LR-N19-0102, Revision 31 to Updated Final Safety Analysis Report, Chapter 4, Reactor

Revision 31 to Updated Final Safety Analysis Report, Chapter 4, Reactor

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SECTION 4 REACTOR TABLE OF CONTENTS Section Title Page 4.1

SUMMARY

DESCRIPTION 4.1-1 4.1.1 Reference for Section 4.1 4.1-4 4.2 MECHANICAL DESIGN 4.2~1 4.2.1 Fuel 4.2-2 4.2.1.1 Design Bases 4.2-2 4.2.1.1.1 Fuel Rods 4.2-3 4.2.1.1.2 Fuel Assembly Structure 4.2-5 4.2.1.2 Design Description 4.2-8 4.2.1.2.1 Fuel Rods 4.2-10 4.2.1.2.2 Fuel Assembly Structure 4.2-11 4.2.1.3 Design Evaluation 4.2-14 4.2.1.3.1 Fuel Rods 4.2-14 4.2.1.3.2 Fuel Assembly Structure 4.2-24 4.2.1.3.3 Operational Experience 4.2-28 4.2.1.3.4 Test Rod and Test Assembly Experience 4.2-28a 4.2.1.4 Testing and Inspection Plan 4.2-28a 4.2.1.4.1 Quality Assurance Program 4.2-28a 4.2.1.4.2 Quality Control 4.2-29 4.2.1.4.3 Onsite Inspection 4.2-32 4.2.2 Reactor Vessel Internals 4.2-33 4.2.2.1 Design Bases 4.2-33 4.2.2.2 Description and Drawings 4.2-34 4.2.2.3 Design Loading Conditions 4.2-40a 4.2.2.4 Design Loading Categories 4.2-42 4.2.2.5 Design Criteria Basis 4.2-43 4.2.3 Reactivity Control System 4.2-44 4-i SGS-UFSAR Revision 11 July 22, 1991

TABLE OF CONTENTS (Cont)

Section 4.2.3.1 Design Bases 4.2-44 4.2.3.1.1 Design Stresses 4.2-44 4.2.3.1.2 Material Compatibility 4.2-45 4.2.3.1.3 Reactivity Control Components 4.2-45 4.2.3.1.4 Control Rod Drive Mechanisms 4.2-48 4.2.3.2 Design Description 4.2-49 4.2.3.2.1 Reactivity Control Components 4.2-51 4.2.3.2.2 Control Rod Drive Mechanism 4.2-56 4.2.3.3 Design Evaluation 4.2-63 4.2.3.3.1 Reactivity Control Components 4.2-63 4.2.3.3.2 Control Rod Drive Mechanism 4.2-73 4.2.3.4 Tests, Verification, and Inspections 4.2-77 4.2.3.4.1 Reactivity Control Components 4.2-77 4.2.3.4.2 Control Rod Drive Mechanism 4.2-79 4.2.4 References for Section 4.2 4.2-81 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 Reactivity Feedbacks (Reactivity Coefficient) 4.3-3 4.3.1.3 Control of Power Distribution 4.3-4 4.3.1.4 Maximum Controlled Reactivity Insertion Rate 4.3-6 4.3.1.5 Shutdown Margins 4.3-7 4.3.1.6 Stability 4.3-8 4.3.1. 7 Anticipated Transients Without Trip 4.3-9 4.3.2 Description 4.3-10 4.3.2.1 Nuclear Design Description 4.3-10 4.3.2.2 Power Distribution 4.3-13 4.3.2.2.1 Definitions 4.3-13 4.3.2.2.2 Radial Power Distribution 4.3-16 4*11 SGS-UFS.AR Revision 6 February 15, 1987

TABLE OF CONTENTS (Cont)

Section 4.3.2.2.3 Assembly Power Distribution 4.3-17 4.3.2.2.4 Axial Power Distribution 4.3-17 4.3.2.2.5 Deleted 4.3-18 4.3.2.2.6 Limiting Power Distribution 4.3-20 4.3.2.2.7 Experimental Verification of Power Distribution Analysis 4.3-26 4.3.2.2.8 Testing 4.3-29 4.3.2.2.9 Monitoring Instrumentation 4.3-29 4.3.2.3 Reactivity coefficients 4.3-30 4.3.2.3.1 Fuel Temperature (Doppler) Coefficient 4.3-30 4.3.2.3.2 Moderator coefficients 4.3-31 4.3.2.3.3 Power Coefficient 4.3-34 4.3.2.3.4 comparison of Calculated and Experimental Reactivity Coefficients 4.3-34 4.3.2.3.5 Reactivity Coefficients Used in Transient Analysis 4.3-35 4.3.2.4 Control Requirements 4.3-35 4.3.2.4.1 Doppler 4.3-36 4.3.2.4.2 variable Average Moderator Temperature 4.3-36 4.3.2.4.3 Redistribution 4.3-37 4.3.2.4.4 Void content 4.3-37 4.3.2.4.5 Rod Insertion Allowance 4.3-37 4.3.2.4.6 Burn up 4.3-38 4.3.2.4.7 Xenon and Samarium Poisoning 4.3-38 4.3.2.4.8 pH Effects 4.3-38 4.3.2.4.9 Experimental Confirmation 4.3-39 4.3.2.5 control 4.3-39 4.3.2.5.1 Chemical Poison 4.3-39 4.3.2.5.2 Rod Cluster Control Assemblies 4.3-40 4.3.2.5.3 Burnable Absorbers 4.3-41 4.3.2.5.4 Peak xenon Startup 4.3-41 4.3.2.5.5 Load Follow Control and Xenon control 4.3-42

• SGS-UFSAR 4-iii Revision 17 october 16, 1998

TABLE OF CONTENTS (Cont)

Section Title Page 4.3.2.5.6 Burnup 4.3-42 4.3.2.6 Control Rod Patterns and Reactivity Worth 4.3-42 4.3.2.7 Criticality of Fuel Assemblies 4.3-45 4.3.2.8 Stability 4.3-46 4.3.2.8.1 Introduction 4.3-46 4.3.2.8.2 Stability Index 4.3-46 4.3.2.8.3 Prediction of the Core Stability 4.3-47 4.3.2.8.4 Stability Measurements 4.3-48 4.3.2.8.5 Comparison of Calculations with Measurements 4.3-50 4.3.2.8.6 Stability Control and Protection 4.3-51 4.3.2.9 Vessel Irradiation 4.3-52 4.3.3 Analytical Methods 4.3-53 4.3.3.1 Fuel Temperature (Doppler) Calculations 4.3-54 4.3.3.2 Macroscopic Group Constants 4.3-55 4.3.3.3 Spatial Few-Group Diffusion Calculations 4.3-58 4.3.3.4 Pin Power Reconstruction 4.3-60 4.3.4 References for Section 4.3 4.3-60 4.4 THERMAL AND HYDRAULIC DESIGN 4.4-1 4.4.1 Design Bases 4.4-1 4.4.1.1 Departure From Nucleate Boiling Design Basis 4.4-2 4.4.1.2 Fuel Temperature Design Basis 4.4-2a 4.4.1.3 Core Flow Design Basis 4.4-3 4.4.1.4 Hydrodynamic Stability Design Bases 4.4-4 4.4.1.5 Other Considerations 4.4-4 4.4.2 Description 4.4-5 4.4.2.1 Summary Comparison 4.4-5 4.4.2.2 Fuel Cladding Temperatures (Including Densification) 4.4-7 4.4.2.2.1 Uranium Dioxide Thermal Conductivity 4.4-9 4-iv SGS-UFSAR Revision 31 December 5, 2019

TABLE OF CONTENTS (Cent)

Section Title 4.4.2.2.2 Radial Power Distribution in UO Fuel Rods 4.4-10 4.4.2.2.3 Gap Conductance 4.4-10 4.4.2.2.4 Surface Heat Transfer Coefficients 4.4-12 4.4.2.2.5 Fuel Clad Temperatures 4.4-12

.4. 4. 2. 3 Critical Heat Flux Ratio or Departure from Nucleate Boiling Ratio and Mixing Technology 4.4-13 4.4.2.3.1 Departure from Nucleate Boiling Technology 4.4-14 4.4.2.3.2 Definition of Departure from Nucleate Boiling Ratio 4.4-lSa 4.4.2.3.3 Mixing Technology 4. 4-17a 4.4.2.3.4 Engineering Hot-Channel Factors 4.4-19 4.4.2.3.5 Effects of Rod Bow on DNBR 4.4-21a 4.4.2.3.6 Transition Core DNB Methodology 4.4-21a 4.4.2.4 Flux Tilt Considerations 4.4-22 4.4.2.5 Void Fraction Distribution 4.4-22

4. 4. 2. 6 Core Coolant Flow Distribution 4.4-23

4. 4. 2. 7 Core Pressure Drops and Hydraulic Loads 4.4-23 4.4.2.7.1 Core Pressure Drops 4.4-23 4.4.2.7.2 Hydraulic Loads 4.4-24

4. 4. 2. 8 Correlation and Physical Data 4.4-25 4.4.2.8.1 Surface Heat Transfer Coefficients 4.4-25 4.4.2.8.2 Total Core apd Vessel Pressure Drop .4. 4-26 4.4.2.8.3 Void Fraction Correlation 4.4-27

4. 4. 2. 9 Thermal Effects of Operational Transients 4. 4-28 4.4.2.10 Uncertainties in Estimates 4.4-29 4.4.2.10.1 Uncertainties in Fuel and Clad Temperatures 4.4-29 4.4.2.10.2 Uncertainties in Pressure Drops 4.4-30 4.4.2.10.3 Uncertainties Due to Inlet Flow Maldistribution 4.4-30 4.4.2.10.4 Uncertainty in DNB Correlation 4.4-30

• SGS-UFSAR 4-v Revision 18 April 26, 2000

TABLE OF CONTENTS (Cant)

Section Title 4.4.2.10.5 Uncertainties in DNBR Calculations 4.4-30 4.4.2.10.6 Uncertainties in Flow Rates 4.4-31 4.4.2.10.7 Uncertainties in Hydraulic Loads 4.4-31 4.4.2.10.6 Uncertainty in Mixing Coefficients 4.4-32 4.4.2.11 Plant Configuration Data 4.4-32

. 4. 4.3 Evaluation 4.4-33 4.4.3.1 Core Hydraulics 4.4-33 4.4.3.1.1 Flow Paths Considered in Core Pressure Drop and Thermal Design 4.4-33 4.4.3.1.2 Inlet Flow Distributions 4.4-34 4.4.3.1.3 Emprical Friction Factor Correlations 4.4-35 4.4.3.2 Influence of Power Distribution 4.4-36 4.4.3.2.1 Nuclear Enthalpy Rise Hot Channel Factor 4.4-37 4.4.3.2.2 Axial Heat Flux Distributions 4.4-38 4.4.3.3 Core Ther.mal Response 4.4-38 4.4.3.4 Analytical Techniques 4.4-39 4.4.3.4.1 Core Analysis 4.4-39 4.4.3.4.2 Fuel Temperatures 4.4-47 4.4.3.4.3 Hydrodynamdc Instability 4.4-47 4.4.3.5 Hydrodynamic and Flow Power Coupled Instability 4.4-47 4.4.3.6 Temperature Transient Effects Analysis 4.4-50 4.4.3.7 Potentially Damaging Temperature Effects During Transients 4.4-51 4.4.3.8 Energy Release During Fuel Element Burnout 4.4-52 4.4.3.9 Energy Release or Rupture of Water-logged Fuel Elements 4.4-53 4.4.3.10 Fuel Rod Behavior Effects from Coolant Flow Blockage 4.4-53 4.4.4 Testing and Verification 4.4-55 4.4.4.1 Tests Prior to Initial Criticality 4.4-55 SGS-UFSAA 4-vi Revision 6 February 15, 1987

TABLE OF CONTENTS (Cont)

Section Title 4.4.4.2 Initial Power and Plant Operation 4.4-55 4.4.4.3 Component and Fuel Inspections 4.4-56 4.4.4.4 Auq.mented Startup Test Program 4.4-56 4.4.5 References for Section 4.4 4.4-56 4.5 RELOAD ANALYSIS 4.5-1 4.5.1 References for Section 4.5 4.5-2

• SGS-UFSAR 4-vii Revision 17 October 16, 1998

LIST OF TABLE:S 4.1-1 Thermal and Hydraulic Design 4.1-2 Analytic Techniques Incore Design 4.1-3 Design Loading Conditions for Reactor Core Components 4.2-l Maximum Deflections Allowed for Reactor Internal Support Structures 4.2-2 Comparison of Single and Double Encapsulated Secondary Source Designs 4.3-1 Reactor Core Description 4.3-2 Nuclear Design Parameters 4.3-3 Reactivity Requirements for Rod Cluster Control Assemblies

4. 3-4 Axial Stability Index-PWR Core With a 12-Foot Height 4.3-5 Typical Neutron Flux Levels at Full Power

~.3-6 Comparison of Measured and Calculated Doppler Defects 4.3-7 Benchmark Critical Experiments SGS-UFSAR 4-viii Revision 18 April 26, 2000

LIST OF TABLES (Cont) 4.3-8 Saxton Core II Isotopics, Rod MY, Axial Zone 6 4.3-9 Critical Boron Concentrations, BOL 4.3-10 Comparison of Measured and Calculated Rod Worth

4. 3-11 Comparison of Measured and Calculated Moderator Coefficients at HZP, BOL 4.4-1 Reactor Thermal and Hydraulic Design Parameters 4.4-2 (This text has been deleted) 4.4-3 Void Fractions at Nominal Reactor Conditions With Design Hot Channel Factors 4.4-4 Comparison of THINC-IV and THINC-I Predictions With Data From Representative Westinghouse Two and Three Loop Reactors 4.5-1 (This text has been deleted) 4.5-2 (This text has been deleted)

• SGS-UFSAR 4-ix Revision 23 October 17 1 2007

LIST OF FIGURES Figure 4.2-1 Fuel Assembly Cross Section - 17 x 17 4.2-2 Standard Fuel Assembly Outline - 17 x 17 4.2-2A 17 x 17 Vantage+/Vantage SH Fuel Assembly Comparison 4.2-28 17 x 17 Standard Robust Fuel Assembly Outline I 4.2-2C 17 x 17 RFA ZIRLOl'M+2 Outline 4.2-3 Standard Fuel Rod Schematic 4.2-3A 17 x 17 Vantage+/Vantage 5H Fuel Rod Comparison 4.2-38 17 x 17 Standard RFA Fuel Rod Schematic I 4.2-3C 4.2-4 17 x 17 RFA ZIRL0'+2 Fuel Rod Schematic Typical Clad and Pellet Dimensions as a Function of Exposure 4.2-5 Plan View 4.2-6 Representative Fuel Rod Internal Pressure and Linear Power Density for the Lead Burnup Rod as a Function of Time 4.2-7 Top Grid to Nozzle Attachment 4.2-8 Lower Core Support Assembly 4.2-9 Elevation View, Grid to Thimble Attachment 4.2-10 Upper Core Support Structure

4. 2-11 Guide Thimble to Bottom Nozzle Joint 4.2-12 Plan View of Upper Core Support Structure 4.2-13 Full Length Rod Cluster Control and Drive Rod Assembly With Interfacing Components SGS-OFSAR 4-x Revision 19 November 19, 2001

LIST OF FIGURES (Cant)

Title 4.2-14 Full Length Rod Cluster Control Assembly Outline 4.2-15 Full Length Absorber Rod 4.2-16 Burnable Absorber Assembly 4.2-17 Pyrex Burnable Poison Rod Cross Section 4.2-17A WABA Rod Cross Section 4.2-18 Primary Source Assembly 4.2-19 Single Encapsulated Secondary Source 4.2-20 Thimble Plug Assembly (Optional Usage}

4.2-21 Full Length Control Rod Drive Mechanism 4.2-22 Full Length Control Rod Drive Mechanism Schematic 4.2-23 Nominal Latch Clearance at Minimum and Maximum Temperature 4.2-24 Control Rod Drive Mechanism Latch Clearance Thermal Effect 4.2-25 Schematic Representation of Reactor Core Model 4.3-1 Production and Consumption of Higher Isotopes 4.3-2 Boron Concentration vs Cycle Burnup (Typical) 4.3-3 Normalized Power Density Distribution At BOL, Unrodded Core, HFP, No Xenon (Typical) 4.3-4 Normalized Power Density, Distribution Near BOL, Unrodded Core, HFP, Equilibrium Xenon (Typical) 4-xi SGS-UFSAR Revision 18 April 26, 2000

LIST OF FIGURES (Cont)

Figure 4.3-5 Normalized Power Density Distribution Near BOL, Group D Inserted, HFP, Equilibrium Xenon (Typical) 4.3-6 Normalized Power Density Distribution Near Middle of Life, Unrodded Core, HFP, Equilibrium Xenon (Typical) 4.3-7 Normalized Power Density Distribution Near EOL, Unrodded Core, HFP, Equilibrium Xenon (Typical) 4.3-8 Typical Axial Power Shapes Occurring at Beginning of Life 4.3-9 Typical Axial Power Shapes Occurring at Middle of Life 4.3-10 Typical Axial Power Shapes Occurring at End of Life

4. 3-11 Maximum F0 -Power vs Axial Height During Normal Operation 4.3-12 Peak Power Density During Control Rod Malfunction Overpower Transients 4.3-13 Peak Linear Power During Boration/Dilution Overpower Transients

4. 3-14 Comparison Between Calculated and Measured Relative Fuel Assembly Power Distribution 4.3-15 Comparison of Calculated and Measured Axial Shape 4.3-16 Measured Values of FQ for Full Power Rod Configuration 4.3-17 Doppler Temperature Coefficient at BOL and EOL, (Typical) 4.3-18 Doppler Power Coefficient-SOL, MOL, EOL, (Typical) 4.3-19 Doppler Power Defect-BOL, MOL, EOL, (Typical) 4.3-20 Moderator Temperature Coefficient-SOL, ARO (Typical) 4-xii SGS-UFSAR Revision 17 October 16, 1998

LIST OF FIGURES (Cont)

Figure Title 4.3-21 Moderator Temperature Coefficient-EOL, ARO (Typical) 4.3-22 Moderator Temperature Coefficient as a Function of Boron Concentration, BOL, ARO (Typical) 4.3-23 Hot Full Power Moderator Temperature Coefficient vs Cycle Burnup (Typical) 4.3-24 Total Power Coefficient-SOL, EOL (Typical) 4.3-25 Total Power Defect-SOL, EOL (Typical) 4.3-26A Rod Cluster Control Assembly Pattern - Unit 1 4.3-26B Rod Cluster Control Assembly Pattern - Unit 2 4.3-27 Accidental Simultaneous Withdrawal of 2 Control Banks EOL, HZP, Banks D&B Moving in the Same Plane 4.3-28 Design Trip Curve 4.3-29 Normalized Rod Worth vs Percent Insertion All Rods But One 4.3-30 Axial Offset vs Time-PWR Core With a 12-Foot Height and 121 Assemblies 4.3-31 XY Xenon Test Thermocouple Response Quadrant Tilt Difference vs Time 4.3-32 Calculated and Measured Doppler Defect and Coefficients at BOL, Two-Loop Plant, 121 Assemblies, 12-Foot Core 4.3-33 Comparison of Calculated and Measured Boron Concentration for 2-Loop Plant, 121 Assemblies, 12-foot Core 4.3-34 Comparison of Calculated and Measured CB 2-Loop with 121 Assemblies, 12-Foot Core 4-xiii SGS-UFSAR Revision 17 October 16, 1998

LIST OF FIGURES (Cant)

Figure Title 4.3-35 Comparison of Calculated and Measured ce 3-Loop Plant, 157 Assemblies, 12-Foot core 4.4-1 Peak Fuel Average and Surface Temperatures During Fuel Rod Lifetime vs Linear Power 4.4-lA Peak Fuel Average and Surface Temperatures During Fuel Rod Lifetime vs Linear Power for Vantage-5H fuel 4.4-2 Peak Fuel Centerline Temperature During Fuel Rod Lifetime vs Linear Power 4 ., 4-2A Peak Fuel Centerline Temperatures During Fuel Rod Lifetime vs Linear Power for Vantage-SH fuel 4.4-3 Thermal Conductivity of uo (Data Corrected to 95% Theoretical 2

Density)

4. 4-4 Axial Variation of Average Clad Temperature for Rod Operating at

'5.43 kW/ft 4.4-SA Comparison of Measured to Predicted 17 x 17 DNB Data 4.4-58 Measured vs Predicted Critical Heat Flux WRB-1 Correlation 4.4-SC Measured Critical Heat Flux - WRB-2 Correlation 4.4-6 TDC vs Reynold's Number for 26-Inch Grid Spacing 4.4-7 Normalized Radial Flow and Enthalpy Distribution at 4-Foot Elevation 4.4-8 Normalized Radial Flow and Enthalpy Distribution at 8-Foot Elevation SGS-UFSAR 4-xiv Revision 18 April 26, 2000

LIST OF FIGURES {Cont)

Figure 4.4-9 Normalized Radial Flow and Enthalpy Distribution at 12-Foot Elevation (Unit 2) 4.4-10 Void Fraction vs Thermodynamic Quality H-HSAT/Hg-HSAT 4.4-11 PWR Natural Circulation Test 4.4-12 Comparison of a Representative Westinghouse Two-Loop Reactor Incore Thermocouple Measurements With THINC-IV Predictions 4.4-13 Comparison of a Representative Westinghouse Three-Loop Reactor Incore Thermocouple Measurements With THINC-IV Predictions 4.4-14 Hanford Subchannel Temperature Data Comparison With THINC-IV 4.4-15 Hanford Subcritical Temperature Data Comparison With THINC-IV 4.4-16 Distribution of Incore Instrumentation - Unit 1 4.4-17 Distribution of Incore Instrumentation - Unit 2 4.5-1 Typical Salem Unit 1 Loading Pattern 4.5-2 Typical Salem Unit 1 Burnable Absorber Configuration 4.5-3 Typical Salem Unit 2 Loading Pattern 4.5-4 Typical Salem Unit 2 Burnable Absorber Configuration 4-xv SGS-UFSAR Revision 23 October 17, 2007

SECTION 4 REACTOR 4.1

SUMMARY

DESCRIPTION This chapter describes the following: 1) the mechanical components of the reactor and reactor core including the fuel rods and fuel assemblies, reactor internals, and the control rod drive mechanisms, 2) the nuclear design, and 3) the thermal-hydraulic design.

The reactor core is comprised of an array of fuel assemblies which are similar in mechanical design and fuel enrichment. The Salem Unit 1 and 2 cores may consist of any combination of fuel designs including Vantage 5H, Vantage+, and standard Robust Fuel Assembly (RFA and RFA-2, which further enhances the anti-fretting characteristics with improved mid grids) as described in Section 4. 2. 1. 2. The most significant difference between the Vantage+ and RFA fuel and the others is the

. TM application of Zlrlo cladding, guide thimble and instrument tubes. The Vantage+

and RFA are modifications of the NRC-approved Vantage 5H fuel assembly design (Reference 1) A detailed description and evaluation of the Vantage+ and RFA features is provided in References 2, 4, 5 and 6.

The core is cooled and moderated by light water at a pressure of 2250 psia in the Reactor Coolant System. The moderator coolant contains boron as a neutron absorber.

The concentration of boron in the coolant is varied as required to control relatively slow reactivity changes including the effects of fuel burnup. Additional boron, in the form of burnable absorber rods and/ or IFBAs, may be employed in the core to establish the desired initial reactivity.

Two hundred and sixty-four fuel rods are mechanically joined in a square array to form a fuel assembly. The fuel rods are supported in intervals along their length by grid assemblies which maintain the lateral spacing between the rods throughout the design life of the assembly. The grid assembly consists of an "egg-crate" arrangement of interlocked straps. The straps contain spring fingers and dimples for fuel rod support as well as coolant mixing vanes. The fuel rods consist of slightly enriched uranium dioxide ceramic cylindrical pellets contained in slightly cold worked Zircaloy-4 or Zirlo TM tubing which is plugged and seal welded at the ends to 4.1-1 SGS-UFSAR Revision 28 May 22, 2015

encapsulate the fuel. All fuel rods are pressurized with helium during fabrication to reduce stresses and strains and to increase fatigue life. In addition, the Zirlorn fuel rods may be oxide coated at the lower end for additional protection against fretting. RFA fuel rods will utilize annular pellets at the top and bottom 6" to provide lower rod internal pressures.

The center position in the assembly is reserved for the in-core instrumentation, while the remaining 24 positions in the array are equipped with guide thimbles joined to the grids and the top and bottom nozzles. Depending upon the position of the assembly in the core, the guide thimbles are used as core locations for rod cluster control assemblies, neutron source assemblies, and burnable absorber rods. The remaining guide thimbles may be fitted with plugging devices to limit bypass flow.

The use of plugging devices is optional.

The bottom nozzle is a box-like structure which serves as a bottom structural element of the fuel assembly and directs the coolant flow distribution to the assembly.

The top nozzle assembly functions as the upper structural element of the fuel assembly in addition to providing a partial protective housing for the rod cluster control assembly or other core components.

The rod cluster control assemblies each consist of a group of individual absorber rods fastened at the top end to a common hub or spider assembly. These assemblies have rods containing absorber material to control the reactivity of the core under operating conditions.

The control rod drive mechanisms are of the magnetic latch type. The latches are controlled by three magnetic coils. They are so designed that upon a loss of power to the coils, the rod cluster control assembly is released and falls by gravity to shut down the reactor.

The components of the reactor internals are divided into three parts consisting of the lower core support structure (including 4.1-2 SGS-UFSAR Revision 18 April 26, 2000

the entire core barrel and thermal shield), the upper core support structure and the in-core instrumentation support structure. The reactor internals support the core, maintain fuel alignment, limit fuel assembly movement, maintain alignment between the fuel assemblies and control rod drive mechanisms, direct coolant flow past the fuel elements and to the pressure vessel head, provide gamma and neutron shielding, and provide guides for the in-core instrumentation.

The nuclear design analyses and evaluation establish physical locations for control rods and burnable absorbers, and physical parameters such as fuel enrichments and boron concentration in the coolant such that the reactor core has inherent characteristics which, together with corrective actions of the Reactor Control, Protection and Emergency Cooling Systems provide adequate reactivity control even if the highest reactivity worth rod cluster control assembly is stuck in the fully withdrawn position. The design also provides for inherent stability against diametral and azimuthal power oscillations.

The thermal-hydraulic design analyses and evaluation establish coolant flow parameters which assure that adequate heat transfer is assured between the fuel cladding and the reactor coolant. The thermal design takes into account local variations in fuel rod dimensions, power generation, flow distribution, and mixing.

The mixing vanes incorporated in the fuel assembly spacer grid design induces additional flow mixing between the various flow channels within a fuel assembly as well as between adjacent assemblies.

Instrumentation is provided in and out of the core to monitor the nuclear, thermal-hydraulic, and mechanical performance of the reactor and to provide inputs to automatic control functions.

The reactor core design together with corrective actions of the Reactor Control, Protection and Emergency Cooling Systems can meet the reactor performance and safety criteria specified in Section 4.2.

4.1-3 SGS-UFSAR Revision 17 October 16, 1998

To illustrate the effects of the change in fuel design, Table 4.1-1 presents principal nuclear, thermal-hydraulic, and mechanical design parameters for the Salem 17 x 17 Vantage 5H, Vantage+, and RFA fuel assemblies.

The effects of fuel densification were evaluated (Reference 3).

The analytical techniques employed in the core design are tabulated in Table 4.1-2.

The loading conditions considered in general for the core internals and components are tabulated in Table 4.1-3. Specific or limiting loads considered for design purposes of the various components are listed as follows: fuel assemblies in Section 4.2.1.1.2; reactor internals in Section 4.2.2.3 and Table 5.1-10; neutron absorber rods, burnable absorber rods, neutron source rods, and thimble plug assemblies (if used) in Section 4.2.3.1.3; control rod drive mechanisms in Section 4.2.3.1.4.

4.1.1 Reference for Section 4.1

1. Davidson, S.L. (Ed.), et al., "Vantage 5H Fuel Assembly Reference Core Report,"

WCAP-10444-P-A and Appendix A, September 1985; Addendum 2-A, March 1986; Addendum 2-A, April 1988.

2. Davidson, S.L., Nuhfer, D.L. (Eds.), "Vantage+ Fuel Assembly Reference Core Report," WCAP-12610-P-A, April 1995.

3. Hellman, J. M. (Ed.), "Fuel Densification Experimental Results and Model for Reactor Application," WCAP-8218-P-A (Proprietary) and WCAP-8219-A (Nonproprietary), March 1975.

4. Letter from W. J. Rinkacs (Westinghouse) to M.M. Mannion (PSE&G), "Westinghouse Fuel Features Recommendation for Cycle 11", July 22, 1998.

5. Letter from W. J. Rinkacs (Westinghouse) to M.M. Mannion (PSE&G), "Westinghouse Generic Safety Evaluation for the 17x17 Standard Robust Fuel Assembly", October 1, 1998.

6. Garde, A., et al., "Westinghouse Clad Corrosion Model for ZIRLO and Optimized ZIRLO," WCAP-12610-P-A & CENPD-404-P-A, Addendum 2-A, October 2013.

4.1-4 SGS-UFSAR Revision 28 May 22, 2015

TABLE 4.1-1 I

THERMAL AND HYDRAULIC DESIGN Reactor Core Heat Output, MWt 3459 6

Reactor Core Heat Output, 10 Btu/hr 11, 806 Heat Generated in the Fuel, % 97.4 Nominal System Pressure, psia 2250 Assumed Initial System Pressure for DNB Transients, psia 2218 (STDP(l))

2250 (RTDP( 2 ))

Minimum DNBR for Design Transients V-5H( 3 ) 1. 24 (RTDP)

RFA(7) 1. 24 (RTDP, Typ)

RFA 1. 22 (RTDPI Thm)

ONE Correlation WRB-1 WRB-2 Coolant Flow Total Thermal Design Flow Rate, 10 6 lb/hr 125.3(S}

Effective Flow Rate for Heat 6

Transfer, 10 lb/hr 116.3(S}

Effective Flow Area for Heat 2

Transfer, ft V-5H( 3 ) 51.3 RFA 51.1 Average Velocity Along Fuel Rods, ft/sec V-5H( 3 )

RFA 14.1 14.2 I

Average Mass Velocity, 10 6 lb/hr-tt 2 I

V-5H( 3 ) 2.27(S)

RFA 2.28{S)

• SGS-UFSAR Page 1 of 5 Revision 19 November 19, 2001

TABLE 4.1-1 (Continued)

THERMAL AND HYDRAULIC DESIGN Coolant Temperature Nominal Inlet, °F 542.7(S)

Average Rise in Vessel, °F 70.4 ( 5 )

Average Rise in Core, °F Average in Core, °F 582.4(S)

Average in Vessel, °F 577.9(S)

Heat Transfer 2

Active Heat Transfer Surface Area, ft 59,700 Heat Flux Hot Channel Factor, FQ 2.40 Average Heat Flux, Btu/hr-ft 2 192,470 Maximum Heat Flux for Normal Operation, 461,930 2

Btu/hr-ft Average Thermal Output, kW/ft 5.52 Maximum Thermal Output for Normal 13.3 Operation, kW/ft Peak Linear Power for Determination <22.4 of Protection Setpoints, kW/ft Peak Fuel Center Temperature at Maximum Thermal <4700 Output for Maximum Overpower Trip Point, °F Page 2 of 5 SGS-UFSAR Revision 19 November 19, 2001

TABLE 4.1-1 (Continued)

THERMAL AND HYDRAULIC DESIGN Fuel Assemblies Design RCC Canless Number of Fuel Assemblies 193 uo Rods per Assembly 264 2

Rod Pitch, in 0. 4 96 Overall Dimension, in 8.426 X 8.426 Weight of Fuel (as uo ) in Core, lbs STD, VSH, V+ 222,739 2

RFA (9 ) 217 I 565 Weight of Zircaloy in Core, lbs All STD 50913 All VSH, V+ 52541 All RFA 53847 Number of Grids per Assembly STD 8 Inconel VSH 2 Inconel (Top & Bottom) 6 4 (Mid Grids)

V+ 2 Inconel (Top & Bottom) 6 Zirlo' (Mid Grids)

RFA 2 Inconel (Top & Bottom) 1 Inconel (Protective Grid) 6 Zirlo' (Mid Grids) 3 Zirlo' (Intermediate Flow Mixing Grids)

Loading Technique 3 Region Non-uniform rue.:. Rods Number in Core 50,952 Outside Diameter, in 0.374 Ciametral Gap, in 0.0065 Clad Thickness, in 0.0225

lad Material STD, VSH Zircaloy-4 V+,RFA Zirlo' Page 3 of 5 SGS-UFSAR Revision 18 April 26, 2000

TABLE 4.1-1 (Continued)

THERMAL AND HYDRAULIC DESIGN Fuel Pellets Material uo2 Sintered Density, % of Theoretical Diameter, in 95.5 0.3225(lO)

I RFA Annular Pellet I.D., in 0.155{ll)

Length, in STD 0.530 V-5H( 3 ) 0.387 RFA Solid 0.387 RFA Annular 0.462 or 0.500( 12 }

Rod Cluster Control Assemblies Neutron Absorber Ag-In-Cd Cladding Material Type 316L Ionnitride Surface Clad Thickness, in 0.0185 Number of Clusters 53 Number of Absorbers per Cluster 24 Core Structure Core Barrel, ID I OD, in 148.0 I 152.5 Thermal Shield, ID I OD, in 158.5 I 164.0 Nuclear Design Parameters:

Structure Characteristics Core Diameter, in (Equivalent) 132.7 Core Average Active Fuel Height, in 143.7 Page 4 of 5 SGS-UFSAR Revision 20 May 6, 2003

TABLE 4.1-1 (Continued)

THERMAL AND HYDRAULIC DESIGN Reflector Thickness and Composition Top - Water Plus Steel, in ~10 Bottom - Water Plus Steel, in -10 Side - Water Plus Steel, in -15 H20/U, Molecular Ratio, Lattice (cold) 2.41 (1) Standard Thermal Design Procedure.

(2) Revised Thermal Design Procedure.

(3) Also valid for V+ assemblies without Intermediate Flow Mixing Grids.

(4) Deleted (5) For analyses where high average core temperature is bounding.

(6) Deleted (7) With Intermediate Flow Mixing Grids.

(8) Deleted (9) With annular axial blankets.

(10) Applicable to solid or annular pellets.

(11) Top and bottom 6n of RFA fuel stack height.

(12) Starting with Unit 1 Region 17 and Unit 2 Region 15.

Page 5 of 5 SGS-UFSAR Revision 20 May 6, 2003

TABLE 4.1-2 ANALYTIC TECHNIQUES IN CORE DESIGN Section Analysis Technique Computer Code Referenced Mechanical Design of Core Internals Loads, Deflections, and Static and Dynamic Blowdown code, FORCE, Stress Analysis Modeling Finite element structural analysis code, and others Fuel Rod Design Fuel Performance Characteristics Semi-empirical thermal Westinghouse fuel rod 4.2.1.3.1 (temperature, internal pressure, model of fuel rod with design model 4.3.3.1 clad stress, etc.) consideration of fuel 4.4.2.2 density changes, heat 4.4.3.4.2 transfer, fission gas release, etc.

Nuclear Design

1) Cross Sections and Group Microscopic data Modified ENDF/B library 4.3.3.2 Constants Macroscopic constants LEOPARD/CINDER type or 4.3.3.2 for homogenized core PHOENIX-P regions Group constants for HAMMER-AIM or PHOENIX-P 4.3.3.2 control rods with self-shielding PARAGON or NEXUS 4.3.3.2

2) X-Y and X-Y-Z Power 2-Group Diffusion TURTLE (2-D) or 4.3.3.3 Distributions, Fuel Depletion, Theory ANC(2-D or 3-D)

Critical Boron Concentrations, x-y and X-Y-Z Xenon Distributions, Reactivity Coefficients

3) Axial Power Distributions 1-D, 2-Group Diffusion PANDA or APOLLO 4.3.3.3 Control Rod Worths, and Theory Axial Xenon Distribution 1 of 2 SGS-UFSAR Revision 31 December 5, 2019

Analysis

• Technique TABLE 4.1-2 (Cont)

Computer Code Section Referenced

4) Fuel Rod Power Integral Transport Theory LASER 4.3.3.1 Effective Resonance Monte Carlo Weighting REPAD Temperature Function Thermal-Hydraulic Design

1) Steady-state Subchannel analysis of THINC-IV 4.4.3.4.1 local fluid conditions in rod bundles, including inertial and crossflow resistance terms, solu-tion progresses from core-wide to hot assembly to hot channel

2) Transient DNB Analysis Subchannel analysis of THINC-I (THINC-III) 4.4.3.4.1 local fluid conditions in rod bundles during transients by including accumulation terms in conservation equations; solution progresses from core-wide to hot assembly to hot channel SGS-UFSAR 2 of 2 Revision 6 February 15, 1987

TABLE 4.1-3

~ DESIGN LOADING CONDITIONS FOR REACTOR CORE COMPONENTS

1. Fuel Assembly Weight

2. Fuel Assembly Spring Forces

3. Internals Weight

4. Control Rod Scram (equivalent static load)

5. Differential Pressure

6. Spring Preloads

7. Coolant Flow Forces (static)

8. Temperature Gradients

9. Differences in thermal expansion

a. Due to temperature differences

b. Due to expansion of different materials

10. Interference between components

~ 11. Vibration (mechanically or hydraulically induced)

12. One or more loops out of service

13. All operational transients listed in Table 5.1-10.

14. Pump overspeed

15. Seismic loads (operation basis earthquake and design basis earthquake)

16. Blowdown forces (due to cold and hot leg break)

• SGS-UFSAR 1 of 1 Revision 6 February 15, 1987

4.2 MECHANICAL DESIGN The plant conditions for design are divided into four categories in accordance with their anticipated frequency of occurrence and risk to the public:

Condition I - Normal Operation; Condition II - Incidents of Moderate Frequency; Condition III - Infrequent Incidents; Condition IV- Limiting Faults.

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

1. The mechanical design of the reactor core components and their physical arrangement, together with corrective actions of the Reactor Control, Protection, and Emergency Cooling Systems (when applicable) assure that:

a. Fuel damage* is not expected during Condition I and Condition II events. It is not possible, however, to preclude a very small number of rod failures. These are within the capability of the Plant Cleanup System and are consistent with the plant design bases.

b. The reactor can be brought to a safe state following a Condition III event with only a small fraction of fuel rods damaged*

although sufficient fuel damage might occur to preclude resumption of operation without considerable outage time.

c. The reactor can be brought to a safe state and the core can be kept subcritical with acceptable heat transfer geometry following transients arising from Condition IV events.

• Fuel damage as used here is defined as penetration of the fission product barrier (i.e., the fuel rod clad).

4.2-1 SGS-UFSAR Revision 6 February 15, 1987

2. The fuel assemblies are designed to accommodate expected conditions for design for handling during assembly inspection and refueling operations and shipping loads.

3. The fuel assemblies are designed to accept control rod insertions in order to provide the required reactivity control for power operations and reactivity shutdown conditions.

4. All fuel assemblies have provisions for the insertion of in-core instrumentation necessary for plant operation.

5. The reactor internals, in conjunction with the fuel assemblies, direct reactor coolant through the core to achieve acceptable flow distribution and to restrict bypass flow so that the heat transfer performance requirements can be met for all modes of operation. In addition, the internals provide core support and distribute coolant flow to the pressure vessel head so that the temperature differences between the vessel flange and head do not result in leakage from the flange during the Condition I and II modes of operation. Required inservice inspection can be carried out as the internals are removable and provide access to the inside of the pressure vessel.

4.2.1 Fuel 4.2.1.1 Design Bases The fuel rod and fuel assembly design bases are established to satisfy the general performance and safety criteria presented in Section 4.2 and specific criteria noted below. The same design bases apply to the 17 x 17 standard (STD), 17 x 17 Vantage 5H, 17 x 17 Vantage+ and 17 x 17 Standard Robust Fuel Assembly (RFA) designs.

4.2-2 SGS-UFSAR Revision 18 April 26, 2000

4.2.1.1.1 Fuel Rods The integrity of the fuel rods is ensured by designing to prevent excessive fuel temperatures, excessive internal rod gas pressures due to fission gas releases, and excessive cladding stresses and strains. This is achieved by designing the fuel rods so that the following conservative design bases are satisfied during Condition I and Condition II events over the fuel lifetime:

1. Fuel Pellet Temperatures The center temperature of the hottest pellet is to be below the melting temperature of the uo2 (melting point of 5080°F(l) unirradiated and reducing by 58°F per 10,000 MWD/MTU). While a limited amount of center melting can be tolerated, the design conservatively precludes center melting. A calculated centerline fuel temperature of 4700°F has been selected as an overpower limit to assure no fuel melting. This provides sufficient margin for uncertainties, as described in Sections 4.4.1.2 and 4.4.2.10.1.

2. Internal Gas Pressure - The internal gas pressure of the lead rod in the reactor will be limited to a value below that which would cause (a) the diametral gap to increase due to outward clad creep during steady-state operation, and (b) extensive departure from nucleate boiling (DNB) propagation to occur.

3. Clad Stress - The effective clad stresses are less than that which would cause general yield of the clad. While the clad has some capability for accommodating plastic strain, the yield strength has been accepted as a conservative design basis.

4. Clad Tensile Strain - The clad tangential strain range is less than one percent. The clad strain design basis addresses slow transient strain rate mechanisms where the clad effective stress never reaches the yield 4.2-3 SGS-UFSAR Revision 6 February 15, 1987

strength due to stress relaxation. The 1 percent strain limit has been established based upon tensile and burst test data from irradiated clad. Irradiated clad properties are appropriate due to irradiation effects on clad ductility occurring before strain-limiting fuel clad interaction during a transient event can occur.

5. Strain Fatigue - The cumulative strain fatigue cycles are less than the design strain fatigue life. This basis is consistent with proven practice.

Radial, tangential, and axial stress components due to pressure differential and fuel clad contact pressure are combined into an effective stress using the maximum-distortion-energy theory. The von Mises' criterion is used to evaluate if the yield strength has been exceeded. The von Mises' criterion states that an isotropic material under multiaxial stress will begin to yield plastically when the effective stress (i.e., combined stress using maximum-distortion-energy theory) becomes equal to the material yield stress in simple tension as determined by an uniaxial tensile test. Since general yielding is to be prohibited, the volume average effective stress determined by integrating across the clad thickness increased by an allowance for local nonuniformi ty effects before it is compared to the yield strength. The yield strength correlation is that appropriate for irradiated clad since the irradiated properties are attained at low exposure whereas the fuel/clad interaction conditions which can lead to minimum margin to the design basis limit always occurs at much higher exposure.

The detailed fuel rod design established such parameters as pellet size and density, clad-pellet diametral gap, gas plenum size, 4.2-4 SGS-UFSAR Revision 17 October 16, 1998

and helium pressure. The design also considers effects such as fuel density changes, fission gas release, clad creep, and other physical properties which vary with burnup.

Irradiation testing and fuel operational experience has verified the adequacy of the fuel performance and design bases. This experience and testing are discussed in References 2 and 3. Fuel experience and testing results, as they become available, are used to improve fuel rod design and manufacturing processes and assure that the design bases and safety criteria are satisfied.

The safety evaluation of the fuel rod internal pressure design basis is presented in Reference 4.

4.2.1.1.2 Fuel Assembly Structure Structural integrity of the fuel assemblies is assured by setting limits on stresses and deformations due to various loads and by determining that the assemblies do not interfere with the functioning of other components. Three types of loads are considered.

1. Nonoperational loads such as those due to shipping and handling

2. Normal and abnormal loads which are defined for Conditions I and II

3. Abnormal loads which are defined for Conditions III and IV.

These criteria are applied to the design and evaluation of the top and bottom nozzles, the guide thimbles, the grids, and the thimble joints.

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

4.2-5 SGS-UFSAR Revision 25 October 26, 2010

1. Nonoperational 4g axial and 6g lateral loading with dimensional stability.

2. Normal Operation (Condition I) and Incidents Moderate Frequency (Condition II).

For the normal operating (Condition I) and upset conditions (Condition I I) , the fuel assembly component structural design criteria are classified into two material categories, namely, austenitic steels and Zircaloy. The stress categories and strength theory presented in the ASME Boiler and Pressure Vessel Code, Section III, are used as a general guide. The maximum shear-theory (Tresca criterion) for combined stresses is used to determine the stress intensities for the austenitic steel components. The stress intensity is defined as the numerically largest difference between the various principal stresses in a three-dimensional field. The allowable stress intensity value for austenitic steels, such as nickel-chromium-iron alloys, is given by the lowest of the following:

a. 1/3 of the specific minimum tensile strength or 2/3 of the specified minimum yield strength at room temperature

b. 1/3 of the tensile strength or 90 percent of the yield strength at temperature but not to exceed 2/3 of the specified minimum yield strength at room temperature.

4.2-6 SGS-UFSAR Revision 11 July 22, 1991

The stress limits for the austenitic steel components follow:

Stress Intensity Limits Categories Limit General Primary Membrane Stress Intensity Sm Local Primary Membrane Stress Intensity 1.5 Sm Primary Membrane plus Bending Stress Intensity 1.5 Sm Total Primary plus Secondary Stress Intensity 3.0 Sm TM The Zircaloy and ZIRLO structural components which consist of guide thimble and fuel tubes are in turn subdivided into two categories because of material differences and functional requirements. The fuel tube design criteria are covered separately in Section 4.2.1.1.1. The maximum stress theory is used to evaluate the guide thimble design. The maximum stress theory assumes that yielding due to combined stresses occur where one of the principal stresses are equal to the simple tensile or compressive yield stress. The Zircaloy and TM ZIRLO unirradiated properties are used to define the stress limits.

Abnormal loads during Conditions III and IV worst case represented by combined seismic and blowdown loads.

1. Deflections of components cannot interfere with the reactor shutdown or emergency cooling of the fuel rods.

2. The fuel assembly component stresses under faulted conditions are evaluated using primarily the methods outlined in Appendix F of the ASME Pressure Vessel Code Section 3. Since the current analytical methods utilize elastic analysis, the stress allowables are defined as the smaller value of 2.4 Sm or 0.70 Su for primary membrane and 3.6 Sm or 1.05 Su for primary membrane plus primary bending. For the austenitic steel fuel assembly components, the stress intensity is defined in accordance with the rules described in the previous section TM for normal operating conditions. For the Zircaloy and ZIRLO components the stress limits are set at two-thirds of the material yield strength, Sy, at reactor operating temperature. This results in Zircaloy stress intensity limits being the smaller of 1.6 Sy or 0.70 Su for primary membrane and 2. 4 Sy or 1. 05 Su for primary membrane plus TM bending. For conservative purposes, the Zircaloy and ZIRLO unirradiated properties are used to define the stress limits. The grid component strength criteria are based on experimental tests. The grid component strength criterion is based on the lower 95 percent confidence level on the true mean from distribution of grid crush strength data at temperature.

4.2-7 SGS-UFSAR Revision 21 December 6, 2004

4.2.1.2 Design Description Fuel assembly and fuel rod design data are given in Tables 4.1-1 and 4.3-1. Two hundred sixty-four fuel rods, twenty-four guide thimble tubes, and one instrumentation thimble tube are arranged within a supporting structure to form a fuel assembly. The instrumentation thimble is located in the center position and provides a channel for insertion of an in-core neutron detector if the fuel assembly is located in an instrument core position. The guide thimbles provide channels for insertion of either a rod cluster control assembly, a neutron source assembly, a burnable absorber assembly or a plugging device (if used),

depending on the position of the particular fuel assembly in the core.

Figure 4.2-1 shows a cross section of a fuel assembly array, and Figure 4.2-2 shows a standard fuel assembly full length view. The fuel rods are loaded into the fuel assembly structure so that there is clearance between the fuel rod ends and the top and bottom nozzles.

The design changes from the 17 x 17 STD design to the Vantage 5H design include reduced guide thimble and instrumentation tube diameters, and replacement of the six intermediate (mixing vane) Inconel grids with Zircaloy grids. The debris filter bottom nozzle (DFBN) design has been incorporated into the Vantage 5H fuel assembly. The DFBN is similar to the standard bottom nozzle design except that it is thinner and has a new pattern of smaller flow holes.

The DFBN helps to minimize passage of debris particles that could cause fretting damage to fuel rod cladding.

The Vantage+ assembly skeleton is identical to that previously described for Vantage 5H except for those modifications necessary to accommodate intended fuel operation to higher burnup levels. The Vantage+ assembly skeleton is made of low cobalt material and the spring height is slightly increased for the reduction in fuel assembly height. The modifications consist of the use of TM ZIRLO guide thimbles as necessary and small skeleton dimensional alterations to provide additional fuel assembly and rod growth space at the extended burnup levels. The Vantage+ fuel assembly is 0.200 inch shorter than the Vantage 5H assembly. The grid centerline elevations of the Vantage+ are identical to those of the Vantage 5H assembly, except for the top grid. The Vantage+ top grid has been moved down by the same 0.200 inch. The RFA assembly provides further design enhancements from the Vantage+ design including modified Low Pressure Drop mid-grids and IFMs, thicker guide tubes, and a protective bottom TM grid. The RFA continues to utilize the ZIRLO fuel rods and skeletons.

4.2-8 SGS-UFSAR Revision 21 December 6, 2004

However, since the Vantage+ and RFA fuel are intended to replace either the Westinghouse LOPAR or Vantage 5H, their assembly exterior envelope is equivalent in design dimensions, and the functional interface with the reactor internals is also equivalent to those of previous Westinghouse fuel assembly designs. Also, the Vantage+ and RFA are designed to be mechanically and hydraulically compatible with the LOPAR and Vantage 5H, and the same functional requirements and design criteria as previously established for the Westinghouse Vantage 5H fuel assembly remain valid for the Vantage+ and RFA (References 18, 21). The fuel rod clad oxidation and hydriding design criteria for ZIRLO cladding has been updated in Reference 27. The Vantage+ and RFA design parameters are provided in Table 4.1-1. Figure 4.2-2A compares the Vantage+ and the Vantage 5H fuel assembly designs. The RFA design is shown in Figure 4. 2-2B.

The feed fuel loaded into Salem Unit 1 (starting with Regions 17A & 17B) and Salem Unit 2 (starting with Regions 14A & 14B) contains further enhancements to the RFA fuel assembly design in order to reduce performance limitations associated with rod internal pressure. The design changes from the 17 x 17 RFA design to the 17 x 17 RFA Z IRL0'+2 design include: an increased fuel rod length, increased fuel assembly length and elimination of the external grip top end plug. The RFA ZIRL0'+2 fuel assembly is 0. 2 inches longer than the RFA assembly due to an increase in thimble and guide tube lengths. The small dimensional alterations to the skeleton provide rod growth space for the longer fuel rods. The grid centerline elevations of the RFA ZIRL0'+2 design are identical to those of the RFA fuel assembly, except for the top grid. The RFA ZIRL0'+2 top grid has been moved upward by 0. 2 inches. The top nozzle spring height is slightly decreased due to the increase in overall assembly height. The RFA ZIRL0'+2 assembly design is shown in Figure 4.2-2C.

The feed fuel loaded into Salem Unit 1 (starting with Regions 19A and 19B) and Salem Unit 2 (starting with Regions 17A and 17B) contains further enhancements to the RFA mid grid design in order to further reduce grid to rod fretting.

The design changes from the RFA mid grid to the RFA-2 mid grid are changes to the spring window cut-outs and the spring and dimple contact areas. Compared to RFA mid grid, RFA-2 mid grid offers improved resistance to fuel rod fretting without significantly affecting any other thermal-hydraulic or mechanical TM performance. Both RFA and RFA-2 mid grids are made from ZIRLO material and both mid grids have the same mass and volume.

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 the upper ends of the fuel assemblies. The upper core plate then bears downward against the fuel assembly top nozzle via the holddown springs to hold the fuel assemblies in place.

4.2.1.2.1 Fuel Rods The Vantage+ and RFA fuel rod designs represent a modification to the Vantage

. TM TM 5H fuel rod in claddlng of ZIRLO as compared to Zircaloy-4. ZIRLO is a zirconium alloy similar to Zircaloy-4, which has been specifically developed to enhance corrosion resistance. The Vantage 5H, Vantage+, and RFA fuel rods contain enriched uranium dioxide fuel pellets, and Integral Fuel Burnable Absorber ( IFBA) coating on some of the fuel pellets. Additionally, the RFA fuel rods contain axial blankets (top and bottom 6 inches) of annular pellets.

The annulus represents a 25% void by volume of the pellet. Schematics of the fuel rods are shown in Figures 4.2-3, 4.2-3A, and 4.2-3B.

4.2-9 SGS-UFSAR Revision 28 May 22, 2015

The Vantage+ and RFA fuel rods have the same wall thickness as the Vantage 5H. The Vantage+ fuel rod length is shorter to provide the required fuel rod growth room. To offset the reduction in plenum length the Vantage+ fuel rod has a variable pitch plenum spring. The variable pitch plenum spring provides the same support as the Vantage 5H plenum spring, but with less spring turns which means less spring volume. The RFA fuel rod is longer than the Vantage+ and Vantage 5H designs but maintains sufficient room for fuel rod growth. This provides additional plenum length to accommodate fission gas release associated with high burnup. In addition, the RFA utilizes the variable pitch plenum spring and axial blankets of annular pellets to provide additional plenum margin.

The 6 inches (top and bottom) of annular pellets effectively provide 3 additional inches of plenum volume beyond the current VANTAGE+ design. The bottom end plug has an internal grip feature to facilitate rod loading on the Vantage+, RFA, and Vantage 5H designs and provides appropriate lead-in for the removable top nozzle reconstitution feature. The RFA design also incorporates the external grip top end plug into its fuel rod design. This feature provides additional capability to reposition the fuel during manufacturing and will simplify reconstitution of the fuel rod from the top of the assembly. The Salem Vantage+ and RFA fuel rods also may posses a zirconium dioxide (Zr0 ) coating on the 2

bottom outside surface of the fuel rod. This fuel feature may be incorporated as an option to be determined on a reload basis. The protective oxide coating covers the bottom end plug, the bottom end plug weldment and a portion of the cladding. A metallurgical-bonded layer of zro uniformly covers a minimum of 4.5 inches of the bottom end of 2

the fuel rod. The minimum coating length was chosen to ensure that the coating would extend through the top of the current bottom Inconel structural grid, independent of the fuel rod loading position or fuel assembly design. The coating, which is 2 to 6 microns thick, provides a hard, wear resistant, surface layer of ZrO for additional debris damage resistance, thereby improving fuel 2

reliability. This extra layer of oxide coating provides additional rod fretting wear protection. The RFA design incorporates the protective bottom grid and modified fuel rod bottom end plug. These enhanced debris mitigating features, described in Section 4.2.1.2.2, diminish the need for the oxide coating. Therefore, for core designs utilizing the RFA, the oxide coating may be incorporated as an option to be determined on a reload basis.

The RFA ZIRL0'+2 fuel rod design represents a modification to the RFA fuel rod in terms of increased fuel rod length and increased fuel assembly length. The RFA ZIRL0'+2 fuel rod length is 0. 2 inches longer than the RFA fuel rod length. Sufficient fuel rod growth margin is accomplished by increasing the length of the instrument thimble and guide thimbles. The RFA ZIRL0'+2 design incorporates a shorter (0.12 inches) external top end plug without a gripper feature. There is no longer any functional purpose for this top end plug gripper feature as fuel rods can still be handled through the top of the assembly for reconstitution. Both the longer fuel rod and shorter external top end plug provide additional plenum length to accommodate fission gas release. The plenum spring free length was increased to accommodate the increase in fuel rod plenum length, while the spring rate was decreased to maintain the same fuel stack hold-down force as in the RFA fuel rod design. A schematic of the RFA ZIRL0'+2 fuel rod is shown in Figure 4.2-3C.

The Salem Vantage 5H, Vantage+, and RFA fuel uses a standardized fuel pellet design which is a refinement to the chamfered pellet design. The standard design helps to improve manufacturability while maintaining or improving performance (e.g., improved pellet chip resistance during manufacturing and handling) .

The Vantage 5H, Vantage+, and RFA IFBA coated fuel pellets are identical to the enriched uranium dioxide pellets except for the addition of a thin boride 4.2-10 SGS-UFSAR Revision 21 December 6, 2004

coating on the pellet cylindrical surface. Coated pellets occupy the central portion of the fuel column. The number and pattern of IFBA rods within an assembly vary depending on specific application. The ends of the coated and uncoated pellets are dished to allow for greater axial expansion at the pellet centerline and void volume for fission gas release. To accommodate a six inch axial blanket length at the top and bottom of the Robust Fuel Assembly (RFA) fuel rods, the annular pellets have a longer length than the solid pellets to obtain an integer number of fuel pellets which will equal a six inch length.

To avoid overstressing of the cladding or seal welds, void volume and clearances are provided within the rods to accommodate fission gases released from the fuel, differential thermal expansion between the cladding and the fuel, and fuel density changes during burnup. Shifting the fuel within the cladding during handling or shipping prior to core loading is prevented by a stainless steel helical spring which bears on top of the fuel. At assembly the pellets are stacked 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 process in order to minimize compressive clad stresses and creep due to coolant operating pressures. Fuel rod pressurization is dependent on the planned fuel burnup as well as other fuel design parameters and fuel characteristics (particularly densification potential).

4.2.1.2.2 Fuel Assembly Structure The fuel assembly structure consists of a bottom nozzle, top nozzle, guide thimbles, and grids, as shown on Figures 4.2-2 and 4.2-2A.

Bottom Nozzle The bottom nozzle is a box-like structure which serves as a bottom structural element of the fuel assembly and directs the coolant flow distribution to the assembly. The square nozzle is fabricated from Type 304 stainless steel and consists of a perforated plate and four angle legs with bearing plates as shown on Figure 4.2-2. The legs form a plenum for the inlet coolant flow to the fuel assembly. The plate itself acts to prevent a downward ejection of the fuel rods from the fuel assembly. The bottom nozzle is fastened to the fuel assembly guide tubes by locked screws which penetrate through the nozzle and mate with an inside fitting in each guide tube.

The debris filter bottom nozzle (DFBN) design was introduced into the Salem fuel assemblies to help reduce the possibility of fuel rod damage due to debris-induced fretting. The Vantage+ and RFA assemblies have a low cobalt 4.2-11 SGS-UFSAR Revision 28 May 22, 2015

stainless steel DFBN. The DFBN design incorporates a modified flow hole size and pattern, as described below, and a decreased nozzle height and thinner top plate to accommodate the high burnup fuel rods. The DFBN retains the design reconstitution feature which facilitates easy removal of the nozzle from the fuel assembly.

The relatively large flow holes in a conventional bottom nozzle were replaced with a new pattern of smaller flow holes in the DFBN. The holes are sized to minimize passage of debris particles large enough to cause damage. The holes were also sized to provide sufficient flow area, comparable pressure drops, and continued structural integrity of the nozzle. Tests to measure pressure drop and demonstrate structural integrity have been performed to verify that the DFBN is totally compatible with the current design.

Salem Unit 2 Region 23 and Salem Unit 1 Region 26 are the first regions to implement the Standardized Debris Filter Bottom Nozzle (SDFBN) design into Salem RFA fuel assemblies. The SDFBN is designed to have a loss coefficient that is the same, independent of supplier. The SDFBN eliminates the side skirt communication flow holes as a means of improving the debris mitigation performance of the bottom nozzle. This nozzle meets all of the applicable mechanical and thermal-hydraulic design criteria. The RFA fuel assembly with the SDFBN is illustrated in Figure 4.2-2C.

Coolant flow through the fuel assembly is directed from the plenum in the bottom nozzle upward through the penetrations in the plate to the channels between the fuel rods . The penetrations in the plate are positioned between the rows of the fuel rods.

Axial loads (holddown) imposed on the fuel assembly and the weight of the fuel assembly is transmitted through the bottom nozzle to the lower core plate.

Indexing and positioning of the fuel assembly is controlled by alignment holes in two diagonally opposite bearing plates which mate with locating pins in the lower core plate. Any lateral loads on the fuel assembly are transmitted to the lower core plate through the locating pins.

Top Nozzle The top nozzle assembly functions as the upper structural element of the fuel assembly in addition to providing a partial protective housing for the rod cluster control assembly or other components. It consists of an adapter plate, enclosure, top plate, and pads. The integral welded assembly has holddown springs mounted on the assembly as shown on Figure 4.2-2. The reconstitutable top nozzle (RTN) design contains hold-down springs and screws made of Inconel 718 and Inconel 600, respectively, whereas other components are made of Type 304 stainless steel. The feed fuel loaded into Salem Unit 2 (starting with Regions 14A and 14B) and Salem Unit 1 (starting with Regions 17A and 17B) contains recons ti tutable top nozzles with hold-down screws made of Inconel 718, as opposed to Inconel 600, in order to increase resistance to stress corrosion cracking.

Vantage+, RFA, and Vantage 5H fuel assemblies use the reconstitutable top nozzle (RTN). The RTN design for the Vantage 5H fuel assembly differs from the conventional design in two ways: 1) a groove is provided in each thimble thru-4.2-12 SGS-UFSAR Revision 28 May 22, 2015

hole in the nozzle plate to facilitate attachment and removal, and; 2) the nozzle plate thickness is reduced to provide additional axial space for fuel rod growth. Additional details of this design feature, the design bases and evaluation of the reconstitutable top nozzle are given in Section 2.3.2 in Reference 15.

The square adapter plate is provided with round and obround penetrations to permit the flow of coolant upward through the top nozzle. Other round holes are provided to accept sleeves which are welded to the adapter plate and mechanically attached to the thimble tubes. The ligaments in the plate cover the tops of the fuel rods and prevent their upward ejection from the fuel assembly. The enclosure is a sheet metal shroud which sets the distance between the adapter plate and the top plate. The top plate has a large square hole in the center to permit access for the control rods and the control rod spiders. Holddown springs are mounted on the top plate and are fastened in place by screws and clamps located at two diagonally opposite corners. The clamps are attached to the nozzle by a specific arrangement of tack welds or tack weld(s) in combination with a stainless steel clamp screw, depending on the manufacturing process in place at the time a given fuel region was built.

The spring screws apply a load directly to the base of the hold-down springs.

The clamps do not have any bearing surfaces that load the spring to the nozzle, but primarily provide a stationary location for attachment of lock wires that prevent rotation of the spring screws. On the other two corners, integral pads are positioned which contain alignment holes for locating the upper end of the fuel assembly.

Salem Units 1 and 2 later implemented the Westinghouse Integral Nozzle (WIN) design in RFA-2 fuel assemblies. The WIN design, while similar to the RTN, incorporates design and manufacturing improvements to eliminate the Inconel 718 spring screw for attachment of the holddown springs. In the WIN nozzle, the springs are assembled into the nozzle pad and pinned in place. The WIN design provides a wedged rather than a clamped (bolted) joint to transfer the fuel assembly holddown forces into the top nozzle structure.

A replacement reconstitutable top nozzle (RRTN) design may be used in a reload cycle to replace the original reconstitutable top nozzle (RTN) or the WIN on an irradiated fuel assembly. The mechanical features of the RRTN are the same as those for the RTN (see Figure 4.2-2) or the WIN with some minor dimensional differences in the top nozzle adapter plate thimble hole to facilitate attachment to an irradiated fuel assembly. The RRTN design contains hold-down springs and screws made of Inconel 718, whereas, other components are made of Type 304 stainless steel.

Guide and Instrument Thimbles The guide thimbles are structural members which also provide channels for the neutron absorber rods, burnable poison rods, or neutron source assemblies. Each TM one is fabricated from Zircaloy-4 or ZIRLO tubing having two different diameters. The larger diameter at the top provides a relatively large annular area to permit rapid insertion of the control rods during a reactor trip as well as to accommodate the flow of coolant during normal operation. Four holes are provided on the thimble tube above the dashpot to reduce the rod drop time.

The lower portion of the guide thimbles has a reduced diameter to produce a dashpot action near the end of the control rod travel during normal operation and to accommodate the outflow of water from the dashpot during a reactor trip.

The dashpot is closed at the bottom by means of an end plug which is provided with a small flow port to avoid fluid stagnation in the dashpot volume during normal operation. The top end of the guide thimble is fastened to a tubular insert by three expansion swages. The insert engages into the top nozzle and is secured into position by the lock tube. The lower end of the guide thimble is fitted with an end plug which is then fastened into the bottom nozzle by a locked screw.

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Fuel rod support grids are fastened to the guide thimble assemblies to create an integrated structure. Since welding of the Inconel grid and Zircaloy thimble is not possible, the fastening technique depicted on Figures 4.2-5 and 4.2-9 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 TM thimble tube to the elevation of the zircaloy sleeves that have been welded to the Zircaloy middle grid assemblies. The four-lobed tool forces the thimble and sleeve outward to a predetermined diameter, thus joining the two components.

The top grid-to-thimble attachment for the Vantage 5H, Vantage+, and RFA design is shown on Figure 4.2-7. The Zircaloy or ZIRLOTM thimbles are fastened to the top nozzle inserts by expanding the members as shown on Figure 4. 2-7. The inserts then engage the top nozzle and are secured into position by the insertion of lock tubes.

The bottom grid assembly is joined to the fuel assembly as shown on Figure 4.2-

11. The stainless steel insert is spot welded to the bottom grid and later captured between the guide thimble end plug and the bottom nozzle by means of a stainless steel thimble screw.

The described methods of grid fastening are standard and have been used successfully since the introduction of Zircaloy guide thimbles in 1969.

The central instrumentation thimble of each fuel assembly is constrained by seating in counterbores in each nozzle. This tube is a constant diameter and guides the incore neutron detectors. This thimble is expanded at the top and mid grids in the same manner as the previously discussed expansion of the guide thimbles to the grids.

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With the exception of an increased length above the dashpot, the Vantage+ guide thimbles are identical to those in the Vantage 5H design. For the RFA, the thimble tube thickness has been increased by 25% ( 4 Mils) relative to the V5H and Vantage+ designs. The thimble tube outer diameter of both the upper spans (major diameter section above the dashpot) and the dashpot spans have been increased. The inner diameter of the thimble tube in both the upper spans and in the dashpot spans are unchanged from the current designs. With the thicker guide thimble tube, the cross-sectional area is increased by 26% in the major diameter section and 29% in the dashpot section. The Vantage+, RFA and Vantage 5H guide thimble ID provides adequate clearance for the control rods and sufficient diametral clearance for burnable absorber rods and source rods.

The Vantage+, RFA and Vantage 5H instrumentation tube diameter has sufficient diametral clearance for the flux thimble to traverse the tube without binding.

Grid Assemblies The fuel rods, as shown on Figures 4. 2-2, 4. 2-2A and 4. 2-2B, are supported laterally at intervals along their length by grid assemblies which maintain the lateral spacing between the rods throughout the design life of the assembly.

Each fuel rod is afforded lateral support at six contact points within each grid by the combination of support dimples and springs. The grid assembly consists of individual slotted straps interlocked and welded in an "egg-crate" arrangement to join the straps permanently at their points of intersection.

The straps contain spring fingers, support dimples, and mixing vanes.

The magnitude of the grid restraining force on the fuel rod is set high enough to minimize possible fretting, without overstressing the cladding at the points of contact between the grids and fuel rods. The grid assemblies also allow axial thermal expansion of the fuel rods without imposing restraint sufficient to develop buckling or distortion of the fuel rods.

Up to four types of grid types are used in each fuel assembly: Mid-grids (structural grids with flow mixing vanes), Intermediate Flow Mixing (IFM) grids (non-structural grids with flow mixing vanes), top and bottom structural grids without mixing vanes, and the protective bottom grid ( P-grid) . Table 4.1-1 provides the breakdown of the grid types, number and grid material used in each of the fuel designs. Flow mixing vanes project from the edge of the inner grid strap into the coolant stream to promote mixing of the coolant in the high heat flux regions of the fuel assembly.

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The IFMs are positioned at the mid-spans of the four uppermost mid-grids to further increase the flow turbulence in the axial zone where departure from nucleate boiling (DNB) is limiting. Each IFM grid cell contains four dimples, which are designed to prevent mid-span channel closure in the spans containing IFMs and fuel contact with the mixing vanes.

For the RFA, the modified low pressure drop mid-grids and IFM grids are embossed to accept the larger diameter guide thimble tube.

The P-grid is a partial height grid similar in configuration to the mid-grid, but without mixing vanes. It is located between the bottom Inconel grid and the bottom nozzle, nearly on the surface of the bottom nozzle. The intersections of the inner straps of the P-grid align with the flow holes of the DFBN, effectively bisecting the flow path through the flow hole into four quarters. This provides an effective barrier against small debris. In conjunction with the P-grid, the fuel rod bottom end plug is changed to a longer design such that the portion of the fuel rod engaged in the P-grid and extending up past the top of the P-grid is solid end plug material. This provides a protective zone where trapped debris cannot fret through the fuel rod and cause a failure. The hydraulic effects of the P-grid are minimized by positioning the fuel rods 0.085 inches above the bottom nozzle. The combination of the lowered fuel rod position and longer fuel rod end plug results in no change to the axial fuel stack height from the previous Vantage+

fuel region.

For the application of the P-grid, the bottom Inconel grid was welded to 20 of the 24 inserts. The remaining four inserts are spot-welded to the P-grid at the four outermost corners on the grid diagonal.

Salem Unit 2 Region 23 and Salem Unit 1 Region 26 are the first regions to implement the Westinghouse Robust Protective Grid (RPG) which was developed as a result of observed failures of the P-grid in the field during Post Irradiation Exams (PIE) performed at several different plants. It was determined by Westinghouse that observed failures were the result of two primary issues; 1) fatigue failure within the protective grid itself at the top of the end strap and 2) stress corrosion cracking ( SCC) primarily within the rod support dimples. The RPG implements design changes such as increasing the maximum nominal height of the grid, increasing the amount of material at the ends of the dimple window cutouts, increasing the radii of the dimple window cutouts, and the welding of four additional inserts for a total of 8 welded inserts out of the 24 total inserts in order to help better support the grid.

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The nominal height of the grid was increased to allow "V-notch" window cutouts to be added to help minimize flow-induced vibration caused by vortex shedding at the trailing edge of the inner grid straps. Figure 4.2-2C shows RFA fuel with the increased nominal height due to implementation of the RPG. These design changes incorporated into the RPG design help address the issues of fatigue failures and failures due to stress corrosion cracking. These changes do not impact the thermal hydraulic performance of the RPG as there is no change to the pressure loss coefficient. In addition, the RPG retains the original protective grid function as a debris mitigation feature.

The outside straps on all the grids contain mixing vanes which, in addition to their mixing function, aid in guiding the grids and fuel assemblies past projecting surfaces during handling or during loading and unloading of the core.

During 1989, snag-resistant grids were introduced. These grids contain outer grid straps which are modified to help prevent assembly hangup from grid strap interference during fuel assembly removal. This was accomplished by changing the grid strap corner geometry and the addition of guide tabs on the outer grid strap.

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4.2.1.3 Design Evaluation 4.2.1.3.1 Fuel Rods The fuel rods are designed to assure the design bases are satisfied for Condition I and II events. This assures that the fuel performed, and safety criteria (Section 4.2.1.1) are satisfied.

Materials - Fuel Cladding The desired fuel rod cladding is a material which has a superior combination of neutron economy (low absorption cross section), high strength (to resist deformation due to differential pressures and mechanical interaction between fuel and clad), high corrosion resistance (to coolant, fuel, and fission products) , and high reliability. Zircaloy-4 and ZIRLOTM have this desired combination of cladding properties. There is considerable pressurized water reactor (PWR) operating experience on the capability of Zircaloy as a cladding material (2). Clad hydriding has not been a significant cause of clad perforation since current controls on fuel-contained moisture levels were instituted.

Metallographic examination of irradiated commercial fuel rods have shown occurrences of fuel/ clad chemical interaction. Reaction layers of <1 mil in thickness have been observed between fuel and clad at limited points around the circumference. Westinghouse metallographic data indicates that this interface layer remains very thin even at high burnup. Thus, there is no indication of propagation of the layer and eventual clad penetration.

Stress corrosion cracking is another postulated phenomenon related to fuel/clad chemical interaction. Out-of-reactor tests have shown that in the presence of high clad tensile stress, relatively large concentrations of iodine, or cadmium in solution in liquid cesium can stress corrode zirconium alloy tubing and lead to eventual clad cracking. Extensive post irradiation examination has 4.2-14 SGS-UFSAR Revision 17 October 16, 1998

produced no conclusive evidence that this mechanism is operative in commercial fuel. Creep collapse and creep-down, along with the associated irradiation stability of cladding, have been evaluated using the models described in references 7 and 28. It has been established that the design basis of no clad collapse during planned core life can be satisfied by limiting fuel densification and by having a sufficiently high initial internal rod pressure.

Materials - Fuel Pellets Sintered, high density uranium dioxide fuel is chemically inert, with respect to the cladding, at core operating temperatures and pressures. In the event of cladding defects, the high resistance of uranium dioxide to attack by water protects against fuel deterioration although limited fuel erosion can occur. As has been shown by operating experience and extensive experimental work, the thermal design parameters conservatively account for changes in the thermal performance of the fuel elements due to pellet fracture which may occur during power operation. The consequences of defects in the cladding are greatly reduced by the ability of uranium dioxide to retain fission products including those which are gaseous or highly volatile.

Observations from several operating Westinghouse PWRs (2) have shown that fuel pellets can densify under irradiation to a density higher than the manufactured values. Fuel densification and subsequent incomplete settling of the fuel pellets results in local and distributed gaps in the fuel rods. Fuel densification has been minimized by improvements in the fuel manufacturing process and by specifying a nominal 95.5 percent initial fuel density.

The effects of fuel densification have been taken into account in the nuclear and thermal-hydraulic design of the reactor described in Sections 4.3 and 4.4, respectively.

Materials - Strength Considerations One of the most important limiting factors in fuel element duty is the mechanical interaction of fuel and cladding. This fuel- cladding interaction produces cyclic stresses and strains in the cladding, and these in turn consume cladding fatigue life. The 4.2-15 SGS-UFSAR Revision 30 May 11, 2018

reduction of fuel-cladding interaction is therefore a principal goal of design.

In order to achieve this goal and to enhance the cyclic operational capability of the fuel rod, the technology for using prepressurized fuel rods in Westinghouse PWRs has been developed.

Initially the gap between the fuel and cladding is sufficient to prevent hard contact between the two. However, during power operation a gradual compressive creep of the cladding onto the fuel pellet occurs due to the external pressure exerted on the rod by the coolant. Cladding compressive creep eventually results in hard fuel-cladding contact. During this period of fuel-cladding contact, changes in power level could result in significant changes in cladding stresses and strains. By using prepressurized 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 delays the time at which substantial fuel-cladding interaction and hard contact occur and hence significantly reduces the number and extent of cyclic stresses and strains experienced by the cladding both before and after fuel-cladding contact. These factors result in an increase in the fatigue life margin of the cladding and lead to greater cladding reliability. If gaps should form in the fuel stacks, clad flattening will be prevented by the rod prepressurization so that the flattening time will be greater than the fuel core life.

Steady-State Performance Evaluation In the calculation of the steady-state performance of a nuclear fuel rod, the following interacting factors must be considered:

1. Clad creep and elastic deflection

2. Pellet density changes, thermal expansion, gas release, and thermal properties as a function of temperature and fuel burnup 4.2-16 SGS-UFSAR Revision 6 February 15, 1987

3. Internal pressure as a function of fission gas release, rod geometry, and temperature distribution These effects are evaluated using an overall fuel rod design model (Reference

17) which include appropriate modifications for time dependent fuel densification. With these interacting factors considered, the model determines the fuel rod performance characteristics for a given rod geometry, power history, and axial power shape. In particular, internal gas pressure, fuel and cladding temperatures, and cladding deflections are calculated. The fuel rod is divided lengthwise into several sections and radially into a number of annular zones. Fuel density changes, cladding stresses, strains and deformations, and fission gas releases are calculated separately for each segment. The effects are integrated to obtain the internal rod pressure. The initial rod internal pressure is selected to delay fuel/clad mechanical interaction and to avoid the potential for flattened rod formation. Clad flattening for Salem Nuclear Generating Station (SNGS) fuel is evaluated using the models described in Reference 7.

The gap conductance between the pellet surface and the cladding inner diameter is calculated as a function of the composition, temperature, and pressure of the gas mixture, and the gap size or contact pressure between clad and pellet.

After computing the fuel temperature of each pellet annular zone, the fractional fission gas release is assessed using an empirical model derived from experimental data (17). The total amount of gas released is based on the average fractional release within each axial and radial zone and the gas generation rate which, in turn, is a function of burnup. Finally, the gas released is summed over all zones and the pressure is calculated.

The model shows good agreement in fit for a variety of published and proprietary data on fission gas release, fuel temperatures, and clad deflections ( 17) . Included in this spectrum are variations in power, time, fuel density, and geometry. The in-pile fuel 4.2-17 SGS-UFSAR Revision 22 May 5, 2006

Temperature measurements' comparisons used are shown in Reference 17.

Typical fuel clad inner diameter and the fuel pellet outer diameter as a function of exposure are presented on Figure 4.2-4. The cycle-to-cycle changes in the pellet outer diameter represent the effects of power changes as the fuel is moved into different positions as a result of refueling. The gap size at any time is merely the difference between clad inner diameter and pellet outer diameter. Total clad-pellet surface contact occurs near the end of Cycle 2.

The figure represents hot fuel dimensions for a fuel rod operating at the power level shown on Figure 4.2-6. Figure 4.2-6 illustrates representative fuel rod internal gas pressure and linear power for the lead burnup rod vs. irradiation time. In addition, it outlines the typical operating range of internal gas pressures which is applicable to the total fuel rod population within a region.

The "best estimate" fission gas release model was used in determining the internal gas pressures as a function of irradiation time.

The clad 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 internal gas pressure. Because of the prepressurization with helium, the volume average effective stresses are always less than -10,000 psi at the pressurization level used in this fuel rod design. Stresses due to the temperature gradient are not included in this average effective stress because thermal stresses are, in general, negative at the clad inner diameter and positive at the clad outer diameter and their contribution to the clad volume average stress is small. Furthermore, the thermal stress decreases with time during steady-state operation due to stress relaxation. The stress due to pressure differential is highest in the minimum power rod at the beginning of life (BOL) (due to low internal gas pressure) and the thermal stress is highest in the maximum power rod (due to steep temperature gradient).

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Tensile stresses could be created once the clad has come in contact with the pellet. These stresses would be induced by the fuel pellet swelling during irradiation. As shown on Figure 4. 2-4, there is very limited clad pushout after pellet-clad contact. Fuel swelling can result in small clad strains

(< 1 percent) for expected discharge burnups, but the associated clad stresses are very low because of clad creep (thermal and irradiation-induced creep).

Furthermore, the 1 percent strain criterion is extremely conservative for fuel-swelling driven clad strain because the strain rate associated with solid 7 1 fission products swelling is very slow (-5 x 10- hr- ) In-pile experiments

( 8) have shown that Zircaloy tubing exhibits "superplastici ty" at slow strain rates during neutron irradiation. Uniform clad strains of greater than 10 percent have been achieved under these conditions with no sign of plastic instability.

Transient Evaluation Method Pellet thermal expansion due to power increases is considered the only mechanism by which significant stresses and strains can be imposed on the clad.

Power increases in commercial reactors can result from fuel shuffling (e.g.,

Region 3 positioned near the center of the core for Cycle 2 operation after operating near the periphery during Cycle 1), reactor power escalation following extended reduced power operation, and control rod movement. In 4.2-19 SGS-UFSAR Revision 11 July 22, 1991

the mechanical design model, lead rods are depleted using best estimate power histories as determined by core physics calculations. During the depletion the amount of diametral gap closure is evaluated based upon the pellet expansion-cracking model, clad creep model, and fuel swelling model. At various times during the depletion, the power is increased locally on the rod to the burnup dependent attainable power density as determined by core physics calculations.

The radial, tangential, and axial clad stresses resulting from the power increase are combined into a volume average effective clad stress.

The von Mises' criterion is used to evaluate if the clad yield stress has been exceeded. This criterion states that an isotropic material in multi-axial stress will begin to yield plastically when the effective stress exceeds the yield stress as determined by a uniaxial tensile test. The yield stress correlation is that for irradiated cladding since fuel/clad interaction occurs at high burnup. Furthermore, the effective stress is increased by an allowance, which accounts for stress concentrations in the clad adjacent to radial cracks in the pellet, prior to the comparison with the yield stress.

This allowance was evaluated using a two-dimensional (r, e) finite element model.

Slow transient power increases can result in large clad strains without exceeding the clad yield stress because of clad creep and stress relaxation.

Therefore, in addition to the yield stress criterion, a criterion on allowable clad positive strain is necessary. Based upon high strain rate burst and tensile test data on irradiated tubing, 1 percent strain was determined to be the lower limit on irradiated clad ductility and thus adopted as a design criterion.

In addition to the mechanical design models and design criteria, Westinghouse relies on performance data accumulated through transient power test programs in experimental and commercial reactors, and through normal operation in commercial reactors.

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It is recognized that a possible limitation to the satisfactory behavior of the fuel rods in a reactor which is subjected to daily load follow is the failure of the cladding by low cycle strain fatigue. During their normal residence time in reactor, the fuel rods may be subjected to -1000 cycles with typical changes in power level from 50 to 100 percent of their steady-state values.

The assessment of the fatigue life of the fuel rod cladding is subjected to a considerable uncertainty due to the difficulty of evaluating the strain range which results from the cyclic interaction of the fuel pellets and claddings.

This difficulty arises, for example, from such highly unpredictable phenomena as pellet cracking, fragmentation, and relocation. Nevertheless, since early 1968, Westinghouse has been investigating this particular phenomenon both analytically and experimentally. Strain fatigue tests on irradiated and nonirradiated hydrided Zircaloy-4 claddings were performed which permitted a definition of a conservative fatigue life limit and recommendation of a methodology to treat the strain fatigue evaluation of the Westinghouse reference fuel rod designs.

However, Westinghouse is convinced that the final proof of the adequacy of a given fuel rod design to meet the load follow requirements can only come from in-pile experiments performed on actual reactors. The Westinghouse experience in load follow operation dates back to early 1970 with the load follow operation of the Saxton reactor. Successful load follow operation has been performed on reactor A (300 load follow cycles) and reactor B (150 load follow cycles) . In both cases, there was no significant coolant activity increase that could be associated with the load follow mode of operation.

The following paragraphs present briefly the Westinghouse analytical approach to strain fatigue.

A comprehensive review of the available strain-fatigue models was conducted by Westinghouse as early as 1968. This included the 4.2-21 SGS-UFSAR Revision 6 February 15, 1987

Langer-O'Donnell model (9), the Yao-Munse model, and the Manson-Halford Model.

Upon completion of this review and using the results of the Westinghouse experimental programs discussed below, it was concluded that the approach defined by Langer-0' Donnell would be retained and the empirical factors of their correlation modified in order to conservatively bound the results of the Westinghouse testing program.

The Langer-O'Donnell empirical correlation has the following form:

E 100 1n +

4.jN; 100 RA where:

s 1/2 E ~ s pseudo - stress amplitude which causes a t

.2 failure in Nf cycles (lb/ln )

~s total strain range (in/in) t

.2 E Young's Modulus (lb/ln )

Nf number of cycles to failure RA reduction in area at fracture in a uniaxial tensile test (percent) 2 S endurance limit (lb/in )

e Both RA and S are empirical constants which depend on the type of material, e

the temperature, and the irradiation. The Westinghouse testing program was subdivided in the following subprograms:

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

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Both hydrided and nonhydrided Zircaloy-4 cladding were tested.

2. A biaxial fatigue experiment in gas autoclave on unirradiated Zircaloy-4 cladding both hydrided and nonhydrided.

3. A fatigue test program on irradiated cladding from the CVTR and Yankee Core V conducted at Battelle Memorial Institute.

The results of these test programs provided information of different cladding conditions including the effect of irradiation, hydrogen level, and temperature.

The Westinghouse design equations followed the concept for the fatigue design criterion according to Section 3 of the ASME Boiler and Pressure Vessel code; namely:

1. The calculated pseudo-stress amplitude (S ) has to be multiplied by a

a factor of 2 in order to obtain the allowable number of cycles (Nf

) .

2. The allowable cycles for a given Sa is 5 percent of Nf, or a safety factor of 20 on cycles.

The lesser of the two allowable number of cycles is selected. The cumulative fatigue life fraction is then computed as:

k

<1 1

where: nk number of diurnal cycles of mode k.

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The potential effects of operation with waterlogged fuel are discussed in Section 4.4.3.6. Waterlogging is not considered to be a concern during operational transients.

4.2.1.3.2 Fuel Assembly Structure Stresses and Deflections The potential sources of high stresses in the assembly are avoided by the design. For example, stresses in the fuel rod due to thermal expansion and TM Zircaloy or ZIRLO irradiation growth are limited by the relative motion of the rod as it slips over the grid spring and dimple surfaces. Clearances TM between the fuel rod ends and nozzles are provided so that Zircaloy or ZIRLO irradiation growth will not result in end interferences. As another example, stresses due to holddown springs in opposition to the hydraulic lift force are limited by the deflection characteristic of the springs. Stresses in the fuel assembly caused by tripping of the rod cluster control assembly have little influence on fatigue because of the small number of events during the life of an assembly. Welded joints in the fuel assembly structure are considered in the structural analysis of the assembly. Appropriate material properties of welds are used to ensure the design bases are met. 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.

The fuel assembly design loads for shipping have been established at 4 g axial and 6g lateral. Probes are permanently placed into the shipping cask to monitor and detect fuel assembly displacements that would result from loads in excess of the criteria. Past history and experience have indicated that loads which exceeded 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 that the shipping 4.2-24 SGS-UFSAR Revision 17 October 16, 1998

design limits do not result in impairment of fuel assembly function.

Dimensional Stability The Vantage 5H Mechanical Test Program description and results are given in Reference 16 and are considered to be applicable to Vantage+ as the two assemblies are structurally essentially identical.

The development of the RFA design included a comprehensive set of mechanical and hydraulic tests: pressure drop, assembly vibration, fuel rod vibration, bulge joint strength, grid crush. A description of these tests and results is provided in References 19, 20 and 21.

The coolant flow channels are established and maintained by the structure composed of grids and guide thimbles. 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 assured by the clamping force provided by the grid springs. Lateral motion of the fuel rods is opposed by the spring force and the internal moments generated between the spring and the support dimples. Grid testing is discussed in Reference 10.

No interference with control rod insertion into thimble tubes will occur during a postulated loss-of-coolant accident (LOCA) transient due to fuel rod swelling, thermal expansion, or bowing. In the early phase of the transient following the coolant break, the high axial loads which potentially could be generated by the difference in thermal expansion between fuel clad and thimbles are relieved by slippage of the fuel rods through the grids. The relatively low drag force restraint on the fuel rods will only induce minor thermal bowing not sufficient to close the fuel rod-to-thimble tube gap. This rod-to-grid slip mechanism occurs simultaneously with control rod drop.

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Vibration and Wear The effect of the flow induced vibration on the V5H and Vantage+ 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.

The conclusion that the effect of flow induced vibrations on the fuel assembly and fuel rod is minimal is based on test results and analysis documented in Reference 11.

Full flow vibration tests have been performed on the RFA covering both assembly and fuel rod vibration. These tests have shown the Robust Fuel Design is not susceptible to flow induced assembly vibration and provides improved fuel rod vibration performance over the prior designs. In addition, the RFA-2 mid grid design provides further enhancements that improve fuel rod vibration performance over the RFA design.

The reaction on the grid support due to vibration motions is also correspondingly small and much less than the spring preload. Firm contact is therefore maintained. No significant wear of the cladding or grid supports is expected during the life of the fuel assembly, based on out-of-pile flow tests, performance of similarly designed fuel in operating reactors (2), and design analyses.

Clad fretting and fuel rod vibration have been experimentally investigated as shown in Reference 11.

No significant guide thimble tube wear due to flow-induced vibration of the control rods is predicted. Based on a conservative wear analyses, Westinghouse concluded that the integrity of the guide tube is maintained during normal operation, accident conditions, and nonoperational loading condition for at least 250 weeks (> 3 cycles) of fuel assembly operation. The Nuclear Regulatory Commission (NRC) has concluded ( 12) that the Westinghouse analyses probably accounts for all the major variables in the wear process. However, the NRC requested additional confirmatory information supporting the absence of significant thimble wear (no wear hole formation) for the 17 x 17 fuel assembly design. Examination of 1434 guide thimble tubes in six fuel assemblies examined at Salem Unit 1 shows no wear hole formation. Four of the assemblies had control rods in the parked 4.2-26 SGS-UFSAR Revision 25 October 26, 2010

position (7 1/2 inches into the guide thimble) for Cycle 1, and two assemblies had control rods parked for Cycles 1 and 2. The parked position of the control rods has the greatest potential for causing guide thimble wear due to flow induced vibrations. The results of this surveillance program satisfy the NRC request to verify the wear analysis conclusion of no wear holes.

Evaluation of the Reactor Core for Limited Displacement RPV Inlet and Outlet Nozzle Breaks The STD fuel assembly response resulting from the most limiting main coolant pipe break was analyzed using time history numerical techniques. Since the resulting vessel motion induces primarily lateral loads on the reactor core, a finite element model similar to the seismic model described in Reference 10 was used to assess the fuel assembly deflections and impact forces.

The reactor core finite element model which simulates the fuel assembly interaction during lateral excitation consists of fuel assemblies arranged in a planer array with inter-assembly gaps. For the Salem Station, 15 fuel assemblies which correspond to the maximum number of assemblies across the core diameter were used in the mode. The fuel assemblies and the reactor baffle support are represented by single beam elements as shown on Figure 4.2-25. The time history motion for the upper and lower core plates and the barrel at the upper core plate elevation are simultaneously applied to the simulated reactor core model as illustrated on Figure 4. 2-25. The three time history motions were obtained from the time history analysis of the reactor vessel and internals finite element model.

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The fuel assembly response, namely the displacements and grid impact forces, were obtained from the reactor core model using the core plate and barrel motions resulting from a reactor coolant pump outlet double ended break. The maximum fuel assembly deflection was determined to occur in a peripheral fuel assembly. The fuel assembly stresses resulting from this deflection indicated significant safety margins compared to the allowable values. The grid maximum impact force for both the seismic and lateral blowdown accident conditions occurred at the peripheral fuel assembly locations adjacent to the baffle wall.

The grid impact forces were appreciably lower for fuel assembly locations inward from the peripheral fuel. For the lateral blowdown case, only a small (outer) portion of the core experienced significant grid impact forces.

The maximum grid impact force obtained from the limiting rupture break was found to be less than the minimum grid strength (using the 95 x 95 value as determined by tests at reactor operating safe shutdown temperatures) . The maximum square-root-of-the-sum-of-the-squares combination of the pipe rupture and safe shutdown earthquake loads for the limiting grid location was found to be less than the minimum grid strength.

The major components that determine the structural integrity of the fuel assembly are the grids. Mechanical testing and analysis of the Vantage 5H Zircaloy grid and fuel assembly have demonstrated that the Vantage 5H structural integrity under seismic/LOCA loads will provide margins comparable to the STD 17 x 17 fuel assembly design and will meet all design bases.

Since the Vantage+ assembly is structurally similar to that of Vantage 5H, the seismic and LOCA analysis for the Vantage 5H assembly are applicable to Vantage+ assembly. The use of ZIRLOTM guide thimbles will not affect the seismic and LOCA loads.

A Salem plant specific seismic and LOCA analysis was performed for the RFA (Reference 22). The results of this analysis demonstrate that the RFA and V5H fuel assembly designs are capable of maintaining a coolable core geometry and control rod insertability under the combined seismic and LOCA loading for both a homogeneous and mixed core of fuel assembly designs.

4.2.1.3.3 Operational Experience Westinghouse has had considerable experience with Zircaloy-clad fuel since its introduction in the Jose Cabrera plant in June 1968. This experience is extensively described in Reference 2.

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4.2.1.3.4 Test Rod and Test Assembly Experience.

This experience is presented in Sections 8 and 23 of Reference 3.

4.2.1.4 Testing and Inspection Plan 4.2.1.4.1 Quality Assurance Program The quality assurance program plan of the Westinghouse Nuclear Fuel Division for Salem is summarized in Reference 13.

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The program provides for control over all activities affecting product quality, commencing with design and development, and continuing through procurement, 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 affecting product quality through a formal auditing program.

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

4.2.1.4.2 Quality Control Quality control (QC) philosophy is generally based on the following inspections being performed to a 95 percent confidence that at least 95 percent of the product meets specification, unless otherwise noted.

Fuel System Components and Parts The characteristics inspected depend upon the component parts; the QC program includes dimensional and visual examinations, check audits of test reports, material certification, and nondestructive examination, such as X-ray and ultrasonic.

Pellets Inspection is performed for dimensional characteristics such as diameter, density, length, and squareness of ends. Additional visual inspections are performed for cracks, chips, and surface conditions according to approved standards.

Density is determined in terms 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.

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Rod Inspection The fuel rod inspection consists of the following nondestructive examination techniques and methods, as applicable:

1. Each rod is leak tested using a calibrated mass spectrometer, with helium being the detectable gas.

2. Rod welds are inspected by ultrasonic test or X-ray in accordance with a qualified technique and Westinghouse specifications.

3. All rods are dimensionally inspected prior to final release. The requirements include such items as length, camber, and visual appearance.

4. All fuel rods are inspected by gamma scanning or other approved methods to ensure proper plenum dimensions.

5. All fuel rods are inspected by gamma scanning or other approved methods to ensure that no significant gaps exist between pellets.

6. All fuel rods are active gamma scanned to verify enrichment control prior to acceptance for assembly loading.

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

Assemblies Each fuel assembly is inspected for compliance with drawing and/or specification requirements. Other incore control component inspection and specification requirements are given in Section 4.2.3.4.

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Other Inspections The following inspections are performed as part of the routine inspection operation:

1. Tool and gage inspection and control, including standardization to primary 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 performed of inspection activities and records to ensure that prescribed methods are followed and that records are correct and properly maintained.

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

Process Control To prevent the possibility of mixing enrichments during fuel manufacture and assembly, strict enrichment segregation and other process controls are exercised.

The uranium dioxide powder is kept in sealed containers. The contents are fully identified both by descriptive tagging and preselected color coding. A Westinghouse identification tag completely describing the contents is affixed to the containers before transfer to powder storage. Isotopic content is confirmed by analysis.

Powder withdrawal from storage can be made by only one authorized group, which directs the powder to the correct pellet production line. All pellet production lines are physically separated from each other and pellets of only a single nominal enrichment are produced in a given production line at any given time.

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Finished pellets are placed on trays identified with the same color code as the powder containers and transferred to segregated storage racks within the confines of the pelleting area. Samples from each pellet lot are tested for isotopic content and impurity levels prior to acceptance by QC. Physical barriers prevent mixing of pellets of different nominal densities and enrichments in this storage area. Unused powder and substandard pellets are returned to storage in the original color-coded containers.

Loading of pellets into the clad is performed in isolated production lines, and again only one enrichment is loaded on a line at a time.

A serialized traceability code is placed on each fuel tube to provide unique identification. The end plugs are inserted and then inert-welded to seal the tube. The fuel tube remains coded and traceability identified until just prior to installation in the fuel assembly.

At the time of installation into an assembly, the traceability codes are removed and a matrix is generated to identify each rod in its positions within a given assembly. The top nozzle is inscribed with a permanent identification number providing traceability to the fuel contained in the assembly.

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

4.2.1.4.3 Onsite Inspection Surveillance of fuel and reactor performance is routinely conducted on Westinghouse reactors. Power distribution is monitored using the ex-core fixed and in-core movable detectors, and the BEACON (Best Estimate Analyzer for Core Operations Nuclear) on-line core monitoring system. BEACON is also known as the Power Distribution Monitoring System (PDMS) in the Technical Specifications. Coolant activity and chemistry is followed which permits early detection of any fuel clad defects.

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Visual fuel inspection is routinely conducted during refueling. Additional fuel inspections are dependent on the results of the operational monitoring and the visual inspections.

4.2.2 Reactor Vessel Internals 4.2.2.1 Design Bases The design bases for the mechanical design of the reactor vessel internals components are as follows:

1. The reactor internals, in conjunction with the fuel assemblies, shall direct reactor coolant through the core to achieve acceptable flow distribution and to restrict bypass flow so that the heat transfer performance requirements are met for all modes of operation. In addition, required cooling for the pressure vessel head shall be provided so that the temperature differences between the vessel flange and head do not result in leakage from the flange during reactor operation.

2. In addition to neutron shielding provided by the reactor coolant, a separate thermal shield is provided to limit the exposure of the pressure vessel in order to maintain the required ductility of the material for all modes of operation.

3. Provisions shall be made for installing in-core instrumentation useful for the plant operation and vessel material test specimens required for a pressure vessel irradiation surveillance program.

4. The core internals are designed to withstand mechanical loads arising from operating basis earthquake (OBE), design basis earthquake (DBE), and pipe ruptures and meet the requirement of Item 5 below.

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5. The reactor shall have mechanical provisions which are sufficient to adequately support the core and internals and to assure that the core is intact with acceptable heat transfer geometry following transients arising from abnormal operating conditions.

6. Following the design basis accident (DBA), the plant shall be capable of being shut down and cooled in an orderly fashion so that fuel cladding temperature is kept within specified limits. This implies that the deformation of certain critical reactor internals must be kept sufficiently small to allow core cooling.

The functional limitations for the core structures during the DBA are shown in Table 4.2-1. To ensure column loading of rod cluster control guide tubes, the upper core plate deflection is limited to not exceed the value shown in Table 4.2-1.

4.2.2.2 Description and Drawings The reactor vessel internals are described as follows:

The components of the reactor internals consist of the lower core support structure (including the entire core barrel and thermal shield), the upper core support structure, and the in-core instrumentation support structure. The reactor internals support the core, maintain fuel alignment, limit fuel assembly movement, maintain alignment between fuel assemblies and control rod drive mechanisms, direct coolant flow past the fuel elements, direct coolant flow to the pressure vessel head, provide gamma and neutron shielding, and guides for the in-core instrumentation. The coolant flows from the vessel inlet nozzles down the annulus between the core barrel and the vessel wall and then into a plenum at the bottom of the vessel. It then reverses and flows up through the core support and through the lower core plate. Flow passages in the lower core plate are sized to provide the desired inlet flow distribution to the core. After passing through the 4.2-34 SGS-UFSAR Revision 6 February 15, 1987

core, the coolant enters the region of the upper support structure and then flows radially to the core barrel outlet nozzles and directly through the vessel outlet nozzles. A small portion of the coolant flows between the baffle plates and the core barrel to provide additional cooling of the barrel.

Similarly, a small amount of the entering flow is directed into the vessel head plenum and exits through the vessel outlet nozzles.

All the major material for the reactor internals is Type 304 stainless steel.

Parts not fabricated from Type 304 stainless steel include bolts and dowel pins which are fabricated from Type 316 stainless steel and the radial support clevis inserts and bolts which are fabricated of Inconel. The only stainless steel materials used in the reactor core support structures which have yield strengths greater than 90,000 pounds are the 403 series used for holddown springs. The use of these materials is compatible with the reactor coolant and is acceptable based on the 1971 ASME Boiler and Pressure Vessel Code, Case Number 1337.

All reactor internals are removable from the vessel for the purpose of their inspection as well as the inspection of the vessel internal surface.

Lower Core Support Structure The major containment and support member of the reactor internals is the lower core support structure, shown on Figure 4.2-8. This support structure assembly consists of the core barrel, the core baffle, the lower core plate and support columns, the thermal shield, and the core support which is welded to the core barrel. All the major material for this structure is Type 304 stainless steel.

The lower core support structure is supported at its upper flange from a ledge in the reactor vessel and its lower end is restrained from transverse motion by a radial support system attached to the vessel wall. Within the core barrel are an axial baffle and a lower core plate, both of which are attached to the core barrel wall and form the enclosure periphery of the core.

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The lower core support structure and core barrel serve to provide passageways and direct the coolant flow. The lower core plate is positioned at the bottom level of the core below the baffle plates and provides support and orientation for the fuel assemblies.

The lower core plate is a member through which the necessary flow distribution holes for each fuel assembly are machined. Fuel assembly locating pins (two for each assembly) are also inserted into this plate. Columns are placed between this plate and the core support of the core barrel in order to provide stiffness and to transmit the core load to the core support. Adequate coolant distribution is obtained through the use of the lower core plate and core support.

The one-piece thermal shield is fixed to the core barrel at the top with rigid bolted connections. The bottom of the thermal shield is connected to the core barrel by means of axial flexures. This bottom support allows for differential axial growth of the shield/core barrel but restricts radial or horizontal movement of the bottom of the shield. Rectangular specimen guides in which material samples can be inserted and irradiated during reactor operation are welded to the thermal shield and extended to the top of the thermal shield.

These samples are held in the rectangular specimen guides by a preloaded spring device at the top and bottom.

Vertically downward loads from weight, fuel assembly preload, control rod dynamic loading, hydraulic loads, and earthquake acceleration are carried by the lower core plate into the lower core plate support flange on the core barrel shell and through the lower support columns to the core support and thence through the core barrel shell to the core barrel flange supported by the vessel flange. Transverse loads from earthquake acceleration, coolant cross flow, and vibration are carried by the core barrel shell and distributed between the lower radial support to the vessel wall, and to the vessel flange.

Transverse loads of the fuel assemblies are transmitted to the core barrel shell by direct 4.2-36 SGS-UFSAR Revision 6 February 15, 1987

connection of the lower core plate to the barrel wall and by upper core plate alignment pins which are welded into the core barrel.

The radial support system of the core barrel is accomplished by "key" and "keyway" joints to the reactor vessel wall. At six equally spaced points around the circumference, an Inconel clevis block is welded to the vessel inner diameter. An Inconel insert block is bolted to each of these clevis blocks, and has a "keyway" geometry. Opposite each of these is a "key" which is welded to the lower core support. At assembly, as the internals are lowered into the vessel, the keys engage the keyways in the axial direction. With this design, the internals are provided with a support at the furthest extremity, and may be viewed as a beam fixed at the top and simply supported at the bottom.

Radial and axial expansion of the core barrel are accommodated, but transverse movement of the core barrel is restricted by this design. With this system, cyclic stresses in the internal structures are within the ASME Section III limits. In the event of an abnormal downward vertical displacement of the internals following a hypothetical failure, energy absorbing devices limit the displacement of the core after contacting the vessel bottom head. The load is then transferred through the energy absorbing devices of the lower internals to the vessel.

The energy absorbers are mounted on a base plate which is contoured on its bottom surface to the reactor vessel bottom internal geometry. Their number and design are determined so as to limit the stresses imposed on all components except the energy absorber to less than yield (ASME Code Section III valves).

Assuming a downward vertical displacement, potential energy of the system is absorbed mostly by the strain energy of the energy absorbing devices.

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Upper Core Support Assembly The upper core support assembly, shown on Figures 4.2-10 and 4.2-12, consists of the upper support assembly and the upper core plate between which are contained support columns and guide tube assemblies. The support columns establish the spacing between the upper support assembly and the upper core plate and are fastened at the top and bottom to these plates. The support columns transmit the mechanical loadings between the upper support and upper core plate.

The guide tube assemblies shield and guide the control rod drive shafts and control rods. They are fastened to the upper support and are guided by pins in the upper core plate for proper orientation and support. Additional guidance for the control rod drive shafts is provided by the upper guide tube which is attached to the upper support.

The upper core support assembly, which is removed as a unit during refueling operation, is positioned in its proper orientation with respect to the lower support structure by slots in the upper core plate which engage flat-sided upper core plate alignment pins which are welded into the core barrel. At an elevation in the core barrel where the upper core plate is positioned, the flat-sided pins are located at angular positions of 90 degrees from each other.

As the upper support structure is lowered into the lower internals, the slots in the plate engage the flat-sided pins in the axial direction. Lateral displacement of the plate and of the upper support assembly is restricted by this design. Fuel assembly locating pins protrude from the bottom of the upper core plate and engage the fuel assemblies as the upper assembly is lowered into place. Proper alignment of the lower core support structure, the upper core support assembly, the fuel assemblies and control rods are thereby assured by this system of locating pins and guidance arrangement. The upper core support assembly is restrained from any axial movements by a large circumferential spring which rests between the upper barrel flange and the upper 4.2-38 SGS-UFSAR Revision 6 February 15, 1987

core support assembly. The spring is compressed when the reactor vessel head is installed on the pressure vessel.

Vertical loads from weight, earthquake acceleration, hydraulic loads, and fuel assembly preload are transmitted through the upper core plate via the support columns to the upper support assembly and then into the reactor vessel head.

Transverse loads from coolant cross flow, earthquake acceleration, and possible vibrations are distributed by the support columns to the upper support and upper core plate. The upper support plate is particularly stiff to minimize deflection.

In-Core Instrumentation Support Structures All bottom-mounted in-core instrumentation support structures consist of a system to convey and support flux thimbles penetrating the vessel through the bottom (Figure 7. 7-6 shows the Basic Flux-Mapping System) . Specifically, the flux thimbles enter the reactor vessel through the bottom penetration nozzles.

Conduits extend from the bottom of the reactor vessel down through the concrete shield area and up to a thimble seal line. The minimum bend radii are about 144 inches and the trailing ends of the thimbles (at the seal line) are extracted approximately 15 feet 4.2-39 SGS-UFSAR Revision 11 July 22, 1991

during refueling of the reactor in order to avoid interference within the core.

The thimbles are closed at the leading ends and serve as the pressure barrier between the reactor pressurized water and the containment atmosphere.

Mechanical seals between the retractable thimbles and conduits are provided at the seal line. During normal operation, the retractable thimbles are stationary and move only during refueling or for maintenance, at which time a space of approximately 15 feet above the seal line is cleared for the retraction operation.

The in-core instrumentation support structure is designed for adequate support of instrumentation during reactor operation and is rugged enough to resist damage or distortion under the conditions imposed by handling during the refueling sequence. These are the only conditions which affect the in-core instrumentation support structure. Reactor vessel surveillance specimen capsules are covered in Section 4.5.1.

As part of the conversion to an all bottom mounted in-core instrumentation, the support structures for the original top entry core exit thermocouple system have been removed. Specifically, the five thermocouple columns on the upper support plate have been removed and their corresponding reactor vessel head penetrations have been cut and capped. However, the support bases for the columns were left on the upper support plate so as not to create any flow openings across the upper support plate.

The core exit thermocouple system went from a top entry system to a bottom entry system for the following reasons as stated in letter NPE-85-1035:

1. Resistance readings were taken on the thermocouples and many were found to have shorts or were declared inoperable. These would provide erroneous indication at normal operating conditions.

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2. Thermocouple columns were bent/damaged several times during refueling operations and not all thermocouples were recoverable. The installation of bottom-mounted thermocouples provided an easier method of reactor disassembly and lower probability of damage.

3. Installation of a bottom entry system would also allow the elimination of the five instrument ports and resolve the numerous problems encountered with the qualification of the reference junction boxes and thermocouples on the top-mounted system.

4.2.2.3 Design Loading Conditions The design loading conditions that provide the basis for the design of the reactor internals are:

1. Fuel Assembly Weight

2. Fuel Assembly Spring Forces

3. Internals Weight

4. Control Rod Scram (equivalent static load)

5. Differential Pressure

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7. Coolant Flow Forces (static)

8. Temperature Gradients

9. Differences in Thermal Expansion

a. Due to temperature differences

b. Due to expansion of different materials

10. Interference Between Components

11. Vibration (mechanically or hydraulically induced)

12. One or More Loops Out of Service

13. All Operational Transients Listed in Table 4.1-10

14. Pump Overspeed

15. Seismic Loads (OBE and DBE)

16. Blowdown Forces (due to cold and hot leg break)

Combined seismic and blowdown forces are included in the stress analysis as a design loading condition by assuming the maximum amplitude of each force to act concurrently.

The main objectives of the design analysis are to satisfy allowable stress limits, to assure an adequate design margin, and to establish deformation limits which are concerned primarily with the functioning of the components.

The stress limits are established not only to assure that peak stresses will not reach unacceptable values, but also limit the amplitude of the oscillatory stress component in consideration of fatigue characteristics of the materials.

Both low and high cycle fatigue 4.2-41 SGS-UFSAR Revision 6 February 15, 1987

stresses are considered when the allowable amplitude of oscillation is established.

As part of the evaluation of design loading conditions, extensive testing and inspection are performed from the initial selection of raw materials up to and including component installation and plant operation. Among these tests and inspections are those performed during component fabrication, plant construction, startup and checkout, and during plant operation.

4.2.2.4 Design Loading Categories The combination of design loadings fits into either the normal, upset, or faulted conditions as defined in the ASME Section III Code.

Loads and deflections imposed on components due to shock and vibration are determined analytically and experimentally in both scaled models and operating reactors. The cyclic stresses due to these dynamic loads and deflections are combined with the stresses imposed by loads from component weights, hydraulic forces, and thermal gradients for the determination of the total stresses of the internals.

The reactor internals are designed to withstand stresses originating from various operating conditions summarized in Table 5.1-10.

The scope of the stress analysis problem is very large requiring many different techniques and methods, both static and dynamic. The analysis performed depends on the mode of operation under consideration.

Allowable Deflections For normal operating conditions, downward vertical deflection of the lower core support plate is negligible.

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For the LOCA plus the DBE condition, the deflection criteria of critical internal structures are the limiting values given in Table 4.2-1. The corresponding no loss of function limits are included in Table 4.2-1 for comparison purposes with the allowed criteria.

The criteria for the core drop accident are based upon analyses which have been performed to determine the total downward displacement of the internal structures following a hypothesized core drop resulting from loss of the normal core barrel supports. The initial clearance between the secondary core support structures and the reactor vessel lower head in the hot condition is approximately 1/2 inch. An additional displacement of approximately 3/4 inch would occur due to strain of the energy absorbing devices of the secondary core support; thus the total drop distance is about 1 1/4 inches which is insufficient to permit the grips of the rod cluster control assembly to come out of the guide thimble in the fuel assemblies.

Specifically, the secondary core support is a device which will never be used, except during a hypothetical accident of the core support (core barrel, barrel flange, etc.). There are four supports in each reactor. This device limits the fall of the core and absorbs the energy of the fall which otherwise would be imparted to the vessel. The energy of the fall is calculated assuming a complete and instantaneous failure of the primary core support and is absorbed during the plastic deformation of the controlled volume of stainless steel, loaded in tension. The maximum deformation of this austenitic stainless piece is limited to approximately 15 percent, after which a positive step is provided to ensure support.

4.2.2.5 Design Criteria Basis The basis for the design stress and deflection criteria is identified below.

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Allowable Stress For normal operating conditions, Section III of the ASME Nuclear Power Plant Components Code is used as a basis for evaluating acceptability of calculated stresses. Both static and alternating stress intensities are considered. Under Code Case 1618, bolt material Type 316 stainless steel is now covered in ASME Section III and is so treated. It should be noted that the allowable stresses in Section III of the ASME Code are based on unirradiated material properties.

In view of the fact that irradiation increases the strength of the Type 304 stainless steel used for the internals, although decreasing its elongation, it is considered that use of the allowable stresses in Section III is appropriate and conservative for irradiated internal structures.

The allowable stress limits during the DBA used for the core support structures are based on the January 1971 draft of the ASME Code for Core Support Structures, Subsection NG, and the Criteria for Faulted Conditions.

4.2.3 Reactivity Control System 4.2.3.1 Design Bases Bases for temperature, stress on structural members, and material compatibility are imposed on the design of the reactivity control components.

4.2.3.1.1 Design Stresses The Reactivity Control System is designed to withstand stresses originating from various operating conditions as summarized in Table 5.2-10.

Allowable Stresses: For normal operating conditions, Section III of the ASME Boiler and Pressure Code is used as a general guide.

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Dynamic Analysis: The cyclic stresses due to dynamic loads and deflections are combined with the stresses imposed by loads from component weights, hydraulic forces, and thermal gradients for the determination of the total stresses of the Reactivity Control System.

4.2.3.1.2 Material Compatibility Materials are selected for compatibility in a PWR environment, for adequate mechanical properties at room and operating temperature, for resistance to adverse property changes in a radioactive environment, and for compatibility with interfacing components.

4.2.3.1.3 Reactivity Control Components The reactivity control components are subdivided into two categories:

1. Permanent devices used to control or monitor the core

2. Temporary devices used to control or monitor the core.

The permanent type components are the rod cluster control assemblies, control rod drive assemblies, neutron source assemblies, and thimble plug assemblies.

Although the thimble plug assembly does not directly contribute to the reactivity control of the reactor, it is presented as a Reactivity Control System component in this document because it can be used to restrict bypass flow through those thimbles not occupied by absorber, source, or burnable absorber rods.

The temporary component is the burnable absorber assembly. The design bases for each of the mentioned components are in the following paragraphs.

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Absorber Rods The following are considered design conditions under Subsections NG and NB of the ASME Boiler and Pressure Vessel Code Section III.

1. The external pressure equal to the Reactor Coolant System operating pressure

2. The wear allowance equivalent to 1,000 reactor trips

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

4. Forces imposed on the rods during rod drop

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

6. Radiation exposure for maximum core life

7. Temperature effects at operating conditions The absorber material temperature shall not exceed its melting temperature (1470°F for Ag-In-Cd absorber material) (14)

Burnable Absorber Rods Two kinds of discrete burnable absorber rods may be used at Salem. The first is the borosilicate glass (PYREX) burnable absorber. The second is the Wet Annular Burnable Absorber (WABA) design. See References 23 and 24 of this section for a more detailed discussion of WABA. Reference 24 extends the allowable lifetime of the WABA rods from 18,000 to 40,000 EFPH, and allows utilization of the WABA rods for up to two eighteen month cycles without requiring additional inspections of the WABA rods. Discrete burnable absorbers are utilized to meet nuclear design requirements of UFSAR Section 4. 3. Some comparative data is also shown in UFSAR Table 4.3-1.

The burnable absorber rod clad is designed using Subsections NG and NB of the ASME Boiler and Pressure Vessel Code, Section III, 1973 as a general guide for Conditions I and I I . For abnormal loads during Conditions I I I and IV, Code stresses are not considered limiting. Failures of the burnable absorber rods during these conditions must 4.2-46 SGS-UFSAR Revision 23 October 17, 2007

not interfere with reactor shutdown or emergency cooling of the fuel rods.

The burnable absorber material is nonstructural. The structural elements of the burnable absorber rods are designed to maintain the absorber geometry even if the absorber material is fractured. The PYREX rods are designed so that the borosilicate absorber material is below its softening temperature ( 1492°F* for reference 12.5 weight percent boron rods). The WABA rods utilize an aluminum oxide/boron carbide (Al 20 3 -B 4C) absorber material which is a sintered ceramic and has a very high melting temperature. In addition, the structural elements are designed to prevent excessive slumping.

Neutron Source Rods The neutron source rods are designed to withstand the following:

1. The external pressure equal to the Reactor Coolant System operating pressure

2. An internal pressure equal to the pressure generated by released gases over the source rod life Thimble Plug Assembly If used in the core (optional), the thimble plug assemblies satisfy the following:

1. Accommodate the differential thermal expansion between the fuel assembly and the core internals

• Borosilicate glass is accepted for use in burnable absorber rods if the softening temperature is 1510 + 18°F. The softening temperature is defined in ASTM C 338.

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2. Maintain positive contact with the fuel assembly and the core internals

3. Limit the flow through each occupied thimble to acceptable design value 4.2.3.1.4 Control Rod Drive Mechanisms The mechanisms are Class I components designed to meet the stress requirements for normal operating conditions of Section III of the ASME Boiler and Pressure Vessel Code. Both static and alternating stress intensities are considered.

The stresses originating from the required design transients are included in the analysis.

A dynamic seismic analysis is required on the control rod drive mechanism (CRDM) when a seismic disturbance has been postulated to confirm the ability of the mechanism to meet ASME Code, Section III allowable stresses and to confirm its ability to trip when subjected to the seismic disturbance.

The CRDM design used for the 17 x 17 fuel assembly control rod is identical to the 15 x 15 CRDM. The seismic analysis and response of 17 x 17 CRDM will be identical to those of the 15 x 15 mechanisms.

Control Rod Drive Mechanism Operational Requirements The basic operational requirements for the CRDMs are as follows:

1. 5/8-inch step

2. 150-inch travel

3. 360-pound maximum load

4. Step in or out at 45 inches/minute (72 steps/minute) 4.2-48 SGS-UFSAR Revision 6 February 15, 1987

5. Power interruption shall initiate release of drive rod assembly

6. Trip delay of less than 150 ms - Free fall of drive rod assembly shall begin less than 150 ms after power interruption no matter what holding or stepping action is being executed with any load and coolant temperatures of 100°F to 550°F

7. 40-year design life with normal refurbishment 6

8. 28,00 complete travel excursions which is 13 x 10 steps with normal refurbishment 4.2.3.2 Design Description Reactivity control is provided by neutron absorbing rods and a soluble chemical neutron absorber (boric acid) The boric acid concentration is varied to control long-term reactivity changes such as:

1. Fuel depletion and fission product buildup

2. Cold to hot, zero power reactivity change

3. Reactivity change produced by intermediate-term fission products such as xenon and samarium

4. Burnable absorber depletion The rod cluster control assemblies provide reactivity control for:

1. Shutdown

2. Reactivity changes due to coolant temperature changes in the power range 4.2-49 SGS-UFSAR Revision 11 July 22, 1991

3. Reactivity changes associated with the power coefficient of reactivity

4. Reactivity changes due to void formation If soluble boron were the sole means of control, the moderator temperature coefficient could be positive. It is desirable to have a negative moderator temperature coefficient throughout the entire cycle in order to reduce possible deleterious effects caused by a positive coefficient during loss-of-coolant or loss-of-flow accidents. This is accomplished by installation of burnable absorber assemblies.

The neutron source assemblies and spontaneous fission neutron sources associated with the irradiated fuel assemblies provide a means of monitoring the core during periods of low neutron activity.

The most effective reactivity control components are the rod cluster control assemblies and their corresponding drive rod assemblies which are the only kinetic parts in the reactor. Figure 4.2-13 identified the rod cluster control and drive rod assembly, in addition to the arrangement of these components in the reactor relative to the interfacing fuel assembly, guide tubes, and CRDM.

The guidance system for the control rod cluster is provided by the guide tube as shown on Figure 4.2-13. The guide tube provides two regimes of guidance. In the lower section a continuous guidance system provides support immediately above the core. This system protects the rod against excessive deformation and wear due to hydraulic loading. The region above the continuous section provides support and guidance at uniformly spaced intervals.

The envelope of support is determined by the pattern of the control rod cluster as shown on Figure 4.2-13. The guide tube assures alignment and support of the control rods, spider body, 4.2-50 SGS-UFSAR Revision 25 October 26, 2010

and drive rod while maintaining trip times at or below required limits. In the following paragraphs, each reactivity control component is described in detail.

4.2.3.2.1 Reactivity Control Components Rod Cluster Control Assembly The rod cluster control assemblies are divided into two categories: control and shutdown. The control groups compensate for reactivity changes due to variations in operating conditions of the reactor, i.e., power and temperature variations. Two criteria have been employed for selection of the control groups. First, the total reactivity worth must be adequate to meet the nuclear requirements of the reactor. Second, in view of the fact that some of these rods may be partially inserted at power operation, the total power peaking factor should be low enough to ensure that the power capability is met. The control and shutdown groups provide adequate shutdown margin which is defined as the amount by which the core would be subcritical at hot shutdown if all rod cluster control assemblies are tripped assuming that the highest worth assembly remains fully withdrawn and assuming no changes in xenon or boron concentration.

A rod cluster control assembly comprises a group of individual neutron absorber rods fastened at the top end to a common spider assembly, as illustrated on Figure 4.2-14.

The absorber material used in the control rods is silver-indium-cadmium single piece absorber rod which is essentially "black" to thermal neutrons and has sufficient additional resonance absorption to significantly increase its worth.

The alloy is in the form of extruded rods which are sealed in stainless steel tubes to prevent the rods from coming in direct contact with the coolant. In construction, the silver-indium-cadmium rods are inserted into cold-worked stainless steel tubing which is then sealed at the bottom and the top by welded Type 308L stainless steel end plugs as shown on Figure 4. 2-15. Sufficient diametral and end clearance is provided to accommodate relative thermal expansions. The cladding surface has been ion-nitrided for hardening and corrosion resistance.

The bottom plugs are made bullet-nosed to reduce the hydraulic drag during reactor trip and to guide smoothly into the dashpot section of the fuel assembly guide thimbles. The upper end plug is threaded for assembly to the spider and is machined with a reduced diameter shank to provide flexibility to the joint for any misalignment condition.

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The spider assembly is a one-piece machined casting in the form of a central hub with radial vanes containing cylindrical fingers from which the absorber rods are suspended. Handling detents and detents for connection to the drive rod assembly are machined into the upper end of the hub. A coil spring inside the spider body absorbs the impact energy at the end of a trip insertion. A centerpost which holds the spring and its retainer is threaded into the hub within the skirt and welded to prevent loosening in service. The spider casting material is CF3M cast 316 stainless steel.

The absorber rods are fastened securely to the spider assembly as shown in Figure 4.2-15 to assure trouble free service. The threaded end of the upper end plug is inserted into the bottom of the spider boss hole. A nut is tightened on and welded to the spider boss to prevent loosening. A lock pin is inserted into the aligned holes of the spider base and upper end plug and welded to prevent the end plug and rod from backing off.

The overall length is such that when the assembly is withdrawn through its full travel the tips of the absorber rods remain engaged in the guide thimbles so that alignment between rods and thimbles is always maintained. Since the rods are long and slender, they are relatively free to conform to any small misalignments with the guide thimble.

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Burnable Absorber Assembly Each burnable absorber assembly consists of burnable absorber rods attached to a holddown assembly. Burnable absorber assemblies are shown on Figure 4.2-16.

The PYREX absorber rods consist of borosilicate glass tubes contained within Type 304 stainless steel tubular cladding which is plugged and seal welded at the ends to encapsulate the glass. The glass is also supported along the length of its inside diameter by a thin wall tubular inner liner of Type 304 stainless steel. The top end of the liner is open to permit the diffused helium to pass into the void volume and the liner overhangs the glass. The liner has an outward flange at the bottom end to maintain the position of the liner with the glass. A PYREX burnable absorber rod is shown in longitudinal and transverse cross sections on Figure 4.2-17.

The WABA consist of aluminum oxide/boron carbide pellets contained within Zircaloy tubular cladding which is plugged and seal welded at the ends to encapsulate the pellets. The pellets are also supported along the length of their insider diameter by a thin wall tubular liner of Type 304 stainless steel. The top and bottom of the inner tube, or liner, is open to allow for coolant flow. There is a void above the Al 2 0 3 -B 4 C pellets between the inner and outer tube to contain diffused helium. A cross section drawing of the WABA rod is shown in Figure 4.2-17A.

The rods are statically suspended and positioned in selected guide thimbles within specified fuel assemblies. The absorber rods in each fuel assembly are grouped and attached together at the top end of the rods to a holddown assembly by a flat, perforated retaining plate which fits within the fuel assembly top nozzle and rests on the adaptor plate. The retaining plate (and the absorber rods) is held down and restrained against vertical motion through a spring pack which is attached to the plate and is compressed by the upper core plate when the reactor upper internals assembly is lowered into the reactor. This arrangement assures that the absorber rods cannot be ejected from the core by flow forces. Each rod is permanently attached to the base plate by a nut which is lock welded into place.

The clad in the rod assemblies is Zircaloy or slightly cold worked Type 304 stainless steel. All other structural materials are Type 304 or 308 stainless steel except for the springs which are Inconel 718. The borosilicate glass tube and I or aluminum oxide/boron carbide pellets provide sufficient boron content to meet the criteria discussed in Section 4.3.1.

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Neutron Source Assembly A neutron source assembly can be used to provide a base neutron level to monitor core multiplication to changes in core reactivity. Since there is very little neutron activity when the core is subcri tical, such as refueling and approach to criticality, neutron source assemblies are placed in the reactor if required to provide a count rate greater than 1 cps in Mode 3 prior to starting the approach to critical on the source range monitors. The source range monitors, which receive their signal from the source range detectors, are used primarily when the core is subcritical and during special subcritical modes of operation. In addition to having a greater than 1 cps count rate, the source channels should maintain a signal to noise ratio of at least two in Mode 3 prior to beginning the approach to critical.

The source assembly also supplements detection of changes in the core multiplication factor during core loading, refueling and approach to criticality. This can be done since the multiplication factor is related to an inverse function of the detector count rate. Therefore a change in the multiplication factor can be detected during addition of fuel assemblies while loading the core, a change in control rod positions, and changes in boron concentration.

Both primary and secondary neutron source rods are used. The primary source rod, containing a radioactive material, spontaneously emits neutrons during the initial core loading and reactor startup. After the primary source rod decays beyond the desired neutron flux level, neutrons are then supplied by the secondary source rod. The secondary source rod contains a stable material which must be activated by neutron bombardment during reactor operation. The activation results in the subsequent release of neutrons. This becomes a source of neutrons during periods of low neutron flux, such as during refueling and subsequent startups.

The initial reactor core employs four source assemblies, two primary source assemblies, and two secondary source assemblies. Each primary source assembly contains one primary source rod and between 0 and 23 burnable absorber rods. Each secondary source assembly contains a symmetrical grouping of 4 or 6 secondary source rods and between 0 and 2 0 burnable absorber rods. The 4 rodlet secondary source utilizes a single encapsulated design. The 6 rodlet secondary source utilizes a double encapsulated design which provides additional margin against source material leakage. Locations not filled with a source or burnable absorber rod may contain a thimble plug (optional). Source assemblies are shown on Figures 4.2-18 and 4.2-19. A comparison of the single and double encapsulated secondary source design is provided in Table 4.2-2.

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Following the initial operating cycle, the actual source range count rate will depend on the core loading and the outage duration. Normally for subsequent reloads, the primary sources are removed and the secondary sources continue functioning.

The core loading determines the placement of the secondary sources and other inherent neutron sources (i.e., spontaneous fission from the irradiated fuel) relative to the location of the source range detectors. For secondary source 124 assemblies made of antimony-beryllium (Sb-Be, gamma-neutron reaction), sb has a 60.2 day half-life. Thus, extended outages may impact the effectiveness of the secondary sources. For such extended outages, inherent neutron sources are sufficient to provide the required greater than 1 cps count rate in Mode 3 prior to beginning the approach to critical. Likewise, if a reload design has sufficient spontaneous fission neutrons to ensure the minimum required count rate response on the source range channels, then there is no need to install the secondary source assemblies.

The primary and secondary source rods both utilize the same cladding material as the absorber rods. The single encapsulated secondary source rods contain approximately 500 grams of Sb-Be pellets in one rod. The double encapsulated secondary source rods contain approximately 338 grams of Sb-Be pellets in one rod. The primary source rods contain capsules of Californium source material and alumina spacer rods to position the source material within the cladding.

The rods in each assembly are permanently fastened at the top end to a holddown assembly, which is identical to that of the burnable absorber assemblies.

The other structural members are constructed of Type 304 stainless steel except for the springs. The springs exposed to the reactor coolant are wound from an age hardened nickel base alloy for corrosion resistance and high strength. The springs, when contained within the rods where corrosion resistance is not necessary, are oil tempered carbon steel.

Thimble Plug Assembly Thimble plug assemblies may be used in order to limit bypass flow through the rod cluster control guide thimbles in fuel assemblies which do not contain either control rods, source rods, or burnable absorber rods.

The thimble plug assemblies as shown on Figure 4. 2-20 consist of a flat base plate with short rods suspended from the bottom surface and a spring pack assembly.

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The 24 short rods, called thimble plugs, project into the upper ends of the guide thimbles to reduce the bypass flow area. Similar short rods are also used on the source assemblies and burnable absorber assemblies to plug the ends of all vacant fuel assembly guide thimbles. At installation in the core, the thimble plug assemblies interface with both the upper core plate and with the fuel assembly top nozzles by resting on the adaptor plate. The spring pack is compressed by the upper core plate when the upper internals assembly is lowered into place. Each thimble plug is permanently attached to the base plate by a nut which is locked to the threaded end of the plug by a small lock-bar welded to the nut.

All components in the thimble plug assembly, except for the springs, are constructed from Type 304 stainless steel. The springs are wound from an age hardened nickel base alloy for corrosion resistance and high strength.

4.2.3.2.2 Control Rod Drive Mechanism All parts exposed to reactor coolant are made of metals which resist the corrosive action of the water. Three types of metals are used exclusively:

stainless steels, Inconel and cobalt based alloys. Wherever magnetic flux is carried by parts exposed to the main coolant, 400 series stainless steel is used. Cobalt based alloys are used for the pins and latch tips. Inconel is used for the springs of both latch assemblies and Type 304 stainless steel is used for all pressure containing parts. Hard chrome plating provides wear surfaces on the sliding parts and prevents galling between mating parts.

A position indicator assembly slides over the CRDM rod travel housing. It detects the drive rod assembly position by means of 42 discrete coils that magnetically sense the entry and presence of the rod drive line through its center line over the normal length of the drive rod travel.

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Control Rod Drive Mechanism Control rod drive mechanisms are located on the dome of the reactor vessel.

They are coupled to rod control clusters which have absorber material over the entire length of the control rods and derive their name from this feature. The CRDM is shown on Figure 4.2-21 and schematically on Figure 4.2-22.

The primary function of the CRDM is to insert or withdraw rod control clusters within the core to control average core temperature and to shut down the reactor.

The CRDM is a magnetically operated jack. A magnetic jack is an arrangement of three electro-magnets which are energized in a controlled sequence by a power cycler to insert or withdraw rod control clusters in the reactor core in discrete steps.

The CRDM consists of four separate subassemblies. They are the pressure vessel, coil stack assembly, the latch assembly, and the drive rod assembly.

1. The pressure vessel includes a latch housing and a rod travel housing which are connected by a threaded, seal welded, maintenance joint which facilitates replacement of the latch assembly. The closure at the top of the rod travel housing is a threaded plug with a canopy seal weld for pressure integrity.

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The latch housing is the lower portion of the vessel and contains the latch assembly. The rod travel housing is the upper portion of the vessel and provides space for the drive rod during its upward movement as the control rods are withdrawn from the core.

2. The coil stack assembly includes the coil housings, an electrical conduit and connector, and three operating coils: 1) the stationary gripper coil, 2) the moveable gripper coil, and 3) the lift coil.

The coil stack assembly is a separate unit which is installed on the drive mechanism by sliding it over the outside of the latch housing. It rests on the base of the latch housing without mechanical attachment.

Energizing of the operation coils causes movement of the pole pieces and latches in the latch assembly.

3. The latch assembly includes the guide tube, stationary pole pieces, moveable pole pieces, and two sets of latches: 1) the moveable gripper latch, and 2) the stationary gripper latch.

The latches engage grooves in the drive rod assembly. The moveable gripper latches are moved up or down in 5/8 inch steps by the lift pole to raise or lower the drive rod assembly while the moveable gripper latches are repositioned for the next 5/8 inch step.

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4. The drive rod assembly includes a flexible coupling, a drive rod, a disconnect button, a disconnect rod, and a locking button.

The drive rod has 5/8 inch grooves which receive the latches during holding or moving of the drive rod. The flexible coupling is attached to the drive rod and produces the means for coupling to the rod control cluster assembly.

The disconnect button, disconnect rod, and locking button provide positive locking of the coupling to the rod control cluster assembly and permits remote disconnection of the drive rod.

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The CRDM is a trip design. Tripping can occur during any part of the power cycler sequencing if power to the coils is interrupted. The CRDM is threaded and seal welded on an adapter on top of the reactor vessel and is coupled to the rod control cluster assembly directly below.

The mechanism is capable of handling a 360 pound load, including the drive rod weight, at a rate of 45 inches per minute. Withdrawal of the rod control cluster is accomplished by magnetic forces while insertion is by gravity.

The mechanism internals are designed to operate in 650°F reactor coolant. The pressure vessel is designed to contain reactor coolant at 650°F and 2500 psia.

The three operating coils are designed to operate at 392°F with forced air cooling required to maintain that temperature.

The CRDM shown schematically on Figure 4.2-22 withdraws and inserts its control rod as electrical pulses are received by the operator coils. An ON or OFF sequence, repeated by silicon controlled rectifiers in the power programmer, causes either withdrawal or insertion of the control rod. Position of the control rod is measured by 42 discrete coils mounted on the position indicator assembly surrounding the rod travel housing. Each coil magnetically senses the entry and presence of the top of the ferro-magnetic drive rod assembly as it moves through the coil center line.

During plant operation the stationary gripper coil of the drive mechanism holds the control rod withdrawn from the core in a static position until the movable gripper coil is energized.

Rod Cluster Control Assembly Withdrawal The control rod is withdrawn by repetition of the following sequence of events:

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1. Movable Gripper Coil (B) - ON The latch locking plunger raises and swings the movable gripper latches into the drive rod assembly groove. A 1/16 inch axial clearance exists between the latch teeth and the drive rod.

2. Stationary Gripper Coil (A) - OFF The force of gravity, acting upon the drive rod assembly and attached control rod, causes the stationary gripper latches and plunger to move downward 1/16 inch until the load of the drive rod assembly and attached control rod is transferred to the movable gripper latches. The plunger continues to move downward and swings the stationary gripper latches out of the drive rod assembly groove.

3. Lift Coil (C)- ON The 5/8 inch gap between the movable gripper pole and the lift pole closes and the drive rod assembly raises one step length (5/8 inch).

4. Stationary Gripper Coil (A) - ON The plunger raises and closes the gap below the stationary gripper pole. The three links, pinned to the plunger, swing the stationary gripper latches into a drive rod assembly groove. The latches contact the drive rod assembly and lift it (and the attached control rod) 1/16 inch. The 1/16 inch vertical drive rod assembly movement transfers the drive rod assembly load from the moveable gripper latches to the stationary gripper latches.

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The latch locking plunger separates from the movable gripper pole under the force of a spring and gravity. Three links, pinned to the plunger, swing the three movable gripper latches out of the drive rod assembly groove.

6. Lift Coil (C) - OFF The gap between the movable gripper pole and lift pole opens. The movable gripper latches drop 5/8 inch to a position adjacent to a drive rod assembly groove.

7. Repeat Step 1 The sequence described above (1 through 6) is termed as one step or one cycle. The control rod moves 5/8 inch for each step or cycle.

The sequence is repeated at a rate of up to 72 steps per minute and the drive rod assembly (which has a 5/8 inch groove pitch) is raised 72 grooves per minute. The control rod is thus withdrawn at a rate up to 45 inches per minute.

Rod Cluster Control Assembly Insertion The sequence for control rod insertion is similar to that for control rod withdrawal, except the timing of lift coil (C) ON and OFF is changed to permit lowering the control rod.

1. Lift Coil (C) - ON The 5/8 inch gap between the movable gripper and lift pole closes.

The movable gripper latches are raised to a position adjacent to a drive rod assembly groove.

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2. Movable Gripper Coil (B) - ON The latch locking plunger raises and swings the movable gripper latches into a drive rod assembly groove. A 1/16 inch axial clearance exists between the latch teeth and the drive rod assembly.

3. Stationary Gripper Coil (A) - OFF The force of gravity, acting upon the drive rod assembly and attached control rod, causes the stationary gripper latches and plunger to move downward 1/16 inch until the load of the drive rod assembly and attached control rod is transferred to the movable gripper latches. The plunger continues to move downward and swings the stationary gripper latches out of the drive rod assembly groove.

4. Lift Coil (C) -OFF The force of gravity separates the movable gripper pole from the lift pole and the drive rod assembly and attached control rod drop down 5/8 inch.

5. Stationary Gripper (A) - ON The plunger raises and closes the gap below the stationary gripper pole. The three links, pinned to the plunger, swing the three stationary gripper latches into a drive rod assembly groove. The latches contact the drive rod assembly and lift it (and the attached control rod) 1/16 inch. The 1/16 inch vertical drive rod assembly movement transfers the drive rod assembly load from the movable gripper latches to the stationary gripper latches.

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The latch locking plunger separates from the movable gripper pole under the force of a spring and gravity. Three links, pinned to the plunger, swing the three movable gripper latches out of the drive rod assembly groove.

7. Repeat Step 1 The sequences are repeated, as for control rod withdrawal, up to 72 times per minute which give a control rod insertion rate of 45 inches per minute.

Holding and Tripping of the Control Rods During most of the plant operating time, the CRDMs hold the control rods withdrawn from the core in a static position. In the holding mode, only one coil, the stationary gripper coil (A), is energized on each mechanism. The drive rod assembly and attached control rod hang suspended from the three latches.

If power to the stationary gripper coil is cut off, the combined weight of the drive rod assembly and the rod cluster control assembly is sufficient to move latches out of the drive rod assembly groove. The control rod falls by gravity into the core. The trip occurs as the magnetic field, holding the stationary gripper plunger half against the stationary gripper pole, collapses and the stationary gripper plunger half is forced down by the weight acting upon the latches. After the drive rod assembly is released by the mechanism, it falls freely until the control rods enter the buffer section of their thimble tubes.

4.2.3.3 Design Evaluation 4.2.3.3.1 Reactivity Control Components The components are analyzed for loads corresponding to normal, upset, emergency, and faulted conditions. The analysis performed 4.2-63 SGS-UFSAR Revision 6 February 15, 1987

depends on the mode of operation under consideration.

The scope of the analysis requires many different techniques and methods, both static and dynamic.

Some of the loads that are considered on each component where applicable are as follows:

1. Control Rod Trip (equivalent static load)

2. Differential Pressure

3. Spring Preloads

4. Coolant Flow Forces (static)

5. Temperature Gradients

6. Differences in thermal expansion

a. Due to temperature differences

b. Due to expansion of different materials

7. Interference Between Components

8. Vibration (mechanically or hydraulically induced)

9. All Operational Transients Listed in Table 4.1-10

10. Pump Overspeed

11. Seismic Loads (OBE and DBE)

The main objective of the analysis is to satisfy allowable stress limits, to assure an adequate design margin, and to establish deformation limits which are concerned primarily with the 4.2-64 SGS-UFSAR Revision 6 February 15, 1987

functioning of the components. The stress limits are established not only to assure that peak stresses will not reach unacceptable values, but also limit the amplitude of the oscillatory stress component in consideration of fatigue characteristics of the materials. Standard methods of strength of materials are used to establish the stresses and deflections of these components. The dynamic behavior of the reactivity control components has been studied using experimental test data (11) and experience from operating reactors.

The design of reactivity component rods provides a sufficient cold void volume within the burnable absorber and source rods to limit the internal pressures to a value which satisfies the criteria in Section 4.2.3.1. The void volume for the helium in the burnable absorber rods is obtained through the use of glass in tubular form which provides a central void along the length of the rods.

Helium gas is not released by the neutron absorber rod material; thus the absorber rod only sustains an external pressure during operating conditions.

The internal pressure of source rods continues to increase from ambient until end-of-life at which time the internal pressure never exceeds that allowed by the criteria in Section 4.2.3.1. The stress analysis of reactivity component rods assumes 100 percent gas release to the rod void volume, considers the initial pressure within the rod, and assumes the pressure external to the component rod is zero.

Based on available data for properties of the borosilicate glass and on nuclear and thermal calculations for the burnable absorber rods, gross swelling or cracking of the glass tubing is not expected during operation. Some minor creep of the glass at the hot spot on the inner surface of the tube could occur but would continuously until the glass came in contact with the inner liner. The wall thickness of the inner liner is sized to provide adequate support in the event of slumping and to collapse locally before rupture of the exterior cladding if unexpected large volume changes due to swelling or cracking should occur. The top of the inner liner is open to 4.2-65 SGS-UFSAR Revision 11 July 22, 1991

allow communication to the central void by the helium which diffuses out of the glass.

Sufficient diametral and end clearances have been provided in the neutron absorber, burnable absorber, and source rods to accommodate the relative thermal expansions between the enclosed material and the surrounding clad and end plugs. There is no bending or warping induced in the rods although the clearance offered by the guide thimble would permit a postulated warpage to occur without restraint on the rods. Bending, therefore, is not considered in the analysis of the rods. The radial and axial temperature profiles have been determined by considering gap conductance, thermal expansion, and neutron and/or gamma heating of the contained material as well as gamma heating of the clad. The maximum neutron absorber material temperature was found to be less than 850°F which occurs axially at only the highest flux region. The maximum borosilicate glass temperature was calculated to be about 1200°F and takes place following the initial rise to power. The glass temperature then decreases rapidly for the following reasons: (1) reduction in power generation due to depletion; (2) better gap conductance as the helium produced diffuses to the gap; and (3) external gap reduction due to borosilicate glass creep. Rod, guide thimble, and dashpot flow analysis performed indicates that the flow is sufficient to prevent coolant boiling and maintain clad temperatures at which the clad material has adequate strength to resist coolant operating pressures and rod internal pressures.

Analysis on the rod cluster control spider indicates the spider is structurally adequate to withstand the various operating loads including the higher loads which occur during the drive mechanism stepping action and rod drop.

The reactivity control component materials selected are considered to be the best available from the standpoint of resistance to irradiation damage and compatibility to the reactor environment. The materials selected partially dictate the reactor environment 4.2-66 SGS-UFSAR Revision 25 October 26, 2010

(e.g., Cl control in the coolant). The current design type reactivity controls have been in service for more than 10 years with no apparent degradation of construction materials.

With regard to the materials of construction exhibiting satisfactory resistance to adverse property changes in a radioactive environment, it should be noted that on work on breeder reactors in current design, similar materials are being applied. At high fluences the austenitic materials increase in strength with a corresponding decreased ductility (as measured by tensile tests) but energy absorption (as measured by impact tests) remains quite high. Corrosion of the materials exposed to the coolant is quite low and proper control of Cl and 0 2

in the coolant will prevent the occurrence of stress corrosion. All of the austenitic stainless steel base materials used are processed and fabricated to preclude sensitization.

Analysis of the rod cluster control assemblies shows that if the drive mechanism housing ruptures, the rod cluster control assembly will be ejected from the core by the pressure differential of the operating pressure and ambient pressure across the drive rod assembly. The ejection is also predicted on the failure of the drive mechanism to retain the drive rod/ rod cluster control assembly position. It should be pointed out that a drive mechanism housing rupture will cause the ejection of only one rod cluster control assembly with the other assemblies remaining in the core.

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Ejection of a burnable absorber or thimble plug assembly (if used) is conceivable based on the postulation that the holddown bar fails and that the base plate and burnable absorber rods are severely deformed. In the unlikely event that failure of the holddown bar occurs, the upward displacement of the burnable absorber assembly only permits the base plate to contact the upper core plate. Since this displacement is small, the major portion of the burnable absorber material remains positioned within the core. In the case of the thimble plug assembly, the thimble plugs will partially remain in the fuel assembly guide thimbles thus maintaining a majority of the desired flow impedance. Further displacement or complete ejection would necessitate the square base plate and burnable absorber rods be forced, thus plastically deformed, to fit up through a smaller diameter hole. It is expected that this condition requires a substantially higher force or pressure drop than that of the holddown bar failure.

Experience with control rods, burnable absorber rods, and source rods is discussed in Reference 2.

The mechanical design of the reactivity control components provides for the protection of the active elements to prevent the loss of control capability and functional failure of critical components. The components have been reviewed for potential failure and consequences of a functional failure of critical parts. The results of the review are summarized below.

Rod Cluster Control Assembly

1. The basic absorbing material is sealed from contact with the primary coolant and the fuel assembly and guidance surfaces by a high quality stainless steel clad.

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Potential loss of absorber mass or reduction in reactivity control material due to mechanical or chemical erosion or wear is therefore reliably prevented.

2. A breach of the cladding for any postulated reason does not result in serious consequences. The absorber material, silver-indium-cadmium, is relatively inert and would still remain remote from high coolant velocity regions. Rapid loss of material resulting in significant loss of reactivity control material would not occur.

3. The individually clad absorber rods are doubly secured to the retaining spider finger by a threaded top end plug secured by a nut welded to the finger and a welded lock pin.

It should also be noted that in several instances of control rod jamming caused by foreign particles, the individual rods at the site of the jam have borne the full capacity of the CRDM and higher impact loads to dislodge the jam without failure. The guide tube card/guide thimble arrangement is such that large loads are required to buckle individual control rods. The conclusion to be drawn from this experience is that this joint is extremely insensitive to potential mechanical damage. A failure of the joint would result in the insertion of the individual rod into the core.

This results in reduced reactivity which is a fail safe condition.

Further information is given in Reference 2.

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4. The spider is a one-piece machined casting and includes the radial vanes and fingers. Reliability is increased by not using brazed joints. Casting allows the rod holes to be drilled to the positional tolerances prior to assembly to ensure the rods will align with the guide cards.

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5. The spider hub being of a one-piece machined casting is very rugged and of extremely low potential for damage. It is difficult to postulate any condition to cause failure. Should some unforeseen event cause fracture of the hub above the vanes, the lower portion with the vanes and rods attached would insert by gravity into the core causing reactivity decrease. The rod could then not be removed by the drive line, again a fail safe condition. Fracture below the vanes cannot be postulated since all loads, including scram impact, are taken above the vane elevation.

6. The rod cluster control rods are provided a clear channel for insertion by the guide thimbles of the fuel assemblies. All fuel rod failures are protected against by providing this physical barrier between the fuel rod and the intended insertion channel.

Distortion of the fuel rods by bending cannot apply sufficient force to damage or significantly distort the guide thimble. Fuel rod distortion by swelling, though precluded by design, would be terminated by fracture before contact with the guide thimble occurs. If such were not the case, it would be expected that a force reaction at the point of contact would cause a slight deflection of the guide thimble. The radius of curvature of the deflected shape of the guide thimbles would be sufficiently large to have a negligible influence on rod cluster control insertion.

Burnable Absorber Assemblies The burnable absorber assemblies are static temporary reactivity control elements. The axial position is assured by the holddown assembly which bears against the upper core plate. Their lateral position is maintained by the guide thimbles of the fuel assemblies.

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The individual rods are shouldered against the underside of the retainer plate and securely fastened at the top by a threaded nut which is then locked in place. The square dimension of the retainer plate is larger than the diameter of the flow holes through the core plate. Failure of the holddown bar or spring pack therefore does not result in ejection of the burnable absorber rods from the core.

The only incident that could potentially result in ejection of the burnable absorber rods is a multiple fracture of the retainer plate. This is not considered credible because of the light loads borne by this component. During normal operation the loads borne by the plate are approximately 5 lbs per rod, or a total of 100 lbs. distributed at the points of attachment. Even a multiple fracture of the retainer plate would result in jamming of the plate segments against the upper core plate, again preventing ejection. Excessive reactivity increase due to burnable poison ejection is therefore prevented.

The same type of stainless steel clad used on rod cluster control rods is also used on the burnable absorber rods. In this application there is even less susceptibility to mechanical damage since these are static assemblies. The guide thimbles of the fuel assembly afford the same protection from damage due to fuel rod failures as that described for the rod cluster control rods.

The consequences of clad breach are also similarly small. The absorber material is borosilicate glass which is maintained in position by a central hollow tube. In the event of a hole developing in the clad for any postulated reason the expected consequence is only the loss of the helium produced by the absorption process into the primary coolant. The glass is chemically inert and remains remote from high coolant velocities; therefore significant loss of absorber material resulting in reactivity increase is not expected.

4.2-72 SGS-UFSAR Revision 11 July 22, 1991

Drive Rod Assemblies All postulated failures of the drive rod assemblies either by fracture or uncoupling, lead to the fail safe condition. If the drive rod assembly fractures at any elevation, that portion remaining coupled falls with and is guided by the rod cluster control assembly. This always results in reactivity decrease for the control rods.

4.2.3.3.2 Control Rod Drive Mechanism Material Selection All pressure-containing materials comply with Section III of the ASME Pressure Vessel Code, and, with the exception of the needle vent valve, will be fabricated from austenitic (Type 304) stainless steel or CF-8 stainless steel.

The vent valve is a modified austenitic stainless steel cap screw.

Magnetic pole pieces are fabricated from Type 410 stainless steel. All nonmagnetic parts, except pins and springs, are fabricated from Type 304 stainless steel. Haynes 25 is used to fabricate link pins. Springs are made from Inconel-X. Latch arm tips are clad with Stelli te to provide improved wearability. Hard chrome plate and Stellite are used selectively for bearing and wear surfaces.

At the start of the development program, a survey was made to determine whether a material better than Type 410 stainless steel was available for the magnetic pole pieces. Ideal material requirements are as follows:

1. High magnetic saturation value 4.2-73 SGS-UFSAR Revision 22 May 5, 2006

2. High permeability

3. Low coercive force

4. High resistivity

5. High Curie temperature

6. Corrosion resistant

7. High impact strength

8. Nonoriented

9. High machinability

10. Radiation damage After a comprehensive material trade-off study was made it was decided that the Type 410 stainless steel was satisfactory for this application.

The cast coil housings require a magnetic material. Both low-carbon cast steel and ductile iron have been successfully tested for this application. The choice, made on the basis of cost, indicated that ductile iron will be specified on the CRDM. The finished housings are zinc plated to provide corrosion resistance.

Coils are wound on bobbins of molded Dow Corning 302 material, with double glass-insulated copper wire. Coils are then vacuum impregnated with silicon varnish. A wrapping of mica sheet is secured to the coil outer surface. The result is a well-insulated coil capable of sustained operation at 200°C.

The drive shaft assembly utilized a Type 410 stainless steel drive rod. The coupling is machined from Type 403 stainless steel. Other 4.2-74 SGS-UFSAR Revision 6 February 15, 1987

parts are Type 304 stainless steel with the exception of the springs which are Inconel-X and the locking button which is Haynes 25.

Radiation Damage As required by the equipment specification, the CRDMs are designed to meet a radiation requirement of 10 Rads/Hr. Materials have been selected to meet this 6

requirement. The above radiation level which amount to 1. 7 53 x 10 Rads in 20 years will not degrade control rod drive mechanism life. Control rod drive mechanisms at Yankee Rowe which have been in operation since 1960 have not experienced problems due to radiation.

Positioning Requirements The mechanism has a step length of 5/8 inches which determines the positioning capabilities of the control rod drive mechanism. (Note: Positioning requirements are determined by reactor physics.)

Elevation of Materials Adequacy The ability of the pressure housing components to perform throughout the design lifetime as defined in the equipment specification is confirmed by the stress analysis report required by the ASME Boiler and Pressure Vessel Code,Section III. Internals components subjected to wear will withstand a minimum of 3,000,000 steps without refurbishment as confirmed by life tests.

Results of Dimensional and Tolerance Analysis With respect to the CRDM systems as a whole, critical clearances are present in the following areas:

1. Latch assembly (Diametral clearances) 4.2-75 SGS-UFSAR Revision 20 May 6, 2003

2. Latch arm-drive rod clearances

3. Coil stack assembly-thermal clearances

4. Coil fit in coil housing The following write-up defines clearances that are designed to provide reliable operation in the CRDM in these four critical areas. These clearances have been proven by life tests and actual field performance at operating plants.

Latch Assembly - Thermal Clearances The magnetic jack has several clearances where parts made of Type 410 stainless steel fit over parts made from Type 304 stainless steel. Differential thermal expansion is therefore important. Minimum clearances of these parts at 650°F minimum clearance is 0.0045 inch, and at the maximum expected operating temperatures of 550°F is 0.0057 inch.

Latch Arm - Drive Rod Clearances The CRDM incorporates a load transfer action. The movable or stationary gripper latch is not under load during engagement, as previously explained, due to load transfer action.

Figure 4.2-23 shows latch clearance variation with the drive rod as a result of minimum and maximum temperatures. Figure 4.2-24 shows clearance variations over the design temperature range.

Coil Stack Assembly - Thermal Clearances The assembly clearance of the coil stack assembly over the latch housing was selected so that the assembly could be removed under all anticipated conditions of thermal expansion.

4.2-76 SGS-UFSAR Revision 6 February 15, 1987

At 70°F the inside diameter of the coil stack is 7.308/7.298 inches. The outside diameter of the latch housing is 7.260/7.270 inches.

Thermal expansion of the mechanism due to operating temperature of the CRDM results in minimum inside diameter of the coil stack being 7. 310 inches at 222°F and the maximum latch housing diameter being 7.302 inches at 532°F.

Under the extreme tolerance conditions listed above, it is necessary to allow time for a 70°F coil housing to heat during a replacement operation.

Four coil stack assemblies were removed from four hot CRDMs mounted on 11.035-inch centers on a 550°F test loop, allowed to cool, and then replaced without incident as a test to prove the proceeding.

Coil Fit in Coil Housing Control rod drive mechanism and coil housing clearances are selected so that coil heatup results in a close or tight fit. This is done to facilitate thermal transfer and coil cooling in a hot CRDM.

4.2.3.4 Tests, Verification, and Inspections 4.2.3.4.1 Reactivity Control Components Tests and inspections are performed on each reactivity control component to verify the mechanical characteristics. In the case of the rod cluster control assembly, prototype testing has been conducted, and both manufacturing test/inspections and functional testing at the plant site are performed.

4.2-77 SGS-UFSAR Revision 6 February 15, 1987

During the component manufacturing phase, the following requirements apply to the reactivity control components to assure the proper functioning during reactor operation:

1. All materials are procured to specifications to attain the desired standard of quality.

2. All clad/end plug welds are checked for integrity by visual inspection and X-ray, and are helium leak checked. All the seal welds in the neutron absorber rods, burnable poison rods, and source rods are checked in this manner.

3. To assure proper fitup with the fuel assembly, the rod cluster control, burnable poison, and source assemblies are installed in the fuel assembly and checked for binding in the dry condition.

The rod cluster control assemblies (RCCA) are functionally tested following core loading, but prior to criticality to demonstrate reliable operation of the assemblies. Each assembly is operated (and tripped) one time at no flow/cold conditions and one time at full flow/hot conditions. In addition, selected assemblies, amounting to about 15 to 20 percent of the total assemblies are operated at no-flow/operating temperature conditions and full flow/ambient conditions. Also the slowest rod and the fastest rod are tripped 10 times at no-flow/ambient conditions and at full flow/operating temperature conditions. Thus each assembly is tested a

4.2-78 SGS-UFSAR Revision 17 October 16, 1998

minimum of 2 times or up to 14 times maximum to ensure that the assemblies are properly functioning.

In order to demonstrate continuous free movement of the RCCAs and to ensure acceptable core power distributions during operations, partial movement checks are performed on every RCCA, as required by the Technical Specifications. In addition, periodic drop tests of the RCCAs are performed at each refueling shutdown to demonstrate continued ability to meet trip time requirements.

If an RCCA cannot be moved by its mechanism, adjustments in the boron concentration of the coolant ensure that adequate shutdown margin would be achieved following a trip. Thus, inability to move one RCCA can be tolerated.

More than one inoperable RCCA could be tolerated but would impose additional demands on the plant operator. Therefore, the number of inoperable RCCAs has been limited to one.

4.2.3.4.2 Control Rod Drive Mechanism Quality assurance procedures during production of CRDMs include material selection, process control, mechanism component tests during production, and hydrotests.

After all manufacturing procedures had been developed, several prototype CRDMs and drive rod assemblies were life tested with the entire drive line under environmental conditions of temperature, pressure, and flow. All acceptance tests were of duration equal to or greater than service required for the plant operation. All drive rod assemblies tested in this manner have shown minimal wear damage.

These tests include verification that the trip time achieved by the CRDMs met the design requirement of 2. 7 seconds from beginning of decay of stationary gripper coil voltage to dashpot entry. Trip time requirement will be confirmed for each CRDM at periodic intervals after initial reactor operation. In

addition, 4.2-79 SGS-UFSAR Revision 16 January 31, 1998

a Technical Specification has been set to ensure that the trip time requirement is met.

It is expected that all CRDMs will meet specified operating requirements for the duration of plant life with normal refurbishment. However, a Technical Specification pertaining to an inoperable RCCA has been set.

If an RCCA cannot be moved by its mechanism, adjustments in the boron concentration ensure that adequate shutdown margin would be achieved following a trip. Thus, inability to move one RCCA can be tolerated. More than one inoperable RCCA could be tolerated, but would impose additional demands on the plant operator. Therefore, the number of inoperable RCCAs has been limited to one.

In order to demonstrate continuous free movement of the RCCAs and to ensure acceptable core power distributions during operation, partial-movement checks are performed on every RCCA at least every 31 days during reactor critical operation. In addition, periodic drop tests of the full length RCCAs are performed at each refueling shutdown to demonstrate continued ability to meet trip time requirements, to ensure core subcri ticali ty after reactor trip, and to limit potential reactivity insertions from a hypothetical RCCA ejection.

During these tests the acceptable drop time of each assembly is not greater than 2.7 seconds, at full flow and operating temperature, from the beginning of decay of stationary gripper coil voltage to dashpot entry.

To confirm the mechanical adequacy of the fuel assembly and RCCA, functional test programs have been conducted on a full scale control rod. The prototype assembly was tested under simulated conditions of reactor temperature, pressure, and flow for approximately 1000 hours0.0116 days <br />0.278 hours <br />0.00165 weeks <br />3.805e-4 months <br />. The prototype mechanism accumulated about 3,000,000 steps and 600 trips. At the end of the test the CRDM was still operating satisfactorily. A correlation was developed to predict the amplitude of flow excited vibration of individual fuel rods and fuel assemblies. Inspection of the drive 4.2-80 SGS-UFSAR Revision 16 January 31, 1998

line components did not reveal significant fretting. The control rod free fall time against 125 percent of nominal flow was less than 1. 5 seconds to the dashpot; about 10 feet of travel.

Actual experience on the Ginna, Mihama No. 1, Point Beach No. 1, and H. B. Robinson plants indicates excellent performance of CRDMs.

All units are production tested prior to shipment to confirm ability of CRDMs to meet design specification-operational requirements. Periodic tests are also conducted during plant operation to confirm brake core operation.

During refueling, tests are also conducted to confirm condition to stator windings.

4.2.4 References for Section 4.2

1. Christensen, J. A.; Allio, R. J.; and Biancheria, A. "Melting Point of Irradiated uo ," WCAP-6065, February 1965.

2

2. Skari tka, J., "Operational Experience with Westinghouse Cores," WCAP-8183 (latest revision - updated annually) .

3. Eggleston, F. T., "Safety Related Research and Development for Westinghouse Pressurized Water Reactor - Program Summaries, Winter 1976 -

Summer 1978," WCAP-8768, Revision 2, October 1977.

4. Risher, D. H. (Ed.), "Safety Analysis for the Revised Fuel Rod Internal Pressure Design Basis," WCAP-8963-P-A, (Proprietary) and WCAP-8964-A, (Nonproprietary) August 1977.

5. Not used.

4.2-81 SGS-UFSAR Revision 22 May 5, 2006

6. Not used.

7. George, R. A., Lee, Y. C.; and Eng, G. H., "Revised Clad Flattening Model," WCAP-8377 (Proprietary) and WCAP-8381 (Nonproprietary), July 1974.

8. Watkins, B. and Wood, D. S. "The Significance of Irradiation - Induced Creep on Reactor Performance of a Zirca1oy-2 Pressure Tube",

Applications - Related Phenomena for Zirconium and its Alloys, ASTM STP 458, American Society for Testing and Materials, pp. 226-240, 1969.

9. O'Donnell, W. J. and Langer, B. F., "Fatigue Design Basis for Zircaloy Components," Nuclear Science and Engineering, 20, 1-12, 1964.

10. Gesinski, L.; Chiang, D.; and Nakazato, S., "Safety Analysis of the 17 x 17 Fuel Assembly For Combined Seismic and Loss-of-Coolant Accident,"

WCAP-8236 (Proprietary) and WCAP-8288, (Westinghouse Nonproprietary),

December 1973 and Addendum 1.

11. Demario, E. E. and Nakazato, S., "Hydraulic Flow Test of the 17 x 17 Fuel Assembly," WCAP-8278 (Proprietary) and WCAP-8279 (Nonproprietary),

February 1974.

12. NUREG-0641, "Control Rod Guide Tube Wear in Operating Reactors," U.S.

NRC, Division of Operating Reactors, April 1980.

13. Moore, J., "Nuclear Fuel Division Quality Assurance Program Plan," WCAP-7800, Revision 5A, November 1979.

14. Cohen, J., "Development and Properties of Silver Base Alloys as Control Rod Materials for Pressurized Water Reactors," WAPD-214, December 1959.

15. Davidson, S. L. (Ed.), et al, "Vantage 5 Fuel Assembly Reference Core Report," WCAP-10444-P-A, September 1985.

4.2-82 SGS-UFSAR Revision 25 October 26, 2010

16. Davidson, S. L. (Ed.), et al, "Vantage 5H Fuel Assembly," WCAP-10444-P-A, Addendum 2-A, February 1989.

17. Foster, J. P., Sidener, S., Westinghouse Improved Performance Analysis and Design Model (PAD 4.0), WCAP-15063-P-A, Revision 1, with Errata, July 2000.

18. Davidson, S. L. and Ryan, T. L., VANTAGE+ Fuel Assembly Reference Core Report, WCAP-12610-P-A, April 1995.

19. Letter from W. J. Rinkacs (Westinghouse) to M. M. Mannion (PSE&G),

Westinghouse Fuel Features Recommendation for Cycle 11, July 22, 1998

20. Letter from W. J. Rinkacs (Westinghouse) to T. K. Ross (PSE&G), Design Reviews for Fuel Feature Changes Proposed for Cycle 11, July 22, 1998

21. Letter from W. J. Rinkacs (Westinghouse) to M. M. Mannion (PSE&G),

Westinghouse Generic Safety Evaluation for the 17x17 Standard Robust Fuel Assembly, October 1, 1998

22. Letter from B. W. Gergos (Westinghouse) to T. K. Ross (PSE&G), Seismic/

LOCA Analysis of the Robust Fuel Assemblies, February 4, 1999

23. Iorii, J. A. and Petrarca, D. J., Westinghouse Wet Annular Burnable Absorber Evaluation Report, WCAP-10021-P-A, Revision 1, October 1983

24. Letter from I. R. Williamson (Westinghouse) to F. D. Rankowski (PSEG),

Extended Lifetime Wet Annular Burnable Absorber, NF-PSE-06-5, February 16, 2006.

25. Davidson, S.L. (Ed.), Westinghouse Fuel Criteria Evaluation Process, WCAP-12488-A, October 1994.

26. Sepp, H.A., Addendum 1 to WCAP-12488-A Revision to Design Criteria, WCAP-12488-A, Addendum 1-A, Revision 1, January 2002.

27. Garde, A., et al., Westinghouse Clad Corrosion Model for ZIRLO and Optimized ZIRLO, WCAP-12610-P-A & CENPD-404-P-A, Addendum 2-A, October 2013.

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

4.2-83 SGS-UFSAR Revision 30 May 11, 2018

TABLE 4.2-1 MAXIMUM DEFLECTIONS ALLOWED FOR REACTOR INTERNAL SUPPORT STRUCTURES No-Loss-of-Allowable Function Deflections Deflections Component (in.) (in.)

Upper Barrel radial inward 4.1 8.2 radial outward 0.5 1.0 Upper Package 0.10 0.15 Rod Cluster Guide Tubes 1.00 1.75

• SGS-UFSAR 1 of 1 Revision 6 February 15, 1987

TABLE 4.2-2 COMPARISON OF SINGLE AND DOUBLE ENCAPSULATED SECONDARY SOURCE DESIGNS PA.RAMETE:R SINGLE ENCAPSULA'J:ED DOUBLE ENCAPSt.T.LATED Number of rodlets 4 6 Outer Clad OD, in. 0.361 +I- 0.001 0.381 +I- 0.001 Outer Clad ID, in. 0.344 +!- 0.0005 0.344 +I- 0.0005 Inner Clad OD, in. N/A b.344 +I- 0.001 Inner Clad ID, in. N/A 0.297 +I- 0.0005 Pellet OD, in. 0.338 +0.002/-0.001 0.292 +/- 0.001 Pellet Stack Length, in. 88.00 86.00 Pellet Stack Weight, grams 500/535 338 +/- 10 Sp=in; Clip Material Carbon steel - plated 410 Stainless steel Outer Pressuriza~ion, psig 625 +/- 50 625 +/- 50 Inner ~ressurization, psig N/A 250 +/- 20 1 of 1 SGS-UFSAR Revision 18 April 26, 2000

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4.3 NUCLEAR DESIGN 4.3.1 Design Bases This section describes the design bases and functional requirements used in the nuclear design of the Fuel and Reactivity Control System and relates these design bases to the General Design Criteria (GDC) in 10CFRSO Appendix A. Where appropriate, supplemental criteria such as the Final Acceptance Criteria for Emergency Core Cooling Systems are addressed. Before discussing the nuclear design bases, it is appropriate to briefly review the four major categories ascribed to conditions of plant operation.

The full spectrum of plant conditions is divided into four categories, in accordance with the anticipated frequency of occurrence and risk to the public:

1. Condition I - Normal Operation,

2. Condition II - Incidents of Moderate Frequency,

3. condition III - Infrequent Faults,

4. Condition IV - Limiting Faults.

In general, the Condition I occurrences are accommodated with margin between any plant parameter and the value of that parameter which would require either automatic or manual protective action. condition II incidents are accommodated with, at most, a shutdown of the reactor with the plant capable of returning to operation after corrective action. Fuel damage* is not expected during Condition I and Condition II events. It is not possible, however,

• Fuel damage as used here is defined as penetration of the fission product barrier (i.e., the fuel rod clad).

4.3-1 SGS-UFSAR Revision 6 February 15, 1987

to preclude a very small number of rod failures. These are within the capability of the plant cleanup system and are consistent with the plant design basis.

Condition III incidents shall not cause more than a small fraction of the fuel elements in the reactor to be damaged, although sufficient fuel element damage might occur to preclude immediate resumption of operation. The release of radioactive material due to Condition III incidents should not be sufficient to interrupt or restrict public use of these areas beyond the exclusion radius.

Furthermore, a Condition III incident shall not, by itself, generate a Condition IV fault or result in a consequential loss of function of the Reactor Coolant System (RCS) or reactor containment barriers.

Condition IV occurrences are faults that are not expected to occur but are defined as limiting faults which must be designed against. Condition IV faults shall not cause a release of radioactive material that results in an undue risk to public health and safety.

The core design power distribution limits related to fuel integrity are met for Condition I occurrences throuqh conservative design and maintained by the action of the Control system. The requirements for Condition Il occurrences are met by providing an adequate protection system which monitors reactor parameters. The Control and Protection Systems are described in section 7, and the consequences of Condition II, III, and IV occurrences are given in Section 15.

4.3.1.1 Fuel Burnup Basis The fuel rod design basis is described in section 4.2. The nuclear design basis is to install sufficient reactivity in the fuel to attain the average region discharge burnup 4.3-2 SGS-UFSAR Revision 17 October 16, 1998

values given in reference 28. The above, along with the design basis in Section 4.3.1.3, Control of Power Distribution, satisfies GOC-10.

Discussion Fuel burnup is a measure of fuel depletion which represents the integrated energy output of the fuel (MWD/MTU} and is a convenient means for quantifying fuel exposure criteria.

The core design lifetime or design discharge burnup is achieved by installing sufficient initial excess reactivity in each fuel region and by following a fuel replacement program (such as that described in section 4.3.2) that meets all safety-related criteria in each cycle of operation.

Initial excess reactivity installed in the fuel, although not a design basis, must be sufficient to maintain core criticality at full power operating conditions throughout cycle life with equilibrium xenon, samarium, and other fission products present. The end-of-design cycle life is defined to occur when the chemical shim concentration is essentially zero with control rods present to the degree necessary for operational requirements. In terms of chemical shim boron concentration this represents approximately 10 ppm with no control rod insertion.

A limitation on initial installed excess reactivity is not required other than as is quantified in terms of other design bases such as core negative reactivity feedback and shutdown margin discussed below.

4.3.1.2 Negative Reactivity Feedbacks (Reactivity Coefficient)

The fuel temperature coefficient will be negative and the moderator temperature coefficient of reactivity will be non-positive for power operating conditions, thereby providing 4.3-3 SGS-UFSAR Revision 17 October 16, 1998

negative reactivity feedback characteristics. The design basis meets GDC-11.

Whe~ compensation for a rapid increase in is there are two or effects. These are the resonance absorption effects (Doppler) associated with changing fuel temperature and the spectrum effect resulting from changing moderator density. These basic physics characteristics are often identified by reactivity coefficients. The use of slightly enriched uranium ensures that the Doppler coefficient of reactivity is negative. This coefficient provides the most reactivity compensation. The core is also designed to have an overall moderator coefficient of so that average coolant or void content slower effect. Nominal power is only in a range of overall non-positive moderator coefficient. The non-positive moderator temperature coefficient can be achieved through use of fixed burnable absorber, integral fuel burnable absorber (IFBA) and/or control rods by limiting the reactivity held down by soluble boron.

Burnable absorber content (quantity and distribution) is not stated as a design basis other than as it relates to accomplishment of a non-positive moderator coefficient at power conditions discussed above.

4.3.1.3 Control of Power Distribution Basis The nuclear design basis is that, with at least a 95 percent confidence level:

1. The fuel will not be operated at greater than 13.3 kW/ft under normal conditions an allowance 4.3-4 SGS-UFSAR Revision 25 October 26, 2010

of 0.6 percent for calorimetric error and including densification effects.

2. Under abnormal conditions including the maximum overpressure condition, the fuel peak power will not cause melting as defined in Section 4. 4. 1.2.

3. The fuel will not operate with a power distribution that violates the departure from nucleate boiling (DNB) design basis (i.e., the DNB ratio I

(DNBR) shall not be less than the safety limit, as discussed in Section 4.4.1.1) under Condition I and II events including the maximum overpower condition.

4. Fuel management will be such as to produce rod powers and burnups consistent with the assumptions in the fuel rod mechanical integrity analysis of Section 4.2.

The above basis meets GDC-10.

Discussion Calculation of extreme power shapes which affect fuel design limits is performed with proven methods and verified frequently with measurements from operating reactors. The conditions under which limiting power shapes are assumed to occur are chosen conservatively with regard to any permissible operating state.

To ensure that the axial profile meets with the linear heat rate limit and the DNB limit, ex-core detector signals are used to provide a top to bottom flux difference, AI, which is input, through F{AI), into both the overpower AT and overtemperature AT trip points.

Even though there is good agreement between measured peak power calculations and measurements, a nuclear uncertainty margin is applied to calculated peak local power. Such a margin is provided 4.3-5 SGS-UFSAR Revision 20 May 6, 2003

both for the analysis of normal operating states and for anticipated transients.

4.3.1.4 Maximum Controlled Reactivity Insertion Rate The maximum reactivity insertion rate due to withdrawal of rod cluster control assemblies or by boron dilution is limited. This limit, expressed as a maximum reactivity change rate (75 pcmjsec)* is set such that peak heat generation rate and DNBR do not exceed the maximum allowable at overpower conditions. This satisfies GDC-25.

The maximum reactivity worth of control rods and the maximum rates of reactivity insertion employing control rods are limited so as to preclude rupture of the coolant pressure boundary or disruption of the core internals to a degree which would impair core cooling capacity due to a rod withdrawal or ejection accident (See Section 15).

Following any condition IV event (rod ejection, steamline break, etc.) the reactor can be brought to the shutdown condition and the core will maintain acceptable heat transfer geometry. This satisfies GDC-2B.

Discussion Reactivity addition associated with an accidental withdrawal of a control bank (or banks) is limited by the maximum rod speed (or travel rate) and by the worth of the bank(s). For this reactor the maximum control rod speed is 45 inches per minute and the maximum rate of reactivity change considering two control banks moving is less than 75 pcm per second.

5

• l pcm = 10 4p (See footnote Table 4.3-2) 4.3-6 SGS-UFSAR Revision 6 February 15 1 1987

4.3.1.5 Minimum shutdown margin as in the Technical is at any power operating condition in the hot shutdown condition and in the cold shutdown condition.

In all analyses involving reactor trip, the single, highest worth Rod Cluster Control Assembly (RCCA) is postulated to remain um::ripped in its full-out position (stuck rod criterion). This satisfies GDC-26.

Two independent control systems are provided, namely control rods and soluble boron in the coolant. The Control Rod System can compensate for the reactivity effects of the fuel and water temperature changes accompanying power level changes over the range from full-load to no-load, In addition, the Control Rod System provides the minimum shutdown margin under Condition I events and is capable of making the core subcritical rapidly enough to prevent exceeding fuel damage limits assuming that the worth control rod is stuck out upon

'I'he Boron can for all xenon burnout and will maintain the reactor in cold shutdown. Thus, backup and emergency shutdown provisions are provided by a mechanical and a chemical shim control system which satisfies GDC-26.

When fuel assemblies are in the pressure vessel and the vessel head is not in kef£ will be maintained at or below 0.95 with control rods and soluble boron. Further, the fuel will be 4.3-7 SGS-UFSAR Revision 25 October 26, 2010

maintained sufficiently subcritical that removal 1 of all RCCAs will not result in criticality.

ANS Standard NIB. 2 specifies a keff not to exceed 0. 95 in spent fuel storage racks and transfer equipment flooded with pure water and a keff not to exceed 0. 98 in normally dry new fuel storage racks assuming optimum moderation. No criterion is given for the refueling operation; however a 5 percent margin, which is consistent with spent fuel storage and transfer and 3 percent below the new fuel storage, is adequate for the controlled and continuously monitored operations involved.

4.3.1.6 Stability Basis The core will be inherently stable to power oscillations of the fundamental mode. This satisfies GDC-12.

Discussion Oscillations of the total power output of the core, from whatever cause, are readily detected by the loop temperature sensors and by the nuclear instrumentation. The core is protected by these systems and a reactor trip would occur if power increased unacceptably, preserving the design margins to fuel design limits.

The stability of the Turbine/Steam Generator/Core Systems and the Reactor Control System is such that total core power oscillations are not normally possible. The redundancy of the protection circuits ensures an extremely low probability of exceeding design power levels.

Basis Spatial power oscillations within the core, with a constant core power output, should they occur, can be reliably and readily detected and suppressed.

4.3-8 SGS-UFSAR Revision 6 February 15, 1987

Discussion The core is designed so that diametral oscillations due to spatial xenon effects are self-damping and no operator action or control action is required to suppress them. The stability of diametral oscillations is so great that this excitation is highly improbable. Convergent azimuthal oscillations can be excited by prohibited motion of individual control rods. Such oscillations are readily observable and alarmed, using the ex-core long ion chambers.

Indications are also continuously available from in-core thermocouples and loop temperature measurements. The BEACON core monitoring system (Reference 34) uses nearly continuous measurements of the excore detector ion chambers, core thermocouples, control rod indications, and loop temperature measurements for updating of the BEACON core model using the 3D-ANC (Advanced Nodal Code) neutronics solution (Reference 36) to provide nearly continuous power distribution monitoring. Moveable in-core detectors can be activated to provide more detailed and accurate information for calibration of the thermocouple data and measured spatial power distribution data used by BEACON.

If BEACON were to become inoperable, the in-core flux mapping process would be used in place of BEACON to monitor core power distribution. In all presently proposed cores these horizontal plane oscillations are self-damping by virtue of reactivity feedback effects designed into the core.

However, axial xenon spatial power oscillations may occur late in core life.

The control bank and ex-core detectors are provided for control and monitoring of axial power distributions. Assurance that fuel design limits are not exceeded is provided by reactor overpower ~T and overtemperature ~T trip functions which use the measured axial power imbalance as an input.

4.3.1.7 Anticipated Transients Without Trip The effects of anticipated transients with failure to trip are not considered in the design bases of the plant. Analysis has shown that the likelihood of such a hypothetical event is negligibly small. Furthermore, analysis of the consequences of a hypothetical failure to trip following anticipated transients has shown that no significant core damage would result and system peak pressures would be limited to acceptable values and no failure of the RCS would result (1).

4.3-9 SGS-UFSAR Revision 27 November 25, 2013

4.3.2 Description The reactor cores consist of a specified number of fuel rods which are held in bundles by spacer grids and top and bottom The fuel rods are constructed o.:f .

Z~Lrca ..J oy or . l o TM Zlr cy l '~Ln d ' l n.ca tubes 00 fuel 2

pellets. 'l.'he bundleR, known as fuel as.<;;emblies, are arranged in a pattern which approximates a circular cylinder.

E:ach fuel assembly contains a 17 :x: n rod array composed of 264 fuel rods, 24 rod cluster control {RCC) th:l.mbles and an in-cor.:*e instrumentation thimble.

Figure 4. 2-1 shows a cross sectional view of a 17 x 17 fuel assembly and the related RCC locations. Further details of the fuel assembly are given in Section 4.2.1.

For initial core loading, the fuel rods within a given assembly have the same uranium enrichrnen*t in both the .r.adial and cndal J<'uel af.;scmblies of three different enrichments are used in the initial core loadings to establish a favorable radial power cU.stribution. 1~wo regions consisting of two lower enrichments are so as to form a checkerboard pattern in the central portion of the core. The third region is arranged around the periphery of the core and contains the highest enrjchment. The enrichments for the first cores are shown in Table 4.3-1.

A ~ypical reJ.oad pattern is a low leakage loading pattern or a low-low leakage A low loading pattern has either burned (depleted) or fresh

(.feed) assemblies arranged on the core periphery. A low-low leakage pat tern has only burned (depleted) assemblies on the core periphery. The reload cores a:re normally designed to be able to operate approximately eighteen months between refuelings. The typical loading patterns for both Salem Units are shown in Figures 4.5-1 and 4.5-3.

SGS-UE'SAR 4.3-10 Revision 23 October 17, 2007

• The core average enrichment is determined by the amount of fissionable material required to provide the desired cycle energy requirements. The physics of the burnout process is such that operation of the reactor the amount of fuel available due to the of neutrons by the U-235 atoms and their fission. The rate of U-235 is proportional to the power level at which the reactor is In addition, the fission process results in the formation of fission products, some of which absorb neutrons. These effects, depletion and the buildup of fission products, are partially offset by the buildup of plutonium shown on Figure 4.3-1 for the 17 x 17 fuel assembly, which occurs due to t~e non-fission absorption of neutrons in U-238. Therefore, at the beginning of any cycle a reactivity reserve to the depletion of the fissionable fuel and the buildup of fission over the life must be "built" into the reactor. This excess is controlled by boron dissolved 1n the coolant and burnable absorbers.

'::'he concentration of boric acid in the primary coolant is varied to provide control and to compensate for long-term reactivity requirements. The concentration of the soluble neutron absorber is varied c:o compensate for reactivity changes due to fuel burnup, fission product poisoning including xenon and samarium, burnable absorber depletion, and the cold-to-operating moderator temperature Using its normal makeup path, the Chemical and Volume Control (CVCS) is of at a rate of 30 when the reactor coolant boron concentration is 1000 ppm and approximately 35 pcm/min when the reactor coolant boron concentration is 100 pprn. The peak burnout rate for xenon is 25 pcm/min. Rapid transient reactivity requirements and safety shutdown requirements are me~ with control rods.

As the boron concentration is increased, the moderator ~emperature coefficient becomes less The use of a soluble absorber alone would result in a moderator coefficient at (BOL). Therefore, burnable absorbers are used to reduce the soluble boron concentration to ensure that the moderator temperature coefficient is for 9ower operating conditions. During the absorber content in the burnable absorbers is depleted thus adding positive reactivity to offset some of the negative reactivity from fuel depletion and fission product build'Jp. The

4. 3-11 SGS-UFSAR Revision 25 October 26, 2010

depletion rate of the burnable absorbers is not critical since chemical shim is always available and flexible enough to cover any possible deviations in the expected burnable absorber depletion rate. Figure 4.3-2 is a graph of a typical core depletion with burnable absorbers.

As stated previously, the purpose of burnable absorbers (integral and discrete) is to provide enough boron loading in the core to decrease core soluble boron concentrations to the point that a negative moderator temperature coefficient is maintained at all hot operating conditions. IFBA boron loadings in the range from 1. 57 to 2. 35 mg/inch per rod is typical. Two different discrete burnable absorbers (borosilicate glass-PYREX) and Wet Annular Burnable Absorbers (WABA) may be used. Both discrete absorbers have similar burnout characteristics and provide similar benefits to core design development which include reduction of the number and loading of the IFBA, which is a fuel performance (rod internal pressure, corrosion, and DNB propagation) benefit.

Typical stack lengths for IFBA and WABA range from 108 to 132 inches. PYREX rods are 139 inches in length. Design data for IFBA, WABA, an PYREX can be seen in Table 4. 3-1. Reference 32 of this section provides a more detailed discussion of WABA.

In addition to reactivity control, the burnable absorbers, both discrete {WABA and/or PYREX) and integral (IFBA), are strategically located to provide a favorable radial power distribution. Reload loading patterns utilize various I

burnable absorber types and distributions. These are determined on a cycle-specific basis. The typical burnable absorber loading patterns are shown in Figures 4.5-2 and 4.5-4.

Tables 4.3-1 through 4.3-3 contain a summary of the reactor core design parameters for a typical fuel cycle, including reactivity coefficients, delayed neutron fraction, and neutron lifetimes. Some of these parameters change on a reload basis. The current values or allowable range of values may be found in the appropriate Reload Safety Evaluation (RSE)/Safety Assessment (SA) or Nuclear Design Report (NDR)/Curvebook (CB)/Plant Operations Package (POP), see Section 4.5. The RSE/SA is a cycle-specific document which evaluates the fuel assembly design and core loading pattern configuration for plant safety. The NDR/CB/POP are cycle-specific documents which include such information as power distributions, reactivity coefficients, boron and control rod worths under a variety of plant operating conditions. There is sufficient information in the RSE/SA and NDR/CB/POP to permit an independent calculation of the nuclear performance characteristics of the core.

4.3-12 SGS-UFSAR Revision 23 October 17, 2007

4.3.2.2 Power Distribution The accuracy of power distribution calculations has been confirmed through approximately one thousand flux maps during some 20 years of operation under conditions very similar to those expected for the plant described herein.

Details of this confirmation are given in Reference 2 and in Section 4.3.2.2.7.

4.3.2.2.1 Definitions Power distributions are quantified in terms of hot channel factors. These factors are a measure of the peak pellet power within the reactor core and the total energy produced in a coolant channel and are expressed in terms of quantities related to the nuclear or thermal design, namely:

Power density is the thermal power produced per unit volume of the core (kW/liter).

Linear oower density is the thermal power produced per unit length of active fuel (kW/ft). Since fuel assembly geometry is standardized, this is the unit of power density most commonly used. For all practical purposes it differs from kW/liter by a constant factor which includes geometry effects and the fraction of the total thermal power which is generated in the fuel rod.

Average linear power density is the total thermal power produced in the fuel rods divided by the total active fuel length of all rods in the core.

2 1 Local heat flux is the heat flux at the surface of the cladding (Btu-ft- -hr- ).

For nominal rod parameters this differs from linear power density by a constant factor.

Rod power or rod integral power is the length integrated linear power density in one rod (kW).

4.3-13

, SGS-UFSAR Revision 17 October 16, 1998

Average rod power is the total thermal power produced in the fuel rods divided by the number of fuel rods (assuming all rods have equal length).

The hot channel factors used in the discussion of power distributions in this '-

section are defined as follows:

FQ, Heat Flux Hot Channel Factor, is defined as the maximum local heat flux on the surface of a fuel rod divided by the average fuel rod heat flux, allowing for manufacturing tolerances on fuel pellets and rods.

FNQ, Nuclear Heat Flux Hot Channel Factor, is defined as the maximum local fuel rod linear power density divided by the average fuel rod linear power density, assuming nominal fuel pellet and rod parameters.

E F Q, Engineering Heat Flux Hot Channel Factor. is the allowance on heat flux required for manufacturing tolerances. The engineering factor allows for local variations in enrichment, pellet density and diameter, surface area of the fuel rod, and eccentricity of the gap between pellet and clad.

Combined statistically the net effect is a factor of 1.03 to be applied to fuel rod surface heat flux.

N F 4H, Nuclear Enthalpy Rise Hot Channel Factor. is defined as the ratio of the integral of linear power along the rod with the highest integrated power to the average rod power.

Manufacturing tolerances, hot channel power distribution and surrounding channel power distributions are treated explicitly in the calculation of the departure from nucleate boiling ratio described in Section 4.4.

4. 3-14 SGS-UFSAR Revision 6 February 15, 1987

It is convenient for the purposes of discussion to define sub factors of F ;

0 however, design limits are set in terms of the total peaking factor .

F0 Total peaking factor or heat flux hot-channel factor Maximum kW/ft Average kW/ft without densification effects where:

F N and FE are defined above.

Q Q F~ the measurement uncertainty; 1.05 if a full core flux map was performed with the movable detectors.

u )

(1.0+-Q- if BEACON (PDMS) was used for power distribution 100.0 monitoring; where U0 is defined by Equation 5-19 of Reference 34 for the use of BEACON.

Note: All peaking factor uncertainties for power distribution measurements are located in the Core Operating Limits Report for each specific cycle.

N F XY ratio of peak power density to average power density in the horizontal plane of peak local power .

• SGS-UFSAR 4.3-15 Revision 19 November 19, 2001

P(Z) = ratio of the power per unit core height in the horizontal plane at height Z to the average value of power per unit core height.

4.3.2.2.2 Radial Power Distribution The power shape in horizontal sections of the core at full power is a function of the fuel and burnable absorber loading patterns and the presence or absence of a single bank of control rods. Thus, at any time in the cycle a horizontal section of the core can be characterized as unrodded or with group D control rods. These two situations combined with burnup effects determine the radial power shapes which can exist in the core at full power. The effect on radial power shapes of power level, xenon, samarium, and moderator density effects are considered also but these are quite small. The effect of non-uniform flow distribution is negligible. While radial power distributions in various planes of the core are often illustrated, the core radial enthalpy rise distribution as determined by the integral of power up each channel is of greater interest.

Figures 4.3-3 through 4.3-7 show 4.3-16

. SGS-UFSAR Revision 17 October 16, 1998

representative low leakage radial power distributions for one eighth of the core for representative operating conditions. These conditions are ( 1) Hot Full Power (HFP) at BOL - unrodded - no xenon, (2) HFP at BOL - unrodded - equilibrium xenon, {3) HFP at BOt - Bank D in - equilibrium xenon, (4) HFP at Middle-of-Life (MOL) - unrodded - equilibrium xenon, and (5) HFP at EOL - unrodded - equilibrium xenon.

The radial power distribution, and hence radial enthalpy rise distribution, changes on a cycle-specific basis. The appropriate NOR should be referenced.

see section 4.5 for the current cycles* predicted radial enthalpy distribution.

Since the position of the hot channel varies from time to time a single reference radial design power distribution is selected for DNB calculations. This reference power distribution is chosen conservatively to concentrate power in one area of the core, minimizing the benefits of flow redistribution. Assembly powers are normalized to core average power.

4.3.2.2.3 Assembly Power Distribution since the detailed power distribution surrounding the hot channel varies based on within-assembly design features and time in life, a conservatively flat assembly power distribution is assumed in the DNB analysis, described in Section 4.4, with the rod of maximum integrated power artificially raised to the design value of F:a* The design value of F~K is determined during the RSE. The current cycles* F~K ~alue is given in Section 4.5. Care is taken in the nuclear design of all fuel cycles and all operating conditions to ensure that a flatter assembly power distribution does not occur with limiting values to F~H*

4.3.2.2.4 Axial Power Distribution The shape of the power profile in the axial or vertical direction is largely under the control of the operator through either the manual operation of the control rods or automatic motion of control rods responding to manual operation of the eves. Nuclear 4.3-17 SGS-UFSAR Revision 17 October 16, 1998

effects which cause variations in the axial power shape include moderator density, Doppler effect on resonance absorption, spatial xenon, burnable absorbers, and burnup. Automatically controlled variations in total power output and control rod motion are also important in determining the axial power shape at any time. Signals are available to the operator from the ex-core ion chambers which are long ion chambers outside the reactor vessel running parallel to the axis of the core. Separate signals are taken from the top and bottom halves of the chambers. The difference between top and bottom signals from each of four pairs of detectors is displayed on the control panel and called the flux difference, 4!. Calculations of core average peaking factor for many plants and measurements from operating plants under many operating situations are associated with either 41 or axial offset in such a way that an upper bound can be placed on the peaking factor. For these correlations axial offset is defined as:

axial offset and ~t and ~b are the top and bottom detector readings.

Representative axial power shapes from Reference 3 for BOL, MOL, and EOL conditions are shown on Figures 4.3-8 through 4.3-10.

These figures cover a wide range of axial offset including values not permitted at full power.

4.3.2.2.5 Deleted 4.3-18 SGS-UFSAR Revision 17 October 16, 1998

THIS PAGE INTENTIONALLY BLANK 4.3-19 SGS-UFSAR Revision 17 October 16, 1998

4.3.2.2.6 Limiting Power Distribution According to the ANS classifications of plant conditions (see section 15),

condition I occurrences are those which are expected frequently or regularly in the course of power operation, maintenance, or maneuvering of the plant. As such, Condition I occurrences are accommodated with margin between any plant parameter and the value of that parameter which would require either automatic or manual protective action. In as much as Condition I occurrences occur frequently or regularly, they must be considered from the point of view of affecting the consequences of fault conditions (Conditions II, III, and IV). In this regard, analysis of each fault condition described is generally based on a conservative set of initial conditions corresponding to the most adverse set of conditions which can occur during Condition I operation.

The list of steady state and shutdown conditions, permissible deviations (such as one coolant loop out of service) and 4.3-20 SGS-UFSAR Revision 17 October 16, 1998

operational transients is given in Section 15. Implicit in the definition of normal operation is proper and timely action by the reactor operator. That is, the operator follows recommended operating procedures for maintaining appropriate power distributions and takes any necessary remedial action when alerted to do so by the plant instrumentation. Thus, as stated above, the worst or limiting power distribution which can occur during normal operation is to be considered as the starting point for analysis of Condition II, III, and IV events.

Improper procedural actions or errors by the operator are assumed in the design as occurrences of moderate frequency (Condition II). Some of the consequences which might result are listed in Section 15.1. Therefore, the limiting power shapes which result from such Condition I I events are those power shapes which deviate from the normal operating condition at the recommended axial offset band, e.g., due to lack of proper action by the operator during a xenon transient following a change in power level brought about by control rod motion. Power shapes which fall in this category are used for determination of the reactor protection system setpoints so as to maintain margin to overpower or DNB limits.

The means for maintaining power distributions within the required hot channel factor limits are described in the core Operating Limit Report ( COLR). A complete discussion of power distribution control in Westinghouse Pressurized Water Reactors (PWRs) is included in Reference 5. Detailed information on the design constraints on local power density in a Westinghouse PWR, on the defined operating procedures and on the measures taken to preclude exceeding design limits is presented in the Westinghouse topical report on peaking factors (Reference 6). The following paragraphs summarize these reports and describe the calculations used to establish the upper bound on peaking factors.

The calculations used to establish the upper bound on peaking factors, FQ and F~, include all of the nuclear effects which 4.3-21 SGS-UFSAR Revision 17 October 16, 1998

influence the radial and/or axial power distributions throughout core life for various modes of operation including load follow, reduced power operation, and axial xenon transients.

Radial power distributions are calculated for the full power condition, and fuel and moderator temperature feedback effects are included for the average enthalpy plane of the reactor. The steady state nuclear design calculations are done for normal flow with the same mass flow in each channel and flow redistribution effects neglected. The effect of flow redistribution is calculated explicitly where it is important in the DNB analysis of accidents. The effect of xenon on radial power distribution is small (compare Figures 4.3-3 and 4.3-4, respectively) but is included as part of the normal design process. Radial power distributions are relatively fixed and easily bounded with upper limits.

The core average axial profile, however, can experience significant changes which can occur rapidly as a result of rod motion and load changes and more slowly due to xenon distribution. For the study of points of closest approach to axial power distribution limits, several thousand cases are examined. Since the properties of the nuclear design dictate what axial shapes can occur, boundaries on the limits of interest can be set in terms of the parameters which are readily observed on the plant. Specifically, the nuclear design parameters which are significant to the axial power distribution analysis are:

1. Core power level

2. Core height

3. Coolant temperature and flow

4. Coolant temperature program as a function of reactor power 4.3-22 SGS-UFSAR Revision 17 October 16, 1998

5. Fuel cycle lifetimes

6. Rod bank worths

7. Rod bank overlaps

8. Burnable absorber length and placement Normal operation of the plant assumes compliance with the following conditions:

1. Control rods in a single ba~k move together with no individual rod insertion differing by more than 12 steps above 85% RTP or more than I

18 steps at or below 85% RTP from the bank demand position.

Reference 35 documents Salem specific analysis performed to allow up to an 18 step rod misalignment at or below 85% RTP conditions.

2. Control banks are sequenced with overlapping banks.

3. The control bank insertion limits are not violated.

4, Axial power distribution procedures, which are given in terms of flux difference control and control bank position, are observed .

The axial power distribution procedures referred to above are part of the required operating procedures which are followed in normal operation. They require control of the axial offset {flux difference divided by fractional power) at all power levels within a permissible operating band of a target value corresponding to the equilibri~ full power value. In the first cycle, the target value changes from about -10 percent to 0 percent linearly through the life of the cycle. Target values in a reload cycle vary based on previous I

cycle length and number of fresh assemblies. These cycle-specific target values are available in the appropriate NDR (see Section 4.5). This minimizes xenon transient effects on the axial power distribution since the procedures essentially keep the xenon distribution in phase with the power distribution.

Calculations are performed for normal operation of the reactor including load following maneuvers. Beginning, middle, and end-of-cycle conditions are included in the calculations. Different histories of operation are assumed prior to calculating the effect of load follow transients on the axial power

• SGS-UFSAR 4.3-23 Revision 23 October 17, 2007

distribution. These different histories assume base loaded operation and extensive load following. The calculated points have been synthesized from axial calculations combined with radial factors appropriate for rodded and unrodded planes. The calculated values have been increased by a factor of 1.05 for conservatism and a factor of 1.03 for the engineering factor, F~

Figure 4. 3-11 represents these results as an upper bound envelope on local power density versus elevation in the core. It should be emphasized that this envelope is a conservative representation of the bounding values of local power density. Expected values are considerably smaller and, in fact, less conservative bounding values may be justified with additional analysis or surveillance requirements.

Finally, as previously discussed, this upper bound envelope is based on procedures of load follow which require the operator to operate within an allowed deviation from a target equilibrium value of axial flux difference, observing certain D bank insertion limits. These procedures are detailed in Technical Specifications and are predicated only upon ex-core surveillance supplemented by the normal monthly full core map requirement and by computer based alarms on deviation and time of deviation from the allowed flux difference band.

Allowing for fuel densification effects, the average kW/ft for both Units 1 and 2 is 5.52. From Figure 4.3-11, the conservative upper bound value of normalized local power density, including allowances for densification effects, is 2. 40 corresponding to a peak local power density of 13.3 kW/ft at 100.6 percent power for Units 1 and 2.

4.3-24 SGS-UFSAR Revision 20 May 6, 2003

To determine Reactor Protection System setpoints, with respect to power distributions, three categories of events are considered, namely rod control equipment malfunctions, operator errors of commission, and operator errors of omission.

The first category comprises uncontrolled rod withdrawal (with rods moving in the normal bank sequence) for full length rod banks. Also included are motions of the full length rod banks below their insertion limits, which could be caused, for example, by uncontrolled dilution or primary coolant cooldown.

Power distributions were calculated throughout these occurrences assuming short-term corrective action, that is no transient xenon effects were considered to result from the malfunction. The event was assumed to occur from typical normal operating situations which did include normal xenon transients.

It was further assumed in determining the power distributions that total power level would be limited by reactor trip to below 118 percent. Since the study is to determine protection limits with respect to power and axial offset, no credit was taken for trip setpoint reduction due to flux difference. Results are given on Figure 4.3-12 in units of kW/ft. The peak power density which can occur in such events, assuming reactor trip at or below 118 percent, is less than that required for centerline melt including uncertainties and densification effects. The second category, also appearing in Figure 4.3-12, assumes that the operator mis-positions the full length rod bank in violation of the insertion limits and creates short-term conditions not included in normal operating conditions.

The third category assumes that the operator fails to take action to correct a flux difference violation. Representative results shown on Figure 4.3-13 are FQ multiplied by 100.6 percent power including an allowance for calorimetric error. The figure shows that provided the assumed error in operation does not continue for a period which is long compared to the xenon time constant, the maximum 4.3-25 SGS-UFSAR Revision 19 November 19, 2001

local power does not exceed 22.4 kW/ft including the above factors. However, the COLR restrict AI at 100. 6 percent power such that the peak linear power density is less than 18 kW/ft. These events are considered Condition II events.

It should be noted that a reactor overpower accident is not assumed to occur coincident with an independent operator error.

Analyses of possible operating power shapes show that the appropriate hot channel factors FQ and F~ for peak local power density and for DNB analysis at full power are the values given in the COLR. The current unit and cycle's COLR reference is given in Section 4.5. Typical values are given in Table 4.3-2.

FQ can be increased with decreasing power as shown in the Technical Specifications. Increasing F~ with decreasing power is permitted by the DNB protection setpoints and allows radial power shape changes with rod insertion to the insertion limits as described in Section 4.4.3.2. It has been detsrmined that provided the above Conditions I through IV are observed, the Technical Specification limits are met.

When a situation is possible in normal operation which could result in local power densities in excess of those assumed as the pre-condition for a subsequent hypothetical accident, but which would not itself cause fuel failure, administrative controls and alarms are provided for returning the core to a safe condition. These alarms are described in detail in Sections 7 and 16.

4.3.2.2.7 Experimental Verification of Power Distribution Analysis This subject is discussed in depth in Reference 2. A summary of this report is given here.

In a measurement of peak local power density, FQ' with the moveable detector system described in Section 7.6, the following uncertainties have to be considered:

4.3-26 SGS-UFSAR Revision 19 November 19, 2001

1. Reproducibility of the measured signal

2. Errors in the calculated relationship between detector current and local flux

3. Errors in the calculated relationship between detector flux and peak rod power some distance from the measurement thimble.

The appropriate allowance for 1 above has been quantified by repetitive measurements made with several inter-calibrated detectors by using the common thimble features of the In-core Detector System. This system allows more than one detector to access any thimble. Errors in category 2 above are quantified to the extent possible, by using the fluxes measured at one thimble location to predict fluxes at another location which is also measured. Local power distribution predictions are verified in critical experiments on arrays of rods with simulated guide thimbles, control rods, burnable poisons, etc. These critical experiments provide quantification of error of types 2 and 3 above.

Reference 2 describes critical experiments performed at the Westinghouse Reactor Evaluation Center and measurements taken on two Westinghouse plants with In-core Detector Systems of the same type as used in the plant described herein. The report concludes that the uncertainty associated with the peak nuclear heat flux factor, FQ' is 4. 58 percent at the 95 percent confidence level with only 5 percent of the measurements greater than the inferred value.

This is the equivalent of a 2cr limit on a normal distribution and is the uncertainty to be associated with a full core flux map with moveable detectors reduced with a reasonable set of input data incorporating the influence of burnup on the radial power distribution. The uncertainty is usually rounded up to 5 percent.

For use of the PDMS (BEACON) system, the measurement uncertainty {0 0 ) is defined in Equation 5-19 of Reference 34. On a cycle specific basis, peaking factor measurement uncertainties can be found in the COLR. The BEACON measured power distribution must be calibrated to flux map measurements during the initial ascension in power above 25 % power and at 180 EFPD intervals from the initial calibration flux map .

• SGS-UFSAR 4.3-27 Revision 19 November 19, 2001

In comparing measured power distributions (or detector currents) against the calculations for the same situation, it is not possible to subtract out the detector reproducibility. Thus a comparison between measured and predicted power distributions has to include some measurement error. Such a comparison is given on Figure 4.3-14 for one of the maps used in Reference 2. Since the first publication of the report, hundreds of maps have been taken on these and other reactors. The results confirm the adequacy of the 5 percent uncertainty allowance on the calculated F .

0 A similar analysis for the uncertainty in FAH (rod integral power) measurements I using the

3. 60 percent in-core at the movable detector system results equivalent of a 20' confidence in level.

an allowance of For historical reasons, an 8 percent uncertainty factor is allowed in the nuclear design basis; that is, the predicted rod integrals at full power must not exceed the Technical Specification I COLR FNAH limit less 8 percent. This 8 percent may be reduced in final design to 4 percent to allow a wider range of acceptable axial power distributions in the DNB analysis and still meet the design bases of Section 4.3.1.3.

For use of the PDMS (BEACON) system, the F~ measurement uncertainty (UAHl is defined in Equation 5-19 of Reference 34. On a cycle specific basis, peaking factor measurement uncertainties can be found in the COLR. The BEACON measured power distribution must be calibrated to flux map measurements during the initial ascension in power above 25 % power and at 180 EFPD intervals from the initial calibration flux map.

A measurement in the second cycle of a 121 assembly, 12 foot core is compared with a simplified one dimensional core average axial calculation on Figure 4. 3-15. This calculation does not give explicit representation to the fuel grids.

The accumulated data on power distributions in actual operation is basically of three types:

1. Much of the data is obtained in steady state operation at constant power in the normal operating configuration.

2. Data with unusual values of axial offset have been obtained in the past as part of the multi-point ex-core detector calibrqtion exercise which is performed monthly.

3. Special tests have been performed in load follow and other transient xenon conditions which have yielded useful information on power distributions.

4.3-28 SGS-UFSAR Revision 19 November 19, 2001

These data are presented in detail in Reference 6. Figure 4.3-16 contains a summary of measured values of FQ as a function of axial offset for five plants from that report.

4.3.2.2.8 Testing A very extensive series of physics tests is performed on first cores. These tests are described in Section 14. Since not all limiting situations can be created at BOL, the main purpose of the tests is to provide a check on the calculational methods used in the predictions for the conditions of the test.

Low power physics tests are performed at the beginning of each reload cycle to confirm the validity of the cycle-specific design models that are used in the calculations supporting the Reload Safety Evaluation. The physics tests measure the critical boron concentration, isothermal temperature coefficient and control rod worth and compare the results to predictions calculated by the design models. Additionally, moveable in-core detectors are used to measure nuclear and safety related parameters during the initial power ascension and periodically throughout the cycle operation. The measurements are used to ensure that the measured parameters are within the limiting values contained in the Technical Specifications, as well as to compare the corre.sponding parameters to the design models for validation .

4.3.2.2.9 Monitoring Instrumentation The adequacy of instrument numbers, spatial deployment, required correlations between readings and peaking factors, calibration and errors are described in References 2, 5, and 6. The relevant conclusions are summarized here in Section 4.3.2.2.7.

Provided the limitations given in Section 4.3.2.2.6 on rod insertion and flux difference are observed, the Ex-core Detector System provides adequate monitoring of power distributions.

The addition of BEACON {PDMS) provides nearly continuous on-line monitoring of the core power distribution through thermocouple measurements, rod position indications, loop temperature measurements, and excore detector readings which are fed from the plant computer. These readings are calibrated to in-core flux map data during the initial power ascension above 25 % power conditions and at 180 EFPD intervals thereafter.

Further details of specific limits on the observed rod positions and flux difference are given in the COLR (See Section 4.5).

Limits for alarms, reactor trip, etc. are given in the Technical Specifications. Descriptions of the systems provided are given in Section 7.7 .

4.3-29 SGS-UFSAR Revision 19 November 19, 2001

4.3.2.3 Reactivity Coefficients The kinetic characteristics of the reactor core determine the response of the core to changing plant conditions or to operator adjustments made during normal operation, as well as the core response during abnormal or accidental transients. These kinetic characteristics are quantified in reactivity coefficients. The reactivity coefficients reflect the changes in the neutron multiplication due to varying plant conditions such as power, moderator, or fuel temperatures, or less significantly due to a change in pressure or void conditions. Since reactivity coefficients change during the life of the core, ranges of coefficients are employed in transient analysis to determine the response of the plant throughout life. The results of such simulations and the reactivity coefficients used are presented in Section 15. The analytical methods and calculational models used in calculating the reactivity coefficients are given in Section 4. 3. 3. These models have been confirmed through extensive testing of more than thirty cores similar to the plant described herein; results of these tests are discussed in Section 4.3.3.

Quantitative information for calculated reactivity coefficients, including fuel Doppler coefficient, moderator coefficients (density, temperature, pressure, void) and power coefficient is given in the following sections.

4.3.2.3.1 Fuel Temperature (Doppler) Coefficient The fuel temperature (Doppler) coefficient is defined as the change in reactivity per degree change in effective fuel temperature and is primarily a measure of the Doppler broadening of U-238 and Pu-240 resonance absorption peaks. Doppler broadening of other isotopes such as 0-236, Np-237 etc. are also considered but their contributions to the Doppler effect is small. An increase in fuel temperature increases the effective resonance absorption cross sections of the fuel and produces a corresponding reduction in reactivity.

SGS-OSFAR 4.3-30 Revision 6 February 15, 1987

The fuel temperature coefficient is calculated by performing two-group three-dimensional calculations using ANC (Reference 31). The fuel temperature coefficient is calculated by subtracting the MTC from the ITC from HZP to HFP.

A typical Doppler temperature coefficient is shown on Figure 4. 3-17 as a function of the effective fuel temperature (at SOL and EOL conditions). The effective fuel temperature is lower than the volume averaged fuel temperature since the neutron flux distribution is non-uniform through the pellet and gives preferential weight to the surface temperature. A typical Doppler-only contribution to the power coefficient, defined later, is shown on Figure 4.3-18 as a function of relative core power. The integral of the differential curve on Figure 4.3-18 is the Doppler contribution to the power defect and is shown on Figure 4.3-19 as a function of relative power. The Doppler-only power coefficient and defects were calculated using three-dimensional ANC (Reference 31). The Doppler coefficient becomes more negative as a function of life as the Pu-240 content increases, thus increasing the Pu-240 resonance absorption but less negative as the fuel temperature changes with burnup as described in Section 4.3.3.1. The upper and lower limits of Doppler coefficient used in accident analyses are given in Section 15. The Doppler-only coefficient and defect change slightly on a cycle-specific basis. The appropriate NDR should be referenced for the current cycle's information (see Section 4.5).

4.3.2.3.2 Moderator Coefficients The moderator coefficient is a measure of the change in reactivity due to a change in specific coolant parameters such as density, temperature, pressure, or void. The coefficients so obtained are moderator density, temperature, pressure, and void coefficients.

Moderator Density and Temperature Coefficients The moderator temperature (density) coefficient is defined as the change in reactivity per unit change in the moderator temperature (density}. Generally, the effect of the changes in moderator 4.3-31 SGS-UFSAR Revision 17 October 16, 1998

density as well as the temperature are considered together. An increase in moderator density results in more moderation and hence an increase in reactivity.

Therefore, the moderator density coefficient is positive. As temperature increases, density decreases (for a constant pressure); hence the moderator temperature coefficient becomes more negative. An increase in coolant temperature, keeping the density constant, leads to a hardened neutron spectrum and results in an increase in resonance absorption in U-238, Pu-240 and other isotopes. The hardened spectrum also causes a decrease in the fission-to-capture ratio in U-235 and Pu-239. Both of these effects make the moderator temperature coefficient more negative. Since water density changes more rapidly with temperature as temperature increases, the moderator temperature (density) coefficient becomes more negative (positive) with increasing temperature.

The soluble boron used in the reactor as a means of reactivity control also has an effect on moderator density coefficient since the soluble boron poison density as well as the water density is decreased when the coolant temperature rises.

A decrease in the soluble poison concentration introduces a positive component in the moderator temperature coefficient.

Thus, if the concentration of soluble poison is large enough, the net value of the coefficient may be positive. With the burnable absorbers present, however, the initial hot boron concentration is sufficiently low that the moderator temperature coefficient is negative at operating temperatures. The effect of control rods is to make the moderator coefficient more negative by reducing the required soluble box:,on concentration and by increasing the "leakage" of the core.

With burnup, the moderator temperature coefficient becomes more negative primarily as a result of boric acid dilution but also to an extent from the effects of the buildup of plutonium and fission products.

4.3-32 SGS-UFSAR Revision 17 October 16, 1998

The moderator coefficient is calculated for the various plant conditions discussed above by performing two-group three-dimensional calculations, varying the moderator temperature (and density) by about +/- 5°F about each of the mean temperatures. The moderator temperature coefficient is shown as a function of core temperature and boron concentration for the unrodded and rodded core on Figures 4. 3-20 through 4. 3-22. The temperature range covered is from cold ( 68°F) to about 600°F. The contribution due to Doppler coefficient (because of change in moderator temperature) has been subtracted from these results. Figure 4. 3-23 shows the hot, full power moderator temperature coefficient plotted as a function of first cycle lifetime for the just critical boron concentration condition based on the design boron letdown condition.

The moderator coefficients presented here are calculated on a corewise basis, since they are used to describe the core behavior in normal and accident situations when the moderator temperature changes can be considered to affect the whole core. Moderator temperature coefficients change on a cycle-specific basis.

The appropriate NOR should be referenced for the current cycle' s*information (see section 4.5).

Moderator Pressure Coefficient The moderator pressure coefficient relates the change in moderator density, resulting from a reactor coolant pressure change, to the corresponding effect an neutron production. This coefficient is of much less significance in comparison with the moderator temperature coefficient. A change of 50 psi in pressure has approximately the same effect on reactivity as a half-degree change in moderator temperature. This coefficient can be determined from the moderator temperature coefficient by relating change in pressure to the corresponding change in density. The moderator pressure coefficient is negative aver a portion of the moderator temperature range at beginning-of-life (-0.004 pcm/psi, BOL) but is always positive at operating conditions and becomes more positive during life

(+0.3 pcm/psi, EOL).

4.3-33 SGS-UFSAR Revision 17 October 16, 1998

Moderator Void Coefficient The moderator void coefficient relates the change in neutron multiplication to the presence of voids in the moderator. In a PWR this coefficient is not very significant because of the low void content in the coolant. The core void content is less than one-half of one percent and is due to local or statistical boiling. The void coefficient at BOL varies from 50 pam/percent void at low temperatures to -250 pcmjpercent void at EOL and at operating temperatures. The negative void coefficient at operating temperature becomes more negative with fuel burnup.

4.3.2.3.3 Power Coefficient The combined effect of moderator temperature and fuel temperature change as the core power level changes is called the total power coefficient and is expressed in terms of reactivity change per percent power change. A typical plot of the power coefficient at BOL and EOL conditions is given on Figure 4.3-24. It becomes more negative with burnup reflecting the combined effect of moderator and fuel temperature coefficients with burnup. A typical plot of the power defect (integral reactivity effect) at BOLand EOL is given on Figure 4.3-25. The power coefficient and defect change on a cycle-specific basis. The appropriate NOR should be referenced for the current cycle (see section 4.5).

4.3.2.3.4 Comparison of Calculated and Experimental Reactivity Coefficients section 4.3.3 describes the comparison of calculated and experimental reactivity coefficients in detail. Based on the data presented there, the accuracy of the current analytical model is:

+/- 0.2 percent 4p for Doppler and power defect i 2 pcm/°F fer the moderator coefficient Experimental evaluation of the calculated coefficients are done during the physics startup tests described in Section 14.

4.3-34 SGS-UFSAR Revision 17 October 16, 1998

4.3.2.3.5 Reactivity Coefficients used in Transient Analysis Table 4. 3-2 gives the representative ranges for the reactivity coefficients used in transient analysis. The exact values of the coefficient used in the analysis depend on whether the transient of interest is examined at the BOL or EOL, whether most negative or the most positive (least negative) coefficients are appropriate, and whether spatial nonuniformity must be considered in the analysis. Conservative values of coefficients, considering various aspects of analysis, are used in the transient analysis. This is described in section 15.

The values listed in Table 4.3-2 and illustrated on Figures 4.3-17 through 4.3-25 apply to a typical reload cycle. The coefficients appropriate for use in subsequent cycles depend on the core

• s operating history, the number and enrichment of fresh fuel assemblies, the loading pattern of burned and fresh fuel, the number and location of any burnable absorber rods, etc. The need for a reevaluation of any accident in a subsequent cycle is contingent upon whether or not the coefficients for that cycle fall within the identified range used in the analysis presented in Section 15. Control rod requirements for typical Unit 1 and Unit 2 reload cores are given in Table 4.3-3.

4.3.2.4 Control Requirements To ensure the shutdown margin stated in the Technical Specifications under conditions where a cooldown to ambient temperature is required, concentrated soluble boron is added to the coolant. Typical boron concentrations for several core conditions are listed in Table 4.3-2. For all core conditions including refueling, the boron concentration is well below the solubility limit. The RCCAs are employed to bring the reactor to the hot 4.3-35 SGS-UFSAR Revision 17 october 16, 1998

shutdown condition. 'l'he minimum required shutdown margin is given in the Technical Specifications.

The ability to accomplish the shutdown for hot conditions is demonstrated in Table 4.3-3 by comparing the difference between the RCCA's reactivity available with an allowance for the worst stuck rod with that required for control and protection purposes. The shutdown margin includes an allowance of 10 percent for analytic uncertainties (see Section 4.3.2.4.9). The largest reactivity control requirement appears at the EOL when the moderator temperature coefficient reaches its peak negative value as reflected in the larger power defect.

The control rods are required to provide sufficient reactivity to account for the power defect from full power to zero power and to provide the required shutdown margin. The reactivity addition resulting from power reduction consists of contributions from Doppler, moderator temperature, flux: redistribution, and reduction in void content as discussed below.

4.3.2.4.1 Doppler The Doppler effect arises from the broadening of U-238 and Pu-240 resonance peaks with an increase in effective pellet temperature. This effect is most noticeable over the range of zero power to full power due to the large pellet temperature increase with power generation.

4.3.2.4.2 Variable Average Moderator Temperature When the core is shut down to the hot, zero power condition, the average moderator temperature changes from the equilibrium full load value determined by the steam generator and turbine characteristics (steam pressure, heat transfer, tube fouling, etc) to the equilibrium no load value, which is based on the steam generator shell side design pressure. The design change in 4.3-36 SGS-UFSAR Revision 17 October 16, 1998

temperature is conservatively increased by 4°F to account for the control dead band and measurement errors.

Since the moderator coefficient is negative, there is a reactivity addition with power reduction. The moderator coefficient becomes more negative as the fuel depletes because the boron concentration is reduced. This effect is the major contributor to the increased requirement at EOL.

4.3.2.4.3 Redistribution During full power operation the coolant density decreases with core height, and this, together with partial insertion of control rods, results in less fuel depletion near the top of the core. Under steady state conditions, the relative power distribution will be slightly asymmetric towards the bottom of the core.

On the other hand, at hot zero power conditions, the coolant density is uniform up the core, and there is no flattening due to Doppler. The result will be a flux distribution which at zero power can be skewed toward the top of the core.

The reactivity insertion due to the skewed distribution is calculated with an allowance for the most adverse effects of xenon distribution.

4.3.2.4.4 Void Content A small void content in the core is due to nucleate boiling at full power. The void collapse coincident with power reduction makes a small reactivity contribution.

4.3.2.4.5 Rod Insertion Allowance At full power, the control bank is operated within a prescribed band of travel to compensate for small periodic changes in boron concentration, changes in temperature, and very small changes in the xenon concentration not compensated for by a change in boron 4.3-37 SGS-UFSAR Revision 6 February 15, 1987

concentration. When the control bank reaches either limit of this band, a change in boron concentration is required to compensate for additional reactivity changes. Since the insertion limit is set by a rod travel limit, a conservatively high calculation of the inserted worth is made which exceeds the normally inserted reactivity.

4.3.2.4.6 Burnup Excess reactivity is installed at the beginning of each cycle to provide sufficient reactivity to compensate for fuel depletion and fission products throughout the cycle. This reactivity is controlled by the addition of soluble boron to the coolant and by burnable absorbers. The soluble boron concentration for several core configurations, and unit boron worths are given in Table 4.3-2.

Since the excess reactivity for burnup is controlled by soluble boron and/or burnable absorbers, it is not included in control rod requirements.

4.3.2.4.7 Xenon and Samarium Poisoning Changes in xenon and samarium concentrations in the core occur at a sufficiently slow rate, even following rapid power level changes, that the resulting reactivity change is controlled by changing the soluble boron concentration.

4.3.2.4.8 pH Effects Changes in reactivity due to a change in coolant pH, if any, are sufficiently small in magnitude and occur slowly enough to be controlled by the Boron System.

Further details are available in Reference 8.

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4.3.2.4.9 Experimental Confirmation a normal shutdown, the total core cool down with a stuck rod has been measured on a 121 10-foot core, and a 121 12-foot core. In each case, the core was allowed to cool down until it reached criticality simulating the steamline break accident. For the 10-foot core, the total reactivity change associated with the cooldown was over-predicted by about 0. 3 percen: L'1p with respect to the measured result.

This represents an error of about 5 percent in the total change and is about half the allowance  :':or this For the 12-foot core, the difference between the measured and was an even smaller 0.2 percent L'1p. These measurements and others demonstrate the ability of the methods described in Section 4. 3. 3 to accurately predict the total shutdown reactivity of the core.

4.3.2.5 Core is controlled by means of a chemical dissolved in the coolant, RCCAs, and burnable absorber rods as described below.

4.3.2.5.1 Chemical Poison Boron in solution as boric acid is used to control relatively slow reactivity changes associated with:

The moderator temperature defect in going from cold shutdown at ambient temperature to the hot temperature at zero power

2. The transient xenon and samarium such as that following power or in RCC 4.3-39 SGS-Uf'SAR Revision 25 October 26, 201C

3. The excess reactivity required to compensate for the effects of fissile inventory depletion and buildup of long-life fission products

4. The burnable absorber depletion Typical boron concentrations for various core conditions are presented in Table 4.3-2.

4.3.2.5.2 Rod Cluster Control Assemblies Fifty-three RCCAs are employed. These are used for shutdown and for control purposes to offset fast reactivity changes associated with:

1. The required shutdown margin in the hot zero power, stuck rod condition

2. The reactivity compensation as a result of an increase in power above hot zero power (power defect including Doppler, and moderator reactivity changes)

3. Unprogrammed fluctuations in boron concentration, coolant temperature, or xenon concentration (with rods not exceeding the allowable rod insertion limits)

4. Reactivity ramp rates resulting from load changes The allowed control bank reactivity insertion is limited at full power to maintain shutdown capability. As the power level is reduced, control rod reactivity requirements are also reduced and more rod insertion is allowed. The control bank position is monitored and the operator is notified by an alarm if the limit is approached. The determination of the insertion limit uses conservative xenon distributions and axial power shapes. In addition, the RCCA withdrawal pattern determined from these analyses is used in determining power distribution factors and in 4.3-40

' SGS-UFSAR Revision 17 October 16, 1998

determining the maximum worth of an inserted RCCl\ ejection accident. For further discussion refer -:::o the Technical Specifications on rod insertion limits.

Power rod ection, and rod are based on the arrangement of the shutdown and control groups of the RCCAs shown on

4. 3-26A and B. All shutdown RCCAs are withdrat-vn before li'lithdrawal of the control banks is  ::_ni tiated. In going from zero to 100 percent power, control banks A, B, C, and D are v.Jithdrawn sequentially. The limits of rod posi ticns are provided in the COLR (see Section 4. 5)

• Further discussion on the basis of rod insertion limits are provided in the Technical Specifications.

4.3.2.5.3 Burnable Absorbers The burnable absorbers provide control of the excess reactivity available during the fuel cycle. In doing so, the modera~or temperature coefficient is prevented from being positive at normal operating conditions.

They perform this function by reducing the requirement for soluble poison in the moderator at the beginning of the fuel cycle as described previously. The number of burnable absorber rods per is shown in Section 4. 5 for a The boron in the burnable absorbers is with but at a slow rate so that the critical cor:centration of soluble boron is such that the moderator temperature coefficient remains non-at all times for power operating conditions.

4.3.2.5.4 Peak Xenon Startup Compensation for the peak xenon buildup is accomplished using the Boron Control System. Startup frorr the peak xenon condition is accomplished with a combination of rod motion and boron dilution.

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The boron dilution may be made at any time, even during the shutdown period, provided the shutdown margin is maintained.

4.3.2.5.5 Load Follow Control and Xenon Control During load follow maneuvers, power changes are accomplished using control rod motion and dilution or boration by the Boron System as required. Control rod motion is limited by the control rod insertion limits as provided in the COLR (see Section 4.$) and discussed previously in this section. Reactivity changes due to the changing xenon concentration can be controlled by rod motion and/or changes in soluble boron concentration.

4.3.2.5.6 Burnup Control of the excess reactivity for burnup is accomplished using soluble boron and/or burnup absorbers. The boron concentration must be limited during operating conditions to ensure the moderator temperature coefficient is non-positive. Sufficient burnable absorber is installed at the beginning of a cycle to give the desired cycle lifetime without exceeding the boron concentration value which would yield a positive MTC per Technical Specifications. The practical minimum boron concentration is 10 ppm.

4.3.2.6 Control Rod Patterns and Reactivity Worth The RCCAs are designated by function as the control groups and the shutdown groups. The terms "group" and "bank" are used synonymously throughout this report to describe a particular grouping of control assemblies. The RCCA pattern is displayed on Figures 4.3-26A and B which is not expected to change during the 1 ife of the plant. The control banks are labeled A, B, C, and D and the shutdown banks are labeled SA, SB, SC, and SD. Each bank, although operated and controlled as a unit, is comprised of two subgroups. The axial position of the RCCAs may be controlled manually or automatically. The RCCAs are all dropped into the core following actuation of reactor trip signals.

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.~--

Two criteria have been employed for selection of the control groups. First, the total reactivity worth must be adequate to meet the requirements specified in Table 4.3-3. Second, in view of the fact that these rods may be partially inserted at power operation, the total power peaking factor should be low enough to ensure that the power capability requirements are met. Analyses indicate that the first requirement can be met either by a single group or by two or more banks whose total worth equals at least the required amount. The axial power shape would be more peaked following movement of a single group of rods worth 3 to 4 percent Ap; therefore, four banks (described as A, B, C, and D on Figures 4. 3-26A and B) each worth approximately 1 percent Ap have been selected.

The position of control banks for criticality under any reactor condition is determined by the concentration of boron in the coolant. On an approach to criticality, boron is adjusted to ensure that criticality will be achieved with control rods above the insertion limit set by shutdown and other considerations (See the Technical Specifications).

Early in the cycle there may also be a withdrawal limit at low power to maintain a negative moderator temperature coefficient. Usual practice is to adjust boron to ensure that the rod position lies within the so-called maneuvering band, that is such that an escalation from zero power to full power does not require further adjustment of boron concentration.

Ejected rod worths are given in Section 15.3.1.6 for several different conditions. Experimental confirmation of these worths can be found by reference to startup test reports (9).

Allowable deviations due to misaligned control rods are discussed in the Technical Specifications.

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A representative calculation for two banks of control rods withdrawn simultaneously (rod withdrawal accident) is given on Figure 4.3-27.

Calculation of control rod reactivity worth versus time following reactor trip involves both control rod velocity and differential reactivity worth. The rod position versus time of travel after rod release assumed is given on Figure 4.3-28 for Vantage+ fuel. The drop time to the dashpot increases from 2.2 to 2.7 seconds for Vantage 5H, Vantage+, and RE'A, with the other times increasing proportionately. For nuclear design purposes, the reactivity worth versus rod position is calculated. by a series of steady state calculations at yarious control rod positions assuming all rods out of the core as the initial position in order to minimize the initial reactivity insertion rate. Also, to be conservative, the rod of highest worth is assumed stuck out of the core and the flux distribution (and thus reactivity importance) is assumed to be skewed to the bottom of the core. The result of these calculations is shown on Figure 4.3-29.

The shutdown groups provide additional negative reactivity to assure an adequate shutdown margin. Shutdown margin is defined as the amount by which the core would be subcritical at hot shutdown if all RCCAs are tripped, but assuming that the highest worth assembly remains fully withdrawn and no changes in xenon or bo:ron take place. The loss of control rod wort;h due to the material irradiation is negligible since only D bank rods may be in the core under normal operating conditions.

The values given in Table 4.3-3 show that the available reactivity in withdrawn RCCAs provides the design basis minimum shutdown margin allowing for the highest worth cluster to be at its fully withdrawn position. An allowance for uncertainty in the calculated worth of N-1 rods is made before determination of the shutdown margin.

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4.3.2.7 Criticality of Fuel Assemblies Criticality of fuel assemblies outside of the reactor is precluded by adequate design of fuel transfer and fuel storage facilities and by administrative control procedures. This section identifies those criteria important to criticality analyses.

New fuel is generally stored in fuel facilities with no water present but which are designed so as to prevent accidental criticality even if unborated water is present.

In the analysis for the storage facilities, the fuel assemblies are assumed to be in their most reactive condition, namely fresh or undepleted and with no control* rods or removable neutron absorbers present. Assemblies cannot be closer than the by the except in cases such as in fuel shipping containers where are carried out to establish the of the design. The mechanical integrity of the fuel assembly is assumed and no credit is taken for neutron absorption properties of the storage facility unless specifically included in the design. For full flooding with unborated water, the fuel assembly spacing of the facility provides essentially full nuclear isolation, and for the array is no greater than keff for the single most reactive fuel The criterion for full flooding is ~ 0.95. For the analysis of new (dry) fuel storage racks, an additional criterion of keff ~ 0.98 is confirmed for optimum (low density) moderation conditions. These fresh fuel rack (J]l include allowances for uncertainties, biases, and manufacturing tolerances and assure 95 percent probability I 95 percent confidence level that the keff critically design criteria are met.

The fuel assembly (17 x 17 fuel rods) of standard design and 3.5 w/o enriched uranium without a control rod or burnable absorbers, fully flooded and reflected with cold clean water, has a keff of about 0. 85. Two such fuel assemblies spaced 1 inch apart with parallel axes 9.5 inches apart have a of about 0.99. Three such fuel assemblies 1 inch with axes would be supercritical.

An infinite number of dry fuel assemblies of this design would have a keff

< 0.80.

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4.3.2.8 Stability 4.3.2.8.1 Introduction The stability of the PWR cores against xenon~induced spatial oscillations and the control of such transients are discussed extensively (References 5, 10, 11,

12) . A surrunary of these reports is in the discussion and the bases are given in Section 4.3.1.6.

In a large reactor core, xenon-induced oscillations can take place with no corresponding change in the total power of the core. The oscillation may be caused by a power shift in the core which occurs rapidly by comparison with the xenon-iodine time constants. Such a power shift occurs in the axial direction when a plant load change is made by control rod motion and results in a change in the moderator density and fuel distributions. Such a power shift could occur in the diametral of the core as a result of abnormal control action.

Due to the negative power coefficient of reactivity, PWR cores are inherently stable to oscillations in total power. Protection against total power stabilities is provided by the Control and Protection System as described in Chapter 7. Hence, the discussion on the core stability will be limited here to xenon-induced oscillations.

4.3.2.8.2 Stability Index Power distributions, either in the axial direction or in the X-Y plane, can undergo oscillations due to perturbations introduced in the equilibrium distributions without changing the total core power. The overtones in the current PWRs, and the stability of the core against xenon-induced oscillations can be determined in terms of the eigenvalues of the first flux overtones.

either in the axial direction or in the X-Y the I Ct of the first flux harmonic as 4.3-46 SGS-UFSAR Revision 17 October 16, 1998

e

• b + ic, (Reference 10}

then b is defined as the stability index and T *2~/C as the oscillation period of the first harmonic. The time-dependence of the first harmonic o4l in the power distribution can now be represented as bt ae cos ct, (Reference 12) where A and a are constants. The stability index can also be obtained approximately by:

where A , A are the successive peak amplitudes of the oscillation and T is the n n+ 1 time period between the successive peaks.

4.3.2.8.3 Prediction of the core Stability The stability of the core described herein (i.e., with 17 x 17 fuel assemblies) against xenon-induced spatial oscillations is expected to be equal to or better than that of earlier designs. The prediction is based on a comparison of the parameters which are significant in determining the stability of the core against the xenon-induced oscillations, namely ( 1) the overall core size is unchanged and spatial power distributions will be similar, ( 2) the moderator temperature coefficient is expected to be similar, and (3) the Doppler coefficient of reactivity is expected to be similar at full power.

Analysis of both the axial and X-Y xenon transient tests, discussed in Section 4.3.2.8.5, shows that the calculational model is adequate for the prediction of core stability.

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4.3.2.8.4 Stability Measurements Axial Measurements TWo axial xenon transient tests conducted in a PWR with a core height of 12 feet and 121 fuel assemblies are reported in Reference 13, and will be briefly discussed here. The tests were performed at approximately 10 percent and 50 pe1.*cent of cycle 1" ~':"

Both a free-running oscillation test and a controlled test were performed during the first test. The second test at mid-cycle consisted of a free-running oscillation test only. In each of the free-running oscillation tests, a perturbation was introduced to the equilibrium power distribution through an impulse motion of the control bank 0, and the subsequent oscillation was monitored to measure the stability index and the oscillation period. In the controlled test conducted early in the cycle, the part-length rods were used to follow the oscillations to maintain an axial offset (AO) within the prescribed limits. The AO of power was obtained from the ex-core ion chamber readings (which had been calibrated against the in-core flux maps) as a function of time for both free-running tests as shown on Figure 4.3-30.

The total core power was maintained constant during these spatial xenon tests, and the stability index and the oscillation period were obtained from a least-square fit of the AO data in the form of Equation 2. The AO of power is the quantity that properly represents the axial stability in the sense that it essentially eliminates any contribution from even order harmonics including the fundamental mode. The conclusions of the tests are the following:

1. The core was stable against induced axial xenon transients both at the core a~erage burnups of 1550 MWO/MTU and 7700 MWO/MTU. The measured stability indices are -0.041 hr-l for the first test (Curve 1 of 4.3-48 SGS-UFSAR Revision 17 October 16, 1998

Figure 4.3-30) and -0.014 hr- 1 for the second test (Curve 2 of Figure 4.3-30). The corresponding oscillation periods are 32.4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> and 27.2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. respectively.

2. The reactor core becomes less stable as fuel burnup progresses and the axial stability index was essentially zero at 12,000 MWD/MTO.

Measurements in the X-Y Plane Two X-Y xenon oscillation tests were performed at a PWR plant with a core height of 12 feet and 157 fuel assemblies. The first test was conducted at a core average burnup of 1540 MWD /MTU and the second at a core average burnup of 12900 MWD/MTU. Both of the X-Y xenon tests show that the core was stable in the X-Y plane at both burnups. The second test shows that the core became more stable as the fuel burnup increased and all Westinghouse PWRs with 121 and 157 assemblies are expected to be stable throughout their burnup cycles.

In each of the two X-Y tests, a perturbation was introduced to the equilibrium power distribution through an impulse motion of one RCCA located along the diagonal axis. Following the perturbation, the uncontrolled oscillation was monitored using the moveable detector and thermocouple system and the ex-core power range detectors. The quadrant tilt difference is the quantity that properly represents the diametral oscillation in the X-Y plane of the reactor core in that the differences of the quadrant average powers over two symmetrically opposite quadrants essentially eliminates the contribution to the oscillation from the azimuthal mode. The quadrant tilt difference data were fitted in the form of Equation 2 through a least-square method. A stability index of -0.076 hr -l with a period of 29.6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> was obtained from the thermocouple data shown on Figure 4.3-31.

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It was observed in the second X-Y xenon test that the PWR core with 157 fuel assemblies had become more stable due to an increased fuel depletion, and the stability index was not determined.

4.3.2.8.5 comparison of Calculations with Measurements The analysis of the axial xenon transient tests was performed in an axial slab geometry using a flux synthesis technique. The direct simulation of the AO data was carried out using the PANDA Code (Reference 14). The analysis of the x-Y xenon transient tests was performed in an X-Y geometry using a modified TURTLE code (Reference 7). Both the PANDA and TURTLE codes solve the two-group time-dependent neutron diffusion equation with time-dependent xenon and iodine concentrations. The fuel temperature and moderator density feedback is limited to a steady state model. All the X-Y calculations were performed in an average enthalpy plane.

The basic nuclear cross sections used in this study were generated from a unit cell depletion program which was evolved from the codes LEOPARD (Reference 15) and CINDER (Reference 16). The detailed experimental data during the tests including the reactor power level, enthalpy rise, and the impulse motion of the control rod assembly, as well as the plant follow burnup data, were closely simulated in the study.

The results of the stability calculation for the axial tests are compared with the experimental data in Table 4.3-4. The calculations show conservative results

-1 for both of the axial tests with a margin of approximately 0.01 hr in the stability index.

An analytical simulation of the first X-Y xenon oscillation test shows a

-1 calculated stability index of -0.081 hr , in good agreement with the measured

-1 value of -0.076 hr

• As indicated earlier, the second X-Y xenon test showed that the core had become more stable compared to the first test and no evaluation of the stability index was attempted. This increase in the core 4.3-50 SGS-UFSAR Revision 17 October 16, 1998

stability in the X-Y plane due to increased fuel burnup is due mainly to the increased magnitude of the negative moderator temperature coefficient.

Previous studies of the physics of xenon oscillations, including three-dimensional analysis, are reported in the series of topical reports (References 10, 11, 12). A more detailed description of the experimental results and analysis of the axial and X-Y xenon transient tests is presented in Reference 13 and Section 1 of Reference 17.

4.3.2.8.6 Stability Control and Protection The Ex-core Detector System is utilized to provide indications of xenon-induced spatial oscillations. The readings from the ex-core detectors are available to the operator and also form part of the protection system.

Axial Power Distribution For maintenance of proper axial power distributions, the operator is instructed to maintain an axial offset within a prescribed operating band, based on the ex-core detector readings. Should the axial offset be permitted to move far enough outside this band, the protection limit will be reached and the power will be automatically cut back.

Twelve-foot PWR cores become less stable to axial xenon oscillations as fuel burnup progresses. However, free xenon oscillations are not allowed to occur except for special tests. The full length control rod banks present in all modern Westinghouse PWRs are sufficient to dampen and control any axial xenon oscillations present. Should the axial offset be inadvertently permitted to move far enough outside the control band due to an axial xenon oscillation, or any other reason, the protection limit on axial offset will be reached and the power will be automatically cut back.

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Radial Power Distribution The core described herein is calculated to be stable against X-Y xenon induced oscillations at all times in life.

The X-Y stability of large PWRs has been further verified as part of the startup physics test program for PWR cores with 193 fuel assemblies. The measured x-Y stability of the cores with 157 and 193 assemblies was in good agreement with the calculated stability discussed in Sections 4. 3. 2. 8. 4 and 4. 3. 2. 8. 5. In the unlikely event that X-Y oscillations occur, backup actions are possible and would be implemented, if necessary, to increase the natural stability of the core.

This is based on the fact that several actions could be taken to make the moderator temperature coefficient more negative, which will increase the stability of the core in the X-Y plane.

Provisions for protection against nonsymmetric perturbations in the x-Y power distribution that could result from equipment malfunctions are made in the protection system delllign. This includes control rod drop, rod misalignment, and asymmetric loss-of-coolant flow.

A more detailed discussion of the power distribution control in the PWR cores is presented in Reference 5.

4.3.2.9 Vessel Irradiation A brief review of the methods and analyses used in the determination of neutron and gamma ray flux attenuation between the core and the pressure vessel is given below. A more complete discussion is given in the pressure vessel irradiation and surveillance program.

The materials that serve to attenuate neutrons originating in the core and gamma rays, from both the core and structural component consist of the core baffle, core barrel, the neutron pads, and 4.3-52

. SGS-UFSAR Revision 6 February 15, 1987

associated water annuli, all of which are within the region between the core and the pressure vessel .

In general, few-group neutron diffusion theory codes are used to determine fission power density distributions within the active core, and the accuracy of these analyses is verified by in-core and/or BEACON power distribution measurements on operating reactors. Region and rodwise power sharing information from the core calculations is then used as source information in two-dimensional transport calculations which compute the flux distributions through the reactor.

The neutron flux distribution and spectrum in the various structural components varies significantly from the core to the pressure vessel. Representative values of the neutron flux distribution and spectrum are presented in Table 4. 3-5. The values listed are based on time averaged equilibrium cycle reactor core parameters and power distributions, and thus, are suitable for long-term nvt projections and for correlation with radiation damage estimates.

The irradiation surveillance program utilizes actual test samples to verify the accuracy of the calculated fluxes at the vessel .

4.3.3 Analytical Methods Calculations required in nuclear design consist of three distinct types, which are performed in sequence:

1. Determination of effective fuel temperatures

2. Generation of macroscopic few-group parameters

3. Space-dependent, few-group diffusion calculations These calculations are carried out by computer codes which can be executed individually; however, at Westinghouse most of the codes required have been linked to form an automated design sequence

• SGS-UFSAR 4.3-53 Revision 19 November 19, 2001

which minimizes design time, avoids errors in transcription of data, and standardizes the design methods.

4.3.3.1 Fuel Temperature <Doppler) Calculations Temperatures vary radially within the fuel rod depending on the heat generation rate in the pellet; the conductivity of the materials in the pellet, gap, and clad; and the temperature of the coolant.

The fuel temperatures for use in most nuclear design Doppler calculations are obtained from a simplified version of the Westinghouse fuel rod design model described in Section 4.2.1.3.1, which considers the effect of radial variation of pellet conductivity, expansion-coefficient and heat generation rate, elastic deflection of the clad, and a gap conductance which depends on the initial fill gas, the hot open gap dimension, and the fraction of the pellet over which the gap is closed. The fraction of the gap assumed closed represents an empirical adjustment used to produce good agreement with observed reactivity data at BOL.

Further gap closure occurs with burnup and accounts for the decrease in Doppler defect with burnup which has been observed in operating plants. For detailed calculations of the Doppler coefficient, a more sophisticated temperature model is used which accounts for the effects of fuel swelling, fission gas release, and plastic clad deformation.

Radial power distributions in the pellet as a function of burnup are obtained from LASER {Reference 18) calculations.

The effective U-238 temperature for resonance absorption is obtained from the radial tempe.t'ature distribution by applying a radially dependent weighting function. The weighting function was determined from REPAD (Reference 19) Monte Carlo calculations of resonance escape probabilities in several steady state and transient temperature distributions. In each case a flat pellet temperature was determined which produced the same resonance escape probability as the actual distribution. The weighting function was empirically determined from these results.

The effective Pu-240 temperature for resonance absorption is determined by a convolution of the radial distribution of Pu-240 number densities from LASER burnup calculations and the radial weighting function. The resulting temperature is burnup dependent, but the difference between U-238 and Pu-240 temperatures, in terms of reactivity effects, is small.

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The effective pellet temperature for pellet dimensional change is that value which produces the same outer pellet radius in a virgin pellet as that obtained from the temperature model. The effective clad temperature for dimensional change is its average value.

The temperature calculational model has been validated by plant Doppler defect data as shown in Table 4.3-6 and Doppler coefficient data as shown on Figure 4.3-32. Stability index measurements also provide a sensitive measure of the Doppler coefficient near full power (see Section 4.3.2.8). It can be seen that Doppler defect data is typically within 0.2 percent of prediction.

ALPHA/PHOENIX/ANC has two Doppler models - a Doppler power model and a Doppler temperature model. The default Doppler model in APA is the temperature model and is based on a fit of fast absorption cross sections against the fuel temperature at 0, 1, and 2 times the reference power. In NEXUS/ANC9, the effects of fuel temperature are captured on all the cross sections directly, as it is one of the fundamental parameters used to fit cross sections.

4.3.3.2 Macroscopic Group Constants There are lattice codes which have been used for the generation of macroscopic group constants needed in the spatial, few-group diffusion codes. One is a version of the LEOPARD and CINDER codes, which has historically been the source of the macroscopic group constants. The others are PHOENIX-P and PARAGON, which are used in present reload designs (Reference 30 and 38). The NEXUS methodology (Reference 39) is a reparameterization of the PARAGON nuclear data output. The NEXUS methodology provides a linkage between PARAGON and ANC, establishing a new code system, while still using PARAGON.

Macroscopic few-group constants and analogous microscopic cross sections (needed for feedback and microscopic depletion calculations) were previously generated for fuel cells by a version of the LEOPARD (Reference 15) and CINDER (Reference 16) codes, which are linked internally and provide burnup dependent cross sections. Normally a simplified approximation of the main fuel chains is used; however, where needed, a complete solution for all the significant isotopes in the fuel chains from Th-232 to Cm-244 is available (Reference 20).

Fast and thermal cross section library tapes contain microscopic cross sections taken for the most part from the ENDF/B (Reference 21) library, with a few exceptions where other data provide better agreement with critical experiments, isotopic measurements, and plant critical boron values. The effect on the unit fuel cell of non-lattice components in the fuel assembly is obtained by supplying an appropriate volume fraction of these materials in an extra region which is homogenized with the unit cell in the fast (MUFT) and thermal (SOFOCATE) flux calculations. In the thermal calculation, the fuel rod, clad, and moderator are homogenized by energy-dependent disadvantage factors derived from an analytical fit to integral transport theory results.

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Group constants for discrete burnable absorber cells, guide thimbles, instrument thimbles, and interassembly gaps are generated in a manner analogous to the fuel cell calculation. Reflector group constants are taken from infinite medium LEOPARD calculations. Baffle group constants are calculated from an average of core and radial reflector microscopic group constants for stainless steel.

Group constants for control rods are calculated in a linked version of the HAMMER (Reference 22) and AIM (Reference 23) codes to provide an improved treatment of self shielding in the broad resonance structure of these isotopes at epithermal energies than is available in LEOPARD. The Doppler broadened cross sections of the control rod materials are represented as smooth cross sections in the 54-group LEOPARD fast group structure and in 30 thermal groups.

The four-group constants in the rod cell and appropriate extra region are generated in the coupled space-energy transport HAMMER calculation. A corresponding AIM calculation of the homogenized rod cell with extra region is used to adjust the absorption cross sections of the rod cell to match the reaction rates in HAMMER. These transport-equivalent group constants are reduced to two-group constants for use in space-dependent diffusion calculations. In discrete X-Y calculations only one mesh interval per cell is used, and the rod group constants are further adjusted for use in this standard mesh by reaction rate matching the standard mesh unit assembly to a fine-mesh unit assembly calculation.

Validation of the cross section method is based on analysis of critical experiments as shown in Table 4.3-7, isotopic data as shown in Table 4.3-8, plant critical boron (C ) values at HZP, BOL, as shown in Table 4.3-9 and at B

HFP as a function of burnup as shown on Figures 4.3-33 through 4.3-35. Control rod worth measurements are shown in Table 4.3-10.

Confirmatory critical experiments on discrete burnable absorbers are described in Reference 24.

PHOENIX-P has been approved by the USNRC as a lattice code for the generation of macroscopic and microscopic few group cross sections for PWR analysis (Reference 30). PHOENIX-P is a two-dimensional, multigroup, transport-based lattice code capable of providing all necessary data for PWR analysis. Since it is a dimensional lattice code, PHOENIX-P does not rely on predetermined spatial/spectral interaction assumptions for the heterogeneous fuel lattice and can provide a more accurate multigroup flux solution than versions of LEOPARD/CINDER.

The solution for the detailed spatial flux and energy distribution is divided into two major steps in PHOENIX-P (Reference 30).

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First, a two-dimensional fine energy group nodal solution is obtained, coupling individual subcell regions (pellet, clad, and moderator) as well as surrounding pins, using a method based on Carlvik's collision probability approach and heterogeneous response fluxes which preserve the heterogeneity of the pin cells and their surroundings. The nodal solution provides an accurate and detailed local flux distribution, which is then used to homogenize the pin cells spatially to fewer groups.

Then, a standard S4 discrete ordinates calculation solves for the angular distribution, based on the group-collapsed and homogenized cross-sections from the first step. These S4 fluxes normalize the detailed spatial and energy nodal fluxes, which are then used to compute reaction rates, power distributions and to deplete the fuel and burnable absorbers. A standard B1 calculation evaluates the fundamental mode critical spectrum, providing an improved fast diffusion coefficient for the core spatial codes.

PHOENIX-P employs a multiple energy group library consisting of 42 or more energy groups derived mainly from ENDF/B files. This library was designed to capture the integral properties of the multigroup data properly during group collapse and to model important resonance parameters properly. It contains all neutronics data necessary for modeling fuel, fission products, cladding and structural materials, coolant, and control and burnable absorber materials present in the PWRs.

Group constants for burnable absorber cells, control rod cells, guide thimbles and instrumentation thimbles, or other non-fuel cells, can be obtained directly from PHOENIX-P without any adjustments such as those required in the cell or ID lattice codes.

PARAGON is a two-dimensional multi-group neutron (and gamma) transport code.

It is an improvement over the Westinghouse licensed code PHOENIX-P (Reference 30). The main difference between PARAGON (Reference 38) and PHOENIX-P resides in the flux solution calculation. PHOENIX-P uses a nodal cell solution coupled to an S4 transport solution as described in Reference 38. PARAGON uses the Collision Probability theory within the interface current method to solve the integral transport equation. Throughout the whole calculation, PARAGON uses the exact heterogeneous geometry of the assembly and the same energy groups as in the cross-section library to compute the multi-group fluxes for each micro-region location of the assembly.

In order to generate the multi-group data that will be used by a core simulator code PARAGON goes through four steps of calculations: resonance self-shielding, flux solution, homogenization and burnup calculation.

PARAGON can provide nuclear data, both cross sections and pin power information, to a core simulator code such as ANC.

4.3-57 SGS-UFSAR Revision 31 December 5, 2019

The NEXUS methodology (Reference 39) is a reparameterization of the PARAGON nuclear data output (cross sections) and a new reconstruction approach with the ANC core simulator code to simplify the use of this code system for design use.

The NEXUS methodology provides a linkage between PARAGON and ANC, establishing a new code system, while still using PARAGON. The NEXUS approach is to account for the spectral changes by parameterizing the cross section output of PARAGON, such that the cycle specific boron letdown curves do not need to be provided in the analysis. The parameterization adequately accounts for the relevant neutronic effects of temperature feedback, fuel depletion, burnable poisons, boron concentrations, and fission products.

The NEXUS methodology (Reference 39) approach uses a calculational matrix of lattice code calculations performed with PARAGON to form a set of data in order to parameterize the cross sections according to a spectral index (SI), the moderator temperature (Tm), and the fuel temperature (Tf). These parameters, in conjunction with nuclide tracking during irradiation, allow for feedback-free cross sections, and correction functions to be generated. The lattice calculations are performed using a base-line reference depletion case with several branches to account for the effects of different local conditions, thus providing a data set that covers a wide range of potential local conditions ranging from those typical of a cold shutdown reactor condition to full power conditions. The SI is based on the ratio of the epithermal to thermal neutron flux and is a measure of the local neutron spectrum. The Tm and Tf dependences account for changes in moderation and resonance absorption respectively. These parameters are used to develop a series of correction factors to account for these physical effects using a multivariable least-squares technique. The correction factors are dependent on the differences between the nodal values for these parameters and the values used in the reference lattice depletion calculations. The effects of xenon, actinides, other fission products, and burnable absorbers are directly accounted for by first tracking the number density of each isotope directly, thereby accounting for explicitly for fuel depletion. The macroscopic cross sections themselves are reconstructed based on these number densities and the microscopic cross section for each particular isotope given the nodal conditions. The microscopic cross sections in these cases are adjusted by correction functions based on local nodal parameters.

4.3.3.3 Spatial Few-Group Diffusion Calculations Historically, spatial few-group diffusion calculations consisted primarily of two-group X-Y calculations using an updated version of the TURTLE code and two-group axial calculations using an updated version of the PANDA code.

Discrete X-Y calculations (1 mesh per cell) were carried out to determine critical boron concentrations and power distributions in the X-Y plane.

4.3-58 SGS-UFSAR Revision 31 December 5, 2019

An axial average in the X-Y plane was obtained by synthesis from unrodded and rodded planes. Axial effects in unrodded depletion calculations were accounted for by the axial buckling, which varies with burnup and is determined by radial depletion calculations which are matched in reactivity to the analogous R-Z depletion calculation. The moderator coefficient is evaluated by varying the inlet temperature in the same X-Y calculations used for power distribution and reactivity predictions.

Validation of the reactivity calculations is associated with the validation of the group constants themselves, as discussed in Section 4.3.3.2. Validation of the Doppler calculations is associated with the fuel temperature validation discussed in Section 4.3.3.1. Validation of the moderator coefficient calculations is obtained by comparison with plant measurements at hot zero power conditions as shown in Table 4.3-11.

Axial calculations are used to determine differential control rod worth curves (reactivity versus rod insertion) and axial power shapes during steady state and transient xenon conditions (flyspeck curve). Group constants are obtained from the three-dimensional nodal model by flux-volume weighting on an axial slice-wise basis. Radial bucklings are determined by varying parameters in the buckling model while forcing the one-dimensional model to reproduce the axial characteristics (axial offset, mid-plane power) of the three-dimensional model.

Recent few-group spatial calculations have input PHOENIX-P or PARAGON supplied two-group cross-sections to the Advanced Nodal Code (ANC). ANC is a two-group, two or three-dimensional nodal code capable of determining typical nuclear design analyses, such as critical boron concentrations, average assembly and pin powers, control rod worths, reactivity coefficients, assembly and pin burnups and axial power distributions. Through the use of advanced nodal techniques, ANC is able to produce solutions similar to the fine mesh, finite difference diffusion theory codes such as TURTLE/TORTISE. ANC has been benchmarked against TORTISE (an improved version of TURTLE) for normal and off-normal conditions, such as ejected rod, stuck rod and dropped rod (Reference 31). The qualification of the PHOENIX-P/ANC methodology against measured data is given in Reference 30. The qualification of the PARAGON/ANC methodology against measured data is given in Reference 38. The qualification of the NEXUS/ANC methodology against measured data is given in Reference 39. The qualification of new pin power recovery methodology can be found in Reference 40.

Validation of the spatial codes for calculating power distributions involves the use of in-core and ex-core detectors and the BEACON core monitoring system (PDMS) and is discussed in Section 4.3.2.2.7. Note that BEACON (Reference 37) was affirmed for continued use with the USNRC approved Westinghouse design model methodologies PHOENIX-P/ANC, PARAGON/ANC, and NEXUS/ANC.

4.3-59 SGS-UFSAR Revision 31 December 5, 2019

4.3.3.4 Pin Power Reconstruction The conventional methodology implemented in ANC calculates the homogeneous pin power distribution and applies the group-wise pin power form factors (these were referred to as pin factors) to obtain the final pin power. The conventional methodology has shown historically that it can predict the pin power with high accuracy for traditional PWR cores, which are operated without significant insertion of control rod banks. With the introduction of new PWR core designs control rods may be inserted into the core during operation, which may significantly change the heterogeneity of the fuel assemblies. Since the conventional methodology used in ANC does not include the control rod history effect on the pin factors, the pin power distribution is not as accurate when control rods are inserted for significant periods of time during operation.

This is particularly true for high-worth control rods. Moreover, because the control rod insertion and withdrawal strategy is not pre-determined, conventional pin power methodology has difficulty in capturing the heterogeneity change and the accumulated history impact on the pin power distribution. This limitation is overcome by the new methodology (Reference 40), which directly follows the history of each individual fuel rod in ANC and computes the fuel rod macroscopic cross-sections based on the fuel rod history and the local spectrum. Therefore, the new methodology enables ANC to calculate the effect of control rod insertion during operation on pin power distribution while maintaining the same accuracy as the conventional method for a traditional core.

Based on comparison with measured data it is estimated that the accuracy of current analytical methods is:

+/- 0.2 percent for Doppler defect

-5

+/- 2 x 10 /F for moderator coefficient

+/- 50 ppm for critical boron concentration with depletion

+/- 3 percent for power distributions

+/- 0.2 percent for rod bank worth

+/- 4 pcm/step for differential rod worth

+/- 0.5 pcm/ppm for boron worth

+/- 0.1 percent for moderator defect 4.3.4 References for Section 4.3

1. "Westinghouse Anticipated Transients Without Reactor Trip Analysis," WCAP-8330, August 1974.

2. Langford, F. L. and Nath, R. J., Jr., "Evaluation of Nuclear Hot Channel Factor Uncertainties," WCAP-7308-L, April 1969 (Westinghouse Proprietary) and WCAP-7810 (Non-Proprietary), December 1971.

4.3-60 SGS-UFSAR Revision 31 December 5, 2019

3. McFarlane, A. F., "Core Power Capability in Westinghouse PWRs," WCAP-7267-L, October 1969 (Proprietary) and WCAP-7809 (Non-Proprietary), December 1971.

4. Hellman, J. M. (Ed.), "Fuel Densification Experimental Results and Model for Reactor Application," WCAP-8218-P-A (Proprietary) and WCAP-8219-A (Non-Proprietary), March 1975.

5. Moore, J. S., "Power Distribution Control of Westinghouse Pressurized Water Reactors," WCAP-7208, September 1968 (Proprietary) and WCAP-7811 (Non-Proprietary), December 1971.

6. McFarlane, A. F., "Power Peaking Factors," WCAP-7912-P-A, (Proprietary) and WCAP-7912-A, (Non-Proprietary), January 1975.

7. Altomare, S. and Barry, R. F., "The TURTLE 24.0 Diffusion Depletion Code,"

WCAP-7213-P-A (Proprietary) and WCAP-7758 (Non-Proprietary), February 1975.

8. Cermak, J. 0., et al, "Pressurized Water Reactor pH - Reactivity Effect,"

Final Report, WCAP-3696-8 (EURAEC-2074), October 1968.

9. Outzs, J. E., "Plant Startup Test Report, H. B. Robinson Unit No. 2,"

WCAP-7844, January 1972.

10. Poncel, C. G. and Christie, A. M., "Xenon-Induced Spatial Instabilities in Large PWRs," WCAP-3680-20 (EURAEC-1974), March 1968.

11. Skogen, F. B. and McFarlane, A. F., "Control Procedures for Xenon-Induced X-Y Instabilities in Large PWRs," WCAP-3680-21, (EURAEC-2111), February 1969.

12. Skogen, F. B. and McFarlane, A. F., "Xenon-Induced Spacial Instabilities in Three-Dimensions," WCAP-3680-22 (EURAEC-2116), September 1969.

13. Lee, J. C., et al, "Axial Xenon Transient Tests at the Rochester Gas and Electric Reactor," WCAP-7964, June 1971.

14. Altomare, S. and Minton, G., "The PANDA Code," WCAP-7048-P-A (Proprietary) and WCAP-7757-A (Non-Proprietary), February 1975.

4.3-61 SGS-UFSAR Revision 31 December 5, 2019

15. Barry, R. F., "LEOPARD - A Spectrum Dependent Non-Spatial Depletion Code for the IBM-7094," WCAP-3269-26, September 1963.

16. England, T. R., "CINDER - A One-Point Depletion and Fission Product Program," WAPD-TM-334, August 1962.

17. Kubit, C. J., "Safety Related Research and Development for Westinghouse Pressurized Water Reactors, Program Summaries, Spring-Fall 1973," WCAP-8204, October 1973.

18. Poncelot, C. G., "LASER - A Depletion Program for Lattice Calculations Based on MUFT and THEMOS," WCAP-6073, April 1966.

19. 0lhoeft, J. E., "The Doppler Effect for a Non-Uniform Temperature Distribution in Reactor Fuel Elements," Final Report, WCAP-2048, July 1962.

20. Nodvik, R. J., et al, "Supplementary Report on Evaluation of Mass Spectrometric and Radiochemical Analysis of Yankee Core I Spent Fuel, Including Isotopes of Elements Thorium Through Curium," WCAP-6086, August 1969.

21. Drake, M. K. (Ed.), "Data Formats and Procedure for the ENDF Neutron Cross Section Library," 8NL-50274, ENDF-102, Vol. 1, 1970.

22. Suich, J. E. and Honeck, H. C., "The HAMMER System, Heterogeneous Analysis by Multigroup Methods of Exponentials and Reactors," DP-1064, January 1967.

23. Flatt, H. P. and Baller, D. C., "AIM-5, A Multigroup, One Dimensional Diffusion Equation Code," NAA-SR-4694, March 1960.

24. Barry, R. F., "Nuclear Design of Westinghouse Pressurized Water Reactors with Burnable Poison Rods," WCAP-7806, December 1971.

25. Strawbridge, L. E. and Barry, R. F., "Criticality Calculations for Uniform Water-Moderated Lattices," Nuclear Science and Engineering 23, 58, 1965.

26. Nodvik, R. J., "Saxton Core II Fuel Performance Evaluation," WCAP-3385-56, Part II, "Evaluation of Mass Spectrometric and Radiochemical Materials Analyses of Irradiated Saxton Plutonium Fuel," July 1970.

4.3-62 SGS-UFSAR Revision 31 December 5, 2019

27. Leamer, R. D., et al, "PUO2-UO2 Fueled Critical Experiments," WCAP-3726-1, July 1967.

28. Davidson, S. L., Ed., et al., "Extended Burnup Evaluation of Westinghouse Fuel," WCAP-10125-P-A, Appendix B, December 1985.

29. Henderson, W. B., "Results of the Control Rod Worth Program," WCAP-9217, October 1977.

30. Nguyen, T. Q. et al., "Qualification of the PHOENIX-P/ANC Nuclear Design System for Pressurized Water Reactor Cores," WCAP-11596-P-A, June 1988.

31. Liu, Y. S., et al., "ANC: A Westinghouse Advanced Nodal Computer Code,"

WCAP-10966-A, September 1986.

32. Iorii, J. A. and Petrarca, D. J., Westinghouse Wet Annular Burnable Absorber Evaluation Report, WCAP-10021-P-A, Revision 1, October 1983

33. Bradfute, J. L., et al, Criticality Analysis of the Salem Units 1 and 2 Fresh Fuel racks, NFU-VTDWW-94-083-00, January 1994

34. C.L. Beard and T. Morita, "BEACON Core Monitoring and Operations Support System", WCAP-12472-P-A, August, 1994.

35. T. R. Wathey, Conditional Extension of the Rod Misalignment Technical Specification for Salem Units 1 and 2, WCAP 14962/14963, August 1997.

36. T. Morita and W. H. Slagle, BEACON Core Monitoring and Operations Support System (WCAP-12472-P-A), Addendum 1-A, January 2000.

37. W. A. Boyd, BEACON Core Monitoring and Operation Support System, (WCAP-12472-P-A), Addendum 4, September 2012.

38. W. H. Slagle, Qualification of the Two-Dimensional Transport Code Paragon (WCAP-16045-P-A), Revision 0, August 2004.

39. W. H. Slagle, Qualification of the NEXUS Nuclear Data Methodology (WCAP-16045-P-A), Addendum 1-A Revision 0, August 2007.

40. Zhang B., et al, Qualification of the New Pin Power Recovery Methodology (WCAP-10965-P-A), Addendum 2-A Revision 0, September 2010.

4.3-63 SGS-UFSAR Revision 31 December 5, 2019

TABLE 4.3-1 REACTOR CORE DESCRIPTION Active Core Equivalent Diameter, in. 132.7 Core Average Active Fuel Height, First Core (Hot), in. 143.7 Height-to-Diameter Ratio 1.09 2

Total Cross Section Area, ft 96.06 H 0/U Molecular Ratio, Lattice (Cold) 2. 41 2

Reflector Thickness and Composition Top - Water plus Steel, in. -vlQ Bottom - Water plus Steel, in. -10 Side - Water plus Steel, in. -15 Fuel Assemblies Number 193 Rod Array 17 X 17 Rods per Assembly 264 Rod Pitch, in. 0.496 Overall Transverse Dimensions, in. 8.426 X 8.426 Fuel Weight (as uo ), lb 222,739{V5H,V+) 217,565 (RFA) 2 Zircaloy Weight, lb 52,541 (V5H, V+) 53,847 (RFA)

Number of Grids per Assembly V5H 2 Inconel (Top & Bottom) 6 Zircaloy-4 (Mid Grids)

V+ 2 Inconel (Top & Bottom) 6 Zirlo' {Mid Grids)

RFA 2 Inconel (Top & Bottom) 1 Inconel (Protective Grid) 6 Zirlo' (Mid Grids) 3 Zirlo' (Intermediate Flow Mixing Grids)

Weight of Grids (Effective in Core), lb 2324 (VSH, V+) 3248 (RFA)

Number of Guide Thimbles per Assembly 24 Composition of Guide Thimbles Zircaloy-4 (V5H)

.1 Z~r o TM (V+, RFA) 1 of 3 SGS-UFSAR Revision 20 May 6, 2003

TABLE 4.3-1 (Cont.)

REACTOR CORE DESCRIPTION Dia. of Guide Thimbles (upper ) , in. 0.442 ID X 0.474 OD (V5H, V+)

0.442 ID X 0.482 OD (RFA)

Dia. of Guide Thimbles (lower part), in. 0.397 ID X 0.429 OD V5H, V+)

0.397 ID X 0.439 OD (RFA)

Dia. of Instrument Guide Thimbles, in. 0.442 ID X 0.474 OD V5H, V+)

0.442 ID X 0.482 OD (RFA)

Fuel Rods Number 50,952 Outside Diameter, in. 0. 374 Diameter Gap, in. 0.0065 Clad Thickness, in. 0.0225 Clad Material Zircaloy-4 (V5H)

Zirlo' (V+, RFA)

Fuel Pellets Material uo 2 Sintered

• (1)

Dens~ty 95,5 Fuel Enrichments w/o(l)

Region 1 Region 2 4.6 Region 3A 4.2 Region 3B 4.6 Diameter, in. 0.3225 RFA Annular Pellet I.D., in. ( 2 ) 0.155 Length, in. 0.530 (STD) 0.387 (VSH, V+)

0.387 (RFA solid) ' 21

0. 462 or (2) 0.500 (RFA annular)

Mass of U0 2

Per Foot of Fuel Rod, lb/ft o. 364 (VSH, V+)

0.355 (RFA)

Rod Cluster Control Assemblies Neutron Absorber Ag-In-Cd Composition, percent 80, 15, 5 Diameter, in. 0.381 Density, lb/in. 0.367 Clad Material Type 316L, Ionnitride Surface Clad Thickness, in. 0.0185 Number of Clusters, full length 53 Number of Absorber Rods per Cluster 24 Full Length Assembly Weight (dry), lb 149

( 1}

Burnable Absorber Rods Material (PYREX) Borosilicate Glass Outside Diameter, in. 0.381 Inner Tube, OD, in. 0.1815 2 of 3 SGS-UFSAR Revision 23 October 17, 2007

TABLE 4.3-1 (Cont.)

REACTOR CORE DESCRIPTION Clad Material Stainless Steel Inner Tube Material Stainless Steel Boron Loading (w/o B o in glass rod) 12.5 2 3 Weight of Boron - 10 per foot of rod 1 lb/ft 0.00419 Material (WABA) Al o - B C Compound 2 3 4 B C Density (Fraction of Theoretical) 0.7 4

Absorber I.D. 1 in. 0.278 Absorber O.D. 1 in. 0.318 BA Clad Material Zircaloy Inner Clad Thickness/ in. 0.021 Inner Clad O.D., in. 0. 267 Outer Clad Thickness/ in. 0. 026 Outer Clad 0. D. I in. 0.381 Gap Material Helium Material ZrB 2

Content 1.570 to 2.355 mg B10 Jin. ( 1 )

Excess Reactivity Maximum Core Reactivity (Cold, Zero

.. .f c yc 1 e) ( 3)

Power, Beg~nn~ng o 1. 200 I

( 1)

Typical reload values. Current values are given in the appropriate unit and Core Loading Plan.

(2}

Robust Fuel Assembly (RFA) uses annular pellets at the top & bottom 6" of the fuel stack height. Middle 132" of fuel stack height is solid pellets.

The RFA annular pellet length starting with Salem Unit 1 Region 17 and Salem Unit 2 Region 15 is 0.500 inches.

(3)

Typical reload value. This parameter is function of energy requirements and number of burnable absorbers used.

3 of 3 SGS-UFSAR Revision 24 May 11, 2009

TABLE 4.3-2 NUCLEAR DESIGN PARAMETERS Core Average Linear Power 1 kW/ft 1 including densification effects 5.52 Total Heat Flux Hot Channel Factor, FQ

2. 4 0 1

Nuclear Enthalpy Rise Hot Channel Factor 1 ( )

RFA 1. 65, V5H 1. 57 Reactivity Coefficients Doppler Coefficient See Figures 4.3-17 and 4.3-18 Moderator Temperature Coefficient at Operating Conditions, pcm/°F(2) 0 to -44 Boron Coefficient in Primary Coolant 1 pcm/ppm -16 to -6 Rodded Moderator Density Coefficient 5

at Operating Conditions, pcm/gm/cc  ::; +0.52 X 10

( 1) Cycle-specific values based on accident analysis. The Core Operating Limits Report (COLR) contains the current cycle limits.

(2) Note: 1 pcm ~ (percent mille) 10 -5 Ap where Ap is calculated from two statepoint values of keff by ln (k 2 /k 1 l.

1 of 2 SGS-UFSAR Revision 24 May 11, 2009

TABLE 4.3-2 (Cont'd)

NUCLEAR DESIGN PARAMETERS

~ eff BOL, (EOL) 0.0075 (0.0044)

Control Rod Worths Rod Requirements See Table 4.3-3 Maximum Bank Worth, pcm < 2000 Maximum Ejected Rod Worth See Chapter 15

. (3)

Boron Concentratlons Refueling CB, ARI (K<0.95) 2 2050 To control at HZP, ARO, (K=l.O) 1700 1950 To control at HFP, ARO, (K=l.O):

0 MWD/MTU, No Xenon 1400 - 1700 150 MWD/MTU, Eq Xenon 1000 - 1250 1000 MWD/MTU, Eq Xenon 1000 - 1250 (3)

Typical reload values. Current cycle values are given in the appropriate NOR or COLR.

2 of 2 SGS-UFSAR Revision 24 May 11, 2009

TABLE 4.3-3 REACTIVITY REQUIREMENTS FOR ROD CLUSTER CONTROL ASSEMBLIES Reactivity Effects, Beginning-of-Life (IJ End-of Life (lJ Control requirements Fuel temperature (Doppler), %Ap 1. 32 1. 30 Moderator temperature, %Ap 0.11 1. 25 Void, %Ap 0.01 0.05 Redistribution, %Ap 0.50 0.85 Rod Insertion Allowance, %Ap 0.50 0.50 Rod Misalignment Relaxation 0.12 0.12 Penalty, %Ap

2. Total Control, %Ap 2.56 4.07

3. Estimated Rod Cluster Control Assembly Worth (53 Rods)

a. All full length assemblies inserted, %Ap 8.595 8.00

b. All but one (highest worth) assemblies inserted, %Ap 6.85 6.30

4. Estimated Rod Cluster Control Assembly credit with 10 percent adjustment to accommodate uncertainties (3b - 10 percent) , %Ap 6.17 5.67

5. Shutdown margin available (4-2),

1.60(2 )

%Ap 3.61 (2)

Typical reload values.

cycle specific rack-up.

See Plant Operations Package Table 10 for limiting End-of-Life The design basis minimum shutdown is 1.3 percent.

I 1 of 1 Revision 24 SGS-UFSAR May 11, 2009

TABLE 4.3-4 AXIAL STABILITY INDEX PWR CORE WITH A 12-FOOT HEIGHT Burnup Fz CB Stability Index (hr-1)

(MWD/T) (ppm) Exp Calc 1550 1. 34 1065 -0.041 -0.032 7700 1.27 700 -0.014 -0.006 Difference: +0.027 +0.026 1 of 1 SGS-UFSAR Revision 6 February 15, 1987

TABLE 4.3-5 TYPICAL NEUTRON FLUX LEVELS (n/cm 2-sec) AT FULL POWER E > 1.0 Mev 5.53 Kev < E 0.625 ev < E E < .625 ev

-< 1.0 Mev < 5.53 Kev (nv) 0 Core Center 6.51 X 10 13 1.12 X 1014 8.50 X 10 13 3.00 X 10 13 Core Outer Radius 3.23 X 10 13 5.74 X 10 13 4.63 X 10 13 8.60 X 1012 at Midheight Core Top, on Axis 1.53 X 1013 2.42 X 10 13 2.10 X 10 13 1.63 X 1013 Core Bottom, on Axis 2.36 X 1013 3.94 X 10 13 3.50 X 10 13 1.46 X 10 13 Pressure Vessel 2. 77 X 10 10 5. 75 X 10 10 6.03 X 10 10 8.38 X 1010 Inner Wall, Azimuthal Peak, Core Midheight 1 of 1 SGS-UFSAR Revision 6 February 15, 1987

TABLE 4.3-6 COMPARISON OF MEASURED AND CALCULATED DOPPLER DEFECTS Core Burnup Plant Fuel TyPe (MWD/MTU) Measure (pcm)(1) Calculated (pcm) 1 Air-filled 1800 1700 1710 2 Air-filled 7700 1300 1440 3 Air and 8460 1200 1210 helium filled

( 1 ) pcm = 105 x ln k /k 12

• SGS-UFSAR 1 of 1 Revision 6 February 15, 1987

TABLE 4.3-7 BENCHMARK CRITICAL EXPERIMENTS of ( 1 ) No. of LEOPARD Keff Experimental Sucklings Al Clad 14 1. 0012 SS Clad 19 0.9963 Borated H o 7 0.9989 2

Total 40 0.9985 C-Metal Al Clad 41 0.9995 Unclad 20 0.9990 Total 61 0.9993 Grand Total 101 0.9990 (1) in Reference 25.

1 of 1 SGS-UFSAR Revision 25 October 26, 2010

TABLE 4.3-8 SAXTON CORE II ISOTOPICS ROD MY, AXIAL ZONE 6 LEOPARD Atom Ratio Measured(!) 2a Precision (%) Calculation U-234/U 4.65 X 10-5 +/- 29 4.6o x 10-s U-235/U 5. 74 X 10- 3 +/- 0.9 5.73 X 10- 3 4

U-236/U 3.55 X 10 -4 +/- 5.6 3.74 X 10 U-236/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

-2 +/- 0.3

-2 Pu-241/Pu 6.014 X 10 5.74 X 10 5.81 X 10 -3 10 -3 Pu-242/Pu +/- 0.9 5.38 X Pu/U( 2 ) 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 (1) Reported in Reference 26 (2) Weight ratio .

• SGS-UFSAR 1 of 1 Revision 6 February 15, 1987

TABLE 4.3-9 CRITICAL BORON CONCENTRATIONS, 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

• SGS-UFSAR 1 of 1 Revision 6 February 15, 1987

TABLE 4.3-10 COMPARISON OF MEASURED AND CALCULATED ROD WORTH 2-Loop Plant, 121 Assemblies, 10-foot Core Measured (pcm) Calculated (pcm)

Group B 1885 1893 Group A 1530 1649 Shutdown 3050 2917 ESADA Critical ( 1 ), 0.69 inch Pitch, 240 2 w/o Puo , 8 percent Pu ,

2 9 Control Rods 6.21 in. rod separation 2250 2250 2.07 in. rod separation 4220 4160 1.38 in. rod separation 4100 4010 (1) Reported in Reference 27 .

• SGS-UFSAR 1 of 1 Revision 6 February 15, 1987

TABLE 4.3-11 COMPARISON OF MEASURED AND CALCULATED MODERATOR COEFFICIENTS AT HZP, BOL Plant Type/ Measured a. (1) Calculated a.

Control Bank Configuration (pcm/oF) lSO (pcmfOF) lSO 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 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

• (1) Isothermal coefficients, which include the Doppler effect in the fuel.

aiso = 10 5

In k2/kl I aT°F

• SGS-UFSAR 1 of 1 Revision 6 February 15, 1987

9 0

- --=

8 -~

II

!i 7 -8  %

0 liiC c::n liiC en 6 -l2 en I.U cr.. """'

0 a..

0 1-0 en 0

• 16 cn 5

.,7 Q:

I.U

/

II

(,:)

~ -20 1.1..

/ 1.1..

0 1.1..

0 z z: 3 -2ij 0 0 .....

I a..

=

1-I

(,)

I cn Q

0 2 -28 z:

0 Ql::

cr.. u

• t -32 0 -36 0 8 12 16 20 2" 28 32 36 "0 8URNUP.{ GWO/ MTU)

REVISION17 OCTOBER16 1998 PUBLICSERVICEELECTRICAND GAS COMPANY ProductionandConsumption ofHigherIsotopes SALEMNUCLEARGENERATING STATION UpdatedFSAR Fig 4.3*1

0 0

0 0

N 0

0 0

a,)

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PUBLICSERVICEELECTRICAND GAS COMPANY SALEMNUCLEARGENERATING STATION UpdatedFSAR Figure4.3-2

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SALEMNUCLEARGENERATING STATION UpdatedFSAR Figure 4.3-3

H G F E 0 c B A

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SALEMNUCLEARGENERATING STATION UpdatedFSAR Figure4.3-4

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• GDDDDDDD

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SALEMNUCLEARGENERATING STATION Updated FSAR Figure4.3*5

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BGGGBBD

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UpdatedFSAR Figure4.3-6

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• BDDDDDDD

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Updated FSAR Figure4.3*7

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z AO =-8.81 P = POWER D =FRACTIONAL INSERTIONOF BANK0 AO =AXIALOFFSET 0.5 0 80 90 100 0 10 20 30 40 50 60 70 PERCENTOF ACTIVECORE HEIGHTFROM BOTTOM TypicalAxialPowerShapesOcurring at PUBLICSERVICEELECTRICAND GAS COMPANY of Life Beginning SALEMNUCLEARGENERATINGSTATION UpdatedFSAR Fig4.3-8 REVISION17 OCTOBER161998

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0 10 20 30 40 50 60 70 80 90 100 PE'RCENTOF ACTIVECORE HEIGHTFROM BOTTOM PUBLICSERVICEELECTRICAND GAS COMPANY TypicalAxialPowerShapesOcurring at Middleof Ufe SALEMNUCLEARGENERATING STATION UpdatedFSAR Fig4.3-9 REVISION17 OCTOBER16 1998

2.0 1.5 a:

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P=POWER 0 = FRACTIONAL INSERTIONOF BANK0 A.O. = AXIALOFFSET 0.5 0.0 0 10 20 30 40 50 60 70 80 90 100 PERCENTOF ACTIVECORE HEIGHTFROM BOTTOM PUBLICSERVICEELECTRICANDGAS COMPANY TypicalAxialPowerShapesOcurringatEnd of Life SALEMNUCLEARGENERATING STATION UpdatedFSAR Fig4.3-10 REVISION17 OCTOBER16 1998

2.8 . . . . . - - - - - - - - - - - - - - - - - - - ,

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20 (j) 0 15 5

0

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~I REVISION17 OCTOBER16 1998 PeakLinearPowerDuringBoration/Dilution PUBLICSERVICEELECTRICAND GAS COMPANY Overpower Transients SALEMNUCLEARGENERATING STATION UpdatedFSAR Fig 4.3-13

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• BOL,MOL,EOL PUBLICSERVICEELECTRICAND GASCOMPANY (Typical)

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REVISION17 OCTOBER16 1998 Moderator Temperature Coemcient

• BOL PUBLICSERVICEELECTRICANDGAS COMPANY (Typical)

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REVISION17 OCTOBER16 1998 Moderator TemperatureCoefficient-EOL PUBLICSERVICEELECTRICAND GAS COMPANY (Typical)

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REVISION17 OCTOBER161998 Moderator Temperature Coefficient as a Function PUBLICSERVICEELECTRICAND GAS COMPANY ofBoronConcentration -BOL (Typical)

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SOL, EOL (Typical)

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DOPPLER COEFFICI EMT { pcm/' p )

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REVISION17 OCTOBER161998 Calculated and Measured DopplerDefectand PUBLICSERVICEELECTRICAND GAS COMPANY Coefficients at BOL,TwoLoopPlant,121 Assemblies SALEMNUCLEARGENERATINGSTATION 12-FtCore UpdatedFSAR FIG.4.3-32

1~00 ~--------------------------------------------------~

1200 1000 800 Q.

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200 0 ~--~----~----~----~----~--~~--~----~~~~

0 2000 6000 8000 I0000 12000 I ~000 16000 18000 BURMUP. t.ftiO /MTU REVISION17 OCTOBER16 1998 Comparison of Calculated and MeasuredBoron PUBLICSERVICEELECTRICAND GAS COMPANY Concentration for2-LoopPlant,121 Assemblies, 12*FtCore SALEMNUCLEARGENERATING STATION UpdatedFSAR Fig4.3*33

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0 0

~

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4.4 THERMAL AND HYDRAULIC DESIGN 4.4.1 Design Basis The overall objective of the thermal and hydraulic design of the reactor core is to provide adequate heat transfer which is compatible with the heat generation distribution in the core such that heat removal by the Reactor Coolant System (RCS) or the Emergency Core Cooling System (ECCS) (when applicable) assures that the following performance and safety criteria requirements are met:

1. Fuel damage* is not expected during normal operation and operational transients (Condition I) or any transient conditions arising from faults of moderate frequency (Condition II) It is not possible, however, to preclude a very small number of rod failures. These will be within the capability of the Plant Cleanup System and are consistent with the plant design bases.

2. The reactor can be brought to a safe state following a Condition III event with only a small fraction of fuel rods damaged* although sufficient fuel damage might occur to preclude resumption of operation without considerable outage time.

3. The reactor can be brought to a safe state and the core can be kept subcritical with acceptable heat transfer geometry following transients arising from Condition IV events.

• Fuel damage as used here is defined as penetration of the fission product barrier (i.e., the fuel rod clad).

4.4-1 SGS-UFSAR Revision 6 February 15, 1987

In order to satisfy the above criteria, the following design bases have been established for the thermal and hydraulic design of the reactor core.

4.4.1.1 Departure From Nucleate Boiling Design Basis Basis Departure from nucleate boiling (DNB) will not occur on at least 95 percent of the limiting fuel rods during normal operation and operational transients and any transient conditions arising from faults of moderate frequency (Condition I and II events) at a 95 percent confidence level.

Discussion The design method employed to meet the DNB design basis for the Vantage 5H, Vantage+, and RFA fuel assemblies is the Revised Thermal Design Procedure (RTDP), Reference 96. With the RTDP methodology, uncertainties in plant operating parameters, nuclear and thermal parameters, fuel fabrication parameters, computer codes and DNB correlation predictions are considered statistically to obtain DNB uncertainty factors. Based on the DNB uncertainty factors, RTDP design limit DNBR values are determined such that there is at least a 95 percent probability at a 95 percent confidence level that DNB will not occur on the most limiting fuel rod during normal operation and operational transients, and during transient conditions arising from faults of moderate frequency (Condition I and II events as defined in ANSI N18.2).

Uncertainties in the plant operating parameters (pressurizer pressure, primary coolant temperature, reactor power, and reactor coolant system flow) have been evaluated at Salem Units 1 and 2, Reference 114. Uncertainties in the power calorimetric at a 1.4% uprated reactor power have been evaluated at Salem Units 1 and 2, Reference 116. Only the random portion of the plant operating parameters uncertainties is included in the statistical combination.

Instrument bias is treated as a direct DNBR penalty. Since the parameter uncertainties are considered in determining the RTDP design limit DNBR values, the plant safety analyses are performed using input parameters at their nominal values.

4.4-2 SGS-UFSAR Revision 19 November 19, 2001

The RTDP design limit DNBR values are 1.24 for both typical and thimble cells of the Vantage 5H and Vantage+ fuels and 1. 24 and 1. 22 for the typical and thimble cells, respectively for RFA fuel.

The design limit DNBR values are used as a basis for the technical specifications and for consideration of the applicability of unreviewed safety questions as defined in 10 CFR 50.59.

To maintain DNBR margin to offset DNB penalties such as those due to fuel rod bow (see paragraph 4. 4. 2. 3. 5) and transition cores (see paragraph 4. 4. 2. 3. 6) ,

safety analyses were performed to DNBR limits higher than the design limit DNBR values. The difference between the design limit DNBRs and the safety analysis limit DNBRs results in available DNBR margin. The net DNBR margin, after consideration of all penalties, is available for operating and design flexibility.

The option of thimble plug removal has been included in all of the DNBR analyses performed for the Vantage 5H, Vantage+, and RFA fuel. The primary impact of thimble plug removal on the thermal hydraulic analysis is an increase in core bypass flow. Bypass flow is assumed to be ineffective for core heat removal. The increased bypass flow is included in all of the flow and DNBR values presented in Table 4.4-1.

Operation with thimble plugs in place reduces the core bypass flow through the fuel assembly thimble tubes. The reduction in core bypass flow for operation with the thimble plugs in place is a DNBR benefit. The increased margin associated with the use of a full compliment of thimble plugs can be used to offset DNBR penalties.

The Standard Thermal Design Procedure (STDP) is used for those analyses where RTDP is not applicable. In the STDP methodology, the parameters used in the analysis are treated in a conservative way from a DNBR standpoint. The parameter uncertainties are applied directly to the plant safety analyses input values to give the lowest minimum DNBR. The DNBR limit for STDP is the appropriate DNB correlation limit increased by sufficient margin to offset the applicable DNBR penalties.

4.4-2a SGS-UFSAR Revision 19 November 19, 2001

By preventing DNB, adequate heat transfer is assured between the fuel clad and the reactor coolant, thereby preventing clad damage as a result of inadequate cooling. Maximum fuel rod surface temperature is not a design basis as it will be within a few degrees of coolant temperature during operation in the nucleate boiling region. Limits provided by the nuclear control and protection systems are such that this design basis will be met for transients associated with Condition II events including overpower transients. There is an additional large DNBR margin at rated power operation and during normal operating transients.

4.4.1.2 Fuel Temperature Design Basis Basis During modes of operation associated with Condition I and Condition II events, the maximum fuel temperature shall be less than the melting temperature of uranium dioxide 4.4-2b SGS-UFSAR Revision 18 April 26, 2000

melting temperature for at least 95 percent of the peak kW/ft fuel rods will not be exceeded at the 95 percent confidence level. The melting temperature of U0 2 is taken as 5080DF (1) unirradiated and decreasing 58DF per 10,000 MWD/MTU.

By precluding U0 2 melting, the fuel geometry is preserved and possible adverse effects of molten U0 2 on the cladding are eliminated. To preclude center melting and as a basis for overpower protection system setpoints, a calculated centerline fuel temperature of 4700DF has been selected as the overpower limit.

This provides sufficient margin for uncertainties in the thermal evaluations as described in Section 4.4.2.10.1.

Discussion Fuel rod thermal evaluations are performed at rated power, maximum overpower and during transients at various burnups. These analyses assure that this design basis, as well as the fuel integrity design bases given in Section 4.2, are met. They also provide input for the evaluation of Condition III and IV faults given in Section 15.

4.4.1.3 Core Flow Design Basis Basis A minimum of 92.8 percent of the thermal flow rate will pass through the fuel rod region of the core and be effective for fuel rod cooling. Coolant flow through the thimble tubes as well as the leakage from the core barrel-baffle region into the core are not considered effective for heat removal.

Discussion Core cooling evaluations are based on the thermal flow rate (minimum flow) entering the reactor vessel. A maximum of 7.2 percent of this value is allotted as bypass flow. This includes rod cluster control (RCC) guide thimble cooling flow, 4.4-3 SGS-UFSAR Revision 18 April 26, 2000

head cooling flow, baffle leakage, and leakage to the vessel outlet nozzle.

The maximum bypass flow fraction of 7.2 percent assumes no plugging devices or burnable absorbers in the RCC guide thimble tubes, which do not contain RCC rods.

4.4.1.4 Hydrodynamic Stability Design Bases Basis Modes of operation associated with Condition I and II events shall not lead to hydrodynamic instability.

4.4.1.5 Other Considerations The above design bases together with the fuel clad and fuel assembly design bases given in Section 4. 2. 1. 1 are sufficiently comprehensive so additional limits are not required.

Fuel rod diametral gap characteristics, moderator-coolant flow velocity and distribution, and moderator void are not inherently limiting. Each of these parameters is incorporated into the thermal and hydraulic models used to ensure the above mentioned design criteria are met. For instance, the fuel rod diametral gap characteristics change with time (see Section 4.2.1.3.1) and the fuel rod in teg ri t y is evaluated on that basis . The effect of the moderator flow velocity and distribution (see Section 4.4.2.3) and moderator void distribution (see Section 4.4.2.5) are included in the core thermal (THINC) evaluation and thus affect the design bases.

Meeting the fuel clad integrity criteria covers possible effects of clad temperature limitations. As noted in Section 4.2.1.3.1, the fuel rod conditions change with time. A single clad temperature limit for Condition I or Condition II events is not appropriate since of necessity it would be overly conservative. A clad temperature limit is applied to the loss-of-coolant accident (LOCA) (Section 15.3.1), control rod ejection accident (2), and locked rotor accident (3).

4.4-4 SGS-UFSAR Revision 18 April 26, 2000

4.4.2 Description 4.4.2.1 Summary Comparison The Salem Unit 1 and Unit 2 reactors are designed to the appropriate DNBR limit as well as no fuel centerline melting during normal operation, operational transients, and faults of moderate frequency. The values of the thermal and hydraulic design parameters are presented in Table 4. 4-1, with comparisons between Vantage 5H, Vantage+, and RFA. Values specific to Standard Thermal Design Procedure and Revised Thermal Design Procedure are provided where appropriate.

4.4-5 SGS-UFSAR Revision 19 November 19, 2001

(The text on this page has been deleted) 4.4-6 SGS-UFSAR Revision 18 April 26, 2000

The effects on DNB of fuel densification have been evaluated utilizing the methods and models described in detail in Reference 7 and summarized in the following sections. The net effect of fuel densification is a reduction of 0.2 percent in the DNBR due to a slight increase in the linear power generation rate as described in Section 4.4.2.2.

4.4.2.2 Fuel and Cladding Temperatures (Including Densification)

Consistent with the thermal-hydraulic design bases described in Section 4.4.1, the following discussion pertains mainly to fuel pellet temperature evaluation.

A discussion of fuel clad integrity is presented in Section 4.2.1.3.1.

The thermal-hydraulic design assures that the maximum fuel temperature is below the melting point of uo2 (melting point of 5080DF (1) unirradiated and decreasing 58DF per 10,000 MWD/MTU) To preclude center melting and as a basis for overpower protection system setpoints, a calculated centerline fuel temperature of 4700DF has been selected as the overpower limit. This provides sufficient margin for uncertainties in the thermal evaluations as described in Section 4.4.2.10.1. The temperature distribution within the fuel pellet is predominantly a function of the local power density and the uo2 thermal conductivity. However, the computation of radial fuel temperature distributions combines crud, oxide, clad, gap, and pellet conductances. The factors which influence these conductances, such as gap size (or contact pressure), internal gas pressure, gas composition, pellet density, and radial power distribution within the pellet, etc, have been combined into a semi-empirical thermal model (see Section 4.2.1. 3.1) which includes a model for time dependent fuel densification described in References 7 and 108. This thermal model enables the determination of these factors and their net 4.4-7 SGS-UFSAR Revision 18 April 26, 2000

effects on temperature profiles. The temperature predictions have been compared to inpile fuel temperature measurements (8 through 14) and melt radius data (15,16) with good results.

Effect of Fuel Densification on Fuel Rod Temperatures Fuel densification results in fuel pellet shrinkage. This affects the fuel temperatures in the following ways:

1. Pellet radial shrinkage increases the pellet diametral gap which results in increased thermal resistance of the gap, and thus, higher fuel temperatures (See Section 4.2.1.3.1)

2. Pellet axial shrinkage may produce pellet to pellet gaps which result in local power spikes and thus, higher total heat flux hot channel factor, F, and local fuel temperatures.

Q

3. Pellet axial shrinkage will result in a fuel stack height reduction and an increase in the linear power generation rate (kW/ft) for a constant core power level. Using the methods described in Section 5.3 of Reference 7, the increase in linear power for the fuel rod specifications listed in Table 4.3-1 is 0.2 percent.

As described in Reference 7, fuel rod thermal evaluations (fuel centerline, average and surface temperatures) are determined throughout the fuel rod lifetime with consideration of time dependent densification. Maximum fuel average and surface temperatures, shown on Figure 4.4-1 for standard fuel and Figure 4. 4-1A for Vantage 5H, Vantage+, and RFA fuel as a function of linear power density (kW/ft), are peak values attained during the fuel lifetime.

Figure 4.4-2 for standard fuel and Figure 4.4-2A for Vantage 5H, Vantage+, and RFA fuel present the peak value of fuel centerline temperature versus linear power density which is attained during the fuel lifetime.

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The maximum pellet temperatures at the hot spot during full power steady state and at the maximum overpower trip point are shown in Table 4.4-1. The principal factors which are employed in the determination of the fuel temperature are discussed below.

4.4.2.2.1 Uranium Dioxide Thermal Conductivity The thermal conductivity of uo 2 was evaluated from data reported in References 17 through 29.

At the higher temperatures, thermal conductivity is best obtained by utilizing the integral conductivity to melt which can be determined with more certainty.

From an examination of the data, it has been concluded that the best estimate for the value of D 2800DC Kdt is 93 watts/em.

0 This conclusion is based on the integral values reported in References 15 and 29 through 33.

The design curve for the thermal conductivity is shown on Figure 4.4-3. The section of the curve at temperatures between ODC and 1300DC is in excellent agreement with the recommendation of the International Atomic Energy Agency panel (34) The section of the curve above 1300DC is derived for an integral value of 93 watts/cm(15, 29, 33).

Thermal conductivity for uo2 at 95 percent theoretical density can be represented best by the following equation:

1 -13 3 K + 8.775 X 10 T (4.4-1) 11.8 + 0.0238T where:

K watts/em-De 4.4-9 SGS-UFSAR Revision 18 April 26, 2000

T DC.

4.4.2.2.2 Radial Power Distribution in U0 2 Fuel Rods An accurate description of the radial power distribution as a function of burnup is needed for determining the power level for incipient fuel melting and other important performance parameters such as pellet thermal expansion, fuel swelling, and fission gas release rates.

This information on radial power distributions in U0 2 fuel rods is determined with the neutron transport theory code, LASER. The LASER code has been validated by comparing the code predictions on radial burnup and isotopic distributions with measured radial microdrill data (35, 36). A "radial power depression factor", f, is determined using radial power distributions predicted by LASER. The factor f enters into the determination of the pellet centerline temperature, T, relative to the pellet surface temperature, T, through the c s expression:

T c

D q'f k(T) dT (4.4-2)

T 40 s

where:

k(T) the thermal conductivity for UO with a uniform density distribution q' the linear power generation rate.

4.4.2.2.3 Gap Conductance The temperature drop across the pellet-clad gap is a function of the gap size and the thermal conductivity of the gas in the gap. The gap conductance model is selected such that when combined with 4.4-10 SGS-UFSAR Revision 18 April 26, 2000

the UO thermal conductivity model, the calculated fuel centerline temperatures reflect the inpile temperature measurements.

The temperature drop across the gap is calculated by assuming an annular gap conductance model of the following form:

Kg as h ( 4. 3 - 3) 5

+ 14.4 X 10-G 2

or an empirical correlation derived from thermocouple and melt radius data:

4.0 h 150 0 Kgas + (4.4- 3a)

0. 006 + 125 where:

K thermal conductivity of the gas mixture including a correction gas factor ( 37) for the accommodation coefficient for light gases (e.g., helium), Btu/hr-ft-DF.

D diametral gap size, ft.

The larger gap conductance value from these two equations is used to calculate the temperature drop across the gap for finite gaps.

For evaluations in which the pellet-clad gap is closed, a contact conductance is calculated. The contact conductance between UO and Zircaloy has been measured and found to be dependent on the contact pressure, composition of the gas at the interface and the surface roughness ( 37,38) . This information together with the surface roughness found in Westinghouse reactors leads to the following correlation:

4.4-11 SGS-UFSAR Revision 6 February 15, 1987

K gas h 0.6P + ( 4. 4-4) 6 14.4 X 10 where:

2 h contact conductance, Btu/hr-ft -DF p contact pressure, psi K thermal conductivity of gas mixture at the interface gas including a correction factor (37) for the accommodation coefficient for light gases (e.g., helium, Btu/hr-ft-DF).

4.4.2.2.4 Surface Heat Transfer Coefficients The fuel rod surface heat transfer coefficients during subcooled forced convection and nucleate boiling are presented in Section 4.4.2.8.1.

4.4.2.2.5 Fuel Clad Temperatures The outer surface of the fuel rod at the hot spot operates at a temperature of approximately 660DF for steady state operation at rated power throughout core life due to the onset of nucleate boiling. Initially (beginning-of-life), this temperature is that of the clad metal outer surface.

During operation over the life of the core, the buildup of oxides and crud on the fuel rod surface causes the clad surface temperature to increase. Allowance is made in the fuel center melt evaluation for this temperature rise. Since the thermal-hydraulic design basis limits DNB, adequate heat transfer is provided between the fuel clad and the reactor coolant so that the core thermal output is not limited by considerations of the 4.4-12 SGS-UFSAR Revision 6 February 15, 1987

clad temperature. Figure 4. 4-4 shows the axial variation of average clad temperature for the average power rod both at beginning and end-of-life.

Treatment of Peaking Factors The total heat flux hot channel factor, FQ' is defined by the ratio of the maximum to core average heat flux. As presented in Table 4.3-2 and discussed in Section 4.3.2.2.1, the design value FQ for normal operation is 2. 4 0, including fuel densification effects.

This results in peak local power of 13.3 kW/ft at full power conditions. As described in Section 4.3.2.2.6, the peak local power at the maximum overpower trip point is <22.4 kW/ft. The centerline temperature at this kW/ft must be below the melt temperature over the lifetime of the rod, including allowances for uncertainties. The melt temperature of unirradiated uo 2 is 5080DF (1) and decreases by 58DF per 10,000 MWD/MTU. From Figure 4. 4-2 for standard fuel and Figure 4.4-2A for Vantage 5H, Vantage+, and RFA fuel, it is evident that the centerline temperatures at the maximum overpower trip points for both units are far below those required to produce melting. Fuel centerline temperature at the maximum overpower trip point is presented in Table 4.4-1.

4.4.2.3 Critical Heat Flux Ratio or Departure from Nucleate Boiling Ratio and Mixing Technology The minimum DNBRs for the rated power, design overpower, and anticipated transient conditions are given in Table 4.4-1. The core average DNBR is not a safety-related i tern as it is not directly related to the minimum DNBR in the core, which occurs at some elevation in the limiting flow channel. Similarly, the DNBR at the hot spot is not directly safety-related. The minimum DNBR in the limiting flow channel will be downstream of the peak 4.4-13 SGS-UFSAR Revision 19 November 19, 2001

heat flux location (hot spot) due to the increased downstream enthalpy rise.

DNBRs are calculated by using the correlation and definitions described in the following Sections 4.4.2.3.1 and 4.4.2.3.2. The THINC-IV computer code (discussed in Section 4.4.3.4.1) is used to determine the flow distribution in the core and the local conditions in the hot channel for use in the DNB correlation. The use of hot channel factors is discussed in Section 4.4.3.2.1 (nuclear hot channel factors) and in Section 4.4.2.3.4 (engineering hot channel factors) .

4.4.2.3.1 Departure from Nucleate Boiling Technology The primary DNB correlation used for the analysis of the Vantage 5H and Vantage+ fuel is the WRB-1 correlation (Reference 89). The Primary DNB correlation used for the analysis of the RFA fuel (with Intermediate flow Mixer grids) is the WRB-2 correlation (Reference 97).

The WRB-1 correlation was developed based exclusively on the large bank of mixing vane grid rod bundle CHF data (over 1100 points) that Westinghouse has collected. The WRB-1 correlation, based on local fluid conditions, represents the rod bundle data with better accuracy over a wide range of variables than the previous correlation used in design. This correlation accounts directly for both typical and thimble cold wall effects, uniform and non-uniform heat flux profiles, and variations in rod heated length and in grid spacing.

The Applicable range of parameters for the WRB-1 correlation is:

Pressure 1440 ~ 2490 psia Local Mass Velocity 0.9 ~ Gloc/10 6 ~ 3.7 lb/ft 2 -hr Local Quality -0.2 ~ Xloc ~ 0.3 Heated Length, Inlet to CHF Location Lh ~ 14 feet Grid Spacing 13 ~ g~ ~ 32 inches Equivalent Hydraulic Diameter 0.37 ~de~ 0.60 inches Equivalent Heated Hydraulic Diameter 0.46 ~ dh ~ 0.59 inches Figure 4.4-5B shows measured critical heat flux plotted against predicted critical heat flux using the WRB-1 correlation.

A 95/95 correlation limit DNBR of 1.17 for the WRB-1 correlation has been approved by the NRC for Vantage 5H and Vantage+ fuel (Reference 92).

The WRB-2 DNB correlation was developed to take credit for the RFA Intermediate flow mixer (IFM) grid design. A limit of 1.17 is also applicable for the WRB-2 correlation. Figure 4. 4-5C shows measured critical heat flux plotted against predicted critical heat flux using the WRB-2 correlation.

4.4-14 SGS-UFSAR Revision 25 October 26, 2010

The applicable range of parameters for the WRB-2 correlation is Pressure 1440 ~ P ~ 2490 psia Local Mass Velocity 0.9 ~ Gloc/10 6 ~ 3.7 lb/ft 2 -hr Local Quality -0.1 ~ xl= ~ 0.3 Heated Length, Inlet to CHF Location Lh ~ 14 feet Grid Spacing 10 ~ g~ ~ 26 inches Equivalent Hydraulic Diameter 0.33 ~ de ~ 0.51 inches Equivalent Heated Hydraulic Diameter 0.45 ~ dh ~ 0.66 inches The use of the WRB-2 correlation has been conservatively modified to utilize a penalty above a certain high quality threshold within approved ranges (Reference 117)

The W-3 correlation (References 4 and 98) is used for all fuel types where the primary DNBR correlations are not applicable. The WRB-1 and WRB-2 correlations were developed based on mixing vane data and therefore are only applicable in the heated rod spans above the first mixing vane grid. The W-3 correlation, which does not take credit for mixing vane grids, is used to calculate the DNBR value in the heated region below the first mixing vane grid. In addition, the W-3 correlation is applied in the analysis of accident conditions where the system pressure is below the range of the primary correlations. For system pressures in the range of 500 to 1000 psia, the W-3 correlation limit is 1.45, Reference 99. for system pressures greater than 1000 psia, the W-3 correlation limit is 1.30. A cold wall factor (Reference 100) is applied to the W-3 DNB correlation to account for the presence of the unheated thimble surfaces.

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References 90 and 91 document the approval of the NRC that a 95/95 limit DNBR of 1.17 is appropriate for the STD and optimized fuel assemblies.

The use of the WRB-2 correlation with a 95/95 DNBR limit of 1.17 is applicable to the RFA fuel (References 111 and 112).

4.4.2.3.2 Definition of Departure from Nucleate Boiling Ratio The DNB heat flux ratio (DNBR) as applied to the standard fuel utilizing the W-3 "R" grid DNB correlation when all flow cell walls are heated is:

q" X F' DNB, N s DNBR (4.4-5) q" loc where:

q" DNB, EU q" ( 4. 4-6)

DNB,N F and q"DNB,EU is the uniform DNB heat flux as predicted by the W-3 DNB correlation (42) when all flow cell walls are heated.

F is the flux shape factor to account for non-uniform axial heat flux distributions (42) with the "C" term modified as in Reference 39.

F is the modified spacer factor described in Reference 5 using an axial grid s

spacing coefficient, K = 0.046, and a thermal diffusion coefficient (TDC) of s

0.038, based on the 26-inch grid 4.4-15a SGS-UFSAR Revision 18 April 26, 2000

spacing data. Since the actual grid spacing is approximately 20 inches, these values are conservative since the DNB performance was found to improve and TDC increases as axial grid spacing is decreased (References 40 and 43).

q" is the actual local heat flux.

loc The DNBR as applied to this design when a cold wall is present is:

4.4-15b SGS-UFSAR Revision 11 July 22, 1991

q" X F' DNB,N,CW S ( 4. 4-7)

DNBR q"

loc where:

q" x CWF DNB,EU,Dh q" ( 4. 4-8)

DNB,N,CW F where:

q" is the uniform DNB heat flux as predicted by the W-3 cold wall DNB,EU,Dh DNB correlation (39) when not all flow cell walls are heated (thimble cold wall cell)).

Values of minimum DNBR provided in Section 4.4.3.3 for STD fuel are the limiting value obtained by applying the above two definitions of DNBR to the appropriate cell (typical cell with all walls heated, or a thimble cold wall cell with a partial heated wall condition) . Approximately 15 percent in DNBR margin has been retained in all DNB analyses performed for this application.

Specifically, all DNBRs computed by Equations 4.4-5 and 4.4-7 have been multiplied by 0.86. Hence, if the value 1.30 is quoted, the actual calculated number (using either Equation 4.4-5 or 4.4-7) is 1.51. The basis for retaining this margin is discussed in detail in Section 4.4.2.1.

Histograms of both the "R" grid (15 x 15 geometry) data obtained from "R" grid rod bundle DNB tests (44,45) and the 17 x 17 geometry data obtained from similar tests (4,5) satisfy the criterion of being obtained from a normal distribution just as does the data used to develop the original W-3 DNB correlation with slight differences in the means and standard deviations from those of the W-3. However, the probability distribution curves for the "R" grid data and the 17 x 17 data (including the 0.88 multiplier) when compared to that of the W-3 correlation show that 4.4-16 SGS-UFSAR Revision 11 July 22, 1991

the approach which was valid for the original W-3 DNB correlation is conservatively applicable for the "R" and 17 x 17 data (with the 0.88 multiplier for the 17 x 17 data)

Standard fuel will continue to be designed to a minimum DNBR of 1. 30 for the peak rod or rods in the core. Based on the W-3 statistics, the proportion of such peak rods that will not reach DNB is 0. 95 or greater at a 95 percent confidence level.

The DNB heat flux ratio (DNBR) as applied to the Vantage 5H and Vantage+

designs utilizing the WRB-1 DNB correlation, and as applied to the RFA design utilizing the WRB-2 DNB correlation is:

q" DNB, N DNBR -

q" loc where:

q" DNB, EU q"

DNB,N F and q" is the uniform critical heat flux as predicted by the WRB-1 DNB DNB,EU correlation (Reference 89) or the WRB-2 correlation (Reference 97).

The procedures used in the evaluation of DNB margin for these applications show that the calculated minimum DNBR for the peak rod or rods in the core will be above the appropriate DNBR limit during Class I and II incidents, even when all the engineering hot channel factors described in Section 4.4.2.3.4 occur simultaneously in these channels. In reality the probability of this simultaneous occurrence is negligibly small and substantial increases in local heat flux or coolant temperature could be tolerated without violation of the design basis.

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4.4.2.3.3 Mixing Technology The rate of heat exchange by mixing between flow channels is proportional to the difference in the local mean fluid enthalpy of the respective channels, the local fluid density, and flow velocity. The proportionality is expressed by the dimensionless thermal diffusion coefficient, (TDC), which is defined as:

w' TDC DODD ( 4. 4-9)

OVa where:

w' flow exchange rate per unit length, lbs/ft-sec D fluld. . denslty, lbm I ft 3 v fluid velocity, ft/sec 4.4-17a SGS-UFSAR Revision 11 July 22, 1991

THIS PAGE INTENTIONALLY BLANK 4.4-17b SGS-UFSAR Revision 11 July 22, 1991

. 2 a lateral flow area between channels per unlt length, ft /ft The application of the TDC in the THINC analysis for determining the overall mixing effect or heat exchange rate is presented in Reference 46.

Various mixing tests have been performed by Westinghouse at Columbia University

( 43) . These series of tests, using the "R" mixing vane grid design on 13-, 26-and 32-inch grid spacing, were conducted in pressurized water loops at Reynolds numbers similar to that of a PWR core under the following single and two phase (subcooled boiling) flow conditions:

Pressure 1500 to 2400 psia Inlet temperature 332DF to 642DF 6 2 Mass velocity 1.0 to 3.5 x 10 lbm/hr-ft 5

Reynolds number 1.34 to 7.45 x 10 Bulk outlet quality -52.1 to -13.5 percent TDC is determined by comparing the THINC code predictions with the measured subchannel exit temperatures. Data for 26-inch axial grid spacing are presented on Figure 4.4-6 where the TDC is plotted versus the Reynolds number.

The thermal diffusion coefficient is found to be independent of Reynolds number, mass velocity, pressure, and quality over the ranges tested. The two-phase data (local, subcooled boiling) fell within the scatter of the single phase data. The effect of two-phase flow on the value of TDC has been demonstrated by Cadek (43), Rowe and Angle (47, 48), and Gonzalez, Santalo, and Griffith (49). In the subcooled boiling region the values of TDC were indistinguishable from the single phase values. In the quality region, Rowe and Angle show that in the case with rod spacing similar to that in 4.4-18 SGS-UFSAR Revision 6 February 15, 1987

pressurized water reactor (PWR) core geometry, the value of TDC increased with quality to a point and then decreased, but never below the single phase value.

Gonzalez, Santalo, and Griffith showed that the mixing coefficient increased as the void fraction increased.

The data from these tests on the "R" grid showed that a design TDC value of

0. 038 (for 26-inch grid spacing) can be used in determining the effect of coolant mixing in the THINC analysis.

A mixing test program similar to the one described above was conducted at Columbia University for the 17 x 17 geometry and mixing vane grids on 26-inch spacing (50). The mean value of TDC obtained from these tests was 0.059, and all data was well above the current design value of 0.038.

Since the actual reactor grid spacing is approximately 20 inches for the Vantage 5H and Vantage+ designs, and approximately 10 inches for the RFA design, additional margin is available, as the value of TDC increases as grid spacing decreases (43).

4.4.2.3.4 Engineering Hot-Channel Factors The total hot-channel factors for heat flux and enthalpy rise are defined as the maximum-to-core average ratios of these quanti ties. The heat flux hot-channel factor considers the local maximum linear heat generation rate at a point (the "hot spot") , and the enthalpy rise hot-channel factor involves the maximum integrated value along a channel (the "hot-channel").

Each of the total hot-channel factors considers a nuclear hot-channel factor (See Section 4.4.3.2) describing the neutron power distribution and an engineering hot-channel factor, which allows for variations in flow conditions and fabrication tolerances. The engineering hot-channel factors are made up of subfactors which account for the influence of the variations of fuel pellet diameter, density, enrichment, and eccentricity; fuel 4.4-19 SGS-UFSAR Revision 25 October 26, 2010

rod diameter, pitch, and bowing; inlet flow distribution; flow redistribution; and flow mixing.

E Heat Flux Engineering Hot-Channel Factor, FQ The heat flux engineering hot channel factor is used to evaluate the maximum liners heat generation rate of the core. This subfactor is determined by statistically combining the fabrication variations for fuel pellet diameter, density, and enrichment has a value of 1.03 at the 95 percent probability level with 95 percent confidence. As shown in Reference 15, no DNB penalty need be taken for the short, relatively low-intensity heat flux spikes caused by variations in the above parameters, as well as fuel pellet eccentricity and fuel rod diameter variation.

E Enthalpy Rise Engineering Hot-Channel Factor, F DH The effect of variations in flow conditions and fabrication tolerances on the hot-channel enthalpy rise is directly considered in the THINC core thermal sub channel analysis (See Section 4.4.3.4.1) under any reactor operating condition. The items considered contributing to the enthalpy rise engineering hot-channel factor are discussed below:

Pellet Diameter, Density and Enrichment Variations in pellet diameter, density, and enrichment are considered statistically in establishing the limit DNBRs (see paragraph 4. 4. 1. 1) for the Revised Thermal Design Procedure (Reference 96) employed in this application.

Uncertainties in these variables are determined from sampling of manufacturing data.

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(This text has been deleted)

Inlet Flow Maldistribution The consideration of inlet flow maldistribution in core thermal performances is discussed in Section 4.4.3.1.2. A design basis of 5 percent reduction in coolant flow to the hot assembly is used in the THINC-IV analysis.

Flow Redistribution The flow redistribution accounts for the reduction in flow in the hot-channel resulting from the high flow resistance in the channel due to the local or bulk boiling. The effect of the nonuniform power distribution is inherently considered in the THINC analysis for every operating condition which is evaluated.

Flow Mixing The subchannel mixing model incorporated in the THINC Code and used in reactor design is based on experimental data (51) discussed in Section 4.4.3.4.1. The mixing vanes incorporated in the spacer grid design induce additional flow mixing between the various flow channels in a fuel assembly as well as between adjacent assemblies. This mixing reduces the enthalpy rise in the hot-channel resulting from local power peaking or unfavorable mechanical tolerances.

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4.4.2.3.5 Effects of Rod Bow on DNBR The phenomenon of fuel rod bowing, as described in Reference 93, must be accounted for in the DNBR safety analysis of Condition I and Condition II events for each plant application. Applicable generic credits for margin resulting from retained conservatism in the evaluation of the DNBR and/or margin obtained from measured plant operating parameters (such as FN or core DH flow) which are less limiting than those required by the plant safety analysis -- can be used to offset the effect of rod bow.

For the safety analysis of the Salem units, sufficient DNBR margin was maintained (see Paragraph 4.4.1.1) to accommodate the full and low flow rod bow DNBR penalties identified in Reference 94. The referenced penalties are applicable to the Vantage 5H and Vantage+ fuel assembly analyses using the WRB-1 DNB correlation and to RFA assembly analyses using the WRB-2 DNB correlation.

This penalty is the maximum rod bow penalty at an assembly average burn up of 24,000 MWD/MTU. For burnups greater than 24,000 MWD/MTU, credit is taken for the effect of FDH burndown, due to the decrease in fissionable isotopes and the buildup of fission product inventory. Therefore, no additional rod bow penalty is required at burnups greater than 24,000 MWD/MTU.

In the upper spans of the RFA fuel assembly, additional restraint is provided with the Intermediate Flow Mixer (IFM) grids such that the grid-to-grid spacing in those spans with IFM grids is approximately 10 inches compared to approximately 20 inches in the other spans. Using the NRC approved scaling factor results in predicted channel closure in the limiting 10 inch spans of less than 50 percent closure. Therefore, no rod bow DNBR penalty is required in the 10 inch spans of the RFA safety analyses.

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The introduction of new fuel assembly design features in a core reload will result in mixed cores. It can typically take two or three cycles to completely transition to a full core of similar fuel assembly features. Generally, mixed (or transition) core impacts are a result of differences in pressure drop and localized flow distribution between adjacent assemblies with different grid types. The Westinghouse transition core DNB methodology is described in References 101, 102 and 103.

The transition from the 17 x 17 Standard design to Vantage 5H has been previously evaluated in References 104 and 105. With the implementation of the RFA design, IFM grids will be located in spans between the mid grids, where no grids exist in the Vantage 5H or Vantage+ assemblies. Test and analyses have confirmed that the RFA design is hydraulically compatible with the Vantage 5H and Vantage+ fuel designs (References 97, 111 and 115).

Transition cores are analyzed as if they were full cores of one assembly type (full core of Vantage 5H or RFA) and applying an applicable transition core DNBR penalty (Reference 106). The transition core DNBR penalty is a function of the number of new type fuel assemblies in the core. Therefore, the penalty is reduced each subsequent cycle. Since the Vantage+ assembly is dimensionally and hydraulically identical to Vantage 5H, introduction of this design resulted in no transition impact.

4.4-21b SGS-UFSAR Revision 18 April 26, 2000

4.4.2.4 Flux Tilt Considerations Significant quadrant power tilts are not anticipated during normal operation since this phenomenon is caused by some asymmetric perturbation. A dropped or misaligned rod cluster control assembly (RCCA) could cause changes in hot-channel factors; however, these events are analyzed separately in Section 15.

This discussion will be confined to flux tilts caused by x-y xenon transients, inlet temperature mismatches, enrichment variations within tolerances, etc.

N The design value of the enthalpy rise hot-channel factor F DH, which includes an 8 percent uncertainty (as discussed in Section 4. 3. 2. 2. 7) is assumed to be sufficiently conservative that flux tilts up to and including the alarm point (see Section 16.3. 10, Technical Specifications) will not result in values of greater than that assumed in this submittal. The design value of FQ does not include a specific allowance for quadrant flux tilts.

4.4.2.5 Void Fraction Distribution The calculated core average and the hot-subchannel maximum and average void fractions are presented in Table 4.4-3 for operation at full power. The void fraction distribution in the core at various radial and axial locations is presented in Reference 52. The void models used in the THINC-IV computer code are described in Section 4.4.2.8.3.

Since void formation due to subcooled boiling is an important promoter of interassembly flow redistribution, a sensitivity study was performed with THINC-IV using the void model referenced above (52)

The results of this study showed that because of the realistic crossflow model used in THINC-IV, the minimum DNBR in the hot-channel is relatively insensitive to variations in this model.

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The range of variations considered in this sensitivity study covered the maximum uncertainty range of the data used to develop each part of the void fraction correlation.

4.4.2.6 Core Coolant Flow Distribution Assembly average coolant mass velocity and enthalpy at various radial and axial core locations are given below. Coolant enthalpy rise and normalized core flow distributions are shown for the 4-foot elevation (1/3 of core height) on Figure 4.4-7, and 8-foot elevation (2/3 of core height) on Figure 4.4-8, and at the core exit on Figure 4. 4-9. These distributions are representative of a Westinghouse 4-loop plant. The THINC Code analysis for this case utilized a uniform core inlet enthalpy and inlet flow distribution.

4.4.2.7 Core Pressure Drops and Hydraulic Loads 4.4.2.7.1 Core Pressure Drops The analytical model and experimental data used to calculate the pressure drops shown in Table 4.4-1 are described in Section 4.4.2.8. The core pressure drop includes the eight grid fuel assembly, core support plate, and holddown plate pressure drops. The full power operation pressure drop values shown in the tables are the unrecoverable pressure drops across the vessel, including the inlet and outlet nozzles, and across the core. These pressure drops are based on the best estimate flow (most likely value for actual plant operating conditions) as described in Section 5.1.1. Section 5.1.1 also defines and describes the thermal design flow (minimum flow) which is the basis for reactor core thermal performance and the mechanical design flow (maximum flow) which is used in the mechanical design of the reactor vessel internals and fuel assemblies. Since the best estimate flow is that flow which is most likely to exist in an operating plant, the 4.4-23 SGS-UFSAR Revision 19 November 19, 2001

calculated core pressure drops in Table 4.4-1 are based on this best estimate flow rather than the thermal design flow.

Uncertainties associated with the core pressure drop values are discussed in Section 4.4.2.10.

4.4.2.7.2 Hydraulic Loads The fuel assembly holddown springs, Figure 4.2-2, are designed to keep the fuel assemblies in contact with the lower core plate under all Condition I and II events with the exception of the turbine overspeed transient associated with a loss-of-external load. The holddown springs are designed to tolerate the possibility of an over deflection associated with fuel assembly liftoff for this case and provide contact between the fuel assembly and the lower core plate following this transient. More adverse flow conditions occur during a LOCA as discussed in Section 15.

Hydraulic loads at normal operating conditions are calculated considering the mechanical design flow which is described in Section 5.1 and accounting for the minimum core bypass flow based on manufacturing tolerances. Core hydraulic loads at cold plant startup conditions are adjusted to account for the coolant density difference. Conservative core hydraulic loads for a pump over speed transient, which create flow rates 20 percent greater than the mechanical design flow, are evaluated to be greater than twice the fuel assembly weight.

Core hydraulic loads were measured during the prototype assembly tests described in Section 1. 5. Reference 6 contains a detailed discussion of the results.

The Vantage 5H, Vantage+ and RFA designs have been shown to be hydraulically compatible in References 111 and 115.

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4.4.2.8 Correlation and Physical Data 4.4.2.8.1 Surface Heat Transfer Coefficients Forced convection heat transfer coefficients are obtained from the familiar Dittus-Boelter correlation (53), with the properties evaluated at bulk fluid conditions:

0.8 ( c ]0.4

~e = 0.023 ( D~G ]  : (4.4-10) where:

2 h heat transfer coefficient, Btu/hr-ft -DF De equivalent diameter, ft 2

K thermal conductivity, Btu/hr-ft -DF G mass veloclty,. lb I hr-ft 2 D dynamic viscosity, lb/ft-hr c heat capacity, Btu/lb-DF p

This correlation has been shown to be conservative (54) for rod bundle geometries with pitch to diameter ratios in the range used by PWRs.

The onset of nucleate boiling occurs when the clad wall temperature reaches the amount of superheat predicted by Thorn's (55) correlation. After this occurrence the outer clad wall temperature is determined by:

0 5 DT (0.072 exp (-P/1260)) (q") " (4.4-11) sat where:

DT wall superheat, T T, OF sat w sat 4.4-25 SGS-UFSAR Revision 18 April 26, 2000

2 q" wall heat flux, Btu/hr-ft p pressure, psia T outer clad wall temperature, DF w

T saturation temperature of coolant at P, DF sat 4.4.2.8.2 Total Core and Vessel Pressure Drop Unrecoverable pressure losses occur as a result of viscous drag (friction) and/ or geometry changes (form) in the fluid flow path. The flow field is assumed to be incompressible, turbulent, single-phase water. These assumptions apply to the core and vessel pressure drop calculations for the purpose of establishing the primary loop flow rate. Two-phase considerations are neglected in the vessel pressure drop evaluation because the core average void is negligible (See Section 4.4.2.5 and Table 4.4-3. Two phase flow considerations in the core thermal subchannel analyses are considered and the models are discussed in Section 4.4.3.1.3. Core and vessel pressure losses are calculated by equations of the form:

2 L D V DP (K+F D DDDDDDDDDD (4.4-12)

L D 2 g (144) ec where:

.2 unrecoverable pressure drop, lbf/ln 3

D fluid density, lb /ft m

L length, ft D equivalent diameter, ft e

v fluid velocity, ft/sec 4.4-26 SGS-UFSAR Revision 11 July 22, 1991

lb - ft m

g 32.174 DDDDDDDDD c 2 lb -sec f

K form loss coefficient, dimensionless F friction loss coefficient, dimensionless Fluid density is assumed to be constant at the appropriate value for each component in the core and vessel. Because of the complex core and vessel flow geometry, precise analytical values for the form and friction loss coefficients are not available. Therefore, experimental values for these coefficients are obtained from geometrically similar models.

Values are quoted in Table 4. 4-1 for unrecoverable pressure loss across the reactor vessel, including the inlet and outlet nozzles, and across the core.

The results of full scale tests of core components and fuel assemblies were utilized in developing the core pressure loss characteristic. The pressure drop for the vessel was obtained by combining the core loss with correlation of 1/7th scale model hydraulic test data on a number of vessels (56,57) and from loss relationships (58) Moody (59) curves were used to obtain the single phase friction factors.

Tests of the primary coolant loop flow rates will be made (see Section 4.4.4.1) prior to initial criticality to verify that the flow rates used in the design, which were determined in part from the pressure losses calculated by the method described here, are conservative.

4.4.2.8.3 Void Fraction Correlation There are three separate void regions considered in flow boiling in a PWR as illustrated on Figure 4.4-10. They are the wall void region (no bubble detachment), the subcooled boiling region (bubble detachment), and the bulk boiling region.

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In the wall void region, the point where local boiling begins is determined when the clad temperature reaches the amount of superheat predicted by Thorn's (55) correlation (discussed in Section 4.4.2.8.1). The void fraction in this region is calculated using Maurer's (60) relationship. The bubble detachment point, where the superheated bubbles break away from the wall, is determined by using Griffith's (61) relationship.

The void fraction in the subcooled boiling region (that is, after the detachment point) is calculated from the Bowring (62) correlation. This correlation predicts the void fraction from the detachment point to the bulk boiling region.

The void fraction in the bulk boiling region is predicted by using homogeneous flow theory and assuming no slip. The void fraction in this region is therefore a function only of the thermodynamic quality.

4.4.2.9 Thermal Effects of Operational Transients DNB core safety limits are generated as a function of coolant temperature, pressure, core power, and axial power imbalance. Steady-state operation within these safety limits ensures that the minimum DNBR is not less than the appropriate DNBR limit. Figure 15. 1-1 shows the DNBR limit lines and the resulting overtemperature D-T trip lines (which become part of the Technical Specifications), plotted as D-T vs T-average for various pressures.

This system provides adequate protection against anticipated operational transients that are slow with respect to fluid transport delays in the primary system. In addition, for fast transients, e.g., uncontrolled rod bank withdrawal at power incident (Section 15. 1. 2), specific protection functions are provided as described in Section 7.2 and the use of these protection functions is described in Section 15 (See Table 15.1-2).

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The thermal response of the fuel rod is discussed in Section 4.4.3.7.

4.4.2.10 Uncertainties in Estimates 4.4.2.10.1 Uncertainties in Fuel and Clad Temperatures As discussed in Section 4.4.2.2, the fuel temperature is a function of crud, oxide, clad, gap, and pellet conductances. Uncertainties in the fuel temperature calculation are essentially of two types: fabrication uncertainties such as variations in the pellet and clad dimensions and the pellet density; and model uncertainties such as variations in the pellet conductivity and the gap conductance. These uncertainties have been quantified by comparison of the thermal model to the inpile thermocouple measurements (8 through 14), by out-of-pile measurements of the fuel and clad properties ( 17 through 2 8) , and by measurements of the fuel and clad dimensions during fabrication. The resulting uncertainties are then used in all evaluations involving the fuel temperature.

The effect of densification on fuel temperature uncertainties is presented in Reference 7 and also included in the calculation of total uncertainty.

In addition to the temperature uncertainty described above, the measurement uncertainty in determining the local power and the effect of density and enrichment variations on the local power are considered in establishing the heat flux hot-channel factor. These uncertainties are described in Section

4. 3. 2. 2.

Reactor trip setpoints as specified in the Technical Specifications include allowance for instrument and measurement uncertainties such as calorimetric error, instrument drift, and channel reproducibility, temperature measurement uncertainties, noise, and heat capacity variations.

Uncertainty in determining the cladding temperature results from uncertainties in the crud and oxide thicknesses. Because of the excellent heat transfer between the surface of the rod and the 4.4-29 SGS-UFSAR Revision 18 April 26, 2000

coolant, the film temperature drop does not appreciably contribute to the uncertainty.

4.4.2.10.2 Uncertainties in Pressure Drops Core and vessel pressure drops based on the best estimate flow, as discussed in Section 5.1, are quoted in Table 4.4-1. The uncertainties quoted are based on the uncertainties in both the test results and the analytical extension of these values to the reactor application.

A major use of the core and vessel pressure drops is to determine the primary system coolant flow rates as discussed in Section 5.1. In addition, as discussed in Section 4. 4. 4 .1, tests on the primary system prior to initial criticality will be made to verify that a conservative primary system coolant flow rate has been used in the design and analyses of the plant.

4.4.2.10.3 Uncertainties Due to Inlet Flow Maldistribution The effects of uncertainties in the inlet flow maldistribution criteria used in the core thermal analyses are discussed in Section 4.4.3.1.2.

4.4.2.10.4 Uncertainty in DNB Correlation The uncertainty in the DNB correlation (Section 4. 4. 2. 3) can be written as a statement on the probability of not being in DNB based on the statistics of the DNB data. This is discussed in Section 4.4.2.3.2.

4.4.2.10.5 Uncertainties in DNBR Calculations The uncertainties in the DNBRs calculated by THINC analysis (see Section

4. 4. 3. 4. 1) due to uncertainties in the nuclear peaking factors are accounted for by applying conservatively high values of the nuclear peaking factors and including measurement error 4.4-30 SGS-UFSAR Revision 11 July 22, 1991

Allowances in the statistical evaluation of the DNBR limit (see paragraph 4.4.1.1) using the Revised Thermal Design Procedure (Reference 96). In addition, conservative values for the engineering hot channel factors are used as discussed in Section 4.4.2.3.4.

The results of a sensitivity study (52) with THINC-IV show that the minimum DNBR in the hot channel is relatively insensitive to variations in the core-N wide radial power distribution (for the same value of F DH).

The ability of the THINC-IV computer code to accurately predict flow and enthalpy distributions in rod bundles is discussed in Section 4.4.3.4.1 and in Reference 63. Studies have been performed (52) to determine the sensitivity of the minimum DNBR in the hot-channel to the void fraction correlation (also see Section 4.4.2.8.3); the inlet velocity and exit pressure distributions assumed as boundary conditions for the analysis; and the grid pressure loss coefficients. The results of these studies show that the minimum DNBR in the hot channel is relatively insensitive to variations in these parameters. The range of variations considered in these studies covered the range of possible variations in these parameters.

4.4.2.10.6 Uncertainties in Flow Rates The uncertainties associated with loop flow rates are discussed in Section 5.1.

For core thermal performance evaluations, a thermal design loop flow is used which is less than the best estimate loop flow (approximately 4 percent). In addition, another 7. 2 percent of the thermal design flow is assumed to be ineffective for core heat removal capability because it bypasses the core through the various available vessel flow paths described in Section 4.4.3.1.1.

4.4.2.10.7 Uncertainties in Hydraulic Loads As discussed in Section 4. 4. 2. 7. 2, hydraulic loads on the fuel assembly are evaluated for a pump overspeed transient which creates flow rates 20 percent greater than the mechanical design 4.4-31 SGS-UFSAR Revision 18 April 26, 2000

flow. The mechanical design flow is greater than the best estimate or most likely flow rate value for the actual plant operating condition.

4.4.2.10.8 Uncertainty in Mixing Coefficient The value of the mixing coefficient, TDC, used in THINC analyses for this application is 0.038. The mean value of TDC obtained in the "R" grid mixing tests described in Section 4.4.2.3.3 was 0.042 (for 26-inch grid spacing). The value of 0.038 is one standard deviation below the mean value, and~ 90 percent of the data gives values of TDC greater than 0.038 (43).

The results of the mixing tests done on 17 x 17 geometry, as discussed in Section 4.4.2.3.3, had a mean value of TDC of 0.059 and standard deviation of D

= 0.007. Hence the current design value of TDC is almost 3 standard deviations below the mean for 26-inch grid spacing.

4.4.2.11Plant Configuration Data Plant configuration data for the thermal-hydraulic and fluid systems external to the core are provided in the appropriate Sections 5, 6, and 9.

Implementation of the ECCS is discussed in Section 15. Some specific areas of interest are the following:

1. Total coolant flow rates for the RCS are provided in Table 5.1-1. Flow rates employed in the evaluation of the core are presented in Section

4. 4.

2. Total RCS volume including pressurizer and surge line, RCS liquid volume including pressurizer water at steady state power conditions are given in Table 5.1-1.

3. The flow path length through each volume may be calculated from physical data provided in the above referenced tables.

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4. The height of fluid in each component of the RCS may be determined from the physical data presented in Section 5.5. The components of the RCS are water filled during power operation with the pressurizer being approximately 60 percent water filled.

5. The elevation of components of the RCS relative to the reactor vessel are shown in Section 5.1. Components of the ECCS are to be located so as to meet the criteria for net positive suction head (NPSH) described in Section 6.1.

6. Line lengths and sizes for the Safety Injection System are determined so as to guarantee a total system resistance which will provide, as a minimum, the fluid delivery rates assumed in the safety analyses described in Section 15.

4.4.3 Evaluation 4.4.3.1 Core Hydraulics 4.4.3.1.1 Flow Paths Considered in Core Pressure Drop and Thermal Design The following flow paths or core bypass flow are considered:

1. Flow through the spray nozzles into the upper head for head cooling purposes

2. Flow entering into the RCC guide thimbles to cool the control rods

3. Leakage flow from the vessel inlet nozzle directly to the vessel outlet nozzle through the gap between the vessel and the barrel 4.4-33 SGS-UFSAR Revision 6 February 15, 1987

4. Flow entering into the core from the baffle-barrel region through the gaps between the baffle plates The above contributions are evaluated to confirm that the design value of core bypass flow is met. The design value of core bypass flow for Salem Units 1 and 2 is equal to 7.2 percent of the total vessel flow. Of the total allowance, 4.0 percent is associated with the core and the remainder associated with the internals. Calculations have been performed accounting for drawing tolerances and uncertainties in pressure losses. Based on these calculations, the core bypass flow for Salem Units 1 and 2 is no greater than the value quoted above.

Flow model test results for the flow path through the reactor are discussed in Section 4.4.2.8.2.

4.4.3.1.2 Inlet Flow Distributions Data has been considered from several 1/7 scale hydraulic reactor model tests (56, 57,64) in arriving at the core inlet flow maldistribution criteria to be used in the THINC analyses (See Section 4.4.3.4.1). THINC-I analyses made using this data have indicated that a conservative design basis is to consider a 5 percent reduction in the flow to the hot assembly (65). The same design basis of 5 percent reduction to the hot assembly inlet is used in THINC-IV analyses.

The experimental error estimated in the inlet velocity distribution has been considered as outlined in Reference 52 where the sensi ti vi ty of changes in inlet velocity distributions to hot channel thermal performance is shown to be small. Studies (52) made with the improved THINC model (THINC-IV) show that it is adequate to use the 5 percent reduction in inlet flow to the hot 4.4-34 SGS-UFSAR Revision 18 April 26, 2000

assembly for a loop out of service based on the experimental data in References 56 and 57.

The effect of the total flow rate on the inlet velocity distribution was studied in the experiments of Reference 56. As was expected, on the basis of the theoretical analysis, no significant variation could be found in inlet velocity distribution with reduced flow rate.

(This text has been deleted) 4.4.3.1.3 Empirical Friction Factor Correlations Two empirical friction factor correlations are used in the THINC-IV computer code (described in Section 4.4.3.4.1).

The friction factor in the axial direction, parallel to the fuel rod axis, is evaluated using the Novendstern-Sandberg correlation (66). This correlation consists of the following:

1. For isothermal conditions, this correlation uses the Moody (59) friction factor including surface roughness effects.

2. Under single-phase heating conditions a factor is applied based on the values of the coolant density and viscosity at the temperature of the heated surface and at the bulk coolant temperature.

3. Under two-phase flow conditions the homogeneous flow model proposed by Owens (67) is used with a modification to account for a mass velocity and heat flux effect.

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The flow in the lateral directions, normal to the fuel rod axis, views the reactor core as a large tube bank. Thus, the lateral friction factor proposed by Idel'chik (58) is applicable. This correlation is of the form:

-0.2 F ARe (4.4-13)

L L where:

A is a function of the rod pitch and diameter as given in Reference 58.

Re is the lateral Reynolds number based on the rod diameter.

L Extensive comparisons of THINC-IV predictions using these correlations to experimental data are given in Reference 63, and verify the applicability of these correlations in PWR design.

4.4.3.2 Influence of Power Distribution The core power distribution which is largely established at beginning of life by fuel enrichment, loading pattern, and core power level is also a function of variables such as control rod worth and position, and fuel depletion throughout lifetime. Radial power distributions in various planes of the core are often illustrated for general interest; however, the core radial enthalpy rise distribution as determined by the integral of power up each channel is of greater importance for DNB analyses. These radial power distributions,

. N .

characterlzed by F DH (deflned in Section 4.3.2.2.2) as well as axial heat flux profiles are discussed in the following two sections.

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N 4.4.3.2.1 Nuclear Enthalpy Rise Hot Channel Factor, F DH" Given the local power density q' (kw/ft) at a point x, y, z in a core with N fuel rods and height H, H

J Max q' (x, y, z) dz hot rod power 0 (4.4-14) average rod power 1

-n LanrodJo q' (x, y, z) dz H

where:

x,y are the position coordinates of the hot rod.

0 0

.N The way in whlch F DH is used in the DNB calculation is important. The location of minimum DNBR depends on the axial profile and the value of DNBR depends on the enthalpy rise to that point. Basically, the maximum value of the rod integral is used to identify the most likely rod for minimum DNBR. An axial power profile is obtained which, when normalized to the design value of FN DH' recreates the axial heat flux along the limiting rod. The surrounding rods are assumed to have the same axial profile with rod average powers which are typical of distributions found in hot assemblies. In this manner worst case axial profiles can be combined with worst case radial distributions for reference DNB calculations.

N It should be noted again that F DH is an integral and is used as such in the DNB calculations. Local heat fluxes are obtained by using hot channel and adjacent channel explicit power shapes which take into account variations in horizontal power shapes throughout the core. The sensitivity of the THINC-IV analysis to radial power shapes is discussed in Reference 52.

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For operation at a fraction P of full power, the design FNDH used is given by:

N RTP F

DH

= F DH (1 + PFDH (1-P)) (4.4-15)

RTP where F is the limit at Rated Thermal Power (RTP) specified in the COLR DH and PFDH is the Power Factor Multiplier for FNDH specified in the COLR N

The permitted relaxation of F DH is included in the DNB protection setpoints and allows radial power shape changes with rod insertion to the insertion limits (68,88), thus allowing greater flexibility in the nuclear design.

4.4.3.2.2 Axial Heat Flux Distributions As discussed in Section 4.3.2.2, the axial heat flux distribution can vary as a result of rod motion, power change, or due to spatial xenon transients which may occur in the axial direction. Consequently, it is necessary to measure the axial power imbalance by means of the ex-core nuclear detectors (as discussed in Section 4.3.2.2.7) and protect the core from excessive axial power imbalance. The Reactor Trip System provides automatic reduction of the trip setpoint in the overtemperature DT channels on excessive axial power imbalance; that is, when an extremely large axial offset corresponds to an axial shape which could lead to a DNBR which is less than that calculated for the reference DNB design axial shape.

The reference DNB design axial shape used for the Reactor Trip System application is a chopped cosine shape with a peak average value of 1.55.

4.4.3.3 Core Thermal Response A general summary of the steady-state thermal-hydraulic design parameters including thermal output, flow rates, etc, is provided in Table 4.4-1 for all loops in operation.

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As stated in Section 4.4.1, the design bases of the application are to prevent DNB and to prevent fuel melting for Condition I and II events. The protective systems described in Section 7 (Instrumentation and Controls) are designed to meet these bases. The response of the core to Condition II transients is given in Section 15.

4.4.3.4 Analytical Techniques 4.4.3.4.1 Core Analysis The objective of reactor core thermal design is to determine the maximum heat removal capability in all flow subchannels and show that the core safety limits, as presented in the Technical Specifications, are not exceeded while compounding engineering and nuclear effects. The thermal design takes into account local variations in dimensions, power generation, flow redistribution, and mixing. THINC-IV is a realistic three-dimensional matrix model which has been developed to account for hydraulic and nuclear effects on the enthalpy rise in the core (52, 63) . The behavior of the hot assembly is determined by superimposing the power distribution among the assemblies upon the inlet flow distribution while allowing for flow mixing and flow distribution between assemblies. The average flow and enthalpy in the hottest assembly is obtained from the core-wide, assembly-by-assembly analysis. The local variations in power, fuel rod and pellet fabrication, and mixing within the hottest assembly are then superimposed on the average conditions of the hottest assembly in order to determine the conditions in the hot channel.

The following sections describe the use of the THINC Code in the thermal-hydraulic design evaluation to determine the conditions in the hot channel and to assure that the safety-related design bases are not violated.

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Steady State Analysis The THINC-IV computer program as approved by the Nuclear Regulatory Commission (NRC) (69) is used to determine coolant density, mass velocity, enthalpy, vapor void, static pressure, and DNBR distributions along parallel flow channels within a reactor core under all expected operating conditions. The core region being studied is considered to be made up of a number of contiguous elements in a rectangular array extending the full length of the core. An element may represent any region of the core from a single assembly to a subchannel.

The momentum and energy exchange between elements in the array are described by the equations for the conservation of energy and mass, the axial momentum equation, and two lateral momentum equations which couple each element with its neighbors. The momentum equations used in THINC-IV are similar to the Euler Equations (70) except that frictional loss terms have been incorporated which represent the combined effects of frictional and form drag due to the presence of grids and fuel assembly nozzles in the core. The crossflow resistance model used in the lateral momentum equations was developed from experimental data for flow normal to tube banks (58, 71) The energy equation for each element also contains additional terms which represent the energy gain or loss due to the crossflow between elements.

The unique feature in THINC-IV is that lateral momentum equations, which include both inertial and crossflow resistance terms, have been incorporated into the calculational scheme. This differentiates THINC-IV from other thermal-hydraulic programs in which only the lateral resistance term is modeled. Another important consideration in THINC-IV is that the entire velocity field is solved, en masse, by a field equation while in other codes such as THINC-1 (46) and COBRA (72) the solutions are obtained by stepwise integration throughout the array.

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The resulting formulation of the conservation equations is more rigorous for THINC-IV; therefore, the solution is more accurate. In addition, the solution method is complex and some simplifying techniques must be employed. Since the reactor flow is chiefly in the axial direction, the core flow field is primarily one-dimensional and it is reasonable to assume that the lateral velocities and the parameter gradients are larger in the axial direction than the lateral direction. Therefore, a perturbation technique can be used to represent the axial and lateral parameters in the conservation equations. The lateral velocity components are regarded as perturbed quantities which are smaller than the unperturbed and perturbed component with the unperturbed component equaling the core average value at a given elevation and the perturbed value is the difference between the local value and the unperturbed component. Since the magnitudes of the unperturbed and perturbed parameters are significantly different, they can be solved separately. The unperturbed equations are one-dimensional and can be solved with the resulting solutions becoming the coefficients of the perturbed equations. An iterative method is then used to solve the system of perturbed equations which couples all the elements in the array.

Three THINC-IV computer runs constitute one design run: a core wide analysis, a hot assembly analysis, and a hot subchannel analysis. While the calculational method is identical for each run, the elements which are modeled by THINC-IV change from run-to-run. In the core wide analysis, the computational elements represent full fuel assemblies. In the second computation the elements represent a quarter of the hot assembly. For the last computation, a quarter of the hot assembly is analyzed and each individual subchannel is represented as a computational element.

The first computation is a core-wide, assembly-by-assembly analysis which uses an inlet velocity distribution modeled from experimental reactor models (56,57,64) (see Section 4.4.3.1.2) In the core-wide analysis the core is considered to be made up of a number of contiguous fuel assemblies divided axially into 4.4-41 SGS-UFSAR Revision 6 February 15, 1987

increments of equal length. The system of perturbed and unperturbed equations are solved for this array giving the flow, enthalpy, pressure drop, temperature, void fraction in each assembly. The system of equations is solved using the specified inlet velocity distribution and a known exit pressure condition at the top of the core. This computation determines the interassembly energy and flow exchange at each elevation for the hot assembly.

THINC-IV stores this information then uses it for the subsequent hot assembly analysis.

In the second computation, each computational element represents one-fourth of the hot assembly. The inlet flow and the amount of momentum and energy interchange at each elevation are known from the previous core wide calculation. The same solution technique is used to solve for the local parameters in the hot one-quarter assembly.

While the second computation provides an overall analysis of the thermal and hydraulic behavior of the hot quarter assembly, it does not consider the individual channels in the hot assembly. The third computation further divides the hot assembly into channels consisting of individual fuel rods to form flow channels. The local variations in power, fuel rod and pellet fabrication, fuel rod spacing and mixing (engineering hot-channel factors) within the hottest assembly are imposed on the average conditions of the hottest fuel assembly in order to determine the conditions in the hot-channel. The engineering hot-channel factors are described in Section 4.4.2.3.4.

Experimental Verifications An experimental verification (63) of the THINC-IV analysis for core wide, assembly-to-assembly enthalpy rises as well as enthalpy rise in a nonuniformly heated rod bundle have been obtained.

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In these experimental tests, the system pressure, inlet temperature, mass flow rate, and heat fluxes were typical of present PWR core designs.

During the operation of a reactor, various core monitoring systems obtain measured data indicating the core performance. Assembly power distributions and assembly mixed mean temperature are measured and can be converted into the proper three-dimensional power input needed for the THINC programs. This data can then be used to verify the Westinghouse thermal-hydraulic design codes.

One standard startup test is the natural circulation test in which the core is held at a very low power (~2 percent) and the pumps are turned off. The core will then be cooled by the natural circulation currents created by the power differences in the core. During natural circulation, a thermal siphoning effect occurs resulting in the hotter assemblies gaining flow, thereby creating significant interassembly crossflow. As described in the preceding section the most important feature of THINC-IV is the method by which crossflow is evaluated. Thus, tests with significant crossflow are of more value in the code verification. Interassembly crossflow is caused by radial variations in pressure. Radial pressure gradients are in turn caused by variations in the axial pressure drops in different assemblies. Under normal operating conditions (sub cooled forced convection) the axial pressure drop is due mainly to friction losses. Since all assemblies have the same geometry, all assemblies have nearly the same axial pressure drops and crossflow velocities are small. However, under natural circulation conditions (low flow) the axial pressure drop is due primarily to the difference in elevation head (or coolant density) between assemblies (axial velocity is low and therefore axial friction losses are small) . This phenomenon can result in relatively large radial pressure gradients and, therefore, higher crossflow velocities than at normal reactor operating conditions.

4.4-43 SGS-UFSAR Revision 19 November 19, 2001

The in-core instrumentation was used to obtain the assembly-by-assembly core power distribution during a natural circulation test. Assembly exit temperatures during the natural circulation test on a 157 assembly, three loop plant were predicted using THINC-IV. The predicted data points were plotted as assembly temperature rise versus assembly power and a least squares fitting program used to generate an equation which best fit the data. The result is the straight line presented on Figure 4. 4-11. The measured assembly exit temperatures are reasonably uniform, as indicated on this figure, and are predicted closely by the THINC-IV code. This agreement verifies the lateral momentum equations and the crossflow resistance model used in THINC-IV. The larger crossflow resistance used in THINC-I reduces flow redistribution, so that THINC-IV gives better agreement with the experimental data.

Data has also been obtained for Westinghouse plants operating from 67 percent to 101 percent of full power. A representative cross section of the data obtained from a two-loop and a three-loop reactor were analyzed to verify the THINC-IV calculational method. The THINC-IV predictions were compared with the experimental data as shown on Figures 4.4-12 and 4.4-13. The predicted assembly exit temperatures were compared with the measured exit temperatures for each data run. The standard deviation of the measured and predicted assembly exit temperatures were calculated and compared for both THINC-IV and THINC-I and are given in Table 4. 4-4. As the standard deviations indicate, THINC-IV generally fits the data somewhat more accurately than THINC-I. For the core inlet temperatures and power of the data examined, the coolant flow is essentially single phase. Thus, one would expect little inter-assembly crossflow and small differences between THINC-IV and THINC-I predictions as seen in the tables. Both codes are conservative and predict exit temperatures higher than measured values for the high powered assemblies.

An experimental verification of the THINC-IV subchannel calculation method has been obtained from exit temperature 4.4-44 SGS-UFSAR Revision 6 February 15, 1987

measurements in a non-uniformly heated rod bundle (73). The inner nine heater rods were operated at approximately 20 percent more power than the outer rods to create a typical PWR intra-assembly power distribution. The rod bundle was divided into 36 subchannels and the temperature rise was calculated by THINC-IV using the measured flow and power for each experimental test.

Figure 4. 4-14 shows, for a typical run, a comparison of the measured and predicted temperature rises as a function of the power density in the channel.

The measurements represent an average of two to four measurements taken in various quadrants of the bundle. It is seen that the THINC-IV results predict the temperature gradient across the bundle very well. On Figure 4.4-15, the measured and predicted temperature rises are compared for a series of runs at different pressures, flows, and power levels.

Again, the measured points represent the average of the measurements taken in the various quadrants. It is seen that the THINC-IV predictions provide a good representation of the data.

Extensive additional experimental verification is presented in Reference 63.

The THINC-IV analysis is based on a knowledge and understanding of the heat transfer and hydrodynamic behavior of the coolant flow and the mechanical characteristics of the fuel elements. The use of the THINC-IV analysis provides a realistic evaluation of the core performance and is used in the thermal analyses as described above.

Transient Analysis The THINC-IV thermal-hydraulic computer code does not have a transient capability. Since the third section of the THINC-I program (46) does have this capability, this code (THINC-III) continues to be used for transient DNB analysis.

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The conservation equations needed for the transient analysis are included in THINC-III by adding the necessary accumulation terms to the conservation equations used in the steady-state (THINC-I) analysis. The input description must now include one or more of the following time-dependent arrays:

1. Inlet flow variation

2. Heat flux distribution

3. Inlet pressure history At the beginning of the transient, the calculation procedure is carried out as in the steady-state analysis. The THINC-III code is first run in the steady-state mode to ensure conservatism with respect to THINC-IV and in order to provide the steady-state initial conditions at the start of the transient. The time is incremented by an amount determined either by the user or by the program itself. At each new time step the calculations are carried out with the addition of the accumulation terms which are evaluated using the information from the previous time step. This procedure is continued until a preset maximum time is reached.

At preselected intervals, a complete description of the coolant parameter distributions within the array as well as DNBR are printed out. In this manner the variation of any parameter with time can be readily determined.

At various times during the transient, steady state THINC-IV is applied to show that the application of the transient version of THINC-I is conservative.

The THINC-III code does not have the capability for evaluating fuel rod thermal response. This is treated by the methods described in Section 15.1.9.

4.4-46 SGS-UFSAR Revision 6 February 15, 1987

4.4.3.4.2 Fuel Temperatures As discussed in Section 4.4.2.2, the fuel rod behavior is evaluated utilizing a semi-empirical thermal model which considers, in addition to the thermal aspects, such items as clad creep, fuel swelling, fission gas release, release of absorbed gases, cladding corrosion and elastic deflection, and helium solubility.

A detailed description of the thermal model can be found in Reference 3 with the modifications for time-dependent densification given in Reference 7.

4.4.3.4.3 Hydrodynamic Instability The analytical methods used to access hydraulic instability are discussed in Section 4.4.3.5.

4.4.3.5 Hydrodynamic and Flow Power Coupled Instability Boiling flow may be susceptible to thermodynamic instabilities (74). These instabilities are undesirable in reactors since they may cause a change in thermo-hydraulic conditions that may lead to a reduction in the DNB heat flux relative to that observed during a steady flow condition or to undesired forced vibrations of core components. Therefore, a thermo-hydraulic design criterion was developed which states that modes of operation under Condition I and II events shall not lead to thermo-hydrodynamic instabilities.

Two specific types of flow instabilities are considered for Westinghouse PWR operation. These are the Ledinegg or flow excursion type of static instability and the density wave type of dynamic instability.

A Ledinegg instability involves a sudden change in flow rate from one steady state to another. This instability occurs (74) when 4.4-47 SGS-UFSAR Revision 6 February 15, 1987

the slope of the reactor coolant system pressure drop-flow rate curve DDP (ODD I becomes algebraically smaller than the loop supply DG internal (pump head) pressure drop-flow rate curve ---- I oM The oG external DDP DDP criterion for stability is thus ODD> DOD DG DG internal external The Westinghouse pump head curve has a negative slope (DDP/DG/external < 0) whereas the Reactor Coolant System pressure drop-flow curve has a positive slope (DDP/DG/internal > 0) over the Condition I and Condition II operational ranges. Thus, the Ledinegg instability will not occur.

The mechanism of density wave oscillations in a heated channel has been described by Lahey and Moody (75) Briefly, an inlet flow fluctuation produces an enthalpy perturbation. This perturbs the length and the pressure drop of the single phase region and causes quality or void perturbations in the two-phase regions which travel up the channel with the flow. The quality and length perturbations in the two-phase region create two-phase pressure drop perturbations. However, since the total pressure drop across the core is maintained by the characteristics of the fluid system external to the core, then the two-phase pressure drop perturbation feeds back to the single phase region. These resulting perturbations can be either attenuated or self-sustained.

A simple method has been developed by Ishii (76) for parallel closed channel systems to evaluate whether a given condition is stable with respect to the density wave type of dynamic instability. This method had been used to assess the stability of 4.4-48 SGS-UFSAR Revision 6 February 15, 1987

typical Westinghouse reactor designs (77,78,79) under Condition I and II operation. The results indicate that a large margin to density wave instability exists, e.g., increases on the order of 200 percent of rated reactor power would be required for the predicted inception of this type of instability.

The application of the method of Ishii (76) to Westinghouse reactor designs is conservative due to the parallel open channel feature of Westinghouse PWR cores. For such cores, there is little resistance to lateral flow leaving the flow channels of high power density. There is also energy transfer from channels of high power density to lower power density channels. This coupling with cooler channels has led to the opinion that an open channel configuration is more stable than the above closed channel analysis under the same boundary conditions. Flow stability tests (80) have been conducted where the closed channel systems were shown to be less stable than when the same channels were cross connected at several locations. The cross connections were such that the resistance to channel crossflow and enthalpy perturbations would be greater than that which would exist in a PWR core which has a relatively low resistance to crossflow.

Flow instabilities which have been observed have occurred almost exclusively in closed channel systems operating at low pressures relative to the Westinghouse PWR operating pressures. Kao, Morgan, and Parker (81) analyzed parallel closed channel stability experiments simulating a reactor core flow. These experiments were conducted at pressures up to 2200 psia. The results showed that for flow and power levels typical of power reactor conditions, no flow oscillations could be induced above 1200 psia.

Additional evidence that flow instabilities do not adversely affect thermal margin is provided by the data from the rod bundle DNB tests. Many Westinghouse rod bundles have been tested over wide ranges of operating conditions with no evidence of premature DNB or of inconsistent data which might be indicative of flow instabilities in the rod bundle.

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In summary, it is concluded that thermo-hydrodynamic instabilities will not occur under Condition I and II modes of operation for Westinghouse PWR reactor designs. A large power margin, greater than doubling rated power, exists due to predicted inception of such instabilities. Analysis has been performed which shows that minor plant-to-plant differences in Westinghouse reactor designs such as fuel assembly arrays, core power flow ratios, fuel assembly length, etc, will not result in gross deterioration of the above power margins.

4.4.3.6 Temperature Transient Effects Analysis Waterlogging damage of a fuel rod could occur as a consequence of a power increase on a rod after water has entered the fuel rod through a clad defect.

Water entry will continue until the fuel rod internal pressure is equal to the reactor coolant pressure. A subsequent power increase raises the temperature and, hence, could raise the pressure of the water contained within the fuel rod. The increase in hydrostatic pressure within the fuel rod then drives a portion of the water from the fuel rod through the water entry defect. Clad distortion and/or rupture can occur if the fuel rod internal pressure increase is excessive due to insufficient venting of water to the reactor coolant. This occurs when there is both a rapid increase in the temperature of the water within the fuel rod and a small defect. Zircaloy clad fuel rods which have failed due to waterlogging ( 82, 83) indicate that very rapid power transients are required for fuel failure. Normal operational transients are limited to about 40 cal/gm-min. (peak rod) while the Spert tests (82) indicate that 120 to 150 cal/gm are required to rupture the clad even with very short transients

( 5. 5 msec. period) . Release of the internal fuel rod pressure is expected to have a minimal effect on the Reactor Coolant System (82) and is not expected to result in failure of additional fuel rods (83). Ejection of fuel pellet fragments into the coolant stream is not expected (82,83). A clad breech due to water logging is thus expected to be similar to any fuel rod failure mechanism which exposes fuel pellets to the reactor 4.4-50 SGS-UFSAR Revision 6 February 15, 1987

coolant stream. Waterlogging has not been identified as the mechanism for clad distortion or perforation of any Westinghouse Zircaloy-4 clad fuel rods.

High fuel rod internal gas pressure could cause clad failure. One of the fuel rod design bases (Section 4.2.1.1.1) is that the fuel rod internal gas pressure is limited to a value below that which could cause ( 1) the diametral gap to increase due to outward cladding creep during steady-state operation, and (2) extensive DNB propagation to occur. During operational transients, fuel rod clad rupture due to high internal gas pressure is precluded by meeting the above design basis.

4.4.3.7 Potentially Damaging Temperature Effects During Transients The fuel rod experiences many operational transients (intentional maneuvers) during its residence in the core. A number of thermal effects must be considered when analyzing the fuel rod performance.

The clad can be in contact with the fuel pellet at some time in the fuel lifetime. Clad-pellet interaction occurs if the fuel pellet temperature is increased after the clad is in contact with the pellet. Clad-pellet interaction is discussed in Section 4.2.1.3.1.

The potential effects of operation with waterlogged fuel are discussed in Section 4.4.3.6 which concluded that waterlogging is not a concern during operational transients.

Clad flattening, as noted in Section 4.2.1.3.1, has been observed in some operating power reactors. Thermal expansion (axial) of the fuel rod stack against a flattened section of clad could cause failure of the clad. This is no longer a concern because clad flattening is precluded during the fuel residence in the core (See Section 4.2.1.3.1).

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There can be differential thermal expansion between the fuel rods and the guide thimbles during a transient. Excessive bowing of the fuel rods could occur if the grid assemblies did not allow axial movement of the fuel rods relative to the grids. Thermal expansion of the fuel rods is considered in the grid design so that axial loads imposed on the fuel rods during a thermal transient will not result in excessively bowed fuel rods (See Section 4.2.1.2.2) 4.4.3.8 Energy Release During Fuel Element Burnout As discussed in Section 4. 4. 3. 3 the core is protected from going through DNB over the full range of possible operating conditions. At full power nominal operation, the probability of a rod going through DNB is less than 0.1 percent at a 95 percent confidence level. In the extremely unlikely event that DNB should occur, the clad temperature will rise due to the steam blanketing at the rod surface and the consequent degradation in heat transfer. During this time there is a potential for a chemical reaction between the cladding and the coolant. However, because of the relatively good film boiling heat transfer following DNB, the energy release resulting from this reaction is insignificant compared to the power produced by the fuel.

DNB With Physical Burnout - Westinghouse (73) has conducted DNB tests in a 25-rod bundle where physical burnout occurred with one rod. After this occurrence, the 25-rod test section was used for several days to obtain more DNB data from the other rods in the bundle. The burnout and deformation of the rod did not affect the performance of neighboring rods in the test section during the burnout or the validity of the subsequent DNB data points as predicted by the W-3 correlation. No occurrences of flow instability or other abnormal operation were observed.

4.4-52 SGS-UFSAR Revision 11 July 22, 1991

DNB With Return to Nucleate Boiling - Additional DNB tests have been conducted by Westinghouse (84) in 19 and 21 rod bundles. In these tests, DNB without physical burnout was experienced more than once on single rods in the bundles for short periods of time. Each time, a reduction in power of approximately 10 percent was sufficient to reestablish nucleate boiling on the surface of the rod. During these and subsequent tests, no adverse effects were observed on this rod or any other rod in the bundle as a consequence of operating in DNB.

4.4.3.9 Energy Release or Rupture of Waterlogged Fuel Elements A full discussion of waterlogging, including energy release, is contained in Section 4.4.3.6. It is noted that the resulting energy release is not expected to affect neighboring fuel rods.

4.4.3.10Fuel Rod Behavior Effects from Coolant Flow Blockage Coolant flow blockages can occur within the coolant channels of a fuel assembly or external to the reactor core. The effects of fuel assembly blockage within the assembly on fuel rod behavior are more pronounced than external blockages of the same magnitude. In both cases the flow blockages cause local reductions in coolant flow. The amount of local flow reduction, where it occurs in the reactor, and how far along the flow stream the reduction persists are considerations which will influence the fuel rod behavior. The effects of coolant flow blockages in terms of maintaining rated core performance are determined both by analytical and experimental methods. The experimental data are usually used to augment analytical tools such as computer programs similar to the THINC-IV program. Inspection of the DNB correlation (Section 4. 4. 2. 3 and References 42, 98, 97 and 113) shows that the predicted DNBR is dependent upon the local values of quality and mass velocity.

The THINC-IV code is capable of predicting the effects of local flow blockages on DNBR within the fuel assembly on a subchannel basis, regardless of where the flow blockage occurs. In Reference 4.4-53 SGS-UFSAR Revision 18 April 26, 2000

63, it is shown that for a fuel assembly similar to the Westinghouse design, THINC-IV accurately predicts the flow distribution within the fuel assembly when the inlet nozzle is completely blocked. Full recovery of the flow was found to occur about 30 inches downstream of the blockage. With the reference reactor operating at the nominal full power conditions specified in Table 4.4-1 the effects of an increase in enthalpy and decrease in mass velocity in the lower portion of the fuel assembly would not result in the reactor reaching the DNBR limit.

From a review of the open literature it is concluded that flow blockage in "open lattice cores" similar to the Westinghouse cores causes flow perturbations which are local to the blockage. For instance, A. Oktsubo et al (85) show that the mean bundle velocity is approached asymptotically about 4 inches downstream from a flow blockage in a single flow cell. Similar results were also found for 2 and 3 cells completely blocked. Basmer et al (86) tested an open lattice fuel assembly in which 41 percent of the subchannels were completely blocked in the center of the test bundle between spacer grids. Their results show the stagnant zone behind the flow blockage essentially disappears after 1.65 L/De or about 5 inches for their test bundle. They also found that leakage flow through the blockage tended to shorten the stagnant zone or in essence the complete recovery length. Thus, local flow blockages within a fuel assembly have little effect on subchannel enthalpy rise. The reduction in local mass velocity is then the main parameter which affects the DNBR. If the Salem Units 1 and 2 were operating at full power and nominal steady state conditions as specified in Table 4.4-1, a reduction in local mass velocity of 80 percent in the Vantage 5H and Vantage+ fuel and 60 percent in the RFA fuel would be required to reduce the DNBR to the DNBR limit. The above mass velocity effect on the DNB correlation was based on the assumption of fully developed flow along the full channel length. In reality a local flow blockage is expected to promote turbulence and thus would likely not affect DNBR at all.

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Coolant flow blockages induce local cross flows as well as promote turbulence.

Fuel rod behavior is changed under the influence of a sufficiently high crossflow component. Fuel rod vibration could occur, caused by this crossflow component, through vortex shedding or turbulent mechanisms. If the crossflow velocity exceeds the limit established for fluid elastic stability, large amplitude whirling results. The limits for a controlled vibration mechanism are established from studies of vortex shedding and turbulent pressure fluctuations. Crossflow velocity above the established limits can lead to mechanical wear of the fuel rods at the grid support locations. Fuel rod wear due to flow induced vibration is considered in the fuel rod fretting evaluation (Section 4. 2) .

4.4.4 Testing and Verification 4.4.4.1 Tests Prior to Initial Criticality A reactor coolant flow test, as noted in Item 5 of Table 13.3-1, is performed following fuel loading but prior to initial criticality. Coolant loop pressure drop data is obtained in this test. This data, in conjunction with coolant pump performance information, allow determination of the coolant flow rates at reactor operating conditions. This test verifies that proper coolant flow rates have been used in the core thermal and hydraulic analysis.

4.4.4.2 Initial Power and Plant Operation Core power distribution measurements are made at several core power levels (see Section 4. 3. 2. 2. 7) . These tests are used to ensure that conservative peaking factors are used in the core thermal and hydraulic analysis.

Additional demonstration of the overall conservatism of the THINC analysis was obtained by comparing THINC predictions to in-core thermocouple measurements.

These measurements were performed on the Zion reactor (Reference 87) .

(Distribution of in-core 4.4-55 SGS-UFSAR Revision 6 February 15, 1987

instrumentation for Salem Units 1 and 2 is shown on Figures 4.4-16 and 4.4-17, respectively). No further in-pile testing is planned.

4.4.4.3 Component and Fuel Inspections Inspections performed on the manufactured fuel are delineated in Section 4.2.1.4. Fabrication measurements critical to thermal and hydraulic analysis are obtained to verify that the engineering hot channel factors employed in the design analyses (Section 4.2.3.4) are met.

4.4.4.4 Augmented Startup Test Program Salem Unit No. 1 will participate in the augmented startup test program as necessary in accordance with the program outlined in the Westinghouse Topical Report, Augmented Startup and Cycle 1 Physics Program, WCAP-8575.

4.4.5 References for Section 4.4

1. Christensen, J. A.; Allio, R. J.; and Biancheria, A., "Melting Point of Irradiated U0 2 , " WCAP-6065, February 1965.

2. Risher, D. H., Jr., "An Evaluation of the Rod Ejection Accident in Westinghouse Pressurized Water Reactors Using Spatial Kinetics Methods,"

WCAP-7588, Revision 1, December 1971.

3. Supplemental information on fuel design transmitted from R. Salvatori, Westinghouse NES, to D. Knuth, AEC, as attachments to letters NS-SL-518

( 12/22/72) (NS-SL-521 ( 12/29/72), NS-SL-524 ( 12/29/72) and NS-SL-543 (1/12/73), (Proprietary), and supplemental information on fuel design transmitted from R. Salvatori, Westinghouse NES, to D. Knuth, AEC, as attachments to letters NS-SL-527 (1/2/73) and NS-SL-544 (1/12/73).

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4. Hill, K. W.; Motley, F. E.; Cadek, F. F.; and Wenzel, A. H., "Effect of 17 x 17 Fuel Assembly Geometry on DNB," WCAP-8296-P-A (Proprietary) and WCAP-8297-A (Nonproprietary), February 1975.

5. Motley, F. E.; Wenzel, A. H.; and Cadek, F. F., "Critical Heat Flux Testing of 17 x 17 Fuel Assembly Geometry with 22 Inch Grid Spacing,"

WCAP-8536 (Proprietary) and WCAP-8537 (Nonproprietary), May 1975.

6. Nakazato, S. and DeMario, E. E., "Hydraulic Flow Test of the 17 x 17 Fuel Assembly," WCAP-8278, February 1974 (Proprietary) and WCAP-8279, February 1974.

7. Hellman, J. M. (Ed.), "Fuel Densification Experimental Results and Model for Reactor Application," WCAP-8218-P-A (Proprietary) and WCAP-8219-A (Nonproprietary), March 1975.

8. Kjaerheim, G. and Rolstad, E., In Pile Determination of UO Thermal Conductivity, Density Effects and Gap Conductance," HPR-80, December 1967.

9. Kj aerheim, G. , "In-Pile Measurements of Centre Fuel Temperatures and Thermal Conductivity Determination of Oxide Fuels," Paper IFA-175 presented at the European Atomic Energy Society Symposium on Performance Experience of Water-Cooled Power Reactor Fuel, Stockholm, Sweden, October 21-22, 1969.

10. Cohen, I.; Lustman, B.; and Eichenberg, J. D., "Measurement of the Thermal Conductivity of Metal-Clad Uranium Oxide Rods During Irradiation," WAPD-228, 1960.

11. Clough, D. J. and Sayers, J. B., "The Measurement of the Thermal Conductivity of uo 2 under Irradiation in the Temperature Range 1500-1600DC," AERE-R-4690, UKAEA Research Group, Harwell, December 1964.

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12. Stora, J. P.; DeBernardy DeSigoyer B.; Delamas, R.; Deschamps, P.; Ringot, C.; and Lavaud, B., "Thermal Conductivity of Sintered Uranium Oxide under In-Pile Conditions," EURAEC-1095, 1964.

13. Devold, I., "A Study of the Temperature Distribution in uo 2

Reactor Fuel Elements," AE-318, Aktiebolaget Atomenergi, Stockholm, Sweden, 1968.

14. Balfour, M. G.; Christensen, J. A.; and Ferrari, H. M., "In-Pile Measurement of uo Thermal Conductivity," Final Report, WCAP-2923, 1966.

2

15. Duncan, R.N., "Rabbit Capsule Irradiation of UO," CVTR Project, CVNA-142, June 1962.

16. Nelson, R. C.; Coplin, D. H.; Lyons, M. F.; and Weidenbaum, B., "Fission Gas Release from uo2 Fuel Rods with Gross Central Melting, II GEAP-4572, July 1964.

17. Howard, V. C. and Gulvin, T. G., "Thermal Conductivity Determinations of Uranium Dioxide by a Radial Flow Method," UKAEA IG-Report 51, November 1960.

18. Lucks, C. F. and Deem, H. W., "Thermal Conductivity and Electrical Conductivity of uo II In Progress Reports Relating to Civilian 2'

Applications, BMI-1448 (Rev.) for June 1960; BMI-1489 (Rev.) for December 1960 and BMI-1518 (Rev.) for May 1961.

19. Daniel, J. L.; Matolich, J., Jr.; and Deem, H. W., "Thermal Conductivity of U0 ," HW-69945, September 1962.

2

20. Feith, A. D., "Thermal Conductivity of uo2 by a Radial Heat Flow Method, II TID-21668, 1962.

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21. Vogt, J.; Grandell, L.; and Runfors, U., "Determination of the Thermal Conductivity of Unirradiated Uranium Dioxide," AB Atomenergi Report RMB-527, 1964, Quoted by IAEA Technical Report Series No. 59, "Thermal Conductivity of Uranium Dioxide."

22. Nishij ima, T.; Kawada, T.; and Ishihata, A., "Thermal Conductivity of Sintered uo 2 and A1 u at High Temperatures," J. American Ceramic Society, 2 3 48, 31-34, 1965.

23. Ainscough, J. B. and Wheeler, M.J., "The Thermal Diffusivity and Thermal Conductivity of Sintered Uranium Dioxide," in "Proceedings of the Seventh Conference on Thermal Conductivity," p. 467, National Bureau of Standards, Washington, 1968.

24. Godfrey, T. G.; Fulkerson, W.; Killie, J. P.; Moore, J. P.; and McElroy, D.

L., "Thermal Conductivity of Uranium Dioxide and Armco Iron by an Improved Radial Heat Flow Technique," ORNL-3556, June 1964.

25. Stora, J. P. et al, "Thermal Conductivity of Sintered Uranium Oxide Under In-Pile Conditions," EURAEC-1095, August 1964.

26. Bush, A. J., "Apparatus for Measuring Thermal Conductivity to 2500DC,"

Westinghouse Research Laboratories Report 64-1P6-401-R3 (Proprietary),

February 1965.

27. Asamoto, R. R.; Anselin, F. L.; and Conti, A. E., "The Effect of Density on the Thermal Conductivity of Uranium Dioxide," GEAP-5493, April 1968.

28. Kruger, 0. L., "Heat Transfer Properties of Uranium and Plutonium Dioxide,"

Paper 11-N-68F presented at the Fall Meeting of Nuclear Division of the American Ceramic Society, Pittsburgh, September 1968.

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29. Gyllander, J. A., "In-Pile Determination of the Thermal Conductivity of uo 2

in the Range 500-2500DC," AE-411, January 1971.

30. Lyons, M. F. et al, "UO Powder and Pellet Thermal Conductivity During 2

Irradiation," GEAP-5100-1, March 1966.

31. Coplin, D. H. et al, "The Thermal Conductivity of uo 2 by Direct In-Reactor Measurements," GEAP-5100-6, March 1968.

32. Bain, A. S., "The Heat Rating Required to Produce Center Melting in Various uo 2 Fuels," ASTM Special Technical Publication, No. 306, pp. 30-46, Philadelphia, 1962.

33. Stora, J. p., "In-Reactor Measurements of the Integrated Thermal Conductivity of uo - Effect of Porosity," Trans. ANS, 13, 137-138, 1970.

2

34. International Atomic Energy Agency, "Thermal Conductivity of Uranium Dioxide," Report of the Panel held in Vienna, April 1965, IAEA Technical Reports Series, No. 59, Vienna, The Agency, 1966.

35. Poncelet, C. G., "Burnup Physics of Heterogeneous Reactor Lattices," WCAP-6069, June 1965.

36. Nodvick, R. J., "Saxton Core II Fuel Performance Evaluation,"

WCAP-3385-56, Part II, "Evaluation of Mass Spectrometric and Radiochemical Analyses of Irradiated Saxton Plutonium Fuel," July 1970.

37. Dean, R. A., "Thermal Contact Conductance Between UO and Zircaloy-2,"

CVNA-127, May 1962.

38. Ross, A. M. and Stoute, R. L., "Heat Transfer Coefficient Between uo and 2

Zircaloy-2," AECL-1552, June 1962.

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39. Tong, L. S., "Boiling Crisis and Critical Heat Flux," AEC Critical Review Series, TID-25887, 1972.

40. Motley, F. E. and Cadek, F. F., "DNB Test Results for New Mixing Vane Grids (R)," WCAP-7695-P-A (Proprietary) and WCAP-7958-A (Nonproprietary),

January 1975.

41. Motley, F. E. and Cadek, F. F., "Application of Modified Spacer Factor to L. Grid Typical and Cold Wall Cell DNB," WCAP-7988-P-A (Proprietary) and WCAP-8030-A (Nonproprietary), January 1975.

42. Tong, L. S., "Prediction of Departure from Nucleate Boiling for an Axially Nonuniform Heat Flux Distribution,"

Journal of Nuclear Energy, 21, 241-248, 1967.

43. Cadek, F. F.; Motley, F. E.; and Dominicis, D. P., "Effect of Axial Spacing on Interchannel Thermal Mixing with the R Mixing Vane Grid,"

WCAP-7941-P-A (Proprietary) and WCAP-7959-A (Nonproprietary), January 1975.

44. Motely, F. E. and Cadek, F. F., "DNB Tests Results for New Mixing Vane Grids (R), "WCAP-7695-P-A, (Proprietary), and WCAP-7958-A, (Nonproprietary), January 1975.

45. Motely, F. E. and Cadek, F. F., "DNB Test Results for R Grid Thimble Cold Wall Cells," WCAP-7695-AI-P-A, Addendum I, (Proprietary), and WCAP-7958-A1-A Addendum I (Nonproprietary), January 1975.

46. Chelemer, H.; Weisman, J.; and Tong, L. S., "Subchannel Thermal Analysis of Rod Bundle Cores," WCAP-7015, Revision 1, January 1969.

4.4-61 SGS-UFSAR Revision 6 February 15, 1987

4 7. Rowe, D. S. and Angle, C. W., "Cross flow Mixing Between Parallel Flow Channels During Boiling, Part II Measurement of Flow and Enthalpy in Two Parallel Channels," BNWL-371, Part 2, December 1967.

48. Rowe, D. S. and Angle, C. W., "Crossflow Mixing Between Parallel Flow Channels During Boiling, Part III Effect of Spacers on Mixing Between Two Channels," BNWL-371, Part 3, January 1969.

49. Gonzalez-Santalo, J. M. and Griffith, P., "Two-Phase Flow Mixing in Rod Bundle Subchannels," ASME Paper 72-WA/NE-19.

50. Motely, F. E.; Wenzel, A. H.; and Cadek, F. F., "The Effect of 17 x 17 Fuel Assembly Geometry on Interchannel Thermal Mixing," WCAP-8298-P-A (Proprietary) and WCAP-8299-A (Nonproprietary), January 1975.

51. Cadek, F. F., "Interchannel Thermal Mixing with Mixing Vane Grids," WCAP-7667-P-A (Proprietary), and WCAP-7755-A (Nonproprietary), January 1975.

52. Hochreiter, L. E. and Chelemer, H., "Application of the THINC IV Program to PWR Design," WCAP-8054 (Proprietary) and WCAP-8195 (Nonproprietary)

September 1973.

53. Dittus, F. W. and Boetler, L. M. K., "Heat Transfer in Automobile Radiators of the Tubular Type," California University Publication in Eng., 2, No. 13, 443-461, 1930.

54. Weisman, J., "Heat Transfer to Water Flowing Parallel to Tube Bundles,"

Nuclear Science and Engineering, 6, 78-79, 1959.

55. Thorn, J. R. S.;, Walker, W. M.; Fallon, T. A.; and Reising, G. F. S.,

"Boiling in Subcooled Water During Flowup Heated Tubes or Annuli,"

Proc. Instn. Mech. Engrs., 180, Pt. C., 226-246, 1965-66.

4.4-62 SGS-UFSAR Revision 6 February 15, 1987

56. Hestroni, G., "Hydraulic Tests of the San Onofre Reactor Model," WCAP-3269-8, June 1964.

57. Hestroni, G., "Studies of the Connecticut-Yankee Hydraulic Model," NY0-3250-2, June 1965.

58. Idel' chik, I. E., "Handbook of Hydraulic Resistance," AEC-TR-6630, 1960.

59. Moody, L. F., "Friction Factors for Pipe Flow,"

Transaction of the American Society of Mechanical Engineers, 66, 671-684, 1944.

60. Mauer, G. W., "A Method of Predicting Steady State Boiling Vapor Fractions in Reactor Coolant Channels," WAPD-BT-19, pp. 59-70, June 1960.

61. Griffith, P.; Clark, J. A.; and Rohsenow, W. M., "Void Volumes in Subcooled Boiling Systems," ASME Paper No. 58-HT-19.

62. Bowring, R. W., "Physical Model, Based on Bubble Detachment, and Calculation of Steam Voidage in the Subcooled Region of a Heated Channel," HPR-10, December 1962.

63. Hochreiter, L. E.; Chelemer, H.; and Chu, P. T., "THINC-IV: An Improved Program for Thermal-Hydraulic Analysis of Rod Bundle Cores," WCAP-7956, June 1973.

64. Carter, F. D., "Inlet Orificing of Open PWR Cores," WCAP-9004, January 1969 (Proprietary) and WCAP-7836 (Nonproprietary), January 1972.

65. Shefcheck, J., "Application of the THINC Program to PWR Design," WCAP-7359-L, August 1969 (Proprietary) and WCAP-7838, January 1972.

4.4-63 SGS-UFSAR Revision 6 February 15, 1987

66. Novendstern, E. H. and Sandberg, R. 0., "Single Phase Local Boiling and Bulk Boiling Pressure Drop Correlations," WCAP-2850, April 1966 (Proprietary) and WCAP-7916 (Nonproprietary), April 1966.

67. Owens, w. L., Jr., "Two-Phase Pressure Gradient," In:

International Developments in Heat Transfer, Part II pp. 363-368, ASME, New York, 1961.

68. McFarlane, A. F., "Power Peaking Factors," WCAP-7912-P-A, (Proprietary) and WCAP-7912-A (Nonproprietary), January 1975.

69. Letter from J. F. Stolz (NRC) to C. Eicheldinger (Westinghouse),

Subject:

Staff Evaluation of WCAP-7956, WCAP-8054, WCAP-8567, and WCAP-8762, April 19, 1978.

70. Vallentine, H. R., "Applied Hydrodynamics," Butterworth Publishers, London, 1959.

71. Kays, W. M. and London, A. L., "Compact Heat Exchangers," National Press, Palo Alto, 1955.

72. Rowe, D. W., "COBRA-III, a Digital Computer Program for Steady State and Transient Thermal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements," BNWL-B-82, 1971.

73. Weissan, J.; Wenzel, A. H.; Tong, L. S.; Fitzsimmons, D.; Thorne, W.; and Batch, J., "Experimental Determination of the Departure from Nucleate Boiling in Large Rod Bundles at High Pressures,"

Chern. Eng. Prog. Symp. Ser. 64, No. 82, 114-125, 1968.

74. Boure, J. A.; Bergles, A. E., and Tong, L. S., "Review of Two Phase Flow Instability," Nuclear Engineering Design, 25 p. 165-192, 1973.

4.4-64 SGS-UFSAR Revision 6 February 15, 1987

75. Lahey, R. T. and Moody, F. J., "The Thermal Hydraulics of a Boiling Water Reactor," American Nuclear Society, 1977.

76. Saha, P.; Ishii, M.; and Zuber, N., "An Experimental Investigation of the Thermally Induced Flow Oscillations in Two-Phase Systems," Journal of Heat Transfer, pp. 616-622, November 1976.

77. Virgil C. Summer FSAR, Docket No. 50-395.

78. Byron/Braidwood FSAR, Docket No. 50-456.

79. South Texas FSAR, Docket No. 50-498.

80. Kakac, S.; Verziroglu, T. N.; Akyuzlu, K.; and Berkol, 0.; "Sustained and Transient Boiling Flow Instabilities in a Cross-Connected Four-Parallel-Channel Upflow System," Pro c. of 5th International Heat Transfer Conference, Tokyo, September 3-7, 1974.

81. Kao, H. S.; Morgan, T. D.; and Parker, W. B., "Prediction of Flow Oscillation in Reactor Core Channel," Trans. ANS, Vol. 16, pp. 212-213, 1973.

82. Stephan, L. A., "The Effects of Cladding Material and Heat Treatment on the Response of Waterlogged UO Fuel Rods to Power Bursts," IN-ITR-111, January 1970.

83. Western New York Nuclear Research Center Correspondence with the AEC on February 11 and August 27, 1971, Docket 50-57.

84. Tong, L. S. et al, "Critical Heat Flux (DNB) in Square and Triangular Array Rod Bundles," presented at the Japan Society of Mechanical Engineers Semi-International Symposium held at Tokyo, Japan, September 4-8, 1967, pages 25-34.

4.4-65 SGS-UFSAR Revision 6 February 15, 1987

85. Ohtsubo, A. and Uruwashi, s. , "Stagnant Fluid due to Local Flow Blockage," Journal of Nuclear Science and Technology 9, No. 7, 433-434, 1972.

86. Basmer, P.; Kirsh, D.; and Schultheiss, G. F., "Investigation of the Flow Pattern in the Recirculation Zone Downstream of Local Coolant Blockages in Pin Bundles," Atomwirtschaft, 17, No. 8, 416-417, 1972. (In German)

87. Burke, T. M.; Meye, C. E.; and Shefcheck, J., "Analysis of Data from the Zion (Unit 1) THINC Verification Test," WCAP-8453-A, May 1976.

88. PSE&G Memorandum, E. A. Liden to S. A. Varga, November 18, 1982.

89. Motley, F. E., Hill, K. W., Cadek, F. F. and Shefcheck, J., "New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids," WCAP-8762 (Proprietary) and WCAP-8763 (non-Proprietary), July 1984.

90. Letter, D. F. Ross, Jr. (NRC) to D. B. Vassala (NRC), "Topical Report Evaluation for WCAP-8762," April 10, 1978.

91. Letter, R. L. Tedesco (NRC) toT. M. Anderson (Westinghouse), "Acceptance for Referencing Topical Report, WCAP-9401 (Proprietary) /WCAP-9402 (Non-Proprietary)," May 1981.

92. Davidson, S. L. (ed.), et al, "Vantage 5H Fuel Assembly," WCAP-10444-P-A, Addendum 2, April 1988 and Letter from W. J. Johnson (Westinghouse) to M.

W. Hodges (NRC), NS-NRC-88-3363, dated July 29, 1988; "Supplemental Information for WCAP-10444-P-A Addendum 2, "Vantage 5H Fuel Assembly."

93. Skaritka, J., (Ed.), "Fuel Rod Bow Evaluation," WCAP-8691 Revision 1, (Proprietary), July 1979.

4.4-66 SGS-UFSAR Revision 11 July 22, 1991

94. "Partial Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1," letter, E. P. Rahe, Jr. (Westinghouse to J. R. Miller (NRC), NS-EPR-2515, dated October 9, 1981; "Remaining Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1," letter, E.

P. Rahe, Jr. (Westinghouse) to J. R. Miller (NRC), NS-EPR-2572, dated March 16, 1982.

95. Letter from C. Berlinger (NRC) to E. P. Rahe, Jr. (Westinghouse),

"Request for Reduction in Fuel Assembly Burn up Limit for Calculation of Maximum Rod Bow Penalty," June 18, 1986.

96. Friedland, A. J. and Ray, S., "Revised Thermal Design Procedure," WCAP-11397-P-A, April 1989.

97. Davidson, S. L. and Kramer, W. R. (Ed.), "Reference Core Report VANTAGE 5 Fuel Assembly," WCAP-10444-P-A, September 1985.

98. Tong, L. S., "Critical Heat Fluxes in Rod Bundles, Two Phase Flow and Heat Transfer in Rod Bundles," Annual Winter Meeting ASME, November 1968, p. 3146.

99. Scherder, W. J. (Ed.) and McHugh, C. J. (Ed.), "Reactor Core Response to Excessive Secondary Steam Releases," WCAP-9226-P-A, Revision 1, 1998.

100. Motley, F. E., Cadek, F. F., "DNB Test Results for R-Grid Thimble Cold Wall Cells," WCAP-7695-L, Addendum 1, October 1972.

101. Davidson, S. L. and Iorii, J. A., "Reference Core Report 17x17 Optimized Fuel Assembly," WCAP-9500-A, May 1982.

102. Letter from E. P. Rahe (Westinghouse) to Miller (NRC) dated March 19, 1982, NS-EPR-2573, WCAP-9500 and WCAPS-9401/9402 NRC SER Mixed Core Compatibility Items.

4.4-66a SGS-UFSAR Revision 25 October 26, 2010

103. Letter from C. 0. Thomas (NRC) to Rahe (Westinghouse), "Supplemental Acceptance No. 2 for Referencing Topical Report WCAP-9500," January 1983.

104. Letter from W. J. Johnson (Westinghouse) to M. W. Hodges (NRC),

VANTAGE 5 DNB Transition Core Effects." NS-NRC-87-32 68, October 2, 1987.a 105. Letter from M. W. Hodges (NRC) to W. J. Johnson (Westinghouse), NRC SER on VANTAGE 5 Transition Core Effects, February 24, 1998.

106. Schueren, P. and McAtee, K. R., "Extension of Methodology for Calculating Transition Core DNBR Penalties," WCAP-11837-P-A, January 1990.

107. Letter from C. Berlinger )NRC) to E. P. Rahe Jr. (Westinghouse),

"Request for reduction in Fuel Assembly Burnup Limit Calculation of Maximum Rod Bow Penalty," June 18, 1986.

108. Weiner, R. A., et al., "Improved Fuel Performance Models for Westinghouse Fuel Rod Design and Safety Evaluations," WCAP-10851-P-A, August 1988.

109. Leech, W. J., et al., "Revised PAD Code Thermal Safety Model," WCAP-8720, Addendum 2, October 1982.

110. Friedland, A. J. and Ray, S., "Improved THINC IV Modeling for PWR Core Design," WCAP-12330-P-A.

111. Davidson, S. L., Nuhfer, D. L. (Eds.), "VANTAGE + Fuel Assembly Reference Core Report," WCAP-12610 and Appendices A through D, June 1990.

4.4-66b SGS-UFSAR Revision 18 April 26, 2000

112. Davidson, S. L. and Ryan, T. L., "VANTAGE+ Fuel Assembly Reference Core Report," WCAP-12610-P-A, April 1995.

113. Motley, F. E., et al., "New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids,"

WCAP-8762-P, July 1984.

114. C. F. Ciocca, "Westinghouse Revised Thermal Design Procedure Instrument Uncertainty Methodology for Salem Units 1 & 2," WCAP-13651 (Proprietary), August 1993.

115. Letter from W. J. Rinkacs (Westinghouse) toM. M. Mannion (PSE&G),

"Westinghouse Generic Safety Evaluation for the 17xl7 Standard Robust Fuel Assembly," October 1, 1998.

116. M. D. Coury, C. R. Tuley, T. P. Williams, "Power Calorimetric for the 1.4% Uprating for Public Service Electric & Gas Company, Salem Units 1 and 2", WCAP-15553 (Proprietary), September, 2000.

117. Letter from Gon Maxwell (Westinghouse) to Doug Tisdel (PSEG), PSE-15-61, "Westinghouse Resolution Plan and Technical Basis for NSAL-14-5, Lower Than Expected Critical Heat Flux Results Obtained During DNB Testing,"

December 3, 2015.

4.4-66c SGS-UFSAR Revision 29 January 30, 2017

THIS PAGE INTENTIONALLY BLANK 4.4-66d SGS-UFSAR Revision 18 April 26, 2000

TABLE 4.4-1 REACTOR THERMAL AND HYDRAULIC DESIGN PARAMETERS Reactor Core Heat OUtput, MWt Reactor Core Heat Output, BTU/hr Heat Generated in Fuel 3459 11,806 97.4 X 10 6

I System Pressure, Nominal psia 2250 System Pressure, Minimum Steady State psia 22U (STDP 111 only) coolant Flow 6

Total Thermal Flow Rate, lb/hr 125.3 X 10 6

Effective Flow Rate for Heat Transfer, lb/hr 116.3 X 10 Effective Flow Area for Heat Transfer, ft 2 V-SH, V+ 51.3 RFA< 2 J Sl.l Average Velocity Along Fuel Rods, ft/sec V-SH, V+ 14.1 RFA 14.2 I

Average Mass Velocity, lb/hr-ft2 V-SH,V+ 2.27 x 10 6 6

RFA 2.28 X 10 Coolant Temperature Nominal Inlet, deg-F 542.7 Average Rise in Vessel, deg-F 70.4 Average Rise in core, deg-F 75.2 1 of 3 SGS-UFSAR Revision 19 November 19, 2001

TABLE 4.4-1 (Cont.)

Average in Core, deg-F 582.4 Average in Vessel, deg-F 577.9 Heat Transfer 2

Active Heat Transfer Surface Area, ft 59,700 2

Average Heat Flux, BTU/hr-ft 192,470 Maximum Heat Flux, For normal operation, BTU/hr-ft 2

461,930 (4 )

Average Thermal Output, kW/ft 5.52 Maximum Thermal Output, For normal operation, kW/ft 13.3 (4 )

Peak Linear Power for determination of protection setpoints, kW/ft < 22.4 (S)

Peak at Thermal Output Maximum for maximum Overpower Trip, deg-F <4700 6

Pressure Drop Across Core, psi Full core V-5H, V+ 22.2 ( )

6 Full core RFA with DFBN 24.7 ( )

9 Full core RFA with SDFBN 24.5 ( )

Minimum DNBR at Normal Conditions Typical Flow Channel V-5H,V+ 2.44 RFA 2.64 Thimble (Cold Wall) Flow Channel V-5H,V+ 2.32 RFA 2.62 2 of 3 SGS-UFSAR Revision 28 May 22, 2015

TABLE 4.4-1 (Cont.)

DNBR Correlation (7 ) V-5H,V+ WRB-1 RFA WRB-2 DNBR Correlation Limit( 7 ) WRB-1 1.17 WRB-2 1.17 DNBR Design Limit (B) WRB-1 1.24(RTDP(J) ,Typical)

WRB-1 1.24(RTDP, Thimble)

WRB-2 1.24(RTDP, Typical)

WRB-2 1.22(RTDP, Thimble)

Notes:

l)Standard Thermal Design Procedure 2)All Parameters for RFA include Intermediate Flow Mixing (IFM) grids 3)Revised Thermal Design Procedure 4)Associated with FQ limit of 2.40 5)See Section 4.3.2.2.6 6)Based on a best estimate reactor flow rate of 93,300 gpm/loop 7)See Section 4.4.2.3.1 8)See Section 4.4.1.1 9)Based on a best estimate reactor flow rate of 94,800 gpm/loop and a Tavg of 566.0 F.

3 of 3 SGS-UFSAR Revision 28 May 22, 2015

TABLE 4.4-2 THERMAL-HYDRAULIC DESIGN PARAMETERS FOR ONE OF FOUR COOLANT LOOPS OUT OF SERVICE (This Table has been deleted)

SGS-UFSAR Revision 18 April 26, 2000

• VOID FRACTIONS AT NOMINAL REACTOR CONDITIONS WITH DESIGN HOT CHANNEL FACTORS Average Maximum Core 0.24%

Hot Subchannel 5.8% 20.H 1 of 1 SGS-UFSAR Revision 19 November 19, 2001

TABLE 4.4-4 COMPARISON OF THINC-IV AND THINC-1 PREDICTIONS WITH DATA FROM REPRESENTATIVE WESTINGHOUSE TWO AND THREE LOOP REACTORS Improvement (°F)

Measured Inlet a rms (oF) a (°F) for THINC-IV Power %Full Reactor (MWt) Power Temp (°F) THINC-I THINC-IV over THINC-1 Ginna 847 65.1 543.7 1.97 1.83 0.14 854 65.7 544.9 1.56 1.46 0.10 857 65.9 543.9 1.97 1.82 0.15 947 72.9 543.8 1.92 1. 74 0.18 961 74.0 543.7 1.97 1. 79 0.18 1091 83.9 542.5 1. 73 1.54 0.19 1268 97.5 542.0 2.35 2.11 0.24 1284 98.8 240.2 2.69 2.47 0.22 1284 98.9 541.0 2.42 2.17 0.25 1287 99.0 544.4 2.26 1.97 0.29 1294 99.5 540.8 2.20 1.91 0.29 1295 99.6 542.0 2.10 1.83 0.27 Robinson 1427.0 65.1 548.0 1.85 1.88 0.03 1422.6 64.9 549.4 1.39 1.39 0.00 1529.0 88.0 550.0 2.35 2.34 0.01 2207.3 100.7 534.0 2.41 2.41 o.oo 2213.9 101.0 533.8 2.52 2.44 0.08 1 of 1 SGS-UFSAR Revision 6 February 15, 1987

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1.017 1.111 1.002 0.994 1.147 1.064 1.181 0.991 0.998 0.988 I

1.163 1.206 1.160 1.187 0.990 0.986 0.990 0.998 I

1.225 1.170 1.188 1.065 1.246 0.985 0.989 0.988 0.998 0.981 I

I 1.128 1.165 1.096 1.088 0.921 0.971 0.992 0.990 0.995 0.996 1.009 1.006 I

I 1.027 1.028 0.993 0.978 0.825 0.471 1.000 1.000 1.003 1.004 1.018 1.018 I

I 0.718 0.782 0.666 0.565 FOR RADIALPOWER DISTRIBUTION 1.022 1.019 1.021 1.019 NEAR BEGINNINGOF LIFE,HOT FULL POWER,EOUILIBRIUM* XENON I

N J

CALCULATEDFD.H = 1.35 Normalized Radial Flow and Enthalpy PUBLIC SERVICE ELECTRIC AND GAS COMPANY Distribution at 8*Ft Elevation SALEM NUCLEAR GENERATING STATION Updated FSAR FIG. 4.4-8 REVISION 6 FEBRUARY 15, 1987

• I I

KEY:

-I-1.087 -I- .1h/.1h

• I- I- 1- 1-0.995 GIG I

1.017 1.112 0.999 0.993 I

I 1.149 1.065 1.182 0.991 0.996 0.989 1.164 1.209 1.161 1.188 0.990 0.980 0.991 0.990 1.228 1.172 1.190 1.066 1.247 0.987 0.990 0.989 0.997 0.987 I

1.129 1.165 1.096 1.089 0.919 0.969 0.993 0.991 0.995 0.996 1.005 1.003 I

I 1.026 1.027 0.991 0.976 0.823 0.472 1.000 1.000 1.002 1.003 1.011 1.030 I

FOR RADIALPOWER DISTRIBUTION 0.714 0.779 0.663 0.563 NEAR BEGINNING OF LIFE,HOT FULL 1.016 1.013 1.019 1.025 POWER,EQUILIBRIUM XENON N

I CALCULATED FaH = 1.35 Normalized Radial Flow and Enthalpy PUBLIC'SERVICE ELECTRIC AND GAS COMPANY Distribution at 12-Ft Elevation (Unit 2)

SALEM NUCLEAR GENERATING STATION Updated FSAR FIG. 4.4-9 REVISION 6 FEBRUARY15,1987

• X 0

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FEBRUARY15.1987

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03UnS'V'3W-3UnJ.VU3dW3J. J.IX3A18W3SS'V' 03.101031-Jd Comparison of a Representative W Two-Loop Reactor PUBLIC SERVICE ELECTRIC AND GAS COMPANY lncore Thermocouple Measurements with SALEM NUCLEAR GENERATING STATION THINC- nz Predictions Updated FSAR FIG. 4.4-12 REVISION 6 FEBRUARY15,1987

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[!]0 [!] 0 15

[Q) THERMOCOUPLELOCATION******* __ **__ **58

(!] INCORE DETECTOR LOCATION__ **_******* 58 (Q] FLOW MIXINGDEVICE LOCATION*..*.....52 THERE ARE 58 LOCATIONSDESIGNATEDAS THERMOCOUPLEAND INCORE DETECTOR LOCATIONS.THE MINUMUM NUMBER.OF FLUX THIMBLESAND THERMOCOUPLESIN USE, AS WELL AS THE MIIMUM NUMBER PER QUADRANT, IS CONTROLLED*ay THE TECHNlCALSPECIFICATIONS.

Solem NuclearGeneratin_g_Station PSEG Nuclear,LLC DISTRIBUTIONOF INCORE INSTRUMENTATION UNIT 1 SALEM NUCLEAR GENERATINGSTATION UpdatedFSAR Figure4.4-16

RPN M LK J HG FE 0 CBA I

0 [!] 0 [!]

0 0 [!] ~ 0 0 2 j'

0 0 0 [!] 000 3

' [!] [!] [!] 0 0 4 00 0 [!] 0 0 [!] 0 [!] [!] 0 5

[!] [!] [!] [!] 0 [!] 6 0 [!] 0 [!] 0 [!] 7 e [!]0 0 [!] 0 [!] 0 [!] 0 [!] 0 e 0 0 00 [!] 0 0 e

0 [!] 0 [!] 10 0 [!] 0 0 0 [!] 0 11 0 [!] [!] [!] 0 12 00 [!] 0 e 0 [!] 13

(!] 0 [!] [!] iO [!] 14

[!] 0 [!] 0 15 (Q] THERMOCOUPLELOCATION***************

58

(!] INCORE DETECTOR LOCATION*******.****

58 (Q]FLOW MIXINGDEVICE LOCATION********.

58 THERE ARE 58 LOCATIONSDESIGNATEDAS THERMOCOUPLEAND INCORE DETECTOR LOCATIONS.THE MINIMUM NUMBER,OF FLUX.THIMBLESAND THERMOCOUPLESIN USE, AS WELL AS THE MINIMUMNUMBER PER QUADRANT IS CONtROLLEDBY THE TECHNICALSPECIFICATIONS.

ReV1s10n Dec.6, 2~04 Salem NuclearGenerotil)gStation PSEG Nuclear,LLC DISTRIBUTIONOF INCORE INSTRUMENT ATION UNIT 2 SALEM NUCLEARGENERATINGSTATION UpdatedFSAR Figure4.4-17

4.5 RELOAD ANALYSES A Reload Evaluation (RSE)/Safety Assessment (SA) is performed for each cycle using methodology described in Reference 1. Based on this methodology, those events analyzed and reported in the Salem UFSAR, as well as limits given in the Technical Specification or Core Operating Limits Report (COLR) that could potentially be affected by the fuel reload are addressed. These UFSAR analyses and limits contain assumptions which involve parameters whose values are core design dependent. Hence, those parameters sensitive to reload core designs are considered, i.e., core criticality, power distributions, shutdown margin, etc. In addition, changes in fuel assembly design (mechanical, nuclear, and thermal-hydraulic) that could potentially affect the events analyzed are also addressed. The RSE/SA results are used as input into the 10CFR50. 59 process to determine if an Unresolved Safety Question (USQ) exists or a Technical Specification needs to be modified for a reload cycle.

As part of the Reload Safety Evaluation process, the values for the parameters defined in the COLR are determined. The COLR contains specific parameter values which were contained in the Technical Specifications:

Beginning and End of Cycle Moderator Temperature Moderator Temperature Surveillance Limit, Control Rod Insertion Limits, Axial Flux Difference Range, Heat Flux Hot Channel Factor (FQ{z)), Nuclear Enthalpy Rise Hot Channel Factor (FAH), and Boron Concentration .

The Nuclear Design Report (NOR), Curvebook (CB), and Plant Operations Package (POP) are cycle-specific documents which contain the best estimate of the reload's nuclear characteristics. Typical nuclear characteristics consist of boron concentrations, reactivity coefficients, boron and control rod worths. Typical fuel assembly design information consists of assembly enrichments, fuel pellet and rod characteristics, and burnable absorber types.

NOR, CB, and POP data is used to support plant operation and to compare measured and predicted plant data.

Typical Salem Unit 1 cycle 1 s loading pattern and burnable absorber configuration are given as 4.5-1 and 4.5-2, respectively. Typical Salem Unit 2 cycle s loading pattern and burnable absorber configuration are 1

given as Figures 4.5-3 and 4.5-4, respectively.

The cycle specific RSE/SA, COLR, and NDR/CB/POP for Salem Units 1 and 2 can be found in the appropriate site specific document control system .

• SGS-UFSAR 4.5-1 Revision 23 October 17, 2007

4.5.1 References for Section 4.5

1. Davidson, S. L. 1 (Ed.), et. al. 1 "Westinghouse Reload Safety Evaluation Methodology," WCAP-9273-NP-A, July 1985.

4.5-2 SGS-UFSAR Revision 17 October l61 1998

TABLE 4.5-1 (This Table has been deleted}

1 of 1 SGS-UFSAR Revision 23 October 17, 2007

TABLE 4.5-2 (This Table has been deleted) 1 of 1 SGS-UFSAR Revision 23 October 17, 2007

r-------------------------------------------

1 I

I R p N M L K J H G F E D c B A I

I I

I I Af'Jq AF20 AF54 AF72 32 AF52 AF18 I L-q L-10 L-15 D-6 -11 E-10 D-8 l I AF03Af't;qAH44 AH60 AH63 AG70 AH65 AH7l AH47 AFS4 AF08 I N-ll H-3 FEED FEED FEED H-14 FEED FEED FEED M-12 C-11 2 I AF01 AF67 AH4q AG36 AH04 AG27 AG2<i' AG21 AG16 AG42 AH!5!5 AF71 AF07 I E-3 M-6 FEED M-3 FEED P-1 J-14 B-11 FEED D-3 FEED F-4 L-3 3 I AF67 AH5S AG56 Atflq AGsq AH2<1' AG48 AH30 AG68 AH08 AG60 AH30 AF55 I D-12 FEED F-2 FEED N-11 FEED F-14 FEED C-11 FEED B-11 FEED C-8 4 I AF12 AH57 AG43 AH38 AGlMAHtiJq AG12 AH24 AGfR AH25 AGil AH21 AG31 AH45 AF21 I H~4 FEED N-4 FEED H-7 FEED K~q FEED F-q FEED G-8 FEED C-4 FEED J-15 5 I AF50 AH72 AH22 AGS6 AH15 AGll.lAG64 AG41 AG52 AG88 AH18 AG65 AHIJ2 AH5q AF51 I F-5 FEED FEED F-3 FEED M-5 r+q G-14 o-q E-4 FEED K-3 FEED FEED K-5 6 I f:Fl7 AH73 AG24 AH28 AG82 AG71 AH14 AG18 AHil AG46 AG17 AH35 AG22 AH61 AF26 I 7 E-8 FEED E-2 FEED G-6 G-4 FEED E-12 FEED J-4 J-6 FEED L-2 FEED E-5 I

I qeo AF65 AG45 AtM0 AG57 AH33 AG32 AG16 AH01 AGI6 AG3<1' AH27 AG51 AG44 AG55 Af&q 8 K-4 B-8 e-q B-6 FEED B-7 P-5 FEEC M-11 P-q FEED P-11 P-7 P-8 F-12 I

AF47 AH68 AF23 AH41 AF1!5 AG62 AH26 AG03 AH23 AH32 AG28 AHsq AF3<1' I

I L-U FEED E-14 FEED G-10 G-12 FEED L-4 FEED ~~ FEED L-14 FEED L-8 AF66 AH66 AH10 AG67 AH37 AGJq AG68 AG3l AG63 AG05 AH08 AG5q AH17 AH67 AF41 q

I 10 I

F-11 FEED FEED F-13 FEED L-12 M-7 J-2 D-7 D-11 FEED K-13 FEED FEED K-11 I AE33 AH!52 AG37 AH31 AG14 AH20 AF20 AH34 AG13 AH36 AG07 AH40 AG38 AH!53 AF27 l1 I G-ll FEED N-12 FEED J-8 FEED K-7 FEED F-7 FEED H-q FEED C-12 FEED H-12 I AF62 AH50 AG4q AH06 AF60 AG3CJ AG47 AH06 AG61 AH12 AG50 AH51 AF61 12 I N-8 FEED P-6 FEED N-6 FEED K-2 FEED C-6 FEED K-14 FEED M-4 I AF06 AF70 AH46 AG33 AH07 AG2!5AG34 AG26 AH13 AG3!5 AH48 AF66 AF05 13 I E-13 K-12 FEED M-13 FEED P-5 G-2 B-15 FEED D-13 FEED P-10 L-13 I AF04 AF60 AG54 AH68 AH70 AH54 AF68 AFI2 I

I

-~~~

E-5 D-4 FEED FEED H-2 FEED FEED FEED H-13 C-5 AF31 AF28 AF25 AFSB AF24 AF11 AF10 14 15 I M-8 L-6 H-5 M-10 E-ll E-6 E-7 I

rv:wlASSEMBLY IDENTIFIER 1 ~ PREVIOUS CYCLE LOCATION I

I rAFl REGlON 18A L_J C4.3q8w; o>

fAGl REGION L_j (4.S01W/o )

tqs fAHlREGION L_J (4.S00W/o )

20C I

I I

rAFlREGION 188 L_j(4.5q7W/o) fAHlREGION L_J (3.600 W/o )

20A I

I fAGlREGION lqA fAHlREGION 2BB 1 I L_j(4.402W/ o> L_j(4.200W/o ) I I I

~-------------------------------------------

ReVISion23, October 17,2fi/J7 Solem NuclearGeneratingStation PSEG Nuclear* LLC TYPICALSALEM UNIJ 1LOADING PATTERN SALEM NUCLEAR GENERATINGSTATION UpdatedFSAR Figure4.5-1

© 2000PSEG NtK:Iecr,LLC. All Ri~ts Reserved.

R PNMLK J HG FE D C B A RCCA 641 1041 1041 RCCA 1041 RCCA 641 2 10-41 12W 12W RCCA &41 RCCA 1281 RCCA 6SSA RCCA 1281 RCCA 641 RCCA 3 64( RCCA 12W 12W RCCA 12W 12W 6.4(

1281 1281 1281 1281 RCCA 12W 641 RCCA 1281 12W 12W 12W RCCA 6.4[

1281 RCCA 128[ 1281 5 RCCA 12W 12W 12W 12W RCCA 104[ 1281 1281 RCCA RCCA RCCA 1281 1281 1041 6 12W RCCA 1041 RCCA 1281 12W 12W 12W RCCA 1041 7 1281 1281 1281 RCCA RCCA 12W RCCA RCCA RCCA 12W RCCA RCCA 8 1281 1281 1281 1041 RCCA 12W 12W 12W RCCA 12W RCCA 1041 1281 1281 1281 1281 RCCA 12W t2W RCCA RCCA 12W 12W RCCA 1041 1281 1281 RCCA 1281 1281 1041 10 12W 12W 12W 12W 12W 641 RCCA 1281 11 1281 RCCA 1281 1281 1281 RCCA 641 641 RCCA 128112W t2W 12W 12W 1281 RCCA 1281 1281 RCCA 6.41 12 RCCA &41 RCCA 12W RCCA 6SSA RCCA 12W RCCA 6.41 RCCA 13 1281 1281 641 RCCA RCCA 1041 1041 RCCA 1041 1041 6.41 14 15 LEGEND FUEL ASSEMBLY ORIENT AliON

~ COMPONENT TYPE

• REFERENCE HOLE

~ NUMBER OF FRESH IFBA RODS o CORE PIN HOLE

/ HOL DOWN BAR CORE COMPONENT TYPES NOTE:FIGURES ARE TOP VIEW RCCA-CONTROL OR SHUTDOWN RCCA

• • w *NUMBER OF RODLETS ON WABA ASSEMBLY

• SSA-NUMBER OF RODLETS ON SECONDARY SOURCE ASSEMBLY COPTIONAL COMPONENT)

---

~

Rev1s1on25, October26, 2010:

Solem NuclearGeneratingStation PSEG Nuclear,LLC TYPICALSALEM UNIT 1BURNABLE RR Fl RATION SALEM NUCLEAR GENERATINGSTATION UpdatedFSAR Figure4.5-2 2000PSEC Nucle<r.llC. All RightsReserved. r

r-------------------------------------------

RPNML K J HG FED C 8 A 168 168 168 17A 168 168 168 1 NS43 NS4q NS65 NTIS NS68 NS37 NS41 168 168 188 188 188 188 188 188 188 168 169 2 NS158 NS31 NU34 NU43 NUl50 INU!52 NU61 NU62 NU36 NS33 NS!57 168 17A 188 178 18B 17A 17A 17A 188 178 188 158 168 3 NS!54 NT27 NU42 NT68 NU70 NT33 NT20 NUJq NU72 NT63 NU67 NT26 NS62 168 188 178 18A 17A leA 178 168 17A 178 178 188 168 4 NS32 NU45 NU75 NU03 NT25 NU08 NT46 NUll NT36 NU16 NT70 NU56 NS27 168 188 178 18A 17A 18A 178 18A 178 ISA 17A 18A 178 188 168 NS46 NU37 NT66 NU17 NT16 NU22 NT 53 NUfJ2 NT45 NU24 NT01 NU2fJ NT5q NU38 NS45 168 188 188 17A leA 17A 178 17A 178 17A 18A 17A 188 188 168 6 NS48 NU57 NU73 fNT2q NU25 NT1!5 NT65 NT07 NT 57 NTlC) NR30 NT28 NS74 NSl53 NS47 168 188 17A 18A 178 178 178 18A 178 178 178 18A 17A 188 168 7 NS70 NU!54 NT02 NU04 NT50 NT73 NT52 NU31 NT47 NT72 ~T4q NU13 NT18 NU4C) NS5C) 17A 188 17A 178 leA 17A leA 148 18A 17A 18A 178 17A 18B 17A 8 NT34 NUt51 NT22 NT51 NU14 NT12 NU21 NP78 NU32 NT03 NUll NT44 NT13 NU41 NT40 168 188 17A 18A 178 178 178 18A 178 178 178 18A 17A 188 168 NS64 NU60 NT11 NU12 NT42 NT62 NT48 NU23 NT56 NT71 NT55 NU07 NT04 NU48 NS66 168 188 188 17A ISA 17A 178 17A 178 17A 18A 17A 18B 188 168 NS40 NU47 NU75 NT05 NU28 NT24 NTt58 NT14 NT76 NT06 ~U2q NT38 NUBCJ NU65 NSt52 1GB 16B 178 18A 17A 18A 178 18A 178 18A 17A 18A 178 188 16B 11 NS38 NUJCJ NT74 NU0q NTfiCJ NU26 NT !54 NUll NT43 NU27 NT21 NU15 NT61 NU33 NS51 16B 188 178 18A 17A 188 178 168 15A 168 1158 168 168 12 NS35 NS64 NT67 NU06 NT17 NUI!S NT41 NlJlq NT10 NU18 NT64 NU5q NS34 168 17A 188 178 188 17A 17A 17A 188 178 188 17A 168 13 NP76 NT32 NU68 NT60 NU76 NT37 NT30 NT23 NU71 NTSq NU58 NT31 NS74 168 169 188 188 188 188 188 188 188 169 1GB 14 NS75 NS2CJ NU4l!J NU!55 NU63 NU4G NU66 NU44 NU35 NS26 NS56 16B 168 168 17A 1GB 148 148 NS50 NS42 NS60 NT35 NS56 NSJq NS44 FUEL ASSEMBLY ORIENTATION LEGEND

• REFERENCE HOLE o CORE PIN HOLE fRl REGION IDENTIFIER / HOLDOWN BAR

~ FUEL ASSEMBLY IDENTIFIER NOTE1FIGURES ARE TOP VIEW

~-------------------------------------------

Revu1Jon23, October 17,2'l!IJ7 Salem Nuclear Generating Station PSEG Nuclear,LLC TYPICALSALEM UNIT 2 LOADING PATTERN SALEM NUCLEAR GENERATINGSTATION UpdatedFSAR Figure4.5~3 CO 2000 PS£G Nlx:la<r, llC. All Ri!llts Reserved.

I I

I I R p N M L K J H 0 F E 0 c B I

I I

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I 801 RCCA 8W 8W 8W 8W 1041 12W 1041 RCCA 1041 12W 1041 RCCA 801 RCCA I

4W 4W I 321 RCCA 8W 8W 8W 321 1041 1281 1041 1281 8W 1041 RCCA I

RCCA 8W RCCA 4W RCCA RCCA 4W RCCA I

I 801 1281 1281 RCCA 128[ 1~:1 ~fA I 8W 801 RCCA 1041 1281 8W 801 7 1041 RCCA I

I RCCA 8w RCCA 1281 RCCA 1281 RCCA 8w RCCA RCCA q0o 8W 801 RCCA 1041 1041 801 8w 8 I

I 801 RCCA 8W 8W 10.41 1281 104[ RCCA 801 I

RCCA 8W 12W 4W RCCA RCCA 4W 12W 8W RCCA I 801 1281 1281 RCCA 1281 1281 801 10 I

8W 8W I 321 RCCA 10.41 8W .4W 1281 4W aw 8W RCCA 321 1041 11 1281 1281 I

aw 8W RCCA aw RCCA aw I RCCA 801 RCCA 1041 RCCA 1041 1041 1041 RCCA 801 12 I aw RCCA SSSA RCCA aw RCCA I 12W 801 RCCA 1281 1281 801 12W 13 I

I 321 RCCA 801 RCCA 801 RCCA 321 RCCA 1.4 801 801 801 I

I 8W 15 I

I I

I

LEGEND FUEL ASSEMBLY ORIENTATION IITYPEICOMPONENT TYPE

• REFERENCE HOLE o CORE PIN HOLE I 0001 NUMBER OF FRESH IFBA RODS

/ HOLDOWN BAR

coRE COMPONENT TYPES NOTE:F[GURESARE TOP V[EW JRCCA-CONTROLOR SHUTDOWN RCCA 1

• • w -NUMBER OF RODLETS ON WABA ASSEMBLY 1 .::_ssA

_:-NUMBER_ OF !ODLETS_ON _:ECONDARY_ SOURCE_ASSEMBLY~OPT_!_ONAL _:oMPONENT~ __ _

Rev1s1on25. October26, 2010 PSEG Nuclear,L L C SALEM NUCLEAR GENERATINGSTATION UpdatedFSAR Figure4.5*4 2000PSEG t<<Jcle<r.LLC. All RightsReserved.

'