ML13004A058

Units 1 and 2, Updated Final Safety Analysis Report, Revision 14, Chapter 4.0 - Reactor

ML13004A058
Person / Time
Site:	Byron, Braidwood
Issue date:	12/14/2012
From:	Exelon Generation Co
To:	Office of Nuclear Reactor Regulation, Office of Nuclear Material Safety and Safeguards
References
RS-12-221
Download: ML13004A058 (302)

Text

B/B-UFSAR 4.0-i REVISION 2 - DECEMBER 1990 CHAPTER 4.0 - REACTOR TABLE OF CONTENTS PAGE 4.0 REACTOR 4.1-1 4.1

SUMMARY

DESCRIPTION 4.1-1

4.1.1 References

4.1-3 4.2 FUEL SYSTEM DESIGN 4.2-1 4.2.1 Design Bases 4.2-2 4.2.1.1 Cladding 4.2-2 4.2.1.2 Fuel Material 4.2-3 4.2.1.3 Fuel Rod Performance 4.2-4 4.2.1.4 Spacer Grids 4.2-4 4.2.1.5 Fuel Assembly 4.2-5 4.2.1.6 Core Components 4.2-7 4.2.1.7 Testing, Irradia tion Demonstration and Surveillance 4.2-9 4.2.2 Description and Design Drawings 4.2-10 4.2.2.1 Fuel Rods 4.2-10a 4.2.2.2 Fuel Assembly Structure 4.2-11a 4.2.2.2.1 Bottom Nozzle 4.2-11a 4.2.2.2.2 Top Nozzle 4.2-12 4.2.2.2.3 Guide and Instrument Thimbles 4.2-12a 4.2.2.2.4 Grid Assemblies 4.2-13a 4.2.2.3 Core Components 4.2-14 4.2.2.3.1 Rod Cluster Control Assembly 4.2-14 4.2.2.3.2 Burnable Absorber Assembly 4.2-15 4.2.2.3.3 Neutron Source Assembly 4.2-16 4.2.2.3.4 Thimble Plug Assembly 4.2-17 4.2.3 Design Evaluation 4.2-17 4.2.3.1 Cladding 4.2-17 4.2.3.2 Fuel Material Consideration 4.2-22 4.2.3.3 Fuel Rod Performance 4.2-23 4.2.3.4 Spacer Grids 4.2-29 4.2.3.5 Fuel Assembl 4.2-29 4.2.3.6 Reactivity Contr ol Assembly and Burnable Absorber Rods 4.2-30

4.2.4 Testing

and Inspection Plan 4.2-33 4.2.4.1 Quality Assurance Plan 4.2-33 4.2.4.2 Quality Control 4.2-33 4.2.4.3 Core Com ponent Testing and Inspection 4.2-36 4.2.4.4 Tests and Inspections by Others 4.2-38 4.2.4.5 Onsite Inspection 4.2-38 4.2.5 References 4.2-38

B/B-UFSAR 4.0-ii REVISION 7 - DECEMBER 1998 TABLE OF CONTENTS (Cont'd)

PAGE 4.3 NUCLEAR DESIGN 4.3-1 4.3.1 Design Bases 4.3-1 4.3.1.1 Fuel Burnup 4.3-2 4.3.1.2 Negative Rea ctivity Feedbacks (Reactivity Coefficient) 4.3-2 4.3.1.3 Control of Power Distribution 4.3-3 4.3.1.4 Maximum Controll ed Reactivity Insertion Rate 4.3-4 4.3.1.5 Shutdown Margins 4.3-5 4.3.1.6 Stability 4.3-6 4.3.1.7 Anticipated Transients Without Trip 4.3-7 4.3.2 Description 4.3-7 4.3.2.1 Nuclear Design Description 4.3-7 4.3.2.2 Power Distributions 4.3-9 4.3.2.2.1 Definitions 4.3-9 4.3.2.2.2 Radial Power Distributions 4.3-11 4.3.2.2.3 Assembly Power Distributions 4.3-12 4.3.2.2.4 Axial Power Distributions 4.3-12 4.3.2.2.5 Deleted 4.3-13 4.3.2.2.6 Limiting Power Distributions 4.3-14 4.3.2.2.7 Experimental V erification of Power Distribution Analysis 4.3-21 4.3.2.2.8 Testing 4.3-23 4.3.2.2.9 Monitoring I nstrumentation 4.3-23 4.3.2.3 Reactivity Coefficients 4.3-23 4.3.2.3.1 Fuel Temperature (Do ppler) Coefficient 4.3-24 4.3.2.3.2 Moderator Coefficients 4.3-25 4.3.2.3.3 Power Coefficient 4.3-26 4.3.2.3.4 Comparison of Ca lculated and Experimental Reactivity Coefficients 4.3-27 4.3.2.3.5 Reactivity Coeff icients Used in Transient Analysis 4.3-27 4.3.2.4 Control Requirements 4.3-28 4.3.2.4.1 Doppler 4.3-28 4.3.2.4.2 Variable Average Mod erator Temperature 4.3-28 4.3.2.4.3 Redistribution 4.3-29 4.3.2.4.4 Void Content 4.3-29 4.3.2.4.5 Rod Insertion Allowance 4.3-29 4.3.2.4.6 Burnup 4.3-29 4.3.2.4.7 Xenon and Samarium Poisoning 4.3-30 4.3.2.4.8 pH Effects 4.3-30 4.3.2.4.9 Experimental Confirmation 4.3-30 4.3.2.4.10 Control 4.3-30 4.3.2.4.11 Chemical Poisoning 4.3-30 4.3.2.4.12 Rod Cluster Control Assemblies 4.3-31 4.3.2.4.13 Reactor Coolant Temperature 4.3-31 4.3.2.4.14 Burnable Absorber Rods 4.3-32 4.3.2.4.15 Peak Xenon Startup 4.3-32 4.3.2.4.16 Load Follow Control and Xenon Control 4.3-32

B/B-UFSAR 4.0-iii REVISION 2 - DECEMBER 1990 TABLE OF CONTENTS (Cont'd)

PAGE 4.3.2.4.17 Burnup 4.3-33 4.3.2.5 Control Rod Patt erns and Reactivity Worth 4.3-33 4.3.2.6 Criticality of t he Reactor During Refueling and Criticality of Fuel Assemblies 4.3-35 4.3.2.7 Stability 4.3-37 4.3.2.7.1 Introduction 4.3-37 4.3.2.7.2 Stability Index 4.3-37 4.3.2.7.3 Prediction of the Core Stability 4.3-38 4.3.2.7.4 Stability Measurements 4.3-38 4.3.2.7.5 Comparison of Calculations with Measurements 4.3-40 4.3.2.7.6 Stability Control and Protection 4.3-41 4.3.2.8 Vessel I rradiation 4.3-42 4.3.3 Analytical Methods 4.3-43 4.3.3.1 Fuel Temperature (Do ppler) Calculations 4.3-43 4.3.3.2 Macroscopic Group Constants 4.3-44 4.3.3.3 Spatial Few-Group Diffusion Calculations 4.3-45 4.3.4 Changes 4.3-46

4.3.5 References

4.3-47

4.4 THERMAL

AND HYDRAULIC DESIGN 4.4-1 4.4.1 Design Basis 4.4-1 4.4.1.1 Departure from Nucleate Boiling Design Basis 4.4-1 4.4.1.2 Fuel Temperature Design Basis 4.4-3 4.4.1.3 Core Flow Design Basis 4.4-3 4.4.1.4 Hydrodynamic Sta bility Design Basis 4.4-4 4.4.1.5 Other Considerations 4.4-4 4.4.2 Description 4.4-4 4.4.2.1 Summary Comparison 4.4-4 4.4.2.2 Critical Heat Fl ux Ratio or Departure from Nucleate Boilin g Ratio and Mixing Technology 4.4-5 4.4.2.2.1 Departure from Nucleate Boiling Technology 4.4-5 4.4.2.2.2 Definition of Departure from Nucleate Boiling Ratio 4.4-6 4.4.2.2.3 Mixing Technology 4.4-7 4.4.2.2.4 Hot Channel Factors 4.4-8 4.4.2.2.5 Effects of Rod Bow on DNBR 4.4-10 4.4.2.2.6 Transition Core 4.4-10 4.4.2.3 Linear Heat Generation Rate 4.4-10a 4.4.2.4 Void Fraction Distribution 4.4-10a 4.4.2.5 Core Coolant Flow Distribution 4.4-11 4.4.2.6 Core Pressur e Drops and Hydraulic Loads 4.4-11 4.4.2.6.1 Core Pressure Drops 4.4-11 4.4.2.6.2 Hydraulic Loads 4.4-11 4.4.2.7 Correlation and Physical Data 4.4-12 4.4.2.7.1 Surface Heat Transfer Coefficients 4.4-12 4.4.2.7.2 Total Core and V essel Pressure Drop 4.4-13

B/B-UFSAR 4.0-iv TABLE OF CONTENTS (Cont'd)

PAGE 4.4.2.7.3 Void Fraction Correlation 4.4-14 4.4.2.8 Thermal Effects of Operational Transients 4.4-15 4.4.2.9 Uncertainties in Estimates 4.4-15 4.4.2.9.1 Uncertainties in Fuel and Cladding Temperatures 4.4-15 4.4.2.9.2 Uncertainties in Pressure Drops 4.4-16 4.4.2.9.3 Uncertainties Due to Inlet Flow Maldistribution 4.4-16 4.4.2.9.4 Uncertainty in DNB Correlation 4.4-16 4.4.2.9.5 Uncertainties in DNBR Calculations 4.4-16 4.4.2.9.6 Uncertainties in Flow Rates 4.4-17 4.4.2.9.7 Uncertainties in Hydraulic Loads 4.4-18 4.4.2.9.8 Uncertainty in Mixing Coefficient 4.4-18 4.4.2.10 Flux Tilt Consideration 4.4-19 4.4.2.11 Fuel and Cladding Temperatures 4.4-19 4.4.2.11.1 UO 2 Thermal Conductivity 4.4-20 4.4.2.11.2 Radial Power Distribution in UO 2 Fuel Rods 4.4-21 4.4.2.11.3 Gap Conductance 4.4-22 4.4.2.11.4 Surface Heat Transfer Coefficients 4.4-23 4.4.2.11.5 Fuel Clad Temperatures 4.4-23 4.4.2.11.6 Treatment of Peaking Factors 4.4-23 4.4.3 Description of the T hermal and Hydraulic Design of the Reactor Coolant System 4.4-24 4.4.3.1 Plant Configuration Data 4.4-24 4.4.3.2 Operating Restrictions on Pumps 4.4-24 4.4.3.3 Power-Flow Operating Map (BWR) 4.4-25 4.4.3.4 Temperature-Power Operating Map 4.4-25 4.4.3.5 Load-Following Characteristics 4.4-25 4.4.3.6 Thermal and Hydr aulic Characteristics Summary Table 4.4-25 4.4.4 Evaluation 4.4-25 4.4.4.1 Critical Heat Flux 4.4-25 4.4.4.2 Core Hydraulics 4.4-25 4.4.4.2.1 Flow Paths Con sidered in Core Pressure Drop and Thermal Design 4.4-25 4.4.4.2.2 Inlet Flow Distributions 4.4-26 4.4.4.2.3 Empirical Friction Factor Correlations 4.4-27 4.4.4.3 Influence of Power Distribution 4.4-27 4.4.4.3.1 Nuclear Enthal py Rise Hot Channel Factor H N F 4.4-28 4.4.4.3.2 Axial Heat Flux Distributions 4.4-29 4.4.4.4 Core Thermal Response 4.4-30 4.4.4.5 Analytical Techniques 4.4-30 4.4.4.5.1 Core Analysis 4.4-30 4.4.4.5.2 Steady-State Analysis 4.4-31 4.4.4.5.3 Experimental Verification 4.4-32 4.4.4.5.4 Transient Analysis 4.4-32 4.4.4.6 Hydrodynamic a nd Flow Power Coupled Instability 4.4-33

B/B-UFSAR 4.0-v REVISION 9 - DECEMBER 2002 TABLE OF CONTENTS (Cont'd)

PAGE 4.4.4.7 Fuel Rod Behavior Effects from Coolant Flow Blockage 4.4-35 4.4.5 Testing and Verification 4.4-36 4.4.5.1 Tests Prior to Initial Criticality 4.4-36 4.4.5.2 Initial Power and Plant Operation 4.4-36 4.4.5.3 Component and Fuel Inspection 4.4-36

4.4.6 Instrumentation

Requirements 4.4-37 4.4.6.1 Incore Instrumentation 4.4-37 4.4.6.2 Overtemperat ure and Overpower T Instrumentation 4.4-37 4.4.6.3 Instrumentation to Limit Maximum Power Output 4.4-38 4.4.6.4 Loose Parts Monitoring System 4.4-38a 4.4.6.4.1 Containment Building Equipment 4.4-39 4.4.6.4.2 Auxiliary Electrical Equipment Room Equipment 4.4-40 4.4.6.4.3 Basis for Alarm Settings 4.4-40 4.4.6.4.4 Operability After Oper ational Basis Earthquake 4.4-41 4.4.6.4.5 Operating Procedures 4.4-42 4.4.6.4.6 Testing 4.4-44 4.4.6.4.7 Training 4.4-44 4.4.7 Reload Safety Evaluations 4.4-44 4.4.8 References 4.4-40 4.4A ADDITIONAL INFORMATI ON ON THE PLANT SPECIFIC APPLICATION OF THE WES TINGHOUSE IMPROVED THERMAL DESIGN PROCEDURE TO BYRON/BRAIDWOOD 4.4A-1

4.5 REACTOR

MATERIALS 4.5-1 4.5.1 Control Rod System S tructural Materials 4.5-1 4.5.1.1 Materials Sp ecifications 4.5-1 4.5.1.2 Austenitic Stainless Steel Components 4.5-3 4.5.1.3 Other Materials 4.5-4 4.5.1.4 Cleaning and Cleanliness Control 4.5-4 4.5.2 Reactor Internals Materials 4.5-4 4.5.2.1 Materials Specification 4.5-4 4.5.2.2 Controls on Welding 4.5-5 4.5.2.3 Nondestructive E xamination of Wrought Seamless Tubular Products and Fittings 4.5-5 4.5.2.4 Fabrication and Proc essing of Austenitic Stainless Steel Components 4.5-5

4.6 FUNCTIONAL

DESIGN OF REACTIVITY CONTROL SYSTEMS 4.6-1

4.6.1 Information

for Control Rod Drive System (CRDS) 4.6-1 4.6.2 Evaluation of the CRDS 4.6-1 4.6.3 Testing and Verification of the CRDS 4.6-1 4.6.4 Information for Comb ined Performance of Reactivity Systems 4.6-2

4.6.5 Evaluation

of Combined Performance 4.6-2 4.6.6 References 4.6-3

B/B-UFSAR 4.0-vi REVISION 6 - DECEMBER 1996 CHAPTER 4.0 - REACTOR LIST OF TABLES NUMBER TITLE PAGE 4.1-1 Reactor Design Table 4.1-4 4.1-2 Analytical Techniques in Core Design 4.1-10 4.1-3 Design Loading Conditi ons Considered for Reactor Core Components 4.1-13 4.3-1 Nuclear Design Key Safety Parameters 4.3-50 4.3-2 Reactivity Requi rements for Rod Cluster Control Assemblies 4.3-52 4.3-3 Benchmark Critical Experiments 4.3-53 4.3-4 Axial Stability Inde x Pressurized Water Reactor Core With a 12-Foot Height 4.3-54 4.3-5 Typical Neutron Flux Levels (n/cm 2-sec) 4.3-55 at Full Power 4.3-6 Comparison of Measured and Calculated Doppler Defects 4.3-56 4.3-7 Saxton Core II Isotopics Rod MY, Axial Zone 6 4.3-57 4.3-8 Critical Boron Concentrations, (ppm) HZP, BOL 4.3-58 4.3-9 Benchmark Critic al Experiments, B 4 C Control Rod Worth 4.3-59 4.3-10 Comparison of Measured and Calculated Moderator 4.3-61 Coefficients at HZP, BOL 4.4-1 Thermal and Hydraulic Data 4.4-52 4.4-2 Void Fractions at Nominal Reactor Conditions 4.4-55

B/B-UFSAR 4.0-vii REVISION 6 - DECEMBER 1996 CHAPTER 4.0 - REACTOR LIST OF FIGURES NUMBER TITLE 4.2-1 17 x 17 VANTAGE 5/VANT AGE+ Fuel Assembly Cross Section 4.2-2 Deleted 4.2-2a 17 x 17 VANTAGE 5 Fuel Assem bly Outline 4.2-2b 17 x 17 VANTAGE+

with Protective Grid 4.2-3 Deleted 4.2-3a 17 x 17 VANTAGE 5/VANTAGE+ Fuel Rod Assembly Comparison 4.2-4 Plan View of Mid Grid and IFM Grid to Guide Thimble Joint (Bottom View) 4.2-5 Elevation View of Mid and IFM Grid to Guide Thimble Joint 4.2-6 Top Grid and Reconstitut able Top Nozzle Attachment Detail 4.2-7 Guide Thimble to Bottom Grid and Nozzle Joint 4.2-8 Rod Cluster Control and Drive Rod 4.2-9 Rod Cluster Cont rol Assembly Outline 4.2-10 Absorber Rod Design 4.2-11 Burnable Absorber Assembly 4.2-12 Wet Annular Burn able Absorber Rod 4.2-13 Burnable Absor ber Rod Sections 4.2-14 Primary Source Assembly 4.2-15 Secondary Source Assembly (Four Secondary Source Rods) 4.2-15a Secondary Source Assembly (Six Seconda ry Source Rods) 4.2-16 Thimble Plug Assembly 4.3-1 Typical Fuel L oading Arrangement 4.3-2 Production and Consump tion of Higher Isotopes 4.3-3 Critical Boron Concentra tion Versus Cycle Burnup 4.3-4 Typical Discrete Burnable Absorber Rod Arrangement Within An Assembly 4.3-5 Integral Fuel Burnable Absorber Rod Arrangement Within an Assembly 4.3-6 Burnable Absorber Loading Pattern (Typical) 4.3-7 Normalized Power Density Dis tribution Near B eginning of Life, Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-8 Normalized Power Density Dis tribution Near B eginning of Life, Bank D at Insertion Limit, Hot Full Power, Equilibrium Xenon 4.3-9 Normalized Power Density Dis tribution Near Middle of Life, Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-10 Normalized Power Density Dis tribution Near E nd of Life, Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-11 Normalized Power Density Dis tribution Near E nd of Life, Bank D at Insertion Limit, Hot Full Powe r, Equilibrium Xenon B/B-UFSAR 4.0-vii (Cont'd) REVISI ON 6 - DECEMBER 1996 LIST OF FIGURES (Cont'd)

NUMBER TITLE 4.3-12 Rodwise Power Distribution in a Typical Assembly (Assembly F-11) Near Beginning of Li fe, Hot Full Power, Equilibrium Xenon, Unrodded Core 4.3-13 Rodwise Power Di stribution in a Typical (Assembly F-11) Near End of Life, Hot Full Power, Equilibrium Xenon, Unrodded Core 4.3-14 Typical Axial Po wer Shapes Occurring At Beginning of Life 4.3-15 Typical Axial Po wer Shapes Occurring at Middle of Life 4.3-16 Typical Axial Power Shapes Occur ring at End of Life 4.3-17 Deleted 4.3-18 Deleted 4.3-19 Deleted 4.3-20 Maximum F Q X Power Versus Ax ial Height D uring Normal Core Operation

B/B-UFSAR 4.0-viii REVISION 9 - DECEMBER 2002 LIST OF FIGURES (Cont'd)

NUMBER TITLE 4.3-21 Peak Linear Power Du ring Control Rod Malfunction Overpower Transients 4.3-22 Peak Linear Power Du ring Boration/Deboration Overpower Transients 4.3-23 Typical Comparis on Between Calcu lated and Measured Relative Fuel Assembly Power Distribution 4.3-24 Comparison of Calculated and M easured Axial Shape 4.3-25 Comparison of Calculated and M easured Peaking Factors, [F Q x PREL] MAX Envelope as a Function of Core Height 4.3-26 Doppler Temper ature Coefficient at BOL and EOL 4.3-27 Doppler Only Power Coefficie nt BOL and EOL 4.3-28 Doppler Only Power Defect BOL and EOL 4.3-29 Moderator Temperature Co efficient BOL, All Rods Out 4.3-30 Moderator Temperature Co efficient EOL, All Rods Out 4.3-31 Moderator Temperature Co efficient As a Function of Boron Concentration BO L, All Rods Out 4.3-32 Hot Full Power Moderator Temperature Coefficient Versus Critical Boron Concentration 4.3-33 Total Power Co efficient BOL and EOL 4.3-34 Total Power De fect BOL and EOL 4.3-35 Rod Cluster Cont rol Assembly Pattern 4.3-36 Accidental Sim ultaneous Withdrawal of Two Control Banks EOL, Hot Z ero Power, Banks C and B Moving in Same Plane 4.3-37 Deleted 4.3-38 Deleted 4.3-39 Axial Offset V ersus Time PWR Core With a 12-foot Height and 1 21 Assemblies 4.3-40 XY Xenon Test Thermocoup le Response Quadrant Tilt Difference Versus Time 4.3-41 Calculated and Measured Dopp ler Defect and Coefficients at BOL, 2-L oop Plant, 121 Assemblies 12-foot Core 4.3-42 Comparison of Calcul ated and Measured Boron Concentration for 2-Loop Plant, 121 Assemblies, 12-foot Core 4.3-43 Comparison of Ca lculated and Measured C B 3-Loop Plant, 157 Assem blies, 12-foot Core 4.3-44 Comparison of Ca lculated and Measured C B 4-Loop Plant, 193 Assemblies, 12-foot Core 4.4-1 Deleted 4.4-2 Measured Versus Predicted Critical Heat Flux WRB-1 Correlation 4.4-2a Measured Versus Predicted Critical Heat Flux WRB-2 Correlation 4.4-3 TDC Versus Reynolds Nu mber For 26" Grid Spacing 4.4-4 Normalized Radial Flow a nd Enthalpy Rise Distribution at 4 ft Elevation

B/B-UFSAR 4.0-viii(Cont'd) REVISI ON 2 - DECEMBER 1990 LIST OF FIGURES (Cont'd)

NUMBER TITLE 4.4-5 Normalized Radial Flow a nd Enthalpy Rise Distribution at 8 Ft Elevation 4.4-6 Normalized Radial Flow a nd Enthalpy Rise Distribution at 12 Ft Elevation

B/B-UFSAR 4.0-ix REVISION 9 - DECEMBER 2002 LIST OF FIGURES (Cont'd)

NUMBER TITLE 4.4-7 Void Fraction Versus Thermodynamic Quality H-HSAT/H g-H SAT 4.4-8 Thermal Conduc tivity of UO 2 (Data Corrected to 95% Theoretical Density) 4.4-9 Reactor Coolant System Temperature P ercent Power Map 4.4-9a Deleted 4.4-10 Distribution of Incore Instrumentation 4.4-11 100 Percent Power Shap es Evaluated at Conditions Representative of Loss of Flow all Sha pes Evaluated with F N H= 1.49

B/B UFSAR 4.1-4 REVISION 9 - DECEMBER 2002 TABLE 4.1-1 REACTOR DESIGN TABLE THERMAL AND HYDRAULIC DESIGN PARAMETERS

1. Reactor Core Heat Output, (100%), MWt 3586.6 2. Reactor Core Heat Output, 10 6 BTU/Hr 12238.2 3. Heat Generated in Fuel, % 97.4 4. Core Pressure, Nominal psia 2270

5. Pressurizer Pressure, Nominal, psia 2250

6. Minimum DNBR at Nominal Conditions Typical Flow Channel 2.25 Thimble (Cold Wall) Flow Channel 2.16 7. Minimum DNBR for Design Transients Typical Flow Channel 1.33 Thimble Flow Channel 1.33 8. DNB Correlation(c) WRB-2 COOLANT FLOW

9. Total Vessel F low Rate, 10 6 lbm/hr (based on Minimum Measured Flow) 141.8 (based on Thermal Design Flow) 137.2 10. Effective Flow Rate for Heat Transfer, 10 6 lbm/hr (based on TDF) 126.6 11. Effective Flow Area for Heat Transfer, ft 2 54.1 12. Average Velocity along Fuel Rods, ft/sec (based on TDF) 15.1 B/B UFSAR 4.1-5 REVISION 9 - DECEMBER 2002 TABLE 4.1-1 (Cont'd)

REACTOR DESIGN TABLE THERMAL AND HYDRAULIC DESIGN PARAMETERS

COOLANT TEMPERATURE

°F 13. Average Mass Velocity, 10 6 lbm/hr-ft 2 (based on TDF) 2.33 14. Nominal Inlet 556.7 15. Average Rise in Vessel 62.6 16. Average Rise in Core 66.3 17. Average in Core 591.7 18. Average in Vessel 588.0 HEAT TRANSFER

19. Active Heat Transfer Surface Area, ft 2 57505 20. Average Heat F lux, BTU/hr-ft 2 207327 21. Maximum Heat Flux for Normal Operation,(f) BTU/hr-ft 2 539050 22. Average Linear Power, kw/ft 5.73 23. Peak Linear Power for Normal Operation,(f) kw/ft 14.9 (not including 2% calorimetric uncertainty) 24. Peak Linear Power Resulting from (g) Overpower Transients/Operator Errors (assuming a maximum overpower of 118%), kw/ft <22.4 (Centerline melt will not be exceeded)

25. Peak Linear Power for Prevention of Centerline Melt, kw/ft 22.4

26. Power Density, kw per kg Uranium 108.7 (hot) 27. Specific Power, kw per kg Uranium 44.1 B/B UFSAR 4.1-6 REVISION 9 - DECEMBER 2002 TABLE 4.1-1 (Cont'd)

REACTOR DESIGN TABLE THERMAL AND HYDRAULIC DESIGN PARAMETERS

28. Temperature at Peak Linear Power for Prevention of Centerline Melt, °F 4700 29. Pressure Drop Across Core, psi (i) 27.5 + 2.7 Across Vessel, including nozzle, psi 46.1 +/- 4.6 30. Design RCC Canless 17 x 17

31. Number of Fuel Assemblies 193 32. UO 2 Rods per Assembly 264 33. Rod Pitch, in. 0.496 34. Overall Dimensions, in. 8.426 x 8.426 35. Fuel Weight (as UO 2), lb 204,236 36. Clad Weight, lb 43,376 37. Number of Grids per Assembly 11/12* 38. Composition of Grids 2 End Grids - Inconel 718 6 Intermediate and 3 flow mixer grids Zircaloy 4/

ZIRLO*

39. Loading Technique 3 Region Nonuniform

• VANTAGE+ includes one Inconel protect ive bottom grid.

B/B UFSAR 4.1-7 REVISION 9 - DECEMBER 2002 TABLE 4.1-1 (Cont'd)

REACTOR DESIGN TABLE THERMAL AND HYDRAULIC DESIGN PARAMETERS

CORE MECHANICAL DESIGN PARAMETRS

FUEL RODS

40. Number 50,952 41. Outside Diameter, in. 0.360 42. Diametral Gap, in. 0.0062 43. Cladding Thickness, in. 0.0225 44. Cladding Material Zircaloy-4/ ZIRLO

FUEL PELLETS

45. Material UO 2 Sintered 46. Density (% of Theoretical) 95 47. Diameter, in. 0.3088 48. Length, in. - Midzone Enriched Fuel 0.370 - Blanket Fuel 0.462/0.500 ROD CLUSTER CONTROL ASSEMBLIES

49. Neutron Absorber Ag-In-Cd or Hafnium

50. Cladding Material Type 304 SS-Cold Worked

51. Cladding Thickness, in. 0.0185 52. Number of Clusters 53 53. Number of Absorber Rods per Cluster 24

B/B UFSAR 4.1-8 REVISION 9 - DECEMBER 2002 TABLE 4.1-1 (Cont'd)

REACTOR DESIGN TABLE THERMAL AND HYDRAULIC DESIGN PARAMETERS

CORE STRUCTURE

54. Core Barrel, ID/OD, in. 148.0/152.5 55. Thermal Shield Neutron Pad Design STRUCTURE CHARACTERISTICS

56. Core Diameter, in. (Equivalent) 132.7

57. Core Height, in. (Active Fuel, Cold Dimensions) 144

REFLECTOR THICKNESS AND COMPOSITION

58. Top - Water plus Steel, in.

~10 59. Bottom - Water p lus Steel, in.

~10 60. Side - Water plus Steel, in.

~15 61. H 2 O/U Molecular Ratio, Cell (Cold) 2.73

62. H 2 O/U Molecular Ratio, Core Average, Cold (first core) 3.16

FEED ENRICHMENT, W/O

63. Typical Split Batch Typical Enriched Zone 4.40 to 4.95 Axial Blanket Range 0.74 to 3.20 ____________________ (a) Deleted.

(b) Deleted.

(c) The W-3 correlation is used for analysis of some accidents outside the ra nge of applicati on for the WRB-2 DNB correlation.

B/B UFSAR 4.1-9 REVISION 9 - DECEMBER 2002 TABLE 4.1-1 (Cont'd)

REACTOR DESIGN TABLE (d) Deleted.

(e) Deleted.

(f) This limit is associ ated with the value of Q T F = 2.60. (g) See Subsection 4.3.2.2.6.

(h) See Subsecti on 4.4.2.11.6.

(i) Based on best estima te reactor flow rate.

B/B-UFSAR

4.1-10 REVIS ION 9 - DECEMBER 2002 TABLE 4.1-2 ANALYTICAL TECHN IQUES IN CORE DESIGN SECTION ANALYSIS TECHNIQUE COMPUTER CODE REFERENCED Fuel Rod Design Fuel Performance Semiempirical thermal Westinghouse fuel rod 4.2.1.1 Characteristics Model design model 4.2.3.1 (temperature, Model of fuel rod with 4.2.3.2 internal pressure consideration of fuel 4.3.3.3 cladding stress, density changes, heat 4.3.3.1 etc.) transfer, fission gas 4.4.2.11 release, etc.

Nuclear Design

1. Cross Sections Microscopic data Modified ENDF/B library 4.3.3.2 and Group Macroscopic constants LEOPARD/CINDER type 4.3.3.2 Constants for homogenized core PHOENIX - P

• 4.3.4 regions Group constants for HAMMER-AIM* 4.3.3.2 control rods with self-shielding

2. X-Y Power 2-D and 3-D, 2-Group TURTLE* 4.3.3.3 Distributions, Diffusion Theory 4.3.4 Fuel Depletion Critical Boron Nodal Code PALADON* 4.3.3.3 Concentrations, X-Y Xenon ANC* Distributions, 4.3.4 Reactivity Coefficients, and Control Rod Worths B/B-UFSAR

4.1-11 REVIS ION 9 - DECEMBER 2002 TABLE 4.1-2 (Cont'd)

SECTION ANALYSIS TECHNIQUE COMPUTER CODE REFERENCED 3. Axial Power 1-D, 2-Group APOLLO* 4.3.4 Distributions Diffusion Theory Control Rod Worths, and 2-D and 3-D 2-Group Nodal PALADON* 4.3.3.3 Axial Xenon Analysis Code Distribution 4. Fuel Rod Power Integral Transport Theory LASER* 4.3.3.1 Effective Monte Carlo Weighting REPAD* Resonance Function Temperature 5. Criticality of 1-D, Multi-group Transport AMPX SYSTEM* 4.3.2.6 Reactor and Theory of Codes 4.3.4 Fuel Assemblies 3-D Monte Carlo KENO-IV Thermal Hydraulic Design

1. Steady-state Subchannel analysis of THINC-IV* 4.4.4.5 local fluid conditions in rod bundles including inertial and crossflow resistance terms, solution progresses from core-wide to hot assembly to hot channel B/B-UFSAR

4.1-12 REVISION 2 - DECEMBER 1990 TABLE 4.1-2 (Cont'd)

SECTION ANALYSIS TECHNIQUE COMPUTER CODE REFERENCED 2. Transient Subchannel analysis of THINC-I (THINC-III)* 4.4.4.5.4 Departure from local fluid conditions Nucleate Boiling in rod bundles during Analysis transients by including accumulation terms in conservation equations; solution progresses from core-wide to hot assembly to hot channel

____________________

• Commonwealth Edison internally applies an in-house naming convention for some or all of these code packages.

B/B-UFSAR 4.1-13 TABLE 4.1-3 DESIGN LOADING CONDITIONS CONSIDERED FOR REA CTOR CORE COMPONENTS

1. Fuel Assembly Weight

2. Fuel Assembly Spring Forces 3. Internals Weight

4. Control Rod Trip (equi valent static load) 5. Differential Pressure

6. Spring Preloads

7. Coolant Flow Forces (static) 8. Temperature Gradients

9. Differences in T hermal Expansion a. Due to tempera ture differences b. Due to expansion of different materials 10. Interference B etween Components 11. Vibration (mechanically or hydraulically induced) 12. One or More Loops Out of Service

13. All Operational Tran sients Listed in Table 5.2-1 14. Pump Overspeed

15. Seismic Loads (operati on basis earthquake and safe shutdown earthquake) 16. Blowdown Forces (due to cold and hot leg break)

B/B-UFSAR 4.2-1 REVISION 12 - DECEMBER 2008 4.2 FUEL SYSTEM DESIGN The plant design conditions are divi ded into four categories in accordance with their anticipated frequency of occurrence and risk to the public: C ondition I - normal op eration; Condition II - incidents of mo derate frequency; Co ndition III - infrequent incidents; Condition IV - limiting fau lts. The bases and description of plant operation and eve nts involving each condition are gi ven in Chapter 15.0.

The reactor is designed so t hat its components meet the following performance and safety criteria:

a. The mechanical desig n of the reactor core components and their physical arr angement, together with corrective actions of the re actor control, protection, and emergency coolin g systems (when ap plicable) assure that: 1. Fuel damage (Note: F uel damage as used here is defined as penetration of the fission product barrier, i.e., the fuel rod clad) is not expected during Condition I and Condition II events. It is not p ossible, however, to preclude a very small nu mber of rod failures.

These are within the c apability of the plant cleanup system and are consistent with plant design bases. The number of rod failures is small enough such that t he dose limits given in 10 CFR 100 and 10 CFR 50.67 will not be exceeded.

2. The reactor can be bro ught to a safe state following a Condition III event with only a small fraction of fuel r ods damaged. (Note:

Fuel damage as used here is defined as penetration of the f ission product barrier, i.e., the fuel rod clad). T he extent of fuel damage might preclude immed iate resumption of operation.

3. The reactor can be broug ht to a safe state and the core can be kept s ubcritical with acceptable heat transfer geometry following transients arising from Conditi on IV events.

b. The fuel assemblies are designed to withstand loads induced during shipp ing, handling, a nd core loading without exceeding th e criteria of Subs ection 4.2.1.5.

c. The fuel assemblies are designed to accept control rod insertions in order to provide t he required reactivity control for power operations and

B/B-UFSAR 4.2-2 REVISION 10 - DECEMBER 2004 reactivity shutd own conditions (if in such core locations).

d. All fuel assemblies have provisions for the insertion of incore instrumentat ion necessary for plant operation (if in suc h core locations).

e. The reactor internals, in conjunction wi th the fuel assemblies and incore co ntrol components, direct coolant through the core.

This achieves acceptable flow distribution and re stricts bypass flow so that the heat transfer performance requirements can be met for all modes of operation.

4.2.1 Design

Bases For both the VANTAGE+ and the VANTAGE 5 Fuel a ssemblies, the fuel rod and fuel assembly de sign bases are established to satisfy the general performance and safety criteria pres ented in this section. Design values for the properties of the materi als which comprise the fuel rod, fuel ass embly and incore contr ol components are given in Reference 2 for Zircaloy clad fuel and in Reference 24 for ZIRLO clad fuel.

Other supplement ary fuel design criteria/limits are given in References 25 a nd 28. Reference 28 is applicable for new fuel r eloads after September 2003.

4.2.1.1 Cladding

a. Material and M echanical Properties Zircaloy-4 and ZIRLO com bine low absorption cross section, high corrosion resi stance to coolant, fuel and fission products, high s trength and ductility at operating temperatures, and high reliability.

Reference 1 documents the operating experience with Zircaloy-4 and ZIRLO as clad materia l, and References 2 and 4 provide their mechanical pro perties with due consideration of temperature and irradiation effects.

b. Stress-Strain Limits Cladding stress - The von Mises criterion is used to calculate the ef fective stresses.

The cladding stresses under Condition I a nd II events are less than the Zircaloy 0.2% o ffset yield stress, with due consideration of temperature a nd irradiation effects.

While the cladding h as some capability for accommodating plasti c strain, the yi eld stress has been accepted as a con servative design basis.

Cladding tensile strain -

The total tensile creep strain is less than 1%

from the unirradiated condition. The elastic tensile strain during a transient is less than 1%

from the p retransient value. This limit is consistent with proven practice.

B/B-UFSAR 4.2-3 REVISION 6 - DECEMBER 1996 c. Vibration and Fatigue Strain fatigue - The cumulative stra in fatigue cycles are less than the design str ain fatigue life. This basis is consistent wi th proven practice.

Vibration - Potential for fr etting wear of the clad surface exists due to flow induced v ibrations. This condition is taken into account in the design of the fuel rod support system.

The clad wear depth is limited to acceptable va lues by the grid support dimple and spring design.

d. Chemical Propert ies of the Cladding This is discussed in Ref erence 2 for Zircaloy and Reference 24 for ZIRLO.

4.2.1.2 Fuel Material

a. Thermal Physical Properties The thermal-physical properties of UO 2 are described in Reference 2 w ith due consideratio n of temperature and irradiation effects.

Fuel pellet temperatures - T he center temperature of the hottest pellet is be low the melting temperature of the UO 2 melting point of 2805

°C (Reference 2) unirradiated and decreasing by 32

°C per 10,000 MWD/MTU. While a limited amount of center melting can be tolerated, th e design conserv atively precludes center melting.

A calculated fu el centerline temperature of 4700

°F has been selected as an overpower limit to assure no fuel melting. This provides sufficient margin for uncerta inties as described in S ubsection 4.4.2.9.

Fuel pellet density - The nominal design density of the fuel is 95%

of theoretical.

b. Fuel Densification a nd Fission P roduct Swelling The design bases and m odels used for fuel densification and swelling a re provided in References 3 and 4. c. Chemical Properties Reference 2 provides the bas is for justi fying that no adverse chemical interaction s occur between the fuel and its adjacent material.

B/B-UFSAR 4.2-4 REVISION 10 - DECEMBER 2004 4.2.1.3 Fuel Rod Performance The detailed fuel ro d design establish es such parameters as pellet size and densit y, cladding-pellet dia metral gap, gas plenum size, and helium prepressurization level.

The design also considers effects such as fu el density chang es, fission gas release, cladding creep, and other physical properti es which vary with burnup. The integrity of the fuel rods is ensured by designing to prevent excessi ve fuel temperatur es, excessive internal rod gas pressures due to fission ga s releases, and excessive cladding str esses and strains. Th is is achieved by designing the fuel rods to satisfy the conse rvative design bases in the following subse ctions during Conditio n I and Condition II events over the fuel lifetime. For each design basis, the performance of the limit ing fuel rod must no t exceed the limits specified.

a. Fuel Rod Models The basic fuel rod models an d the ability to predict operating characteristics are given in Reference 4 and Subsection 4.2.3.

b. Mechanical Design Limits Fuel rod design methodol ogy has been introduced (Reference 26) that redu ces the densif ication power spike factor to 1.0 and demonstrates that clad flattening will not occur in Westinghouse fuel designs. The rod internal gas pressur e shall remain below the value which causes t he fuel-cladding d iametral gap to increase due to outward cladding cre ep during steady-state operation.

Rod pressure is also limited such that extensive DNB propagation shall not occur during normal operation and accident events (Reference 12). F or the Section 15.3.3 Locked Rotor Analysis, a small number of rods are predicted to experience DN B; however, it has been determined that an R CS temperature red uction or an RCS flow increase is sufficient to show no rods in DNB.

By taking credit for one of these parameters and precluding rods in DNB, the DNB prop agation analysis for the locked rotor eve nt is inherently met.

4.2.1.4 Spacer Grids

a. Material Properties and Mechanical Design Limits Two types of spacer (str uctural) grids a re used in each fuel assembly. The top, bottom, and protective bottom grids are made of Incon el 718. The others are made of Zircal oy-4 or ZIRLO.

B/B-UFSAR 4.2-4a REVISION 10 - DECEMBER 2004 Lateral loads resulting from a seismic or LOCA event will not cause unaccep tably high plastic grid deformation. Each fuel assembly's geometry will be maintained such that the fuel rods remain in an array amenable to cooling. The be havior of the grids under loading has been studied exper imentally to establish strength criteria. For the Zircaloy, ZIRLO, and Inconel grids, the limit is the 95%

confidence level of the true mean as taken from the dis tribution of measurements of buckli ng loads at operating temperature.

B/B-UFSAR 4.2-5 REVISION 6 - DECEMBER 1996 b. Vibration and Fatigue The grids provide suffic ient fuel rod support to limit fuel rod vibration and maintain cladding fretting wear to within acceptable limits.

4.2.1.5 Fuel Assembly

a. Structural Design As previously di scussed in Subse ction 4.2.1, the structural integrity of the fuel assemblies is assured by setting design limits on stresses and deformations due to va rious nonoperational, operational and accident loa ds. These limits are applied to the desig n and evaluation of the top and bottom nozzles, guide th imbles, grids, and the thimble joints.

The design bases for e valuating the structural integrity of the fuel assemblies are:

1. Nonoperational - 6g late ral and traverse and 4g longitudinal loading with dimensional stability (Reference 17).

2. Normal and abn ormal loads for Conditions I and II - the fuel assembly component structural design criteria are es tablished for the two primary material categor ies, namely austenitic steels, Zircaloy, and ZIRLO. The stress categories and s trength theory p resented in the ASME Boiler and Pressu re Vessel Code,Section III, are used as a general guide.

For austenitic steel str uctural components, the Tresca criterion is used to determine the stress intensities. The design stress intensity value, S m , is given by the lowe st of the following:

One-third of the speci fied minimum tensile strength or two-thirds of the specified minimum yield strength at room temperature.

One-third of the tensile strength or 90% of the yield strength at operat ing temperature, but not to exceed two-thirds of the spec ified minimum yield strength at room temperature.

The stress intensity lim its are given below. All stress nomenclature is per the ASME Boiler and Pressure Vessel Code ,Section III.

B/B-UFSAR 4.2-6 REVISION 6 - DECEMBER 1996 Stress Intensity Limits Categories Limits General Primary Membrane Stress Intensity S m Local Primary Membrane Stress Intensity 1.5 S m Primary Membrane plus Primary Bending Stress Intensity 1.5 S m Total Primary plus S econdary Stress Intensity Range 3.0 S m The Zircaloy and ZIRLO struc tural components, which consist of guide thimbles, i nner six grids and fuel rods are in turn subdivi ded into two categories because of material di fferences and functional requirements. The fuel rod and grid design criteria are covered separately in Subsections 4.2.1.1 and 4.2.1.4, respectively. For the guide th imble design, the stress intensities, the design stress intensities, and the str ess intensity limits are calculated using the same methods as for the austenitic steel structu ral components. For conservative purpose s, the unirradiated properties of Zircaloy and Z IRLO are used.

3. Abnormal loads d uring Conditions III or IV - worst case represented by seismic loads, or blowdown loads during a LOCA event.

Deflections or failures of components cannot interfere with the react or shutdown or emergency cooling of the fuel rods.

The fuel assembly structural component s tresses under faulted conditions are evaluated usi ng primarily the methods outlined in Ap pendix F of the ASME Boiler and Pressure Vessel Code ,Section III.

For the austenitic steel fuel assembly components, the stress intensity and the design stress intensity value, S m , are defined in accordance with the rules described in the previous section for normal operating conditions. S ince the current analytical methods utilize elastic analysis, the stress intensity limits are defined as the smaller value of 2.4 S m or 0.70 S u for primary me mbrane and 3.6 S m or 1.05 S u for primary membrane p lus primary bending.

For the Zircaloy and ZIRLO components, the stress intensities are defined in accordance with the rules described in the previous section for normal operating conditions, and th e design str ess intensity values, S m , are set at two-thirds of the material yield strength, S y , at reactor operati ng temperature.

This results in Zircaloy and ZIRLO stress intensity limits B/B-UFSAR 4.2-7 REVISION 9 - DECEMBER 2002 being the smaller of 1.6 S y or 0.70 S u for primary membrane and 2.4 S y or 1.05 S u for primary membrane plus bending.

For conservative purposes, the Zircaloy and ZIRLO unirr adiated properti es are used to define the stress limits.

b. Thermal-hydraulic Design This topic is covered in Section 4.4

4.2.1.6 Core Components The core components consists of the rod cluster control assemblies (RCCAs), the primary and secondary source assemblies, the thimble plug assemblies and the bu rnable absorber assemblies.

A description of these components is pro vided in Section 4.2.2.

a. Thermal-Physical Propert ies of the Absorber Material The absorber material for the RCCA is either Ag-In-Cd, or Hafnium.

The thermal-physical properties of Ag-In-Cd are described in Reference 2, and Hafnium properties are described in Reference 16.

The absorber mat erial temperature shall not exceed its minimum melting temperature (1454

°F for Ag-In-Cd and 3913°F for Hafnium).

The burnable absorbe r material is ei ther borosilicate glass in burnable absorber rods, or aluminum oxide boron carbide pellets for th e wet annular burnable absorber (WABA) rods.

The thermal-physical properties of the borosilica te glass are described in Reference 2, and those of the WABA in Reference 15.

The burnable absorbe r rods are desig ned so that the borosilicate glass tempe rature is below its minimum softening temperature of 1492

°F (for reference 12.5 weight percent boron). The so ftening temperature is defined in accordanc e with ASTM C 33

8. In addition, the structural elements are designed to prevent excessive slumping.

The WABA rods are design ed so that the maximum temperature is less than 1200

°F. This ensures that the helium gas r elease will not exce ed 30% for the WABA rod mechanical design life. This also assures that the Zircaloy clad s train limit is satisfied.

B/B-UFSAR 4.2-8 REVISION 2 - DECEMBER 1990 b. Compatibility of the Abs orber and Cladding Materials The control rod and borosili cate burnable absorber rod cladding is cold dra wn type 304 st ainless steel tubing, and the WABA rod cla dding is Zircaloy-4.

Extensive in-reactor e xperience and available quantitative information sho w that reaction rates between 304 stainless st eel and water or any contacting metals are negligible at operational temperatures (Refere nces 2 and 16).

c. Cladding Stres s-Strain Limits For Conditions I and II, the stress categories and strength theory presented in the ASME Boiler and Pressure Vessel Code, Se ction III, Subsection NG-3000, are used as a g eneral guide.

The code methodology is applied, as with fuel assembly structural design, where possible. For Conditions III and IV, code stresses are not limiting.

Failures of the burnable abs orber rods during these conditions must not interfere with r eactor shutdown or cooling of the fuel rods.

The deformation or f ailure of the control rod cladding must not prevent reactor shutd own or cooling of the fuel rods. A breach in the cladding does not result in serious conse quences because either the Ag-In-Cd or Hafnium material is relatively inert.

The mechanical design ba ses for the control rods are consistent with the load ing conditions of the ASME Boiler and Pressure Vessel Code, Section III:

1. External pressure equal to the r eactor coolant system operating pressure with appropriate allowance for overpr essure transients.

2. Wear allowance equivalent to 1000 full power reactor trips.

3. Bending of the rod due to a mi salignment in the guide tube.

4. Forces imposed on the rods during rod drop.

5. Loads imposed by the accelerations of the control rod drive mechanism.

6. Radiation exposure d uring maximum core life.

B/B-UFSAR 4.2-9 REVISION 6 - DECEMBER 1996 7. Temperature effects from room to operating conditions.

The burnable absorber as semblies, thimble plug assemblies, and sour ce assemblies ar e static core components. The mecha nical design of these components satisfies the following:

a. Accommodate the diff erential thermal expansion between the fuel assembly and the core internals.

b. Maintain positive co ntact with the fuel assembly and the core internals.

The design evaluation of the core components is discussed in S ubsection 4.2.3.6.

d. Irradiated Behav ior of Absorber Material Operating experience and/or testing evaluation of the effects of irradiation u pon the properties of Ag-In-Cd and Hafnium have shown that in-pile corrosion behavior is simila r to out-of-pile behavior and that, for low oxygen content water, corrosion rates are low (Refer ences 2 and 16).

4.2.1.7 Testing, Irradiation D emonstration a nd Surveillance

An extensive testing p rogram was conducted to verify the adequacy of the predicted fuel performance and the design bases.

Reference 19 provides a descript ion of the tests performed and a summary of t he results.

In addition, in-plan t irradiation demons tration programs have been completed on VANTAGE 5 and VANTAGE+ fue l designs. The objectives of the demonstration programs were to confirm the adequacy of the design and to obtain early performance information. The VANTAGE 5 demo nstration assemblies operated for 3 cycles (including 2- to 18-m onth cycles). Examinations performed at the refueling o utages confirmed excellent performance of the dem onstration assemblies.

The improved corrosion resistance of ZIRLO cladding has b een shown with high burnups in the BR-3 and North Anna demonstra tion assemblies.

Cladding corrosion mea surements showed that the reduced corrosion exhibited by the ZIRLO clad ro ds was better than anticipated.

Full production regions of the VANTAGE+ desi gn have been placed in operation.

B/B-UFSAR 4.2-9a REVISION 13 - DECEMBER 2010 Beginning with Byron U nit 2 Cycle 16, one Lead Test Assembly (LTA) is inserted into a non-limitin g location for r epresentative testing purposes for one cycle. This LTA co nsists of fresh fuel rods in four alloys of AXIOM cladding al ong with 16 twice-burned fuel rods in AXIOM c ladding selected from previously irradiated LTAs. The four alloys are zirconium bas ed with vary ing nominal weight percent composi tion of niobium, t in, iron, chromium, copper, vanadium, and/or nickel.

The LTA is identical to the current 17X17 VANTAGE+ fuel design with the exception that all the fuel rods ar e clad in AXIOM fuel cladding. This LTA is only applicable to Byron Unit 2 Cycle 16.

In one or more of Br aidwood Unit 1 Cycles 15 through 17, eight Lead Use Assemblies (L UAs) will be inser ted into non-limiting core locations f or lead use testing purposes. These LUAs are 17x17 standard latti ce fuel assemblies speci fically designed for Westinghouse-type reacto rs and designated as Advanced Mark-BW(A).

The LUAs will not be placed in core locations with any core components (e.g., RC CA, WABA, secondary sour ces, thimble plugs).

The design of these LUAs inclu de the following features:

1. The LUA features MONOBLOCŽ g uide tubes, which have a constant outer diame ter along with a gradually tapered inner diameter to fo rm a dashpot region.

2. The LUA features Inconel 718 H igh Mechanical Performance (HMP) top and bottom end grids.

3. The LUA features a welded ca ge design with the Mid-Span Mixing Grids directl y welded to the gu ide tubes, the Intermediate Spacer Grids we lded to the guide tubes using weld tabs, and upper and lower HMP end grids.

4. The LUA features fuel rods t hat are not seated on the bottom nozzle at the beginning of life.

5. The LUA features Uranium-Gadol inia fuel rods and axial blanket fuel regions.

6. The LUA features a Modular Quick-Disconnect (QD) sub-assembly for connection of the guide t ubes to the top nozzle. 7. The LUA features the FUELGUA RDŽ bottom n ozzle design.

These LUAs are o nly applicable to Braidwood Unit 1.

B/B-UFSAR 4.2-10 REVISION 6 - DECEMBER 1996

4.2.2 Description

and Design Drawings The fuel assembly, fuel rod, and core compon ent design data are given in Tables 4.1-1 and 4.3-1.

NRC approval of the VANTAGE 5 design is given in R eference 19 and in Reference 24 for the VANTAGE+ design.

Each fuel assembly consists of 264 fuel rods, 24 guide thimble tubes, and 1 instrumenta tion thimble tube arranged within a supporting structure.

The instrumentation t himble is located in the center position and provides a channel for insertion of an incore neutron detecto r, if the fuel ass embly is located in an instrumented core po sition. The guide t himbles provide channels for insertion of either a rod cluster control assembly, a neutron source assembly, a bur nable absorber assembl y, or a thimble plug assembly, depending on the position of t he particular fuel assembly in the core.

Figure 4.2-1 shows a crosssection of the fuel assembly array; F igures 4.2-2a and 4.2-2b show fuel assembly full length outlines.

The fuel rods are loaded into the fuel assembly structure so that the re is clearance between the fuel rod ends and the top and bottom nozzles. The fu el rod also has an oxide coating at the bottom of the rod to provide additional rod fretting wear protection.

Figures 4.2-2a and 4.2-2b show the VAN TAGE 5 and VANTAGE+

assembly designs with their resp ective overall h eight and grid elevation dimensions.

The design changes between the VANTAGE 5 and VANTAGE+ designs i nclude a slightly shorter fuel rod for the VANTAGE 5 fuel rod design to accommodate extended burnup growth.

The VANTAGE 5 and VANT AGE+ designs also incorporate three Zircaloy or ZIRLO interm ediate flow mixing (IFM) grids. The DFBN is similar to the OFA design used in first core Cycle 1 fuel except it is lower in height and has a n ew pattern of smaller flow holes in its thinner top plate. This design minimizes passage of debris pa rticles which could cause fretting damage to fuel rod cladding.

Additional debris protec tion is provided by the protective grid asse mbly and an elongated fuel rod bottom end plug, which is described in Subsection 4.2.2.2.4.

The VANTAGE+ assembly skeleton is identi cal to that previously described for VANTAGE 5, except for those mo difications necessary to accommodate t he intended fuel operation to higher burnups.

The modifications consist of the use of ZIRLO gu ide thimbles and small skeleton d imensional alteratio ns to provide additional fuel assembly and rod growth space at the extended bu rnup levels. The VANTAGE+ fuel assembly is shorter than the VANTAGE 5 fuel assembly. The grid ce nterline elevations of the VANTAGE+ are identical to those of the VANTAGE 5 fuel assembly, except for the top grid. The VANTAGE+

top grid has been lo wered. Because the VANTAGE+ fuel is intended to rep lace the VANTAGE 5 fuel, the VANTAGE+ exterior assembly env elope is equivalent in design dimensions, and the functional i nterface with the reactor internals is also equivalent to those of previ ous Westinghouse

B/B-UFSAR 4.2-10a REVISION 10 - DECEMBER 2004 fuel designs. Also, the VANTAGE+ fuel assem bly is designed to be mechanically and hydraul ically compatible with t he VANTAGE 5 fuel assembly. The same functional requirements and design criteria previously establish ed for the Westing house VANTAGE 5 fuel assembly remain valid for the VANTAGE+ fuel as sembly. The VANTAGE 5 and VA NTAGE+ fuel assembly des igns are provided in Figures 4.2-2a a nd 4.2-2b, respectively.

Each fuel assembly is installed vertically in the reactor vessel and stands upright on the lower core plate, which is fitted with alignment pins to locate and orient the assembly.

After all fuel assemblies are set in place, the upper support structure is installed. Alignment pins, built into the upper core plate, engage and locate th e upper ends of the fuel assemblies. The upper core plate then be ars downward against t he holddown springs on the top nozzle of each fuel assembly to hold the fuel assemblies in place.

Visual confirmation of t he orientation of th e fuel assemblies within the core is provided by an engraved identification number on a corner clamp on the top n ozzle and an ind exing hole in the opposite corner clamp.

4.2.2.1 Fuel Rods The VANTAGE 5 and VANTAG E+ fuel rods consist of uranium dioxide ceramic pellets contained in slightly cold wor ked Zircaloy-4 or ZIRLO tubing, which is p lugged and seal welded at the ends to encapsulate the fuel. S chematics of the fuel rods are shown in Figure 4.2-3a. The VANT AGE+ fuel rod repres ents a modification to the VANTAGE 5 fuel rod intended to su pport operation for fuel clad in place of the Zircaloy-4 clad. T he ZIRLO alloy is a zirconium alloy simi lar to Zircaloy-4, which has been specifically developed to enhance corros ion resistance. The VANTAGE+ fuel rods con tain, as in VANTAGE 5, enriched uranium dioxide fuel pellets, and an integral fu el burnable absorber (IFBA) coating on so me of the enriched fuel pellets.

The VANTAGE+ fuel rod has the sa me clad wall t hickness as the VANTAGE 5 design.

The VANTAGE 5 fuel rod length is shorter to provide room for the r equired fuel r od growth. To offset the reduction in the ple num length, the VANTAGE+

fuel rod has a variable pitch plenum sp ring. The variable pitch plenum spring provides the same support as the regular plenum spring, but with fewer spring turns, which translates to less spring volume. The bottom end plug has an i nternal grip feature to facilitate fuel rod loading on both designs (VANTAGE+

and VANTAGE 5) and provides appropriate lead-in for the removabl e top nozzle reconstitution feature. The VANTAGE+ f uel rod also has an oxide coating at the bottom end of the fu el rod. The extra l ayer of oxid e coating provides additional debr is-induced, rod-fretti ng wear protection.

B/B-UFSAR 4.2-11 REVISION 9 - DECEMBER 2002 The axial blankets are typically a nominal 6 inches or 8 inches of natural or slightly enriched fuel p ellets at each end of the fuel rod pellet stack.

Axial blankets reduce neutron leakage and improve fuel utilization.

The axial bla nkets utilize chamfered pellets which are physic ally different (length) than the enriched pellets to help prevent accidental mixing during manufacturing.

The axial blanket may contain an annulus providing additional plenum space to reduce the rod i nternal pressure.

The IFBA coated fuel p ellets are identical to the enriched uranium dioxide pellets except for the a ddition of a thin zirconium diboride (ZrB

2) coating on t he pellet cylindrical surface. This coating may be app lied with a linear boron-10 loading (mg/in) that is greater than the ori ginal IFBA design for added flexibility in the core de sign. Coated pel lets occupy the central portion of the fuel column (up to 132 inches). The number and pattern of IFBA rods within an assembly may vary depending on the specific ap plication. The ends of the IFBA enriched coated pellets, like the en riched uncoated pellets, are also dished to allow for greater axial e xpansion at the pellet centerline and void vo lume for fission g as release. An evaluation and test pr ogram for the IFBA design features is given in Section 2.5 of Re ference 19. New sta ndard IFBA patterns have been incorporated in the core designs.

As a result of reconst itution activities perfo rmed during unit outages, leaking fuel rods may be replaced with either filler rods fabricated from stainless s teel, Zircaloy-4, or ZIRLO in accordance with cycle-specific reloa d analyses.

Void volume and cleara nces are provided within the rods to accommodate fission gases releas ed from the fuel, differential thermal expansion between the cl adding and the f uel, and fuel density changes during i rradiation, thus, avoi ding overstressing of the cladding or s eal welds. Shifting of the fuel within the cladding during handling or shipping prior to core loading is prevented by a stainle ss steel helical spring which bears on top of the fuel. Du ring assembly, the pelle ts are stack ed in the cladding to the required fuel height, the spring is then inserted into the top end of the fuel tube and the end plugs pressed into the ends of the tube and welded.

All fuel rods are internally pressurized with helium during the welding p rocess in order to minimize compressive cladding st resses and pre vent cladding flattening due to cool ant operating pressures.

4.2.2.2 Fuel A ssembly Structure The fuel assembly structure cons ists of a bottom nozzle, top nozzle, guide and inst rument thimbles, and grids as shown in Figure 4.2-2a.

B/B-UFSAR 4.2-11a REVISION 7 - DECEMBER 1998 4.2.2.2.1 Bottom Nozzle The bottom nozzle serv es as the bottom structural element of the fuel assembly and directs the coolant flow dis tribution to the assembly. The squar e nozzle is fabricat ed from Type 304 equivalent stainless s teel and consists of a perforated plate and four angle legs with bearing plates as shown in Figure 4.2-2a.

The legs form a plenum for the inlet coolant flow to the fuel assembly. The plate also prevents accid ental downward ejection of the fuel rods from the fuel assembly.

The bottom nozzle is fastened to the fuel a ssembly guide thimbles by locked screws which penetrate thro ugh the nozzle and mate with a threaded plug in each guide th imble. The bottom nozzle may be removed, as necessary to support fuel reco nstitution, by the removal of the locking screws.

Upon completion of reconsti tution activities, a circular locking cap, lo cated around the thimb le screw head, will be crimped into mati ng lobes on the nozz le, thus securing the locking screws in place.

Coolant flows from t he plenum in the bottom no zzle upward through the penetrations in the plate to the channel s between the fuel rods. The pen etrations in the plate are posit ioned between the rows of the fuel rods.

The VANTAGE 5 and VANTAG E+ designs include u se of the DFBN to reduce the possi bility of fuel rod damage due to debris-induced fretting. The relativ ely large flow hol es in a conventional bottom nozzle are repl aced with a new pa ttern of smaller flow holes for the DFBN. T he holes are sized to mi nimize passage of debris particles large enough to cause damag e while providing sufficient flow area, comparable pressure drop, and continued structural integrity of the nozzle. Tes ts to measure pressure drop and demonstrate s tructural integrity ve rified that the low cobalt 304 stainless steel DFBN is total ly compatible with the VANTAGE 5 and VANTAGE+ designs.

Changes in design compared to the 17x17 OFA bottom nozzle design for the Cycle 1 fuel inv olve: 1) a modified flow hole size and pattern as described above, 2) a decreased nozzle height and thinner top plate to a ccommodate the extende d burnup fuel rod, and 3) increased lead-in chamfers fo r the core pin interface to improve handling. T he DFBN also has a r econstitution design feature which facilitates easy removal of th e nozzle from the fuel assembly in the same mann er as all previous Byron/Braidwood Stations fuel designs.

Axial loads (holddown) imposed on the fu el assembly and the weight of the fuel a ssembly are transmit ted through the bottom nozzle to the lower core plate. Indexing and positioning of the fuel assembly are cont rolled by alignment ho les in two diagonally opposite bearing plates whic h mate with locating

B/B-UFSAR 4.2-12 REVISION 9 - DECEMBER 2002 pins in the lower core plate. Lateral loads on the fuel assembly are transmitted to the lower core plate through the locating pins. 4.2.2.2.2 Top Nozzle The top nozzle assembly func tions as the upper structural element of the fuel assembly and provides a part ial protective housing for the rod cluster control assembly or other core components.

It consists of an adapter plate, enclosure, top plat e, and pads.

Holddown springs are mounted on the top nozzle, as shown in Figures 4.2-2 and 4.2-2a.

The springs are made of Inconel-718, the screws are made of Inconel-718 or Incone l-600, and the top nozzle is made of Type 304 stainless steel.

The VANTAGE+ fuel assembly uses the same top nozzle design as the VANTAGE 5. The design bases and evalu ation of the reconstit uted top nozzle are given in Subsection 2.3.2 of Reference 19.

The reconstitutable top nozzle for the VANTA GE 5 and VANTAGE+

fuel assembly differs from the conventional OFA design in two ways: a groove is p rovided in each thimble throughhole in the nozzle plate to facilita te attachment and remo val; and the nozzle plate thickness is reduced to provide ad ditional axial space for fuel rod growth.

In the VANTAGE 5 and VANTAGE+ reconstitutabl e top nozzle design, a stainless steel nozz le insert is mechanically connected to the top nozzle adapter plate by means of a preform ed circumferential bulge near the top of the insert. The insert engages a mating groove in the wall of the adapter plate thimble tube throughhole.

The insert has four eq ually spaced axial slo ts which allow the insert to deflect inwardly at the elevation of t he bulge, thus permitting the installat ion or removal of the no zzle. The insert bulge is positively held in the adapter plate ma ting groove by placing a lock tube with a uniform ID identi cal to that of the thimble tube into the insert.

To remove the top nozzle, a to ol is first inserted through the lock tube and expanded r adially to engage the bottom edge of the tube. An axial force is then exerted on the t ool which overrides the local lock tube defo rmations and withdraws t he lock tube from the insert. After the lock tubes have been wi thdrawn, the nozzle is removed by raising it off the upper slotted ends of the nozzle inserts which deflect in wardly under the axial lift load. With the top nozzle removed, direct access is provi ded for fuel rod examination or replaceme nt. Reconstitution is completed by the remounting of the nozzle and the insertion of new lock tubes.

The design bases and e valuation of the reconstitutable top nozzle are given in Section 2.3.2 in Reference 19.

The square adapter plate in both the conventional and VANTAGE 5 RTN designs is provided with round penetrati ons and semicircular ended slots to permit the flow of coolant upward through the top nozzle. The l igaments in

B/B-UFSAR 4.2-12a REVISION 6 - DECEMBER 1996 the plate cover the tops of the fuel rods and prevent their upward ejection from the fuel assembly.

The enclosure is a box-like structure which sets the distance betwe en the adapter plate and the top plate. The top plate has a la rge square hole in the center to per mit access for the control rods and the control rod spiders and static core component assemblies.

Holddown springs are mounted on the top plate and are fastened in place by bolts and clamps located at two diagonally opposite corners. On the other two corne rs, integral pads are positioned which contain alignment holes for locating t he upper end of the fuel assembly.

4.2.2.2.3 Guide and In strument Thimbles The guide thimbles are structural members which also provide channels for the neutron absorber rods, burnable abs orber rods, neutron source, or thimb le plug assemblies.

Each thimble is fabricated from Zircaloy

-4 or ZIRLO tubing h aving two different diameters. The tube diameter at the top section provides the annular area necessary to permit rapid contr ol rod insertion during a reactor trip. The lower po rtion of the guide thimble is swaged to a smaller di ameter to redu ce diametral clearances and produce a dashpot action near the end of the control rod travel during normal trip opera tion. Holes are provi ded on the thimble tube above the d ashpot to reduce the rod drop time. The dashpot is closed at the bot tom by means of an end plug which is provided with a small flow po rt to avoid fluid st agnation in the dashpot volume during normal ope ration. The top end of the guide thimble is fastened to a tubular nozzle insert by three expansion swages.

The insert engages i nto the top nozzle a nd is secured into position by a lock tube, as shown in Figure 4.2-6.

The lower end of the guide t himble is fitted with an end plu g, which is then fastened into the bottom noz zle by a locked screw.

B/B-UFSAR 4.2-13 REVISION 6 - DECEMBER 1996 Each grid is fastened to the g uide thimble ass emblies to create an integrated structure. Th e fastening meth od depicted in Figures 4.2-4 and 4.2-5 is used for all but the top and bottom grids in a fuel assembly.

An expanding tool is inserted into the inner diameter of the Zircaloy or ZIRLO thimble tube to the elevation of the Zircaloy sleeves that have be en welded into the n ine Zircaloy grid assemblies (six stru ctural and three flow mi xer). The four-lobed tool forces the thimble and sleeve outward to a predetermined diameter, thus joining the two components.

VANTAGE 5 and VANTAG E+ fuel assembly t op grid-to-thimble attachments are identical and are shown on F igure 4.2-6. The Zircaloy or ZIRLO thimbl es are fastened to the top nozzle inserts by expanding the members as shown in Figure 4.2-

6. The inserts then engage the top nozz le and are secured into position by the insertion of lock tubes.

The bottom grid assemb ly is joined to the assembly as shown in Figure 4.2-7. The stainless steel insert is spotwel ded to the bottom grid and later ca ptured between t he guide thimble end plug and the bottom nozzle by means of a stai nless steel thimble screw.

The described methods of grid and nozzle insert fastening have been mechanically tested and fou nd to meet all applicable design criteria.

The VANTAGE 5/VANTAGE+ g uide thimbles are id entical. Both the VANTAGE 5 and VANTAG E+ guide thimble tube ID provide an adequate nominal diametral clea rance for the control ro ds. The thimble tube ID also pro vides sufficient diametral c learance for burnable absorber rods, sourc e rods, and thim ble plugs.

The VANTAGE 5 and VANTAG E+ instrumentation tub es are identical in design and both allo w sufficient diametral cle arance for the flux thimble to traverse the tube without binding.

The central instrumentation tu be of each fuel assembly is constrained by seating in counterbores in each nozzle. This tube is a constant diameter a nd guides the incore neutron detector thimbles. This tube is expanded at the top and mid-grids in the same manner as the p reviously discussed expa nsion of the guide thimbles to the grids.

4.2.2.2.4 Grid Assemblies The fuel rods, as shown in Figure 4.2-2a, are supported at intervals along their le ngth by structural g rid assemblies which maintain the lateral spacing bet ween the rods. Each fuel rod is supported within each grid cell by a combination of support dimples and springs.

The magnitude of grid spring force on the fuel rods is set high enough to minimize possible fretting, B/B-UFSAR 4.2-13a REVISION 7 - DECEMBER 1998 without overstressing the cladding at the contact points. All grid assemblies allow axial thermal expansion of the fuel rods without imposing restrai nt sufficient to develop buckling or distortion.

The top and bottom grids are made of Inc onel-718 strap material, chosen for its s trength and high corrosi on resistance. These nonmixing vane grids a re identical in the VA NTAGE 5 and VANTAGE+

designs.

The six intermediate (mi xing vane) grids are made of Zircaloy straps or ZIRLO, chosen for low neutron absorp tion properties and corrosion resistance, and are identical in the VANTAGE 5 and VANTAGE+ designs. Inner straps include mixing vanes which project into the coolant stream and promote mixing of the coolant. In addition to the a nti-snag feature, the intermediate grids incorporate the sa me grid cell support configuration as the top and bottom Inconel g rids (six support locations per cell:

four dimples and two springs

). The Zircaloy and ZIRLO grid interlocking strap joints and gr id/sleeve joints are fabricated by laser welding, whereas the Inconel grid joi nts are brazed.

The intermediate flow mixer (IFM) grids are located in the three uppermost spans between the Zircaloy or ZIRLO mixing vane structural grids and incorporate a similar mixing vane array.

Their prime function is mid-span flow mixing in the hottest fuel assembly spans. Each IFM grid cell cont ains four dimples which are designed to prev ent mid-span channel closure in the spans containing IFMs and fuel rod contact with the mixing vanes. This simplified cell arrang ement allows short gri d cells so that the IFM grid can accomplish its flow mixing objective with minimal pressure drop.

The IFM grids are fabric ated from Zircaloy or ZIRLO and assembled in the same manner as the six intermediate (mixing vane) grids.

These grids are not inte nded to be structura l members. The outer strap was designed similar to the other grids to preclude grid hang-up and damage dur ing fuel handling.

Impact tests have been performed to show that a coolable geometry is assured at the IFM and structural grid el evation during ass ured at the IFM and structural grid elevat ion during seismic

/LOCA events. The VANTAGE 5 grid a ssembly design bases and elevations are given in Section 2.3.5 of Reference 19, and the V ANTAGE+ in Sec tion 2.3 of Reference 24.

Reload fuel assemblies i ncorporate a bottom protective grid and modifications to the top and b ottom fuel rod end plug. The protective grid illustra ted in Figure 4.

2-2b is a partial height grid, similar in configu ration to the interm ediate flow mixing grid, fabricated of Inco nel without mixing van es, and positioned on the top plate of the bottom nozzle. In conjunction with the protective grid, both the bottom and the top fuel rod end plugs were elongated. The protective grid and elo ngated bottom end plug provide a zone below the active fuel in which debris can be trapped.

B/B-UFSAR 4.2-14 REVISION 2 - DECEMBER 1990 4.2.2.3 Core Components 4.2.2.3.1 Rod Cluster Control Assembly The rod cluster control assembli es are used for shutdown and control purposes to offs et fast reactivity changes. Figure 4.2-8 illustrates the rod cluster control asse mbly location in the reactor relative to the in terfacing fuel assemblies and guide tube assemblies.

A rod cluster contro l assembly is comp rised of a group of individual neutron abs orber rods fastened at the top end to a common spider as sembly, as illustrat ed in Figure 4.2-9.

The absorber materials used in the contr ol rod design are either: (1) Ag-In-C d alloy extruded rods or (2) a solid hafnium bar. The abso rber materials are ess entially "black" to thermal neutrons and have suff icient additio nal resonance absorption to significan tly increase their wor th. For both the Ag-In-Cd alloy and the hafnium design, the material is sealed in cold-worked type 304 stainless steel tubes to prevent the absorber material from coming in direct contact with the coolant (Figure 4.2-10).

Sufficient diametral and end clearances are provided to accomm odate relative the rmal expansions and material swelling, as shown in Subsection 4.2.3.6.

Enhanced Performance Rod Cluster Control Assembl ies (EP-RCCA's) which use silver-indiu m-cadmium (Ag-In-Cd) w ill be utilized.

These have a thin chrome elect roplate applied ov er the length of absorber rodlet cladding in contact with the reactor internal guides to pro vide increased res istance to cladding wear. In addition, the absorber diameter is reduced slightly at the lower extremi ty of the rodlets in order to accommodate absorber swelling and minimize cladding intera ction. The absorber rod cladding material is a very hig h purity 10% cold worked type 304 stainl ess steel tubing.

Because of its r elatively high yield strength (minimum of 62,000 psi at 600

°F), use of this mate rial results in a design with practical wall th ickness that meets ASM E Section III type stress criteria for stre sses induced by oper ating conditions.

The high purity stainless steel has a signif icant reduction in cobalt content as compar ed to earlier design.

The chrome plate further reduces the ef fluence of cobalt into the coolant, thereby benefiting the ALARA conditions.

This high purity cladding is also very resistant to irradiation assisted stress corrosion cracking.

As the bottom 12 inches of the silver absorb er material sees a much higher fluence than the rest of the absorber material, an additional 5 mil diametral gap h as been introduc ed in that area to allow for more irradiatio n induced swelling without imposing

B/B-UFSAR 4.2-14a REVISION 2 - DECEMBER 1990 significant hoop stresses in t he cladding. This modification was made based on extensive exam ination of irr adiated material which indicated that clad cracking could occ ur as a result of absorber/clad interfer ence. The gap size is small enough so that heat transfer from the absorber to the coolant is sufficient to maintain a substantial margin against absorber rod melting. The bottom end plugs are bullet-nosed to reduce the hydraulic drag during reactor trip and to guide the absorber rods smoothly into the dashpot section of the fuel assembly guide thimbles.

The spider assembly is in the form of a central hub with radial vanes supporting fingers from which the absorber rods are suspended. Handling d etents and detents for connection to the drive rod assembly are machined into the upper end of the hub.

A coil spring in side the spider body a bsorbs the impact energy at the end of a trip i nsertion. The radial vanes are joined to the hub by welding a nd brazing, and the fing ers are joined to the vanes by brazing. A centerp ost, which hol ds the spring and B/B-UFSAR 4.2-15 REVISION 6 - DECEMBER 1996 its retainer, is threaded into t he hub within the skirt and welded to prevent loos ening in service. All components of the spider assembly are made from Types 304 and 308 stainless steel, except for the r etainer which is of 17-4 pH material and the springs which are Inconel-718 alloy.

The absorber rods ar e fastened secur ely to the spider. The rods are first threaded into the spid er fingers and then pinned to maintain joint tightness, after which the pi ns are welded in place. The end plug b elow the pin posit ion is designed with a reduced section to permit flexing of the rods to correct for small misalignments.

The overall length is su ch that when the ass embly is withdrawn through its full travel, the tips of the absorber rods remain engaged in the g uide thimbles so that al ignment between rods and thimbles is always mai ntained. Since the rods are long and slender, they are relatively free to conform to any small misalignments with t he guide thimble.

4.2.2.3.2 Burnable Absorber Assembly Each burnable absorber assembly consists of borosilicate or WABA burnable absorber ro ds attached to a h olddown assembly.

Conventional burnabl e absorber assembl ies (containing borosilicate absorber) a re shown in Figure 4.2

-11. WABA rods may be used in place of the borosilicate a bsorber rods.

The borosilicate absorber rods c onsist of boro silicate glass tubes contained within T ype 304 stainless st eel tubular cladding which is plugged and seal welded at the ends to encapsulate the glass. The glass is also supp orted along the length of its inside diameter by a thin wall tubular i nner liner. The top end of the liner is open to permit the diffused helium to pass into the void volume and the liner overhangs the glas

s. The liner has an outward flange at the bottom end to maintain the position of the liner with the glass. A typical borosilicate burnable absorber rod is shown in longitu dinal and transverse cross-sections in Figure 4.2-13.

A WABA rod (Figure 4.2-12) consists of annular pellets of alumina-boron carbide (A1 2 0 3-B 4 C) burnable absorber material contained within two concentric Zircaloy tubes.

These Zircaloy tubes which form the inn er and outer clad fo r the WABA rod, are plugged and welded at ea ch end to encapsulate the annular stack of absorber material.

The assembled r od is then internally pressurized to 650 psig and seal welded.

The absorber stack lengths are positioned a xially within the WABA rods by the use of Zircaloy bottom-end spacers. The spacer in the lower portion of the WABA rod was lengthened to a ccount for the ZIRLO guide thimbles. The burna ble absorber centerline is aligned with the fuel centerline at hot full power condit ions at the BOL. An annular plenum is provid ed within the rod to accommodate the helium gas released fr om absorber material depletion during irradiation. The reactor coolant flows inside the inner tube and outside the outer tube of the annular ro

d. Further design details are given in Sec tion 3.0 of Reference 15.

B/B-UFSAR 4.2-15a REVISION 2 - DECEMBER 1990 The burnable absorbe r rods are statica lly suspended and positioned in select ed guide thimbles within the fuel assemblies. The absorber rods in each assembly are attached together at the top end of the rods to a hol d down assembly by a flat, perforat ed retaining plate whi ch fits within the fuel assembly top nozzle and rests on the ada pter plate. The absorber rod assembly is held do wn and restrained against vertical motion through a spring pack which is attached to the plate and is compres sed by the upper c ore plate when the reactor upper intern als assembly is lowered into the reactor.

This arrangement ensures that the absorber r ods cannot be ejected from the core by flow fo rces. Each rod is permanently attached to the base plate by a nut which is c rimped in place.

The borosilicated rod cladding is slightly cold worked Type 304 stainless steel, and the WABA rod cladding is Zircaloy-4. All other structural materia ls are Types 304 or 308 stainless steel except for the s prings which are Inconel-718. T he borosilicate glass tube provides su fficient boron con tent to meet the criteria discussed in Subsection 4.3.1.

B/B-UFSAR 4.2-16 REVISION 6 - DECEMBER 1996 4.2.2.3.3 Neutron Source Assembly The purpose of the neutr on source assembly is to provide a base neutron level to ensure that the neutron detectors are operational and respondi ng to core multiplication neutrons.

Both primary and secondary neutr on source rods a re used. The primary source rod, containing a radio active material (californium-252), s pontaneously emits n eutrons during initial core loading and reactor startup.

After the pri mary source rod decays beyond the desired neutron flux level, ne utrons are then supplied by the secondary source rod. The secon dary source rod contains a stable material (Sb-Be), which is activated by neutron bombardment duri ng reactor operation. This becomes a source of neutrons dur ing periods of l ow neutron flux, such as during refueling and s ubsequent startups.

Four source assemblies are installed in reactor core for the initial fuel cycle:

two primary source assemblies and two secondary source assemblies. Ea ch primary source assembly contains one primary source rod and a number of burnable absorber rods. Each secondary source as sembly contains a grouping of four or six secondary source rods.

In both types of assemblies locations not filled with source or burnable absorber rods contain a thimble plug. The s ource assemblies are shown in Figures 4.2-14, 4.2-15, and 4.2-15a. After the initial fuel cycle, only the secondary sources are used.

The source assemblies contain a holddown ass embly identical to that of the burnable absorber assembly.

The primary and secondary source rods have the same cladding m aterial as the absorber rods. The seco ndary source rods contain pellets stacked to a height of approximately 88 inch es. The primary source rods contain caps ules of californium so urce material and alumina spacer pellets to position the s ource materi al within the cladding. The r ods in each assemb ly are permanently fastened at the top end to a holddown assembly.

B/B-UFSAR 4.2-17 REVISION 2 - DECEMBER 1990 The structural members a re constructed of Ty pe 304 stainless steel except for the springs.

The springs exposed to the reactor coolant are Inconel-718.

4.2.2.3.4 Thimble Plug Assembly Thimble plug assemblies may be used if desir ed to limit bypass flow through the guide thimbles in fuel asse mblies which do not contain either control r ods, source rods, or burnable absorber rods.

The thimble plug assembly, as shown in Figure 4.2-16, consists of a flat base plate with sh ort rods suspend ed from the bottom surface and a spring p ack assembly. The 24 short rods, called thimble plugs, project into the upper ends of the guide thimbles to reduce the bypass flow. Each thimble plug is permanently attached to the base plate by a nut which is crimped into the threaded end of the plug. Similar short rods are also used on the source asse mblies and bur nable absorber assemblies to plug t he ends of all vacant fuel assem bly guide thimbles. When in the core, the thi mble plug assemblies interface with both the upper core plate and with the fuel assembly top nozzles by resting on the adapter plate.

The spring pack is compressed by the upper core p late when the upper internals assembly is lowered into place.

All components in the thimble plug assem bly, except for the springs, are constructed from Ty pe 304 stainless steel. The springs are Inconel-718.

4.2.3 Design

Evaluation The fuel assemblies, fuel rods, and incore contr ol components are designed to satisfy the performance and safety criteria of Section 4.2, the mecha nical design bases of Subsection 4.2.1, and other interfacin g nuclear and ther mal-hydraulic design bases specified in Sections 4.3 and 4.4.

Effects of accident conditions II, III, IV, or anticipated trans ients without trip on fuel integrity are presented in Chapter 15.0 or supporting topical reports.

4.2.3.1 Cladding

a. Vibration and Wear Fuel rod vibrations are flow induced.

The effect of the vibration on the fuel assembly and individual fuel rods is minimal.

The cyclic stress range associated with deflections of such small magnitude is insignificant and has no effect on the structural integrity of the fuel rod.

No significant wear of the cladding or grid supports is expected during the life of the fuel assembly.

Fuel vibration has been experimentally inves tigated as shown in Reference 7.

B/B-UFSAR 4.2-18 REVISION 3 - DECEMBER 1991 b. Fuel Rod Internal Pres sure and Cladd ing Stresses The burnup dependent f ission gas release model (Reference 4) is used in det ermining the internal gas pressures as a function of irradiation time.

The fuel rod has been design ed to ensure that the maximum internal pressure of the fuel rod will not exceed the value whi ch would cause an increase in the fuel cladding diamet ral gap and extensive DNB propagation during n ormal operation.

The cladding stresses at a constant local fuel rod power are low. Compressive stresses are created by the pressure differential between the coolant pressure and the rod interna l gas pressure. Because of the prepressurization with helium, the volume average effective stresses are always less than approximately 10,000 psi at the pressurization level used in this fuel rod de sign. Stresses due to the temperature gradient are not included in this average effective stress because thermal stresses are, in general, negative at the cladding inside diameter and positive at the cladding outside diameter and their contr ibution to t he cladding volume average stres s is small. Furthermore, the thermal stress decreases with time during steady-state operati on due to stress relaxation.

The stress due to pressure differential is highest in the minimum p ower rod at the beginning-of-life due to low internal gas pres sure, and the thermal stress is highest in the maximum power rod due to the steep temp erature gradient.

The internal gas pressure at beginning-of-life is approximately 850 psia at operating temperature for a typical lead burnup fu el rod. The total tangential stress at the cla dding inside diameter at beginning-of-life is app roximately 15,200 psi compressive (~13,200 psi due to P and ~1,400 due to T) for a low power rod, operating at 5 kW/ft, and approximately 14,600 psi com pressive (~11,000 psi due to P and 3,600 psi due to T) for a high power rod operating at 15 kW/ft.

However, the volume average effective stress at beginning-of-life is between approximatel y 7,500 psi (high power rod) and approximately 10,000 psi (lo w power rod). These stresses are substantially b elow even the unirradiated cladding strength (~55,500 p si) at a typical cladding mean operating t emperature of 700

°F. Tensile stresses could be cr eated once the cladding has come in contact with the pellet. These

B/B-UFSAR 4.2-19 REVISION 6 - DECEMBER 1996 stresses would be induced by the fuel pellet swelling during irradiation.

Fuel swelling can result in small cladding str ains (<1% for expected discharge burnups), but the associated cladding stresses are very low be cause of cladding creep (thermal and irradiation-induced creep). The 1%

strain criterion is extr emely conservative for fuel-swelling driven cladding strain because the strain rate associated with so lid fission products swelling is very slow.

c. Materials and Chemical Evaluation Zircaloy-4 cladding and ZIRLO cladding have high corrosion resistance to the coolant, fuel, and fission products. As shown in Reference 1, there is PWR operating experience on the capability of Zircaloy and ZIRLO as a clad ding material. Controls on fuel fabricat ion specify maximum moisture levels to preclude clad ding hydriding.

Metallographic exami nations of irradia ted commercial fuel rods have shown occurrences of fuel-clad chemical interaction. React ion layers of <1 mil in thickness have been observed between fuel and clad at limited points arou nd the circumference.

Metallographic data indicate that this interface layer remains very t hin even at high burnup. Thus, there is no indication of pr opagation of the layer and eventual cla dding penetration.

d. Fretting Cladding fretting has been experimentally investigated as shown in Reference 7. No significant fretting of the cladding is expected during the life of the fuel assembly.

e. Stress Corrosion Stress corrosion cracking is another postulated phenomenon related to fuel/clad chemical interaction. Out-of-pile te sts have sho wn that in the presence of high c ladding tensil e stresses, large concentrations of selected fission products (such as iodine) can chemica lly attack t he Zircaloy and ZIRLO tubing and can lead to eventual cladding cracking. Extensive pos tirradiation examination has produced no in-pile evidence that this mechanism is operative in c ommercial fuel.

B/B-UFSAR 4.2-20 REVISION 9 - DECEMBER 2002 f. Cycling and Fatigue A comprehensive review of the availa ble strain fatigue models was conducted by Westinghouse as early as 1968. This r eview included the Langer-O'Donnell mod el (Reference 8,) the Yao-Munse model and the Manson-Halford model. Upon completion of this review a nd using the results of the Westinghouse experimental pr ograms discussed below, it was concluded that the approach defined by Langer-O'Donnell would be retained and the empirical factors of the ir correlation modified in order to conservativ ely bound the results of the Westinghouse tes ting program.

The Westinghouse testing pro gram was subdivided into the following subprograms:

1. A rotating bend fatigue experiment on unirradiated Zircaloy-4 specimens at room temperature and at 725

°F. Both hydrided and nonhydrided Zircaloy-4 c ladding were tested.

2. A biaxial fatigue ex periment in gas autoclave on unirradiated Zircaloy

-4 cladding, both hydrided and nonhydrided.

3. A fatigue te st program on irra diated cladding from the CVS and Yankee Core V conducted at Battelle Memorial Institute.

The results of these test programs provided information on differe nt cladding conditions including the effects of irradiation, of hydrogen leve ls and of temperature.

The design equations followed the concept for the fatigue design crite rion according to the ASME Boiler and Pressure Vess el Code,Section III.

It is recognized that a possible limitation to the satisfactory behavior of the fuel rods in a reactor which is subjected to da ily load follow is the failure of the c ladding by low cycle strain fatigue. During their n ormal residence time in reactor, the fuel rods may be subjected to approximately 1000 cycles with typical changes in power level from 50% to 100%

of their steady-state values. The assessment of th e fatigue life of the fuel rod cladding is subject to a con siderable uncertainty due to the difficulty of evaluating the strain range which results from the cyclic interaction of the fuel pellets and cla dding. This difficulty

B/B-UFSAR 4.2-21 REVISION 6 - DECEMBER 1996 arises, for example, from such highly unpredictable phenomena as pellet crac king, fragmentation, and relocation. Nevertheless, since early 1968, this particular phenomenon has been investigated analytically and experim entally. Strain fatigue tests on irradiated and noni rradiated hydrided Zr-4 claddings were performed, which permitted a definition of a conservative fatigue life limit and recommendation on a methodol ogy to treat the strain fatigue evaluation of the Westinghouse reference fuel rod designs.

It is believed t hat the final proof of the adequacy of a given fuel rod design to meet the load follow requirements can only come from incore experiments performed on actual reactors. Experience in load follow operation dat es back to early 1970 with the load follow operation of the Saxton reactor.

Successful load foll ow operation has been performed on reactor A (>400 load follow cycles) a nd reactor B

(>500 load follow cycles).

In both cases, there was no significant c oolant activity incr ease that could be associated with the load follow m ode of operation.

g. Rod Bowing For Zircaloy-4 grid fuel ass emblies, the largest contributors to signific ant rod bow are high end grid forces (Inconel g rids) and low stiffness Zircaloy-4 grid springs.

The VANTAGE 5 and VANTAGE+

fuel assembly designs have low spring fo rces on the top Inconel grids. This reduc es the end loadings on the fuel rod brought about by fuel rod growth. The Zircaloy-4 or ZIRLO mid-grid design has a very high spring stiffness. T his design o ffers high resistance to fuel rod r otation within a grid and still has low spring force to allow the rods to slip freely thru the grids. This design reduces the rod bow of the VANTAGE 5 and VAN TAGE+ (or any Zircaloy grid design) to values as good or better than all Inconel gridded assemblies.

The current conservative NRC approved methodology for comparing the magnitude of rod bow between two different fuel assembly designs is given in Reference 14. Based on this approved methodology, a comparison of L 2/I (where I = the fuel rod bending moment of inertia and L

= span lengt h) and the initial rod-to-rod g ap for both the 17 x17 VANTAGE 5 and VANTAGE+ fuel assembly designs, shows that the amount of rod bow at any given burnup is essentially the same for both 17x17 VANTAGE 5 and VANTAGE+ fuel assemblies.

B/B-UFSAR 4.2-22 REVISION 6 - DECEMBER 1996 The effects of rod bow on DNBR are des cribed in Subsection 4.4.2.2.5.

Thus, for a given burnup, the rod bow effects to be applied to the VANTA GE+ fuel assembl ies are the same as those applied to the VANTAGE 5 17 x17 fuel.

h. Consequences of Power-Coolant Mismatch This subject is discussed in Chapter 15.0.

i. Irradiation Stabilit y of the Cladding As shown in Reference 1, the re is considerable PWR operating experience on the capability of Zircaloy as a cladding ma terial and for Z IRLO to date.

Extensive experience with irradiated Z ircaloy-4 is summarized in Reference 2 and in Appendixes A through E in Ref erence 24 for ZIRLO.

j. Creep Collapse and Creepdown This subject and the associated irradiation stability of cladding have been evaluated using the model described in Reference 26. It has been established that clad collap se has been eliminated from the design basis.

4.2.3.2 Fuel Mat erial Consideration

a. Dimensional Stab ility of the Fuel The mechanical desig n of the fuel rods accounts for the differential thermal expan sion of the fuel and the cladding, and for the pellets densification effect. b. Potential for Chemical Interaction Sintered, high density u ranium dioxide fuel reacts only slightly with the cladding at core operating temperatures and pressur es. In the event of cladding defects, the hi gh resistance of uranium dioxide to attack by w ater protects against fuel deterioration, altho ugh limited fuel erosion can occur. The effects of w ater-logging on fuel behavior are discussed in Subsection 4.2.3.3.

c. Thermal Stability

As has been shown by o perating experience and extensive experimental w ork, the thermal design parameters conservatively ac count for changes in the thermal performance of the fuel elements due to

B/B-UFSAR 4.2-23 REVISION 1 - DECEMBER 1989 pellet fracture which may occur duri ng power operation. Observations from several operating Westinghouse PWRs (Reference

6) have shown that fuel pellets can densify under irradiation to a density higher than the manufactured values. Fuel densification and subsequent settling of the fuel pellets can result in lo cal and distributed gaps in the fuel rods. Fuel densifi cation has b een minimized by improvements in the fuel manufactur ing process and by specifying a nominal 95% initial fuel density.

The evaluation of fuel densi fication effects and their consideration in fuel design a re described in References 3 and 4.

d. Irradiation Stability

The treatment of fuel swelli ng and fission gas release is described in Reference 4.

4.2.3.3 Fuel Rod Performance The initial step in fuel rod des ign evaluation f or a region of fuel is to determine the limiting rod(s). Limiting rods are defined as those rods whose predicted performa nce provides the minimum margin to each of the design criteri

a. For a number of design criteria, the l imiting rod is the lead burnup rod of a fuel region. In oth er instances, it m ay be the maximum power or the minimum b urnup rod. For the most part, no single rod will be limiting with respec t to all design criteria.

After identifying th e limiting rod(s), a worst-case evaluation is made which ut ilizes the limiting rod design basis power history and considers the effects of mod el uncertainties and dimensional variations.

Furthermore, to ver ify adherence to the design criteria, the conservative ca se evaluation also considers the effects of postu lated transient power increases, which are achievable during operation consis tent with Conditions I and II. These transient pow er increases can affect both rod average and local power levels. The analyti cal methods used in the evaluation result in performance parameters which demonstrate the fuel rod behavior.

Examples of parameters considered include rod internal pressure, fuel temperature, cladd ing stress, and cladding strain. In fuel rod de sign analyses, t hese performance parameters provide the basis for comparison be tween expected fuel rod behavior and the correspondi ng design criteria limits.

In calculating the s teady-state performa nce of a nuclear fuel rod, the following interacti ng factors are considered:

B/B-UFSAR 4.2-24 REVISION 2 - DECEMBER 1990 a. Cladding creep and elastic deflection;

b. Pellet density changes, thermal expansion, gas release, and thermal pro perties as a function of temperature and fuel burnup; and

c. Internal pressure as a function of f ission gas release, rod geometry, and temperature distribution.

These effects are evaluated using a fuel rod design model (Reference 4). The model modifi cations for time dep endent fuel densification are given in Reference 4.

With these interacting factors considered, the model determines the fuel rod performance characteri stics for a given rod geometry, power history, and axial power shape.

In particular, internal gas pressure, fuel and cla dding temperature, and cladding deflections are calculated. The fuel rod is divided into several axial sections and radially into a n umber of annular zones. Fuel density changes are calc ulated separately for each segment. The effects are integrated to obtain the internal rod pressure.

The initial rod internal press ure is selected to delay fuel-clad mechanical interaction and to avoid the potential for flattened rod formation. It is limited, however, by the design criteria for the rod internal pressure given in Subsections 4.2.1.3 and 4.2.3.1.b.

The gap conductance between the pell et surface and the cladding inner diameter is calcul ated as a function of the composition, temperature, and pressure of the gas mixture, and the gap size or contact pressure be tween cladding and pelle

t. After computing the fuel temperature for each pellet annular zone, the fractional fission g as release is assess ed using an empirical model derived from exp erimental data (Refere nce 4). The total amount of gas released is based on the average fractional release within each axial and radial zone an d the gas generation rate, which in turn is a function of bur nup. Finally, the gas released is summ ed over all zones and the pressure is calculated.

The code shows good agreement and fit for a variety of published and proprietary data on fission ga s release, fuel temperatures, and cladding defle ctions (Reference 4). Included in this spectrum are v ariations in power, time, fuel density, and geometry. In-pile fuel temperature meas urement comparisons are shown in Reference 4.

a. Fuel-Cladding Me chanical Interaction One factor in fuel eleme nt duty is potential mechanical interacti on of fuel and c ladding. This fuel-clad interaction pr oduces cyclic stresses and strains in the cladd ing, and these in turn consume

B/B-UFSAR 4.2-25 clad fatigue life. The reduction of fuel-clad interaction is, therefore, a goal of design. In order to achieve this goal and to enhance the cyclic operational c apability of the fuel rod, the technology for u sing prepressurized fuel rods in Westinghouse PWRs ha s been developed.

Initially, the gap betwe en the fuel and cladding is sufficient to prevent hard contact between the two. However, d uring power oper ation, a gradual compressive creep of the cladding onto the fuel pellet occurs due to the ext ernal pressure exerted on the rod by the coolan

t. Cladding compressive creep eventually results in the fuel-clad contact.

During this period of fu el-clad contact, changes in power level could result in changes in cladding stresses and strains. By using prepre ssurized fuel rods to partially offset the effect of the coolant external pressure, the rate of cladding creep toward the surface of the fuel is reduced. Fuel rod prepressurization de lays the time at which fuel-clad contact occurs and hence, significantly reduces the number and e xtent of cyclic stresses and strains experienced by t he cladding both before and after fuel-clad contact.

These factors result in an increase in the fatigu e life margin of the cladding and lead to gre ater cladding reliability.

If gaps should form in t he fuel stacks, cladding flattening will be p revented by the rod prepressurization so that the flattening time will be greater than the fuel core life.

A two dimensional (r, ) finite element model has been developed to investigate the effects of radial pellet cracks on stress concentrations in the cladding. Stress conc entration, herein, is defined as the difference betw een the maximum cladding stress in the -direction and the mean cladding stress. The fir st case has the fuel and cladding in mechanical equilibrium, and as a result the stress in the cladding is close to zero. In subsequent cases, the pellet power is increased in steps and the result ant fuel thermal expansion imposes tensile stress in the cladding.

In addition to uniform cladding stresses, stress concentrations develop in the c ladding adjacent to radial cracks in the pellet. Thes e radial cracks have a tendency to open during a power i ncrease but the frictional forces between fuel and clad ding oppose the opening of these cracks and result in localized increases in cladding stress. As the power is further increased and large tens ile stresses exceed the ultimate tensile strength of UO 2 , additional cracks

B/B-UFSAR 4.2-26 REVISION 6 - DECEMBER 1996 in the fuel are crea ted which limit the magnitude of the stress concentrati on in the cladding.

As part of the s tandard fuel rod design analysis, the maximum stress conce ntration evaluated from finite element calcu lations is added to the volume average effective stre ss in the cladding as determined from one di mensional stress/strain calculations. The resultant claddin g stress is then compared to the temperature dependent Zircaloy/ZIRLO yield stress in order to assure that the stress/strain criteria are satisfied.

Pellet thermal expan sion due to powe r increases is considered the only mech anism by which significant stresses and strains can be imposed on the cladding.

Power increases in commercial reacto rs can result from fuel shuffling, r eactor power escalation following extended reduc ed power operation, and control rod movement. In the mechanical design model, lead rods are dep leted using best estimate power histories as determined by core physics calculations. During the depletion, the amount of diametral gap closure is eva luated based upon the pellet expansion-cracking model, cladding creep model, and fuel swelling model. At various times during depletion, the po wer is increased locally on the rod to the burnup de pendent attainable power density, as determined by co re physics calcu lations.

The radial, tangential, and ax ial cladding stresses resulting from the power increase are combined into a volume average effec tive cladding stress.

The von Mises criter ion is used to evaluate whether the cladding yield stress has been exceeded. The yield stress correlation is that for irradiated cladding since fuel-clad int eraction occurs at high burnup. Furthermore, the effective stress is increased by an allowance, which accounts for stress concentrations in the cladding adjac ent to radial cracks in the pe llet, prior to the comparison with the yield stress. This allowance was evaluated using a two-dimensional (r,) finite ele ment model.

Slow transient power increases can r esult in large cladding strains without exc eeding the cladding yield stress because of cladding creep and stress relaxation. The refore, in addit ion to the yield stress criterion, a criterio n on allowable cladding strain is necessary. Ba sed upon high strain rate burst and tensile test data on irradiated tubing, 1%

strain was determined to be a conservative lower

B/B-UFSAR 4.2-27 REVISION 8 - DECEMBER 2000 limit for irradiated cla dding ductiblity and thus adopted as a d esign criterion.

b. Irradiation Experience

Westinghouse fuel op erational experience is presented in Reference 1.

Additional test assembly and test rod experiences are given in Sections 8 and 23 of Reference 6.

c. Fuel and Cladding Temperature The methods used for e valuation of fuel rod temperatures are presented in Subsection 4.4.2.11.

d. Water-logging Water-logging damage of a defective fuel rod has occasionally been postul ated as a me chanism for subsequent rupture of the cl adding. Such damage has been postulated as a con sequence of a power increase on a rod after water has entered such a rod through a cladding defect of a ppropriate size. Rupture is postulated upon powe r increase if th e rod internal pressure increase is exc essive due to insufficient venting of water to the reactor coolant. Local cladding deformations ty pical of water-logging bursts have never been o bserved in commercial Westinghouse fuel. Expe rience has shown that the small number of rods whi ch have acquired cladding defects, regardless of prima ry mechanism, remain intact and do not progre ssively distort or restrict coolant flow. In fact, such sma ll defects are normally observed through reductions in coolant activity to be progressively closed upon further operation due to the buildup of zirconium oxide and other substances. Secon dary failures which have been observed in defected rods are attributed to hydrogen embrittlement of the cladding.

Postirradiation examin ations point to the hydriding failure mechanism rath er than a waterlogging mechanism; the secondary failure occurs as axial cracks in the claddi ng and are similar regardless of the primary failure mechanis

m. Such cracks do not result in flow blockage.

Hence, the presence of such fuel, the q uantity of which must be maintained below Technical Spec ification limits, does not in any way exacerbate the e ffects of any postulated transients.

Zircaloy clad fuel rods whic h have failed due to water-logging (Reference 9) in dicate that very rapid power transients are require d for fuel failure.

Normal operation al transients are limited

B/B-UFSAR 4.2-28 to about 40 cal/gm-min. (pea k rod), while the Spert tests (Reference 10) ind icate that 120 cal/gm to 150 cal/gm are r equired to ruptu re the cladding even with very short tra nsients (5.5 msec period).

e. Potentially Damaging T emperature Effects During Transients The fuel rod experiences man y operational transients (intentional maneuve rs) during its r esidence in the core. A number of t hermal effects m ust be considered when analyzing the f uel rod performance.

The cladding can be in conta ct with the fuel pellet at some time in the fuel lifetime. Clad-pellet interaction occurs if the fuel pellet temperature is increased aft er the cladding is in contact with the pellet. Clad-pellet int eraction is discussed in Subsection 4.2.3.3.a.

The potential effects of ope ration with waterlogged fuel are discussed in Su bsection 4.2.3.3.d in which it is concluded that waterlogging is not a concern during operation al transients.

Clad flattening, as shown in Reference 5, has been observed in some operating p ower reactors. Thermal expansion (axial) of the fuel pellet stack against a flattened section of c ladding could cause failure of the cladding. This is no longer a concern because cladding flattening is precluded during the fuel residence in the core (see Subs ection 4.2.3.1).

Potential differential t hermal expansion between the fuel rods and the gu ide thimbles during a transient is considered in the design. Excessive bowing of the fuel rods is precluded because the grid assemblies allow ax ial movement of the fuel rods relative to the grids.

Specifically, thermal expansion of the fuel ro ds is considered in the grid design so that axial loads imposed on the fuel rods during a th ermal transient will not result in excessively bowed fuel rods.

f. Fuel Element Burnout a nd Potential Energy Release As discussed in Subsecti on 4.4.2.2, the core is protected from DNB over the full range of possible operating conditions. In the extremely unlikely event that DNB shoul d occur, the cladding temperature will rise due to the ste am blanketing at the rod surface and the consequent degradation in heat transfer. During this t ime, there is potential for chemical reaction between the claddi ng and the

B/B-UFSAR 4.2-29 REVISION 6 - DECEMBER 1996 coolant. However, because of the relatively good film boiling heat transfer following DNB, the energy release resulting from this reaction is insignificant compar ed to the power produced by the fuel.

g. Coolant Flow Blockag e Effects on Fuel Rods This evaluation is presented in Subs ection 4.4.4.7.

4.2.3.4 Spacer Grids The coolant flow channels are es tablished and maintained by the structure composed of grids and guide thimbl es. The lateral spacing between fuel rods is provided and controlled by the support dimples of adjacent grid cells.

Contact of the fuel rods on the dimples is m aintained through th e clamping f orce of the grid springs. Lateral motion of the fuel rods is opposed by the spring force and the internal moments gene rated between the spring and the s upport dimples.

Grid testing is discussed in Reference 7.

The fuel assembly component stre ss levels are limited by the grid design. For example, s tresses in the f uel rod due to thermal expansion and Zircaloy or ZIRLO irra diation growth are limited by the r elative motion of the rod as it slip s over the grid spring and dimple surfaces.

4.2.3.5 Fuel Assembly

a. Loads Applied by Core Restrain System The upper core plate bears downward against the fuel assembly top nozzle springs.

The springs are designed to accommodate the differential thermal expansion and irradiation growth between the fuel assembly and the core internals.

b. Analysis of Accident Loads As shown in Reference 7 and in Appendix A of Reference 19, grid crush ing tests and seismic and LOCA evaluations show that the fuel assembly will maintain a geometry that is capable of being cooled under the worst-case acciden t Condition IV event.

References 22 and 23 doc ument the acceptability of fuel assemblies in the two Byron Unit 1 locations that have only a single upper core plate fuel locating pin. Referen ce 21 documents the acceptability of fuel assemb lies in the six Byron Unit 2 locations whi ch have only a sin gle upper core plate fuel locating pin.

B/B-UFSAR 4.2-29a REVISION 5 - DECEMBER 1994 A prototype fuel assembly has been subjected to column loads in excess of those expected in no rmal service and faulted conditions (Reference 7).

The VANTAGE 5 test program description is given in Appendix A of Reference 19.

No interference between control rod insertion and thimble tubes will occur during a safe shutdown earthquake (SSE).

B/B-UFSAR 4.2-30 REVISION 2 - DECEMBER 1990 Stresses in the fuel assembl y caused by tripping of the rod cluster control assembly have little influence on fatigue bec ause of the small number of events during the life of an assembly. Assembly components and prototype fuel assemblies made from production parts have been subjected to structural tests to verify that the design bases requirements are met (Reference 7).

c. Loads Applied in Fuel Handling The fuel assembly design loads for shipping have been established at 6g later al and traverse and 4g longitudinal. Accelerom eters are permanently placed into the shipping con tainers to monitor and detect fuel assembly acc elerations that would exceed the criteria.

Past history and experience have indicated that loads which exceed the allowable limits rarely occur.

Exceeding the limits requires reinspection of the fuel assembly for damage. Tests on various fuel assembly components such as the grid assembly, sleeves, inserts and structure joints have been performed to assure tha t the shipping design limits do not result in impairment of fuel assembly function.

4.2.3.6 Reactivity Con trol Assembly and Burn able Absorber Rods

a. Internal Pressure and Cl adding Stresses During Normal, Transient and Ac cident Conditions The designs of the burna ble absorber and source rods provide a sufficient cold void volume to accommodate the internal press ure increase during operation. This is not a concern for the Ag-In-Cd or hafnium absorber rod because no gas is released by the absorber material.

For the borosilicate bur nable absorber rod, the use of glass in tubular form provides a central void volume along the len gth of the rods.

For the wet annular burnable absorber (WABA) rod, the use of annular aluminum oxide-boron carbide pellets provides two concentric void volumes surrounding the pellets. For the source rods a void volume is provided in the cladding in order to limit the internal pressure increase until end-of-life (see Figures 4.2-14 a nd 4.2-15).

The stress analysis of the burnable absorber and source rods assumes 100 perc ent gas release to the rod void volume in addition to the initi al pressure within the rod.

The stress anal ysis of the WABA rod assumes a conservative 30 percent helium gas

B/B-UFSAR 4.2-31 release to the void volu me in addition to the initial pressure within the rod.

During normal transi ent and accident conditions the void volume limits the internal pressu res to values which satisfy the criteria in Subsection 4.2.1.6.

These limits are establi shed not only to assure that peak stresses do not reach unacceptable values, but also to limit the amplitude of the oscillatory stress compo nent in consideration of the fatigue characterist ics of the materials.

Rod, guide thimble, and dashpot flow analyses indicate that the flow is sufficient to prevent coolant boiling. Therefore, clad temperatures at which the clad material has adequate strength to resist coolant opera ting pressures a nd rod internal pressures are maintained.

b. Thermal Stability of the Absorber Material, Including Phase Changes and Thermal Expansion The radial and axial tem perature profiles have been determined by considering gap conductance, thermal expansion, and n eutron or gamma heating of the contained material as we ll as gamma heating of the clad.

The maximum temperature of the absorber material (whether Ag-In-Cd or Hafnium) was calculated to be substantially less than the material's melting point, and occurs axially at only the highest flux region. The thermal exp ansion properties of the absorber material and the phase changes are discussed in Ref erences 2 and 16.

The maximum temperature of the borosilicate glass was calculated to be about 1300

°F and takes place following the initial ri se to power. As the operating cycle proceeds, the glass temperature decreases for the following reasons: (1) reduction in power generation due to boron 10 depletion, (2) better gap condu ctance as the he lium produced diffuses to the gap, and (3) external gap reduction due to borosilicate glass creep. The maximum absorber temperature for the WABA rod is less than 1200°F and takes pla ce following the initial rise to power. As the oper ating cycle proceeds, the absorber temperature rapidly decreases below 1000

°F.

Sufficient diametral and end clearances have been provided in the neutron absorber, borosilicate

B/B-UFSAR 4.2-32 burnable poison, WABA, and sou rce rods to accommodate the relative thermal exp ansions between the enclosed material and the surrounding clad and end plug.

c. Irradiation Stability of the Absorber Ma terial, Taking into Consideration Gas Release and Swelling The irradiation stability of the absorber material is discussed in References 2 and 16. Irradiation produces no deleterious effe cts in the absorber material.

Gas release is not a concern for the absorber rod because no gas is re leased by the absorber material.

Sufficient diametral and end clearances are provided to accommodate s welling of the a bsorber material.

Based on experience with bor osilicate glass, and on nuclear and thermal calculat ions, gross swelling or cracking of the glas s tubing is not expected during operation. Some minor creep of the glass at the spot, on the inner surface of the tube, could occur but would continue only until the glass came in contact with the inn er liner. The wall thickness of the inner liner is sized to provi de adequate support in the event of slumping, and to collapse locally before rupture of the exterior cladding if unexpected large volume chan ges, due to swelling or cracking, should occur.

The ends of the inner liner are open to allow helium, which diffuses out of the glass, to occupy the central void.

The WABA rod cladding and rod initial internal pressure have been designed so that the clad will not rely upon the pellets for suppor t under all Condition I and II event

s. The WABA rod design precludes irradiation-induce d matrix damage and B 4 C particle swelling which may lead to gross pellet disintegration or creep.

Rodlet prepressurization will support the outer clad against irradiation-induced creep collap se in the event of 0% gas release (worst case) from the absorb er for the design life. Ca lculations also verify the clad integrity under Condition I and II circumstances with a conservat ive maximum gas rele ase of 30%.

The maximum outer clad strain due to creep has been demonstrated to be less than the all owable strain limit.

B/B-UFSAR 4.2-33 d. Potential for Chemical Interaction The structural materials selected have good resistance to irradiation damage and are compatible with the reactor environment.

Corrosion of the materials exposed to the coolant is quite low and proper control of c hloride and oxygen in the coolan t will prevent the occurrence of stress corrosion. The potential for interference with rod cluster control movement due to possible corrosion phenomena is very low.

4.2.4 Testing

and Inspection Plan 4.2.4.1 Quality Assurance Plan

The Quality Assurance Pr ogram Plan of the Westinghouse Nuclear Fuel Division, as su mmarized in Refere nce 11, has been developed to serve t he division in plann ing and monitoring its activities for t he design and manufacture of nuclear fuel assemblies and assoc iated components.

The program provides for control over all activities affecting product quality, commencing wi th design and development and continuing through p rocurement, materials handling, fabrication, testing and inspection, storage, and transportation. The program also provides for the indoctrination and training of personnel and for the auditing of activities a ffecting product quality through a formal auditing program.

Westinghouse drawings and pr oduct, process, and material specifications identify the inspections to be performed.

4.2.4.2 Quality Control

Quality control philosop hy is generally base d on the following inspections being performed to a 95% confidence that at least 95% of the product m eets specification, unless otherwise noted.

a. Fuel System Co mponents and Parts The characteristics inspecte d depend upon the component parts and in clude dimensions, visual appearance, audits of te st reports, material certification, and nondestructive examination such as X-ray and u ltrasonic tests.

All material used is a ccepted and released by Quality Control.

B/B-UFSAR 4.2-34 REVISION 3 - DECEMBER 1991 b. Pellets Inspections are performed for dimensional characteristics such as diameter, de nsity, length, and squareness of ends.

Additional visual inspections are performed for cracks, chips, and surface conditions according to approved standards.

Density is determined in ter ms of weight per unit length and is plotted on zone charts used in controlling the process.

Chemical analyses are taken on a specified sample basis throughout pellet production.

c. Rod Inspection Fuel rod, control rod, b orosilicate burn able poison, WABA, and source rod i nspection consists of the following nondestructive exa mination techniques and methods, as applicable.

1. Leak Testing Each rod is tested usi ng a calibrated mass spectrometer with heli um being the detectable gas. 2. Enclosure Welds

Rod welds are inspected by ultrasonic or x-ray in accordance with Wes tinghouse specifications.

3. Dimensional All fuel rods are dimens ionally inspected prior to final release. The requirements include such items as length, cambe r, weld diameter, and visual appearance.

4. Plenum Dimensions All fuel rods are inspec ted by gamma scanning, or other approved meth ods as discussed in Subsection 4.2.4.4 to ensure proper plenum dimensions.

5. Pellet-to-Pellet Gaps All fuel rods are inspec ted by gamma scanning or other methods as discussed in Subsection 4.2.4.4 to ensure that no sign ificant gaps exist between pellets.

B/B-UFSAR 4.2-35 REVISION 2 - DECEMBER 1990 6. All fuel rods are acti ve gamma scanned to verify enrichment control p rior to acceptance for assembly loading.

7. Traceability Traceability of rods and associated rod components is established by Quality Control.

d. Assemblies Each fuel, control, boro silicate burnable absorber, WABA, and source rod a ssembly is inspected for compliance with drawing and/or specification requirements. Other core component in spection and specification requirements a re given in Subsection 4.2.4.3.

e. Other Inspections The following inspections ar e performed as part of the routine inspection operation:

1. Tool and gauge inspect ion and control including standardization to prima ry and/or secondary working standards.

Tool inspection is performed at prescribed intervals on all serialized tools.

Complete records are kept of calibration and conditions of tools.

2. Audits are per formed of inspec tion activities and records to assure th at prescribed methods are followed and that re cords are correct and properly maintained.

3. Surveillance inspection where appropriate, and audits of outside cont ractors are performed to ensure conformance with specified requirements.

f. Process Control

To prevent the possibili ty of mixing enrichments during fuel manufacture and assembly, strict enrichment segre gation and other process controls are exercised.

The UO 2 powder is kept in se aled containers or is processed in a c losed system. The containers are either fully ide ntified both by descriptive tagging and preselected colo r coding or, for the closed system, the material is monitored by a computer data management info rmation system. For the sealed

B/B-UFSAR 4.2-36 REVISION 7 - DECEMBER 1998 container system, a West inghouse identification tag completely describing the co ntents is affixed to the containers before transf er to powder storage.

Isotopic content is co nfirmed by analysis.

Powder withdrawal fr om storage can be made by only one authorized group , which directs the powder to the correct pellet produ ction line. All pellet production lines are phy sically separated from each other and pellets of o nly a single nominal enrichment and d ensity are pro duced in a given production line at any given time.

Finished pellets are placed on trays and transferred to segregated st orage racks within the confines of the pelleting area.

Samples from each pellet lot are tested for p hysical and chemical properties including isotopic content and impurity levels prior to acceptance by Quality Contr ol. Physical barriers prevent mixing of pellets of dif ferent nominal designs and enrichment in th is storage area.

Unused powder and subst andard pellets a re returned to storage for disposition.

Loading of pellets into the cladding is performed in isolated production lines an d again only one density and enrichment is loaded on a line at a time.

A serialized traceability code is placed on each fuel tube which identifies the enrichmen

t. The end plugs are inserted a nd then welded to seal the tube.

The fuel tube remains co ded and traceability identified until just pr ior to installation in the fuel assembly.

At the time of installat ion into an assembly, a matrix is generated to i dentify each rod in its position within a given assembly. The top nozzle is inscribed with a permanent identification number providing traceability to th e fuel conta ined in the assembly.

Similar traceability is provided for burnable poison, source and contr ol rods as required.

4.2.4.3 Core Component T esting and Inspection Tests and inspections were perfo rmed on each c ore component to verify the mecha nical characteristic

s. In the c ase of the rod cluster control assembly, prototype testing had been conducted and both manufacturing test/insp ections and func tional testing at the plant site were performed.

B/B-UFSAR 4.2-37 REVISION 7 - DECEMBER 1998 During the compo nent manufacturing p hase, the following requirements applied to the core components to assure the proper functioning during r eactor operation:

a. All materials were procured to specifications to attain the desired standard of quality.

b. Each spider was proof tested by applying a 5000 pound load to the spider bod y, so that a pproximately 310 pounds were appl ied to each vane. This proof load provided a bend ing moment at th e spider body approximately equivalent to 1.4 times the load caused by the accele ration imposed by the control rod drive mechanism.

c. All rods were checked for integrity by t he methods described in Sub section 4.2.4.2.c.

d. To assure proper fitup with the fuel assembly, the rod cluster control, borosilicate burn able absorber, WABA, and source assemblies were installed in the fuel assembly and checked fo r restriction or binding in the dry condi tion. Also a straig htness of 0.01 in/ft was required on the entire inserted length of

each rod assembly.

The rod cluster control assemblies were func tionally tested, following core loadi ng but prior to in itial criticality to demonstrate reliable ope ration of the assembli es. Each assembly was operated at no f low/cold conditions.

In addition, each assembly was operated (and t ripped) at full flow/operating temperature conditions.

Those control rods whose drop times fell outside the two-sigma limit of the drop time data for all control rods were tested a s ufficient number of times (3 times) to reasonably ensure proper performance duri ng subsequent plant operations.

In order to demonstr ate continuous free movement of the rod cluster control assembli es and to ensure acceptable core power distributions during operation, partial movement checks are performed on the rod cluster c ontrol assemblies as required by the Technical Specif ications. In addition, periodic drop tests of the rod cluster c ontrol assemblies are performed after each refueling shutdown to demonstrate continued ability to meet trip time requirements, to ensure c ore subcriticali ty after reactor trip, and to limit potential reactivity insert ions from a hypothetical rod cluster con trol assembly ejection.

If a rod cluster control assembly cannot be moved by its mechanism, adjustments in the boron concentr ation ensure that adequate shutdown margin would be achieved follo wing a trip.

Thus, inability to move one rod cluster cont rol assembly can be tolerated. More than one in operable rod cluster control B/B-UFSAR 4.2-38 REVISION 10 - DECEMBER 2004 assembly could be tole rated, but would impos e additional demands on the plant operato

r. Therefore, the allowable number of inoperable rod c luster control assemblies has been limited to one.

4.2.4.4 Tests and In spections by Others If any tests and ins pections are to be p erformed on behalf of Westinghouse, Westinghou se will review and a pprove the quality control procedures, insp ection plans, etc. to be utilized to ensure that they are e quivalent to the d escription provided above and are perfor med properly to meet all Westinghouse requirements.

4.2.4.5 Onsite Inspection Detailed written procedu res are used by the station staff for the postshipment inspection of a ll new fuel and associated components such as c ontrol rods and other inserts. The procedures are speci fic and have been fi eld tested. This process is subject to QA aud it and inspe ction under the applicable portions of the approved QA progr am to ensure proper implementation and c ompliance with com mitments. This is discussed in the QA Topical Re port, NO-AA-10.

A master fuel handling procedure specifies the seq uence in whi ch handling and inspection takes place.

4.2.5 References

1. Slagle, W. H., "Operational Experien ce with Westinghouse Cores," WCAP-8183 (L atest Revision).

2. Beaumont, M. D., et al., (Ed.), "Propert ies of Fuel and Core Component Materials

," WCAP-9179, Revision 1 (Proprietary) and WCAP

-9224, July, 1978.

3. Hellman, J. M., (Ed.

), "Fuel Densifica tion Experimental Results and Model fo r Reactor Operation," WCAP-8218-P-A, March 1975 (Proprietary) and WCAP-8219-A, March 1975.

4. Weiner, R. A., et al., "Improved Fuel Pe rformance Models for Westinghouse Fuel Rod Design and Safety Evaluations," WCAP-11873-P-A (Propriet ary) and WCAP-10873-A (Nonproprietary), August 1988.

5. Not used.

6. Eggleston, F., " Safety Related Research and Development for Westinghouse Pressurized Wat er Reactors - Program Summaries, WCAP-8768, Re vision 2, October 1978.

7. Davidson, S.L., et al., (Ed.), "Verification Testing and Analyses of the 17x17 Optimized Fuel Ass embly," WCAP-9401 (Proprietary) and WCAP

-9402-A, August 1981.

B/B-UFSAR 4.2-39 REVISION 3 - DECEMBER 1991 8. O'Donnell, W. J. and Langer, B. F., "Fatigue Design Basis for Zircaloy Component s," Nuclear Science and Engineering , 20, 1-12, 1964.

9. Western New York Nuclear Res earch Center Correspondence with the AEC on February 11 and August 27, 1971, Docket 50-57. 10. Stephan, L. A., "The Effects of Claddi ng Material and Heat Treatment on the Response of Waterlogged UO 2 Fuel Rods to Power Bursts," IN-ITR-11 1, January 1970.

11. Dollard, W. J., "Nuclear Fue l Division Quali ty Assurance Program Plan," WCAP-7800, Revision 4-A, April 1975.

12. Risher, D. et al., "Safe ty Analysis for the Revised Fuel Rod Internal Pressure De sign Basis," WCAP-8963 (Proprietary), Novem ber 1976, WCAP-896 4, August 1977.

13. Davidson, S. L., Iorii, J.

A., "Reference Core Report 17 x 17 Optimized Fuel Assemb ly," WCAP-9500-A, May 1982.

14. Skaritka, J., (Ed.), "Fuel Rod Bow Evalu ation," WCAP-8691, Rev. 1, July 1979.

15. Skaritka, J., "W estinghouse Wet Annular Burnable Absorber Evaluation Report," WC AP-10021, February 1982.

16. Beaumont, M.D, et al., (Ed.), "Hafnium," Appendix A to WCAP-9179, Revision 1 (Proprietary) and WCAP-9224, February 1981. 17. Letter from E.P.Rahe , Jr. (Westinghouse) to L.E. Phillips (NRC) dated April 12, 1984 (NS-EPR-2893).

Subject:

Fuel Handling Load Criter ia (6g vs. 4g).

18. Davidson, S. L., (Ed.), et al., "Extended Burnup Evaluation of Westinghouse Fuel

," WCAP-10126-O-A (Nonproprietary), December 1985.

19. Davidson, S. L., (Ed.), "Reference Core Report-VANTAGE 5 Fuel Assembly," WCAP-104 44-P-A, September 1985.

20. Letter from R. A Chr zanowski (CECo) to T.

E. Murley (NRC) dated July 31, 1 989.

Subject:

Byron Station Units 1 and 2 application for Amendm ent to Facility Operating Licenses NPF-37 and NPF-66.

21. Rarig, B.E., " Fuel Assembly Alig nment Pin Removal," Westinghouse SECL 90-561, Re vision 1, November 7, 1990.

B/B-UFSAR 4.2-40 REVISION 10 - DECEMBER 2004 22. Rarig, B. E., "R emoval of Six Fuel Assembly Alignment Pins," Westinghouse SECL 93-054, Revision 1, March 22, 1993.

23. Humphries, B. S., "Fuel Assembly Alignment Pin Damage -

Final Configuration," Westinghouse CAE-9 3-149, March 30, 1993. 24. Davidson, S. L., and Nuhfer, D. L. (Eds.), " VANTAGE+ Fuel Assembly Reference Core Report," WCAP-12610-A and Appendices A through D, June 1990.

25. Davidson, S. L., "Westin ghouse Fuel Criteria Evaluation Process," WCAP-12488-1-A, October 1994.

26. Kersting, P. J., et al., "As sessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel," WCAP-13589-A, March 1995.

27. Swogger, J. W., "Extended Life Wet Ann ular Burnable Absorber," Westingho use 99CB-G-0162, O ctober 1, 1999.

28. Sepp, H. A., "Re vision to Design Crite ria," WCAP-12488-A, Addendum 1-A, Revision 1, January 2002.

B/B-UFSAR 4.3-50 REVISION 9 - DECEMBER 2002 TABLE 4.3-1

NUCLEAR DESIGN KEY SAFETY PARAMETERS

SAFETY PARAMETER VANTAGE 5 Reactor Core Power (MWt) 3586.6 Core Average Coolant Temperature 579.5 - 592.7 FP (°F) Coolant System Pressure (psia) 2250 Core Average Linear Heat Rate 5.73 (kW/ft) Most Positive MTC (pcm/

°F)** +7.0 to 0.00 Most Positive MDC(k/g/cc) 0.54 Doppler Temperature Coefficient -.91 to -2.9 (pcm/°F) Doppler Power Coefficient -9.55 to -6.05 (pcm/% Power) Least Negative Doppler Power Coefficient -19.4 to -12.6 (pcm/% Power) Most Negative Beta-Effective .0044 - .0075

Boron Worth (pcm/ppm) -5 to -16 Shutdown Margin (%delta-rho) 1.3

Nuclear Design F N H 1.574* Total Heat Flux Hot Channel Factor, F Q 2.60

• For VANTAGE 5/VANTAGE+, the va lues in Subsecti on 4.3.2.2.6 include the 1.08 uncer tainty allowance.

• • Control rod withdrawal limits may be required to preclude an MTC more positive than the Tec hnical Specifi cations limit.

Note: 1 pcm = (% mille rho) = 10

-5 where is calculated from two statepoint values of Keff by n (k1/k2)

B/B-UFSAR 4.3-51 REVISION 6 - DECEMBER 1996 TABLE 4.3-1 (Cont'd)

Original Negative Positive Boron Concentrations (ppm) MTC Design MTC Design Zero Power, keff = 1.00, Cold, Rod Cluster Control Assemblies Out 1258 1890 Design Basis Ref ueling Boron Concentration 2000 2300 Zero Power, keff < 0.95, Cold, Rod Cluster Control Assemblies I n, 100 ppm allowance included 1351 2011 Zero Power, No Xenon, keff = 1.00, Hot, Rod Cluster Control Assemblies Out 1259 2056 Full Power, No Xenon, keff = 1.0, Hot, Rod Cluster Control Assemblies Out 1128 1923 Full Power, Equi librium Xenon, k eff = 1.0, Hot, Rod Cluster Control Assemblies Out 833 1590 Reduction with Fuel Burnup Reload Cycle, ppm/GWd/Mtu*** See Figure 4.3-3

• • • Gigawatt, Day (GWd) = 100 0 Megawatt Day (1000 MWd).

Burnable absorber rods m ay reduce the boron depletion rate.

B/B-UFSAR 4.3-52 REVISION 9 - DECEMBER 2002 TABLE 4.3-2 REACTIVITY REQUIREME NTS FOR ROD CLUSTER CONTROL ASSEMBLIES TYPICAL END OF LIFE (EQUILIBRIUM CYCLE) REACTIVITY EFFECTS

1. Control requirements

Fuel temperature (Doppler), % 0.94 Moderator temperature, % 0.75 Void, % 0.05 Redistribution, %0.90 Rod Insertion Allowance, % 0.50 2. Total Control Re quirements, % 3.14 3. Estimated Ag-In-Cd or Hafnium Rod Cluster Control Assembly Worth (53 Rods) a. All full length assemblies 6.50 inserted, % b. All but one (highest worth) 5.57 assemblies inserted, % 4. Estimated Rod Cluster Control Assembly credit with 7 pe rcent adjustment to accommodate uncertai nties (3b - 7 percent), % 5.18 5. Shutdown margin available (4-2), % 2.04*

• The design basis minimum sh utdown margin is 1.3 %.

B/B-UFSAR

4.3-53

REVISION 2 - DECEMBER 1990 TABLE 4.3-3 BENCHMARK CRITICAL EXPERIMENTS(35, 37) GENERAL ENRICHMENT SEPARATING DESCRIPTION W/O U235 REFLECTOR MATERIAL SOLUBLE BORON (ppm) 1. UO 2 rod lattice 2.46 water water 0

2. UO 2 rod lattice 2.46 water water 1037

3. UO 2 rod lattice 2.46 water water 764

4. UO 2 rod lattice 2.46 water B4C pins 0

5. UO 2 rod lattice 2.46 water B4C pins 0

6. UO 2 rod lattice 2.46 water B4C pins 0

7. UO 2 rod lattice 2.46 water B4C pins 0

8. UO 2 rod lattice 2.46 water B4C pins 0

9. UO 2 rod lattice 2.46 water water 0

10. UO 2 rod lattice 2.46 water water 143

11. UO 2 rod lattice 2.46 water stainless steel 514

12. UO 2 rod lattice 2.46 water stainless steel 217

13. UO 2 rod lattice 2.46 water borated aluminum15

14. UO 2 rod lattice 2.46 water borated aluminum92

15. UO 2 rod lattice 2.46 water borated aluminum395

16. UO 2 rod lattice 2.46 water borated aluminum121

17. UO 2 rod lattice 2.46 water borated aluminum487

18. UO 2 rod lattice 2.46 water borated aluminum197

19. UO 2 rod lattice 2.46 water borated aluminum634

20. UO 2 rod lattice 2.46 water borated aluminum320

21. UO 2 rod lattice 2.46 water borated aluminum72 22. U metal cylinders 93.2 bare air 0 23. U metal cylinders 93.2 bare air 0 24. U metal cylinders 93.2 bare air 0 25. U metal cylinders 93.2 bare air 0 26. U metal cylinders 93.2 bare air 0

27. U metal cylinders 93.2 bare air 0 28. U metal cylinders 93.2 bare plexiglass 0 29. U metal cylinders 93.2 paraffin plexiglass 0 30. U metal cylinders 93.2 bare plexiglass 0 31. U metal cylinders 93.2 paraffin plexiglass 0 32. U metal cylinders 93.2 paraffin plexiglass 0 33. U metal cylinders 93.2 paraffin plexiglass 0 B/B-UFSAR 4.3-54 REVISION 3 - DECEMBER 1991 TABLE 4.3-4

AXIAL STABILITY INDEX PR ESSURIZED WATER REACTOR CORE WITH A 12-FOOT HEIGHT

BURNUP C B STABILITY INDEX (hr

-1) (MWd/Mtu) F Z (ppm) Exp Calc 1550 1.34 1065 -0.041 -0.032

7700 1.27 700 -0.014 -0.006

Difference: +0.027 +0.026

B/B-UFSAR

4.3-55 R

EVISION 2 - DECEMBER 1990 TABLE 4.3-5

TYPICAL NEUTRON FLUX LEVELS (n/cm 2 -sec) AT FULL POWER

E > 1.0 Mev 5.53 Kev < E

.625 ev E E < .625 ev 1.0 Mev < 5.53 Kev (nv) 0 Core Center 6.51 x 10 13 1.12 x 10 14 8.50 x 10 13 3.00 x 10 13 Core Outer Radius at Midheight 3.23 x 10 13 5.74 x 10 13 4.63 x 10 13 8.60 x 10 12 Core Top, on Axis 1.53 x 10 13 2.42 x 10 13 2.10 x 10 13 1.63 x 10 13 Core Bottom, on Axis 2.36 x 10 13 3.94 x 10 13 3.50 x 10 13 1.46 x 10 13 Pressure Vessel Inner Wall, Azimuthal Peak, Core Midheight 2.77 x 10 10 5.75 x 10 10 6.03 x 10 10 8.38 x 10 10

B/B-UFSAR 4.3-56 REVISION 2 - DECEMBER 1990 TABLE 4.3-6

COMPARISON OF MEASUR ED AND CALCULATE D DOPPLER DEFECTS

CORE BURNUP CALCULATED PLANT FUEL TYPE (MWd/Mtu) MEASURED (pcm) (pcm) 1 Air-filled 1800 1700 1710

2 Air-filled 7700 1300 1440

3 Air and 8460 1200 1210 helium-filled

B/B-UFSAR 4.3-57 REVISION 2 - DECEMBER 1990 TABLE 4.3-7

SAXTON CORE II ISOTOPICS ROD MY, AXIAL ZONE 6

LEOPARD ATOM RATIO MEASURED*

2 PRECISION (%)

CALCULATION U-234/U 4.65 x 10

-5 +/-29 4.60 x 10-5 U-235/U 5.74 x 10

-3 +/-0.9 5.73 X 10-3 U-236/U 3.55 x 10

-4 +/-5.6 3.74 x 10-4 U-238/U 0.99386

+/-0.01 0.99385 Pu-238/Pu 1.32 x 10

-3 +/-2.3 1.222 x 10

-3 Pu-239/Pu 0.73971

+/-0.03 0.74497 Pu-240/Pu 0.19302

+/-0.2 0.19102 Pu-241/Pu 6.014 x 10

-2 +/-0.3 5.74 x 10-2 Pu-242/Pu 5.81 x 10

-3 +/-0.9 5.38 x 10-3 Pu/U** 5.938 x 10

-2 +/-0.7 5.970 x 10

-2 Np-237/U-238 1.14 x 10

-4 +/-15 0.86 x 10-4 Am-241/Pu-239 1.23 x 10

-2 +/-15 1.08 x 10-2 Cm-242/Pu-239 1.05 x 10

-4 +/-10 1.11 x 10-4 Cm-244/Pu-239 1.09 x 10

-4 +/-20 0.98 x 10-4

• Reported in Reference 29

• • Weight ratio B/B-UFSAR 4.3-58 REVISION 2 - DECEMBER 1990 TABLE 4.3-8 CRITICAL BORON CONCENTRATION S, (ppm) HZP, BOL

PLANT TYPE MEASURED CALCULATED 2-Loop, 121 Assemblies 10 foot core 1583 1589 2-Loop, 121 Assemblies 12 foot core 1625 1624 2-Loop, 121 Assemblies 12 foot core 1517 1517 3-Loop, 157 Assemblies 12 foot core 1169 1161 3-Loop, 157 Assemblies 1344 1319 12 foot core 4-Loop, 193 Assemblies 1370 1355 12 foot core 4 Loop, 193 Assemblies 1321 1306 12 foot core

B/B-UFSAR 4.3-59 REVISION 2 - DECEMBER 1990 TABLE 4.3-9 BENCHMARK CRITICAL EXPERIMENTS B 4 C CONTROL ROD WORTH

WREC NO. OF NO. OF MEASURED(a) CALCULATED CRITICAL FUEL CONTROL WORTH, WORTH, EXPERIMENT RODS RODS (in.) % % 2A 888 12 .395 OD B 4C 8.20 8.37 3B 888 12 .232 OD B 4C 4.81 4.82 4B 884 16 .232 OD B 4C 6.57 6.35 5B 945 16 .232 OD B 4C 5.98 5.83

____________________

(a)The measured wort h was derived from the calculated value of ln (k 1/k 2), where k 1 and k 2 were calculated with the measured buckling before and after insert ion of the contr ol rods, which replace fuel rods in arrays at the center of the experiment.

The standard deviation in the measured w orth is about 0.3% based on the uncertainties in the measured axial bucklings.

AG-IN-CD COMPARISON OF MEASURED AND CALC ULATED ROD WORTH 4-LOOP PLANT, 193 ASSEMBLIES, 12-FOOT CORE MEASURED (pcm) CALCULATED (pcm) Bank D 1403 1366 Bank C 1196 1154 All Rods In Less One 6437 6460

ESADA Critical*, 0.69 Inch Pitch, 2 w/o PuO 2 , 8% Pu240 9 Control Rods 6.21 inch rod separation 2250 2250 2.07 inch rod separation 4220 4160 1.38 inch rod separation 4100 4019

____________________

• Reported in Reference 30.

B/B-UFSAR 4.3-60 REVISION 2 - DECEMBER 1990 TABLE 4.3-9 (Cont'd)

BENCHMARK CRITICAL EXPERIMENT HAFNIUM CONTROL ROD WORTH

CONTROL NO. OF MEASURED (b) CALCULATED (b) ROD FUEL WORTH WORTH CONFIGURATION RODS (PPM B-10) (PPM B-10) 9 Hafnium Rods, 1192 138.3 141.0 0.341" OD

____________________ (b)Calculated and measured wor ths are given in terms of an equivalent change in B-10 concentration.

B/B-UFSAR 4.3-61 REVISION 2 - DECEMBER 1990 TABLE 4.3-10 COMPARISON OF MEASURED A ND CALCULATED MODERATOR COEFFICIENTS AT HZP, BOL

PLANT TYPE/ MEASURED iso* CALCULATED iso CONTROL BANK CON FIGURATION (pcm/

°F) (pcm/°F) 2-loop, 121 assemblies, 12 foot core D at 180 steps +0.85 +1.02

D in, C at 180 steps -2.40 -1.90 C and D in, B at 165 steps -4.40 -5.58 B, C, and D in A at 174 steps -8.70 -8.12 3-loop, 157 assemblies, 12 foot core D at 160 steps -0.50 -0.50

D in, C at 190 steps -3.01 -2.75 D in, C at 28 steps -7.67 -7.02 B, C and D in -5.16 -4.45 4-loop, 193 assemblies, 12 foot core ARO -0.52 -1.2 D in -4.35 -5.7 D + C in -8.59 -10.0 D + C + B in -10.14 -10.55 D + C + B + A in -14.63 -14.45

• Isothermal coefficients, which include the Do ppler effect in the fuel. F T /k)k (ln 10 = 1 2 5 iso°

B/B-UFSAR 4.4-52 REVISION 9 - DECEMBER 2002 TABLE 4.4-1

THERMAL AND HY DRAULIC DATA

THERMAL AND HYDRAULIC DESIGN PARAMETERS Reactor Core Heat Output, (100%), MWt 3586.6 Reactor Core Heat Output, 10 6 Btu/Hr 12238.2 Heat Generated in Fuel, % 97.4 Core Pressure, Nominal, psia 2270 Pressurizer Pressure, psia 2250 Minimum DNBR at Nominal Conditions Typical Flow Channel 2.25 Thimble (Cold Wall) Flow Channel 2.16 Minimum DNBR for Design Transients Typical Flow Channel 1.33 Thimble Flow Channel 1.33 DNB Correlation(c) WRB-2 COOLANT FLOW (d) Total Vessel F low Rate, 10 6 lbm/hr (based on Minimum Measured Flow) 141.8 (based on Thermal Design Flow) 137.2 Effective Flow Rate for Heat Transfer, 10 6 lbm/hr (based on TDF) 126.6 Effective Flow Area for Heat Transfer, ft 2 54.1 Average Velocity along Fuel Rods, ft/sec (based on TDF) 15.1 Average Mass Velocity, 10 6 lbm/hr-ft 2 (based on TDF) 2.33 B/B-UFSAR 4.4-53 REVISION 9 - DECEMBER 2002 TABLE 4.4-1 (Cont'd)

THERMAL AND HYDRAULIC DESIGN PARAMETERS COOLANT TEMPERATURE, °F Nominal Inlet 556.7 Average Rise in Vessel 62.6 Average Rise in Core 66.3 Average in Core 591.7 Average in Vessel 588.0 HEAT TRANSFER Active Heat Transfer Surface Area, ft 2 57505 Average Heat F lux, BTU/hr-ft 2 207327 Maximum Heat F lux for Normal (f) Operation, BTU/hr-ft 2 539050 Average Linear Power, kW/ft 5.73 Peak Linear Power for Normal Operation,(f) kw/ft 14.9 Peak Linear Power Resulting from (g) Overpower Transients/Operator Errors (assuming a maximum overpower of

118%), kW/ft (centerline melt will not be exceeded) <22.4 Peak Linear Power for Prevention of(h) Centerline Melt, kW/ft 22.4 Temperature at Peak Linear Power for Prevention of Centerline Melt, °F 4700 Pressure Drop Across Core, psi (i) 27.5 +/- 2.7 Across Vessel, inclu ding nozzle, psi 46.1 +/- 4.6 B/B-UFSAR 4.4-54 REVISION 9 - DECEMBER 2002 TABLE 4.4-1 (Cont'd)

(a) Deleted.

(b) Deleted.

(c) The W-3 correlation is used for analysis of some accidents outside the range of app lication for the WRB-2 DNB correlation.

(d) Deleted.

(e) Deleted.

(f) This limit is associ ated with the value of F T Q = 2.60.

(g) See Subsection 4.3.2.2.6.

(h) See Subsecti on 4.4.2.11.6.

(i) Based on best estima te reactor flow rate.

B/B-UFSAR 4.4-55 REVISION 7 - DECEMBER 1998 TABLE 4.4-2 VOID FRACTIONS AT NOMI NAL REACTOR CONDITIONS

AVERAGE MAXIMUM Core (VANTAGE 5 1.1%

/VANTAGE +)

Hot Subchannel (VANTAGE 5 9.6% 27.9%

/VANTAGE +)

B/B-UFSAR

Attachment 4.4A

Additional Information On the Plant Specific Application of t he Westinghouse Improved Thermal Design Procedure To Byron/Braidwood

B/B-UFSAR 4.4A-1 4.4A Additional Infor mation On the Plant Sp ecific Application of the Westinghouse Improved Thermal Design Procedure To Byron/Braidwood The NRC Safety Evaluation Re port on WCAP-9500 entitled Reference Core R eport 17x17 Optimize d Fuel Assembly noted the specific plants using the Westinghouse Improved Thermal Design Proc edure (ITDP) must supply additional informati on on the plant sp ecific application of the ITDP to p erform thermal-hydra ulic analyses.

Thus, Byron/ Bra idwood specific resp onses to NRC information requests a re provided below.

4.4A.1 Request 1 Provide the sensitiv ity factors (S i) and their range of applicability.

4.4A.2 Response 1 The sensitivity factors (S i) and their range of applicability are given in T able 1 of Reference 2 for Byron/Braidwood. Please note that these values are the same as those used in WCAP-9500 with the exception of the range for vessel flow.

The range on flow for Byron/

Braidwood has been exten ded down to 273270 gpm (70%

flow) with no change in the correspond ing sensitivity factor being required.

4.4A.3 Request 2 If the S i values used in the B yron/Braidwood analyses are different than those use d in WCAP-9500, then the applicant must reevaluate th e use of an uncertainty allowance for application of equation 3-2 of WCAP-8567, "Improved Thermal Design Procedure" and the linearity assumption must be validated.

4.4A.4 Response 2 The S i values used in Byron/Br aidwood analys es are the same as those used in WCAP-9500. Therefore, reevaluating the use of an uncertainty a llowance for application of equation 302 of WCAP-8567, "Improved Thermal Design Procedu re" and the li nearity assumption is not required.

4.4A.5 Request 3 Provide and justify the varian ces and distributions for input parameters.

B/B-UFSAR 4.4A-2 4.4A.6 Response 3 The distribution assumed for the input parameters such as pressurizer pressure, core average temperature, reactor power, and R CS flow are normal, two-sided 95+%

probability distributions.

The variances of the se parameters fo r Byron/Braidwood are consistent with the varian ces calculated in the generic response.

Specifically, the uncertainties for pressurizer pressure and cor e average temperature are identical to the generic res ponse since the sensors, process racks, and compu ter and readout devices are standard Westinghouse su pplied NSSS equipment.

Variances in reactor pow er and reactor coolant system flow are calculated based on equation 4 and equation 8 respectively in Refere nce 1. As can be seen from the equations, both prim ary and secondary side parameters are measured for power and flow calorimetrics. The error allowances for the parameters measured by Westinghouse sup plied equipment are identical to those used in the generic submittal (Reference 1). Two input parameters are measu red by non-Westi nghouse supplied instruments. Th ese are feedwater temperature and feedwater pressure. As expe cted, the error allowances for these instruments vary slightly from those used in Reference 1. The error allowances for feedwater temperature and pressure were statistically combined (as described in Reference 1) to get the total channel allowance for each parameter.

The feedwater pressure error allowance w as calculated to be less than the error allowance used in Reference 1.

Therefore, the error contribution to the reactor power and flow uncertainties from fe edwater pressu re is less than that used in the generic response.

Similarly, the error s for feedwater temperature were combined to get the total channel allowance. The total allowance was found to be slightly hig her than that used to calculate RCS flow uncert ainty in Reference 1.

However, the error allowance from feedwater temperature is very small relati ve to the other co ntributing errors and in fact this small additio nal error is absorbed in the statistical comb ination. Therefore, the flow uncertainty calculated in Re ference 1 is applicable for Byron/Braidwood.

As stated in Reference 1, the flow cal orimetric can be performed one of several way

s. Commonwealth Edison plans to do a precis ion flow calorimetric at the beginning of the cycle and normalize the loop elbow

B/B-UFSAR 4.4A-3 taps. For monthly surve illance to assure plant operation consistent with th e ITDP assum ptions, the loop flows will be read o ff the plant proce ss computer. The total flow uncertain ty associated with this method was calculated in Reference 1 and is applicable to the Byron/Braidwood units.

It is to be noted that the total channel allowance for feedwater temperature was calc ulated to be l ess than the error assumed for the re actor power uncertainty calculation in Reference 1.

Therefore, the power uncertainty for Byro n/Braidwood is bounded by the uncertainty calculated in the generic response.

4.4A.7 Request 4 Justify that the nor mal conditions used in the analyses bound all permitted modes of plant operation.

4.4A.8 Response 4 This item was addressed in R eference 1 and is applicable to the Byron/B raidwood units.

4.4A.9 Request 5 Provide a discussion of what code uncertainties, including their values, are included in the DNBR analyses.

4.4A.10 Response 5 The uncertainties included in the ITDP DNBR analyses for Byron/Braidwood are given in Table 1 of Reference 2. As a result of these values bei ng different from those used in WCAP-9500, the Design DNBR Limits also differ. The calculation of the D esign limit DNBRs for the Typical and Thimble cells are given in Reference 2, Tables 2 and 3 respectively.

Since the Design DNBR Limits given in Table 2 and 3 are different fr om those originally given, Section 4.4 has been rev ised to inco rporate the Reference 2 values.

4.4A.11 Request 6 Provide a block diag ram depicting sens or, processing equipment, computer and readout devices for each parameter channel used in the uncertainty analysis.

Within each element of the b lock diagram identify the accuracy, drift, range, span, operating limits, and setpoints. Identify the ove rall accuracy of each channel transmitter to final output and specify the minimum acceptable accur acy for use with the new procedure. Also identify th e overall accura cy of the

B/B-UFSAR 4.4A-4 final output value and maximum accuracy requirements for each input channel for this final output device.

4.4A.12 Response 6

Block diagrams are n ot provided in t his response.

However, as in the gen eric response, a table is provided in Reference 2 givin g the error breakd own from sensor to computer and readout devices.

This table is abbreviated though, giving only the error breakdowns for instruments that differ from those in Ta ble 4, "Typi cal Instrument Uncertainties," of Ref erence 1. As noted earlier, these instruments are those that measure fee dwater temperature and pressure.

4.4A.13 Request 7

If there are any changes to the THINC-IV correlation, or parameter values outside of previously demonstrated acceptable ranges, t he staff requires a reevaluation of the sensitivity factors and of t he use of equation 3-2 of WCAP-8567.

4.4A.14 Response 7 For Byron/Braidwood, the THINC-IV code and WRB-1 DNB Correlation are the same as that used in WCAP-9500.

Therefore, reevaluat ing the sensitivity factors and the use of equation 3-2 of WCAP-8567 is not required.

References

1. Westinghouse letter, NS-EPR-2577, E.

P. Rahe to C. H. Berlinger (NRC), March 31, 1982, proprietary.

2. General Electric Compa ny letter transmitting improved thermal design info rmation to the NRC (to be written), proprietary.

B/B-UFSAR 4.5-1 4.5 REACTOR MATERIALS Section 4.5 provides a discussion of the materials employed in the control rod drive system and the r eactor internals.

A more detailed evaluation of the re actor materials and reactivity control sys tems indicating the degree of conformance with the recommendatio ns of the applicable R egulatory Guides is presented in the Final Safet y Analysis Report as follows:

a. control rod drive mechan ism and reactor internals:

Chapter 3.0, b. control rod drive mechan ism testing:

Chapters 3.0, 14.0, and the Techni cal Specifications, c. control rod drive mech anism and reactor internals materials: Chapter 5.0, d. safety injection sys tem: Chapter 6.0, e. instrumentation for reac tor control and protection:

Chapter 7.0, and

f. failure of the control rod drive mechanism cooling system and chemi cal and volume c ontrol system:

Chapter 9.0.

4.5.1 Control

Rod System Structural Materials 4.5.1.1 Materials Specifications All parts exposed to reactor coolant are mad e of metals which resist the corrosive action of the water.

Three types of metals are used exclusively:

stainless steels, nickel-chromium-iron, and cobalt based alloys. In the case of stainless steels, only austenitic and m artensitic stainless s teels are used. For pressure boundary parts, martens itic stainless s teels are not used in the heat treat ed conditions which ca use susceptibility to stress corrosion cracking or accelerated corrosion in the Westinghouse pressurized wat er reactor water chemistry.

a. Pressure Boundary All pressure contain ing materials co mply with Section III of the ASME Boil er and Pressure Ve ssel Code, and are fabricated from austenitic (Type 304) stainless steel. b. Coil stack assembly The coil housings require a magnetic material. Both low carbon cast steel and ductile iron have been successfully tested for this application. On the

B/B-UFSAR 4.5-2 basis of cost and perfor mance, ductile iron was selected for the control rod d rive mechanism (CRDM).

The finished housings are zinc plated or flame sprayed to provide c orrosion resistance.

Coils are wound on bobbins of molded Dow Corning 302 material, with doubl e glass insulated copper wire.

Coils are then vacuum im pregnated with silicon varnish. A wrapping of mica sheet is secured to the coil outside diameter.

The result is a well insulated coil capable of sustained operation at 200°C. c. Latch assembly Magnetic pole pieces are fabricated from Type 410 stainless steel. All nonmag netic parts, except pins and springs, are fabricated from Type 304 stainless steel. Haynes 25 is used to fabricate link pins.

Springs are made from nickel-chromium-iron alloy (Inconel-750). Latch arm tips are clad with Stellite-6 to provide improv ed wearability. Hard chrome plate and Ste llite-6 are used selectively for bearing and wear surfaces.

d. Drive rod assembly The drive rod assemb ly utilizes a Type 410 stainless steel drive rod and disconnect rod a ssembly. The coupling is machined from ty pe 403 stain less steel.

Other parts are Type 304 sta inless steel with the exception of the springs, which are nickel-chromium-iron alloy, and the locking button, which is Haynes 25; and the belleville washers which are Inconel 718.

Several small parts (s crews and pins) are Inconel 600.

Material specifications for Cl ass 1 components of the CRDM are as follows:

CRDM, upper head SB-166 or SB-167 and SA-182 Grade F304 Latch housing SA-1 82, Grade F304 or SA-351 Grade CF8 Rod travel housing SA-182, Grade F304 or SA-336 Class F8 Cap SA-479, Type 304 Welding materials Stainless Steel Weld Metal Analysis A-8

B/B-UFSAR 4.5-3 4.5.1.2 Austenitic Stain less Steel Components

a. All austenitic stainless steel materials used in the fabrication of CRDM comp onents are p rocessed, inspected and tested to avoid sensitization and prevent intergranula r stress cor rosion cracking.

The rules covering these controls are stipulated in Westinghouse process specifications.

As applicable, these process specif ications supplem ent the equipment specifications and purchase order requirements of every individual austenitic stainless steel component regardless of the AS ME Code Classification.

Westinghouse practice is that austenitic stainless steel materials of p roduct forms with simple shapes need not be corrosion te sted provided that the solution heat treatment is followed by water quenching. Simple shapes are defined as all plates, sheets, bars, pipe and tubes, as wel l as forgings, fittings and other shaped products whi ch do not have inaccessible cavities or chamb ers that would preclude rapid cooling when water que nched. When testing is required the tests are p erformed in accordance with ASTM A 262, Prac tice A or E, as amended by Westinghouse Process S pecification 84201 MW.

If, during the c ourse of fabrication the steel is inadvertently exposed to the sensitization temperature range, 800

°F to 1500°F the material may be tested in accordance with ASTM A 262, as amended by Westinghouse Process Specificatio n 84201 MW to verify that it is not susceptible to intergranular att ack, except that testing is not required for:

1. Cast metal or weld m etal with a ferrite content of 5 percent or more, 2. Material with a carbon c ontent of 0.03%

or less that is subjected to tem peratures in the range of 800°F to 1,500

°F for less than 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />.

3. Material exposed to sp ecial processi ng provided the processing is proper ly controlled to develop a uniform product an d provided t hat adequate documentation exists of service expe rience and/or test data to demonstrate that the processing will not result in increased susceptibility to intergranular st ress corrosion.

If it is not verified th at such material is not susceptible to intergr anular attack, the material

B/B-UFSAR 4.5-4 will be re-solution anne aled and water quenched or rejected.

b. The welding of austenitic stainless steel is controlled to mitigate the occurrence of microfissuring or hot cracking in the weld.

Available data indicates tha t a minimum delta ferrite level expressed in F errite Number (FN), above which the weld metals commonly use d by Westinghouse will not be prone to hot cracking, lies som ewhere between 0 FN and 3 FN. The undi luted weld dep osits of the starting welding materials a re required to contain a minimum of 5 FN.

4.5.1.3 Other Materials

The CRDMs are cleaned prior to delivery in a ccordance with the guidance of ANSI N45.2.1. Westi nghouse personnel do conduct surveillance to ensure that manufacturers an d installers adhere to appropriate requirements.

Haynes 25 is used in small q uantities to fabrica te link pins.

The material is ordered in the solution trea ted and cold worked condition. Stress cor rosion cracking has not been observed in this application over the last 15 years.

The CRDM springs are made fr om nickel-chromium-iron alloy (Inconel-750) ordered to MIL-S-23192 or MIL-N-24114 Class A #1 temper drawn wire. Op erating experience has shown that springs made of this materia l are not subject to stress-corrosion cracking.

4.5.1.4 Cleaning and Cleanliness Control The CRDMs are cleaned prior to delivery in a ccordance with the guidance of ANSI N45.2.1. Mea sures are applied, as appropriate, to apply packaging req uirements to procureme nt orders, to review supplier packaging procedures, to apply prop er cleaning requirements, marking and iden tification and to provide protection to equipment from physical or wea ther damage, to apply special handling precautions and to define sto rage requirements.

Westinghouse quality assurance procedures are described in "Westinghouse Water Re actor Divisions Qualit y Assurance Plan," WCAP-8370, Revision 8A u pdated per letter NS-TMA-2039, From T. M.

Anderson to W. P. Haass, February 8, 1979.

4.5.2 Reactor

Internals Materials 4.5.2.1 Materials Specifications All the major material for the reactor i nternals is Type 304 stainless steel. Parts not fa bricated from Type 304 stainless steel include bolts and dowel pins, which are fabricated from

B/B-UFSAR 4.5-5 REVISION 11 - DECEMBER 2006 Type 316 stainless steel, and ra dial support key bolts, which are fabricated from Inconel-750.

Material specifications for reac tor vessel internals for emergency core cooling systems are listed in Table 5.2-4.

There are no other materials u sed in the reactor internals or core support structu res which are not ot herwise included in ASME Code,Section II I, Appendix I.

4.5.2.2 Controls on Welding The discussions provided in Subs ection 4.5.1 a re applicable to the welding of reactor i nternals and core support components.

4.5.2.3 Nondestructive Examination of Wr ought Seamle ss Tubular Products and Fittings The nondestructive exa mination of wrought seamless tubular products and fittings is in acco rdance with Sect ion III of the ASME Code.

4.5.2.4 Fabrication and Processing of Au stenitic Stainless Steel Components The discussions provided in Subsection 4.5.1.4 are applicable to the cleaning of reactor internals and core s upport structures in accordance with ANSI N45.2.1.

B/B-UFSAR 4.6-1 REVISION 7 - DECEMBER 1998 4.6 FUNCTIONAL DESIGN OF REACTIVITY CONTROL SYSTEMS

4.6.1 Information

for Control Rod Drive System (CRDS)

Figure 4.2-8 pro vides the layout of the CRDS. The CRDS is a magnetically operated jack with no hydra ulic system associated with its functioning.

The control r od drive mec hanism consists of four separa te subassemblies.

a. The pressure vessel which includes the l atch housing and rod travel housings.

b. The coil stack assembly which includes three operating coils: stationary gripper coil, movable gripper coil and lift coil.

c. The latch assembly which includes the guide tube, the stationary and the movab le pole piec es and the stationary and movab le gripper latches.

d. The drive rod assembly which includes the RCC coupling system and the drive rod.

4.6.2 Evaluation

of the CRDS The CRDS has been analyzed in detail in a fa ilure mode and effects analysis (Refere nce 1). This study, and the analyses presented in Chapter 15.0, d emonstrates that the CRDS performs its intended safety function, reactor tr ip, by putting the reactor in a subcritic al condition when a safety system setting is approached, with any assumed credible failure of a single active component. T he essential elements of the CRDS (those required to ensure reactor trip) are isolated fr om nonessential portions of the CRDS (the rod control system).

Despite the extremely low probability of a c ommon mode failure impairing the ability of the rea ctor trip system to perform its safety function, analyses have been performe d in accordance with the requirements of WASH-1270.

These analyses, documented in References 2 and 3, have demonst rated that acc eptable safety criteria would not be exceeded even if the C RDS were rendered incapable of functioning during a reactor tran sient for which their function w ould normally be expected.

The design of the control ro d drive mechanism is such that failure of the contr ol rod drive mechanism c ooling system will, in the worst cas e, result in an individu al control r od trip or a full reactor trip.

4.6.3 Testing

and Veri fication of the CRDS The CRDS was extensively tested prior to its ope ration. These tests may be subdivided into five catego ries: (1) prototype

B/B-UFSAR 4.6-2 REVISION 8 - DECEMBER 2000 tests of components, (2) prototy pe CRDS tests, (3) production tests of components following manufacture and prior to installation, (4) onsite preoper ational tests, a nd (5) initial startup tests.

In accordance with Table 14.2-65, the reactor trip system operation was verified in a startup test. This test ensured that the system operated in accordance with the safety analysis report, design requireme nts, and plant insta llation. A final test was performed in which a manual reactor trip was initiated, (after fuel load but pri or to initial criticalit y) to verify that all rods would fully insert.

The rod cluster control assembli es were dropped and the drops were timed. The time from b eginning of deca y of stationary gripper coil voltage to dashpot entry shall be less than or equal to 2.7 seconds for each rod, the Technical Specification limit.

In compliance with Tables 14.2

-66 and 14.2-66a, all rods falling outside the two-sigma limit were retested a minimum of three times each. Rods we re dropped into represen tative flow condi-tions. In addition, the CRDS is subject to periodic ins ervice tests.

These tests are conducted to verify the oper ability of the CRDS when called upon to function.

4.6.4 Information

for Combined Perfo rmance of Reacti vity Systems As is indicated in Chapt er 15.0, the only po stulated events which assume credit for re activity control systems other than a reactor trip to render t he plant subcritical are the steam line break, feedwater line break, and loss-of-coolant accident. The reactivity control sys tems for which cre dit is taken in these accidents are the reactor trip system and the safety injection system (SIS). Note that no cred it is taken for the boration capabilities of the ch emical and volume cont rol system (CVCS) as a system in the analys is of transients present ed in Chapter 15.0.

The adverse boron dilution possi bilities due to the operation of the CVCS are investigated in C hapter 15.0. Prior proper operation of the CVCS has been presumed as an initial condition to evaluate transients, and appropriate Techni cal Specifications have been prepared to ensure the correct operation or remedial action.

4.6.5 Evaluation

of Combined Performance The evaluations of the steam line break, feedwater line break, and the loss-of-coolant accident, which presume the combined actuation of the reactor trip system to the CRDS and the SIS, are presented in Chapter 15.0.

Reactor trip signals and safety injection signals for these events are generated from functionally diverse sensors and actuate diverse means of reactivity control, i.e., control rod insertion and injection of soluble poison.

B/B-UFSAR 4.6-3 REVISION 7 - DECEMBER 1998 Nondiverse but r edundant types of equipm ent are utilized only in the processing of the incoming sensor signals into a ppropriate logic, which ini tiates the protective ac tion. In particular, note that protection from eq uipment failures is provided by redundant equipment and periodic testing. Eff ects of failures of this equipment have been extensively investiga ted as reported in Reference 4. The failure mode and effects ana lysis described in this reference v erifies that any single failure will not have a deleterious effect on the engine ered safety feat ures actuation system.

4.6.6 References

1. Shopsky, W. E., "Failu re Mode and Effects Analysis (FMEA) of the Solid State Full Length Rod Control Syst em," WCAP 8976, August 1977.

2. "Westinghouse Anticipa ted Transients Without Trip Analysis," WCAP-8330, August 1974.

3. Gangloff, W. C.

and Loftus, W. D., "An E valuation of Solid State Logic Reactor Pr otection in Antici pated Transients," WCAP-7706-L (Proprietary) and WCAP-7706 (Nonproprietary), July 1971.

4. Eggleston, F. T., Rawlins, D.

H. and Petrow, J. R., "Failure Mode and Effects Analysis (FMEA) of the Engineering Safeguard Features Actua tion System," WCAP-8584 (Proprietary) and WC AP-8760 (Nonpropriet ary), April 1976.

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