps
294
4-4.6. Prototypical core: Kinetic and Safety parameters
4-4.6.1. Prototypical core kinetic and reactivity coefficients
The core behavior could be evaluated trough the reactivity coefficients. There
are several reactivity coefficients. There are several reactivity coefficient that usually
used for a reactor safety analysis [236]. We have already discussed above the
influence of the increased enrichment and burnable absorbers on the MASLWR core
reactiv
r the prototypical core at the EOC.
ity coefficients. At the previous chapters the reactivity coefficients were
evaluated for a fixed values of the soluble boron concentration. For the prototypical
core we will evaluate reactivity coefficients for the boron-free core and for the core
with a critical boron concentration (soluble boron concentration required for Keff = 1).
The critical soluble boron concentration as a function of depletion is presented
at the figure 4-4.6.1. The decreased power at the EOC leads to the increase of the
critical soluble boron concentration (red curve). This effect will cause the changes in a
reactivity coefficients evaluated fo
Figure 4-4.6.1 Critical soluble boron concentration
295
s
for a pr
Now let’s consider an effective fraction of the delayed neutrons for the
MASLWR core (figure 4-4.6.2) in comparison with other core options described
above. Here and for all reactivity coefficient graphs solid black line represents result
ototypical core with 0.0 ppm of soluble boron for a HFP depletion cycle down
to 50 MWD/kgHM. The dashed black line represents a prototypical (final) core with a
critical boron concentration and reduced power at EOC. The colors of the remaining
curves remain the same as in previous section of this chapter.
Figure 4-4.6.2
As we can see from this picture, the effective fraction of delayed neutrons for a
prototypical core with a “programmed” burnup is slightly smaller than for a non-
poisoned 8.0% core. The main reasons for such decrease is a combination of spectrum
harden
due to use of burnable absorbers).
ing in a core with high concentration of the gadolinium burnable absorbers, and
due to a specific burnup-pattern. In a prototypical core, the usage of the of the fuel at
the core periphery is greater than in a non-poisoned core, so for the same core average
burnup final core have lower amount of U-235 on the periphery (due to depletion) and
a greater amount of plutonium in a center core (
The coolant temperature related reactivity coefficients curves are given at the
figure 4-4.6.3.
296
Figure 4-4.6.3 Reactivity coefficients associated with moderator temperature.
297
ed
by the
erature divided by the requested change in temperature.
s we may see from the figure above, the values of all three reactivity
coefficients are different, but the general tendency and behavior of these coefficients is
OC the boron free prototypical core has a greater magnitude
of the
boron decreases the magnitude of the moderator
reactiv
, the value of
UDC is
riod of
effective gadolinium in a core, which is obvious from the graphs.
hanges of the power level an related changes of the flow rate at the EOC
increases the value of the UDC.
The isothermal temperature coefficient (ITC) is the reactivity change
associated with a uniform change in the fuel and moderator inlet temperatures divid
change in the averaged moderator temperature. The moderator temperature
coefficient (MTC) is the reactivity change associated with a change in the moderator
inlet temperature divided by the change in the averaged moderator temperature. The
inlet temperature coefficient is the reactivity associated with a uniform change in the
inlet temp
A
the same. At the BOC-M
moderator reactivity coefficients than boron free non-poisoned 8.0% core and
core with a standard burnable absorbers. The curves for a boron free cores are crossing
at ~ 25 MWD/kgHM. After that point magnitude of moderator reactivity coefficients
for a prototypical core became smaller than fro a non-poisoned 8.0% core.
The use of the soluble
ity coefficients of a prototypical core, so the ITC, MTC curve are relatively flat
at the BOC up to ~ 36 MWD/kgHM. After that point the concentration of the soluble
boron decreases with increase of burnup, so the magnitude of the moderator reactivity
coefficients increasing with burnup, and reaching valued of boron-free core.
The uniform Doppler coefficient (UDC) is the reactivity change associated
with a uniform change in the fuel temperature divided by the change in the averaged
fuel temperature. The values of the UDC as a function of depletion for a different core
options are shown on the figure 4-4.6.4. As we may see from this figure
insensitive to the soluble boron concentration. The presence of Gadolinium is
increasing the magnitude of the UDC. The prototypical core has a greater amount of
Gadolinium than core with a standard burnable absorbers, and longer pe
C
298
Figure 4-4.6.4
The behavior of the power reactivity coefficient (figure 4-4.6.5) is somehow
similar to the UDC. Initially power reactivity coefficient of prototypical core is higher
than for 8.0 non-poisoned core, but after ~ 25 MWD/kgHM it becomes smaller for
prototypical core than for a non-poisoned core. The power reactivity coefficient for a
prototypical core with critical boron concentration is smaller than for boron free core.
The decrease of the power and flow at the EOC lead to significant increase of the
power reactivity coefficient.
Figure 4-4.6.5
299
re,
than for 8.0 non-poisoned core or core with a standard burnable absorbers.
The pressure coefficient is the reactivity change associated with a perturbation
of the primary system pressure divided by the pressure change. The behavior of the
pressure reactivity coefficient with a burnup for a prototypical core (figure 4-4.6.6) is
more complicated, due to a combination of the three dimensional effects caused by
“burnup-programming” implemented in the design. The magnitude of the changes for
pressure reactivity coefficient during the cycle is smaller for the prototypical co
Figure 4-4.6.6
The flow coefficient is the reactivity change associated with a uniform
perturbation of the flow density divided by the percent change in flow. The results for
the flow reactivity coefficient calculations are presented in the figure 4-4.6.7. The 3D
burnable poisons distribution and burnup programming allow flattening the flow
reactivity coefficient curve over burnup. However at the BOC the value of the flow
reactivity coefficient is greater for the prototypical core than for a core with standard
burnable absorbers or for non-poisoned core. The use of the soluble boron decreases
the value of the flow reactivity coefficient, and make reactor less sensitive to the flow
oscillations. The operation at 50% power 50% flow mode, lead to a significant
increase of the flow reactivity coefficient. So, the evaluation of the power-flow
correlations and related reactivity feedbacks should be performed.
300
Figure 4-4.6.7
The boron coefficient is the reactivity change associated with a uniform
perturbation of the boron concentration divided by the boron change. The results of
the soluble boron worth evaluation are presented in a figure 4-4.6.8 A and B.
Figure 4-4.6.8 A
As we can see from figure 4-4.6.8 A The worth of the soluble boron for a
prototypical core is significantly lower than for a core loaded with 4.5% enriched fuel,
but comparable with other 8.0% enriched fuel core options. The evaluation of the
301
souble boron worth for 8.0% enriched fuel core options is given at figure 4-4.6.8 B. As
we can see from the figure, the initially soluble boron worth for a prototypical core ins
greater than for 8.0% non-pois wth rate of the soluble boron
on is lower for prototypical core, and after MOC the soluble boron
oned core, but than the gro
worth with depleti
worth is smaller for a prototypical core comparing to a non-poisoned core.
Figure 4-4.6.8. B
The reactivity coefficient data is also presented in Table 4-4.6.9,
nits for a bete
in a format of
r
compar
wer reactivity
feedbac
interna ient, and value of this coefficient
show th
the absolute English units (like pcm/F, pcm/psia) and relative standard u
ison with a competing reactor design discussed above.
As we can see from these tables, all temperature and po
ks are negative. It comply with all safety requirements, national and
tional. The positive flow reactivity coeffic
at MASLWR core with a natural circulation is sensitive to the flow rate.
302
Te
4-4.
6Pr
typi
MA
WR
core
rea
ctiv
ity c
oeff
icie
nts
Table 4-4.6.9
abl
*
Pow
er d
efec
t was
cal
cula
ted
for H
Znd
ition
s, as
sum
ing
100%
flow
a H
ZP c
ondi
tions
*
OC
con
dns
am
e50
% o
f nom
lin
a p
ower
50%
of n
omin
al c
ore
aver
age
flow
ot
oca
l SL
P-H
FP c
o*
Eiti
ore
ass
u
.9.
d as
303
ins4-4.6.2. Prototypical core RCS efficiency and shutdown marg
WR Reactor control system options
des. Initially 3
possibl stem was considered:
ystem options
The preliminary evaluation of the MASL
were performed with MCNP [104, 107, 113] and CASMO-4E [113] co
e configurations of the reactor control sy
1) PWR-kind control rod clusters
2) BWR-kind cross-blades control rods
3) Reflector/Absorber rotating drums.
Figure 4-4.6.10 Reactor control s
The evaluation of the multiplication factor of the MASLWR core was
perform
results ean of
ctor cooling down.
provide same level of
the cou possible geometrical locations are
es control rods would
when the control rods
partiall profile when the control
ll require a water gap
betwee aulic model of the reactor
ed in a 2D geometry for a quarter core segment in CASMO-4 [108, 113]. The
shows, that use of the rotating reflector/absorber drums is inefficient m
the reactivity control, and does not allow to provide alone an insertion of the negative
reactivity even for a compensation of the power defect during rea
The use of the cross-blades BWR control rods and cluster PWR control rods
allow similar maximum worth of the reactor control system and
ld shutdown safety margin in a case if all
occupied with a control rods. However the use of the cross blad
lead to the limitation of the cross-flow between fuel assemblies
y inserted, and disturb the incoming of upcoming flow
rods are withdrawn. The use of the cross blade control rods wi
n fuel assemblies, which will complicate thermal hydr
304
complication of the neutronic design, because the compensation
decision to consider a cluster type control rods for a reactor
control
l
during to scram reactor and keep subcritical at the cold
diti
nd
h studies are presented in a figure 4-4.6.11.
core, and lead to the
of thermal flux spikes in the water gaps will be necessary.
Finally we made a
system.
One of the MASLWR core design requirements is ability of the reactor contro
normal operation, ability
con ons at any time during the cycle.
The reactor control system worth was evaluated for a 3 options:
1) 4 control rods in a center fuel assemblies.
2) 12 control rods in a 1st and 2 ring of fuel assemblies.
3) 24 control rods clusters in all fuel assemblies
The results of RCS wort
Figure 4-4.6.11.
305
e
possibl e
ident
during
he preferred
urer
at a
conside tified as of principal
portation of the loaded reactor from
ltiplication factor (around core depletion 37
for maintenance
le.
udies
The blue dots represent
HFP co r for CZP conditions,
0.1 MP
The specific requirements on subcriticality are related to transportation of th
reactor module with fuel, and involve proving reactor subcriticality during any
e transportation accident. For the purposes of the feasibility study, th
following accidental conditions were considered as the worst criticality acc
transportation: the core is filled with fresh cold water, and there is a subsequent
failure of a single control rod cluster to maintain its inserted position. T
level of the reactor subcriticality for these conditions is keff < 0.90
The fact that the reactor module may require refueling service at manufact
ny stage of depletion requires that tsite he transportation accident scenario be full-
red at any depletion step. Three depletion steps were iden
concern for the accident evaluation:
1) The fresh core - trans
manufacturer to customer.
2) The maximum of mu
MWD/kgHM) - the necessity to transport reactor
before the fuel is fully depleted.
3) The end of cycle - particularly the end of the extended cyc
The result of the multiplication factor calculations for these accidental st
for the three RCS options are presented in figure 4-4.6.12.
nditions. The red dots represent the multiplication facto
for coolant pressure and temperature equal to normal atmospheric conditions (20 °C,
a).
306
Figure 4-4.6.12.
The calculations show that the transportation configuration requirements can
. In order to provide an acceptable level of sub critbe met icality with a single control
clurod ster failure, the RCS should have 24 control rods.
307
a
SL
of 60 M ading
pattern will be necessary to avoid exceeding the maximum burnup requirements and
oolant
flow m
es. The coupling
ydraulic codes with SIMULATE-3 will be necessary for more precise
la
f
et of the power
density might be
necessary in order to avoid sub cooled boiling and heat exchange issues.
ed
fuel do ons
profilin r the burnup cycle.
4-4.7. Conclusions
The core design feasibility studies show that it is possible to design
WR core for five effective years of operation, withiMA n the fuel burnup limitation
WD/kgHM fuel assembly average. A specific 3D burnable poison lo
avoid bulk coolant boiling.
The current calculations were performed with an insulated channel c
odel implemented in SIMULATE-3 for a PWR core. The natural circulation in
MASLWR will lead to a significant cross-flow between fuel assembli
of thermal h
calcu tions.
The power density of the MASLWR core is significantly higher than in any o
comp ing design proposed in the literature [3,21,27]. The reduction
and reconsideration of the core geometry and operational parameters
The operational conditions of the MASLWR reactor and use of 8.0% enrich
es not lead to the positive reactivity feedbacks. The use of 3D burnable pois
g allows flattening reactivity coefficients curves ove
308
WR
fuel, an
t of
d
identifi f
w rate
conduc ith high concentration of gadolinium, and
uncerta aration of the cross-sections libraries. Quantifying the
e
to which the conclusions of the feasibility study change if design parameters change.
Chapter 4-5 Results – Core Design Sensitivity Study
In previous chapters we have discussed specific features on the MASL
d the process used to develop the MASLWR core design. The straight burn
core loaded with 8.0% enriched uranium fuel will require a precise burnable poison
loading pattern with burnup programming. It is important to understand the impac
the uncertainty in the initial design data on the characteristics of the design. The
purpose of this study was to demonstrate the feasibility of the core design an
cation of issues and deficiencies related to small reactor core design and use o
8.0% enriched fuel in the core. In this chapter, we consider with the effect of flo
uncertainty in a reactor with natural circulation, the uncertainty of the thermal
tivity of the 8.0% enriched fuel w
inties related to the prep
effect of these uncertainties helps to assess the robustness of the design, and the degre
309
4-5.1.
er
on
ixed
with th are
[248, 249, 430].
, LWR’s use fuel with gadolinium concentrations up to 8.0%, and
l
b th PWR fuel
presenc
d
d
(60 MW specific location, the fuel pin (or fuel assembly)
effective burnup could be between 65-70 MWD/kgHM.
y
ving
ax
el
er limit
specifie
4-5.1.1
Effects of fuel thermal conductivity variations
al conductivity of the uranium nuclear fuel is a function of the fuelThe therm
temp ature, fuel burnup and fuel irradiation history [248, 249, 271]. The fuel pellet
microstructure, fission gas and fission product concentrations and migration rates
through the fuel pellet, and fuel pellet swelling and radiation growth also depend
these parameters [19, 22, 249]. The use of gadolinium as a burnable absorber m
e fuel decreases the fuel thermal conductivity [22, 248]. These effects
extremely important at the EOC, for fuel with high burnup levels and for fuel with a
relatively high initial concentration of gadolinium
Traditionally
specific burnup limits and peaking factor limits are imposed on the fuel pins and fue
assem lies with gadolinium [248, 249, 251-253, 258]. Experiments wi
indicate that a decrease the thermal conductivity of up to 5% is possible due to the
e of gadolinium [25, 254].
MASLWR fuel will use fuel assemblies with 8.0% enriched fuel an
gadolinium concentrations up to 10%. The fuel assembly average burnup is limite
D/kgHM); however, at a
In this sensitivity case study, we have modeled fuel thermal conductivit
changes of ±15%, ±10%, ±5%. The effect of these changes is quantified by obser
the m imum fuel pellet average temperature and core power peaking.
The safety limit we are most concerned with is that the mass-average fu
temp ature should not exceed 1204°C (1477°K) [491, 19] (Or 1200 °C as a
d by Russian manufacturer TVEL [19]) The average fuel temperature in the
most heated fuel pellet as a function of core average burnup is presented in figure
.
310
et as a function of Figure 4-5.1.1 Average fuel temperature in the most heated pellthermal conductivity variation and core average burnup.
This figure shows that the average fuel temperature of the most heated fuel
pellet d el
a
of the p
rmal
conduc
gth variation
oes not exceed the safety limit, even if we assume a 15% decrease of the fu
therm l conductivity. The location of the most heated fuel pellet changes as a function
rogrammed fuel burnup in the MASLWR core.
The increase of the fuel temperature caused by the decrease of the the
tivity leads to an increase of the neutron absorption due to Doppler broadening,
of the resonances, and this affects the fuel burnup pattern and cycle length.
Table 4-5.1.1. Core cycle lenThermal conductivity
variation (%) Multiplication factor
variation (%) Core cycle length
variation (%) ± 5% Less than ± 0.05% + 0.17% / – 0.23% ± 10% Less than ± 0.10% + 0.32% / – 0.32% ± 15% Less than ± 0.15% + 0.49% / – 0.46%
These calculations show that the effect of the fuel thermal conductiv
n on the peaking factor is smaller th
ity
tio actor
output
shown in figure 4-5.1.2, and the variation of the axial offset is shown in figure 4-5.1.3
varia an the precision of the peaking f
data format (i.e. less than 1%). The variation in the multiplication factor is
311
Figure 4-5.1.2 Variation in the multiplication factor with fuel thermal conductivity variation ± 15%
Figure 4-5.1.3 Variation in the axial offset with fuel thermal conductivity variation ± 15%
312
rate. Th n this
104].
ely
ctures
h
structu
ra
l to ± 1%,
,
were ev
boiling
bul
equal f
coolant k
boiling starts in the coolant). This param
coolant ted channel, which
change
t it
occurs % of
na
4-5.2. Effect of coolant flow rate variations
An important characteristic of a natural circulation reactor is the coolant flow
e data for the coolant flow rate used in design calculations performed i
dissertation were obtained from experiments at the OSU scaled test facility [
During the full power experiment, the oscillation of the flow rate was approximat
5% [104]; however, the scaled model did not contain the reactor internal stru
as control rod drives and instrumentation channels) and ex-core support (suc
res in any significant detail. It is reasonable to assume that predicted coolant
flow te might be different from that in the real reactor.
For the purposes of the sensitivity study, flow variations equa
± 2% ± 5%, ± 10% were considered. The influence of the flow variation on the
coolant temperature, multiplication factor, power peaking factors and cycle length
aluated.
In chapter 4-4, we demonstrated that there might be some sub-cooled
k boiling occurring in the MASLWR core model with an insulated channel and and
low through each channel. We measure the impact of flow rate changes and
subsequent changes coolant conditions, by calculating the spatial location where the
reaches saturation temperature in the most heated channel (i.e. where the bul
eter is presented graphically in figure 4-5.2.1.
This figure shows that the decrease of the coolant flow rate to 90% of nominal
flow leads to bulk boiling in the upper part of the most hea
in turn leads to a decrease in moderation in the upper core and potential heat ex
crisis. The base scenario case also has some bulk boiling in the upper core, bu
closer to the exit of the heated channel. An increase in the flow rate to 105
nomi l prevents bulk boiling in the MASLWR core model.
313
boiling in the most heated n and burnup
Figure 4-5.2.1. The axial location of coolant bulk channel as a function of flow rate variatio
Figure 4-5.2.2. The core multiplication factor as a function of flow rate variation and burnup
Variations in the multiplication factor caused by flow changes are relativ
± 0.25% for flow variation ± 10%, and ± 0
ely
small ( .05% for flow variation ± 2%).
Howev
core, w
5.2.3 A-B, 4-5.2.4 A-B).
er, variations in flow rate lead to the redistribution of the heat generation in the
hich are illustrated graphically in a plot of power peaking factors (figure 4-
314
Figure 4-5.2.3.A: The 3PIN peaking factor
as a function of flow rate variations and burnup
Figure 4-5.2.3.B: The 3PIN peaking factor variations as a function of flow rate and burnup
Variations in the 3PIN peaking factor are relatively small for small flow
se.
urve is
substan
of the heat generation in the core.
variations, and the shape of the curve is very similar to that observed in base ca
However for large flow variations, the shape of the 3PIN peaking factor c
tially different from the base case, which indicates a significant redistribution
315
g factor variations Figure 4-5.2.4.A: Radial peakinas a function of flow rate and burnup
4.B: Axial peaking factor variations ction of flow rate and burnup
Figure 4-5.2.nas a fu
Figure 4-5.2.4 A-B shows that changes in the flow rate have very little effect
o
flow v led
boiling in the upper core.
on the radial power peaking; however, the axial power peaking is very sensitive t
ariations, primarily because of moderator density changes and sub-coo
316
4-5.3. fuel
MASLWR core design outside of the well-known and well-verified area of LWR
parame nd cross-sections. There are very few benchmark experiments for LWR
, 58, 60]. There are
several
d burnup [224, 256,
257], b
aluate the effect of the uncertainty in fuel composition in the MASLWR
reactor
er
length
n of the core average
ipl
(See fig
does
riation of the
he output data
at
Effects of fission cross section uncertainty for 8.0%
The use of 8.0% enriched fuel and high concentrations of gadolinium force the
ters a
production code verification currently available [224, 254-257, 53
proposals and programs explaining the necessity of these benchmark
experiments for LWR fuel with increased enrichment and increase
ut the results of these experiments are not available at this time.
To ev
, a ±1% uncertainty in the νΣf (the fission cross section multiplied by the
numb of neutrons produced per fission) is considered in all fuel segments. The cycle
and peaking factor variations are observed in these test problems.
For the ±1% uncertainty in νΣf, the uncertainty in the cycle length can be as
great as ±4%, relative to the base case scenario. The variatio
ication factor reaches a maximum ±0.9% rmult elative to the base case scenario.
ure 4-5.3.1)
A simultaneous variation of the multiplication factor for all fuel segments
not lead to significant changes in the power peaking factors. The va
peaking factor throughout the cycle is smaller than the precision of t
form (less than 1%; see figure 4-5.3.2)
317
factor Figure 4-5.3.1 The core average multiplication as a function of changes in νΣf and fuel burnup
Figure 4-5.3.2 The 3PIN peaking factor as a function of changes in νΣf and fuel burnup
These plots show that simultaneous variations in νΣf lead to significant cycle
length variations, but do not change the heat generation profile.
318
m worth uncertainty at BOC-MOC
creased
effects. ting
evaluate the sensitivity in th
that the
uncerta m worth evaluation is 5%. A series of simulations were
rm KS card
in SIMULATE3) for each burnable absorber segment. In total 13 correction tables
were us worth.
S
card is for each
en
gadolin e
of
effng a gadolinium worth variation 5%
4-5.4. Effects of gadoliniu
The increased concentration of gadolinium in MASLWR fuel and in
fuel enrichment leads to fuel and burnable absorber self-shielding and shadowing
Because the design of the MASLWR core is involves data and opera
conditions outside of the standard range of PWR parameters, it is reasonable to
e design due to gadolinium worth.
For the purposes of the current sensitivity study we have assumed
inty in the gadoliniu
perfo ed for the fuel with a perturbed multiplication factor (using the TAB.X
ed for the simulation of the 5% uncertainty in the gadolinium
The algorithm for calculation of the keff correction coefficients for TAB.XK
shown in figure 4-5.4.1. First, the worth of gadolinium was calculated
segm t as the reactivity difference between non-poisoned fuel and fuel with
ium. A 5% variation in this gadolinium worth is then implemented, and th
multiplication factor correction coefficient is calculated to reflect the 5% variation
gadolinium worth.
Figure 4-5.4.1. Calculation of k correction coefficients, ssumi a
The core average multiplication factor as a function of depletion and
gadolinium worth variation is presented in figure 4-5.4.2.
319
tion of depletion and gadolinium worth variation Figure 4-5.4.2. The core average multiplication factor
as a func
A ±5% variation in the gadolinium worth leads to a ±2% variation in the
multipl t BOC, and an approximately ±1% variation in the
e gadolinium worth leads to changes in the burnup pattern
rnup
.
However, the 3PIN peaking factor remains below 2.0 throughout the depletion cycle.
eas t
generat
ogrammed
sig
ication factor a
multiplication factor at MOC.
This variation in th
and the effectiveness of the fuel burnup programming. The changes in the bu
pattern affects the shape of the 3PIN peaking factor curve (see figure 4-5.4.3)
The axial offset is also sensitive to gadolinium worth (see figure 4-5.4.4.). A
e in the gadolinium worth leads to a more neutrdecr al (centered) core hea
ion.
A ±5 % variation in gadolinium worth does not dramatically affect the core
cycle length. The variation of the cycle length is less than ±0.06%.
In summary, variations in the gadolinium worth affect the fuel pr
burnup strategy, but do not dramatically increase fuel peaking factors and do not lead
nificant change in the cycle length. to a
320
Figure 4-5.4.3. 3PIN Peaking factor as function of variation in gadolinium worth and fuel burnup
Figure 4-5.4.4. Axial offset as function of variation in gadolinium worth and fuel burnup
321
udy,
we est
describ e
the var
ual
lin
factor
less than
al
4-5.5. Effect of variations in residual gadolinium worth at EOC
The residual gadolinium worth is related to the uncertainty in the prediction of
the gadolinium isotopic content in the fuel at the EOC. For the purposes of this st
imate that worth of residual gadolinium has an uncertainty of ±5%. The
methodology of the current simulation is similar to the methodology of the case study
ed in section 4-5.4, with the exception that now we consider no variation in th
gadolinium worth between the beginning and middle of the cycle; we consider only
iation of the residual gadolinium worth.
The multiplication factor variation as a function of depletion and resid
gado ium worth is shown in figure 4-5.5.1. The relative difference in multiplication
caused by variations in the residual gadolinium worth is shown in figure 4-
5.5.2. The variation in the core cycle length caused by residual gadolinium is
0.4% relative to the base case scenario cycle length.
Figure 4-5.5.1 The multiplication factor as a function of depletion and residugadolinium worth
322
as a function of depletion and residual gadolinium worth Figure 4-5.5.2 The relative difference in multiplication factor
The variation in the residual gadolinium worth leads to the insignificant (less
1%
redistri ase case scenario.
than ) variations in the pin peaking factor at the EOC and negligible changes in the
bution of the heat generation and axial offset relative to the b
323
Chapter 5 Discussion, Conclusions, Recommendations
e feasibility of neutronic core design for
The
complex core-loading pattern with 3-D profiling of the burnable absorber load will be
ssa
igh
concentration of burnable absorbers in a MASLWR operational conditions.
5.1.1. Fuel assembly level studies
5.1. Conclusion
The dissertation demonstrates th
MASLWR reactor with off-site refueling and 5 effective years of operation.
nece ry to fit a complex MASLWR design requirements. A complex of the support
studies were performed in order to evaluate the 8.0% fuel behavior with h
nd PWR fuel is increased initial
fuel
temper y). The combined effect of these
factors tics.
burnup,
me
advanta fuel burnup
be on
growth and swelling. Currently, burnup is limited to 60 MWD/kgHM average per fuel
assembly. This means that the spent nuclear fuel of MASLWR with initial enrichment
tor
f-
site ref
pic composition of the irradiated MASLWR fuel will also be
different form the isotopic com
differences are 1) greater weight fractions of U-235 and U-236 at the EOC; 2) lower
The major difference between MASLWR a
enrichment (up to 8.0%) and lower operational parameters (lower power density,
ature, coolant temperature, coolant densit
leads to differences in the core neutronic characteris
The increased enrichment of the MASLWR fuel allows greater fuel
and a longer reactor campaign. The lower operational parameters will also give so
ges for an extended reactor campaign. The main limitation to the
will caused by mechanical properties of the fuel cladding and fuel pellet radiati
of 8.0% still could be critical at this burnup level. This creates challenges for reac
refueling and transportation of spent nuclear fuel (or the fueled reactor in case of of
ueling).
The isoto
position of the spent PWR fuel. The specific
324
of
Pu-239 on of the fission products caused by
be
he
of the
MASL rence in the spent
nuclear sition compared to PWR spent nuclear fuel.
leads to he increase of the initial enrichment will
e i
neutrons and a softer spectrum than in a PWR with the same fuel enrichment. The
will have a slightly harder neutron energy spectrum than a conventional PWR, and this
compo irradiated fuel will also lead to differences in the prompt neutron
me LWR and a conventional
PWR. in
modeli
rol rod worth and
total plutonium weight fraction, but better quality of the plutonium (greater fraction
and Pu-241; 3) slightly different compositi
a longer irradiation period at a lower power density. These differences should
considered during the evaluation of the criticality safety and radiation safety for t
EOC fuel management, storage, transportation and reprocessing. The extension
WR burnup beyond 60 MWD/kgHM will lead to a larger diffe
fuel compo
The increased initial enrichment and difference in the operational parameters
a neutron energy spectrum shift. T
caus ncreased absorption of thermal neutrons, and spectrum hardening. Increased
moderator density (compared to a PWR) will lead to better thermalization of the
evaluation of the combined effect of this competing process shows that MASLWR
fact should be taken into account.
The difference in neutron energy spectrum and the difference in isotopic
sition of the
lifeti and effective delayed neutron fraction between MAS
Some assumptions and models used for the PWR should be corrected
ng MASLWR..
The reactor kinetics and dynamics parameters will also be different for a PWR
and for MASLWR. This means that reactivity coefficients, cont
soluble boron worth should be accurately evaluated for the MASLWR safety analysis.
5.1.2. Burnable absorber studies
The implementation of burnable absorbers is essential in order to compensate
and
technol R does not belong to one
of the w t
is, however, rber technology that can be provided by
initial reactivity reserves. Different types of burnable absorbers materials
ogies are currently available on the market. MASLW
orld’s biggest fuel and reactor vendors (like AREVA, GE, Westinghouse); i
necessary to use burnable abso
325
ide mixed
uniform his
ctions
ughter
products have relatively low absorption cross-sections. This is why 1) gadolinium
effect.
when l rent
lin uel
burnup uring
burnup
e neutron
flux ar
effectiv e
multipl
type) also transform bly, which may
ith
evenly me
ol
rods or soluble boron. These effects do e
reactiv nderstanding of these effects is
ASLWR fuel, operational conditions and
specific requirements, such as off-site refueling and 5-year core lifetime creates an
several of these suppliers. Burnable absorbers based on gadolinium ox
ly with fuel in a fuel pellet can be produced by a variety of vendors, so t
technology was chosen as a baseline for the MASLWR core design feasibility study.
Two gadolinium isotopes, Gd-155 and Gd-157, have absorption cross-se
several orders of magnitude greater than that of uranium-235, but their da
burns out faster than uranium and 2) gadolinium-doped fuel has a great self-shielding
Gadolinium self-shielding increases the gadolinium effective burnout time
arger gadolinium concentrations are used. The use of fuel pins with diffe
gado ium concentrations in the same fuel assembly allows programming of the f
pattern and programming of the multiplication capabilities of the fuel d
.
The great neutron absorption of gadolinium-doped pins depresses th
ound the pin. Combinations of gadolinium-doped pins may lead to shadowing
of the other fuel pins and control rods. Use of this shadowing effect can extend the
e burnout time for gadolinium-doped pins in a fuel assembly, and flatten th
ication factor curve.
Fuel assemblies with built-in shadowing features (called here V-type and W-
the neutron flux distribution in the fuel assem
cause an increase in pin-peaking factor compared to non-poisoned fuel, or fuel w
distributed burnable absorbers. At the same time, this shadowing saves so
uranium in poisoned pins as reserves for use at the middle and end of cycle.
The transformation of the neutron flux in a fuel assembly with burnable
absorbers affects the efficiency of the reactivity control mechanisms such as contr
not necessarily mean a decrease of th
ity control mechanism’s efficiency. The u
necessary for high-quality reactor core design.
The increased enrichment of the M
326
designs vates the study of the fuel behavior and behavior of the burnable
e
data ge may
hm
envelope of the design parameters that is different from traditional PWR or BWR
. This moti
absorbers in fuel up to 8.0% enrichment in a MASLWR operational conditions. Th
nerated in this study creates a basis for a core design feasibility study and
be useful for other designs of small LWRs with extended core life and increased
ent, such as modifications of IRIS, Cenric AREM, and SMART [27, 120, 122].
5.1.3. Reflector Studies
A variety of reflector materials are considered for the MASLWR reactor:
These roper
a
small r
standar
ls
provide d
photon lifetime and plant lifetime management.
void
reflecto
generat tor cooling system is
ess steel reflector with reflector pads will
be use
prospec eel reflector was also made in
r
water, graphite and stainless steel. Different geometries of the reflector are possible.
reflector options have different advantages and disadvantages, so the p
choice of reflector is a complicated task.
The use of a stainless steel reflector may provide improved fuel usage in
eactor core, where neutron leakage is significant. For the MASLWR core, fuel
burnup in the outer fuel assemblies could increase by up to 150% compared to the
d PWR water reflector.
Heating rates in the stainless steel reflector and structures could be as high as
580 kW, requiring proper heat removal mechanisms. Also different reflector materia
different levels of protection of the reactor vessel against neutrons an
s, which could affect the reactor vessel
A heat removal mechanism should be designed for the reflector in order to a
r melting and degradation. The current calculations estimate the heat
ion in the reflector and core barrel, but design of the reflec
out of the scope of the current report.
For the further calculations the stainl
d as a base line preferred option for neutron economy and fuel utilization
tive. The same argument for use of stainless st
other reactor designs [3, 27, 133, 138]. The use of the reflector pads is necessary fo
327
l
assemb
ge
of feasibility study, the bypass flow and reflector heating contribution into energy
r it
should sign (the next stages of
the MASLWR core in order to improve fuel utilization of the outer ring of fue
lies (especially fuel utilization in the pins on the corners of the assemblies).
In order to simplify the neutronic modeling of the MASLWR core at the sta
generation would be considered as insignificant (only at current stage), howeve
be considered at the stage of the detailed core de
design).
5.1.4. Core Design
The core design feasibility studies show that it is possible to design a
WR core for five effective years of operation, within the fuel burnup limitation
MASL
and
avoid b
d channel coolant
m lation in
MASL
of ther draulic codes with SIMULATE-3 will be necessary for more precise
calcula
of
compet power
erational parameters might be
ed
profilin burnup cycle.
of 60 MWD/kgHM fuel assembly average. A specific 3D burnable poison loading
pattern will be necessary to avoid exceeding the maximum burnup requirements
ulk coolant boiling.
The current calculations were performed with an insulate
flow odel implemented in SIMULATE-3 for a PWR core. The natural circu
WR will lead to a significant cross-flow between fuel assemblies. The coupling
mal hy
tions.
The power density of the MASLWR core is significantly higher than in any
ing design proposed in the literature [3, 21, 27]. The reduction of the
density and reconsideration of the core geometry and op
necessary in order to avoid sub cooled boiling and heat exchange issues.
The operational conditions of the MASLWR reactor and use of 8.0% enrich
fuel does not lead to the positive reactivity feedbacks. The use of 3D burnable poisons
g allows flattening reactivity coefficients curves over the
328
ent
rta
re loaded with 8.0% core, drives us into
conditi ified for a
MWD/
of these codes for a MASLWR core should be performed in order use the results of a
calcula e
uncerta
fresh a 7]. Curent
pattern le
for the design of the spent fuel storage, packaging, transportation and reprocessing
nol
MASL ill remain
fu not
rm e the
results
5.2. Recommendations for a future work
A several recommendations could be made upon the result of the curr
tion: disse
The design of the MASLWR reactor co
a new are of the fuel enrichments, burnup levels and new core thermal hydraulic
ons. Today most of the LWR analysis codes are designed and ver
standard LWR enrichment levels (below 5.0%) and burnup levels (below 60
kgHM). The evaluation of the calculation uncertainties and methodology of use
tion for the safety analysis and justification. The evaluation of the cod
inties and code verification will require a cross-reference calculation as well as
verification of the codes against an experimental data [224, 231].
The use of the 8.0% fuel, in a straight-burn core, could create issues with a
nd spent fuel storage, handling and transportation [340, 341, 376, 37
dissertation provides an evaluation of the spent fuel isotopic content and burnup
of the fuel assemblies in a MASLWR core. This information could be valuab
tech ogies associated with MASLWR fuel. The isotopic content of the spent fuel
was also evaluated in order to provide an input for an economical evaluation of the
WR fuel recycling, and estimation of the radioactive wastes that w
after el reprocessing. However evaluations for the back-end of the fuel cycle was
perfo ed in a current research and will require an additional research, wher
of the current dissertation could be used as an impute for calculations.
329
core. T on a secondary circuit
e ss
raulic
m e practical
l.
t
was not performed for the
per
descrip
core wi n a future for a safety evaluation and next steps of the core
design.
parame g designs. The calculations shows the reduction of the
power w might be necessary in order to
er of these
e
installa l
he
nhanced
oo a complex of the
rc
R reactor might be interested in a core design
ome
evaluation of the standard fuel use in a MASLWR core are presented in a current
dissertation, however the design and optimization of the MASLWR core with a partial
The natural circulation raises an issue of the coolant cross-flow in a MASLWR
he core flow rate for a current design is also depends
param ters and geometry of the reactor internals. The proper evaluation of the cro
flow requires 1) a research and development work on coupling of thermal hyd
thcore odel with a code for a neutronic calculations and 2) A set of
experiments at OSU and other test facilities for a natural circulation system mode
The reactivity feedbacks of the MASLWR core were evaluated in a curren
dissertation. However the evaluation of the reactivity feedbacks at the intermediate
ion core parameters stability power levels and evaluat
performed for a coupled system (neutronic –thermal hydraulic coupling with a pro
tion of the natural circulation effects). The core transients analysis for a such
ll be also necessary i
The thermal hydraulic parameters of the MASLWR core are different from
ters of the competin
density with an improvement of the core flo
ensure safety margins and avoid heat exchange crisis. Larger core with a greater
numb of the fuel assemblies could be one of the ways for a resolution
problems. Another way for solution of the heat exchange issues could be th
tion of the pumps for a forced circulation of the coolant at some operationa
modes (like it was proposed for CAREM [27] reactor design). Another option for t
MASLWR core design could be a boron-free core, with natural circulation e
by c lant boiling. Evaluation of all these options will require
resea h work
Some customers of the MASLW
option with on-site refueling and use of the fuel with standard enrichment. S
330
onal calculations, and, probably a different
ns with CASMO-4E performed in the
ce
onto reactor vessel. The choice of the reflector for a prototypical core was made from
differen tion for the MASLWR core reflector. So the additional studies of
the core reflector option from a lifetime management prospective will require an
Finally, as it was mentioned in the early beginning, the current feasibility study
is opening a door to a large variety of studies associated with small, innovative LWR
design. I hope that result of the current research work will be valuable for a variety of
the researchers and engineers involved into innovative small LWR designs.
refueling will require a lot of additi
philosophy of the core design.
The evaluation of the core reflector optio
dissertation shows that different core reflectors lead to a different radiation fluen
the fuel usage prospective, however the plant live-time management could lead to a
t design solu
additional research work.
331
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Accordance with 10 CFR 50.71(e)”, September 1999 312. NRC Regulatory Guide 1.203 “Transient and Accident Analysis Methods”, December 2005 313. NRC Regulatory Guide 1.206 “Combined License Applications for Nuclear Power Plants
(LWR Edition)” June 2007, (Publication of the associated revision to Title 10 Code of Federal Regulations Part 52 is pending. (HTML Version: Hhttp://www.nrc.gov/reading-rm/doc-collections/reg-guides/power-reactors/active/01-206/H)
314. NRC Regulatory Guide 3.71 “Nuclear Criticality Safety Standards for Fuels and Material Facilities”, Rev. 1, October 2005
315. NRC Regulatory Guide 3.54 “Spent Fuel Heat Generation in an Independent Spent Fuel Storage Installation” Rev. 1, January 1999
316. NRC Regulatory Guide 3.49 “Design of an Independent Spent Fuel Storage Installation (Water-Basin Type)”, December, 1981
317. NRC Regulatory Guide 3.60 “Design of an Independent Spent Fuel Storage Installation (Dry Storage)”, March 1987
318. NRC Regulatory Guide 5.54 “Standard Format and Content of Safeguards Contingency Plans for Nuclear Power Plants”, March 1978 (Version for Comment)
319. NRC Regulatory Guide 5.56 “Standard Format and Content of Safeguards Contingency Plans for Transportation”, March 1978 (Version for Comment)
320. NRC Regulatory Guide 5.57 “Shipping and Receiving Control of Strategic Special Nuclear Material”, Rev 1, June 1980
321. NRC Regulatory Guide 5.59 “Standard Format and Content for a Licensee Physical Security Plan for the Protection of Special Nuclear Material of Moderate or Low Strategic Significance” Rev. 1, February 1983
322. NRC Regulatory Guide 5.60 “Standard Format and Content of a Licensee Physical Protection Plan for Strategic Special Nuclear Material in Transit”, April 1980
323. NRC Regulatory Guide 5.63 “Physical Protection for Transient Shipments”, July 1982
350
324. NRC Regulatory Guide 5.67 “Material Control and Accounting for Uranium Enrichment Facilities Authorized To Produce Special Nuclear Material of Low Strategic Significance”, December 1993
325. NRC Regulatory Guide 7.1 “Administrative Guide for Packaging and Transporting Radioactive Material” June 1974
326. NRC Regulatory Guide 7.3 “Procedures for Picking Up and Receiving Packages of Radioactive Material” May 1975
327. NRC Regulatory Guide 7.4 “Leakage Tests on Packages for Shipment of Radioactive Materials” June 1975
328. NRC Regulatory Guide 7.5 “Administrative Guide for Obtaining Exemptions from Certain NRC Requirements over Radioactive Material Shipments”, June 1975 (Revised version O-R May, 1977)
329. NRC Regulatory Guide 7.6 “Design Criteria for the Structural Analysis of Shipping Cask Containment Vessels”, Rev. 1, March 1978
330. NRC Regulatory Guide 7.8 “Load Combinations for the Structural Analysis of Shipping Casks for Radioactive Material”, Rev. 1, March 1989
331. NRC Regulatory Guide 7.9 “Standard Format and Content of Part 71 Applications for Approval of Packages for Radioactive Material”, Rev. 2, March 2005
NRC Publications NUREG 332. NUREG-75/087 “Standard Review Plan for the Review of Safety Analysis Reports for
Nuclear Power Plants” (LWR Edition), US NRC, May 1980
333. NUREG-0800 “Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants” (LWR Edition), US NRC, June 1996
334. NUREG-1275 Vol.14 “Causes and Significance of Design-Based Issues at U.S. Nuclear Power Plants”US NRC, November 2000
335. NUREG-1449 “Shutdown and Low-Power Operation at Commercial Nuclear Power Plants in the United States”, US NRC, September 1993
336. NUREG-1513, R.I. Milstein “Integrated Safety Analysis Guidance Document”, US NRC, May 2001
337. NUREG-1520 “Standard Review Plan for the Review of a License Application for a Fuel Cycle Facility”, US NRC, March 2002
338. NUREG-1536 “Standard Review Plan for Dry Cask Storage Systems “ US NRC, January 1997 339. NUREG-1567 “Standard Review Plan for Spent Fuel Dry Storage Facilities” US NRC, March
2000 340. NUREG-1609 “Standard Review Plan for Transportation Packages for Radioactive Material”,
US NRC March 1999 341. NUREG-1617 “Standard Review Plan for Transportation Packages for Spent Nuclear Fuel”,
US NRC March 2000 342. NUREG-1749 R.O. Meyer “Implications From The Phenomenon Identification and Ranking
Tables (PIRTs) and Suggested Research Activities for High Burnup Fuel”, US NRC, September 2001
343. NUREG-1754 G.M. O’Donnell, H.H. Scott, R.O. Meyer “A New Comparative Analysis of LWR Fuel Designs”, US NRC December 2001
351
344. NUREG-1780, “Regulatory Effectiveness of the Anticipated Transient Without Scram Rule”, US NRC, September 2003
345. NUREG-1860 M. Drouin “Feasibility Study for a Risk-Informed and Performance-Based Regulatory Structure for Future Plant Licensing”, Volumes 1 and 2, US NRC, December 2007
346. NUREG/CP-0172 “Proceedings of the Twenty-Eighth Water Reactor Safety Information Meeting”, Held at Bethesda Marriott Hotel, Bethesda, MD, October 23-25, 2000, Published by US NRC, April 2001
347. NUREG/CP-0174 “Transactions of the Twenty-Ninth Nuclear Safety Research Conference (formerly The Water Reactor Safety Information Meeting)”, Held at Marriott Hotel at Metro Center, Washington D.C., October 22-24, 2001, Published by US NRC, October 2001
348. NUREG/CP-0175 “Proceedings of the Advisory Committee on Reactor Safeguards Workshop on Future Reactors”, June 4-5 2001, Published by US NRC, December 2001
349. NUREG/CP-0180 “Proceedings of the 2002 Nuclear Safety Research Conference”, Held at Marriott Hotel at Metro Center, Washington D.C., October 28-30, 2002, Published by US NRC, March 2003
350. NUREG/CP-0182 “Proceedings of the Advisory Committee on Nuclear Waste Transportation Working Group Meeting”, Volumes 1-4, November 19-20, 2002, Published by US NRC, April 22, 2003
351. NUREG/CP-0185 “Proceedings of the 2003 Nuclear Safety Research Conference”, Published by US NRC, July, 2004
352. NUREG/CP-0192 “Proceedings of the Nuclear Fuels Sessions of the 2004 Nuclear Safety Research Conference”, Held at Marriott Hotel at Metro Center, Washington D.C., October 25-27, 2004, Published by US NRC, October 2005
353. NUREG-BR-0053 “USNRC Regulations Handbook”, Rev.6, US NRC, September 2005 354. NUREG-BR-0175 J.S. Walker “A Short History of Nuclear Regulation, 1946-1999”, US
NRC, January 2000 355. NUREG-BR-0249 “The Atomic Safety and Licensing Board Panel”, US NRC, August 2004 356. NUREG-BR-0280 “Regulating Nuclear Fuel”, US NRC, September 2001 357. NUREG-BR-0292 “Safety of Spent Fuel Transportation”, US NRC, March 2003 358. NUREG-BR-0298 “Nuclear Power Plant Licensing Process”, US NRC, July 2004 359. NUREG-BR-0299 “Web-Based Public Access to ADAMS”, US NRC, November 2002 360. NUREG/IA-0156, L. Yegorova “Data Base on the Behavior of High Burnup Fuel Rods with
Zr-1%Nb Cladding and U02 Fuel (VVER Type) under Reactivity Accident Conditions”, US NRC, July 1999
361. NUREG/IA-0169, V.A. Vinogradov, A.Y. Balykin “Analysis of KS-1 Experimental Data on the Behavior of the Heated Rod Temperatures in the Partially Uncovered VVER Core Model Using RELAP5/MOD3.2” US NRC, November 1999
362. NUREG/IA-0175, A. Avvakumov, V. Malofeev, V. Sidorov “Analysis of Pin-by-Pin Effects for LWR Rod Ejection Accident”, US NRC, March 2000
363. NUREG/IA-0180, H. Kantee “Application of RELAP5/MOD3.1 to ATWS Analysis of Control Rod Withdrawal From 1% Power Level”, US NRC, June 2000
364. NUREG/IA-0195, J.I. Sinchez, C.A. Lage, T. Nfifiez “LBLOCA Analysis in a Westinghouse PWR 3-Loop Design Using RELAP5/MOD3”, US NRC, Jan 2001
365. NUREG/IA-0199, E. Kaplar, L. Yegorova, K. Lioutov, A. Konobeyev, N. Jouravkova, “Mechanical Properties of Unirradiated and Irradiated Zr-1% Nb Cladding - Procedures and
352
Results of Low Temperature Biaxial Burst Tests and axial Tensile Tests”, US NRC, April 2001
366. NUREG/IA-0209, A. Shestopalov, K. Lioutov, L. Yegorova “Adaptation of USNRC's FRAPTRAN and IRSN's SCANAIR Transient Codes and Updating of MATPRO Package for Modeling of LOCA and RIA Validation Cases with Zr-1%Nb (VVER type) Cladding”, US NRC, April 2003
367. NUREG/IA-0211, L. Yegorova, K. Lioutov, N. Jouravkova, A. Konobeev “Experimental Study of Embrittlement of Zr-1%Nb VVER Cladding under LOCA-Relevant Conditions” US NRC, March 2005
368. NUREG/IA-0213, L. Yegorova, K. Lioutov, N. Jouravkova, O. Nechaeva, A. Salatov, V. Smirnov, A. Goryachev, V. Ustinenko, I. Smirnov “Experimental Study of Narrow Pulse Effects on the Behavior of High Burnup Fuel Rods with Zr-1%Nb Cladding and UO2 Fuel (VVER Type) under Reactivity-Initiated Accident Conditions”, US NRC, May 2006
369. NUREG/IA-0215, A. Avvakumov, V. Malofeev, V. Sidorov “Spatial Effects and Uncertainty Analysis for Rod Ejection Accidents in a PWR”, US NRC, September 2007
370. NUREG/CR-4674, Vol.27 “Precursors to Potential Severe Core Damage Accidents: 1998 A Status Report” US NRC, July 2000 (ADAMS: ML003733843)
371. NUREG/CR-5734 “Recommendations to the NRC on Acceptable Standard Format and Content for the Fundamental Nuclear Material Control (FNMC) Plan Required for Low-Enriched Uranium Enrichment Facilities”, US NRC
372. NUREG/CR-6042 “Perspectives on Reactor Safety”, Rev. 2, US NRC, March 2002, (ADAMS: ML021080422)
373. NUREG/CR-6441 “Analysis of Spent Fuel Heat-up Following Loss of Water in a Spent Fuel Pool, Final Report”, US NRC, March 2002, (ADAMS: ML021050336)
374. NUREG/CR-6577 “U.S. Nuclear Power Plant Operating Cost and Experience Summaries”, Supplement 1 and 2, US NRC, Jan 2001, (ADAMS: ML010120458)
375. NUREG/CR-6655 “Sensitivity and Uncertainty Analyses Applied to Criticality Safety Validation” (Vol. 1 - Methods Development; Vol. 2 - Illustrative Applications and Initial Guidance), US NRC, November 1999, (ADAMS: ML003726900, ML003726890)
376. NUREG/CR-6665, C.V. Parks, M.D. DeHart, J.C. Wagner, “Review and Prioritization of Technical Issues Related to Burnup Credit for LWR Fuel”, (ORNL/TM-1999/303), prepared for US NRC by Oak Ridge National Laboratory, Oak Ridge, Tenn., February 2000.
377. NUREG/CR-6683 “A Critical Review of the Practice of Equating the Reactivity of Spent Fuel to Fresh Fuel in Burn-up Credit Criticality Safety Analyses for PWR Spent Fuel Pool Storage”, US NRC, September 2000
378. NUREG/CR-6686 “Experience With the Scale Criticality Safety Cross-Section Libraries”, US NRC, Oct. 2000
379. NUREG/CR-6700 “Nuclide Importance to Criticality Safety, Decay Heating, and Source Terms Related to Transport and Interim Storage of High-Burnup LWR Fuel” US NRC, January 2001 (ADAMS: ML010330186)
380. NUREG/CR-6701 “Review of Technical Issues Related to Predicting Isotopic Compositions and Source Terms for High-Burn-up LWR Fuel”, US NRC, January 2001, (ADAMS: ML010230244)
381. NUREG/CR-6702 “Limited Burnup Credit in Criticality Safety Analysis: A Comparison of ISG-8 and Current International Practice”, US NRC, January 2001, (ADAMS: ML010190276)
353
382. NUREG/CR-6703 “Environmental Effects of Extending Fuel Burn-up Above 60 Gwd/MTU”, US NRC, January 2001, (ADAMS: ML010310298)
383. NUREG/CR-6716 “Recommendations on Fuel Parameters for Standard Technical Specifications for Spent Fuel Storage Casks”, US NRC, March 2001 (ADAMS: ML010820352)
384. NUREG/CR-6742 “Phenomenon Identification and Ranking Tables (PIRTs) for Rod Ejection Accidents in Pressurized Water Reactors Containing High Burnup Fuel” US NRC, September 2001, (ML012890487)
385. NUREG/CR-6743 “Phenomenon Identification and Ranking Tables (PIRTs) for Power Oscillations Without Scram in Boiling Water Reactors Containing High Burnup Fuel” US NRC, September 2001, (ML012850324)
386. NUREG/CR-6744 “Phenomenon Identification and Ranking Tables (PIRTs) for Loss-of-Coolant Accidents in Pressurized and Boiling Water Reactors Containing High Burnup Fuel”, US NRC, December 2001
387. NUREG/CR-6745 “Dry Cask Storage Characterization Project - Phase 1: CASTOR V/21 Cask Opening and Examination”, US NRC, September 2001, (ADAMS: ML013020363)
388. NUREG/CR-6747 “Computational Benchmark for Estimation of Reactivity Margin from Fission Products and Minor Actinides in PWR Burn-up Credit”, US NRC, October 2001 (ADAMS: ML013060035)
389. NUREG/CR-6748 “STARBUCS: A Prototypic SCALE Control Module for Automated Criticality Safety Analyses Using Burn up Credit” US NRC, October 2001 (ADAMS: ML013060098)
390. NUREG/CR-6759 “Parametric Study of Effect of Control Rods for PWR Burn-up Credit” US NRC, February 2002, (ADAMS: ML020810111)
391. NUREG/CR-6760 “Study of the Effect of Integral Burnable Absorbers for PWR Burn-up Credit” US NRC, March 2002, (ADAMS: ML020770436)
392. NUREG/CR-6761, J. C. Wagner, C. V. Parks “Parametric Study of the Effect of Burnable Poison Rods for PWR Burnup Credit” US NRC, March 2002
393. NUREG/CR-6764 “Burn-up Credit PIRT Report”, US NRC, May 2002, (ADAMS: ML021510370)
394. NUREG/CR-6768 “Spent Nuclear Fuel Transportation Package Performance Study Issues Report” US NRC, Manuscript complete: January 31, 2001, Publication Date: Unknown
395. NUREG/CR-6777 “Results and Analysis of The ASTM Round Robin On Reconstitution”, US NRC, August 2002, (ADAMS: ML022540225)
396. NUREG/CR-6781 “Recommendations on the Credit for Cooling Time in PWR Burnup Credit Analyses” US NRC, January 2003, (ADAMS: ML030290585)
397. NUREG/CR-6798 “Isotopic Analysis of High-Burnup PWR Spent Fuel Samples From the Takahama-3 Reactor”, US NRC, January 2003, (ADAMS: ML030570346)
398. NUREG/CR-6800 “Assessment of Reactivity Margins & Loading Curves for PWR Burnup-Credit Cask Designs”, US NRC, March 2003, (ADAMS: ML031110280)
399. NUREG/CR-6801 “Recommendations for Addressing Axial Burnup in PWR Burnup Credit Analyses”, US NRC, March 2003, (ADAMS: ML031110292)
354
400. NUREG/CR-6802 “Recommendations for Shielding Evaluations for Transport & Storage Packages”, US NRC, March 2003, (ADAMS: ML031330514)
401. NUREG/CR-6811 “Strategies for Application of Isotopic Uncertainties in Burnup Credit” US NRC, June 2003, (ADAMS: ML032130638)
402. NUREG/CR-6831 “Examination of Spent PWR Fuel Rods After 15 Years in Dry Storage”, US NRC, September 2003, (ADAMS: ML032731021)
403. NUREG/CR-6832 “Regulatory Effectiveness of Unresolved Safety Issue (USI) A-45, “Shutdown Decay Heat Removal Requirements”, US NRC, August 2003
404. NUREG/CR-6835 “Effects of Fuel Failure on Criticality Safety and Radiation Dose for Spent Fuel Casks” US NRC, September 2003, (ADAMS: ML032880058)
405. NUREG/CR-6842 “Advanced Reactor Licensing: Experience with Digital I&C Technology in Evolutionary Plants”, US NRC, April 2004
406. NUREG/CR-6845 “Sensitivity Analysis Applied to the Validation of the 10B Capture Reaction in Nuclear Fuel Casks” US NRC, August 2004, (ADAMS: ML042530008)
407. NUREG/CR-6888 “Emerging Technologies in Instrumentation and Controls: An Update” US NRC, January 2006
408. NUREG/CR-6890 “Reevaluation of Station Blackout Risk at Nuclear Power Plants” US NRC, January 2006
409. NUREG/CR-6944 “Next Generation Nuclear Plant Phenomena Identification and Ranking Tables (PIRTs)” Vol.1 – Vol. 6, US NRC, March 2008
410. NUREG/CR-6951 G. Radulescu, D. E. Mueller, and J. C. Wagner “Sensitivity and Uncertainty Analysis of Commercial Reactor Criticality for Burnup Credit” US NRC, January 2008.
411. NUREG/CR-6955 J. C. Wagner “Criticality Analysis of Assembly Misload in a PWR Burnup Credit Cask” US NRC, January 2008 (Manuscript complete May 2004)
412. “Advanced Reactor Research Plan”, Rev. 1, Draft, Attachment, Office of Nuclear Regulatory Research, US NRC, June 2002, (ADAMS: ML021760135)
413. “Reactor Concepts Manual 1: Nuclear Power for Electrical Generation” US NRC Technical Training Center, Rev. 0703, Access Electronically
414. “Reactor Concepts Manual 2: The Fission Process and Heat Production” US NRC Technical Training Center, Rev. 0703, Access Electronically
415. “Reactor Concepts Manual 3: Boiling Water Reactor (BWR) Systems” US NRC Technical Training Center, Rev. 0703, Access Electronically
416. “Reactor Concepts Manual 4: Pressurized Water Reactor (PWR) Systems” US NRC Technical Training Center, Rev. 0703, Access Electronically
417. “Reactor Concepts Manual 11: Transportation of Radioactive Materials” US NRC Technical Training Center, Rev. 0703, Access Electronically
ORNL Documents 418. M.D. DeHart, B.L. Broadhead “Investigation of Burnup Credit Issues in BWR Fuel”
Submitted to the ICNC’99 Sixth International Conference on Nuclear Criticality Safety, Palais des congrès, Versailles, France, September 20–24, 1999,
419. J.C. Wagner, C.V. Parks, "Impact of Burnable Poison Rods on PWR Burnup Credit Criticality Safety Analyses," ANS Transactions, 83, 130-134, November 2000. (Best paper award from the Nuclear Criticality Safety Division).
355
420. C.V. Parks, I.C. Gauld, J.C. Wagner, B.L. Broadhead, M.D. DeHart, D.D. Ebert, “Research Supporting Implementation of Burnup Credit in the Criticality Safety Assessment of Transport and Storage Casks”, US NRC, Proceedings of the Twenty-Eighth Water Reactor Safety Information Meeting, Bethesda, Maryland, October 23-25, 2000
421. J.C. Wagner, M.D. DeHart, B.L. Broadhead, “Investigation of Burnup Credit Modeling Issues Associated With BWR Fuel”, ORNL/TM-1999/193, UT-Battelle, Oak Ridge National Laboratory, October 2000.
422. J.C. WAGNER and M.D. DeHart, ”Investigation of BWR Depletion Calculations with SAS2H”, ANS Transactions, 82, 173-176, June 2000.
423. J.C. Wagner, M.D. DeHart, “Review of Axial Burnup Distribution Considerations for Burnup Credit Calculations”, ORNL/TM-1999/246, Lockheed Martin Energy Research Corp., Oak Ridge National Laboratory, March 2000.
424. M.D. DeHart “A Statistical Method for Estimating the Net Uncertainty in the Prediction of K Based on Isotopic Uncertainties” Submitted to the, American Nuclear Society ANS/ENS 2000 International Winter Meeting and Embedded Topical Meetings, November 12–16, 2000, Washington, D.C.
425. C.V. Parks, M.D. DeHart, J.C. Wagner, “Phenomena and Parameters Important to Burnup Credit”, Proceedings of the Technical Committee Meeting on the Evaluation and Review of the Implementation of Burnup Credit in Spent Fuel Management Systems, IAEA-TECDOC-1241, 2001, p. 233-247
426. C.V. Parks, J.C, Wagner, "Issues for Effective Implementation of Burnup Credit," Proc. of the Technical Committee Meeting on the Evaluation and Review of the Implementation of Burnup Credit in Spent Fuel Management Systems, IAEA-TECDOC-1241, 2001, p. 298-308.
427. C.V. Parks, J.C. Wagner, I.C. Gauld, B.L. Broadhead, C.E Sanders, “U.S. Regulatory Research Program for Implementation of Burnup Credit in Transport Casks”, Proceedings of the 13th International Symposium on the Packaging and Transport of Radioactive Materials (PATRAM2001), Chicago, IL, September 3-7, 2001.
428. J.C. Wagner, C.V. Parks, “Critical Review of the Practice of Equating the Reactivity of Spent Fuel to Fresh Fuel in Burnup Credit Criticality Safety Analyses for PWR Spent Fuel Pool Storage”, Nuclear Technology. 136(1), 130-140, October 2001.
429. C.E. Sanders, . J.C. Wagner, “Parametric Study of Control Rod Exposure for PWR Burnup Credit Criticality Safety Analyses”, 2001 ANS Embedded Topical Meeting on Practical Implementation of Nuclear Criticality Safety Analyses, Reno, NV, November 11-15, 2001
430. C.E. Sanders, . J.C. Wagner, "Impact of Integral Burnable Absorbers on PWR Burn-up Credit Criticality Safety Analyses," Proceedings of 2001 ANS Embedded Topical Meeting on Practical Implementation of Nuclear Criticality Safety Analyses, Reno, NV, November 11-15, 2001.
431. J.C. Wagner, “Addressing the Axial Burnup Distribution in PWR Burnup Credit Criticality Safety Analyses” 2001 ANS Embedded Topical Meeting on Practical Implementation of Nuclear Criticality Safety Analyses, Reno, NV, November 11-15, 2001.
432. C.E. Sanders, J.C. Wagner, “Investigation of Average and Pin-Wise Burnup Modeling of PWR Fuel”, ANS Transactions, 86, 2002, (pp. 98-100)
356
433. J.C. Wagner, C.E. Sanders, “Investigation of the Effect of Fixed Absorbers on the Reactivity of PWR Spent Nuclear Fuel for Burnup Credit”, Nuclear Technology, 139(2), August 2002, (pp. 91-126)
434. J.C. Wagner “Evaluation of Burnup Credit for Accommodating PWR Spent Nuclear Fuel in High-Capacity Cask Designs”, Proceedings of the 7th International Conference on Nuclear Criticality Safety (ICNC2003), Tokai-mura, Japan, October 20-24, 2003. (pp. 684-689)
435. J.C. Wagner, “Impact of Soluble Boron Modeling for PWR Burnup Credit Criticality Safety Analyses”, ANS Transactions, 89, 2003, (pp. 120-122).
436. C. V. Parks, C. J. Withee “Recommendations for PWR Storage and Transportation Casks That Use Burnup Credit”, Transactions of “2003 International High-Level Radioactive Waste Management Conference, “Progress Through Cooperation,” Las Vegas, Nevada, March 30–April 2, 2003
437. C.V. Parks, J.C. Wagner "Current Status and Potential Benefits of Burnup for Spent Fuel Transportation," Proceedings of the 14th Pacific Basin Nuclear Conference, March 21-25, Honolulu, Hawaii, ANS Order # 700305, ISBN: 0-89448-679-9 (2004).
438. C.V. Parks, J.C. Wagner “Status of Burnup Credit for Transport of SNF in the United States,” presented at the 14th International Symposium on the Packaging and Transportation of Radioactive Materials, Berlin, Germany, September 20-24, 2004.
439. D.E. Mueller, J.C. Wagner “Application of Sensitivity/Uncertainty Methods to Burnup Credit Criticality Validation,” Proceedings of the IAEA Technical Meeting on Advances in Applications of Burnup Credit to Enhance Spent Fuel Transportation, Storage, Reprocessing and Disposition, August 29-September 2, 2005, London, U.K., IAEA-TECDOC-1547, ISBN 92-0-103307-9, Date of Issue: June 21, 2007.
440. C.V. Parks, C.J. Withee, J.C. Wagner “U.S. Regulatory Recommendations for Actinide-Only Burnup Credit in Transport and Storage Casks”, Proceedings of the IAEA Technical Meeting on Advances in Applications of Burnup Credit to Enhance Spent Fuel Transportation, Storage, Reprocessing and Disposition, August 29-September 2, 2005, London, U.K., IAEA-TECDOC-1547, ISBN 92-0-103307-9, Date of Issue: June 21, 2007.
441. C.V. Parks, J.C. Wagner “A Coordinated U.S. Program to Address Full Burnup Credit in Transport and Storage Casks”, Proceedings of the IAEA Technical Meeting on Advances in Applications of Burnup Credit to Enhance Spent Fuel Transportation, Storage, Reprocessing and Disposition, August 29-September 2, 2005, London, U.K., IAEA-TECDOC-1547, ISBN 92-0-103307-9, Date of Issue: June 21, 2007.
442. J.C. Wagner, D.E. Mueller “Assessment of Benefits for Extending Burnup Credit in Transporting PWR Spent Nuclear Fuel in the USA,” Proceedings of the IAEA Technical Meeting on Advances in Applications of Burnup Credit to Enhance Spent Fuel Transportation, Storage, Reprocessing and Disposition, August 29-September 2, 2005, London, U.K., IAEA-TECDOC-1547, ISBN 92-0-103307-9, Date of Issue: June 21, 2007.
443. J.C. Wagner, D.E. Mueller "Updated Evaluation of Burnup Credit for Accommodating PWR Spent Nuclear Fuel to High-Capacity Cask Designs," presented at the 2005 NCSD Topical Meeting, Knoxville, TN, Sept. 19-22, 2005.
444. C.V. Parks, J.C. Wagner, D.E. Mueller, "Full Burnup Credit in Transport and Storage Casks: Benefits and Implementation," Proceedings of the International High-Level Radioactive Waste Management Conference, Las Vegas, NV, April 30-May 4, 2006, (pp. 1299–1308).
445. I.C. Gauld, S.M. Bowman, B.D. Murphy, P. Schwalbach “Application of ORIGEN to Spent Fuel Sefeguards and Non-proliferation”, Proceedings of INMM 47th Annual Meeting, Nashville, TN, July 16–20, 2006.
357
446. M.D. Muhlheim, R.T. Wood “Design Strategies and Evaluation for Sharing Systems at Multi-Unit Plants” Phase I Report, ORNL/LTR/INERI-BRAZIL/06-01, Oak Ridge National Libratory, August 2007
447. G. Ilas, I.C. Gauld, V. Jodoin “LWR Cross Section Libraries for ORIGEN-ARP in SCALE 5.1” ANS Transactions, 95, 706 (2006).
448. B.L. Broadhead “K-infinite Trends with Burnup, Enrichment, and Cooling Time for BWR Fuel Assemblies”, ORNL/M-6155, Oak Ridge National Laboratory, August 1998.
449. B. D. Murphy, J. Kravchenko, A. Lazarenko, A. Pavlovitchev, V. Sidorenko, A. Chetverikov ”Simulation of Low-Enriched Uranium (LEU) Burnup in Russian VVER Reactors with the HELIOS Code Package” ORNL/TM-1999-168, Oak Ridge National Laboratory, March 2000.
450. I. C. Gauld “SCALE-4 Analysis of LaSalle Unit 1 BWR Commercial Reactor Critical Configurations” ORNL/TM-1999-247, Oak Ridge National Laboratory, March 2000.
451. J.C. Gehin, J.J. Carbajo, R.J. Ellis “Issues in the Use of Weapons-Grade MOX Fuel in VVER-1000 Nuclear Reactors: Comparison of UO2 and MOX Fuels”, ORNL/TM-2004-223, ORNL, October 2004
452. M. D. DeHart “Sensitivity and Parametric Evaluations of Significant Aspects of Burnup Credit for PWR Spent Fuel Packages” ORNL/TM-12973, ORNL, May 1996
453. C.M. Hopper Overview of Sensitivity and Uncertainty Analysis Methods for Establishing Areas of Applicability and Subcritical Margins Submitted to the American Nuclear Society ANS/ENS 2000 International Winter Meeting and Embedded Topical Meetings, Washington, D.C.November 12-16, 2000
454. T. Greifenkamp, K. Clarno, J. Gehin, “Effect of Fuel Temperature on Eigenvalue Calculations”, Proceedings of the 2008 American Nuclear Society National Student Conference “Expanding the Nuclear Family,” February 28 – March 1, 2008, Texas A&M University, College Station, Texas.
455. G. Ilas, I.C. Gauld, “Analysis of Decay Heat Measurements for BWR Fuel Assemblies”, ANS Transactions, 94, 2006, (pp.385–387).
456. B. D. Murphy “Characteristics of Spent Fuel From Plutonium Disposition Reactors, Vol. 4: Westinghouse Pressurized-Water-Reactor Fuel Cycle Without Integral Absorber”, ORNL/TM-13170/V4, Lockheed Martin Energy Research Corp., Oak Ridge National Laboratory, April 1998.
3D Coupling of Neutronics and Thermal Hydraulics 457. P. Kral, J. Hadek, J. Macek “TMI-1 MSLB Coupled 3-D Neutronics / Thermal hydraulics
Analysis: Application of RELAP5-3D and Comparison with Different Codes” Proceedings of RELAP5 International Users Seminar was held in Sun Valley, Idaho, September 5-7, 2001
458. C. Parisi, M. Cherubini, F. D’Auria “Development of a 3D Neutron Kinetic-Thermal-hydraulic model for an RBMK reactor by RELAP5-3D code”, Proceedings of International RELAP5 Users Seminar, 16-18 August in West Yellowstone, Montana, 2006
459. J. Judd “Coupling of RELAP5-3D and SIMULATE-3K for Transient Analysis” Proceedings of International RELAP5 Users Seminar, 7-9 September in Jackson Hole, WY 2005
358
460. “Further Development and Verification/Validation of Computational Programs for Reactor Safety” A report on the status of the development and the scientific continuation of work of reactor safety research funded by the German Government, November 30, 2005
461. P.J. Turinsky “Nuclear Fuel Management optimization: A work in progress” Nuclear Technology, ans International journal of the American Nuclear Society, Vol. 151, No 1, NUTYBB 151(1), July 2005,
462. H.S. Abdel-Khalic, P.J. Turinsky “Adaptive Core Simulation Employing Discrete Inverse Theory – Part 1: Theory” Nuclear Technology, ANS International journal of the American Nuclear Society, Vol. 151, No 1, NUTYBB 151(1), July 2005,
463. H.S. Abdel-Khalic, P.J. Turinsky “Adaptive Core Simulation Employing Discrete Inverse Theory – Part 2: Numerical Experiments” Nuclear Technology, ANS International journal of the American Nuclear Society, Vol. 151, No 1, NUTYBB 151(1), July 2005,
464. R. Gregg, A. Worrall “Effects of Highly Enriched/Highly burnt UO2 fuels of Fuel Cycle costs, Radio toxicity and Nuclear Design Parameters” Nuclear Technology, ans International journal of the American Nuclear Society, Vol. 151, No 2, NUTYBB 151(2), August 2005,
465. A. Hämäläinen "Applying thermal hydraulics modeling in coupled processes of nuclear power plants" Dissertation for the degree of Doctor of Technology, Department of Energy and Environmental Technology at Lappeenranta University of Technology, VTT PUBLICATIONS 578, Lappeenranta, Finland, Nov. 2005
466. H. Okuno Classification of Criticality Calculations with Correlation Coefficient Method and Its Application to OECD/NEA Burnup Credit Benchmarks Phase III-A and II-A” Journal of Nuclear Science and Technology, Vol. 40, No. 7, July 2003 (p. 544–551) (Received Feb-7, 2003; Accepted Apr-18, 2003)
467. DOE Fundamentals Handbook “Nuclear Physics and Reactor Theory” Vol. 1 & 2 DOE-HDBK-1019/1-93, US DOE, Washington D.C., January 1993
468. R. Schenkel "Nuclear Measurements" Presentation Slides for IAEA SCIENTIFIC FORUM 2005 "Nuclear Science: Physics Helping the World", Vienna, Austria. 27 - 28 September 2005
469. Safety Assessment of General Design Aspects of NPPs (Part 1) IAEA Training Course on Safety Assessment of NPPs to Assist Decision Making PowerPoint Slides, Date of Publication is Unavailable.
470. N. Aksan “Overview on Some Aspects of Safety Requirements and Considerations for Future Nuclear Reactors”, IAEA Course on Natural Circulation in Water-Cooled Nuclear Power Plants, International Center for Theoretical Physics (ICTP), Trieste, Italy, June 25-29 2007, Paper ID. T22
471. F. Jatuff “Neutron Capture in 238U in BWR Fuel: PROTEUS Insights in 2005” LRS Scientific Advisory Committee Meeting Paul Scherrer Institute, OVGA/407, February-27 2006
472. A. Waris, R. Kurniadi, Z. Su’ud “Plutonium and Minor Actinides Recycling in Standard BWR using Equilibrium Burnup Model” ITB J. Sci. Vol. 40 A, No. 1, 2008, (pp.15-23)
473. M. Hamasaki, K. Sakashita, T. Natsume “Request from Nuclear Fuel Cycle and Criticality Safety Design” (Publisher and dates are currently unavailable)
474. H. Okuno “Development of a Statistical Method for Evaluation of Estimated Criticality Lower-Limit Multiplication Factor Depending on Uranium Enrichment and H/Uranium-235 Atomic Ratio” Journal of Nuclear Science and Technology, Vol. 44, No. 2, 2007 (p. 137–146) © Atomic Energy Society of Japan
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475. B.T. Rearden, W.J. Anderson, G.A. Harms “Use of Sensitivity and Uncertainty Analysis in the Design of Reactor Physics and Criticality Benchmark Experiments for Advanced Nuclear Fuel”, Nuclear Technology, Vol. 151 August 2005, (pp. 133-158)
476. Transactions of European Nuclear Conference ENC 2007, Brussels, Belgium, Sepember 16-20, 2007
477. Transactions of 2006 International Meeting on LWR Fuel Performance “Nuclear Fuel: Adressing The Future”, TOPFUEL-2006, Salamanca, Spain, October 22-26, 2006
478. “The New Economics of Nuclear Power”, WNA Report, WNU, London, 2005 479. K.S. Smith “Assembly Homogenization Techniques for Light Water Reactor Analysis”
Progress in Nuclear Energy, Vol. 17, No. 3, 1986, (pp. 303-335), Printed in Great Britain 480. R.D. Lawrence “Progress in Nodal Methods for the Solution of the Neutron Diffusion and
Transport Equations” Progress in Nuclear Energy, Vol. 17, No. 3, 1986, (pp. 271-301), Printed in Great Britain
481. R.J.J. Stammler, M.J. Abbate “Methods of Steady-State Reactor Physics in Nuclear Design” Publisher and Date of Publication currently unavailable.
Publications on Russian 482. V.F. Ukraintsev “Reactivity Effects in the Power Reactors”, Training Materials, Obninsk
Institute of Nuclear Power Engineering, Department of professional training and qualification improvement, Obnisk, Russia 2000. (Published on Russian)
483. V.D. Shmelev, Yu.G. Dragunov, V.P. Denisov, I.N. Vasilchenko “WWER Core Designs for Nuclear Power Plants”, IKC “Akademkniga”, Moscow, Russia 2004 (Published on Russian)
484. R.Z. Aminov, V.A. Chrustalev, A.S. Duchovenskiy, A.I. Osadchiy “NNPs with WWER Reactor: Operation Modes, Characteristics, Efficiency”, Energoizdat, Moscow, Russia 1990 (Published on Russian)
485. P.L. Kirillov, Yu.S. Yuriev, V.P. Bobkov “Guide for Thermal-Physics Calculation and Design of the NPP”Energoatomizdat, Moscow, Russia, 1990
486. P.F. Zvifel “Reactor Physics” (Translation from English), Atomizdat, Moscow, Russia, 1977, (Published on Russian)
487. V.I. Naumov “Lectures on Nuclear Reactor Physics: Reactor Kinetics and Safety”, MEPHI, Moscow, Russia 2002 (Published on Russian)
488. V.P. Chromov “Physics and Analysis of Heterogeneous Reactor Core” Collection of the lectures, Department of Experimental and Theoretical Physics of Nuclear Reactors, MEPHI, Moscow, Russia 1999 (Published on Russian).
Publications, Training Materials and Textbooks available in Paper-Copy Only 489. “2005 World Nuclear Industry Handbook”, Nuclear Engineering International, 2005 490. J.J. Duderstadt, L.J. Hamilton “Nuclear Reactor Analysis”
© 1976 John Wiley and Sons, Inc. Published simultaneously in Canada 491. N.E. Todreas, M.S. Kazimi “Nuclear Systems I: Thermal Hydraulic Fundamentals”
Published in 1990 by Taylor and Francis Group, New York, NY, 1990 492. M/ Modarres, M.Kaminskiy, V. Krivtsov “Reliability Engineering and Risk Analysis. A
Practical Guide” Marcell Dekker Inc. New York, 1999
493. M. Benedict, T.H. Pigford, H.W. Levi “Nuclrear Chemical Engineering” Second Edition Published by McGraw-Hill Inc. 1981
360
Policy papers and Energy Strategies 494. T. Ellis “Recommendations for the Increased Utilization of Nuclear Power in the United States
Energy Infrastructure” MIT report for the Washington Internship for Students of Engineering (WISE) program, Sponsored by ANS, October 2004
495. “PGE’s 2006 Integrated Resource Plan” Stakeholder Dialogue No.3 Portland General Electric (PGE), June 12, 2006
496. My Energy Report For WNU-2005 (need to find a copy of the full report and presentation). 497. V. Bragin, J. Carlson, R. Leslie “Building Proliferation-Resistance into the Fuel Cycle”
IAEA-SR-218/52 498. S.V. Mladineo, C.D. Ferguson “Research Memorandum On the Westinghouse AP 1000 Sale to
China and its Possible Military Implications”, Nonproliferation Policy Education Center (NPEC), 2006
499. “Annual Energy Outlook 2007, With Projections to 2030” Energy Information Administration (EAI), US DOE, DOE/EIA-0383(2007), February 2007
500. Harvey W. Graves, Jr. “Nuclear Fuel Management”, 1979, Printed by John Wiley & Sons, Inc.
