ML17151A999

Revision 30 to Updated Final Safety Analysis Report, Chapter 4.0, Reactor

ML17151A999
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Issue date:	03/09/2017
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{{#Wiki_filter:WOLF CREEK TABLE OF CONTENTS CHAPTER 4.0 REACTOR Section Page

4.1

SUMMARY

DESCRIPTION 4.1-1

4.

1.1 REFERENCES

4.1-4

4.2 FUEL SYSTEM DESIGN 4.2-1

4.2.1 DESIGN BASES 4.2-2

4.2.1.1 Cladding 4.2-2 4.2.1.2 Fuel Material 4.2-3 4.2.1.3 Fuel Rod Performance 4.2-4 4.2.1.4 Spacer Grids 4.2-4 4.2.1.5 Fuel Assembly 4.2-5 4.2.1.6 Incore Control Components 4.2-7 4.2.1.7 Surveillance Program 4.2-9

4.2.2 DESIGN DESCRIPTION 4.2-9

4.2.2.1 Fuel Rods 4.2-11 4.2.2.2 Fuel Assembly Structure 4.2-13 4.2.2.3 Incore Control Components 4.2-20

4.2.3 DESIGN EVALUATION 4.2-24

4.2.3.1 Cladding 4.2-25 4.2.3.2 Fuel Materials Considerations 4.2-28 4.2.3.3 Fuel Rod Performance 4.2-28 4.2.3.4 Spacer Grids 4.2-36 4.2.3.5 Fuel Assembly 4.2-36 4.2.3.6 Reactivity Control Assembly and Burnable Absorber Rods 4.2-37

4.2.4 TESTING AND INSPECTION PLAN 4.2-40

4.2.4.1 Quality Assurance Program 4.2-40 4.2.4.2 Quality Control 4.2-40 4.2.4.3 Incore Control Component Testing and Inspection 4.2-43 4.2.4.4 Tests and Inspections by Others 4.2-44 4.2.4.5 Inservice Surveillance 4.2-45 4.2.4.6 Onsite Inspection 4.2-45

4.0-i Rev. 29 WOLF CREEK TABLE OF CONTENTS (Continued)

Section Page

4.

2.5 REFERENCES

4.2-46

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) 4.3-2

4.3.1.3 Control of Power Distribution 4.3-3

4.3.1.4 Maximum Controlled Reactivity Insertion

Rate 4.3-4

4.3.1.5 Shutdown Margins 4.3-5

4.3.1.6 Stability 4.3-6

4.3.1.7 Anticipated Transients Without SCRAM 4.3-6

4.

3.2 DESCRIPTION

4.3-7

4.3.2.1 Nuclear Design Description 4.3-7

4.3.2.2 Power Distributions 4.3-8

4.3.2.3 Reactivity Coefficients 4.3-19

4.3.2.4 Control Requirements 4.3-22

4.3.2.5 Control Rod Patterns and Reactivity

Worth 4.3-27

4.3.2.6 Criticality of the Reactor During Refuel-

ing and Criticality of Fuel Assemblies 4.3-29

4.3.2.7 Stability 4.3-29

4.3.2.8 Vessel Irradiation 4.3-33

4.3.3 ANALYTICAL METHODS 4.3-34

4.3.3.1 Fuel Temperature (Doppler) Calculations 4.3-34

4.3.3.2 Macroscopic Group Constants 4.3-36

4.3.3.3 Spatial Few-Group Diffusion Calculations 4.3-38

4.

3.4 REFERENCES

4.3-39

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 4.4-1

4.4.1.2 Fuel Temperature Design Basis 4.4-2

4.0-ii Rev. 29 WOLF CREEK TABLE OF CONTENTS (Continued)

Section Page

4.4.1.3 Core Flow Design Basis 4.4-3

4.4.1.4 Hydrodynamic Stability Design Basis 4.4-3

4.4.1.5 Other Considerations 4.4-3

4.

4.2 DESCRIPTION

4.4-4

4.4.2.1 Summary Comparison 4.4-4

4.4.2.2 Critical Heat Flux Ratio or Departure from

Nucleate Boiling Ratio and Mixing Technology 4.4-4

4.4.2.3 Linear Heat Generation Rate 4.4-11

4.4.2.4 Two Phase Flow Correlations and Void

Correlations 4.4-11

4.4.2.5 Core Coolant Flow Distribution 4.4-12

4.4.2.6 Core Pressure Drops and Hydraulic Loads 4.4-13

4.4.2.7 Correlation and Physical Data 4.4-14

4.4.2.8 Thermal Effects of Operational Transients 4.4-16

4.4.2.9 Uncertainties in Estimates 4.4-16

4.4.2.10 Flux Tilt Considerations 4.4-19

4.4.2.11 Fuel and Cladding Temperatures 4.4-19

4.4.2.12 Revised Thermal Design Procedure (RTDP) 4.4-22

4.

4.3 DESCRIPTION

OF THE THERMAL AND HYDRAULIC

DESIGN OF THE REACTOR COOLANT SYSTEM 4.4-23

4.4.3.1 Plant Configuration Data 4.4-23

4.4.3.2 Operating Restrictions on Pumps 4.4-23

4.4.3.3 Power-Flow Operating Map (BWR) 4.4-24

4.4.3 4 Temperature-Power Operating Map 4.4-24

4.4.3.5 Load Following Characteristics 4.4-24

4.4.3.6 Thermal and Hydraulic Characteristics

Summary Table 4.4-24

4.4.4 EVALUATION 4.4-24

4.4.4.1 Critical Heat Flux 4.4-24

4.4.4.2 Core Hydraulics 4.4-24

4.4.4.3 Influence of Power Distribution 4.4-26

4.4.4.4 Core Thermal Response 4.4-28

4.4.4.5 Analytical Techniques 4.4-28

4.4.4.6 Hydrodynamic and Flow Power Coupled Insta-

bility 4.4-30

4.0-iii Rev. 29

WOLF CREEK TABLE OF CONTENTS (Continued)

Section Page

4.4.5 TESTING AND VERIFICATION 4.4-32

4.4.5.1 Tests Prior to Initial Criticality 4.4-32

4.4.5.2 Initial Power and Plant Operation 4.4-32

4.4.5.3 Component and Fuel Inspections 4.4-33

4.4.6 INSTRUMENTATION REQUIREMENTS 4.4-33

4.4.6.1 Incore Instrumentation 4.4-33

4.4.6.2 Overtemperature and Overpower DT Instru-

mentation 4.4-33

4.4.6.3 Instrumentation to Limit Maximum Power

Output 4.4-33

4.4.6.4 Loose Parts Monitoring System 4.4-34

4.

4.7 REFERENCES

4.4-36

4.5 REACTOR MATERIALS 4.5-1

4.5.1 CONTROL ROD SYSTEM STRUCTURAL MATERIALS 4.5-1

4.5.1.1 Materials Specifications 4.5-1

4.5.1.2 Fabrication and Processing of Austenitic

Stainless Steel Components 4.5-2

4.5.1.3 Contamination Protection and Cleaning

of Austenitic Stainless Steel 4.5-2

4.5.2 REACTOR INTERNALS MATERIALS 4.5-2

4.5.2.1 Materials Specifications 4.5-2

4.5.2.2 Controls on Welding 4.5-2

4.5.2.3 Nondestructive Examination of Wrought

Seamless Tubular Products and Fittings 4.5-3

4.5.2.4 Fabrication and Processing of Austenitic

Stainless Steel Components 4.5-3

4.5.2.5 Contamination Protection and Cleaning of

Austenitic Stainless Steel 4.5-3

4.6 FUNCTIONAL DESIGN OF REACTIVITY CONTROL SYSTEMS 4.6-1

4.6.1 INFORMATION FOR CONTROL ROD DRIVE SYSTEM

(CRDS) 4.6-1

4.6.2 EVALUATION OF THE CRDS 4.6-1

4.6.3 TESTING AND VERIFICATION OF THE CRDS 4.6-1

4.0-iv Rev. 29 WOLF CREEK TABLE OF CONTENTS (Continued)

Section Page

4.6.4 INFORMATION FOR COMBINED PERFORMANCE OF

REACTIVITY SYSTEMS 4.6-2

4.6.5 EVALUATION OF COMBINED PERFORMANCE 4.6-2

4.

6.6 REFERENCES

4.6-3

4.0-v Rev. 0 WOLF CREEK TABLE OF CONTENTS (Continued)

LIST OF TABLES

Table No. Title

4.1-1 Reactor Design Table

4.1-2 Analytical Techniques In Core Design

4.1-3 Design Loading Conditions for Reactor Core Components

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 Benchmark Critical Experiments

4.3-5 Axial Stability Index Pressurized Water Reactor Core

with a Twelve Foot Height

4.3-6 ARO Moderator Temperature Coefficients versus

Average Moderator Temperature and Burnup

4.3-7 Comparison of Measured and Calculated Doppler Effects

4.3-8 Saxton Core II Isotopics, Rod MY, Axial zone 6

4.3-9 Critical Boron Concentrations, HZP, BOL

4.3-10 Comparison of Measured and Calculated Rod Worth

4.3-11 Comparison of Measured and Calculated Moderator

Coefficients at HZP, BOL

4.0-vi Rev. 29 WOLF CREEK

TABLE OF CONTENTS (Continued)

LIST OF TABLES

Table No. Title

4.4-1 Thermal and Hydraulic Comparison Table

4.4-2 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

4.4-5 Loose Parts Monitoring System

4.0-vii Rev. 11 WOLF CREEK CHAPTER 4 - LIST OF FIGURES

*Refer to Section 1.6 and Table 1.6-3. Controlled drawings were removed from the USAR at Revision 17 and are considered incorporated by reference.

Figure # Sheet Title Drawing #* 4.2-1 Fuel Assembly Cross Section 17 x 17 LOPAR 4.2-1a Fuel Assembly Cross Section 17 x 17 VANTAGE 5H 4.2-2 Typical Fuel Assembly Outline 17 x 17 4.2-2a Fuel Assembly Outline 17 x 17 Vantage 5H with IFM Grids 4.2-2b Fuel Assembly Outline 17 x 17 Vantage 5H with IFMs & PBG 4.2-2c Fuel Assembly Outline 17 x 17 Vantage 5H with Performance+ features (V5HP+) 4.2-2d Fuel Assembly Outline 17 x 17 Vantage 5H with Performance+ features, Zirlo2 (V5HP+Z+2), RFA Z+2 and RFA-2 Z +2 4.2-2e Fuel Assembly Outline 17x17 Standard, Performance & Features, Zirlo +2 , RFA, RFA-2, Win 4.2-3 Fuel Rod Schematic Standard Rod 4.2-3a Fuel Rod Schematic High Burnup Rod 4.2-3b Fuel Rod Schematic Performance+ Zirc-4 4.2-3c Fuel Rod Schematic Performance+, Zirlo 4.2-3d Fuel Rod Schematic Performance+, Zirlo2 4.2-4 Mid Grid Expansion Joint - Plan View 4.2-5 Mid Grid Expansion Joint - Elevation View 4.2-6 Top Grid to Nozzle Attachment, Standard 4.2-6a Thimble/Insert/Top Grid Sleeve Bulge Joint Geometry 4.2-7 Guide Thimble to Bottom Nozzle Joint 4.2-8 Rod Cluster Control and Drive Rod Assembly With Interfacing Components 4.2-9 Rod Cluster Control Assembly Outline 4.2-10 Absorber Rod 4.2-11 Wet Annular Burnable Absorber Assembly 4.2-11a Typical Borosilicate Glass Burnable Absorber Rod Assembly 4.2-12 Wet Annular Burnable Absorber Rod Assembly 4.2-12a Borosilicate Glass Burnable Absorber Rod 4.2-13 Secondary Source Rod Assembly 4.2-13a Double Encapsulated Secondary Source Rod Assembly 4.2-14 Secondary Source Assembly 4.2-14a Typical Primary Source Assembly 4.2-14b Double Encapsulated Secondary Source Assembly 4.2-15 Typical Double Spring Thimble Plug Device 4.2-15a Typical Single Spring Thimble Plug Device 4.3-1 Typical Core Loading Pattern 4.3-2 Production and Consumption of Higher Isotopes

4.0-viii Rev. 29 WOLF CREEK CHAPTER 4 - LIST OF FIGURES

*Refer to Section 1.6 and Table 1.6-3. Controlled drawings were removed from the USAR at Revision 17 and are considered incorporated by reference.

Figure # Sheet Title Drawing #* 4.3-3 Boron Concentration Versus First Cycle Burnup With and Without Burnable Poison Rods 4.3-4 Typical Integral Fuel Burnable Absorber Rod Arrangement Within an Assembly 4.3-5 Typical Integral Fuel Burnable Absorber and Source Assembly Locations 4.3-6 MTC vs Burnup at HFP, ARO, Critical Conditions (Typical) 4.3-7 Deleted 4.3-8 Deleted 4.3-9 Deleted 4.3-10 Deleted 4.3-11 Deleted 4.3-12 Rodwise Power Distribution in a Typical Assembly Near Beginning-of-Life, Hot Full Power, Equilibrium Xenon, Unrodded Core 4.3-13 Rodwise Power Distribution in a Typical Assembly Near End-of-Life, Hot Full Power, Equilibrium Xenon, Unrodded Core 4.3-14 Typical Axial Power Shapes Occurring at Beginning-of-Life 4.3-15 Typical Axial Power Shapes Occurring at Middle-of-Life 4.3-16 Typical Axial Power Shapes Occurring at End-of-Life 4.3-17 Comparison of a Typical Assembly Axial Power Distribution With Core Average Axial Distribution Bank D Slightly Inserted 4.3-18 Deleted 4.3-19 Deleted 4.3-20 Deleted 4.3-21 Maximum F X Q Power Versus Axial Height During Normal Operation 4.3-22 Deleted 4.3-23 Deleted 4.3-24 Comparison Between Calculated and Measured Relative Fuel Assembly Power Distribution 4.3-25 Comparison of Typical Calculated and Measured Axial Shapes 4.3-26 Measured Values of F Q for Full Power Rod Configurations

4.0-ix Rev. 17 WOLF CREEK CHAPTER 4 - LIST OF FIGURES

*Refer to Section 1.6 and Table 1.6-3. Controlled drawings were removed from the USAR at Revision 17 and are considered incorporated by reference.

Figure # Sheet Title Drawing #* 4.3-27 Deleted 4.3-28 Deleted 4.3-29 Deleted 4.3-30 Deleted 4.3-31 Deleted 4.3-32 Deleted 4.3-33 Deleted 4.3-34 Deleted 4.3-35 Deleted 4.3-36 Rod Cluster Control Assembly Pattern 4.3-37 Example Differential and Integral Rod Worth Versus Steps Withdrawn at MOL, HZP, HFP, EQ Xenon Banks D, C, and B Moving With 113 Step Overlap 4.3-38 Design Trip Curve 4.3-39 Typical Normalized Rod Worth Versus Percent

Insertion, All Rods Out But One 4.3-40 Axial Offset Versus Time, PWR Core With A 12 Foot

Height and 121 Assemblies 4.3-41 X-Y Xenon Test Thermocouple Response Quadrant Tilt Difference Versus Time 4.3-42 Calculated and Measured Doppler Defect and Coefficients at BOL, 2-Loop Plant, 121 Assemblies, 12

Foot Core 4.3-43 Comparison of Calculated and Measured Boron Concentration 2-Loop Plant, 121 Assemblies, 12 Foot

Core 4.3-44 Comparison of Calculated and Measured C B , 2-Loop Plant, 121 Assemblies, 12 Foot Core 4.3-45 Comparison of Calculated and Measured C B , 3-Loop Plant, 157 Assemblies, 12 Foot Core 4.3-46 Typical Boron Letdown Curve 4.4-1 Deleted 4.4-2 Deleted 4.4-3 Deleted 4.4-4 TDC Versus Reynolds Number for 26 Inch Grid Spacing 4.4-5 Normalized Radial Flow and Enthalpy Distribution at 4

Foot Elevation

4.0-x Rev. 17 WOLF CREEK CHAPTER 4 - LIST OF FIGURES

*Refer to Section 1.6 and Table 1.6-3. Controlled drawings were removed from the USAR at Revision 17 and are considered incorporated by reference.

Figure # Sheet Title Drawing #* 4.4-6 Normalized Radial Flow and Enthalpy Distribution at 8 Foot Elevation 4.4-7 Normalized Radial Flow and Enthalpy Distribution at

12 Foot Elevation - Core Exit 4.4-8 Void Fraction Versus Thermodynamic Quality 4.4-9 Thermal Conductivity of UO 2 (Data Corrected to 95% Theoretical Density) 4.4-10 Reactor Coolant System Temperature - Percent Power Map 4.4-11 100% Power Shapes Evaluated at Conditions Representative of Loss of Flow, All Shapes Evaluated with F N DH = 1.55 4.4-12 Deleted 4.4-13 Deleted 4.4-14 Deleted 4.4-15 Deleted 4.4-16 Deleted 4.4-17 Deleted 4.4-18 Deleted 4.4-19 Deleted 4.4-20 Deleted 4.4-21 Moveable Detector and Thermocouple Locations 4.4-22 Measured versus Predicted Critical Heat Flux - WRB-2 Correlation

4.0-xi Rev. 17 WOLF CREEK CHAPTER 4.0 REACTOR 4.1

SUMMARY

DESCRIPTION This chapter describes: 1) the mechanical components of the reactor and reactor core, including the fuel rods and fuel assemblies, 2) the nuclear design, and

3) the thermal-hydraulic design.

The reactor core is composed of an array of fuel assemblies that are similar in mechanical design, but different in fuel enrichment. Within each fuel assembly, all rods are of the same enrichment. Three different enrichments were employed in the first core. The enrichments for Cycle 1 at Wolf Creek were 2.10 (Region 1), 2.60 (Region 2), and 3.10 (Region 3) weight percent. The average enrichments were increased in subsequent reloads in order to achieve an

eighteen month cycle. This began in Cycle 2 and Cycle 4 was the first eighteen

month cycle. Enrichments up to 5.0 weight percent may be used for reload fuel

when credit is taken for integral fuel burnable absorbers (IFBA) or 4.6 weight

percent without credit of IFBA. The Westinghouse 17x17 low-parasitic (LOPAR) fuel design was used during cycle 1 and for the fresh fuel loaded in Cycles 2 and 3 as well. Cycle 4 fresh fuel

incorporated the anti-snag grid design into the LOPAR fuel design. Cycle 5

fresh fuel added the reconstitutable top nozzle (RTN) and debris filter bottom

nozzle (DFBN) features to the WCGS fuel design. Cycle 6 fresh fuel incorporated the low pressure drop Zircaloy mid grid feature as described in Reference 1. With the incorporation of the Zircaloy mid grids, the WCGS fuel

design changed from the LOPAR design to the Westinghouse VANTAGE 5H (V5H)fuel

design. Cycle 7 fresh fuel incorporated the Zircaloy Intermediate Mixing Vane

Grids (IFM), as described in Reference 1, to provide additional coolant mixing in the upper fuel regions. An Inconel Protective Bottom Grid (PBG) was added to Cycle 8 fresh fuel to provide an additional debris barrier and increased fretting resistance. Cycle 9 fresh fuel incorporated the Integral Fuel

Burnable Absorber (IFBA) design, as described in Reference 1, as an alternative

to discrete burnable absorbers. Cycle 10 fresh fuel incorporated fully enriched annular axial blankets and the use of Zirlo as the material for the manufacture of the fuel clad, guide thimble and instrumentation tubes, mid

grids, and IFM grids. With the incorporation of the Zirlo material, the WCGS fuel design changed to the Westinghouse VANTAGE 5H with Performance + features (V5H P+) fuel design. The V5H P+ design is the .374 outside diameter rod

equivalent to the VANTAGE+ design discussed in Reference 2. The Cycle 10 fresh

fuel also included 8 demonstration assemblies of the Robust Fuel Assembly (RFA)

design. The differences between the V5H P+ design and the RFA design are

discussed in Reference 4. The Cycle 12 fresh fuel incorporated a revised rod

design that increases the void volume available in the fuel rod and is referred

to as the low rod internal pressure fuel rod design. The low rod internal

pressure fuel rod design is discussed in Reference 5. With the incorporation

of the low rod internal pressure fuel rod design the WCGS fuel design changed

to the Westinghouse VANTAGE 5H with Performance + features, Zirlo +2 (V5H P+Z+2) fuel design. The Cycle 13 fresh fuel incorporated the features of the Robust Fuel Assembly design, including modified mid-grids, modified IFM grids, and thicker wall guide thimble and instrument tubes, into the V5H P+Z +2 design. With the incorporation of these features the WCGS fuel design changed to the Westinghouse Standard Fuel Rod Robust Fuel Assembly Zirlo +2 (STD RFA Z +2 or RFA Z+2) design. The Cycle 13 fresh fuel also included 4 demonstration assemblies that incorporated the RFA-2 mid-grid design and the Integral Clamp Top Nozzle (ICTN) design. The RFA-2 mid-grid is an improved mid-grid that provides increased margin for fretting wear, while maintaining the RFA mid-grids 4.1-1 Rev. 16 WOLF CREEK performance in other areas such as DNB and pressure drop. The key difference

between the RFA and RFA-2 mid-grid design is the increased spring and dimple

contact area with the fuel rod. The complete discussion of the differences

between the modified mid-grid used in the RFA Z +2 design and RFA-2 mid-grid is contained in Reference 6. The ICTN includes a modified top nozzle casting that

includes the spring clamps. The springs are located with pins that are welded

in place (to the integral clamp) but do not react to the spring force. The

ICTN design eliminates the potential for the fracture of the hold down spring

screws by the removal of the spring screws in the ICTN design. The

modification increases the fuel assembly integrity and eliminates the potential

for loose parts from fractured spring screws entering the RCS during normal

operations or during fuel movement during refueling outages. The features of

the Integral Clamp Top Nozzle are discussed in Reference 7. The Cycle 14 fresh fuel incorporated the features of the 17x17 RFA-2 (second

generation Robust Fuel Assembly) design, including modified mid-grids, modified

IFM grids, and thicker wall guide thimble and instrument tubes. The RFA-2

design is identical to the RFA design except for the mid-grid. The key

difference between the RFA and RFA-2 mid-grid design is the increased spring

and dimple contact area with the fuel rod. There is no change to the fuel

assembly length, envelope or fuel rod design relative to the RFA design. The

RFA-2 mid-grid is an improved mid-grid that provides increased margin for

fretting wear while maintaining the RFA mid-grids performance in others areas

such as DNB and pressure drop. The complete discussion of the differences

between the RFA-2 Z +2 modified mid-grid design and the RFA-2 Z +2 mid-grid design is contained in Reference 6. The Cycle 16 fresh fuel incorporates the Westinghouse Integral Nozzle (WIN) top

nozzle and a Performance+ feature of fuel rod oxide coating. The WIN top nozzle

was previously known as the Integral Clamp Top Nozzle (ICTN) and was introduced

in four demonstration assemblies in Cycle 13. The features of the WIN top

nozzle are discussed in Reference 8. The fuel rod has an oxide coating at the

bottom end of the fuel rod. The extra layer of oxide coating provides

additional debris induced rod fretting wear protection. The features of the

fuel rod oxide coating are discussed in Reference 9. The Cycle 21 fresh fuel incorporates a Standardized Debris Filter Bottom Nozzle (SDFBN) and a Robust Protective Grid. The Robust Protective Grid is provided as part of the Combination Grid which also included the bottom grid. This change will impact the location of the Protective Grid centerline in relation to the bottom of the fuel stack and the elevation of the Protective Grid to the bottom of the bottom nozzle. The SDFBN evaluation is discussed in Reference 10 and later in Section 4.2.2.2.1. The Robust Protective Grid is discussed in Reference 11 and Section 4.2.2.2.4. The core may consist of any combination of LOPAR, V5H, V5H P+, RFA, V5H P+ Z +2 , RFA Z+2 and RFA-2 Z +2 fuel assemblies as described in Subsection 4.2.2. The fuel is arranged in a checkered low-leakage pattern. A fuel assembly is composed of 264 fuel rods in a 17 x 17 square array, except

that limited substitution of filler rods for fuel rods may be made (Reference

3). The center position in the fuel assembly is reserved for incore

instrumentation. The additional 24 positions in the fuel assembly have guide

thimbles for the rod cluster control assemblies (RCCAs). The guide thimbles

are joined to the bottom nozzles of the fuel assembly and also serve to support

the fuel grids. The fuel grids consist of an "egg-crate" arrangement of

interlocked straps that maintain lateral spacing between the rods. The straps

have spring fingers and dimples which grip and support the fuel rods. The

grids also have coolant-mixing vanes. The fuel rods consist of slightly

enriched uranium, in the form of cylindrical pellets of uranium dioxide, contained in Zircaloy-4/Zirlo tubing. The tubing is plugged and seal-welded at the ends to encapsulate the fuel. All fuel rods are pressurized internally with helium during fabrication to reduce clad creepdown during operation and

thereby to increase fatigue life. 4.1-2 Rev. 29 WOLF CREEK The bottom nozzle is a box-like structure which serves as the lower structural element of the fuel assembly and directs the coolant flow distribution to the assembly. The top nozzle assembly serves as the upper structural element of the fuel assembly and provides a partial protective housing for the RCCA or other components. The RCCAs consist of 24 absorber rods fastened at the top end to a common hub or spider assembly. Each absorber rod consists of either all hafnium or an

alloy of silver-indium-cadmium clad in stainless steel. The RCCAs are used to

control relatively rapid changes in reactivity and to control the axial power

distribution. The reactor core is cooled and moderated by light water at a pressure of 2250 psia. Soluble boron in the moderator/coolant serves as a neutron absorber. The concentration of boron is varied to control reactivity changes that occur

relatively slowly, including the effects of fuel burnup and transient xenon.

Burnable absorber rods were also employed in the first core and subsequent reloads to limit the amount of soluble boron required and thereby maintain the

desired range of reactivity coefficients. Either the borosilicate glass burnable absorber, the Wet Annular Burnable Absorber (WABA), or the Integral

Fuel Burnable Absorber (IFBA) are included in subsequent reloads. The nuclear design analyses established the core locations for control rods and burnable absorbers and define design parameters, such as fuel enrichments and boron concentration in the coolant. The nuclear design analyses established that the reactor core and the reactor control system satisfy all design

criteria, even if the highest reactivity worth RCCA is in the fully withdrawn

position. The core has inherent stability against diametral and azimuthal

power oscillations. Axial power oscillations which may be induced by load

changes and resultant transient xenon may be suppressed by the use of the

control rods (RCCAs). The thermal-hydraulic design analyses established that adequate heat transfer is provided between the fuel clad and the reactor coolant. The thermal design takes into account local variations in dimensions, power generation, flow distribution, and mixing. The mixing vanes incorporated in the fuel assembly spacer grid design induce additional flow-mixing between the various flow

channels within a fuel assembly as well as between adjacent assemblies. The performance of the core is monitored by fixed neutron detectors outside of the core, movable neutron detectors within the core, and thermocouples at the outlet of selected fuel assemblies. The ex-core nuclear instrumentation provides input to automatic control functions. Table 4.1-1 presents the principal nuclear, thermal-hydraulic, and mechanical design parameters of WCGS. The analytical techniques employed in the core design are tabulated in Table 4.1-2. The mechanical loading conditions considered 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 4.2.1.5 and neutron absorber rods, burnable absorber rods, neutron source rods, and thimble plug devices in Section 4.2.1.6. The

dynamic analyses, input forcing functions, and response loadings are presented

in Section 3.9(N). 4.1-3 Rev. 21 WOLF CREEK 4.

1.1 REFERENCES

1. Davidson, S. L., (Ed.), et al., "VANTAGE 5 Fuel Assembly Reference Core Report," WCAP-10444-P-A, September 1985, and Addendum 2A, February 1989.

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

3. Slagle, W. H. (Ed.), "Westinghouse Fuel Assembly Reconstitution Evaluation Methodology," WCAP-13060-P-A, July 1993.

4. O'Cain, M.B., (Ed.), "17 x 17 Wolf Creek Robust Fuel Assembly Final Design Review Package," DR-97-2 (Proprietary), July 1997.

5. Shah, H., (Ed), "Low RIP Fuel Rod Design," DR-98-04 (Proprietary), October 1998. 6. Seel, D. D. (Ed.), "17 x 17 Robust Fuel Assembly with RFA-2 Mid-Grid Final Design Review Package", DR-01-5 (Proprietary), October 2001.

7. Maurer, B.F., "Generic - Implementation of Integral Clamp Top Nozzle (ICTN) Design Change - 15 x 15 and 17 x 17, 12 Foot Fuel Assemblies",

EVAL-01-067, November 2001.

8. Kitchen, T. J., "Generic - Implementation of the Westinghouse Integral Nozzle (WIN) for 15x15 and 17x17, 12 foot Fuel Assemblies," EVAL-04-10, February 2004.

9. Kitchen, T. J., "Revision to Generic PERFORMANCE+ Fuel Assembly Safety Evaluation - SECL-92-305, Rev. 0," EVAL-03-127, May 2004.

10. Solomon, D. K., "17x17 Standardized Debris Filter Bottom Nozzle with Flow Hole Elimination - Impact on Best Estimate Flow (BEF) Calculations", EVAL-09-8, Revision 9, December 2013.

11. Solomon, D. K., "Generic - 17x17 Robust Protective Grid (RPG)", EVAL 12, Revision 4, October 2012.

4.1-4 Rev. 29 WOLF CREEK TABLE 4.1-1 REACTOR DESIGN TABLE Thermal and Hydraulic Design Parameters WCGS 1. Reactor core heat output, MWt 3,565 2. Reactor core heat output, 10 6 Btu/hr 12,480 3. Heat generated in fuel, % 97.4 4. System pressure, nominal, psia 2,250 5. System pressure, minimum steady state, psia 2,220 6. Minimum departure from nucleate boiling ratio for design transients 1.76 (WRB-2) 1.30 (W-3) 7. DNB correlation WRB-2 or W-3 Coolant Flow 8. Total thermal flow rate, gpm 361,296 9. Effective flow rate for heat transfer, gpm (6.61% bypass flow assumed) 337,414 Rev. 10 WOLF CREEK TABLE 4.1-1 (Sheet 2) Thermal and Hydraulic Design Parameters WCGS Coolant Flow (Continued) 10. Effective flow area for heat transfer, ft 2 51.3 11. Average velocity along fuel rods, ft/sec 14.7 12. Average mass velocity, 10 6 lb m/hr-ft 2 2.31 13. Nominal inlet, F 553.7 14. Average rise in vessel, F 65.6 15. Average rise in core, F 68.6 16. Average in core, F 588.0 17. Average in vessel, F 586.5 Heat Transfer 18. Active heat transfer, surface area, ft 2 59,742 19. Average heat flux, Btu/hr-ft 2 198,340 20. Maximum heat flux for normal operation, Btu/hr-ft 2 460,100 21. Average linear power, kW/ft 5.68 Rev. 13 WOLF CREEK TABLE 4.1-1 (Sheet 3) Thermal and Hydraulic Design Parameters WCGS Heat Transfer (Continued) 22. Peak linear power for normal operation, kW/ft 14.48 23. Peak linear power resulting from overpower transients/operator errors, assuming a maxi-mum overpower of 118%, kW/ft 21.8 a 24. Heat flux hot channel factor, F Q 2.50 b 25. Peak fuel control temperature at peak linear power for prevention of centerline melt, F 4,700 Core Mechanical Design Parameters 26. Number of fuel assemblies 193

27. Designs RCC canless 17 X 17 LOPAR RCC canless 17 X 17 V5H RCC canless 17 X 17 V5H w/IFM RCC canless 17 X 17 V5H w/IFM & PBG RCC canless 17 x 17 V5H P+ RCC canless 17 x 17 V5H P+ Z+2 RCC canless 17 x 17 RFA Z+2 and RFA-2 Z+2 28. UO 2 rods per assembly 264 264 264 264 264 264 264 29. Rod pitch, in. 0.496 0.496 0.496 0.496 0.496 0.496 0.496 30. Overall dimensions, in.

8.426 x 8.426 8.426 x 8.426 8.426 x 8.426 8.426 x 8.426 8.426 x 8.426 8.426 x 8.426 8.426 x 8.426 31. Fuel weight, as UO 2 , lb per assembly (typical)1154 1154 1154 1149 1132 1138 1138 Rev. 21 WOLF CREEK TABLE 4.1-1 (Sheet 4) Core Mechanical Design Parameters 32. Zircaloy/Zirlo weight, lb per assembly (Approx.) 264 270 275 278 275 274 274 33. Number of grids per assembly See Note 1See Note 2 See Note 3 See Note 4 See Note 5 See Note 5 See Note 5 34 Loading technique 3 Region Nonuniform 3 Region Nonuniform 3 Region Nonuniform 3 Region Nonuniform 3 Region Nonuniform 3 Region Nonuniform 3 Region Nonuniform Note 1 8 Total Grids, 1 Inconel Top Grid, 6 Inconel Mid Grids, 1 Inconel Bottom Grid Note 2 8 Total Grids, 1 Inconel Top Grid, 6 Zircaloy Mid Grids, 1 Inconel Bottom Grid Note 3 11 Total Grids, 1 Inconel Top Grid, 6 Zircaloy Mid Grids, 3 Zircaloy IFM Grids, 1 Inconel Bottom Grid Note 4 12 Total Grids, 1 Inconel Top Grid, 6 Zircaloy Mid Grids, 3 Zircaloy IFM Grids, 1 Inconel Bottom Grid, l Inconel Protective Bottom Grid Note 5 12 Total, 1 Inconel Top Grid, 6 Zirlo Mid Grids, 3 Zirlo Intermediate Flow Mixing Grids, 1 Inconel Bottom Grid, 1 Inconel Protective Bottom Grid or 1 Robust Protective Bottom Grid Fuel Rods 35. Total Number of Fuel Rods in the core 50,952 36. Outside diameter, in. 0.374 0.374 0.374 0.374 0.374 0.374 0.374 37. Diametral gap, in. 0.0065 0.0065 0.0065 0.0065 0.0065 0.0065 0.0065 38. Clad thickness, in. 0.0225 0.0225 0.0225 0.0225 0.0225 0.0225 0.0225 39. Clad material Zircaloy-4 Zircaloy-4 Zircaloy-4 Zircaloy-4 Zirlo Zirlo Zirlo Fuel Pellets

40. Material UO 2 sintered 41. Density % of theoretical 95

42. Diameter, in. 0.3225

Rev. 29 WOLF CREEK TABLE 4.1-1 (Sheet 5) Core Mechanical Design Parameters WCGS Fuel Pellets 43. Length, in (range) 0.372 - 0.530 44. Fuel Enrichment, Weight Percent (range) 2.1 - 5.0 45. Deleted Rod Cluster Control Assemblies 46. Number of clusters, full length / part length 53 / - 47. Neutron absorber Full length, Hafnium Ag-In-Cd 48. Cladding Material Type 304 SS-cold worked Type 304 SS-cold worked 49. Clad thicknesses, in 0.0185 0.0185 50. Number of absorber rods per cluster 24 24 Rev. 14 WOLF CREEK TABLE 4.1-1 (Sheet 6) Core Mechanical Design Parameters WCGS Core structure 51. Core barrel, I.D./O.D., in. 148.0/152.5 52. Thermal shield Neutron pad design 53. Baffle thickness, in. 0.88 Structure Characteristics 54. Core diameter, equivalent, in. 132.7 55. Core height, active fuel, in. 143.7 Reflector Thickness and Composition 56. Top, water plus steel, in. 10 57. Bottom, water plus steel, in. 10 58. Side, water plus steel, in. 15 59. H 2 O/U molecular ratio core, lattice, cold 2.41 Notes: (a) See Section 4.3.2.2.6. (b) This is the value of F Q for normal operation. (c) Limited substitution of filler rods for fuel rods is allowed. Rev. 11 WOLF CR EE K TABL E 4.1-2 ANALYTICAL T E CHNIQU E S IN COR E D E SIGN SectionANALYSISTechnique Computer Code ReferencedMechanical design of coreStatic and dynamicBlowdown code,3.7(N).2.1internals, loads,modelingFORC E, finite3.9(N).2deflections, andelement, struc-3.9(N).3stress analysistural analysis code, and others Fuel rod designFuel performance charac-Semiempirical thermalWestinghouse fuel4.2.1.1teristics (temperature,model of fuel rod withrod design model4.2.3.2 internal pressure, cladconsideration of fuel4.2.3.3 stress, etc.)density changes, heat4.3.3.1transfer, fission gas4.4.2.11 release, etc. Nuclear design (initial core design)1.Cross sections andMicroscopic data:Modified ENDF/B4.3.3.2group constantsmacroscopic constants Library L E OPARDfor homogenized coreCIND E R type regions Rev. 11 WOLF CR EE K TABL E 4.1-2 (Sheet 2) Section ANALYSIS Technique Computer Code Referenced Nuclear Design (initial core design)Group constants forHAMM ER-AIM4.3.3.2 control rods with self-shielding2.X-Y power distribu-2-D, 2-group diffusionTURTL E 4.3.3.3tions, fuel depletion,theory critical boron concen-trations, X-Y xenon distributions, reac-

tivity coefficients3.Axial power distribu-1-D, 2-group diffusionPANDA4.3.3.3tions, control rodtheory worths, and axial

xenon distribution4.Fuel rod powerIntegral transportLAS ER4.3.3.1 theory 5.Effective resonanceMonte Carlo weightingR E PADtemperaturefunction Rev. 11 WOLF CR EE K TABL E 4.1-2 (Sheet 3) Section ANALYSIS Technique Computer Code Referenced Nuclear Design (Continued)

6. Criticality of reac- 2-D, 2-group diffusion L E OPARD 4.3.2.6 tor and fuel assem- theoryPDQ blies

7. Vessel irradiation Multigroup spatial DOT 4.3.2.8 dependent transport theory Thermal-hydraulic design (initial core design)

1. Steady state Subchannel analysis of THINC-IV 4.4.4.5.2 local fluid conditions in rod bundles, includ-ing inertial and cross-

flow resistance terms, solution progresses from

core-wide to hot assem-

bly to hot channel

2. Transient departure Subchannel analysis of THINC-I 4.4.4.5.2 from nucleate boil- local fluid conditions (THINC-III) ing analysis in rod bundles during transients by including

accumulation terms in

conservation equations;

solution progresses from

core-wide to hot assembly

to hot channel Rev. 11 WOLF CREEK TABLE 4.1-2 (Sheet 4) Section ANALYSIS Technique Computer Code Referenced Nuclear Design (Reload Pattern Design & Analysis)

1. Cross section and Micro-group neutron spectrum, NEXUS/PARAGON 4.3.3.2 group constants cell average few-group cross sections, assembly average nodal constants

2. Rod worths, Boron 3-D, Diffusion Theory ANC9 4.3.3.3 worths & letdown, - based Nodal Method reactivity coefficients

3. Nodal power distribu- 3-D, Diffusion Theory ANC9 4.3.3.3 tions - based Nodal Method

4. Fuel Pin Powers, 3-D, Diffusion Theory ANC9 4.3.3.3 INCORE constants - based Nodal Method

Thermal-hydraulic design (Reload Pattern Design & Analysis)

1. Steady state and Subchannel analysis of VIPRE-01 4.4.4.5.2 transient departure local fluid conditions from nucleate in rod bundles, including boiling analysis inertial and cross-flow resistance terms

Rev. 29 WOLF CR EE K TABL E 4.1-3 D E SIGN LOADING CONDITIONS FOR R E ACTOR COR E COMPON E NTS 1. Fuel assembly weight

2. Fuel assembly spring forces

3. Internals weight

4. Control rod trip (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 3.9(N)-1

14. Pump overspeed

15. Seismic loads (Operating Basis E arthquake and Safe Shutdown E arthquake)

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

Rev. 0 WOLF CREEK 4.2 FUEL SYSTEM DESIGN The plant design conditions 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; and Condition IV - Limiting Faults. Chapter 15.0 describes bases and plant operation and events involving each condition. The reactor is designed so that its components meet the following performance and safety criteria:

a. 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) ensure that:

1. 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 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 immediate resumption of

operation.

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.

b. The fuel assemblies are designed to withstand loads induced during shipping, handling, and core loading without exceeding the criteria of Section 4.2.1.5.

c. 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 (if in such locations).

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

**  In any case, the fraction of fuel rods damaged must be limited so as to meet the dose guideline of 10 CFR 100.      4.2-1    Rev. 12 WOLF CREEK

d. All fuel assemblies have provisions for the insertion of incore instrumentation necessary for plant operation.

e. The reactor internals, in conjunction with the fuel assemblies and incore control components, direct reactor coolant through the core. This achieves acceptable flow distribution and restricts bypass flow so that the heat transfer performance requirements can be met for all

modes of operation. 4.2.1 DESIGN BASES

The fuel rod and fuel assembly design bases are established to satisfy the general performance and safety criteria presented in this section. Design values for the properties of the materials which comprise the fuel rod, fuel assembly, and incore control components are given in Reference 2 for Zircaloy clad fuel and in Reference 20 for Zirlo clad fuel. Other supplementary fuel design criteria/limits are given in Reference 21. 4.2.1.1 Cladding a. Material and Mechanical Properties Zircaloy-4 and Zirlo combines neutron economy (low absorption cross-section); high corrosion resistance to coolant, fuel, and fission products; and high strength and ductility at operating temperatures. Reference 1 documents the operating experience with Zircaloy-4 and Zirlo as a clad material. Information on the mechanical properties of the cladding is given in References 2 and 20 with due consideration of temperature and irradiation effects. b. Stress-strain limits 1. Clad stress The von Mises criterion is used to calculate the effective stresses. The cladding stresses under Condition I and II events are less than the Zircaloy 0.2% offset yield stress, with due consideration of temperature and irradiation effects. While the cladding has some capability for accommodating plastic strain, the yield stress has been accepted as a conservative design basis.2. Clad tensile strain The total tensile creep strain is less than 1% from the unirradiated condition. The elastic tensile strain during a transient is less than 1% from the pretransient value. This limit is consistent with proven practice. 4.2-2 Rev. 12 WOLF CREEK

c. Vibration and fatigue

1. Strain fatigue

The cumulative strain fatigue cycles are less than the design strain fatigue life. This basis is consistent with proven practice. (Ref. 1).

2. Vibration

Potential fretting wear due to vibration is prevented, ensuring that the stress-strain limits are not exceeded during design life. Fretting of the clad surface can occur due to flow-induced vibration between the fuel rods and fuel assembly grid springs. Vibration and fretting forces vary during

the fuel life due to clad diameter creepdown combined

with grid spring relaxation.

d. Chemical properties Chemical properties of the cladding are discussed in Reference 2 for Zircaloy and Reference 20 for Zirlo.4.2.1.2 Fuel Material

a. Thermal-physical properties The thermal-physical properties of U0 2 are described in Reference 2 with due consideration of temperature and irradiation effects.

Fuel pellet temperatures - The center temperature of the hottest pellet is below the melting temperature of the UO 2 [melting point of 5080 F (Ref. 3) unirradiated and decreasing 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 fuel centerline temperature of 4700 F has been selected as an overpower limit to ensure no fuel melting. This provides sufficient margin for uncertainties, as

described in Section 4.4.2.9. The normal design density of the fuel is 95 percent of theoretical. Additional information on fuel properties

is given in Reference 2.

b. Fuel densification and fission product swelling

The design bases and models used for fuel densification and swelling are provided in Reference 18.

c. Chemical properties Reference 2 provides the justification that no adverse chemical interactions occur between the fuel and its

adjacent material. 4.2-3 Rev. 18 WOLF CREEK 4.2.1.3 Fuel Rod Performance The detailed fuel rod design establishes such parameters as pellet size and density, cladding-pellet diametral gap, gas plenum size, and helium pre-pressurization level. The design also considers the effects such as fuel density changes, fission gas release, cladding creep, and other physical properties which vary with burnup. 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 to satisfy the conservative design bases in the following subsections during Condition I and

Condition II events over the fuel lifetime. For each design basis, the performance of the limiting fuel rod must not exceed the limits specified.

a. Fuel rod models

The basic fuel rod models and the ability to predict operating characteristics are given in References 17, 18, 27, and Section 4.2.3.

b. Mechanical design limits Fuel rod design methodology has been introduced that reduces the densification power spike factor to 1.0 and Reference 19 demonstrates

that clad flattening will not occur in Westinghouse fuel designs. The rod internal gas pressure remains below the value which causes the fuel/clad diametral gap to increase due

to outward cladding creep during steady state operation.

Rod pressure is also limited so that extensive departure

from nucleate boiling (DNB) propagation does not occur during normal operation and any accident event. (Reference 7). 4.2.1.4 Spacer Grids a. Mechanical limits and material properties The grid component strength criteria are based on experimental tests. The limit is established at the lower 95%confidence on the true mean crush strength. This limit is sufficient to ensure that under worst-case combined seismic and blowdown loads from a Condition III and IV, loss-of-coolant accident, the core will maintain a geometry amenable to cooling. As an integral part of the fuel

assembly structure, the grids satisfy the applicable fuel assembly design bases and limits defined in Section 4.2.1.5. The grid material and chemical properties are given in Reference 2 for Zircaloy-4 and Reference 20 for Zirlo. 4.2-4 Rev. 18 WOLF CREEK b. Vibration and fatigue The grids provide sufficient fuel rod support to limit fuel rod vibration and maintain clad fretting wear to within acceptable limits (defined in Section 4.2.1.1). 4.2.1.5 Fuel Assembly a. Structural design Integrity of the fuel assembly structure is ensured by setting design limits on potential stresses and deformations due to various loads and

by preventing the assembly structure from interfering with the

functioning of other components. Three types of loads are considered. 1. Non-operational 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 limits are applied to the design and evaluation of the top and bottom nozzles, guide thimbles, grids, and the thimble joints. The design bases for evaluating the structural integrity of the fuel assemblies are: 1. Nonoperational - 4 g axial and 6 g lateral loading with dimensional stability. 2. For the normal operating and upset conditions, the fuel assembly component structural design criteria are established for the two primary material categories, namely austenitic steels and Zirconium Alloys. 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) One-third of the specified minimum tensile strength or 2/3 of the specified minimum yield strength at room temperature; (b) One-third 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. 4.2-5 Rev. 15 WOLF CREEK The stress limits for the austenitic steel components are given below. All stress

nomenclature is per the ASME Code, Section III. Stress Intensity Limits Categories Limit General primary membrane Sm stress intensity Local primary membrane 1.5 Sm stress intensity Primary membrane plus bending 1.5 Sm stress intensity Total primary plus secondary 3.0 Sm stress intensity The Zircaloy or Zirlo structural components, which consist of guide thimbles, fuel tubes, and mixing grids 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. The

maximum shear theory is used to evaluate the

guide thimble design. For conservative purposes, the Zircaloy and Zirlo unirradiated properties are used to define the stress limits. (c) Abnormal loads during Condition III or IV - worst cases represented by combined seismic and

blowdown loads.

1. Deflections or failures of components cannot interfere with the reactor shutdown or

emergency cooling of the fuel rods.

2. The fuel assembly structural component stresses under faulted conditions are evaluated using primarily the methods

outlined in Appendix F of the ASME Code, Section III. 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 for normal operating conditions. For the Zircaloy and Zirlo components, the stress intensity, Sm, is set as the smaller value of 2/3 of the material yield s t r ength, S y, o r 1/3 of the ult i mate s t r ength, S u, at r eacto r ope r at i ng temperature. This results in Zircaloy and Zirlo stress 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 bending. For conservative purposes, the Zircaloy and Zirlounirradiated properties are used to define the stress limits. 4.2-6 Rev. 18 WOLF CREEK The material and chemical properties of the fuel assembly components are given in

Reference 2 for Zircaloy-4 and Reference 20 for Zirlo. 3. Thermal-hydraulic design This topic is discussed in Section 4.4. 4.2.1.6 Incore Control Components The control components are subdivided into permanent and temporary devices. The permanent type components are the rod cluster control assemblies, secondary neutron source assemblies, and thimble plug devices. The temporary components are the burnable absorber assemblies and the primary neutron source assemblies, which are normally used only in the initial core. Materials are selected for compatibility in a pressurized water reactor 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. Material

properties are given in Reference 2. The design bases for each of the mentioned components are given in the following subsections.

a. Control (neutron absorber) rods Design conditions which are considered under Article NB-3000 of the ASME Code, Section III are as follows:

1. External pressure equal to the reactor coolant system operating pressure with appropriate allowance for

overpressure transients

2. 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 imposed by the control rod drive mechanism

6. Radiation exposure during maximum core life

The control rod cladding is cold drawn Type 304 stainless steel tubing. The stress intensity limit, Sm, for this material is defined as 2/3 of the 0.2 percent offset yield stress. The absorber material temperature does not exceed its melting temperature

• . 7. Temperature effects at operating conditions
• The melting point basis is determined by the nominal material melting point minus uncertainty. 4.2-7 Rev. 18 WOLF CREEK b. Burnable absorber rods (standard and WABA) The cladding for burnable absorber rods is designed as a Class 1 component under Article NB-3000 of the ASME Code, Section III, 1973 for Conditions I and II. For abnormal loads during Conditions III and IV, code stresses are not considered limiting. Failures of the burnable absorber r od s during conditions III and IV do not interfere with reactor shutdown or cooling of the fuel rods. The burnable absorber material is nonstructural. The structural elements of the burnable absorber rod are designed to maintain the absorber geometry even if the absorber material is fractured. In addition, the structural elements are designed to prevent excessive slumping. The standard burnable absorber material is borosilicate glass and is designed so that the absorber material is below its softening temperature (1510 F + 18 F for reference 12.5 w/o boron rod). The softening temperature for borosilicate glass is defined in ASTM C 338. The wet annular burnable absorber (WABA) material is B 4 C contained in an Alumina matrix. Thermal-physical and gas release properties of Al 2 O 3-B 4 C are described in reference 8. The WABA rods are designed so that the absorber temperature does not exceed 1200 F during normal operation or an overpower transient. The 1200 F maximum temperature He gas release in a WABA rod will not exceed 30% (reference 8). c. 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 with appropriate allowance for overpressure transients, and 2. An internal pressure equal to the pressure generated by released gases over the source rod life 4.2-8 Rev. 18 WOLF CREEK

d. Thimble plug device The thimble plug device may be used to restrict bypass flow through those thimbles not occupied by absorber, source, or burnable absorber rods.

The thimble plug devices satisfy the following criteria:

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

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

3. Limit the flow through each occupied thimble to an acceptable design value 4.2.1.7 Surveillance Program Section 4.2.4.5 and Sections 8 and 23 of Reference 9 discuss the testing and fuel surveillance operational experience program that has been and is being

conducted to verify the adequacy of the fuel performance and design bases.

Fuel surveillance and testing results, as they become available, are used to

improve fuel rod design and manufacturing processes and ensure that the design

bases and safety criteria are satisfied. 4.2.2 DESIGN DESCRIPTION The fuel assembly, fuel rod, and incore control component design data are given in Table 4.3-1. Each fuel assembly consists of 264 fuel rods, 24 guide thimble tubes, and one instrumentation thimble tube arranged within a supporting structure. Limited substitution of filler rods for fuel rods may be made. The instrumentation thimble is located in the center position and provides a channel for insertion of an incore neutron detector, if the fuel assembly is located in an instrumented 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 thimble plug device, depending on the position of the particular fuel assembly in the core. Figure 4.2-1 and Figure 4.2-1a show a cross-section of typical fuel assembly arrays, and Figure 4.2-2, Figure 4.2-2a, Figure 4.2-2b, Figure 4.2-2c, and Figure 4.2-2d show a 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. Fuel assemblies are installed vertically in the reactor vessel and stand upright on the lower core plate, which is fitted with alignment pins to locate

and orient each 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 holddown springs on the top nozzle

of each fuel assembly to hold the fuel assemblies in place. 4.2-9 Rev. 18 WOLF CREEK The V5H P+ assembly skeleton is identical to that previously described for V5H except for those modifications necessary to accommodate the intended fuel operation to higher burnups. The modifications consist of the use of Zirloguide thimbles and small skeleton dimensional alterations to provide additional fuel assembly and rod growth space at the extended burnup levels. The V5H P+ fuel assembly is shorter than the V5H fuel assembly. The grid centerline elevations of the V5H P+ are identical to those of the V5H fuel assembly, except for the top grid. The V5H P+ top grid has been lowered. However, since the V5H P+ fuel is intended to replace the V5H fuel, the V5H P+ exterior assembly envelope is equivalent in design dimensions, and the functional

interface with the reactor internals is also equivalent to those of previous

Westinghouse fuel designs. Also, the V5H P+ fuel assembly is designed to be

mechanically and hydraulically compatible with the V5H fuel assembly. The same

functional requirements and design criteria as previously established for the Westinghouse V5H fuel assembly remains valid for the V5H P+ fuel assembly. Figure 4.2-2c shows a full-length view of the V5H P+ fuel assembly design. A comparison between Figure 4.2-2b and Figure 4.2-2c details the small skeleton dimensional alterations mentioned above. The V5H P+Z +2 assembly skeleton is similar to that previously described for V5H P+ except for those modifications necessary to accommodate the low rod internal pressure design and incorporation of a cast top nozzle design. The

modifications consist of the use of longer Zirlo TM guide thimbles and instrument tube and repositioning of the top grid. The V5H P+Z +2 fuel assembly is taller than the V5H P+ fuel assembly and the same height as the V5H fuel

assembly. Operational experience with the ZIRLO TM material has shown that the growth characteristics of ZIRLO TM do not require the shorter skeleton design used with the V5H P+ fuel assembly. The additional height of the V5H P+Z +2 fuel assembly skeleton allows the incorporation of fuel rod design

modifications to accrue rod internal pressure benefits (low rod internal

pressure rod design). The grid centerline elevations of the V5H P+Z +2 are identical to those of the V5H fuel assembly (all grids) and V5H P+ fuel

assembly except for the top grid. Since the V5H P+Z +2 fuel is intended to replace the V5H and V5H P+ fuel, the V5H P+Z +2 exterior assembly envelope is equivalent in design dimensions, and the functional interface with the reactor

internals is also equivalent to those of previous Westinghouse fuel designs.

Also, the V5H P+Z +2 fuel assembly is designed to be mechanically and hydraulically compatible with the V5H and V5H P+ fuel assembly. The same

functional requirements and design criteria as previously established for the

Westinghouse V5H and V5H P+ fuel assemblies remains valid for the V5H P+Z +2 fuel assembly. Figure 4.2-2d shows a full-length view of the V5H P+Z +2 fuel assembly design. A comparison between Figure 4.2-2c and Figure 4.2-2d details

the alterations mentioned above. The RFA Z+2 assembly skeleton is similar to that previously described for the V5H P+ Z+2 except for those modifications made to accommodate a modified mixing vane LPD mid-grid, a modified mixing vane IFM grid, and thicker guide thimble and instrument tubes. The grid changes are designed to improve thermal-

hydraulic performance and the addition of thicker thimble and instrument tubes

reduce the potential for fuel assembly bow and subsequently incomplete rod

insertion (IRI) concerns. The same functional requirements and design criteria

as previously established for the Westinghouse V5H P+Z +2 fuel assembly design remains valid for the RFA Z +2 design. Figure 4.2-2d shows a full-length view of the RFA Z +2 fuel assembly design. The RFA-2 Z +2 assembly skeleton is similar to that previously described for the RFA Z+2 except for the mid-grids. The differences between the RFA and RFA-2 mid-grids are the increased spring and dimple contact area with the fuel rod in the RFA-2 design. The same functional requirements and design criteria as previously established for the Westinghouse RFA Z +2 fuel assembly design remains valid for the RFA-2 Z +2 design. Figure 4.2-2d shows a full-length view of the RFA-2 Z +2 fuel assembly design. 4.2-10 Rev. 18 WOLF CREEK The RFA-2 Z+2 assembly skeleton was modified in Cycle 21 to include a combination bottom grid and Robust Protective Grid as well as a Standardized Debris Filter Bottom Nozzle. The grid change impacts the location of the Protective Grid centerline in relation to the bottom of the fuel stack and the elevation of the Protective Grid to the bottom of the bottom nozzle. Improper orientation of fuel assemblies within the core is prevented by the use of an indexing hole in one corner of the top nozzle top plate. The assembly is

oriented with respect to the handling tool and the core by means of a pin which

is inserted into this indexing hole. Visual confirmation of proper orientation

is also provided by an engraved identification number on the opposite corner

clamp.

4.2.2.1 Fuel Rods

Two types of fuel rod designs may be used in the V5H, V5H P+, V5H P+Z +2 , RFA Z+2 and RFA-2 Z +2 fuel assemblies. The fuel rod designs are referred to as Integral Fuel Burnable Absorber (IFBA) fuel rods and Non-IFBA fuel rods. The

IFBA and Non-IFBA fuel rod designs are identical with the exception of the

items noted in Section 4.2.2.1.2. A reference to fuel rods encompasses both

designs. The fuel rod structure consists of bottom end plug, a fuel tube (clad), uranium dioxide ceramic pellets, a plenum spring and top end plug. A

schematic of the fuel rod is shown in Figure 4.2-3, Figure 4.2-3a, Figure 4.2-

3b, Figure 4.2-3c, and Figure 4.2-3d.

4.2.2.1.1 Non-IFBA Fuel Rods

The LOPAR and V5H fuel rods consist of uranium dioxide ceramic pellets

contained in slightly cold worked Zircaloy-4 tubing, which is plugged, and seal

welded at the ends to encapsulate the fuel. The fuel pellets are right

circular cylinders consisting of slightly enriched uranium dioxide powder, which has been compacted by cold pressing and then sintered to the required

density. The ends of each pellet are dished slightly to allow greater axial

expansion at the center of the pellets.

Void volume and clearances are provided within the rods to accommodate fission

gases released from the fuel, differential thermal expansion between the clad

and the fuel, and fuel density changes during irradiation. Shifting of the

fuel within the clad during handling or shipping prior to core loading is

prevented by a stainless steel helical spring (plenum spring) which bears on

top of the fuel. At assembly, the bottom plug is inserted and welded and the

pellets are stacked in the clad to the required fuel height. The spring is

then inserted into the top end of the fuel tube and the top end plug is pressed

into the end of the tube and welded. All fuel rods are internally pressurized

with helium during the top end plug welding process in order to minimize

compressive clad stresses and prevent clad flattening under coolant operating

pressures. A schematic of the fuel rod is shown in Figure 4.2-3.

The fuel rods are prepressurized and designed so that: 1) the internal gas

pressure mechanical design limit given in Section 4.2.1.3 is not exceeded, 2)

the cladding stress-strain limits (see Section 4.2.1.1) are not exceeded for

Condition I and II events, and 3) clad flattening will not occur during the

fuel core life.

Cycle 2 fresh fuel incorporated a small chamfer on the end of each pellet at

the outer cylindrical surface and an internal gripper bottom end plug. The

internal gripper feature facilitates fuel rod loading and provides appropriate

lead-in for the removable top nozzle reconstitution feature.

Cycle 5 fresh fuel incorporated the high burnup short top and bottom end plug

design with a slightly longer fuel tube. A schematic of the fuel rod is shown

in Figure 4.2-3a.

4.2-11 Rev. 29 WOLF CREEK Cycle 8 fresh fuel incorporated the Performance+ top end plug, Performance+ extended bottom end plug, and variable pitch plenum spring. The extended bottom end plug is used in conjunction with the protective bottom grid discussed in Section 4.2.2.2.4. The variable pitch plenum spring has a smaller wire diameter, coil diameter and shorter free length. The variable pitch plenum spring provides the same support as the regular V5H plenum spring but with fewer turns, which translates into less spring volume and increased void volume in the rod. A schematic of the fuel rod is shown in Figure 4.2-3b. Cycle 10 fresh fuel incorporates the V5H P+ fuel rod. The V5H P+ fuel rod

represents a modification to the V5H fuel rod intended to support extended burnup operation for the fuel clad by using Zirlo in place of the Zircaloy-4 clad. The Zirlo alloy is a zirconium alloy similar to Zircaloy-4, which has been specifically developed to enhance corrosion resistance. The V5H P+ fuel rod has the same clad wall thickness as the V5H design. The V5H P+ fuel tube

is shorter to provide room for the required rod growth at extended burnups.

The V5H P+ fuel rods will contain, as in the V5H design, enriched uranium

dioxide fuel pellets. Schematics of the V5H P+ fuel rods are shown in Figure

4.2-3c.

Cycle 10 fresh fuel (V5H P+) incorporates the use of axial blankets in the fuel

rod. The axial blankets are a nominal 6 inches of unenriched fuel pellets or

fully enriched annular fuel pellets at each end of the fuel rod pellet stack.

Axial blankets reduce neutron leakage and improve fuel utilization. The use of

fully enriched annular fuel pellets in the axial blankets also provides

additional void volume. The axial blankets utilize chamfered pellets which are

physically different in length from the enriched pellets used in the rest of

the pellet stack to help prevent accidental mixing during manufacturing. Axial

blankets continue to be utilized in subsequent fresh fuel designs.

Cycle 12 fresh fuel incorporates the low rod internal pressure fuel rod design

associated with the V5H P+Z +2 fuel assembly design. Operational experience has shown that the ZIRLO TM material growth characteristics will accommodate a taller fuel assembly skeleton and a longer fuel rod than the V5H P+ design, while still allowing extended burnup operation. The V5H P+Z +2 fuel rod represents a modification to the V5H P+ fuel rod intended to provide additional

rod internal void volume to achieve rod internal pressure relief. The

additional void volume is created by the following configuration changes:

1) the V5H P+Z

+2 fuel rod top end plug does not include the external gripper feature of the Performance+ top end plug, resulting in a shorter top end

plug, 2) the V5H P+Z +2 fuel tube is longer than the V5H P+ fuel tube, and

3) the variable pitch plenum spring is longer to accommodate the increased rod

length. The V5H P+Z +2 fuel rods will contain, as in the V5H P+ design, enriched uranium dioxide fuel pellets. Schematics of the V5H P+Z +2 fuel rods are shown in Figure 4.2-3d.

Cycle 13 fresh fuel, RFA Z +2 design, utilizes the same fuel rod design as the V5H P+Z+2 design. Cycle 14 fresh fuel, RFA-2 Z +2 design, utilizes the same fuel rod design as the V5H P+Z+2 and RFA Z+2 design. Cycle 16 fresh fuel incorporates the use of a fuel rod oxide coating on the RFA

Z+2 design. The fuel rod has a very thin oxide coating at the bottom end of the fuel rod. The extra layer of oxide coating provides additional debris

induced rod fretting wear protection.

4.2-12 Rev. 29 WOLF CREEK 4.2.2.1.2 Integral Fuel Burnable Absorber Fuel Rods The Integral Fuel Burnable Absorber (IFBA) fuel rod design for the V5H, V5H P+, V5H P+Z+2 , RFA Z+2 and RFA-2 Z +2 designs are identical to the Non-IFBA fuel rod design for the V5H, V5H P+, V5H P+Z +2 , RFA Z+2 and RFA-2 Z +2 designs, respectively, with the following exceptions: a) Some of the fuel pellets are coated with a thin layer of zirconium diboride (ZrB 2) on the pellet cylindrical surface. b) The helium back fill pressure for the IFBA fuel rod is lower than the Non-IFBA fuel rod.

The zirconium diboride coating is referred to as the Integral Fuel Burnable

Absorber design or IFBA. Other than the zirconium diboride coating, the fuel

pellets for an IFBA rod are identical to the enriched uranium dioxide pellets

described for the Non-IFBA fuel rod. The IFBA pellets are placed in the

central portion of the fuel pellet stack (up to 134 inches). The lower back

fill pressure for the IFBA rod offsets the increased rod pressure at end of

life due to the production and release of helium from the zirconium diboride

coating on the IFBA fuel pellets.

The number and pattern of IFBA rods loaded within an assembly may vary

depending on the specific application. The IFBA design provides an alternate

means of reactivity control as opposed to the discrete burnable absorber

designs discussed in Section 4.2.2.3. An evaluation and test program for the

IFBA design features is given in section 2.5 of Reference 19. Cycle 9 fresh

fuel incorporated the use of the IFBA rod design.

4.2.2.2 Fuel Assembly Structure

The fuel assembly structure consists of a bottom nozzle, thimble screws, top

nozzle, guide thimbles, inserts, lock tubes, and grids, as shown in Figure 4.2-

2, Figure 4.2-2a, Figure 4.2-2b, Figure 4.2-2c, and Figure 4.2-2d.

4.2.2.2.1 Bottom Nozzle

The bottom nozzle serves as the bottom structural element of the fuel assembly

and distributes the coolant flow to the assembly. The bottom nozzle is

fabricated from Type 304 stainless steel. The standard bottom nozzle design

consists of a perforated plate and four angle legs with bearing plates, as

shown in Figure 4.2-2. The plate prevents accidental downward ejection of the

fuel rods from the fuel assembly. The bottom nozzle is fastened to the fuel

assembly guide tubes by locked thimble screws which penetrate through the

nozzle and mate with a threaded plug in each guide tube.

The Cycle 5 fresh fuel design incorporated the Debris Filter Bottom Nozzle (DFBN) to reduce the possibility of fuel rod damage due to debris-induced

fretting. The relatively large flow holes in a conventional nozzle are

replaced with a new pattern of smaller flow holes. The holes are sized to

minimize passage of debris particles large enough to cause damage while

providing sufficient flow area, comparable pressure drop, and continued

structural integrity of the nozzle. The Cycle 6 fresh fuel added a reinforcing

skirt to the DFBN design, as shown in Figure 4.2-2a, Figure 4.2-2b, Figure 4.2-

2c, and Figure 4.2-2d. the reinforcing skirt is located between the angle legs

around the perimeter of the bottom nozzle and contains five holes on each face

to allow lateral fluid flow. The legs and skirt form a plenum for the inlet

coolant flow to the fuel assembly and enhance reliability during postulated

adverse handling conditions while refueling. Tests to measure pressure drop

and demonstrate structural integrity verified that the 304 stainless steel DFBN

is totally compatible with the current design.

The Cycle 21 fresh fuel incorporates a Standardized Debris Filter Bottom Nozzle (SDFBN). The SDFBN has eliminated the side skirt communication flow holes as a means of improving the debris mitigation performance of the bottom nozzle. 4.2-13 Rev. 29 WOLF CREEK This nozzle has been evaluated and meets all of the applicable mechanical design criteria. In addition, there is no adverse effect on the thermal hydraulic performance of the SDFBN either with respect to the pressure drop or with respect to Departure from Nucleate Boiling (DNB). Coolant flows 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 are transmitted through the bottom nozzle to the lower core plate. Indexing and positioning of the fuel assembly are provided by alignment holes in two diagonally opposite bearing plates which mate with locating pins in the lower core plate. Lateral loads on the fuel assembly are transmitted to the lower core plate through the locating pins. 4.2.2.2.2 Top Nozzle

The top nozzle functions as the upper structural element of the fuel assembly

and provides a partial protective housing for the rod cluster control assembly

or other components that are installed in the guide thimble tubes. The top

nozzle consists of an adapter plate, enclosure, top plate, and pads. The top

nozzle assembly consists of holddown springs mounted on the top nozzle as shown

in Figure 4.2-2, Figure 4.2-2a, Figure 4.2-2b, Figure 4.2-2c, Figure 4.2-2d and

Figure 4.2-2e. The springs and spring screws are made of Inconel-718 and

Inconel-600 respectively, whereas other components are made of Type 304

stainless steel.

The standard top nozzle adapter plate is provided with round penetrations and

semicircular ended slots 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 at their upper ends and mechanically attached to the thimble

tubes at the lower end. 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 box-like structure which sets the distance between the adapter plate and

the top plate. The top nozzle has a large square hole in the center to permit

access to the thimble tubes for the control rods and provide a partial

protective housing for the control rod spiders. Holddown springs are mounted

on the standard top nozzle and are retained by spring screws and clamps located

at two diagonally opposite corners. On the other two corners, integral pads

are positioned, which contain alignment holes for locating the upper end of the

fuel assembly. Figure 4.2-6 shows the top nozzle attachment to the thimble

tubes for the standard top nozzle assembly.

Cycle 5 fresh fuel incorporated the reconstitutable top nozzle (RTN) design.

The RTN design for the V5H and V5H P+ fuel assembly differs from the standard

top nozzle design in two ways: a groove is provided in each thimble

throughhole in the nozzle adapter plate to facilitate attachment and removal;

and the nozzle plate thickness is reduced to provide additional axial space for

fuel rod growth.

Cycle 12 fresh fuel incorporates a cast RTN design and shot-peened Inconel-600

spring screws into the top nozzle design. The top nozzle enclosure, top plate

and pads are cast as a single unit and joined with the adapter plate to make

the cast RTN.

Cycle 13 fresh fuel incorporates shot-peened Inconel-718 spring screws into the

cast RTN top nozzle design.

Cycle 14 fresh fuel, RFA-2 Z +2 , utilizes the same top nozzle design as the Cycle 13 fresh fuel, RFA Z +2. In the RTN design, a stainless steel nozzle insert is mechanically connected to

the top nozzle adapter plate by means of a preformed circumferential bulge near 4.2-14 Rev. 29 WOLF CREEK the top of the insert. The insert engages a mating groove in the wall of the adapter plate thimble tube throughhole. The insert has four equally spaced axial slots which allow the insert to deflect inwardly at the elevation of the bulge, thus permitting the installation or removal of the top nozzle. The insert bulge is positively held in the adapter plate mating groove by placing a lock tube with a uniform ID identical to that of the thimble tube into the insert. The inserts are mechanically attached to the thimble tubes at the lower end with three bulge joints. Figure 4.2-6a shows the top nozzle attachment to the thimble tubes for the RTN assembly. Cycle 16 fresh fuel incorporates the Westinghouse Integral Nozzle (WIN) top nozzle. The WIN design differs from the RTN design in the attachment method for the hold down springs. The WIN top nozzle includes a modified top nozzle casting that includes the spring clamps. The springs are located with pins that are welded in place but do not react to the spring force. The WIN top nozzle design eliminates the potential for the fracture of the hold down spring screws by the replacing the spring screws with the spring pins. The

modification increases the fuel assembly integrity and eliminates the potential

for loose parts from fractured spring screws entering the RCS during normal

operations or during fuel movement during refueling outages.

To remove the top nozzle, a tool is first inserted through the lock tube and

expanded radially to engage the bottom edge of the lock tube. An axial force

is then exerted on the tool which overrides the local lock tube deformations

and withdraws the lock tube from the insert. After the lock tubes have been

withdrawn, the top nozzle is removed by raising it off the upper slotted ends

of the nozzle inserts which deflect inwardly under the axial lift load. With

the top nozzle removed, direct access is provided for fuel rod examination or

replacement. Reconstitution is completed by the remounting of the top nozzle

and the insertion of the lock tubes. The design bases and evaluation of the

RTN are given in Section 2.3.2 of Reference 19.

4.2.2.2.3 Guide Thimble and Instrument Tubes

The guide thimbles are structural members which also provide channels for the

neutron absorber rods, burnable absorber rods, neutron source rods, or thimble plug devices. Each thimble is fabricated from Zircaloy-4 or Zirlo tubing having two different diameters.

The Cycle 6 fresh fuel incorporation of the Zircaloy-4 mid grids required a

concurrent incorporation of the VANTAGE 5 (V5) reduced diameter thimble tubes.

The VANTAGE 5 guide thimbles are also referred to as the VANTAGE 5H (V5H) guide

thimble tubes. With the exception of a reduction in the guide thimble diameter

above the dashpot, the V5H and V5H P+ guide thimbles are identical to those in

the LOPAR design. A 0.008 inch reduction to the guide thimble OD and ID is required due to the thicker Zircaloy/Zirlo grid straps. The V5H and V5H P+ guide thimble tube ID provides an adequate nominal diametral clearance of 0.061 inch for the control rods. The scram time to the dashpot for accident analyses

is 2.7 seconds. The reduced V5H and V5H P+ thimble tube ID provides sufficient

diametral clearance for burnable absorber rods, source rods, and any dually compatible thimble plugs. Cycle 10 fresh fuel incorporated the use of Zirlo material for the guide thimble and instrumentation tubes. The V5H P+ assembly design uses guide thimble and instrument tubes which are slightly shorter than

those used in the V5H assembly design. Cycle 12 fresh fuel incorporated

slightly longer guide thimble and instrumentation tubes as part of the V5H

P+Z+2 fuel assembly design (same length as the V5H design).

Cycle 13 fresh fuel incorporated thicker guide thimble and instrumentation

tubes with a larger outer diameter as part of the RFA Z +2 fuel assembly design. The RFA Z+2 guide thimble tube wall thickness is increased approximately 25% to improve stiffness and address incomplete rod insertion (IRI) considerations.

The major and minor (dashpot) OD of the guide thimble tube are increased while

maintaining the same major and minor (dashpot) ID to accommodate the increased

4.2-15 Rev. 29 WOLF CREEK wall thickness. There is no change to the dashpot flow hole diameters or the dashpot transition elevation. Cycle 14 fresh fuel, RFA-2 Z +2 fuel assembly design, utilizes the same guide thimble tube design included in the RFA Z +2 fuel assembly design. The guide thimble diameter at the top section provides the annular area necessary to permit rapid control rod insertion during a reactor trip. The lower portion of the guide thimble reduces to a smaller diameter to produce a dashpot action near the end of the control rod travel during trip operation. The dashpot is provided with a calibrated flow port to decelerate the rod at the end of the travel. The top end of the guide thimble is fastened to an insert (RTN) or top Inconel grid sleeve (Standard Top Nozzle) by three expansion swages. When attaching to a RTN, the insert fits into and is locked into the top nozzle adapter plate using a lock tube. When attaching to a standard top nozzle, the top Inconel grid sleeve is welded to the top nozzle adapter plate. The lower end of the guide thimble is fitted with an end plug

which is then fastened to the bottom nozzle by a crimp-locked thimble screw.

Fuel rod support grids are fastened to the guide thimble assemblies to create an integrated structure. Attachment of the Inconel and Zircaloy or Zirlo grids to the Zircaloy or Zirlo thimbles is performed using the mechanical fastening technique as depicted in Figures 4.2-4 and 4.2-5 except for the bottom grid which is retained by clamping between the thimble end plug and the

bottom nozzle.

An expanding tool is inserted into the inner diameter of the Zircaloy or Zirlo thimble tube at the elevation of the grid sleeves that have been previously attached into the grid assembly. The four-lobed tool forces the thimble and sleeve outward to a predetermined diameter, thus joining the

twocomponents.

When attaching to a standard top nozzle, the top inconel grid sleeve and

thimble tube are joined together using three bulge joint mechanical attachments

as shown in Figure 4.2-6. The sleeve is then welded to the top nozzle adapter

plate. When attaching to a RTN, the thimble tube is joined together with the

top nozzle insert and top Inconel grid sleeve using three bulge joint

mechanical attachments as shown in Figure 4.2-6a. This bulge joint connection

was mechanically tested and found to meet all applicable design criteria.

The intermediate mixing vane Zircaloy grids, incorporated with Cycle 7 fresh

fuel, employ a single bulge connection to the sleeve and thimble as compared to

a three bulge connection used in the top Inconel grid (Figure 4.2-5).

Mechanical testing of this bulge joint connection was also found to be acceptable. Cycle 10 fresh fuel incorporated the use of Zirlo material for the intermediate mixing vane grids.

The bottom grid assembly is joined to the assembly by crimp lock screw, as

shown in Figure 4.2-7. 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 tube of each fuel assembly is constrained by

seating in a counterbore in the bottom nozzle at its lower end and is expanded

at the top and mid grids in the same manner as the previously discussed

expansion of the guide thimbles to the grids. This tube has a constant

diameter and guides the incore neutron detectors.

The V5H, V5H P+, and V5H P+Z +2 instrumentation tube designs have a 0.008 inch diametral decrease compared to the LOPAR assembly instrumentation tube. This 4.2-16 Rev. 29 WOLF CREEK decrease still allows sufficient diametral clearance for the incore neutron detector (max. OD = 0.397 inch) to traverse the tube without binding. The RFA Z+2 and RFA-2 Z +2 instrumentation tube design includes an increased wall thickness consistant with the RFA Z +2 and RFA-2 Z +2 guide thimble tubes. The OD of the tube is increased while maintaining the same ID to accommodate the increased wall thickness. 4.2.2.2.4 Grid Assemblies The fuel rods, as shown in Figure 4.2-2, Figure 4.2-2a, Figure 4.2-2b, Figure 4.2-2c, and Figure 4.2-2d, are supported at intervals along their length by grid assemblies which maintain the lateral spacing between the rods. Each fuel rod is supported within each grid by the combination of support dimples and springs. The grid assembly consists of individual slotted straps assembled and interlocked into an "egg-crate" arrangement with the straps permanently joined at their points of intersection. The top and bottom Inconel (non-mixing vane) grids of the LOPAR, V5H, V5H P+, V5H P+Z+2 , RFA Z+2 and RFA-2 Z +2 assemblies are nearly identical in design. The only differences are: 1) V5H, V5H P+, V5H P+Z +2 , RFA Z+2 and RFA-2 Z +2 top and bottom grids have a snag-resistant design which minimizes assembly interactions

during core loading/unloading, 2) V5H, V5H P+, V5H P+Z +2 , RFA Z+2 and RFA-2 Z +2 top and bottom grids have dimples which are rotated 90 degrees to minimize fuel

rod fretting and dimple cocking, 3) V5H, V5H P+, V5H P+Z +2 and RFA Z +2 top and bottom grid heights have been increased to 1.522 inches, 4) the V5H, V5H P+, V5H P+Z+2 , RFA Z+2 and RFA-2 Z +2 top grid spring force has been reduced to minimize rod bow, and 5) the V5H, V5H P+, V5H P+Z +2 , RFA Z+2 and RFA-2 Z +2 top grid uses 304L stainless steel sleeves.

Cycle 4 fresh fuel incorporated the snag-resistant top and bottom grid design

mentioned above into the fuel design for Wolf Creek.

The LOPAR fuel design utilizes six intermediate (mixing vane) grids made of

Inconel. The snag-resistant design described for the top and bottom grid was

incorporated into the six intermediate grids with Cycle 4 fresh fuel. Cycle 6

fresh fuel incorporated intermediate (mixing vane) grids made of Zircaloy

material rather than Inconel. Cycle 10 fresh fuel incorporated six intermediate (mixing vane) grids made of Zirlo rather than Zircaloy. These Zircaloy and Zirlo grids (known as the V5H Zircaloy grid and V5H P+ Zirlo grid) are designed to give the same pressure drop as the Inconel grid.

Relative to the Inconel grid, the V5H Zircaloy and V5H P+ Zirlo grid strap thickness and strap height are increased for structural performance. In addition to the snag-resistant design noted above, the upstream strap edges of the V5H Zircaloy grid and V5H P+ Zirlo grid are chamfered and a diagonal grid spring is employed to reduce pressure drop. The V5H Zircaloy grids and V5H P+

Zirlo grids incorporate the same grid cell support configuration as the Inconel grids (six support locations per cell: four dimples, and two springs).

The Zircaloy and Zirlo grid interlocking strap joints and grid/sleeve joints are fabricated by laser welding, whereas the Inconel grid joints are brazed. The V5H Zircaloy, V5H P+ Zirlo, RFA Zirlo and RFA-2 Zirlo TM grid have superior dynamic structural performance relative to the Inconel grid. Structural testing was performed and analyses have shown the V5H Zircaloy grid, V5H P+ Zirlo, RFA Zirlo and RFA-2 Zirlo TM seismic/LOCA grid load margin is superior to that of the Inconel grid.

The Intermediate Flow Mixer (IFM) grid in the VANTAGE 5H assembly is an

adaptation of the existing VANTAGE 5 IFM grid design to a 0.374 inch OD

standard fuel rod. As shown in Figure 4.2-2a, Figure 4.2-2b, and Figure 4.2-

2c, IFMs are located in the three uppermost spans between the mid-grids but are

not intended to be structural members. The IFM grid envelope is slightly

smaller than the mid grid. Each IFM grid cell provides four (4) point fuel rod 4.2-17 Rev. 29 WOLF CREEK support. The simplified cell arrangement allows the IFM to accomplish its flow mixing objective with minimal pressure drop. Cycle 7 fresh fuel incorporated the Zircaloy Intermediate Flow Mixer grid. Cycle 10 fresh fuel incorporated the use of Zirlo material in the manufacture of the IFM grids. The Protective Bottom Grid (PBG) is a partial height grid similar in configuration to the IFM Grid, but fabricated of Inconel without mixing vanes. The PBG is positioned directly above the bottom nozzle. As shown in Figures 4.2-2b, 4.2-2c, 4.2-2d, 4.2-3b, 4.2-3c, and 4.2-3d, the fuel rods are positioned close to the bottom nozzle and are modified with a slightly longer bottom end plug. The PBG provides added protection against debris induced fretting by trapping debris below this grid where it can wear against the solid end plug. In addition, the PBG provides improved resistance to grid-rod fretting by means of additional support at the bottom of the fuel rod. Cycle 8 fresh fuel incorporated the protective bottom grid. Cycle 13 fresh fuel incorporated the RFA Z +2 fuel assembly design. RFA Z +2 changes made to the mid-grid include a modified vane pattern (which is now

symmetrical), longer vane geometry, modified spring and dimple geometry, a

narrower spring window cut-out, a longer intersect slot length, opposite hand

spring and the incorporation of the anti-snag outer grid strap design. IFM

modifications include a symmetric vane pattern, longer vane geometry, and a

change to the dimple profile. The Inconel top, bottom, and protective grids

are not changed in the RFA Z +2 design except for new insert tubing for the bottom and protective grids to accommodate the increase in thimble and

instrument tube diameters.

Cycle 14 fresh fuel is the RFA-2 Z +2 fuel assembly design. The RFA-2 Z +2 design changes the mid-grid to include a modified spring and dimple geometry that

increases the line-contact length of the rod-spring and rod-dimple interface.

The RFA-2 Z +2 IFM grid design is not changed relative to the RFA Z +2 IFM grid design. The RFA-2 Z +2 Inconel top, bottom and protective grid designs are not changed relative to the RFA Z +2 Inconel top, bottom and protective grid designs.

The Cycle 21 fresh fuel implemented a combination bottom grid and Robust Protective Grid (RPG). Westinghouse has developed the RPG as a result of observed failures in the field as noted in Post Irradiation Exams (PIE) performed at several different plants. It was determined 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 implemented design changes such as increasing the maximum nominal height of the grid, increasing te ligament length and the radii of the ligament cutouts, and the use of four additional spacers or inserts to help strengthen the grid. 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. The design changes incorporated into the RPG design helped address the issues of fatigue failures and failures due to SCC. It was determined that the above changes do not impact the thermal hydraulic performance of the RPG as there is no change to the loss coefficient. In addition, the RPG retains the original protective grid function as a debris mitigation feature.

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.

4.2.2.2.5 Fuel Assemblies - LOPAR, V5H, and V5H P+, V5H P+Z +2 , RFA Z+2 and RFA-2 Z+2 designs

4.2-18 Rev. 29 WOLF CREEK The initial fuel assembly design used at Wolf Creek was the Westinghouse 17x17 low-parasitic (LOPAR) fuel design. The original LOPAR fuel assembly design is shown in Figure 4.2-1, Figure 4.2-2, and Figure 4.2-3. Westinghouse developed several fuel performance enchancing features which were added to the LOPAR design over a period of several reloads. The major enhancements included: Chamfered pellet design High burnup top and bottom end plug designs Anti-snag grid design Debris Filter Bottom Nozzle (DFBN) design Reconstitutable Top Nozzle (RTN) design These features were gradually added to the base LOPAR design for Cycle 2 through 5 fresh fuel. The actual point that the particular enhancement was incorporated is specified in the appropriate section of the USAR. The Westinghouse VANTAGE 5H fuel design is a variation of the LOPAR design that includes all of the fuel performance enhancements listed above along with the following: VANTAGE 5H (V5H) Zircaloy-4 Mid Grid design VANTAGE 5 Guide Tube design Cycle 6 fresh fuel incorporated the V5H Zircaloy mid grids and the V5 guide tube designs. This marked the point at which the fuel design for Wolf Creek became the Westinghouse VANTAGE 5H fuel design. The V5H fuel assembly design is shown in Figures 4.2-1a, 4.2-2a, 4.2-2b, 4.2-3a, and 4.2-3b. Westinghouse has continued to developed fuel performance enhancing features which were added to the base V5H design over a period of several reloads. The major enhancements include: Zircaloy-4 Intermediate Flow Mixer (IFM) grid design Inconel Protective Bottom Grid (PBG) design Performance+ Extended Bottom End Plug design Performance+ Top End Plug design Variable Pitch Plenum Spring design Integral Fuel Burnable Absorber (IFBA) design These features were gradually added to the base V5H design for Cycle 7 through

9 fresh fuel. The actual point that the particular enhancement was

incorporated is specified in the appropriate section of the USAR.

Westinghouse VANTAGE 5H with Performance+ features (V5H P+) fuel design is a

variation of the V5H design that includes all of the fuel performance

enhancements listed above along with the following:

Zirlo Clad fuel rod design Zirlo guide thimble and instrumentation tube design Zirlo mid grid design Zirlo IFM grid design ZirloFully enriched annular axial blankets Cycle 10 fresh fuel incorporated the performance enhancement features listed

above. This marked the point at which the fuel design became the Westinghouse

VANTAGE 5H with Performance+ features (V5H P+) fuel design. The V5H P+ fuel

assembly design is shown in Figures 4.2-1a, 4.2-2c and 4.2-3c.

Westinghouse VANTAGE 5H with Performance+ features, Zirlo +2 (V5H P+Z+2) fuel design is a variation of the V5H P+ design that includes all of the fuel

performance enhancements listed above along with the following:

4.2-19 Rev. 29 WOLF CREEK Low pressure fuel rod design Cast Reconstitutable Top Nozzle design Shot-peened spring screw design To implement the low rod internal pressure fuel rod design, the following changes were required to the fuel rod and skeleton designs: Performance + Top End Plug design replaced by a shorter top end plug (with no external gripper) design Extended length Zirlo TM fuel rod tube design Extended length Variable Pitch Plenum Spring design Extended length Zirlo TM guide thimble tubes and instrument tubes Cycle 12 fresh fuel incorporated the performance enhancement features listed above. This marked the point at which the fuel design became the Westinghouse VANTAGE 5H with Performance+ features, Zirlo +2 (V5H P+Z+2) fuel design. The V5H P+Z+2 fuel assembly design is shown in Figures 4.2-1a, 4.2-2d and 4.2-3d. Westinghouse Robust Fuel Assembly Zirlo +2 (RFA Z+2) fuel design is a variation of the, V5H P+Z +2 design that includes the fuel performance features of the V5H P+Z+2 design along with the following: Shot-peened Inconel-718 spring screw design, ZIRLO TM thicker thimble and instrument tube design (0.020 in. wall vs. 0.016 in.), Modified Zirlo TM Low Pressure Drop (LPD) structural mid-grid design, Modified Zirlo TM Intermediate Flow Mixing (IFM) grid design. Cycle 13 fresh fuel incorporated the performance enhancement features listed

above. This marked the point at which the fuel design became the Westinghouse

Robust Fuel Assembly Zirlo +2 (RFA Z+2) design. The RFA Z +2 fuel assembly design is shown in Figures 4.2-2d and 4.2-3d.

The Westinghouse second-generation Robust Fuel Assembly Zirlo +2 (RFA-2 Z+2) fuel design is a variation of the RFA Z +2 design that includes the fuel performance features of the RFA Z +2 design along with the following: Modified Zirlo Low Pressure Drop (LPD) structural mid-grid design with increased spring and dimple contact area (RFA-2 mid-grid).

There is no change to the fuel assembly length, envelope or fuel rod design

relative to the RFA Z +2 design. Cycle 14 fresh fuel incorporated the performance enhancement features listed

above. This marked the point at which the fuel design became the Westinghouse

second-generation Robust Fuel Assembly Zirlo +2 (RFA-2 Z+2) design. The RFA-2 Z +2 fuel assembly design is shown in Figures 4.2-2d and 4.2-3d.

Cycle 16 fresh fuel incorporated the performance enhancement features of the

WIN top nozzle and fuel rod oxide coating. The fuel design continues to be the

Westinghouse second-generation Robust Fuel Assembly Zirlo +2 (RFA-2 Z+2) design. The RFA-2 Z +2 fuel assembly design with the WIN top nozzle is shown in Figures 4.2-2e and 4.2-3d.

The Cycle 21 fresh fuel incorporated a Standardized Debris Filter Bottom Nozzle (SDFBN) and a combination bottom grid and Robust Protective Grid (RPG). The fuel design continues to be the Westinghouse second-generation Robust Fuel Assembly Zirlo+2 (RFA-2 Z+2) design. The RFA-2 Z+2 fuel assembly design with the SDFBN and the RPG is shown in Figure 4.2-2f.

Table 4.3-1 provides a comparison of the LOPAR, V5H, V5H P+, V5H P+Z +2 , RFA Z+2 and RFA-2 Z +2 fuel assembly design parameters. 4.2-20 Rev. 29 WOLF CREEK 4.2.2.3 Incore Control Components 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:

a. Fuel depletion and fission product buildup

b. Cold to hot, zero power reactivity change

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

d. Burnable absorber depletion The chemical and volume control system is discussed in Chapter 9.0.

The rod cluster control assemblies provide reactivity control for: a. Shutdown b. Reactivity changes resulting from coolant temperature changes in the power range c. Reactivity changes associated with the power coefficient of reactivity d. Reactivity changes resulting from void formation It is desirable to have a negative moderator temperature coefficient at power levels exceeding 70% rated thermal power (RTP) 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. Since soluble boron alone is insufficient to ensure a negative moderator coefficient, burnable absorber assemblies and/or IFBAs are also used. Burnable absorbers such as WABAs and IFBAs are used to achieve a better power peaking control and a flatter power distribution. Although a negative moderator coefficient is desirable, it is acceptable and in

some cases essential to have the coefficient be slightly positive in an attempt

to extend cycle length. Current WCGS reload cycles are designed to have a small positive moderator temperature coefficient (<3 pcm/ F) at low thermal power (<30% RTP) during the first 25% of the cycle. The addition of excess reactivity to extend cycle length necessitates a greater amount of boric acid, which results in an increase of the moderator temperature coefficient.

The rod cluster control assemblies and their control rod drive mechanisms are

the only moving parts in the reactor. Figure 4.2-8 illustrates the rod cluster

control and control rod drive mechanism assembly, in addition to the

arrangement of these components in the reactor, relative to the interfacing

fuel assembly and guide tubes. In the following paragraphs, each reactivity

control component is described in detail. The control rod drive mechanism

assembly is described in Section 3.9(N).4.

The neutron source assemblies provide a means of monitoring the core during

periods of low neutron level. The thimble plug may be used to limit bypass

flow through those fuel assembly thimbles, which do not contain control rods, burnable absorber rods, or neutron source rods.

4.2.2.3.1 Rod Cluster Control Assembly

The rod cluster control assemblies are divided into two categories: control

and shutdown. The control groups compensate for reactivity changes associated

with variations in operating conditions of the reactor, i.e., power and

temperature variations. Two nuclear design criteria have been employed for

selection of the control group. First, the total reactivity worth must be

adequate to meet the nuclear requirements of the reactor. Second in view of

4.2-21 Rev. 29 WOLF CREEK 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 is met. The control and shutdown group provides adequate shutdown margin. A rod cluster control assembly is composed of 24 neutron absorber rods fastened at the top end to a common spider assembly, as illustrated in Figure 4.2-9. The absorber material used in the control rods is a solid hafnium or Silver-Indium-Cadmium (Ag-In-Cd) bar which is essentially "black" to thermal neutrons and has sufficient additional resonance absorption to significantly increase its worth. The absorber material is sealed in cold worked stainless steel tubes (see Figure 4.2-10). Sufficient diametral and end clearances are provided to accommodate relative thermal expansions. The bottom plugs are 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 absorber rod end plugs are Type 308 stainless steel. The design stresses used for the Type 308 material are the same as those defined in the ASME Code, Section III, for Type 304 stainless steel. At room temperature, the yield and ultimate stresses per ASTM 580 are the same for the two alloys. In view of the similarity of the alloy composition, the temperature dependence of strength for the two materials is also assumed to be the same. The allowable stresses used as a function of temperature are listed in Table 1-1.2 of Section III of the ASME Code. The fatigue strength for the Type 308 material is based on the S-N curve for austenitic stainless steels in Figure 1-9.2 of Section III. The spider assembly is 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. Two coil springs inside the spider

body absorbs the impact energy at the end of a trip insertion. The radial

vanes are joined to the hub by tack welding and brazing, and the fingers are

joined to the vanes by brazing. A centerpost, which holds the spring and its

retainer, is threaded into the hub within the skirt and welded to prevent

loosening in service. All components of the spider assembly are made from

Types 304 and 308 stainless steel except for the retainer, which is of 17-4 PH

material, and the springs, which are Inconel-718 alloy.

The absorber rods are fastened securely to the spider. The rods are first

threaded into the spider fingers and then pinned to maintain joint tightness, after which the pins are welded in place. The end plug below the pin position

is designed with a reduced section to permit flexing of the rods to correct for

small misalignments.

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.

4.2.2.3.2 Burnable Absorber Assembly

(Standard Borosilicate Glass and WABA)

Each burnable absorber assembly consists of burnable absorber rods attached to

a holddown assembly. A burnable absorber assembly is shown in Figure 4.2-11

for the WABA rod and in Figure 4.2-11a for the borosilicate glass absorber rod.

When needed for nuclear considerations, burnable absorber assemblies may be

inserted into selected thimbles within fuel assemblies.

4.2-22 Rev. 29 WOLF CREEK The discrete burnable absorber rods are the wet annular burnable absorber (WABA) rod design and the borosilicate glass rod design. Integral Fuel Burnable Absorber (IFBA) rods, described in Section 4.2.2.1.2, are an alternative burnable absorber that may be used. The borosilicate glass burnable absorber design was used in Cycles 1 and 2 and the WABA design was introduced in Cycle 3. Cycle 9 fresh fuel incorporated the IFBA design. The WABA rod design consists of annular pellets of aluminum oxide-boron carbide (Al 2 O 3-B 4 C) burnable absorber material contained within two concentric Zircaloy tubes. These Zircaloy tubes, which form the inner and outer clad for the annular burnable absorber rod, are plugged, pressurized with helium, and seal welded at the ends to encapsulate the annular stack of absorber material. A Zircaloy spacer tube is placed at the bottom of the pellet stack to position the absorber stack within the WABA rod, and a C-shape Zircaloy spring clip is placed on top of the absorber stack to keep it in position and accommodate absorber stack growth. An annular plenum is provided within the rod to accommodate the helium gas released from the absorber material during boron depletion. The reactor coolant flows inside the inner tubing and outside the outer tubing of the annular rod. A typical WABA rod is shown in a longitudinal cross-section in Figure 4.2-12. The borosilicate glass 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. 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 typical borosilicate glass burnable absorber rod is shown in longitudinal and transverse cross-sections in Figure 4.2-12a. The absorber rods in each burnable absorber assembly are grouped and attached together at the top end of the rods to a hold-down assembly by a flat

perforated retaining plate which fits within the fuel assembly top nozzle and

rests on the adapter plate. The retaining plate and the absorber rods are 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 ensures that the absorber rods

cannot be ejected from the core by flow forces. Each rod is permanently

attached to the baseplate by a nut which is crimped or lock-welded into place. The cladding of the WABA rods is Zircaloy. The cladding of the borosilicate

glass rods is slightly cold worked Type 304 stainless steel. All other

structural materials in the assembly are Type 304 or 308 stainless steel except

for the springs, which are Inconel-718. The aluminum oxide-boron carbide

pellets or the borosilicate glass tubes provide sufficient boron content to

meet the criteria discussed in Section 4.3.1. 4.2.2.3.3 Neutron Source Assembly The purpose of the neutron source assembly is to provide base neutron level to

ensure that the neutron detectors are operational and responding to core

multiplication neutrons. For the first core, a neutron source is placed in the

reactor to provide a positive neutron count of at least 2 counts per second on

the source range detectors attributable to core neutrons. The detectors, called source range detectors, are used primarily when the core is subcritical

and during special subcritical modes of operations.

The source assembly permits detection of changes in the core multiplication

factor during core loading and approach to criticality. This can be done since

the multiplication factor is related to an inverse function of the detector

count rate. Changes in the multiplication factor can be detected during

addition of fuel assemblies while loading the core, changes in control rod

positions, and changes in boron concentration. 4.2-23 Rev. 29 WOLF CREEK The primary source rod, containing a radioactive material, spontaneously emits neutrons during initial core loading, reactor startup, and initial operation of the first core. 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 is activated during reactor operation. The activation results in the subsequent release of neutrons. Four source assemblies were installed in the initial reactor core: two primary source assemblies and two secondary source assemblies. Subsequent cycles (2-

10) utilize only the secondary source assemblies. Each primary source assembly contains one primary source rod and a number of burnable absorber rods. Each secondary source assembly contains four secondary source rods and a number of thimble plugs. A secondary source assembly is shown in Figure 4.2-14 and a primary source assembly is shown in Figure 4.2-14a.

"Double encapsulated" secondary source assemblies are available for use beginning with Cycle 11. Each of the double encapsulated secondary source assemblies contains six double encapsulated secondary source rods and a number of thimble plugs. A double encapsulated secondary source assembly is shown in Figure 4.2-14b.

Neutron source assemblies are positioned at opposite sides of the core. The source assemblies are inserted into the guide thimble tubes in fuel assemblies

at selected unrodded core locations. As shown in Figure 4.2-14 and Figure 4.2-

14b, the secondary source assembly contains a holddown assembly identical to

that of the burnable absorber assembly. The primary and secondary source rods

have the same cladding material as the absorber rods. The secondary source

rods contain Sb-Be pellets stacked to a height of approximately 88 inches. A

secondary source rod assembly is shown in Figure 4.2-13. The double

encapsulated secondary source rods also contain Sb-Be pellets stacked to a

height of approximately 88 inches. A double encapsulated secondary source rod

assembly is shown in Figure 4.2-13a. The primary source rods contain capsules

of californium source material and alumina spacer to position the source

material within the cladding. The rods in each source assembly are permanently

fastened at the top end to a holddown assembly. The other structural members are constructed of Type 304 or Type 308 stainless

steel, except for the springs. The springs exposed to the reactor coolant are

Inconel-718. 4.2.2.3.4 Thimble Plug Device Thimble plug devices may be used 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. A typical thimble plug device is

shown in Figures 4.2-15 and 4.2-15a. The thimble plug devices consist of a flat baseplate with short rods suspended

from the bottom surface and a spring pack assembly. The 24 short rods, called

thimble plugs, project into the upper ends of the guide thimbles to reduce the

bypass flow. Each thimble plug is permanently attached to the baseplate by a nut which is

crimped or lock-welded to the threaded end of the plug. 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. When in the core, the thimble

plug devices interface with both the upper core plate and with the fuel

assembly top nozzles by resting on the adapter plate. The spring pack is

compressed by the upper core plate when the upper internals assembly is lowered

into place. All components in the thimble plug device, except for the springs, are

constructed from Type 304 or Type 308 stainless steel. The springs are

Inconel-718. 4.2-24 Rev. 29

WOLF CREEK 4.2.3 DESIGN EVALUATION The fuel assemblies, fuel rods, and incore control components are designed to satisfy the performance and safety criteria of the introduction to Section 4.2, the mechanical design bases of Section 4.2.1, and other interfacing nuclear and thermal-hydraulic design bases specified in Sections 4.3 and 4.4. Effects of Conditions II, III, IV or anticipated transients without trip on fuel integrity are presented in Chapter 15.0 or supporting topical reports. The initial step in fuel rod design evaluation for a region of fuel is to determine the limiting rod(s). Limiting rods are defined as those rod(s) whose predicted performance provides the minimum margin to each of the design criteria. For a number of design criteria, the limiting rod is the highest burnup rod of a fuel region. In other instances, it may be the maximum power or the minimum burnup rod. For the most part, no single rod is limiting with respect to all design criteria. After identifying the limiting rod(s), a worst-case performance analysis is performed which considers the effects of rod operating history, model

uncertainties, and dimensional variations. To verify adherence to the design

criteria, the evaluation considers the effects of postulated transient power

changes during operation consistent with Conditions I and II. These transient

power increases can affect both rod average and local power levels. Parameters

considered include rod internal pressure, fuel temperature, clad stress, and

clad strain. In fuel rod design analyses, these performance parameters provide

the basis for comparison between expected fuel rod behavior and the

corresponding design criteria limits.

Fuel rod and fuel assembly models used for the performance evaluations are

documented and maintained under an appropriate control system. Materials

properties used in the design evaluations are given in Reference 2.

4.2.3.1 Cladding

a. Vibration and wear

Fuel rod vibrations are flow induced. The effect of the vibration on the fuel assembly and individual fuel rods is minimal. The cyclic stress range associated with deflections of such small magnitude is insignificant and has no effect on the structural integrity of the fuel

rod.

The reaction force on the grid supports due to rod vibration motions is also small and is much less than the spring preload. No significant wear of the clad or grid supports is expected during the life of the fuel assembly.

Clad fretting and fuel vibration have been experimentally investigated, as shown in Reference 10. Hydraulic flow test results of the RFA-2 fuel assembly are discussed in Reference 26.

b. Fuel rod internal pressure and cladding stresses

A burnup dependent fission gas release model (References 18 and 27) is used to determine the internal gas pressures as a function of irradiation time. The plenum height of the fuel rod has been designed to ensure that the maximum internal pressure of the fuel rod will not exceed the value which would cause the fuel/clad diametral gap to increase and extensive DNB propagation during steady state

operation. 4.2-25 Rev. 29 WOLF CREEK 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 approximately 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 inside diameter and positive at the clad outside 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 due to low internal gas pressure, and the thermal stress is highest in the maximum power rod due to steep temperature

gradient.

Tensile stresses can occur once the clad has come into contact with the pellet. These stresses are induced by the fuel pellet swelling during irradiation. Swelling of the fuel pellet 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). The 1-percent strain criterion is extremely conservative for fuel-swelling driven clad strain because the strain rate associated with solid fission products swelling is very slow. A detailed discussion on fuel rod performance is given in Section 4.2.3.3.

c. Materials and chemical evaluation

Zircaloy-4 and Zirlo clad has a high corrosion resistance to the coolant, fuel, and fission products. As shown in Reference 1, there is pressurized water reactor operating experience on the capability of Zircaloy and Zirlo as a clad material. Controls on fuel fabrication specify maximum moisture levels to preclude clad hydriding.

Metallographic examination of irradiated commercial fuel rods has 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. 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.

d. Stress Corrosion

Stress corrosion cracking is another postulated phenomenon related to fuel/clad chemical interaction. Out-of-pile tests have shown that in the presence of high cladding tensile stresses, large concentrations of selected fission products (such as iodine) can chemically attack the Zircaloy and Zirlo tubing and can lead to eventual cladding cracking. Extensive post-irradiation examination has produced no in-pile evidence that this mechanism is operative in commercial fuel. 4.2-26 Rev. 29

WOLF CREEK e. Cycling and Fatigue A comprehensive review of the available strain fatigue models was conducted by Westinghouse as early as 1968. This review included the Langer-O'Donnell model (Reference 12), 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-O'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 Westinghouse testing program was subdivided into the following subprograms:

1. A rotating bend fatigue experiment on unirradiated Zircaloy-4 specimens at room temperature and at 725 F. Both hydrided and nonhydrided Zircaloy-4 cladding were tested.

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

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

The results of these test programs provided information on different

cladding conditions including the effects of irradiation, of hydrogen

levels and of temperature.

The design equations followed the concept for the fatigue design

criterion according to the ASME Boiler and Pressure Vessel Code, Section III.

It is recognized that a possible limitation to the satisfactory

behavior of the fuel rods in a reactor which is subjected to daily

load follow is the failure of the cladding by low cycle strain

fatigue. During their normal residence time in a reactor, the fuel

rods may be subjected to ~1000 cycles with typical changes in power

level from 50% to 100% of their steady-state values.

The assessment of the fatigue life of the fuel rod cladding is subject

to a considerable uncertainty due to the difficulty of evaluating the

strain range which results from the cyclic interaction of the fuel

pellets and cladding. This difficulty arises, for example, from such

highly unpredictable phenomena as pellet cracking, fragmentation, and

relocation. Nevertheless, since early 1968, this particular

phenomenon has been investigated analytically and experimentally.

Strain fatigue tests on irradiated and nonirradiated hydrided Zr-4

claddings were performed, which permitted a definition of a

conservative fatigue life limit and recommendation on a methodology to

treat the strain fatigue evaluation of the Westinghouse reference fuel

rod designs.

It is believed that the final proof of the adequacy of a given fuel

rod design to meet the load follow requirements can only come from

incore experiments performed on actual reactors. Experience in load

follow operation dates back to early 1970 with the load follow

operation of the Saxton reactor. Successful load follow operation has

been performed on reactor A (>400 load follow cycles) and reactor B

(>500 load follow cycles). In both cases, there was no significant

coolant activity increase that could be associated with the load

follow mode of operation.

4.2-27 Rev. 29 WOLF CREEK f. Rod bowing Reference 11 presents the NRC-approved model used for evaluation of fuel rod bowing. The effects of rod bowing on DNBR are described in Section 4.4.2.2.5. Also refer to item e in Section 4.2.3.3.

g. Consequences of power-coolant mismatch

This subject is discussed in Chapter 15.0.

h. Irradiation stability of the cladding

As shown in Reference 1, there is PWR operating experience on the capability of Zircaloy and Zirlo as a cladding material. Extensive experience with irradiated Zircaloy-4 is summarized in Reference 2, and Appendices A through E in Reference 20 for Zirlo. i. Creep collapse and creepdown

This subject and the associated irradiation stability of cladding have been evaluated, using the models described in Reference 19.

4.2.3.2 Fuel Materials Considerations

Sintered, high density uranium dioxide fuel reacts only slightly with the clad

at core operating temperatures and pressures. In the event of clad 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 clad 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

pressurized water reactors (Ref. 9) have shown that fuel pellets can densify

under irradiation to a density higher than the manufactured values. Fuel

densification and subsequent settling of the fuel pellets can result in 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-percent initial fuel density.

The evaluation of fuel densification effects and their consideration in fuel

design are described in References 18 and 27. The treatment of fuel swelling

and fission gas release are described in Reference 18.

The effects of waterlogging on fuel behavior are discussed in Section 4.2.3.3.

4.2.3.3 Fuel Rod Performance

In the calculation of the steady state performance of a nuclear fuel rod, the

following interacting factors must be considered.

a. Clad creep and elastic deflection

b. Pellet density changes, thermal expansion, gas release, and thermal properties as a function of temperature and

fuel burnup

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

4.2-28 Rev. 29 WOLF CREEK These effects are evaluated using fuel rod design models (References 18 and 27) which include appropriate models for time-dependent fuel densification. With the above 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 clad temperatures, and clad deflections are calculated. The fuel rod is divided into several

axial sections and radially into a number of annular zones. Fuel density

changes 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. It is

limited, however, by the design criteria for the rod internal pressure (see

Section 4.2.1.3). The gap conductance between the pellet surface and the clad 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 for each pellet annular zone, the

fractional fission gas release is assessed, using an empirical model derived from experimental data (References 18 and 27). 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 with a variety of published and proprietary data on fission gas release, fuel temperatures, and clad deflections (References 18 and 27). These data include variations in power, time, fuel density, and geometry. a. Fuel/cladding mechanical interaction

One factor in fuel element duty is potential mechanical interaction of fuel and clad. This fuel/clad interaction produces cyclic stresses and strains in the clad, and

these, in turn, consume clad fatigue life. The reduction of fuel/clad interaction is therefore a goal of design. The technology of using prepressurized fuel rods has been

developed to further this objective. The gap between the fuel and clad is initially sufficient to prevent hard contact between the two. However, during power operation a gradual compressive creep of the clad

onto the fuel pellet occurs due to the external pressure exerted on the rod by the coolant. Clad compressive

creep eventually results in fuel/clad contact. Once

fuel/clad contact occurs, changes in power level result

in changes in clad stresses and strains. By using prepressurized fuel rods to partially offset the effect of the coolant external pressure, the rate of clad creep

toward the surface of the fuel is reduced. Fuel rod

prepressurization delays the time at which fuel/clad

contact occurs and hence significantly reduces the extent

of cyclic stresses and strains experienced by the clad both before and after fuel/clad contact. These factors result in an increase in the fatigue life margin of the clad and lead to greater clad 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. 4.2-29 Rev. 18 WOLF CREEK A two-dimensional (r,) finite element model has been developed to investigate the effects of radial pellet cracks on stress concentrations in the clad. Stress

concentration, herein, is defined as the difference between the maximum clad stress in the -direction and the mean clad stress. The first case has the fuel and clad in mechanical equilibrium and, as a result, the

stress in the clad is close to zero. In subsequent

cases, the pellet power is increased in steps, and the resultant fuel thermal expansion imposes tensile stress

in the clad. In addition to uniform clad stresses, stress concentrations develop in the clad adjacent to radial cracks in the pellet. These radial cracks have a

tendency to open during a power increase but the

frictional forces between fuel and clad oppose the

opening of these cracks and result in localized increases

in clad stress. As the power is further increased, large

tensile stresses exceed the ultimate tensile strength of

UO 2 , and additional cracks in the fuel are created which limits the magnitude of the stress concentration in the clad. As part of the standard fuel rod design analysis, the maximum stress concentration evaluated from finite

element calculations is added to the volume-averaged

effective stress in the clad, as determined from one-

dimensional stress/strain calculations. The resultant clad stress is then compared to the temperature-dependent Zircaloy/Zirlo yield stress in order to assure that the stress/strain criteria are satisfied. 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. Such

increases are a consequence of fuel shuffling , reactor power escalation following extended reduced power

operation, and full-length control rod movement. In the

mechanical design model, lead rod burnup values are obtained using best estimate power histories, as

determined by core physics calculations. During burnup, 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 determine if the clad yield stress has been exceeded. This criterion states that an isotropic material in multiaxial stress will begin to yield plastically when the effective stress

exceeds the yield stress, as determined by an axial

tensile test. The yield stress correlation is that for

irradiated cladding, fuel/clad interaction occurs at high 4.2-30 Rev. 18 WOLF CREEK burnup. In applying this criterion, 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, ) 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 strain is necessary. Based upon high

strain rate burst and tensile test data on irradiated tubing, 1-percent strain was determined to be a conservative lower limit on irradiated clad deformation and was thus adopted as a design criterion. A comprehensive review of the available strain-fatigue models was conducted by Westinghouse as early as 1968.

This included the Langer-O'Donnell model (Ref. 12), 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-

O'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: S a =E N 4 f ln 100 100RA + S e where: S a = 1/2 E t = pseudo-stress amplitude which causes failure in N cycles (lb/in.

2) f t = total strain range (in./in.)

E = Young's Modulus (lb/in.

2) N f = number of cycles to failure RA = reduction in area at fracture in a uniaxial tensile test (%)

S e = endurance limit (lb/in.

2) 4.2-31 Rev. 16 WOLF CREEK Both RA and S e are empirical constants which depend on the type of material, the temperature, and irradiation.

The Westinghouse testing program was subdivided into the following subprograms:

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

F. Both hydrided and nonhydrided Zircaloy-4 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 Carolina-Virginia Tube Reactor and Yankee Core V

conducted at Battelle Memorial Institute. The results of these test programs provided information on different cladding conditions, including the effect of

irradiation, hydrogen level, and temperature. The design equations followed the concept for the fatigue design criterion according to the ASME Code, Section III. Namely, 1. The calculated pseudo-stress amplitude (S a) has to be multiplied by a factor of 2 in order to obtain the allowable number of cycles (N f) 2. The allowable cycles for a given S a is 5 percent of N f , maintaining 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: n N k fk l k 1 4.2-32 Rev. 16 WOLF CREEK where: n k = number of diurnal cycles of mode k N fk = number of allowable cycles 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 clad by low-cycle strain fatigue. During their normal

residence time in the reactor, the fuel rods may be subjected to 1,000 cycles or more 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 clad is subject to a considerable uncertainty due to the

difficulty of evaluating the strain range which results

from the cyclic interaction of the fuel pellets and

clad. This difficulty arises, for example, from such

highly unpredictable phenomena as pellet cracking, fragmentation, and relocation. Nevertheless, since early

1968, this particular phenomenon has been investigated

analytically and experimentally (Ref. 12). 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 on a methodology to treat the strain

fatigue evaluation of the Westinghouse reference fuel rod

designs. It is believed that the final proof of the adequacy of a given fuel rod design to meet the load follow

requirements can come only from incore experiments

performed on actual reactors. 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 (~400 load

follow cycles) and reactor B (~500 load follow cycles).

In both cases, there was no significant coolant activity increase that could be associated with the load follow mode of operation.

b. Irradiation experience

Westinghouse fuel operational experience is presented in Reference 1. Additional test assembly and test rod experiences are given in Sections 8 and 23 of Reference

9. 4.2-33 Rev. 16 WOLF CREEK

c. Fuel and cladding temperature The methods used for evaluation of fuel rod temperatures are presented in Section 4.4.2.11.

d. Waterlogging

Local cladding deformations typical for waterlogging

• bursts have never been observed in commercial Westinghouse fuel. Experience has shown that the small

number of rods which have acquired clad defects, regardless of primary mechanism, remain intact and do not

progressively distort or restrict coolant flow. In fact, such small defects are normally observed through

reductions in coolant activity to be progressively closed

upon further operation due to the buildup of zirconium

oxide and other substances. Secondary failures which

have been observed in defected rods are attributed to

hydrogen embrittlement of the cladding. Post-irradiation

examinations point to the hydriding failure mechanism

rather than a waterlogging mechanism; the secondary

failures occur as axial cracks in the cladding and are

similar regardless of the primary failure mechanism.

Such cracks do not result in flow blockage or increase

the effects of any postulated transients. More information is provided in References 15 and 16.

e. 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 earlier in the section.

• Waterlogging damage of a previously defected fuel rod has occasionally been postulated as a mechanism for subsequent

rupture of the cladding. Such damage has been postulated as a consequence of a power increase on a rod after water has entered such a rod through a clad defect of appropriate size. Rupture is postulated upon power increase if the rod

internal pressure increase is excessive due to insufficient

venting of water to the reactor coolant. 4.2-34 Rev. 16 WOLF CREEK The potential effects of operation with waterlogged fuel discussed above concluded that waterlogging is not a

concern during operational transients. Clad flattening, as shown in Reference 6 and 19, has been observed in some operating power reactors. Thermal expansion (axial) of the fuel rod stack against a flattened section of the 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 4.2.3.1). Potential differential thermal expansion between the fuel rods and the guide thimbles during a transient is considered in the design. Excessive bowing of the fuel

rods is precluded because the grid assemblies allow axial

movement of the fuel rods relative to the grids.

Specifically, 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.

f. Fuel element burnout and potential energy release

As discussed in Section 4.4.2.2, the core is protected from DNB over the full range of possible operating conditions. In the extremely unlikely event that DNB

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 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.

g. Coolant flow blockage effects on fuel rods

This evaluation is presented in Section 4.4.4.6. 4.2-35 Rev. 16 WOLF CREEK 4.2.3.4 Spacer Grids 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 maintained through the clamping force of 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 13 (LOPAR), Reference 22 (V5H), References 20 and 23 (V5H P+), and References 24 and 25 (RFA and RFA-2). As shown in Reference 13 (LOPAR), Reference 22 (V5H), and References 20 and 23 (V5H P+), and References 24 and 25 (RFA and RFA-2) grid crushing tests and seismic and loss-of-coolant accident evaluations demonstrate that the grids will maintain a geometry that is capable of being cooled under the worst-case

accident Condition III & IV event. 4.2.3.5 Fuel Assembly 4.2.3.5.1 Stresses and Deflections The fuel assembly component stress levels are limited by the design. For example, stresses in the fuel rod due to axial thermal expansion and 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 between the fuel

rod ends and nozzles are provided so that Zircaloy or Zirlo irradiation growth does not result in rod end interferences. 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.

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 6 g lateral directions. Accelerometers are permanently placed into the shipping cask to monitor and detect fuel assembly accelerations that would

exceed the criteria. Past history and experience have indicated that loads

which exceed the allowable limits rarely occur. Exceeding the limits requires

reinspection of the fuel assembly for damage. Tests on various fuel assembly

components, such as the grid assembly, sleeves, inserts, and structure joints, have been performed to assure that the shipping design limits do not result in

impairment of fuel assembly function. Seismic analysis of the fuel assembly is

presented in Reference 13 (LOPAR), Reference 22 (V5H), References 20 and 23 (V5H P+), and Reference 24 (RFA). Since the RFA-2 mid-grid change has no impact on the seismic/LOCA analysis, the conclusion for the RFA Z +2 design in Reference 24 remains valid for the RFA-2 Z +2 design. 4.2-36 Rev. 18 WOLF CREEK 4.2.3.5.2 Dimensional Stability A prototype fuel assembly has been subjected to column loads in excess of those expected in normal service and faulted conditions (Ref. 13). No interference between control rods and thimble tubes will occur during insertion of the rods following a postulated loss-of-coolant accident 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 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 induce only minor thermal bowing, which is insufficient to close the fuel rod-to-thimble tube gap.Reference 13 (LOPAR), Reference 22 (V5H), References 20 and 23 (V5H P+), and Reference 24 (RFA) shows that the fuel assemblies will maintain a geometry

amenable to cooling during a combined seismic and double-ended loss-of-coolant

accident. Reference 25 shows that the grid crush strength and seismic factor P/K 1/2 improved with the RFA-2 design relative to the RFA design. Since the contact length change has no impact on the fuel assembly models used in the seismic and LOCA evaluation, the seismic and LOCA evaluation for the RFA design is applicable for the RFA-2 design. 4.2.3.6 Reactivity Control Assembly and Burnable Absorber Rods

a. Internal pressure and cladding stresses during normal, transient and accident conditions The designs of the standard burnable absorber, WABA, and source rods provide a sufficient cold void volume to

accommodate the internal pressure increase during

operation. This is not a concern for the standard

absorber rod because no gas is released by the absorber

material. For the standard absorber rod, the use of glass in tubular form provides a central void volume along the

length of the rods (see Figure 4.2-12a). For the WABA

rods, an annular plenum is provided within the rod to

accommodate the helium gas released from the absorber

material during boron depletion (see Figure 4.2-12). For the source rods, a void volume is provided within the

rod in order to limit the internal pressure increase

until end of life (see Figures 4.2-13 and 4.2-13a). The stress analysis of the standard absorber and source rods assumes 100-percent gas release to the rod void

volume, in addition to the initial pressure within the

rod. The stress analysis of the WABA rods assumes a

helium release rate of 30% due to the design of the rod. 4.2-37 Rev. 18 WOLF CREEK During normal transient and accident conditions the void volume limits the internal pressures to values which

satisfy the criteria in Section 4.2.1.6. These limits are established not only to ensure that peak stresses do not reach unacceptable values, but also to limit the

amplitude of the oscillatory stress component in consideration of the fatigue characteristics of the materials. Rod, guide thimble, and dashpot flow analyses indicate that the flow is sufficient to prevent coolant boiling

within the guide thimble. Therefore, clad temperatures at which the clad material has adequate strength to resist coolant operating pressures and rod internal pressures are maintained.

b. Thermal stability of the absorber material, including phase changes and thermal expansion The radial and axial temperature profiles within the source and burnable absorber rods have been determined by

considering gap conductance, thermal expansion, neutron

or gamma heating of the contained material as well as

gamma heating of the clad. The maximum temperature of the silver-indium-cadmium alloy or hafnium control rod absorber material was

calculated and found to be significantly less than the

material melting point, and occurs axially at only the

highest flux region. The thermal expansion properties of the absorber material and the phase changes are discussed in Reference 3. The maximum temperature of the borosilicate glass was calculated to be about 1300 F and takes place following the initial rise to power. As the operating cycle proceeds, the glass temperature decreases for the

following reasons: 1) reduction in power generation due to boron-10 depletion, 2) better gap conductance as the

helium produced diffuses to the gap, and 3) external gap

reduction due to borosilicate glass creep. The maximum temperature of the aluminum oxide-boron carbide burnable absorber pellet is calculated to be less than 1200 F which takes place following the initial rise to power. As the operating cycle proceeds, the burnable absorber pellet temperature decreases for the following

reasons: (1) reduction in heat generation due to B 10 depletion, (2) better gap conductance as the helium

produced diffuses to the gap. 4.2-38 Rev. 16 WOLF CREEK 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 plug.

c. Irradiation stability of the absorber material, taking into consideration gas release and swelling The irradiation stability of the absorber material is discussed in Reference 3 for the Ag-In-Cd and hafnium material. Irradiation produces no deleterious effects in the absorber material.

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

Sufficient diametral and end clearances are provided to

accommodate swelling of the absorber material. Based on experience with borosilicate glass and on nuclear and thermal calculations, 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 continue only 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 ends of the

inner liner are open to allow helium, which diffuses out of the glass, to occupy the central void. The Al 2 O 3-B 4 C WABA pellets are designed such that gross swelling or crumbling of the pellets is not expected during reactor operation. Although some minor cracking

of the pellets may occur due to temperature cycles during

startup and shutdown, this cracking should not affect the

overall absorber stack integrity.

d. Potential for chemical interaction, including possible waterlogging rupture The structural materials selected have good resistance to irradiation damage and are compatible with the reactor environment.

Corrosion of the materials exposed to the coolant is quite low, and proper control of chloride and oxygen in

the coolant will prevent the occurrence of stress

corrosion. The potential for the interference with rod

cluster control movement due to possible corrosion phenomena is very low. 4.2-39 Rev. 16 WOLF CREEK Waterlogging rupture is not a failure mechanism associated with Westinghouse-designed control rods.

However, a breach of the cladding for any postulated reason does not result in serious consequences. The Ag-In Cd and hafnium absorber material are 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. There is extensive U.S. Naval reactor experience with unclad hafnium as an absorber material, and its

corrosion resistance has been excellent, in fact it has been reported to be superior to Zircaloy-2, with respect to corrosion resistance (Ref. 3). 4.2.4 TESTING AND INSPECTION PLAN 4.2.4.1 Quality Assurance Program The quality assurance program plan of the Westinghouse Nuclear Fuel Division is summarized in Reference 14. 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. 4.2.4.2 Quality Control Quality control 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.

a. Fuel system components and parts

The characteristics inspected depend upon the component parts and includes dimensional, visual check, audits of test reports, material certification and nondestructive examination such as X-ray and ultrasonic. All material used in this core is accepted and released by Quality Control.

b. Pellets 4.2-40 Rev. 16 WOLF CREEK 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. c. Rod inspection Fuel rod, control rod, burnable absorber, and source rod inspections consist of the following nondestructive examination techniques and methods, as applicable. 1. Leak testing Each fuel, WABA, and secondary source rod is tested, using a calibrated mass spectrometer, with helium being the detectable gas. 2. Enclosure welds All weld enclosures are nondestructively examined by a qualified volumetric nondestructive examination method (e.g., per ASME 142, x-ray or ultrasonics) in accordance with Westinghouse specifications. 3. Dimensional All rods are dimensionally inspected prior to final release. The requirements include such items as length, camber, and visual appearance. 4. Plenum dimensions All of the fuel rods and burnable absorber rods are inspected by X-ray, gamma scanning, or other approved methods to ensure proper plenum dimensions. 5. Pellet-to-pellet gaps All of the fuel rods are inspected by gamma scanning or other methods to ensure that no significant gaps exist between pellets. 6. Enrichment Deviation All of the fuel rods are gamma scanned to verify enrichment control prior to acceptance for assembly loading. 4.2-41 Rev. 16 WOLF CREEK

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

d. Assemblies

Each fuel, control, burnable absorber and source rod assembly is inspected for compliance with drawing and/or specification requirements. Other incore control

component inspection and specification requirements are

given in Section 4.2.4.3.

e. 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 of calibration and conditions of

tools are kept.

2. Audits of inspection activities and records are performed 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.

f. Process control

To prevent the possibility of mixing enrichments during fuel manufacture and assembly, strict enrichment

segregation and other process controls are exercised. The UO 2 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 and density are produced in a given production line at any given time. 4.2-42 Rev. 16 WOLF CREEK Finished pellets are placed on trays 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 Quality Control. Physical barriers prevent

mixing of pellets of different 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 laser marked on each fuel tube, which identifies the contract and enrichment. The end plugs are inserted and the end plugs are then

inert welded to seal the tube. The code provides a

reference to the fuel contained in the fuel rods. At the time of installation into an assembly, the rod codes are placed into a matrix to identify each rod

in its position within a given assembly. Before a fuel

assembly is Quality Control released, the traceability

codes on the described matrix are checked to ensure that the fuel rods in the assembly are from the correct region. Traceability of all fuel assembly components in

an assembly are permanently maintained and identified

with a unique identification number engraved on the fuel

assembly top nozzle. Similar traceability is provided for burnable absorber rods, source rods, and control rods, as required. 4.2.4.3 Incore Control Component Testing and Inspection 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

tests/inspections and functional testing at the plant site are performed. During the component manufacturing phase, the following requirements apply to the reactivity control components to ensure proper functioning during reactor

operation:

a. All materials are procured to specifications to attain the desired standard of quality. 4.2-43 Rev. 16 WOLF CREEK

b. All spider assemblies are proof tested by applying a load to the spider body so that a specified load with a given tolerance is applied to each vane. This proof load applied to each vane provides a bending moment at the spider body greater than the load caused by the acceleration imposed by the control rod drive

mechanism.

c. All rods are checked for integrity by the methods described in Section 4.2.4.2, item c. d. To ensure proper fitup with the fuel assembly, the rod cluster control, burnable absorber, and source assemblies are installed in the fuel assembly without restriction or binding in the dry condition. In addition, each rod assembly must meet a straightness requirement over the entire inserted length of each rod assembly. Following core loading, but prior to initial criticality, the rod cluster control assemblies were tested to demonstrate reliable operation in accordance with Regulatory Guide 1.68, Appendix A, Section 2.b. This testing is further discussed in Section 14.2.12.3.27.

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 rod cluster control assembly, as required by the technical specifications. In addition, periodic drop tests of the rod cluster control assemblies are performed at each refueling shutdown to demonstrate continued ability to meet trip time requirements. If a 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 rod cluster control assembly can be tolerated. More than one inoperable rod cluster control assembly could be tolerated, but would impose additional demands on the plant operator.

Therefore, the number of inoperable rod cluster control assemblies has been

limited to one. 4.2.4.4 Tests and Inspections by Others If any tests and inspections are to be performed on behalf of Westinghouse, Westinghouse will review and approve the quality control procedures, inspection plans, etc. to be utilized to ensure that they are equivalent to the description provided in Sections 4.2.4.1 through 4.2.4.3 and are performed to

meet all Westinghouse requirements. 4.2-44 Rev. 17 WOLF CREEK 4.2.4.5 Inservice Surveillance Westinghouse has conducted a program to examine detailed aspects of the 17 x 17 fuel assembly. This program is described in Section 23 of Reference 9. Reference 1 is periodically updated in order to provide recent results of operating experience with Westinghouse fuel and incore control components. 4.2.4.6 Onsite Inspection Written procedures are used by the station staff for the post-shipment inspection of all new fuel and associated components, such as control rods, plugs, and inserts. Fuel handling procedures specify the sequence in which

handling and inspection take place. Loaded fuel containers, when received onsite, are externally inspected to ensure that labels and markings are intact and seals are unbroken. After the

containers are opened, the shock indicators attached to the suspended internals

are inspected to determine if movement during transit exceeded design limitations. Following removal of the fuel assembly from the container in accordance with detailed procedures, the fuel assembly plastic wrapper is examined for evidence of damage. The polyethylene wrapper is then removed, and a visual inspection

of the entire bundle is performed. Control rod, source and burnable absorber assemblies usually are shipped in fuel assemblies and are inspected after removal of the fuel assembly from the

container. The control rod assembly is withdrawn a few inches from the fuel

assembly to ensure free and unrestricted movement, and the exposed section is

visually inspected for mechanical integrity, replaced in the fuel assembly and stored with the fuel assembly. Control rod, source or burnable poison assemblies may be stored separately or within fuel assemblies. 4.2-45 Rev. 16 WOLF CREEK 4.

2.5 REFERENCES

1. Slagle, W.H., "Operational Experience with Westinghouse Cores," WCAP-8183 (latest revision).

2. Beaumont, M. D., et al., "Properties of Fuel and Core Component Materials," WCAP-9179, Revision 1 (Proprietary) and WCAP-9224 (Non-Proprietary), July

1978.

3. Beaumont, M. D., et al., "Properties of Fuel and Core Component Materials", WCAP-9179, Revision 1 (Proprietary) and WCAP-9224 (Non-Proprietary) Appendix A, "Hafnium", October 1980. 4. Deleted

5. Deleted

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

7. Risher, D. H., et al., "Safety Analysis for the Revised Fuel Rod Internal Pressure Design Basis," WCAP-8963

(Proprietary), November 1976 and WCAP-8964 (Non-Proprietary), August 1977. 8. Skarita, J., et al., "Westinghouse Wet Annular Burnable Absorber Evaluation Report", WCAP-10021-P-A, Revision 1 (Proprietary), October, 1983. 9. Eggleston, F. T., "Safety-Related Research and Devel- opment for Westinghouse Pressurized Water Reactors, Program Summaries - Winter 1977 - Summer 1978," WCAP-8768, Revision 2, October 1978. 10. Demario, E. E., "Hydraulic Flow Test of the 17 x 17 Fuel Assembly," WCAP-8278 (Proprietary) and WCAP-8279 Non-Proprietary), February 1974. 11. Skaritka, J. (Ed.), "Fuel Rod Bow Evaluation," WCAP- 8691, Rev. 1 (Proprietary) and WCAP-8692, Rev. 1 (Non- Proprietary), July 1979. 4.2-46 Rev. 18 WOLF CREEK 12. O'Donnell, W. J. and Langer, B. F., "Fatigue Design Basis for Zircaloy Components," Nuclear Science and Engineering, 20, 1-12, 1964. 13. Gesinski, L. and Chiang, D., "Safety Analysis of the 17 x 17 Fuel Assembly for Combined Seismic and Loss-of- Coolant Accident," WCAP-8236 (Proprietary) and WCAP-8288 (Non-Proprietary), December 1973 and Addendum 1 (Proprietary) March 1974. 14. Quality Management System, Revision 2.1 15. Stephan, L. A., "The Effects of Cladding Material and Heat Treatment on the Response of Waterlogged UO2 Fuel Rods to Power Bursts," IN-ITR-111, January 1970. 16. Western New York Nuclear Research Center Correspondence With the U.S. Atomic Energy Commission on February 11 and August 27, 1971, Docket No. 50-57. 17. Davidson, S.L., et.al., "Extended Burnup Evaluation of Westinghouse Fuel," WCAP-10125-P-A (Proprietary) and WCAP-10126-A (Non-Proprietary), December 1985. 18. Weiner, R.A., et.al., "Improved Fuel Performance Models for Westinghouse Fuel Rod Design and Safety Evaluations," WCAP-10851-P-A (Proprietary) and WCAP-11873-A (Non-Proprietary), August 1988. 19. Kersting, P.J., et.al., "Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel, WCAP-13589-A (Proprietary), March 1995. 20. Davidson, S.L., Nuhfer, D.L. (Eds.) VANTAGE + Fuel Assembly Reference Core Report, WCAP-12610-A and Appendices A through D, June 1990. 21. Davidson, S.L., Westinghouse Fuel Criteria Evaluation Process, WCAP-12488-A, October 1994. 22. Davidson, S.L., (Ed.), Reference Core Report - VANTAGE 5 Fuel Assembly, WCAP-10444-P-A, September 1985 and Addendum 2A, February

1989. 23. Letter from Slater, J.L., Westinghouse, to Norton, W.B., WCNOC, dated May 5, 1994, Safety Evaluations For Region 10 (SAHF) Fuel

Changes, Westinghouse letter number 94SAP-G-0027 (WCNOC letter

number 94-00490) and Attachment A, SECL-92-305, Performance + Fuel

Features. 24. Kitchen, T. J., 17 x 17 Standard Robust Fuel Assembly (17 x 17 STD RFA), Safety Evaluation Check List, SECL-98-056, September 1998. 25. Seel, D. D. (Ed.), 17x17 Robust Fuel Assembly with RFA-2 Mid-Grid Final Design Review Package, DR-01-5 (Proprietary), October 2001. 26. H. A. Sepp, Fuel Criterion Evaluation Process (FCEP) Notification of the RFA-2 Design (Proprietary, LTR-NRC-01-44, December 19, 2001. 27. Foster, J. P. and Sidener, S., Westinghouse Improved Performance Analysis and Design Model (PAD 4.0), WCAP-15063-P-A, Revision 1 with Errata, July 2000. 4.2-47 Rev. 18

o. 058. GUIDE THIMBLE ' WOLF CREEK 0 FUEL RODS 261J. REQ'D OD = 0.371J. CLAD THICKNESS=

0.0225 CLAD MATERIAL-ZIRC-IJ. FUEL ASSEMBLY WITH ROD CLUSTER CONTROL 0 0 o-+ 00 0 0 0 0 0 0 0 0 0 0 o 0 .. G 0 I 1.973 FUEL ASS'Y AND CONTROL ROD PITCH 8.1J.66 TYP INSTRUMENTATION SHEATH FUEL ASSEMBLY WITHOUT ROD CLUSTER CONTROL Rev. 11 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.2-1 FUEL ASSEMBLY CROSS SECTION 17 X 17 LOPAR 245 GUIDE THIMBLE 8.426 ...--.1 FliEL ASSEMBLY WITH ROO CONTROL CLUSTER 11 .973 FUE:l AND RCCA PITCH fJ> 0 0 e 0 0 0 0 0 0 0 0 0 0 0 0 0 CONTROL CLUSTER ELEMENT \__ FUEL ROO (264 PER FUEL ASSEMBLY) 00 = .374 CLAD THICKNESS = CLAD MATERIAL = ZIRC-*4* "" r 8.466 TYP _l. INSTRUMENTATION SHEATH FUEL ASSEMBLY WITHOUT ROO CLUSTER CONTFIOL. .A GUIDE THIMBLE DIMENSIONS AT TOP NOZZLE ADAPTOR PLATE Rev. 11 DIMENSIONS ARE IN INCHES (NOt.liNAl) WOLIF CREEK U P OAT ED SA F E T Y AN A LYSIS R E P 0 R l FIGUBE 4.2-1A FUEL ASSEMBLY CROSS SECTION 17 x 1"7 VANTAGE 5H Wolf Creek ADAPTER PLATE \ 8 RfF -------r 8. 42" SQ REF Rev. 0 14023 BOTTOM VIEW WOLF CREEK DPDATED SAFETY ANALYSIS REPORT FIGURE 4.2-2 TYPICAL FUEL ASSEMBLY OUTL! NE 17 x 17 7.162 TYP REF TOP VIEW 0.875 il TYP REF 8.404 TYP REF .578 Ill REF 6.750 TYP REF BOTIOM VIEW 8 424 rYP REF 159.975 3.475 ---1t----t-t 1-o....-------------------- 152.200 IT 2 383 GRIDS 1 & 8 1 .522 ---153.60 '---4t5 IFM GRI0 1 s I 1--122 31 1--101 76 I L 1.500 GRIDS 2-7 133.10 1--112 55 92 00 -r In I --5UYO I I I INC 0 0::: 0 8HIIl I=D II II --9 1©11 0::: 0 @_] r--JO J5 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.2-2A FUEl ASSEMBLY OUTLINE 17 X 17 VANTAGE 5H WITH IFM GRIDS Rev. 11 1.522 L 153.60 TOP VIEW 0.875 DIA 1YP REF 8.404 1YP .578 OIA REF 0.875 DIA TYP REF 6.750 TYP 0 8.424 1YP REF REF L.i _j BOTIOM V!EW 159.975 ------------------------*-!1 152.200 F/R LENGTH -------.j*'l*

• II I :I I: 2.383 t.1751Fiol CRI!S L j PBC 12J.31 I...-101.76 81.21 I ' I . I --1 1--1.500 I I CRIDS2-71 I 1JJ.10 1--112.55 92.00 1--71.45 I 1---50.90 I I I 1---29.70 I I 1--6.5J5 ---1 WOLF CREEK FIGURE 4.2-28 FUEL ASSEMBLY OUTLINE Rev. 11 17x17 VANTAGE 5H with IFMs & PBG

n ... nl **l-------

• • -1152.100 REI. 11CG L.EtarN)*

--*--------------------oto----1--(f, !IS) L L REV.18 211 (71,41) IC!ICI. fill WOLF CRBBK: tJPDAT.ID SAPB"'T ANALYSIS RBPOllT FIGURE 4.2-2C FUEL ASSEMBLY OUTLif\E .17x17 VANTAGE 5H WITH PERFORMANCE+ FEATURES CV5H P+) (. 621 )---T-(7. 162) TYP-------! TYP

• I . 875 DIA TYP I r-(8. 424) SQ.----! 1 r-(6. 150) TYP *--1 : I l (7.504) TYP (. 900) TYP I __ =i___L t_ __ (. 580) DIA Q -a

---

(!59. 975) ----------.--------- (3. 475)-t *--, ! .... --(152.800 FUEL ROD LENGTH)-------- r (8. 426) SQ TYP I I I I r ( 8. 41 8) SQ TYP -(8. 386) SQ TYP I 0 0 0 . 875 DIA TYP ! -I 0 -=+ I 837) TYP J (. I I I --r-+(2. 383) (8. 408) so l , I I REV .18 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.2-20 FUEL ASSEMBLY OUTLINE 17x17 VANTAGE 5H WITH PERFORMANCE+ FEATURES, ZIRLo*2 (V5H P+Z *2) RF A z*2 AND RF A-2 z*2 (. 621) TYP (7.504) TYP (. 571 DIA) D (8.424) (6.750) TYP II (159.975) ( 3. 4 7 5) --t----------1 (1.610) ---1-----t-++ 1.270 MIN 1. 7 91 MAX --1-------1 D (8. 386) TYP REV. 21 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.2-2E FUEL ASSEMBLY OUTLINE 17x17 STANDARD,PERFORMANCE* FEATURES Zirlo" 2 , RF A, RF A-2, WIN UPDATED SAFETY ANALYSIS REPORT WOLF CREEK FIGURE 4.2-2F REV. 29 F/A OUTLINE (COMBO GRID AND SDFBN) 17 STD WIN P+ Z+2 RFA TOP AND BOTTOM RECONSTITUTABLE 151.6" (REF) (TYP) WOt..F CREEK P\.UG -SPRING ..--?ELL.£':'5 L GAP ,--Z I RCAI.OY CLAD BOT1'DM END PLUG, DIMENSIONS DEPEND ON DESIGN vARIABLES SUCH AS Rev. 11 oqE-?RESSUR tZlTtON. lMO DISCHARGE 3URNUP UPDA'l'Jm SA!"E'l'Y AHALYSIS REPOR'!' FIGURE 4.2-:5 FUEL Roo SCHEMATIC STANDARD ROD 152.200 FUEL ROD LENGTH 7.440 PLENUM 144.00 ,\ I\, *-ri-/---SEAL WELD .// *-H GIRTH WELD --1-1-__ u_. :s:""' -) ---*-11 !I li ---U0 2 PELLET FUEL STACK LENGTH .J\(' ---1 PELLET -CLAD GAP n-----ZIRCALOY CLAD GIRTH WELD ,.: ...

• BOITOM END PLUG ILLUSTRATES INTERNAL GRIP CONNECTION
• ------------

,...__ SPECIFIC DIMENSIONS DEPEND :lN DESIGN VARIABLES SUCH AS PREPRESSURIZATION, l*ti!)TORY, AND DISCHARGE BURNUP DIMENSIONS ARE IN INCHES (NOMINAl.)

• ------WOLF CREEK UPDATED SAFETY ANALYSIS REPOHT Rev .. 11 *-----***-*---**,*------

.... -FIGURE 4.2-3A FUEL ROD SCHEMATIC HIGH BURI\JUP HOD l 152.200 FUEL ROD LENGTH T 6.940 PLENUM 144.00 SEAL WELD --GIRTH WELD FUEL STACK LENGTH --ZIRCALOY ClJ\D --GIRTH WELD 1":::.., \ BOTTOM END PLUG ILLUSTRATES

• ----*-* -----\[_ INTERNAL CONNECTION SPECIFIC DIMENSIONS DEPEND ON DESIGN VARIABLES SUCH AS PREPRESSURIZATION, POWER HISTORY, AND DISCHARGE BURNUP DIMENSIONS ARE IN INCHES (NOMINAL)

WOLF CREEK FIGURE 4.2-36 FUEL ROD SCHEMATIC PERFOHMANCE + ZIRC-4 Rev. 11 J: ..... I i IY z: ... I ... 11!1 M ' 0 0 !9 REV1B WaLII' caBB:I: UPDATED SAFKTY ANALYSIS RBPOR.T FIGURE 4.2-3C FUEL ROD SCHEMATIC PERFORMANCE+, ZIRLO J: .... ri II I .. ! ... Ill I !i I .. I l ...

• 0 ID -.,. -... REV.18 WOLP CRBU: UPDA'l'BD SAJ'ET1' .ANALYSIS J.UDIOilT FIGURE +.2-.JD FUEL ROD SCHEMATIC PERFORMANCE+

1 ZIRL0 .. 2 GRID STRAP < / OIAMETV ---.,_'\ ' *+ ) / ' / ZIRCALOY SLEEVE /,_ ZIRCALOY THIMBLE -FUEL ROD "" ""-EXPANSION LOBE MID GRID EXPANSION ,JOINT DESIGN WOLF CREEK FIGURE 4.2-4 MID GRID EXPANSION JOINT PLAN VIEW Rev. 9 GUIDE THIMBLE TUBE "'l MID GRID- ... MID GRID SLEEVE EXPANSION LOBE DIMENSIONS ARE IN INCHES (NOMINAL) \'-WELD MIXING VANE WOLF CREEK FIGURE 4 .. 2-5 MID GRID EXPANSION JOINT ELEVATION VIEW Rev. 9 h-*-------------------------------------- WOLF CREEK ZIRCALOY THIMBLE BUTT WELD ALL AROUND TOP NOZZLE ADAPTER PLATE STAINLESS STEEL SLEEVE EXPANSION LOBE TOP GRID Rev. 11 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.2-6 TOP GRID TO NOZZLE ATTACHMENT STANDARD UNASSEMBLED LOCK TUBE ADAPTOR PLATE THRU HOLE NOZZLE INSERT E -E (TYPI::AL) ASSEMBLED WOLF CREEK UPDATED SAFETY ANALYSIS BEI'CIRT FIGURE 4.:2-6A Rev .. *1 11 THIMBLE I INSERT I TOP GRID SLEEVE BULGE JOINT GEOMETRY BOHOM GRID PROTECTIVE GRID THIMBLE TO END PLUG WELD INSERT TO END --"' PLUG WELD SPOT WELDS (TYP 20 of 24 INSERTS) BOTIOM GRID SS INSERT ZIRCALOY GUIDE THIMBLE THREADED END PLUG r;t,"'l?"-+t*<2ltfllt11----- SPOT WELD'S (TYP. 4 of 24 INSERTS) SS BOTIOM NOZZLE TOP PLATE SS THIMBLE SCREW WITH INTEGRAL LOCKING CAP WOLF CREEK FIGURE 4.2-7 GUIDE THIMBLE TO BOTIOM NOZZLE JOINT Hev. 9 .. _________ , _____ , ____ , __ 11._ ..... 1 .. ACC INSI:RT£0 !JRjy( H(IJ ASSDe.Y Fll.l UNGTH RCC ASSDe.Y WID£ TUII[ A55£-Y \ \ \ __ Fl£L A 55£...._ 'f' Rev. 6 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.2-8 ROD CLUSTER CONTROL AND DRIVE ROD ASSEMBLY WITH INTERFACiNG COMPONENTS I I 1.645 MAX 160.949 ---------------11, SPRINGS (2) BODY .362 .365 OIA 150.574 ----------"'"1 I ! I I ABSORBER I I L-142.00__] ABSORBER LENGTH HAFNIUM OR 60% SILVER 15% INDIUM 5% CADMIUM Ct-<EEK FIGURE 4.2-9 ROD CLUSTER CONTROl OUTLiNE Rev. 9 TOP END PLUG2 oo I[_ 11m I .38i OiA NOM TYPE 304 STAINLESS LSTEEL TUBE lD I / ., Ag-ln-Cd OR HAFNIUM BOTTOM END PLUG 142.00 ABSORBER MATERiAL 151.885 WOLF CREEK FIGURE 4.2-10 ABSORBER ROD Rev. 9 I i

---

REV. IS WOLF CREEK UPDATED SAFETY YSIS REPORT 42.-11 WET ANNU.AR BURNABLE ABSORBER Y

"" I I{ 1--;; ...... fo t-" --. --E:::> : -8/ ---1------,--I"IIILICUI 1 sun ruN BURNABLE ABSORBER ROO --.-Wolf Creek 151 60 142.0 REF. r .... -l)!J ABSORBER LENGTH l -ll**-----=::Ji

---

I . -*-----*--t:::;l f-----Rev. 11 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.2-11a TYPICAL BURUS1LICATE GLASS BtlRNi\RI E ARSORRFR ROD ASSEMBLY TOP CONNECTOR SPRING CUP PELLET STACK BOTTOM PLUG -TOP END PLUG CLAD INNER TUBE -Zr4 SPACER -

,_ ___________ ll!!!ll!(j f .381 O.D . .a+--INNER TUBE SECTION A-A BURNABlE ABSORBER ROD Zr-4 149.83 132.00 _____________ ____,,.......,_ ... J WOLF CREEK FIGURE 4.2-12 Rev. 9 WET ANNULAR BURNABLE ABSORBER ROO ASSlMBLY WOLF CREEK 152.350 ------*------------------.-1 ,--STAINLESS STEEL /TOP END PLUG I I I 7.5l8 REF .. r+- REF. ---->--1+----------------- 150.018 REF. TUBE LENGTH BOROS I L I CHE GLASS SECTION A-A IS SHOWN IN FIGURE 4.2-12c BOTTOM END PLUG , I 91! 10.385 I DIA. MAX. I 142.00 REF ABSORBER LENGTH -----------------t' 0.875 REF. t i I f--WOLF CREEK Rev. 11 UPDATED SAFETY ANALYSIS REPORT FIGURE 4.2-12a GLASS .1\.BSORBER ROD I TO: END PlUG buL.:. ___J,I.J.L.II ____ --'/ I HOlDDOVN SPRING .-SPACER SECONDARY SOURCE PELLET STACK BOTTOM END PLUG A A i.oo t--------------------------- 152.300 SECONDARY SOURCE ----+<<< ANiiM0NY-BERYlliUt.4 \Vflllll/l/IJJ .3380.0. SECTION A-A SECONDARY SOURCE CLAD SS-304 .381 0.0. Rev. 11 WOLF CREEK Up DATED SAFETY ANALYSIS REPORT FIGURE SECONDARY SOURCE ROD ASSEMBLY TOP END PLUG TOP END PLUG HOLD DOWN SPRING OUTER TUBE SOURCE 152.358 SECONDARY SOURCE ROD ASSEMBLY c HOLD DOWN SPRING BOTTOM END PLUG BOTTOM END PLUG .381 DIA. iC=:§I 1]1 t CAPSULE CLAD--+---' SPACrn i-ol*>-----,,f"! ...... .334 DIA. SECONDARY SOURCE -------, OUTERCLAD SS-304 .381 O.D. SECONDARY SOURCE CAPSULE CLAD SS-304 .334 O.D. SECONDARY ANTIMONY -BERYLLIUM .292 O.D. SECTION 11 A-N 1 SCALE 8:1 DOUBLE ENCAPSULATED SECONDARY SOURCE ROD 148.000 SECONDARY SOURCE CAPSULE WOLF CREEK UPDATED SAFETY ANALYSIS REPORT Figure 4.2-13A, REV. 13 DOUBLE ENCAPSULATED SECONDARY SOURCE ROD ASSEMBLY

6.320 I l i . ! . ' t---+-1.675 . 885 .190 F !55.900 -----------------------1 4.385 --**r:------------- .-151.515

---

7.942

___ .,...., . r--------8A.oo _____ _, ... SECONDARY SOURCE I 689Mi'9&&AAS3&M!Cl&iS'O&dll

>

I J tfl HI I I L DOUBLE SPRING DESIGN .250 --i 1---/CLAD . SECONDARY PELLET .. 338 0.0. SECTION A-A SECONDARY SOURCE --1 I--.665 SECTION B-B THIMBLE PLUG .424 0.0. SS-304 r Rev 11 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT F lGU RE 4.2-SECONDARY SOURCE ASSEMBLY ...., j _; .. r: =Fr r -w_ wwww'\ ! r-ruuuul7J 1/ f :!II hi II I:) f ""'SPRINGS__} WOLF CREEK -150 4 I I. 5 NOM. I CAL I FORNI UM *)H( 142 REF ABSORBER LENGTH NOTE: ALL DIMENSIONS ARE IN INCHES _I 32.5 NOM. I SPACER I I l Rev. 11 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4. 2-14A TYPICAL PRIMARY SOURCE ASSEMBLY

8.104 SECONDARY SOURCE ------. OUTERCLAO SS-304 . 381 O.D. SECONDARY SOURCE CAPSULE ClAD SS.-304 .334 O.D. . SECONDARY SOURCE ----l ANTIMONY-BERYLLIUM .292 0.0. SECTION "A-A" SCALE 8:1 DOUBLE ENCAPSULATED SECONDARY SOURCE ROD ..... __ _.,_._ _____________ -oot t--.270 I rr88.00l

• -Ill E3 " .. E ... """""" .426.0.0 . THIMBLE PLUG *. SS-304 SECTiON "B*S" SCALE 8:1 THIMBLE PLUG i4. THIMBLE PLUG WOLF CREEK UPDATED SAFETY ANALYSIS REPORT Figure 4.2-148, REV. 13 DOUBLE ENCAPSULATED SECONDARY SOURCE ASSEMBLY ! i ! i I I j i l i i j i i I j ' I L .. -**---**-*---

.. -.. -**-*---.. -*---.. -**---.. -**-**-*-**-**-**-**-**-*--- .. -**---**-*---*-*-**-*--- .. -**-**-*-**-**-**-*-**-**-*-**-*-*-*-*- .. -*-*---*-.. ---*-.. -**-**-*-*-*---**-**- .. -*-**-.. -**-*-**-.. -*-**-**---**-**-**-**-*- .. -*-**-**-*-**-**-* ... k;....---7.94. E.:-----... o * .q4

• • • • -e+ -e<l .-e:<l -= D : -D : -D REV. 18 WOLF CREEK UPDATED SAFETY AN.AL YSIS REPORT FIGURE 4-.2-15 TYPIC.AL DOUBLE SPRING THIMBLE PLUG DEVICE

...., ___ .,. ___ ... REV. 18 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.2-15A TYPICAL SINGLE SPRING THIMBLE PLUG DEVICE WOLF CREEK 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) presented in 10 CFR 50, Appendix A. Where applicable, supplemental criteria such as the "Final Acceptance Criteria for Emergency Core Cooling Systems" are addressed. But, 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:

a. Condition I - Normal Operation

b. Condition II - Incidents of Moderate Frequency

c. Condition III - Infrequent Faults

d. 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 (fuel damage as used here is defined as penetration of the fission product barrier, i.e., the fuel rod clad) is not expected during Condition I and Condition II events. It is not possible, however, to preclude a very small number of rod failures. These are

within the capability of the chemical and volume control system (CVCS) and are

consistent with the plant design basis. Condition III incidents do 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 is not sufficient to

interrupt or restrict public use of those areas beyond the exclusion radius.

Furthermore, a Condition III incident does not by itself generate a Condition

IV fault or result in a consequential loss of function of the reactor coolant

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 do not cause a release of radioactive material that results in exceeding the

limits of 10 CFR 100. The core design power distribution limits related to fuel integrity are met for Condition I occurrences through conservative design and maintained by the action of the control system. The requirements for Condition II occurrences are met by providing an adequate protection system which monitors reactor

parameters. The control and protection systems are described in Chapter 7.0, and the consequences of Condition II, III, and IV occurrences are given in

Chapter 15.0. 4.3-1 Rev. 11 WOLF CREEK 4.3.1.1 Fuel Burnup Basis A limitation on initial installed excess reactivity or average discharge burnup 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. 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 (e.g., the controlling bank at the "bite" position). In terms of chemical shim boron concentration, this represents approximately 10 ppm with no control rod insertion. 4.3.1.2 Negative Reactivity Feedbacks (Reactivity Coefficient) Basis The fuel temperature coefficient will be negative, and the moderator temperature coefficient of reactivity will be nonpositive for full-power

operating conditions, thereby providing negative reactivity feedback

characteristics. The design basis meets GDC-11. Discussion When compensation for a rapid increase in reactivity is considered, there are two major effects. These are the resonance absorption effects (Doppler) associated with changing fuel temperature and the neutron spectrum and reactor

composition change effects 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 rapid reactivity compensation. The core is also designed to have an overall non-positive

moderator temperature coefficient of reactivity during full power operation so that average coolant temperature or void content provides another, slower

compensatory effect. Full power operation is permitted only in a range of overall non-positive moderator temperature coefficient. The desired moderator temperature coefficient can be achieved through use of fixed burnable absorber and/or control rods by limiting the reactivity held down by soluble boron. 4.3-2 Rev. 11 WOLF CREEK Restrictions on burnable absorber content (quantity and distribution) are not applied as a design basis other than as they relate to accomplishment of the

desired moderator temperature coefficient at power operating 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:

a. The fuel will not be operated at greater than 14.48 kW/ft under normal operating conditions, including an allowance

of 2 percent for calorimetric error and not including

power spike factor due to densification.

b. Under abnormal conditions, including the maximum over-

power condition, the fuel peak power will not cause

melting, as defined in Section 4.4.1.2.

c. The fuel will not operate with a power distribution that

violates the departure from nucleate boiling (DNB) design basis (i.e., the DNB Ratio (DNBR) shall not be less than 1.30 for W-3 analyses and 1.76 for WRB-2 analyses, as

discussed in Section 4.4.1) under Condition I and II events,

including the maximum overpower condition.

d. Fuel management will be such as to produce values of fuel rod power and burnup 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.

Even though there is good agreement between calculated peak power and

measurements, a nuclear uncertainty (see Section 4.3.2.2.1) is applied to

calculated peak local power. Such a margin is provided both for the analysis

for normal operating states and for anticipated transients.

4.3-3 Rev. 23 WOLF CREEK 4.3.1.4 Maximum Controlled Reactivity Insertion Rate Basis The maximum reactivity insertion rate due to withdrawal of rod cluster control assemblies at power or by boron dilution is limited. During normal at power

operation, the maximum controlled reactivity insertion rate is less than 35

pcm/sec*. A maximum reactivity change rate of 75 pcm/sec

• for accidental withdrawal of control banks is set such that peak heat generation rate and DNBR

do not exceed the maximum allowable at over-power 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 Chapter 15.0). Following any Condition IV event (rod ejection, steam line break, etc.) the reactor can be brought to the shutdown condition, and the core will maintain acceptable heat transfer geometry. This satisfies GDC-28.

• 1 pcm = 10

-5 p (see footnote to Table 4.3-2). 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/sec. During normal operation at power and with control rod overlap, the maximum reactivity change rate is less than

35 pcm/sec. The reactivity change rates are conservatively calculated assuming unfavorable axial power and xenon distributions. The peak xenon burnout rate is 25

pcm/min, significantly lower than the maximum reactivity addition rate of 35 pcm/sec for normal operation and 75 pcm/sec for accidental withdrawal of two

banks. 4.3-4 Rev. 11 WOLF CREEK 4.3.1.5 Shutdown Margins Basis Minimum shutdown margin as specified in the COLR is required at any power operating condition, in the hot standby condition, hot shutdown condition, and in the cold shutdown condition. In all analyses involving reactor trip, the single, highest worth rod cluster control assembly is postulated to remain untripped in its full out position (stuck rod criterion). This satisfies GDC-26. Discussion Two independent reactivity control systems are provided: 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 acceptable fuel damage limits (very small number of rod failures), assuming that the highest worth control rod is stuck out upon trip. The boron system can compensate for all xenon burnout reactivity changes and will maintain the reactor in the cold shutdown condition. Thus, backup and emergency shutdown provisions are provided by a mechanical and a chemical shim control system which satisfies GDC-26. Basis When fuel assemblies are in the pressure vessel and the vessel head is not in place, k eff will be maintained at or below 0.95 with control rods and soluble boron. Further, the fuel will be maintained sufficiently subcritical that removal of all rod cluster control assemblies will not result in criticality. Discussion ANSI Standard N18.2 specifies a k eff not to exceed 0.95 in spent fuel storage racks and transfer equipment flooded with pure water and a k eff 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 the new fuel

storage, is adequate for the controlled and continuously monitored operations

involved.The boron concentration required to meet the refueling shutdown criteria is specified in the Core Operating Limits Report (COLR). Verification that this

shutdown criteria is met, including uncertainties, is achieved based on

calculations performed with the ANC computer code (Reference 31). The

subcriticality of the core is continuously monitored, as described in the

Technical Specifications. 4.3-5 Rev. 13 WOLF CREEK 4.3.1.6 Stability Basis The core will be inherently stable to power oscillations at the fundamental mode. This satisfies GDC-12. Spatial power oscillations within the core with a constant core power output, should they occur, can be reliably and readily detected and suppressed. 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. The core is designed so that diametral and azimuthal oscillations due to spatial xenon effects are self-damping, and no operator action or control

action is required to suppress them. The stability to 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 excore long ion chambers. Indications are also continuously available from incore thermocouples and loop temperature measurements. Movable incore detectors can

be activated to provide more detailed information. In all 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 excore detectors are provided for control and monitoring

of axial power distributions. Assurance that fuel design limits are not exceeded is provided by reactor OverpowerT and Overtemperature T trip functions which use the measured axial power imbalance as an input. Detection and suppression of xenon oscillations are discussed in Section 4.3.2.7. 4.3.1.7 Anticipated Transients Without SCRAM 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, generic analyses of the consequences of a hypothetical failure to trip following anticipated transients has shown that no significant core damage would result, system peak pressures would be limited to acceptable values, and no failure of the reactor coolant system would result (Ref. 1 and 3). Nevertheless, in accordance with the final USNRC ATWS rule; 10CFR50.62(b) Requirements for Reduction of Risk from Anticipated Transients Without Scram (ATWS) Events for Light-Water-Cooled Nuclear Power Plants, ATWS Mitigation System Actuation Circuitry (AMSAC) has been installed at Wolf Creek (see Section 15.8). The AMSAC system initiates a turbine trip and actuates auxiliary feedwater independent of the reactor trip system. The AMSAC equipment is described in Section 7.7.1.11. 4.3-6 Rev. 13 WOLF CREEK 4.

3.2 DESCRIPTION

4.3.2.1 Nuclear Design Description The reactor core consists of a specified number of fuel rods which are held in bundles by spacer grids and top and bottom fittings. The fuel rods are constructed of Zircaloy or Zirlo cylindrical tubes containing UO 2 fuel pellets. The bundles, known as fuel assemblies, are arranged in a pattern which approximates a right circular cylinder. Each fuel assembly normally contains a 17 x 17 rod array composed of 264 fuel rods, 24 rod cluster control thimbles, and an incore instrumentation thimble.

Figure 4.2-1 shows a cross-sectional view of a 17 x 17 fuel assembly and the

related rod cluster control locations. Further details of the fuel assembly

are given in Section 4.2. Fuel assemblies of different enrichments are used in the WCGS core loadings to establish a favorable radial power distribution. A typical checker-board loading pattern is shown in Figure 4.3-1. The exact reloading pattern, initial and final positions of fuel assemblies, and the number of fresh fuel assemblies

and their placement are dependent on the energy requirement for each cycle, and

burnup and power histories of previous cycles prior to Cycle 4. The core average enrichment is determined by the amount of fissionable material required to provide the desired core life-time and energy requirements. The

physics of the burnout process is such that operation of the reactor depletes

the amount of fuel available due to the absorption of neutrons by the U-235

atoms and their subsequent fission. In addition, the fission process results

in the formation of fission products, some of which readily absorb neutrons.

These effects, depletion and the buildup of fission products, are partially

offset by the buildup of plutonium shown in Figure 4.3-2 for a typical 17 x 17

fuel assembly, which occurs due to the nonfission absorption of neutrons in U-

238. Therefore, at the beginning of any cycle a reactivity reserve equal to

the depletion of the fissionable fuel and the buildup of fission product

poisons over the specified cycle life must be "built" into the reactor. This

excess reactivity is controlled by removable neutron absorbing material in the

form of boron dissolved in the primary coolant and burnable absorber rods or

IFBAs.The concentration of the soluble neutron absorber is varied to 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 change. Figure 4.3-46 shows a typical boron letdown

curve. Using its normal makeup path, the CVCS is capable of inserting negative

reactivity at a rate of approximately 30 pcm/min when the reactor coolant boron

concentration is 1,000 ppm and approximately 35 pcm/min when the reactor

coolant boron concentration is 100 ppm. If the emergency boration path is

used, the CVCS is capable of inserting negative reactivity at a rate of

approximately 65 pcm/min when the reactor coolant concentration is 1,000 ppm

and approximately 75 pcm/min when the reactor coolant boron concentration is

100 ppm. The peak burn-out rate for xenon is 25 pcm/min (Section 9.3.4

discusses the capability of the CVCS to counteract xenon decay). Rapid

transient reactivity requirements and safety shutdown requirements are met with

control rods. 4.3-7 Rev. 12 WOLF CREEK During operation, the absorber content in burnable absorber rods or IFBAs is depleted, thus adding positive reactivity to offset some of the negative

reactivity from fuel depletion and fission product buildup. The depletion rate

of the burnable absorber rods or IFBAs 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 4.3-3 is a plot of typical core depletions with and without burnable absorber rods. Note that even at end-of-life conditions some residual absorber remains in the burnable absorber rods, resulting in a net decrease during the cycle lifetime. In addition to reactivity control, the burnable absorber rods are strategically located to provide a favorable radial power distribution. Figure 4.3-4 shows the burnable absorber distributions within a fuel assembly for several IFBA patterns used in a 17 x 17 array for WCGS. A typical core loading pattern with IFBA for WCGS is shown in Figure 4.3-5. Tables 4.3-2 through 4.3-4 contain a summary of the reactor core design parameters for WCGS Cycle 3, including reactivity coefficients, delayed neutron

fraction, and neutron lifetimes. Sufficient information is included to permit

an independent calculation of the nuclear performance characteristics of the

core.4.3.2.2 Power Distributions The accuracy of power distribution calculations has been confirmed through approximately 1,000 flux maps during some 20 years of operation under

conditions very similar to those expected. 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, relative to the total reactor power

output, 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 power 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 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. Local heat flux is the heat flux at the surface of the cladding (Btu-ft hr-1). For nominal rod parameters, this differs from linear power density by a constant factor. 4.3-8 Rev. 13 WOLF CREEK Rod power or rod integral power is the length integrated linear power density in one rod (kW). 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: F Q , 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. F N Q , 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. F E 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. F E DH , 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

DNBR described in Section 4.4. It is convenient for the purposes of discussion to define subfactors of F Q.However, design limits are set in terms of the total peaking factor.

F Q = Total peaking factor or heat flux hot channel factor

       =Maximum kW/ft Average kW/ft F Q =  F N Q  x  F E Q    F Q= max F N XY (Z)  x P(Z)  x  F N

U x F E Q 4.3-9 Rev. 12 WOLF CREEK where: F N Q and F E Q are defined above F N U = factor for conservatism, assumed to be 1.05 F N XY (Z) = ratio of peak power density to average power density in the horizontal plane of peak local power 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 Distributions The power shape in horizontal sections of the core at full power is a function of the fuel assembly and burnable absorber loading patterns, the control rod pattern, and the fuel burnup distribution. 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. Typical values of radial factor F N XY are given in Table 4.3-2. 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 nonuniform 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. 4.3-10 Rev. 12 WOLF CREEK 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. The radial power

distribution within a fuel rod and its variation with burnup as utilized in thermal calculations and fuel rod design is discussed in Section 4.4. 4.3.2.2.3 Assembly Power Distributions

For the purpose of illustration, typical assembly power distributions from the BOL and EOL conditions are given for the same assembly in Figures 4.3-12 and 4.3-13, respectively. Since the detailed power distribution surrounding the hot channel varies from time to time, a conservatively flat radial 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 N DH. 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 of F N DH.4.3.2.2.4 Axial Power Distributions 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 rods responding to manual operation of the CVCS. Nuclear effects which cause variations in the axial power shape include moderator density, Doppler effect on resonance absorption, spatial distribution

of xenon, and burnup. Automatically controlled variations in total power

output and full length rod motion are also important in determining the axial

power shape at any time. Signals are available to the operator from the excore

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, DI. Calculations of core average peaking factor for many plants and measurements from operating plants under many operating situations are associated with either DI 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 = f t - f b f t + f b and f t and f b are the top and bottom detector readings. Representative axial power shapes for BOL, MOL, and EOL conditions are shown in Figures 4.3-14 through 4.3-16. These figures cover a wide range of axial

offset, including values not permitted at full power. 4.3-11 Rev. 11 WOLF CREEK The radial power distributions involving the partial insertion of control rods represent a synthesis of power shapes from the rodded and unrodded planes. The applicability of the separability assumption upon which this procedure is based

is assured through extensive three-dimensional calculations of possible rodded

conditions. As an example, Figure 4.3-17 compares the axial power distribution

for several assemblies at different distances from inserted control rods with

the core average distribution. The only significant difference from the average occurs in the low power peripheral assemblies, thus confirming the validity of the separability assumption. 4.3.2.2.5 Local Power Peaking

In January 1993 Westinghouse submitted topical report WCAP-13589 (Reference 35) to the NRC; NRC approval of the report was received in January 1995. WCAP-13589 evaluated the densification power spike factor and the clad flattening design criterion based on fuel examination data. This data showed that, for the then current Westinghouse nuclear fuel designs, pellet gaps did not occur which were large enough to permit cladding collapse. The report concluded that a densification power spike factor S(Z) (where Z is axial location in the core) of 1.0 was appropriate for current Westinghouse nuclear fuel designs. Later fuel designs, not covered by the data in WCAP-13589, were evaluated as outlined in Reference 36 to assure that the conclusions of WCAP-13589 applied to those designs as well. The reduced power spike factor of 1.0 is appropriate for use with all Westinghouse fuel in the WCGS core. 4.3.2.2.6 Limiting Power Distributions

According to the ANSI classification of plant conditions (see Chapter 15.0), 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. Inasmuch 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 operational transients is given in Chapter 15.0. 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 actions 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 Conditions II, III, and IV events. 4.3-12 Rev. 11 WOLF CREEK 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 discussed in Chapter 15.0. Therefore, the limiting

power shapes which result from such Condition II events are those power shapes

which deviate from the normal operating condition at the recommended axial

offset band, e.g., 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 Technical Specifications. A complete

discussion of power distribution control in Westinghouse pressurized water

reactors is included in Reference 6. Detailed background information on the

following design constraints on local power density in a Westinghouse pressurized water reactor, the defined operating procedures, and on the

measures taken to preclude exceeding design limits is presented in the

Westinghouse topical report on power distribution control and load following

procedures (Ref. 7). 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, F Q and F N DH , include all of the nuclear effects which 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 including fuel and moderator temperature feedback effects. The steady state nuclear design calculations are done for normal flow with the same mass flow in each channel and flow redistribution effects are 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 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:

a. Core power level

b. Core height

c. Coolant temperature and flow

d. Coolant temperature program as a function of reactor power

e. Fuel cycle lifetimes

f. Rod bank worths

g. Rod bank overlaps 4.3-13 Rev. 11 WOLF CREEK Normal operation of the plant assumes compliance with the following conditions:

a. Control rods in a single bank move together with no individual rod insertion differing by more than 13 steps (indicated) from the bank demand position.

b. Control banks are sequenced with overlapping banks.

c. The control full length bank insertion limits are not violated.

d. Axial power distribution control 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. Briefly

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 equilibrium full power value. 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, peak reactivity, 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 distribution. A finite number of maneuvers each cycle are analyzed to determine the general behavior of the local power density as a function of core elevation. These cases represent many possible reactor states in the life of one fuel cycle, and they have been chosen as sufficiently definitive of the cycle by comparison with much more exhaustive studies performed on some 20 or 30

different, but typical, plant and fuel cycle combinations. The cases are

described in detail in Reference 7, and they are considered to be necessary and

sufficient to generate a local power density limit which, when increased by 5

percent for conservatism, will not be exceeded with a 95-percent confidence

level. Many of the points do not approach the limiting envelope. However, they are part of the time histories which lead to the hundreds of shapes which

do define the envelope. They also serve as a check that the reactor studied is

typical of those studied more exhaustively. Thus it is not possible to single out any transient or steady state condition which defines the most limiting case. It is not even possible to separate out a small number which form an adequate analysis. The process of generating a

myriad of shapes is essential to the philosophy that leads to the required

level of confidence. A maneuver which provides a limiting case for one reactor

fuel cycle (defined as approaching the line of Figure 4.3-21) is not

necessarily a limiting case for another reactor or fuel cycle with different

control bank worths, enrichments, burnup, coefficient, etc. Each shape depends

on the detailed history of operation up to that time and on the manner in which

the operator conditioned xenon in the days immediately prior to the time at which the power distribution is calculated. 4.3-14 Rev. 13 WOLF CREEK The calculated points are synthesized from axial calculations combined with radial factors appropriate for rodded and unrodded planes. In these calculations, the effects on the unrodded radial peak of xenon redistribution

that occurs following the withdrawal of a control bank (or banks) from a rodded

region is obtained. A detailed discussion of this effect may be found in Reference 7. 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 E Q.The envelope drawn over the calculated [max (F Q . Power)] points in Figure 4.3-21 represents 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. For example, Figure 4.3-21 bounds both BOL and EOL conditions but without consideration of

radial power distribution flattening with burnup, i.e., both BOL and EOL points

presume the same radial peaking factor. Inclusion of the burnup flattening

effect would reduce the local power densities corresponding to EOL conditions

which may be limiting at the higher core elevations. Finally, as previously discussed, this upper bound envelope is based on procedures of load follow which require operation within an allowed deviation

from a target equilibrium value of axial offset. These procedures are detailed

in the Technical Specifications and are followed by relying only upon excore

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 linear power at 3565 MWt is 5.68 kW/ft. From Figure 4.3-21, the conservative upper bound value of normalized local power density, including uncertainty allowances, is 2.50, corresponding to a peak linear power of 14.48 kW/ft at 102 percent power. 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. In evaluating these three categories of events, the core is assumed

to be operating within the four constraints described above. The first category comprises uncontrolled rod withdrawal (with rods moving in the normal bank sequence) for full length banks. Also included are motions of the full-length 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 include normal xenon transients. It was

further assumed in determining the power distributions that total core 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. The peak power density which can occur in such events, assuming reactor trip at or below 118 percent, is less than that required for center-line melt, including

uncertainties and densification effects. 4.3-15 Rev. 11 WOLF CREEK The second category assumes that the operator mispositions 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. The results for F Q are multiplied by 102 percent power, including an allowance for calorimetric error. The peak linear power does not exceed the centerline fuel melt kW/ft limit, including the above factors.Since the peak kW/ft is below the centerline fuel melt limit, no flux difference penalties are required for overpower protection. It should be noted that a reactor overpower accident is not assumed to occur coincident with an independent operator error. Additional detailed discussion of these analyses is presented in Reference 7. Analyses of possible operating power shapes show that the appropriate hot channel factors F Q and F H N for peak local power density and for DNB analysis at full power are the values given in Table 4.3-2 and addressed in the Technical Specifications. The maximum allowable F Q can be increased with decreasing power, as shown in the Core Operating Limits Report (COLR). Increasing F H N with decreasing power is permitted by the DNB protection setpoints points and allows radial power shape changes with rod insertion to the insertion limits, as described in Section 4.4.4.3. The allowance for increased F H N permitted is F H N = 1.65 [1 + 0.3 (1-P)]. This becomes a design basis criterion which is used for establishing acceptable control rod patterns and control bank sequencing. Likewise, fuel loading patterns for each cycle are selected with consideration of this design criterion. The worst values of F H N for possible rod configurations occurring in normal operation are used in verifying that this criterion is met. The worst values generally occur when the rods are assumed to be at their insertion limits. Maintenance of constant axial offset control

establishes rod positions which are above the allowed rod insertion limits, thus providing increasing margin to the F H N criterion. As discussed in Section 3.2 of Reference 8, it has been determined that the COLR limits are met, provided the above conditions a and b are observed. These limits are taken as input to the thermal-hydraulic design basis, as described in Section 4.4.4.3.1. When a situation is possible in normal operation which could result in local power densities in excess of those assumed as the precondition 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 Chapter 7.0. 4.3-16 Rev. 12 WOLF CREEK 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 below. It should be noted that power-distribution-related measurements are incorporated into the evaluation of calculated power distribution

information, using an incore instrumentation processing code described in

Reference 9. The measured versus calculational comparison is normally

performed periodically throughout the cycle lifetime of the reactor, as

required by Technical Specifications.

In a measurement of the heat flux hot channel factor, F Q , with the movable detector system described in Sections 7.7.1 and 4.4.6, the following

uncertainties have to be considered:

a. Reproducibility of the measured signal

b. Errors in the calculated relationship between detector

current and local flux

c. Errors in the calculated relationship between detector

flux and peak rod power some distance from the

measurement thimble

The appropriate allowance for category a above has been quantified by

repetitive measurements made with several inter-calibrated detectors by using

the common thimble features of the incore detector system. The WCGS system

allows more than one detector to access any thimble. Errors in category b

above are quantified to the extent possible, by using the detector current

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 absorbers, etc. These critical experiments provide

quantification of errors of categories a and c above.

Reference 2 describes critical experiments performed at the Westinghouse Reactor Evaluation Center and measurements taken on two Westinghouse plants

with incore systems of the same type as used in the WCGS plant. The report

concludes that the uncertainty associated with F Q (heat flux) 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 1.645 limit on a normal distribution and is the uncertainty to be associated with a full core flux map with movable 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.

In comparing measured power distributions (or detector currents) with

calculations for the same operating conditions, it is not possible to isolate

out the detector reproducibility. Thus a comparison between measured and

predicted power distributions has to include some measurement error. Such a

comparison is given in Figure 4.3-24 for one of the maps used in Reference 2.

Since the first publication of Reference 2, 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 Q.

4.3-17 Rev. 23 WOLF CREEK A similar analysis for the uncertainty in F H N (rod integral power) measurements results in an allowance of 3.65 percent at the equivalent of a 1.645 confidence level. For historical reasons, an 8 percent uncertainty factor is allowed in the nuclear design calculational basis; that is, the predicted rod integrals at full power must not exceed the design F H N less 8 percent. A measurement in the second cycle of a 121 assembly, 12 foot, core is compared with a simplified one-dimensional core average axial calculation in Figure 4.3-

25. 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:

a. Much of the data is obtained in steady state operation at

constant power in the normal operating configuration.

b. Data with unusual values of axial offset are obtained as

part of the excore detector calibration exercise which is

performed monthly.

c. Special tests have been performed in load follow and

other transient xenon conditions which have yielded useful information on power distributions.

These data are presented in detail in Reference 8. Figure 4.3-26 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

An extensive series of physics tests was performed on the first core. These

tests and the criteria for satisfactory results are described in Chapter 14.0.

Since not all limiting situations can be created at BOL, the main purpose of

the tests was to provide a check on the calculational methods used in the

predictions for the conditions of the test. Tests performed at the beginning

of each reload cycle are limited to verification of the selected safety-related

parameters of the reload design.

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, 6, and 8. The relevant conclusions are summarized in Sections

4.3.2.2.7 and 4.4.6.

Provided the limitations given in Section 4.3.2.2.6 on rod insertion and flux

difference are observed, the excore detector system provides adequate on-line

monitoring of power distributions. Further details of specific limits on the

observed rod positions and flux difference are given in the Technical

Specifications, together with a discussion of their bases.

Limits for alarms, reactor trip, etc. are given in the Technical

Specifications. Descriptions of the systems provided are given in Section 7.7.

4.3-18 Rev. 11 WOLF CREEK 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 pressure or void conditions, although the latter are

relatively unimportant in the WCGS reactor. 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 Chapter 15.0. The reactivity coefficients are calculated on a corewise

basis by radial and axial diffusion theory methods. The effect of radial and

axial power distribution on core average reactivity coefficients is implicit in

those calculations and is not significant under normal operating conditions.

For example, a skewed xenon distribution which results in changing axial offset

by 5 percent changes the moderator and Doppler temperature coefficients by less

than 0.01 pcm/F and 0.03 pcm/F, respectively. An artificially skewed xenon distribution which results in changing the radial F H N by 3 percent changes the moderator and Doppler temperature coefficients by less than 0.03 pcm/F and 0.001 pcm/F, respectively. The spatial effects are accentuated in some

transient conditions, for example, in postulated rupture of the main steam line

break and rupture of a rod cluster control assembly mechanism housing described in Sections 15.1.5 and 15.4.8, and are included in these analyses.

The analytical methods and calculational models used in calculating the

reactivity coefficients are given in Section 4.3.3. Quantitative information

for calculated reactivity coefficients, including fuel-Doppler coefficient, moderator coefficients (density, temperature, pressure, and void) and power

coefficient is given in the following sections.

4.3.2.3.1 Fuel Temperature (Doppler Power) Coefficient

The fuel temperature (Doppler Power) coefficient (DPC) 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 is also considered but their

contribution 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.

The DPC is calculated by performing three-dimensional calculations using the

ANC computer code (Ref. 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 the

overall value becomes less negative since 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 Chapter 15.0.

4.3-19 Rev. 23 WOLF CREEK 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 degree change in the moderator temperature. Generally, the

effects of the changes in moderator density as well as the temperature are

considered together.

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 density introduces a positive

component in the moderator coefficient. If the concentration of soluble poison

is large enough, the net value of the coefficient may be positive.

With the burnable absorber rods present, however, the initial hot boron

concentration is sufficiently low that the moderator temperature coefficient may be negative at operating temperatures. The effect of control rods is to make the moderator coefficient more negative since the thermal neutron mean

free path, and hence the volume affected by the control rods, increases with an

increase in temperature.

With burnup, the moderator coefficient becomes more negative, primarily as a

result of boric acid dilution, but also to a significant extent from the

effects of the buildup of plutonium and fission products.

The moderator coefficient is calculated for a range of plant conditions by

performing ANC calculations, in which the moderator temperature (and density)

is varied. Typical values for MTC are shown in Table 4.3-6 as a function of

core average temperature, boron concentration, and burnup. Figure 4.3-6 shows

MTC plotted as a function of burnup for conditions of hot full power and just

critical boron concentration. The moderator coefficients presented here are calculated on a corewide 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 entire core.

Moderator Pressure Coefficient The moderator pressure coefficient relates the change in moderator density, resulting from a reactor coolant pressure change, to the corresponding effect

on 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 1/2-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 over a portion of the moderator temperature range at BOL (0.004

pcm/psi, BOL) but is always positive at operating conditions and becomes more

positive during life (+0.3 pcm/psi, EOL).

4.3-20 Rev. 13 WOLF CREEK Moderator Void Coefficient The moderator void coefficient relates the change in neutron multiplication to

the presence of voids in the moderator. In a pressurized water reactor, this

coefficient is not very significant because of the low void content in the coolant. The core void content is less than 1/2 of 1 percent and is due to

local or statistical boiling. The void coefficient varies from 50 pcm/percent

void at BOL and at low temperatures to -250 pcm/percent void at EOL and at

operating temperatures. The 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. It becomes more negative with burnup reflecting the combined effect of moderator and fuel

temperature coefficients with burnup.

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:

a. +10 percent for Doppler and power defect

b. +2 pcm/F for the moderator coefficient

Experimental evaluation of the reactivity coefficients will be performed during

the physics startup tests described in Chapter 14.0.

4.3.2.3.5 Reactivity Coefficients Used in Transient Analysis

Table 4.3-2 gives the limiting values as well for the reactivity coefficients.

The limiting value is used as design limits in the 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 Chapter 15.0.

4.3-21 Rev. 23 WOLF CREEK The limiting values shown in Table 4.3-2 are chosen to encompass the best estimate reactivity coefficients, including the uncertainties given in Section

4.3.3.3 over appropriate operating conditions calculated for this cycle and the

expected values for the subsequent cycles. The most positive, as well as the

most negative, values are selected to form the design basis range used in the transient analysis. A direct comparison of the best estimate and design limit

values can be misleading since, in many instances, the most conservative

combination of reactivity coefficients is used in the transient analysis even

though the extreme coefficients assumed may not simultaneously occur at the

conditions of lifetime, power level, temperature, and boron concentration

assumed in the analysis. 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

Chapter 15.0 with due allowance for the calculational uncertainties given in

Section 4.3.3.3. Control rod requirements are given in Table 4.3-3 for the core described and for an equilibrium cycle, since these are markedly different. These latter numbers are provided for information only since

refueling specifications for subsequent cycles have not yet been established.

4.3.2.4 Control Requirements To ensure the shutdown margin stated in the COLR under conditions where a

cooldown to ambient temperature is required, concentrated soluble boron is

added to the coolant. Boron concentrations for several core conditions are listed in Table 4.3-2; these values were calculated with ANC (Reference 31) for

Cycle 9. For all core conditions including refueling, the boron concentration

is well below the solubility limit. The rod cluster control assemblies are

employed to bring the reactor to the hot shutdown condition. The minimum

required shutdown margin is given in the COLR.

The ability to accomplish the shutdown for hot conditions is demonstrated in

Table 4.3-3 by comparing the difference between the rod cluster control

assembly 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

cross-sections 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.

4.3-22 Rev. 23 WOLF CREEK 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 temperature is conservatively increased to account for the control dead band and measurement errors. Since the moderator coefficient at full-load temperature 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 control 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 toward 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 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 changes in boron concentration, changes in temperature, and very small changes in the xenon concentration not compensated for by a change in boron 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. 4.3-23 Rev. 13 WOLF CREEK 4.3.2.4.6 Burnup The reactor core is composed of an array of fuel assemblies that are similar in mechanical design, but different in fuel enrichment. Within each fuel assembly, all rods are of the same enrichment. Three different enrichments were employed in the first core. Other enrichments are employed in reload fuel. The enrichments for cycle 1 at Wolf Creek were 2.10 (Region 1), 2.60 (Region 2), and 3.10 (Region 3) weight percent. The average enrichment has increased in each successive cycle load in order to achieve an eighteen month cycle. This began in Cycle 2 and Cycle 4 was the first eighteen month cycle. For a 12 month cycle, excess reactivity of approximately 10 percent (hot) is installed at the beginning of the cycle to provide sufficient reactivity to compensate for fuel depletion and fission product buildup throughout the cycle. Excess reactivity of approximately 20 percent (hot) is installed at the beginning of an 18 month cycle to provide sufficient reactivity to compensate for fuel depletion and fission product buildup throughout the cycle. This reactivity is controlled by the addition of soluble boron to the coolant and by

burnable absorber. The soluble boron concentrations for several core configurations are given in Table 4.3-2; these values were calculated with ANC for Cycle 9. Since the excess reactivity for burnup is controlled by soluble boron and/or burnable absorber, 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 can be controlled by changing the soluble boron

concentration (also see Section 4.3.2.4.16). 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 provided in Reference 11. 4.3.2.4.9 Experimental Confirmation

Following a normal shutdown, the total core reactivity change during cooldown with a stuck rod has been measured on a 121 assembly, 10-foot-high core, and 121 assembly, 12-foot-high core. In each case, the core was allowed to cool down until it reached criticality simulating the steam line break accident. For the 10-foot core, the total reactivity change associated with the cooldown is overpredicted by about 0.3 percent with respect to the measured result.

This represents an error of about 5 percent in the total reactivity change and is about half the uncertainty allowance for this quantity. For the 12-foot

core, the difference between the measured and predicted reactivity change was

an even smaller 0.2 percent. These measurements and others demonstrate the ability of the methods described in Section 4.3.3. 4.3.2.4.10 Control Core reactivity is controlled by means of a chemical poison dissolved in the coolant, rod cluster control assemblies, and burnable absorber rods, as described below. 4.3-24 Rev. 11 WOLF CREEK 4.3.2.4.11 Chemical Poison Boron in solution as boric acid is used to control relatively slow reactivity changes associated with:

a. The moderator temperature defect in going from cold shutdown at ambient temperature to the hot operating temperature at zero power

b. The transient xenon and samarium poisoning, such as that following power changes or changes in rod cluster control

position

c. The reactivity effects of fissile inventory depletion and buildup of long-life fission products

d. The burnable absorber depletion The boron concentrations for various core conditions are presented in Table 4.3-2; these values were calculated with ANC for Cycle 9.

4.3.2.4.12 Rod Cluster Control Assemblies

The number of rod cluster control assemblies is shown in Table 4.3-1. The rod cluster control assemblies are used for shutdown and control purposes to offset fast reactivity changes associated with:

a. The required shutdown margin in the hot zero power, stuck rods condition

b. The reactivity compensation as a result of an increase in power above hot zero power (power defect, including Doppler, and moderator reactivity changes)

c. Unprogrammed fluctuations in boron concentration, coolant temperature, or xenon concentration (with rods not

exceeding the allowable rod insertion limits)

d. 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 rod cluster control assembly withdrawal pattern determined from these analyses is used in determining power distribution factors and in determining the maximum worth of an inserted rod cluster control assembly ejection accident.

For further discussion, refer to the COLR on rod insertion limits. 4.3-25 Rev. 13 WOLF CREEK Power distribution, rod ejection, and rod misalignment analyses are based on the arrangement of the shutdown and control groups of the rod cluster control

assemblies shown in Figure 4.3-36. All shutdown rod cluster control assemblies

are withdrawn before withdrawal of the control banks is initiated. In going

from zero to 100-percent power, control banks A, B, C, and D are withdrawn

sequentially. The limits of rod positions and further discussion on the basis for rod insertion limits are provided in the COLR. 4.3.2.4.13 Reactor Coolant Temperature Reactor coolant (or moderator) temperature control has added flexibility in reactivity control of the Westinghouse pressurized water reactor. This feature

takes advantage of the negative moderator temperature coefficient inherent in a

pressurized water reactor to:

a. Maximize return to power capabilities

b. Provide

+-5 percent power load regulation capabilities without requiring control rod compensation

c. Extend the time in cycle life to which daily load follow operations can be accomplished Reactor coolant temperature control supplements the dilution capability of the plant by lowering the reactor coolant temperature to supply positive reactivity through the negative moderator coefficient of the reactor. After the transient is over, the system returns the reactor coolant temperature to the programmed value.Moderator temperature control of reactivity, like soluble boron control, has the advantage of not significantly affecting the core power distribution.

However, unlike boron control, temperature control can be rapid enough to

achieve reactor power change rates of 5 percent/minute. 4.3.2.4.14 Burnable Absorber Rods The standard burnable absorber of WABA rods provide partial control of the excess reactivity available during the first fuel cycle. In doing so, these

rods prevent the moderator temperature coefficient 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 first

fuel cycle, as described previously. For purposes of illustration, a typical burnable absorber rod pattern in the core together with the number of rods per assembly are shown in Figure 4.3-5, while the arrangements within an assembly are displayed in Figure 4.3-4. The reactivity worth of these rods is shown in

Table 4.3-1. The boron in the rods is depleted with burnup but at a

sufficiently slow rate so that the resulting critical concentration of soluble boron is such that the moderator temperature coefficient remains below the safety analysis limit at all times for power operating conditions in the first

cycle. 4.3-26 Rev. 13 WOLF CREEK 4.3.2.4.15 Peak Xenon Startup Compensation for the peak xenon buildup is accomplished, using the boron control system. Startup from the peak xenon condition is accomplished with a combination of rod motion and boron dilution. The boron dilution may be made at any time, including during the shutdown period, provided the shutdown margin is maintained. 4.3.2.4.16 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 on full-length rods, as provided in the COLR and discussed in Sections 4.3.2.4.12 and 4.3.2.4.13. The power distribution is maintained within acceptable limits through location of the full-length rod bank. Reactivity changes due to the changing xenon

concentration can be controlled by rod motion and/or changes in the soluble

boron concentration. Late in cycle life, extended load follow capability is obtained by augmenting the limited boron dilution capability at low soluble boron concentrations by

temporary moderator temperature reductions. Rapid power increases (5 percent/min) from part power during load follow operation are accomplished with a combination of rod motion, moderator temperature reduction, and boron dilution. The rapid power increase is accomplished initially by a combination of rod withdrawal and moderator temperature reduction. As the slower boron dilution takes effect after the

initial rapid power increase, the moderator temperature is returned to the

programmed value. 4.3.2.4.17 Burnup

Control of the excess reactivity for burnup is accomplished, using soluble boron and/or burnable absorber. Sufficient burnable absorber is installed at the beginning of a cycle to give the desired cycle lifetime, without exceeding

the boron concentration limit. The practical minimum boron concentration is in

the range of 0 to 10 ppm. 4.3.2.5 Control Rod Patterns and Reactivity Worth The rod cluster control assemblies 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 rod cluster assembly pattern is displayed in Figure

4.3-36, which is not expected to change during the life of the plant. The

control banks are labeled A, B, C, and D and the shutdown banks are labeled SA, SB, SC, SD, and SE. Each bank, although operated and controlled as a unit, is

composed of two subgroups. The axial position of the control rod banks may be controlled manually or automatically, while the shutdown banks are only controlled manually. The rod cluster control assemblies are all dropped into the core following actuation of reactor trip signals. 4.3-27 Rev. 12 WOLF CREEK 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. Therefore, four banks (described as A, B, C, and D

in Figure 4.3-36) have been selected. Typical control bank worths are shown in Table 4.3-2. 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 COLR). Ejected rod worths are given in Section 15.4.8 for several different

conditions. Allowable deviations due to misaligned control rods are discussed in the Technical Specifications. A representative calculation for three banks of control rods withdrawn simultaneously (rod withdrawal accident) is given in Figure 4.3-37. 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 in Figure 4.3-38. For nuclear design purposes, the reactivity worth versus rod position

is calculated by a series of steady state calculations at various 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 4.3-39.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 rod cluster control

assemblies are tripped, but assuming that the highest worth assembly remains

fully withdrawn and no changes in xenon or boron take place. The loss of

control rod worth due to the material irradiation is negligible, since only

bank D may be in the core under normal operating conditions (near full power). The values given in Table 4.3-3 show that the available reactivity in withdrawn rod cluster control assemblies provides the design bases minimum shutdown

margin, allowing for the highest worth cluster to be at its fully withdrawn

position. An allowance for the uncertainty in the calculated worth of N-1 rods

is made before determination of the shutdown margin. 4.3-28 Rev. 13 WOLF CREEK 4.3.2.6 Criticality of the Reactor During Refueling and Criticality of Fuel Assemblies The basis for maintaining the reactor subcritical during refueling is presented in Section 4.3.1.5, and a discussion of how control requirements are met is given in Sections 4.3.2.4 and 4.3.2.5. Criticality of fuel assemblies outside the reactor is precluded by adequate design of fuel transfer and fuel storage facilities and by administrative control procedures. Sections 9.1.1 and 9.1.2 identify those criteria important to criticality safety analyses. 4.3.2.7 Stability 4.3.2.7.1 Introduction

The stability of the pressurized water reactor cores against xenon-induced spatial oscillations and the control of such transients are discussed

extensively in References 6, 14, 15, and 16. A summary of these reports is

given in the following discussion, and the design 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 temperature distributions. Such a power shift could occur in the diametral plane of the core as a result of abnormal control action. Due to the negative power coefficient of reactivity, pressurized water reactor cores are inherently stable to oscillations in total power. Protection against total power instabilities is provided by the control and protection system, as described in Section 7.7. Hence, the discussion on the core stability will be

limited here to xenon-induced spatial oscillations. 4.3.2.7.2 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 pressurized water reactors and the stability of the core against xenon-

induced oscillations can be determined in terms of the eigenvalues of the first flux overtones. Writing the eigenvalue of the first flux harmonic as: = b + ic [4.3-1] 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 in the power distribution can now be represented as: (t) = A et = ae bt cos ct [4.3-2] 4.3-29 Rev. 11 WOLF CREEK where A and a are constants. The stability index can also be obtained approximately by: b = 1 T ln A n+1 A n [4.3-3] where A n and A n+1 are the successive peak amplitudes of the oscillation and T is the time period between the successive peaks. 4.3.2.7.3 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 for cores of similar size. 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 to

or slightly more negative, and 3) the Doppler coefficient of reactivity is

expected to be equal to or slightly more negative at full power. Analysis of both the axial and X-Y xenon transient tests, discussed in Section 4.3.2.7.5, shows that the calculational model is adequate for the prediction of

core stability. 4.3.2.7.4 Stability Measurements

a. Axial measurements Two axial xenon transient tests conducted in a pressurized water reactor with a core height of 12 feet

and 121 fuel assemblies is reported in Reference 17 and

will be briefly discussed here. The tests were performed

at approximately 10 percent and 50 percent of cycle life. 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 D and the subsequent 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 within the prescribed limits.

The axial offset of power was obtained from the excore

ion chamber readings (which had been calibrated against

the incore flux maps) as a function of time for both

free-running tests, as shown in Figure 4.3-40. 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 squares fit of the axial offset data in the form of Equation [4.3-2]. The axial offset 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: 4.3-30 Rev. 11 WOLF CREEK

1. The core was stable against induced axial xenon transients, both at the core average burnups of 1550

MWD/MTU and 7700 MWD/MTU. The measured stability

indices are -0.041 hr -1 for the first test (Curve 1 of Figure 4.3-40) and -0.014 hr -1 for the second test (Curve 2 of Figure 4.3-40). The corresponding oscillation periods are 32.4 and 27.2 hours, respectively.

2. The reactor core becomes less stable as fuel burnup progresses and the axial stability index was

essentially zero at 12,000 MWD/MTU. However, the

movable control rod system can control axial

oscillations, as described in Section 4.3.2.7.

b. Measurements in the X-Y plane

Two X-Y xenon oscillation tests were performed at a pressurized water reactor 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 12,900 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 pressurized water reactors 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 rod cluster control unit located along the diagonal axis. Following the perturbation, the uncontrolled oscillation was monitored, using the movable

detector and thermocouple system and the excore power

range detectors. The quadrant tilt difference (QTD) 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 QTD data were fitted in the form of Equation [4.3-2] through a least squares method. A stability

index of -0.076 hr -1 with a period of 29.6 hours was obtained from the thermocouple data shown in Figure 4.3-

41.

It was observed in the second X-Y xenon test that the pressurized water reactor core with 157 fuel assemblies

had become more stable due to an increased fuel

depletion, and the stability index was not determined. 4.3-31 Rev. 11 WOLF CREEK 4.3.2.7.5 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 axial offset data was carried out using the PANDA Code (Ref. 18). The analysis of the X-Y xenon transient tests was performed on an X-Y geometry, using a modified TURTLE Code (Ref. 10) concurring with the Advanced Nodal Code (ANC)(Ref. 31). The PANDA, TURTLE, and ANC 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 has evolved from the code ARK(C) which is

essentially a combination of the codes LEOPARD (Ref. 19) and CINDER (Ref. 20).

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 4.3-5. The calculations show conservative

results for both of the axial tests with a margin of approximately -0.01 hr-1

in the stability index. An analytical simulation of the first X-Y xenon oscillation test shows a calculated stability index of -0.081 hr -1 , in good agreement with the measured value of -0.076 hr -1. 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 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 14, 15 and 16. A more detailed description of the experimental results and

analysis of the axial and X-Y xenon transient tests is presented in Reference

17 and Section 1 of Reference 21. 4.3.2.7.6 Stability Control and Protection

The excore detector system is utilized to provide indications of xenon-induced spatial oscillations. The readings from the excore detectors are available to

the operator and also form part of the protection system.

a. 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 excore 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 reduced. 4.3-32 Rev. 11 WOLF CREEK Twelve-foot pressurized water reactor 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 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 reduced.

b. 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 pressurized water reactors has been further verified as part of the startup physics test

program for pressurized water reactor 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, as discussed in Sections 4.3.2.7.4

and 4.3.2.7.5. 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 design. This includes control rod

drop, rod misalignment, and asymmetric loss-of-coolant

flow. A more detailed discussion of the power distribution control in pressurized water reactor cores is presented in References 6 and 7. 4.3.2.8 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 on the pressure vessel irradiation and surveillance program is given in Section 5.3. The materials that serve to attenuate neutrons originating in the core and gamma rays from both the core and structural components consist of the core

baffle, core barrel, neutron pads, and associated water annuli, all of which

are within the region between the core and the pressure vessel. 4.3-33 Rev. 11 WOLF CREEK 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 incore measurements on operating reactors.

Region and rodwise power-sharing information from the core calculations is then

used as source information in two-dimensional S n transport calculations (DOT code) which compute the flux distributions throughout the reactor. The neutron flux distribution and spectrum in the various structural components varies significantly from the core to the pressure vessel. As discussed in Section 5.3, 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:

a. Determination of effective fuel temperatures

b. Generation of macroscopic few-group parameters

c. 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 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. Initial Core: LASER/REPAD Calculations 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 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 gap, 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, such as for use in xenon stability calculations, a more sophisticated temperature model is used which accounts for the effects of fuel swelling, fission gas release, and plastic clad deformation. 4.3-34 Rev. 11 WOLF CREEK Radial power distributions in the pellet as a function of burnup were obtained from LASER (Ref. 22) calculations.

The effective U-238 temperature for resonance absorption was obtained from the

radial temperature distribution by applying a radially dependent weighting

function. The weighting function was determined from REPAD (Ref. 23) 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 was determined by a

convolution of the radial distribution of Pu-240 densities from LASER burnup

calculations and the radial weighting function. The resulting temperature was

burnup dependent, but the difference between U-238 and Pu-240 temperatures, in

terms of reactivity effects, was small.

The effective pellet temperature for pellet dimensional change was that value

which produce the same outer pellet radius in a virgin pellet as that obtained

from the temperature model. The effective clad temperature for dimensional

change was its average value.

The temperature calculational model was validated by plant Doppler defect data, as shown in Table 4.3-7, and Doppler coefficient data, as shown in Figure 4.3-

42. Stability index measurements also provided a sensitive measure of the

Doppler coefficient near full power (see Section 4.3.2.7). It can be seen that

Doppler defect data was typically within 0.2 percent of prediction.

Reload Cores: FIGHT-H Calculations

The FIGHT-H code (Reference 33) performs a calculation of effective

temperatures in fuel rods for use in nuclear design. The fuel model includes

radial variations of heat generation rate, thermal conductivity, and thermal

expansion in the fuel pellet. Pellet-clad gap conductance depends on the kind

of initial fill gas, hot open gap dimension, and fraction of pellet

circumference over which the gap is effectively closed due to pellet cracking.

PHOENIX-P code system, described in Section 4.3.3.2, generates few group cross

sections as a function of burnup, fuel type, and temperature, for use in ANC.

The FIGHT-H code generates temperature dependent number densities for use in

PHOENIX-P. FIGHT-H also generates fuel effective resonance temperatures and

average temperatures as a function of burnup and enrichment for use in PHOENIX-

P. Starting with Cycle 21, fuel temperature coefficients are calculated with NEXUS/PARAGON/ANC (References 4 and 5), as described in Section 4.3.3.2.

Burnup dependence in the fuel model for Doppler defect and coefficients is

based on an empirical model of progressive pellet cracking which was determined

from operating plant measurements.

4.3-35 Rev. 29 WOLF CREEK 4.3.3.2 Macroscopic Group Constants Macroscopic group constants for use in the spatial few group diffusion codes

were generated for the initial core with a linked version of LEOPARD (Reference

19) and CINDER (Reference 20) codes. Cross sections for reloads were previously generated using the PHOENIX-P code (Reference 32). Starting in reload Cycle 21, cross section calculations are done using the PARAGON/NEXUS (References 4 and 5) code package. A description of each code follows.

Macroscopic few-group constants and consistent microscopic cross-sections (needed for feedback and microscopic depletion calculations) are generated for

fuel cells by a version of the LEOPARD (Ref. 19) and CINDER (Ref. 20) 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 (Ref. 24). Fast and thermal cross-

section library tapes contain microscopic cross-sections taken for the most

part from the ENDF/B (Ref. 25) library, with a few exceptions where other data

provided better agreement with critical experiments, isotopic measurements, and

plant critical boron values. The effect on the unit fuel cell of nonlattice

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.

Group constants for 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 (Ref. 26) and AIM (Ref.27) codes to provide a better treatment of self-

shielding in the broad resonance structure of the 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.

4.3-36 Rev. 29 WOLF CREEK Validation of the cross-section method is based on analysis of critical experiments, as shown in Table 4.3-4, isotopic data, as shown in Table 4.3-8, plant critical boron (C B) values at HZP, BOL, as shown in Table 4.3-9, and at HFP as a function of burnup, as shown in Figures 4.3-43 through 4.3-46.

Control rod worth measurements are shown in Table 4.3-10.

Confirmatory critical experiments on burnable poisons are described in

Reference 28.

The PHOENIX-P computer code is a two-dimensional, multigroup, transport based

lattice code and capable of providing all necessary data for PWR analysis.

Being a dimensional lattice code, PHOENIX-P does not rely on pre-determined

spatial/spectral interaction assumption for a heterogeneous fuel lattice, hence, will provide a more accurate multi-group flux solution than versions of

LEOPARD/CINDER. The PHOENIX-P computer code is approved by the USNRC as the

lattice code for generating macroscopic and microscopic few group cross

sections for PWR analysis (Reference 32).

The solution for the detailed spatial flux and energy distribution is divided

into two major steps in PHOENIX-P ( Reference 32). In the first step, a two-

dimensional fine energy group nodal solution is obtained which couples

individual subcell regions (pellet, clad and moderator) as well as surrounding

pins. PHOENIX-P uses a method based on the Carlvik's collision probability

approach and heterogeneous response fluxes which preserves the heterogeneity of

the pin cells and their surroundings. The nodal solution provides accurate and

detailed local flux distribution which is then used to spatially homogenize the

pin cells to fewer groups.

The second step in the solution process solves for the angular flux

distribution using a standard S4 discrete ordinates calculation. This step is

based on the group-collapsed and homogenized cross sections obtained from the

first step of the solution. The S4 fluxes are then used to normalize the

detailed spatial and energy nodal fluxes. The normalized nodal fluxes are used

to compute reaction rates, power distribution and to deplete the fuel and

burnable absorbers. A standard B1 calculation is employed to evaluate the

fundamental mode critical spectrum and to provide an improved fast diffusion

coefficient for the core spatial codes.

The PHOENIX-P code originally employed a 42 energy group library which had been derived mainly from the ENDF/B-V files. Starting in Cycle 11, the PHOENIX code was upgraded to a 70 group library, which was derived from the ENDF/B-VI files.

The PHOENIX-P cross sections library was designed to properly capture integral properties of the multi-group data during group collapse, and enabling proper

modeling of important resonance parameters. The library contains all neutronic

data necessary for modeling fuel, fission produces, cladding and structural, coolant and control/burnable absorber materials present in Light Water Reactor

cores.

Group constants for burnable absorber cells, guide thimbles, instrument

thimbles, control rod cells and other non-fuel cells can be obtained directly

from PHOENIX-P without any adjustments such as those required in the cell or

1-dimensional lattice codes.

PARAGON has been approved by the NRC as the new generation of Westinghouse lattice code (Reference 4). PARAGON is a replacement for PHOENIX-P and its primary use will be to provide the same types of input data that PHOENIX-P generates for use in three dimensional core simulator codes. This includes macroscopic cross sections, microscopic cross sections for feedback adjustments to the macroscopic cross sections, pin factors for pin power reconstruction calculations, discontinuity factors for a nodal method solution, and other data needed for safety analysis or other downstream applications. 4.3-37 Rev. 29 WOLF CREEK PARAGON is based on collision probability - interface current cell coupling methods. PARAGON provides flexibility in modeling that was not available in PHOENIX-P including exact cell geometry representation instead of cylinderization, multiple rings and regions within the fuel pin and moderator cell geometry, and variable cell pitch. The solution method permits flexibility in choosing the quality of the calculation through both increasing the number of regions modeled within the cell and the number of angular current directions tracked at the cell interfaces. The calculation scheme in PARAGON is based on the conventional lattice modules: resonance calculation, flux solution, leakage correction and depletion. The detailed theory of these modules is described in Reference 4. The cross-section resonance calculation module is based on the space dependent Dancoff method (Reference 4); it is a generalization of the PHOENIX-P methodology that permits to subdivide the fuel pin into many rings and therefore generates space dependent self-shielded isotopic cross-sections. The flux solution module uses the interface current collision probability method and permits a detailed representation of the fuel cells (Reference 4). The other two modules (leakage and depletion) are similar to the ones used in PHOENIX-P. The current PARAGON cross section library is a 70-group library, based on the ENDF/B basic nuclear data, with the same group structure as the library currently used with PHOENIX-P. The PARAGON qualification library has been improved through the addition of more explicit fission products and fission product chains (Reference 4). PARAGON is however designed to employ any number of energy groups. The new NEXUS cross-section generation system uses PARAGON as the lattice code (Reference 5). 4.3.3.3 Spatial Few-Group Diffusion Calculations

For the initial core, spatial few-group diffusion calculations primarily

consisted of two-group X-Y calculations using an updated version of the TURTLE

code, and two-group axial calculations were performed using an updated version

of the PANDA code. Spatial few-group diffusion calculations for reload cores

are performed with the ANC code (Advanced Nodal Code) (Reference 31). The

three dimensional nature of ANC provides both the radial and axial power

distributions.

For the initial core, validation of TURTLE reactivity calculations was

associated with the validation of the group constants themselves, as discussed

in Section 4.3.3.2. Validation of the Doppler calculations was associated with

the fuel temperature validation discussed in Section 4.3.3.1. Validation of

the moderator coefficient calculations was 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 and the

radial buckling used in the axial calculation were obtained from the PANDA

radial calculation, in which group constants in annular rings representing the

various material regions in the X-Y plane are homogenized by flux-volume

weighting.

For reload cores, nodal three dimensional calculations are carried out to

determine the critical boron concentrations and power distributions. The

moderator coefficient is evaluated by varying the inlet temperature in the same

calculations used for power distribution and reactivity predictions.

ANC is used in two-dimensional and three-dimensional calculations. ANC can be

used for safety analyses and to calculate critical boron concentrations, control rod worth, reactivity coefficients, etc.

4.3-38 Rev. 29 WOLF CREEK For reload cores, nodal three dimensional calculations are carried out to determine the critical boron concentrations and power distributions. The

moderator coefficient is evaluated by varying the inlet temperature in the same

calculations used for power distribution and reactivity predictions.

ANC is used in two-dimensional and three-dimensional calculations. ANC can be

used for safety analyses and to calculate critical boron concentrations, control rod worth, reactivity coefficients, etc.

Validation of the spatial codes for calculating power distributions involves

the use of incore and excore detectors and is discussed in Section 4.3.2.2.7.

4.

3.4 REFERENCES

1. "Westinghouse Anticipated Transients Without Reactor Trip

Analysis," WCAP-8330, August 1974.

2. Langford, F. L. and Nath, R. J., "Evaluation of Nuclear

Hot Channel Factor Uncertainties," WCAP-7308-L

(Proprietary), April 1969 and WCAP-7810 (Non-Proprietary), December 1971.

3. Letter from T. M. Anderson (Westinghouse) to S. H. Hanauer (USNRC), "ATWS Submittal," NS-TMA-2182, December 1979.

4. Ouisolumen, M. et al., "Qualification of the Two-Dimensional Transport Code PARAGON" WCAP-16045-P-16045-P-A, August 2004.

5. Zhang, B. et al., "Qualification of the NEXUS Nuclear Data Methodology", WCAP-16045-P-A, Addendum 1, November 2005.

6. "Power Distribution Control of Westinghouse Pressurized

Water Reactors," WCAP-7811, December 1971.

7. Morita, T., et al., "Power Distribution Control and Load

Following Procedures," WCAP-8385 (Proprietary) and WCAP-

8403 (Non-Proprietary), September 1974.

8. McFarlane, A. F., "Power Peaking Factors," WCAP-7912-P-A

(Proprietary) and WCAP-7912-A (Non-Proprietary), January

1975.

9. Meyer, C. E. and Stover, R. L., "Incore Power

Distribution Determination in Westinghouse Pressurized

Water Reactors," WCAP-8498, July 1975.

10. Barry, R. F. and Altomare, S., "The TURTLE 24.0

Diffusion Depletion Code," WCAP-7213-P-A (Proprietary)

and WCAP-7758-A (Non-Proprietary), February 1975.

11. Cermak, J. 0., et al., "Pressurized Water Reactor pH -

Reactivity Effect Final Report," WCAP-3696-8 (EURAEC-

2074), October 1968. 4.3-39 Rev. 29 WOLF CREEK 12. Strawbridge, L. E. and Barry, R. F., "Criticality Calculation for Uniform Water-Moderated Lattices," Nucl.

Sci. and Eng. 23, 58 (1965).

13. Dominick, I. E. and Orr, W. L., "Experimental Verification of Wet Fuel Storage Criticality Analyses," WCAP-8682 (Proprietary) and WCAP-8683 (Non-Proprietary), December 1975.

14. Poncelet, C. G. and Christie, A. M., "Xenon-Induced Spatial Instabilities in Large Pressurized Water

Reactors," WCAP-3680-20 (EURAEC-1974), March 1968.

15. Skogen, F. B. and McFarlane, A. F., "Control Procedures for Xenon-Induced X-Y Instabilities in Large Pressurized Water Reactors," WCAP-3680-21 (EURAEC-2111), February

1969.

16. Skogen, F. B. and McFarlane, A. F., "Xenon-Induced Spatial Instabilities in Three-Dimensions," WCAP-3680-22

(EURAC-2116), September 1969.

17. Lee, J. C., et al., "Axial Xenon Transient Tests at the Rochester Gas and Electric Reactor," WCAP-7964, June

1971.

18. Barry, R. F. and Minton, G., "The PANDA Code," WCAP-7048-P-A (Proprietary) and WCAP-7757-A (Non-

Proprietary), February 1975.

19. Barry, R. F., "LEOPARD - A Spectrum Dependent Non-Spatial Depletion Code for the IBM-7094," WCAP-3269-26, September 1963.

20. England, T. R., "CINDER - A One-Point Depletion and Fission Product Program," WAPD-TM-334, August 1962.

21. Eggleston, F. T., "Safety-Related Research and Development for Westinghouse Pressurized Water Reactors, Program Summaries - Winter 1977 - Summer 1978," WCAP-

8768, Revision 2, October 1978.

22. Poncelet, C. G., "LASER - A Depletion Program for Lattice Calculations Based on MUFT and THERMOS," WCAP-

6073, April 1966.

23. Olhoeft, J. E., "The Doppler Effect for a Non-Uniform Temperature Distribution in Reactor Fuel Elements," WCAP-2048, July 1962.

24. Nodvik, R. J., "Supplementary Report on Evaluation of Mass Spectrometric and Radiochemical Analyses of Yankee

Core I Spent Fuel, Including Isotopes of Elements

Thorium Through Curium," WCAP-6086, August 1969.

25. Drake, M. K. (Ed.), "Data Formats and Procedure for the ENDF/B Neutron Cross Section Library," BNL-50274, ENDF-

102, Vol. 1, 1970. 4.3-40 Rev. 11 WOLF CREEK

26. Suich, J. E. and Honek, H. C., "The HAMMER System, Heterogeneous Analysis by Multigroup Methods of

Exponentials and Reactors," DP-1064, January 1967.

27. Flatt, H. P. and Buller, D. C., "AIM-5, A Multigroup, One Dimensional Diffusion Equation Code," NAA-SR-4694, March 1960.

28. "Nuclear Design of Westinghouse Pressurized Water Reactors with Burnable Poison Rods," WCAP-7806, December

1971.

29. 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.

30. Leamer, R. D., et al., "PuO2-UO2 Fueled Critical Experiments," WCAP-3726-1, July 1967.

31. Liu, Y. S., et al., "ANC: A Westinghouse Advanced Nodal Computer Code," WCAP-10965, December 1985.

32. 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. 33. FIGHT-H: 15.0 TR-95-0001 W02: METCOM Manual, Westinghouse Electric Corporation, Rev. 3 10/95. 34. Jackson, E.W., et.al., Qualification of Steady State Core Physics Methodology for Wolf Creek Design and Analysis, TR-91-0018 W01, December 1991. 35. Kersting P.J., et.al., Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel, WCAP-13589-A (Proprietary), March 1995. 36. Letter, N.J. Liparulo (W) to R.C. Jones (NRC),

Subject:

Response to Second Request for Additional Information on WCAP-13589, Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel, AW-94-751, dated November 7, 1994. 4.3-41 Rev. 11 WOLF CREEK TABLE 4.3-1 REACTOR CORE DESCRIPTION Act i ve Co r e Equ i valent d i amete r , i n. 132.7 Act i ve fuel he i ght, i n. 143.7 He i ght-to-d i amete r r at i o 1.0 8 Total c r o ss s ect i on a r ea, ft 2 9 6.0 6 H 2 O/U molecula r r at i o, latt ice, cold 2.41 Reflecto r Th i ckne ss and Compo si t i on Top - wate r plu s s teel, i n. ~10 Bottom - wate r plu s s teel, i n. ~10 Si de - wate r plu s s teel, i n. ~15 Fuel A ss em b l i e s Total Num b e r i n the Co re 193 Fuel A ss em bly Type LOPAR V 5H V 5H V 5H V 5H P+ V 5H P+Z+2 RFA Z+2 and RFA-2 Z+2 (S tanda rd) w/IFM w/IFM & PBG Rod A rr ay 17 x 17 17 x 17 17 x 17 17 x 17 17 x 17 17 x 17 17 x 17 Rod s pe r a ss em b ly 2 6 4 2 6 4 2 6 4 2 6 4 2 6 4 2 6 4 2 6 4 Rod p i tch, i n. 0.49 6 0.49 6 0.49 6 0.49 6 0.49 6 0.49 6 0.49 6 Ove r all t r an s ve rs e d i men si on s , i n. 8.42 6 x 8.42 6 8.42 6 x 8.42 6 8.42 6 x 8.42 6 8.42 6 x 8.42 6 8.42 6x 8.42 6 8.42 6x 8.42 6 8.42 6x 8.42 6 Fuel we i ght, a s U0 2 , l b. pe r 1154 1154 1154 1149 1132 113 8 113 8 a ss em b ly (App r o ximate) Z ir caloy/Z ir lo we i ght, l b. pe r a ss em b ly (App r o xi mate)2 6 4 270 275 27 8 275 274 274 Num b e r of g ri d s pe r a ss em b ly 8 - Type R 8 11 12 12 12 12 Compo si t i on of g ri d s Inconel-71 8 (see note 1)(see note 2)(see note 3)(see note 4) (see note 4) (s ee note 4) We i ght of Z ir caloy/Z ir lo g ri d s i n act i ve co r e r eg ion, N/A 11.70 14. 6 1 14.6 1 14.6 5 14.6 5 14.6 5 l b. pe r a ss em b ly (App r o ximate) We i ght of Inconel g ri d s in act i ve co r e r eg i on, l b. pe r 12.04 2.22 2.22 2.22 2.22 2.22 2.22 a ss em b ly (App r o ximate) Num b e r of gu i de th i m b le s pe r a ss em b ly24 24 24 24 24 24 24 Compo si t i on of gu i de th i m b le s Z ir caloy-4 Z ir caloy-4 Z ir caloy-4 Z ir caloy-4 Z ir lo Z ir lo Z ir lo Num b e r of In s t r ument gu ide 1 1 1 1 1 1 1 th i m b le s pe r a ss em bly Compo si t i on of In s t r ument tu b e Z ir caloy-4 Z ir caloy-4 Z ir caloy-4 Z ir caloy-4 Z ir lo Z ir lo Z ir lo D i amete r of gu i de th i m b le s , 0.450 I.D. x 0.442 I.D. x 0.442 I.D. x 0.442 I.D. x 0.442 I.D. x 0.442 I.D. x 0.442 I.D. x uppe r pa r t (a b ove da s hpot), i n. 0.4 82 O.D. 0.474 O.D. 0.474 O.D. 0.474 O.D. 0.474 O.D. 0.474 O.D. 0.4 8 2 O.D D i amete r of gu i de th i m b le s , lowe r 0.397 I.D. x 0.397 I.D. x 0.397 I.D. x 0.397 I.D. x 0.397 I.D. x 0.397 I.D. x 0.397 I.D. x pa r t (b elow da s hpot), i n. 0.429 O.D. 0.430 O.D. 0.430 O.D. 0.430 O.D. 0.430 O.D. 0.430 O.D. 0.439 O.D. D i amete r of In s t r ument gu i de 0.44 8 I.D. x 0.440 I.D. x 0.440 I.D. x 0.440 I.D. x 0.440 I.D. x 0.440 I.D. x 0.442 I.D. x th i m b le s , full length, i n. 0.4 84 O.D. 0.47 6 O.D. 0.47 6 O.D. 0.47 6 O.D. 0.47 6 O.D. 0.47 6 O.D. 0.4 8 2 O.D. Note (1) E i ght total g ri d s - 1 Inconel Top G ri d, 1 Inconel Bottom G ri d, 6 Z ir caloy M i d G ri d s Note (2) Eleven total g ri d s - 1 Inconel Top G ri d, 1 Inconel Bottom G ri d, 6 Z ir caloy M i d G ri d s , 3 Z ir caloy IFM G ri d s Note (3) Twelve total g ri d s - 1 Inconel Top G ri d, 1 Inconel Bottom G ri d, 6 Z ir caloy M i d G ri d s , 3 Z ir caloy IFM G ri d s , 1 Inconel P r otect i ve Bottom G ri d Note (4) Twelve total g ri d s - 1 Inconel Top G ri d, 1 Inconel Bottom G ri d, 6 Z ir lo M i d G ri d s , 3 Z ir lo IFM G ri d s , 1 Inconel P r otect i ve Bottom G ri d Rev. 21 WOLF CREEK TABLE 4.3-1 (S heet 2) Fuel Rod s Total Num b e r fuel r od s i n the co re 50,952 Fuel Rod Type S tanda r d Pe r fo r mance + (appl i ca b le Fuel A ss em bly Type) (LOPAR and V 5H) (V 5H P+, V 5H P+Z+2 , RFA Z+2 , and RFA-2 Z +2) Out si de d i amete r , in. 0.374 0.374 D i amete r gap, i n. 0.00 6 5 0.00 6 5 Clad th i ckne ss , in. 0.0225 0.0225 Clad mate ri al Z ir caloy-4 Z ir loFuel Pellet s Mate ri al U0 2 si nte r ed Den si ty (pe r cent of theo r et i cal) 95 Fuel en ri chment s (we i ght pe r cent r ange) 2.1-5.0 D i amete r , i n. (Typ i cal) 0.3225 Length, i n. (Typ i cal) 0.3 8 7 Ma ss of U0 2 pe r foot of fuel r od, l b/ft 0.3 6 3 Rod Clu s te r Cont r ol A ss em b l i e s Num b e r of clu s te rs , full length, i n the co r e 53 Neut r on A bs o rb e r Mate ri al Hafn i um Si lve r-Ind i um-Cadn i um D i amete r , in. 0.341 0.341 Den si ty, l b/i n.3 0.454 0.3 6 7 Cladd i ng mate rial Type 304, cold wo r ked s ta i nle ss s teel Type 304, cold wo r ked s ta i nle ss s teel Clad th i ckne ss , i n. 0.01 8 5 0.01 8 5 Num b e r of a bs o rb e r r od s pe r clu s te r 24 24 E x ce ss React i v i ty (In i t i al Co re) Ma xi mum fuel a ss em b ly K (cold, clean un b o r ated wate r) 1.39 Ma xi mum co r e r eact i v i ty (cold, ze r o powe r , b eg i nn i ng of cycle, ze r o s olu b le b o r on) 1.222 Rev. 1 8 WOLF CREEK TABLE 4.3-1 (S heet 3) Deleted Ta b le Rev. 11 WOLF CR EE K TABL E 4.3-2 NUCL E AR D E SIGN PARAM E T E RS Core average linear power, including densification effects, kW/ft 5.68 Total Heat flux hot channel factor, F Q 2.50 Nuclear enthalpy rise hot channel factor, F N 1.65 Reactivity Coefficients (a) Design Limits (Sheet 1 Only) Doppler-only power coefficients, see Figure 15.0-2 (pcm/%power)(b) Upper curve -19.4 to -12.7 Lower curve -10.1 to -6.7

Doppler temperature coefficient (pcm/F)(b) -3.5 to -1.0 Boron coefficient (pcm/ppm)(b) -16 to -5 Moderator temperature coefficient (pcm/F)(b) +6 to -50 Rodded moderator density (pcm/gm/cc)(b) <47,000 Boron coefficient for boron dilution (pcm/ppm)(b) Modes 1, 2, and 3 -12.5 Modes 4 and 5 -14.0 Rev. 13 WOLF CR EE K TABL E 4.3-2 (Sheet 2) Delayed Neutron Fraction and Prompt NeutronLifetimeeff Limits: maximum at BOL, (minimum at E OL) 0.0075 (0.0044) 1* microsecond (Typical reload) 12.7 (BOL) Control Rods Rod requirements See Table 4.3-3 Maximum bank worth, pcm <2000

Maximum ejected rod worth See Chapter 15.0

Bank worth, HZP no overlap (pcm)(b) BOL, Xe free (Reload) E OL, Xe = E qu. (Reload) Bank D 606 571 Bank C 645 999

Bank B 445 716

Bank A 396 5351

Typical Radial Factor (BOL to E OL) Unrodded 1.37 to 1.28

D bank 1.50 to 1.45

D + C banks 1.60 to 1.45

D + C + B banks 1.80 to 1.55 Rev. 13 WOLF CR EE K TABL E 4.3-2 (Sheet 3) Boron Concentrations Zero power, k eff = 0.99, cold (e) 2050 Typical Reload, all RCCA out, No Xenon, 3000 MWD/T Zero power, k eff = 0.99, hot (f) 2310 Typical Reload, all RCCA out, No Xenon, 3000 MWD/T Design basis refueling boron concentration See COLR Zero power, k eff < 0.95, cold (e) 1960 Typical Reload, all RCCA out, No Xenon, 3000 MWD/T Zero power, k eff = 1.00, hot (e) 2150 Typical Reload, all RCCA out, No Xenon, 3000 MWD/T Full power, k eff = 1.0 1950 Typical Reload, all RCCA out, No Xenon, 3000 MWD/T Full power, k eff = 1.0, 1550 Typical Reload, all RCCA out, No Xenon, 3000 MWD/T Reduction with fuel burnup

First cycle (ppm/GWD/MTU)(c) See Figure 4.3-3 Reload cycle (ppm/GWD/MTU) ~ 100 Rev. 13 WOLF CR EE K TABL E 4.3-2 (Sheet 4) NOT E S: (a) Uncertainties are given in Section 4.3.3.3.(b) 1 pcm = (percent mille) 10 -5 where is calculated from two statepoint values of k eff by In (k 1/k 2).(c) Gigawatt day (GWD) = 1000 megawatt day (1,000 MWD). During the first cycle, fixed burnable poison rods are present which significantly reduce the boron letdown rate compared to reload cycles.(d) Deleted (e) Cold means 68°F, 1 atm.(f) Hot means 557°F, 2250 psia. Rev. 11 WOLF CR EE K TABL E 4.3-3 R E ACTIVITY R E QUIR E M E NTS FOR ROD CLUST E R CONTROL ASS E MBLI E S E nd-of-Life Reactivity E ffects Beginning-of-Life E nd-of-Life (Typical (percent) (First Cycle) (First Cycle) Reload Cycle)

1. Control requirements Fuel temperature, Doppler (%) 1.37 1.21 Moderator temperature (%)(a) 0.15 1.15 Redistribution (%) 0.50 0.85 Rod insertion allowance (%) 0.50 0.50

                                                             -----                 -----               -----

2. Total control (%) 2.52 3.71 3.29

3. E stimated rod cluster control assembly worth (53 rods)

a. All full length assemblies inserted

(%) 7.54 7.42 Not Required

b. All but one (highest worth) assemblies inserted (%) 6.46 6.39 5.74 Rev. 13 WOLF CR EE K TABL E 4.3-3 (Sheet 2)

E nd-of-Life Reactivity E ffects Beginning-of-Life E nd-of-Life (Typical (percent) (First Cycle) (First Cycle) Reload Cycle)

4. E stimated rod cluster control assembly credit with 10 percent adjustment to accomodate uncertainties, 3b - 10 precent (%) 5.82 5.75 5.17

5. Shutdown margin available 4-2 (%) 3.30 2.04 1.88 (b)NOT E S: (a) Includes void effects.(b) The design basis minimum shutdown is 1.3 percent.

Rev. 13 WOLF CR EE K TABL E 4.3-4 B E NCHMARK CRITICAL E XP E RIM E NTS Description of Number of L E OPARD k eff Using E xperiments

• E xperiments

E xperimental Bucklings U0 2 Al clad 14 1.0012 SS clad 19 0.9963 Borated H 2 0 7 0.9989 Subtotal 40 0.9985 U-Metal Al clad 41 0.9995

Unclad 20 0.9990 Subtotal 61 0.9993 Total 101 0.9990

• Reported in Reference 12.

Rev. 0 WOLF CR EE K TABL E 4.3-5 AXIAL STABILITY IND E X PR E SSURIZ E D WAT E R R E ACTOR COR E WITH A 12-FOOT H E IGHT Burnup C B Stability Index (hr -1) (MWD/MTU) F Z (ppm)

E xp Calc 1550 1.34 1065 -0.041 -0.032 7700 1.27 700 -0.014 -0.006 5090

• -0.0325 -0.0255 RADIAL STABILITY IND E X 2250
  • -0.068 -0.07
• 4-loop plant, 12-foot core in Cycle 1, axial stability test.
  • 4-loop plant, 12-foot core in Cycle 1, radial (X-Y) stability test.

Rev. 0 WOLF CR EE K Table 4.3-6 ARO Moderator Temperature Coefficients versus Average Moderator Temperature and Burnup Average Moderator Temperature (Deg. F) Power: 0 25% 50% 75% 100% Core Avg Temp.: 557 565.3 573.4 581.4 589.1 GWD/T PPM RCS Avg Temp.: 557 564.3 571.6 579.1 586.5 0.150 2500 4.86 3.89 2.72 1.39 -0.22 0.150 2000 1.20 0.04 -1.35 -2.90 -4.74 0.150 1500 -2.90 -4.28 -5.87 -7.66 -9.72 3 2500 5.60 4.55 3.28 1.82 0.04 3 2000 1.79 0.53 -0.97 -2.66 -4.67 3 1500 -2.50 -3.97 -5.70 -7.61 -9.84 10 1500 -4.83 -6.68 -8.81 -11.15-13.86 10 1000 -9.74 -11.80-14.13-16.67-19.52 10 500 -15.21 -17.49-20.02-22.76-25.78 21.4 1000 -12.28 -14.72-17.51-20.64-24.08 21.4 500 -17.85 -20.51-23.52-26.78-30.22 21.4 0 -24.22 -27.12-30.32-33.65-37.15 ARO, HFP E quilibrium Xenon Rev. 16 WOLF CREEK TABLE 4.3-7 COMPARISON OF MEASURED AND CALCULATED DOPPLER DEFECTS Core Burnup CalculatedPlant Fuel Type (MWD/MTU) Measured (pcm)

• (pcm) l Air-filled 1800 1700 1710 2 Air-filled 7700 1300 1440

3 Air and 8460 1200 1210 helium-filled

• pcm = 10 5 x ln (k 1/k 2) Rev. 0 WOLF CREEK TABLE 4.3-8 SAXTON CORE II ISOTOPICS ROD MY, AXIAL ZONE 6 (a) LEOPARDAtom Ratio Measured 2Precision (%) Calculation U-234/U 4.65 x 10

-5 +29 4.60 x 10 -5 U-235/U 5.74 x 10 -3 +0.9 5.73 x 10 -3 U-236/U 3.55 x 10 -4 +5.6 3.74 x 10 -4U-238/U 0.99386 +0.01 0.99385 Pu-238/Pu 1.32 x 10 -3 +2.3 1.222 x 10 -3Pu-239/Pu 0.73971 +0.03 0.74497Pu-240/Pu 0.19302 +0.2 0.19102 Pu-241/Pu 6.014 x 10 -2 +0.3 5.74 x 10 -2 Pu-242/Pu 5.81 x 10 -3 +0 9 5.38 x 10 -3 (b)Pu/U 5.938 x 10-2 +0.7 5.970 x 10 -2 Np-237/U-238 1.14 x 10 -4 +15 0.86 x 10 -4 Am-241/Pu-239 1.23 x 10 -2 +15 1.08 x 10 -2 Cm-242/Pu-239 1.05 x 10 -4 +10 1.11 x 10 -4 Cm-244/Pu-239 1.09 x 10 -4 +20 0.98 x 10 -4 NOTES: (a) Reported in Reference 29.(b) Weight ratio. Rev. 0 WOLF CREEK TABLE 4.3-9 CRITICAL BORON CONCENTRATIONS, HZP, BOLPlant 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 Rev. 0 WOLF CREEK 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 group 3050 2917 ESADA Critical

• , 0.69" Pitch,2 w/o PuO 2 , 8% Pu-240, 9 Control Rods6.21" rod separation 2250 2250 2.07" rod separation 4220 4160

1.38" rod separation 4100 4019

• Reported in Reference 30.

Rev. 0 WOLF CREEK TABLE 4.3-11 COMPARISON OF MEASURED AND CALCULATED MODERATOR COEFFICIENTS AT HZP, BOL Plant Type/ Measured iso* Calculated isoControl Bank Configuration (pcm/F) (pcm/F)_____ 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 -8.70 -8.12 steps 4-loop, 193 assemblies, 12-foot core ARO -0.52 -1.2

D in -4.35 -5.7

D and C in -8.59 -10.0 D, C, and B in -10.14 -10.55

D, C, B, and A in -14.63 -14.45

• Isothermal coefficients, which include the Doppler effect in the fuel.iso= 10 5 ln k 2 k 1/T°F Rev. 0 R p N 1 2 L 28 FEED 3 18 FEED 4 I 5 6 I I FEED : !8 FEED 28 I 7 18 18 18 ,() 8 90 18 FEED 28 9 FEED 18 FEED to 28 FEED 18 :11 18 FEED :12 :28 FEED :13 28 :14 '15 M L j28 i I 18 FEED FEED 18 28 FEED I 18 I FEED FEED 18 2B FEED FEED 18 18 FEED FEED 18 28 FEED FEED 18 18 FEED 2B Wolf Creek K J 1-f G FEED 18 18 FE I FEED 28 I 1 18 FE I I ! 28 FE I FEED 18 28 18 FEED iEI 28 FE FEED 18 18 FF FEED 2EI 18 H FEED 1B lB Once Burnt 2B Twice Burnt F E 0 G B A I r.---i I 18 FEED 28 EJJ . t8 I I FEED , -+EED 128 rli] 18 128 ! FEED 18 FEED 18 FEE! f: I 18 FEED 28 FEED j18 --

FEED 28 18 r-* -- 18 FEED 18 FEED 28 -* ---28 FEED 18 FEEiJ 18 FE ED ,_.. ---FEED 18 FEED 18 FEED 28 -* 18 FEED 2B FEED 18 r-* FEED 18 FEED FEEIJ 2B 18 FEED 18 2B r-* --FEED 2B '-* Rev. 4 WOLF CFIEEK UPDATED SAFETY ANALYSIS REPOR'r FIGURE 4.3-1 TYPICAL CORE LC>ADING PATTERN __ j 9 8 :::::1 ? I-::::E *-01 ::..: <(/) 6 ILLJ 10-*0 II-0 ,. <:1) ,) l:l::: l.LJ :x: c;,:l ::t: l.L. <:::> :z: a <:::> I f-<J ::::> I c::a <:::> 2 c:.:: CL I 0 0 8 \illOLE' CREEK ..,. ..... -*-----0 8 =) -f-*--*-(, *-12 (/') U.J 0... Cl 1-.. C> *-16 (I') -*' U..l =) *-20 1..1... 1..1... C> z: C> -24 1-0.... :=> (I') z: **28 0 <-:> .--

---

.... --*32 .......... . . ,.,_...-__ I _L_I --36 16 20 28 3:2 36 40 BURNUP (GWD/MTU)

Rev * () ....... WOLF C:REEK OIPm'-TED SAFETY ANALYSIS REPORT FIGURE 4.3-2 P R 0 D U C T I 0 N A N 0 C 0 N S U M P T Li II F HIGHER ISOTOPES 2000 --------*---------*------------------Q,, 1600 :z 0 <( 0::: 1Z 1200 lJ.J u :z 0 u 800 a:: 0 cc --I <( u l.j.00 r-ea:: u 0 0 NO'E: liOT. FULL POWER. RODS OUT BURNUP DIFFERENCE 2000 l.j.()OO 6000 8000 I 0, 000 12, 000 ltj., 000 16, 000 18 . 000 CORE AVERAGE BURNUP (MWD/MTU) Rev. 0 FIGURE '-L3-3 WOLF UPDATli!!D SAFETY ANALYSIS REPORT BORON CONCENTRATION VERSUS CYCLE BURNUP WITH AND WITHOUT I __ BURNABLE PorsoN Roos _ WOLF CREEK I16 IFBA Pattern I32 IFBA Pattern I64 IFBA Pattern - IFBA I- INST. TUBE- GUIDE TUBE I80 IFBA Pattern I104 IFBA Pattern I128 IFBA Pattern Rev. 16 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.3-4 TYPICAL INTEGRAL FUEL BURNABLE ABSORBER ROD ARRANGEMENT WITHIN AN ASSEMBLY WOLF CREEKRPNMLKJHGFEDCBA 13280328032 210480104 6SS10480104 31048010480104 48010410480 53280104801048032 680104104808010410480 7 90 o (N)32104808010432 880104104808010410480 93280104801048032 108010410480 111048010480104 1210480104 6SS10480104 133280328032 14 15 0 o (W)Rev. 13WOLF CREEKUPDATED SAFETY ANALYSIS REPORTFIGURE 4.3-5TYPICAL INTEGRAL FUEL BURNABLE ABSORBER ANDSOURCE ASSEMBLY LOCATIONS MTC(pcm/F) 0 0 I 0 I 0 0 I I 0 I I -s ..... *****-:** ........

• • '! ...... ---. *r ------*****; .............
• r* --. *------;--------*

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• i I I I I I o I I 0 I I 0 I I I 0 0 I *25 . -----..** -:----.... --*-:-------... --** ---. ****: ...........

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.. -..... -:--* .. --.... . I 0 I I I I I I ! ' i ' ' ; *30 .......... ..;. .......... .. --*-----**; ............ r-..

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o o I I I t o o I I o I I o o o o I I t : : : : : : : : : : ' . . . ' ' . ' ' . I I o o o I I 0 I 0 l l \ \ \ \ \ \ l _, ----------*----

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: : : t : : ! ! 0 o 0 0 I I I 0 0 0 o I I 0 I I 0 I I o I I I o I o I 0 0 o 0 I 0 I I 0 I -40 * **--* *------*-----* * * :-----.. ---* r ** * * *-----1 * ** ** ** .. ** ** :* * * *--*-**T* * * ------*l** **-* **--r *---------1-** * ** ----* 1--------*---! i i j i j

_____________________ __ _._ 0 4 6 I 110 17. 14 16 11 20 2Z GWUtr Rev. 11 -------*------------- WOLF CREEK UPDATED SAFETY ANALYSI:S REPORT -------------------- FIGURE 4.3-6 MTC VS BURNUP at HFP, ARO CRITICAL CONDITIONS (TYPICAL)

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-I o 02 I ,02 I o03 I ,02 I ,05 I. 08 I, 02 I. 06 I. 12 I .. 03 I. 08 I. llt I. 17 I. 16 1.03 I o 10 I. 17 I o 18 1 .. 03 I .08 1.. llt I. llf I. 16 1.03 1.08 1 .. llt I

• llf I. 16 I. 03 I. 10 I. 16 I. 18 1.03 1.09 I. llt I. llf I. 16 1.03 1.09 I. 15 I. 15 I. 16 I.Olt I. II X I o 18 I o 18 I.Olt 1.09 I. 15 I. 18 I. 17 I oOlt I o08 I o llt I. 18 I oOlf I o06 I o 10 I. llf I. 16 I "Olt I. 05 I o07 I o 08 I o 10 I o Olt I.OLt I
• 0 Lt I.OLt !. 05 WOLE' CHEEK I. 18 I I. 19 I V<l I. 19 I I. 19 I I. 19 I I. 18 I XI I. 12 I I. 05 I ' I. 18 I. I B I. 18 I. 20 I. 20 D>< I. 17 I. 17 I. 20 I. 18 I o 16 I. 16 I. 19 I. 19 X I. 16 I. 16 I. 17 I. 15 I. II I -.10 1.10 1.12 I. I 0 I. II I. I 2 I
• I I I .09 1.08 I. 0 7 .05 1.05 1.05 I. o:, 1.05 I. 06 I. 06 I .06 I .06 I. 06 Rev. rJ WOLF CR:EEK UPDATED SAFETY ANALYSIS REPORT FIGURE '1.3-12 RODI.,!ISE POWER 0ISTIHBUTION I'J II, TYPICAL ASSEMBLY NEAR OF-LIFE, HOT FULL POWER. EQUILIBRIUM XENON, UNRODDED r---0.98 0.98 0.99 0.98 1.00 I. 03 0.99 I. 0 2 I. 07 I. 00 I. 03 1.09 I. II I. 09 1.0 I I. 06 X I. II 1.11 1.00 I. 08 I. 07 1.08 I. 00 1.08 I. 07 1.08 I. 0 I 1.06 X I .09 I. 10 I. 00 I. 1.08 I. 07 1.08 1.00 I. 1.08 1.07 I. 08 I. 0 I I. 06 I. II I. II I .00 I. 03 1.09 I. II 1.09 0.99 I. 02 1.07 I. II 0.98 1.00 I. 03 1.07 1.09 0.98 0.99 I. 00 I. 02 I. 03 0.98 0.98 0.98 0.99 1.00 WOLE' CREEK -I. 10 I. 08 -* I. 10 I. 08 1.08 I. 110 I. 10 I. 10 I. 08 I .08 I. 10 1.08 1.08 X I. 110 I. 10 I. II I. 08 1.08 I. II I. 07 I .07 I. 08 I .08 I. 06 I. I. I. 0 I I. 00 I. 00 >< I. 10 1.08 I. 10 I. 08 I. 08 [>( I. 10 I. 10 I. 10 I. 08 I. 08 1.11 I. 09 I. 09 1.07 I. 07 I. II I. II 1.08 1.08 [X 1.09 I. 07 I. 03 _,_ I. 06 I. 06 I.O:l 1.02 1.00 0.98 I. 0 I 1.00 1.00 I. 01 I . 00 0.99 0.98 0.'38 0 0 .;1 Rev. 0 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT 1+.3-13 RODWISE POWER UlSTRIBUTION JN A TYPICAL AssEMBLY NEAR l[FE, HOT FULL POWER, XENON, UNHODDED CORE 2.JI I . 5 a:: w 3: 0 CL a I. 0 w N ......1 <t :::E: a:: 0 :z: 0.5 0.0 0 0. I p = D = AO = 0.2 WOLF CREEK P = N 0 R MA l. I Z E D POWER p :: P = 0. I AO = -8.f:l D = I'RA.CT I L INSERT I ON OF BANK D A 0 = A X I A L I' F S ET ( P E R C EN T) 0.3 O.lt 0.6 1,---p = 'J.) I D = 'J. I A'J = 5. 0.7 0.8 0.9 FRACTION OF ACTIVE CORE HEIGHT FRGH BOTTOM Rev. 0 WOLF CREEK UPDA'rED SAFETY RE:PORT FIG U R E 1-L 3-1 4 TYPICAL AXIAL POWER SHAPES 0 C C U R R I N G AT 8 E G I N N I N G-0 F --I_. I F 1:: -----*----------------------

I. 0 0::: w 3: 0 Cl... 0 w N _J <( :::E: 0::: 0 z: VJOLF

2. 0 I. 5 I. 0 0.5 p = 0. 5 D = 0. 55 AO =-26.03 P = NORMH I ZI:D POWEll p D AO = 0.5 -0.21 *-8.20 0 = FIIACTION,H.

INSEilTION OF BANK 0 AO = AXIAL. OF=SET (Pi:RCENT)

0. 0 L----'--------L--.j_

____ _j_ _ _j_ _ __ _J...-._, 0.0 0.1 0.2 0.3 0 0.6 0.7 0.8 0.9 I. 0 FRACTION OF ACTIVE CORE HEIGHT FROM BOTTOM Rev. D WOLF CREEK OPDATED SAFETY ANALYSIS REPORT FIGURE 4.3-15 TYPICAL AXIAL POWER SHAPES OCCURRING AT

---

2.0 I . 5 0::: L..I..J a.. Cl .0 _J <( ::::::: 0::: C> z 0.5 0 0. I WOLF CREEK . ,r-P = 0.'5 I D = 0.62 AO =-25.7 P = ZED POWER D =FRACTIONAL INSERT',ON OF llANK 0 AO = AXIAL OFFSU (I)EI\CENT) 0.2 0.3 0 4 0.5 0.6 O.l 0.8 0.9 I. 0 FRACTION OF ACTIVE CORE HEIGHT FROM BOTTOM Hev .* 0 WOLF CREEK UPDATI!:D SAFETY ANALYSIS REPORT FIGURE Y.3-16 TYPICAL AXIAL POWER SHAPES OCCURRING AT END-OF-LIFE cr: w 3: 0 CL w > < I. 5 1.0 < 0.5 0.0 CORE WOLF COKl AV£ RAGE ADJACENT IO D BANK RlMOVED fROM 0 PERIPHf>-;M lOW POW:R ll.SSt!18LY CORE HEIGHT CORE TOP WOLF CREEK Rev. 0 SAFETY ANALYSIS REPORT FIGURE q_3-17 I i 1_._.: DISTRIBUTION OF A TYP AXIA:... PGwER CORt l\VERAGE CAL b.A.NK D SLIGHTLY ..L..I'HJL...!\ I l..U T (' [" f) T c-f"\ ,. K(Z)-NORMALIZED PEAKING VS. CORE HEIGHT 1.2 ,...----.....------------___,..-----------, 1 t:!. &! .... Q: 0.8 I-***********************************<******************* (.) (!) 2 i: 0.6 L5 a.. c 0.4 I-********************** < Q: 0 2 0.2 !-*************

• ----**.-******************

---

_j

....__ .___1--/o.. 10 12 0 0 2 4 6 8 CORE HEIGHT (FT) F=QT'=Z.50 Elevation (fl:) 0.0 6.0 12.0 k(z) 1.0 1.0 0.925 Rev. 7 WOLFCREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.3-21 MAXIMUM FaX POWER VER:SUS AXIAL HEIGHT DURING NORMAL OPERATION 0.759 0. 79 2 Lt. 35 o. 77Lt I o 255 0.800 I. 2Lt9 3.1.1% -0.5% I. 229 I. 225 -0.3% I. I 09 I. 077 I o I 07 1.092 --0.2% I ,ll% I. 20 2 I. 170 -2.7% 0.523 I. 217 I. 221 o. 5lt8 I. 203 I. 233 L+.6% I

• I. 0 il, I. 229 I. 189 -3. 3*% WOL,E' CREEK r-*-* -* -* . -* . -* . -* I" 2:23 I" * ;7% . . . L-.. --I, 2 I 7 I. 22L+ O.S -----I. 229 I. 2 18 -0,. 9% -----I. 2 17 I. 2 I 0 -0" 6 !1----I. 229 I. 220 -0.7% I. 217 ...,_

I. 2 I I -0. DIFFF.RENCE I _ PEAKING FACfORS F l. 5 z: FN I . 357 6H 2.07 Rev. 0 r-*---------------------------------- WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGUI"XE '1.3-2L! COMPARISON QE*rwEEN CALCULATED AND MEASURED RELATIVE FUEL ASSEMfLY POWER = LJ..l 3: 0 1.4 1.2 1.0 .0.8 LJ..l > t-< __J LJ..l 0::: __J <t >< <( 0.6 0.4 0.2 0 BOTTOM 10 20 30 WOLF CREEK BOL NO XENON POWER AT 70 BANK D I NSI::R I EO CORE AVERAGE 40 50 60 PERCENT OF CORE HEIGHT 70 80 90 WOLF CREEK !00 TOP Rev= 0 UPDATED SAFETY ANALYSIS REPORT FIGURE Lf.3-25 LOMPARISH C1LCULATEO M u r:l A 0 U h t U l A C H A P E .::; ',{<) (. f C REf: K X REACTOR I 0 iUACfO.: u. \1 RE.ACTOR 2

• I{EAC fOF J:) 0 e 3 FQ 0 0 3.0
• 0 * * \7 2.5 \1 'VI< o*
• X 0 \1 0 *
• X \1 X X * *o*
• X X
• i'. . 0 X
• xx
• 0 0 \l,v.sz.o.

0

• 0 oo * **a:9
• 0 6lo 0 !:6_ I. 5
• L I I *r *, 4-5 35 25 15 5 0 5 10 IS 2.0 25 30 AXIAL OFFSET (%) RP.V. () WOLF CREEK OPDATED SAFETY ANALYSIS REPORT FIGURE Lj
• 3-t1E AS U R E 0 VALUES OF ... FOH F U L 1. ro POWER ROD CONFIGJRATIONS R p SA SA D sc B SB c SE SB B SD SA D SA CONTROL BANK A B c D TOTAL WOLF CREEK L K J H G F I 180° I B c B SD SB Sg SE c A c A D A c A c SE sc SB Sg B c B oo NUMBER OF RODS 4 8 8 5 25 E sc so SHUTDOWN BANK D C SA D SD SB SE SB sc D SA 13,557-29 B A I ----*---2 1-r--------3 SA ' B c B SA NUMBER OF RODS 5 6 7 -270°-8 ---10 I i 12 13 14 15 TOTAL 28 Rev
• 0 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT F:i:GURE 4.3-36 Roo CLUSTER CONTROL ASSEMBLY PATTERN Wolf c:reek I' I I -*-I ,.' ,,.* . .J.i ' : ,, I I II 'I ,11 1 , :fl I I* T,1l] i 'II I I I ' 1,1' 'I I 'I I *' 'I I I I, I ,, ,, ' I ! [1 _ _:, I;: '1'1 : I I , ' I Jll: 'I /iii),: I /I' i i ,, I I I *I,' ,, j ! I I I , I* r .. ,. :I , :I I , r 1\ : ,, ' *--2 2000 (,J a. ...., li :I, ' I -*-! rr I i, : I : *1 I .,.;...--: 'I ' I : I ' II; '""\ I J..-I 7 ... \: I' " II -*-{I ' i ' 'I* I :{1 ! ' *I or: ' Iiiii,, ! I / If* I I I'  :: ,_ l!
• s ISOO a:: -' < a:: c.;, T:T*** j ,I i i [:II ' ',, I ' I I Ill : I I ' " I ,, i!' '/ I I, I ' ' ' -c, ' l '-r---.* ' ,. 'I'.., -* I I I I I, 'li /"4. I IIi : I I ,, *  : :' I ' ,, rl' ITf I li I, *, I' 1:1 ' ' : ' ' I I ] 11 1 , T: I "* , I ., ,, WJ -z 1000 ' *I T* r 1 ... \i* I I ,, *-' ,, 'I "\. ' ',, I' , I ' ' I /I *-1il ! ! i I I I i' i 'I , ,, I II il: I lj1 Ill ' i ! I  ! :: 111: . 1: II ' ':1;11 ,i : \J! i I i [I[ /: I I', I*' I' ill ,, J, I 1: 1 !I [I[*-' f1 i 'I :I:

[II! Jl: I i I i!: I i ,I! :I* l!illl i fliT I 11"T1 1 i: I I I II \I. II: :,ill: ,lj I J:' :j;l ii llii I II I 'IIi\ i il ! ' II II;:. I l ' I '...J. :l'['i: *!I, I' !Ill J I T .... II[ I I I : II )! '.I li 'I* *, i: 1 1 1 i :: l T i I )I ! i I ll1 i 1: I i I i 11":71 ill 'I I' *Tf ,:Til: l I ' *II 11': T1T11: T"""' A'fjll I I I I I ' I I ::I /: il/\1 I IIi ........,-,II i I I I' I:.;,' ! 11 ,!, !II , I, ' I 1 1' 1 1 ' 0 *140 ----*120 ... .... 0 .. * ,. STEPS II -ill :I I 1$.0 ,I I ,I I ! !TTl 1 I **/, *' ' ! il !'1 I I I l,i, 1%.$ ! ! II I -----f-i-;----IO.CI --I --------__;__ *: 1 .so l\. li*/. ' f1: -I 5.00 li :!1 i I' I,: ' ' II! i ,, I %.!10 *I i I ' ""'"':' -1"'\. I I ill! .. I I X -O.CM:IO l .. E ____ ___ :.::Jm 3"""' a : :11 : aa 1 : ll DIFFERENTIAL ANO INTEGRAL ROCI VERSUS STEPS Wll'HDRAWN AT EOL HZP. HFP EQ XEN0 1 N, BANKS D. C. AND 8 MOVING 113 lHEP OVERLAP .-. a. ... -\l'l ........ 2 (.,) a. ._ :: -Q:: 0

• Q 0 Q:: ..... c -z ... Q:: ... Q llnv. 4 WOLF CREE:K UPDATED SAFETY ANill.YSIB REPORT FIGUIH: 4.:3-37 EXAMPLE DIFFERENTIAL IIHEGRAL RC:I:I WORTH VERSUS STEPS WITHDRI'WN A1' MOL, HZP, llFP, EO XENON BANKS, D, C, AN[) B !MOVING WITH 113 STEP OVEBLAP 160 150 140 130 120 110 Yi 100 w. J: (.) 90 z --z 80 0 70 (/) 0 60 a. 50 40 30 20 10 0 0 WOLFCREEK

---

THIMBLE 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2 2.2 2.4 2.6 2.8 3 TIME OF TRAVEL (SECONDS)

Rev. 7 WOLFCREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.3-38 DESIGN TRIP CURVE * *

• r-0 I. 0 WOLF CREEK -----------------------------------
• -------.,!

r-0. 5 -a. -a. 0.0 100 50 () ROD POSIT I (PERCENT INSERTED) Rev. 0 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.3-39 TYPICAL NORMALIZED Roo VERSUS PERCENT INSERTION. A_L RODS OUT BUT ONE WOLF CREEK 5 0 tz -5 UJ u c.:: UJ 0... II'\ ........,. -IV i-w U) u.. -15 u.. 0 _J <( >< -20 <( -25 '" ""--CURV F !

• BURNUP CURVE (MWD/MTU)

I 2 1550 7700 PERIOD STABILITY INDEX (HOURS) (HOURS-I) 32.4 27.2 -0.041 -O.OIIt -30 0 72 TiME Rev. 0 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.3-40 n.-r.or'\11 TTfl..AI o:_:o t '--U i r 0 t. I V C. i\ \J U V I J.. 1*1 t. ' : 'l'-1: \ WITH A FOOT HEIGHT AND 121 AooEMBLIES , , I A" MA Co 10 "b""-8 :=T f--6 :z: <( a::: 0 <( ::::>> C> 2 N I-0 :z: <( a::: 0 <( -2 ::::>> C> w -4 u :z: w a::: w -6 LL.. LL.. 0 1--8 _. 1-f---10 :z: <( a::: 0 -12 <( ::::>> C> 16 0

• , * * . \ WOLF CREEK . \ * * *
• 4
• STABILITY INDEX= -0.076 (HR"-1) \._ !c;. ....... * *****&.-..*****.-*

.. .)#,. Nl}3 Nl}l 8 12 16 20 24 28 32 36 40 48 52 56 60 64 HOURS AFTER WITHDRAWAL OF RCC E-i i WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FlGURE 4.3-41 X-Y XENON TEST THE MO 0 ANT j i i V RSUS I It1 c.. -60.0 -50.0 e--4o.o 0 Q. ffi 0 -30.0 w 0 <.:> ffi -20.0 ...J 0 0 -10.0 0 0 10 0 6 CONTROL BANK CONPENSATION BORON COMPENSATION ---EXPERIMENT ---CALCULATION 20 30 40 50 60 PERCENT OF FULL POWER -2400 -2200 -2000 -1800 e 0 Q. -1600 :=-c..> w -----1400 tt 0 I -1200 ffi ...J 0.. -1000 g ...J -800 <C 0<: (!) w -600 1-z: -400 -200 0 Rev. 70 80 90 100 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.3-42 CALCULATED AND DOPPLER DtfECT AND COEFFICIENTS AT BDL. 2-LOOP PLANT, 121 ASSEMBLIES. 12 FooT CoRE () 'J WOLP CREEK 1400 1200 1000 E 800 0.. L a. r co u 600 I '+00 L I" AI f'\111 J.Tr-1"1. ld\Ll<ULHI I:U-200 0 0 2000 4000 6000 8000 10000 I 2000 14000 I 6000 18000 Rev. 0 BURN UP, M\'ID /MTU WOLF CREEK UPDATED SAFETY ANALYSIS REPORT i:"TI"'IIR[ 4 3-ll"Z IJ..UU

• FooT CoRE WOLF CREEK 1300 1200 1000 :::E 900 Q... Q... 800 700 600 500 lWO MEASURED 0 900 1600 2700 3600 4500 BURNUP, MWD/MTU I

6300 7200 8!00 Rev. 0 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT Li I ..LV Ul\ L I * "-" I I "-'"- n n (' _ J-! :-1nD Pi V!!* L_ !-VV' 'L.o HBL!ES, lC: FOOT NT. ......... r\l-11 vun.c. I J WOLF CREEK I IIUU 1000 I 900 800 /MlASURED 700 600 CL. l:l.. . .:fl 500 II"-" "tUU I CALCULATED__/ 300 200 100 0 0 I 000 2000 3000 4000 5000 6000 7000 8000 9000 i 0000 i i 000 I 2000 l 3000 I 4000 BURNUP, Mt'ID/MTU Rev. 0 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT ,, -::-'*' t-I 1__: IIU L L.J. ...__W....._ I ..!.VUl\L. !sV IV t,..-1 c t.. c* r-: (' 2_ =! n n D D! ,"; T C 7 ; i i..._ ,--;, 0 ....... ...., '!' ......; ._ v vi i b_ .-. ;*,; i '!' y; /1, 12 FooT CORE RCS BORON (PPM) 2,000 BOC: Xe=O, Peak Sm 1,600 1,200 I onn L uuu I r 400 I 0 0 4 8 12 16 20 CORE AVERAGE BURNUP (GWD/MTU) Rev. 11 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT TYPICAL BORON LETDOWN CURVE WOLF CREEK 4.4 THERMAL AND HYDRAULIC DESIGN 4.4.1 DESIGN BASES 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 or the emergency core cooling system (when applicable) assures that the following performances and safety criteria requirements are met:

a. Fuel damage (defined as penetration of the fission product barrier, i.e., the fuel rod clad) 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.

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

rods damaged (see above definition) 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. 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 There will be at least a 95% probability that departure for nucleate boiling (DNB) will not occur on 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 95% confidence level. Historically, this criterion has been conservatively met by adhering to the following thermal design basis: there must be at least a 95% probability that the minimum departure from nucleate boiling ratio (DNBR) of the limiting power rod during Condition I and II events is greater than or equal to the DNBR limit of the DNB correlation being used. The DNBR limit for the correlation is established based on the variance of the correlation such that there is a 95% probability with 95% confidence that DNB will not occur when the calculated DNBR is at the DNBR limit. 4.4-1 Rev. 6 WOLF CREEK Discussion The WCGS utilizes the EPRI VIPRE-01 Computer Code (Versatile Internals and Component Program for Reactors, EPRI) with the WRB-2 Critical Heat Flux Correlation for core thermal-hydraulic analysis. The WRB-2 correlation was developed to obtain a more accurate CHF predictor for mixing vane grid fuel assemblies of the same design as the 17X17 standard fuel mixing vane design (Reference 8). The calculated design limit DNBR for the VIPRE-01 code with the WRB-2 CHF correlation is below the 1.17 DNBR design limit for the Westinghouse THINC computer code with the WRB-2 correlation. Therefore, conservative use of a 1.17 DNBR design limit for the VIPRE-01 code with the WRB-2 correlation will be utilized for core DNBR analyses. DNBR margin is maintained for the fuel by ensuring the DNB safety analyses meet a Safety Analysis Limit DNBR of 1.76. The limiting Condition II transient from DNBR perspective, is analyzed using VIPRE-01 code and results in a minimum DNBR greater than 1.76. The Safety Analysis Limit DNBR is set greater than the design limit DNBR (see Section 4.4.2.12) to provide generic DNB margin. For analyses beyond the range of application of the WRB-2 correlation, the W-3

CHF Correlation is used. For the Rod Withdrawal from Subcritical accident

analysis, the design limit DNBR for the VIPRE-01 code with the W-3 correlation

is 1.30. For the Steam Line Break accident analysis, the design limit DNBR

for the VIPRE-01 code with the W-3 correlation is 1.45 (see section 4.4.2.2). 4.4.1.2 Fuel Temperature Design Basis Basis During modes of operation associated with Condition I and Condition II events, there is at least a 95-percent probability that the peak kW/ft fuel rods will

not exceed the U0 2 melting temperature at the 95-percent confidence level. The melting temperature of U0 2 is taken as 5,080°F (Ref. 1), unirradiated and decreasing 58°F 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 4,700°F has been selected as the overpower limit. This provides sufficient margin for uncertainties in the thermal evaluations, as described in Section

4.4.2.9.1. 4.4-2 Rev. 16 WOLF CREEK 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 events given in Chapter 15.0. 4.4.1.3 Core Flow Design Basis

Basis A minimum of 91.6 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 8.4 percent of this value is

allotted as bypass flow. This includes rod cluster control guide thimble

cooling flow, head cooling flow, baffle leakage, and leakage to the vessel

outlet nozzle. 4.4.1.4 Hydrodynamic Stability Design Basis 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, are sufficiently comprehensive so no additional

limits are required. Fuel rod diametrical 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

diametrical gap characteristics change with time (see Section 4.2.3.3), and the fuel rod integrity is evaluated on that basis. The effect of the moderator

flow velocity and distribution (see Section 4.4.2.2) and moderator void

distribution (see Section 4.4.2.4) are included in the core thermal evaluation

and thus affect the design bases. 4.4-3 Rev. 10 WOLF CREEK Meeting the fuel clad integrity criteria covers possible effects of clad temperature limitations. As noted in Section 4.2.3.3, 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 (see Section 15.6.5), control rod ejection accident, and locked rotor accident.4.

4.2 DESCRIPTION

4.4.2.1 Summary Comparison

The design of the WCGS unit described in this report has similar thermal-hydraulic parameters as the Comanche Peak Units 1 and 2 (Docket Nos. 50-445 and 50-446).Values of pertinent design and operating parameters are presented in Table 4.4-

1. The reactor is designed to meet the DNB design basis as no fuel centerline melting during normal operation, operational transients, and faults of moderate frequency.

Fuel densification has been considered in the DNB and fuel temperature evaluations, utilizing the methods and models described in detail in Reference

3.4.4.2.2 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 minimum DNBR in the limiting flow channel will be downstream of the peak heat flux location (hot spot) due to the increased downstream enthalpy rise. DNBRs are calculated by using the correlation and definitions described in Sections 4.4.2.2.1 and 4.4.2.2.2. The VIPRE-01 computer code (discussed in Section 4.4.4.5.1) 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 4.4.4.3.1 (nuclear hot

channel factors) and in Section 4.4.2.2.4 (engineering hot channel factors). 4.4-4 Rev. 11 WOLF CREEK The WRB-2 (Reference 8) correlation has been developed to predict the DNB performance of Westinghouse fuel designs which employ grids with mixing vanes

of the same design as the 17X17 standard fuel mixing vane design. This mixing

vane design is unique in that the mixing vane area to flow area ratio and the

azimuthal extension of the vanes around the rod circumference both differ from

the mixing vane designs of other Westinghouse grids. The WRB-2 correlation is also based entirely on rod bundle data. However, the data used to develop the correlation is essentially the 17X17 "R" type grid subset (both STD and OFA fuel) of the WRB-1 data base plus additional DNB test data obtained to quantify

the increase in DNB performance due to the addition of Intermediate Flow mixing

vane grids (IFM) for the Vantage 5 geometry. The applicable range of variables for the WRB-2 correlation is: Pressure  : 1440 < P < 2490 psia Local Mass Velocity  : 0.9 < G loc/10 6 < 3.7 lb/ft 2-hr Local Quality  : -0.1 < X loc < 0.3 Heated Length, Inlet to  : L h < 14 feet CHF Location Grid Spacing  : 10 < g sp < 26 inches Equivalent Hydraulic Diameter  : 0.37 < d e < 0.51 inches Equivalent Heated Hydraulic  : 0.46 < d h < 0.59 inches Diameter

Figure 4.4-22 shows measured critical heat flux plotted against predicted critical heat flux using the WRB-2 correlation. The 95/95 limit DNBR utilized

in thermal/hydraulic analyses has been conservatively set equal to the

Westinghouse THINC/WRB-2 code design limit DNBR of 1.17, appropriate for 17X17

standard fuel assemblies. For conditions outside the range of applicability of

the WRB-2, the W-3 correlation is used. For the W-3 correlation, the 95/95 limit DNBR is 1.30 at system pressures greater than or equal to 1000 psi. For low pressure application (500-1000

psi), the 95/95 limit DNBR is 1.45 (Reference 87). 4.4.2.2.1 Departure from Nucleate Boiling Technology Early experimental studies of DNB were conducted with fluid flowing inside single heated tubes or channels and with single annulus configurations with one

or both walls heated. The results of the experiments were analyzed, using many different physical models for describing the DNB phenomenon, but all resultant

correlations are highly empirical in nature. The evolution of these

correlations is described by Tong (Ref. 4 and 5), including the W-3 correlation

which is in wide use in the pressurized water reactor industry. 4.4-5 Rev. 10 WOLF CREEK As testing methods progressed to the use of rod bundles instead of single channels, it became apparent that the bundle average flow conditions could not

be used in DNB correlations. As discussed by Tong (Ref. 6) test results showed

that correlations based on average conditions were not accurate predictors of

DNB heat flux, and that a knowledge of the local subchannel conditions within

the bundle is necessary. In order to determine the local subchannel conditions, the VIPRE-01 Code (Ref.

9) was developed. VIPRE-01 has been developed for nuclear power utility thermal/hydraulic analysis applications. It is designed to help evaluate

nuclear reactor core safety limits, including minimum departure from nucleate

boiling ratio (MDNBR), fuel and clad temperature, and coolant state in normal operating steady state and transients and assumed accident conditions. VIPRE-01 was developed on the strengths of the COBRA code series and has gone through extensive benchmarking against COBRA in Reference 9. Calculations

covered a large range of data from comparisons of VIPRE-01 calculations to

simple heat-conduction problems having analytical solutions, to complex

environments involving flow blockage, two phase pressure drop, void fraction

measurements, fuel temperatures and heat transfer. The basic computational philosophy of VIPRE uses the subchannel analysis concept where a problem is divided into a number of quasi-one-dimensional

channels that communicate laterally by diversion crossflow and turbulent

mixing. Conservation equations of mass, axial and lateral momentum, and energy

are solved for the fluid enthalpy, axial flow rate, lateral flow per unit length, and momentum pressure drop. The flow field is assumed to be incompressible and homogeneous, although models are added to reflect subcooled

boiling and co-current liquid/vapor slip. NRC approval of the EPRI VIPRE-01 computer code is given in Reference 90. WCGS Thermal Hydraulic methodology utilizing the VIPRE-01 code is given in Reference 2.4.4.2.2.2 Definition of Departure from Nucleate Boiling Ratio The DNB heat flux ratio (DNBR) as applied to this design when all flow cell walls are heated is: DNBR = q" DNB,N q" loc [4.4-4] where: q " DNB,N = q" DNB,EU F [4.4-5] and q" DNB,EU is the uniform DNB heat flux as predicted by the applicable DNB correlation. 4.4-6 Rev. 10 WOLF CREEK F is the flux shape factor to account for nonuniform axial heat flux distributions (Ref. 85) with the "C" term modified as in Reference 5. q" is the actual local heat flux.

4.4.2.2.3 Mixing Technology The transverse momentum equation in VIPRE-01 includes terms describing the exchange of momentum between channels due to turbulent mixing. Turbulent

mixing is natural eddy diffusion between subchannels which is characterized by

eddy diffusivities. However, for numerical applications, such as VIPRE-01, the

turbulent mixing is represented by an equivalent lateral mass flow rate. This

equivalent lateral flow, defined as W' in VIPRE-01, defines the coolant

exchange rate between adjacent channels and thus, specifies the exchange of

mass, energy, and momentum between channels. Turbulent mixing in VIPRE-01 is accounted for with an empirical relation in which the user must specify a form for the turbulent mixing correlation, the

turbulent mixing coefficient, and the turbulent momentum factor (FTM). There are four correlations available in VIPRE-01 for defining the turbulent crossflow. A sensitivity study was performed in Reference 93, showing that VIPRE-01 in insensitive to which correlation is utilized. The correlation used in WCGS analyses is:

_ W' = *S*G [4.4-9] where: W' = the calculated turbulent crossflow, lbm*ft/sec = the turbulent mixing coefficient (TDC) S = the gap width, ft _ G = the average mass velocity in the channels connected by the gap under consideration lbm/sec The application of the TDC in the VIPRE-01 analysis for determining the overall mixing effect on heat exchange rate is presented in Reference 9andReference

93.As a part of an ongoing research and development program, Westinghouse has sponsored and directed mixing tests at Columbia University (Ref. 12). 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 pressurized water reactor core under the following single

and two phase (subcooled boiling) flow conditions: 4.4-7 Rev. 11 WOLF CREEK Pressure 1,500 to 2,400 psia Inlet temperature 332°to 642°F Mass velocity 1.0 to 3.5 x 10 6 lb /hr-ft 2 Reynolds number 1.34 to 7.45 x 10 5 m 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 in Figure 4.4-4 where the TDC is plotted versus the Reynolds number. TDC 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 investigated by Cadek (Ref. 12), Rowe and Angle (Ref. 13 and 14), and Gonzalez-Santalo and Griffith (Ref. 15). 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 pressurized

water reactor 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 (TDC) 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 thermal-hydraulic analyses. 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 (Ref. 16). 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, additional margin is available for this design, as the value of TDC increases as grid

spacing decreases (Ref. 12). Use of the 0.038 TDC for V5H and RFA fuel with IFM grids is utilized for Wolf Creek. Calculation of the generic DNBR margins for Wolf Creek was performed utilizing this 0.038 TDC (References 11, 92). The turbulent momentum factor (FTM) in VIPRE-01 controls the efficiency of the momentum mixing due to turbulent crossflow between subchannels. An FTM of 0.0

indicates that crossflow mixes enthalpy only, while an FTM of 1.0 indicates

that crossflow mixes momentum equally with enthalpy. Sensitivity studies performed during the VIPRE - 01 qualification effort have shown that VIPRE-01 is relatively insensitive to FTM (Reference 93). However, Reference 9 recommends an FTM of 0.8. 4.4-8 Rev. 16 WOLF CREEK 4.4.2.2.4 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 quantities. 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 is composed of a nuclear hot channel factor (see Section 4.4.4.3) describing the fission 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 rod diameter pitch and

bowing; inlet flow distribution; flow redistribution; and flow mixing. Heat Flux Engineering Hot Channel Factor, F E Q The heat flux engineering hot channel factor is used to evaluate the maximum heat flux. This subfactor is determined by statistically combining the tolerances for the fuel pellet diameter, density, enrichment, eccentricity, and

the fuel rod diameter, and has a value of 1.033. Measured manufacturing data

on recent Westinghouse 17 x 17 fuel were used to verify that this value was not exceeded for 95 percent of the limiting fuel rods at a 95-percent confidence level. Thus, it is expected that a statistical sampling of the fuel assemblies

of this plant will yield a value no larger than 1.033. Enthalpy Rise Engineering Hot Channel Factor, F EH The effect of variations in flow conditions and fabrication tolerances on the hot channel enthalpy rise is directly considered in the VIPRE core thermal subchannel analysis (see Section 4.4.4.5.1) under any reactor operating

condition. The items included in the consideration of the enthalpy rise engineering hot channel factor are discussed below:

a. Pellet diameter, density, and enrichment and fuel rod diameter, pitch, and bowing Design values employed in the

VIPRE analysis related to the above fabrication variations are

based on applicable limiting tolerances so that these design

values are met for 95 percent of the limiting channels at a

95-percent confidence level. Measured manufacturing data on

Westinghouse 17 x 17 fuel show that the tolerances used

in this evaluation are conservative. In addition, each

fuel assembly is inspected to assure that the channel

spacing design criteria are met. The effect of

variations in pellet diameter, enrichment, and density is

considered statistically in establishing the design limit DNBRs (see Subsection 4.4.2.12 for the Revised Thermal Design Procedure (Reference 91) employed in this application.) 4.4-9 Rev. 10 WOLF CREEK

b. Inlet flow maldistribution The consideration of inlet flow maldistribution in core thermal performances is discussed in Section 4.4.4.2.2.

A design basis of 5-percent reduction in coolant flow to the hot assembly is used in the VIPRE analysis.

c. 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 VIPRE analysis for every operating condition which is evaluated.

d. Flow mixing

The subchannel mixing model incorporated in the VIPRE Code and used in reactor design is based on experimental

data (Ref. 11 and 17) discussed in Sections 4.4.2.2.3 and

4.4.4.5.1. 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. 4 4.2.2.5 Effects of Rod Bow on DNBR The phenomenon of fuel rod bowing, as described in Reference 83, 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 DNBR and/or margin obtained from measured plant operating parameters (such as F H N or core 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 WCGS safety analysis, sufficient margin was maintained between the VIPRE-01/WRB-2 design limit DNBR (1.23) and the safety analysis limit DNBR (1.76) to completely offset any DNBR penalties associated with rod bow (a maximum of less than 1.5% for a burnup of 24,000 MWD/MTD, identified in References 83, 84 and 85. 4.4-10 Rev. 16 WOLF CREEK The maximum rod bow penalties accounted for in the design safety analysis are based on an assembly average burnup of 24,000 MWD/MTU (Reference 85). At burnups greater than 24,000 MWD/MTU, credit is taken for the effect F H Nburndown, due to the decrease in fissionable isotopes and the buildup of fission product inventory, and no additional rod bow penalty is required. 4.4.2.3 Linear Heat Generation Rate

The core average and maximum linear heat generation rates are given in Table 4.4-1. The method of determining the maximum linear heat generation rate is

given in Section 4.3.2.2. 4.4.2.4 Two Phase Flow Correlations and Void Correlations Two phase flow is less well understood and considerably more complex than single phase flow and consequently requires more constitutive relations.

VIPRE-01 uses the homogeneous model for two phase flow. It considers the two

phase flow to be a single fluid with the properties (density, viscosity, etc.) of the mixture. This is a fairly reasonable approximation of the flow field at high pressures and high mass velocities, but is less satisfactory at lower

pressures and low mass velocities. The homogeneous model can be modified by

including a two phase flow multiplier in the calculation of the friction

pressure losses. The mixture density for momentum can be adjusted by using the void fractions determined with void fraction/quality relations and subcooled

void correlations. These correlations take into account the effects of

nonhomogeneities in the two phase flow field. The subcooled void correlation in VIPRE-01 is used to model the transition from single-phase to boiling flow for heat transfer from a hot wall. Specifically, the subcooled correlations in VIPRE-01 are used to model boiling which occurs in the proximity of the fuel rod while the bulk flow remains in a subcooled condition. Thus, the subcooled correlation is used to determine a flowing

quality for the coolant even though the bulk fluid temperature remains below

saturation. The flowing quality, supplied by the subcooled void correlation, is then used in a bulk void correlation to calculate the subcooled void. The two phase flow correlations available in VIPRE-01 fall into three categories. They are the two phase friction multipliers, a subcooled void correlation, and a bulk void correlation. An option to supply a hot wall

correction term to the friction factor is also available. 4.4-11 Rev. 10 WOLF CREEK The subcooled void correlation selected for use was the EPRI correlation, which was developed from rod bundle data. The other subcooled void correlation

available in VIPRE-01 is the LEVY correlation. LEVY was developed from data using non-cylindrical geometries. Both correlations employ a two-step method for the determination of the flowing quality. First, the subcooled

correlations determine the point of bubble departure from the heated surface. Secondly, the correlations establish a relationship between the actual local vapor fraction and the corresponding thermal equilibrium value. The EPRI

correlation was determined to be more appropriate for use in the analysis of reactor cores and CHF test sections (Reference 9). The bulk void correlation used was also the EPRI correlation, which should be used when the EPRI subccooled void correlation is selected (Reference 9). This correlation is actually the Zuber-Findlay correlation with coefficients developed for the EPRI model. The two phase friction multiplier used in the thermal-hydraulic analyses is the EPRI correlation, which was developed principally from rod bundle two phase

pressure drop experimental data (Reference 9). This two phase friction flow

correlation demonstrated superior performance over other correlations available

in VIPRE-01 during the VIPRE-01 benchmarking effort. Like the subcooled void

and bulk void correlations, the two phase friction multiplier correlations are

empirical expressions dependent upon the range of conditions and physical

geometries used in the correlation derivation. As such, the selection of

appropriate subcooled void, bulk void, and two phase friction multiplier correlations can be based on those correlations which have consistent assumptions and complimentary bases. 4.4.2.5 Core Coolant Flow Distribution Assembly average coolant mass velocity and enthalpy at various radial and axial core locations for first core near the beginning of core life power distribution are given in Figures 4.4-5 through 4.4-7. Typical coolant

enthalpy rise and flow distributions for the 4-foot elevation (1/3 of core

height) are shown in Figure 4.4-5, for the 8-foot elevation (2/3 of core

height) in Figure 4.4-6, and at the core exit in Figure 4.4-7. The THINC Code analysis for this case utilized a uniform core inlet enthalpy and inlet flow

distribution. No orificing is employed in the reactor design. 4.4-12 Rev. 11 WOLF CREEK 4.4.2.6 Core Pressure Drops and Hydraulic Loads 4.4.2.6.1 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 4.4.2.7. The core pressure drop includes the fuel assembly (including the effect of inserted core components, such as rod cluster controls), lower core plate, and upper core plate pressure drops. The full power operation pressure drop values shown in Table 4.4-1 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 for actual plant operating conditions, as described in

Section 5.1.4. This section 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

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 4.4.2.9.2. 4.4.2.6.2 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

possible overdeflection 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 loss-of-coolant

accident. These conditions are presented in Section 15.6.5. 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 based on the cold mechanical design flow, but are adjusted to account for the coolant density difference. Conservative core hydraulic loads for a pump overspeed transient, which could possibly create flow rates 20 percent greater than the mechanical design flow, are evaluated to be approximately twice the fuel assembly weight. Core hydraulic loads were measured during the prototype assembly tests described in Section 1.5. Reference 19 contains a detailed discussion of the

results. 4.4-13 Rev. 13 WOLF CREEK 4.4.2.7 Correlation and Physical Data 4.4.2.7.1 Surface Heat Transfer Coefficients The VIPRE-01 code contains a set of heat transfer correlations for each of the four regions of the boiling curve. The user can supply a separate heat transfer correlation for use in the single phase forced convection region, the subcooled and saturated nucleate boiling region, the transition boiling region, and the film boiling region. Each correlation is discussed in detail in Reference 9. In the single phase forced convection region, forced convection heat transfer coefficients are obtained from the familiar Dittus-Boelter correlation (Ref. 20), with the properties evaluated at bulk fluid conditions: hD e K = 0.023(D e G m)0.8(C p m K)0.4 [4.4-10] where h = heat transfer coefficient, (Btu/hr-ft 2-F) D e = equivalent diameter, (ft) K = thermal conductivity, (Btu/hr-ft-F) G = mass velocity, (lb m/hr-ft 2) = dynamic viscosity, (lb m/ft-hr) C p = heat capacity, (Btu/lb m-F)This correlation has been shown to be conservative (Ref. 21) for rod bundle

geometries with pitch to diameter ratios in the range used by pressurized water reactors.The onset of nucleate boiling occurs when the clad wall temperature reaches the amount of superheat predicted by Thom's (Ref. 22) correlation. After this

occurrence, the outer clad wall temperature is determined by: T sat = [0.072 exp (-P/1260)] (q") 0.5 [4.4-11] where:T sat = wall superheat, T w - T sat (F) q" = wall heat flux, (Btu/hr-ft

2) P = pressure, (psia)

T w = outer clad wall temperature, (F) T sat = saturation temperature of coolant at P, (F) 4.4-14 Rev. 7 WOLF CREEK VIPRE benchmarking studies indicate that the Thom correlation provided the best agreement with experimental data. The WRB-2 CHF correlation was used to define the peak of the boiling curve. The VIPRE-01 code has not been approved for licensing analyses for conditions in which the heat transfer mode is beyond the point of CHF on the boiling curve (Reference 90). The Condie-Bengston correlation was used to define the transition boiling region and the Groenveld 5.7 correlation used in the film

boiling region. These correlations are input only to complete the heat transfer correlation set. 4.4.2.7.2 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 Table 4.4-3). Two-phase flow considerations in the core thermal subchannel analyses are considered, and the models are discussed in Section 4.4.4.2.3. Core and vessel

pressure losses are calculated by equations of the form: DP L =K+F L D e v 2 2g c (144) [4.4-12] where:P L = unrecoverable pressure drop, (lb f/in 2) = fluid density, (lb m /ft 3) L = length, (ft) D e = equivalent diameter, (ft) V = fluid velocity, (ft/sec) g c = 32.174, (lb m-ft/lb f-sec 2) 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. 4.4-15 Rev.7 WOLF CREEK 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 (Ref. 23 and 24) and form loss relationships (Ref. 25). Moody (Ref. 26) curves were used to obtain the single phase friction factors. Tests of the primary coolant loop flow rates were made (see Section 4.4.5.1) 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, were conservative. 4.4.2.8 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 DNB design basis is met. Figure 15.0-1 shows the DNBR limit lines and the resulting Overtemperature T trip lines (which become part of the Technical Specifications), plotted as T, versus T avg 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 (see Section 15.4.2) specific protection functions

are provided as described in Section 7.2, and the use of these protection

functions are described in Chapter 15.0. 4.4.2.9 Uncertainties in Estimates 4.4.2.9.1 Uncertainties in Fuel and Clad Temperatures

As discussed in Section 4.4.2.11, 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 inpile measurements, (Ref. 30 through 36), by out-of-pile measurements of the fuel and clad properties (Ref. 37 through 48), 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 also included in the calculation of the total

uncertainty. 4.4-16 Rev.7 WOLF CREEK 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 4.3.2.2.1. 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 coolant, the film temperature drop does

not appreciably contribute to the uncertainty. 4.4.2.9.2 Uncertainties in Pressure Drops Core and vessel pressure drops based on the best estimate flow, as described 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 4.4.5.1, tests on the primary system prior to initial criticality were made to verify that a conservative primary system coolant flow

rate has been used in the design and analyses of the plant. 4.4.2.9.3 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 4.4.4.2.2. 4.4.2.9.4 Uncertainty in DNB Correlation

The uncertainty in the DNB correlation (see Section 4.4.2.2) 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 4.4.2.2.2. 4.4.2.9.5 Uncertainties in DNBR Calculations

The uncertainties in the DNBRs calculated by VIPRE-01 analysis (see Section 4.4.4.5.1) 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 allowances. In 4.4-17 Rev.7 WOLF CREEK addition, conservative values for the engineering hot channel factors are used as discussed in Section 4.4.2.2.4. The results of a sensitivity study (Ref.

18, 93) with VIPRE-01 show that the minimum DNBR in the hot channel is relatively insensitive to variations in the core-wide radial power distribution (for the same value of F H N).The ability of the VIPRE-01 Code to accurately predict flow and enthalpy distributions in rod bundles is discussed in Section 4.4.4.5.1. Studies have been performed (Ref. 18, 93) to determine the sensitivity of the minimum DNBR in the hot channel to the void fraction correlation (also see Section 4.4.2.4); 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. 4.4.2.9.6 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. In addition, another 8.4

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 4.4.4.2.1. 4.4.2.9.7 Uncertainties in Hydraulic Loads

As discussed in Section 4.4.2.6.2, hydraulic loads on the fuel assembly are evaluated for a pump overspeed transient which creates flow rates 20 percent

greater than the mechanical design flow. As stated in Section 5.1, the

mechanical design flow is greater than the best estimate or most likely flow

rate value for the actual plant operating condition. 4.4.2.9.8 Uncertainty in Mixing Coefficient

The value of the mixing coefficient, TDC, used in VIPRE analyses for this application is 0.038, approved for grid spacing < 22 in.The results of the mixing tests done on 17 x 17 geometry, as discussed in Section 4.4.2.2.3, had a mean value of TDC of 0.059 and standard deviation of 0.007. Calculation of generic DNBR margin was done utilizing a 0.038 TDC. 4.4-18 Rev. 10 WOLF CREEK 4.4.2.10 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 could cause changes in hot channel factors. However, these events are analyzed separately in Chapter 15.0. This discussion will be confined to flux tilts caused by x-y xenon transients, inlet temperature mismatches, enrichment variations within tolerances, and so forth. The design value of the enthalpy rise hot channel factor F H N, which includes an 8-percent uncertainty (as discussed in Section 4.3.2.2.7), is sufficiently conservative such that flux tilts up to and including the alarm point (see the Technical Specifications) will not result in values of F H N greater than that assumed in the limiting analysis. The design value of F Q does not include a specific allowance for quadrant flux tilts. When the indicated quadrant power tilt ratio exceeds 1.02, corrective action must be taken. The procedure to be followed is explained in detail in the Technical Specifications. The quadrant power tilt ratio limit assures that the radial power distribution satisfies the design values used in the power capability analysis. 4.4.2.11 Fuel and Cladding Temperatures 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 4.2.3.1. The thermal-hydraulic design assures that the maximum fuel temperature is below the melting point of UO 2 (see Section 4.4.1.2). To preclude center melting and as a basis for overpower protection system setpoints, a calculated centerline

fuel temperature of 4,700°F has been selected as the overpower limit. This provides sufficient margin for uncertainties in the thermal evaluations as described in Section 4.4.2.9.1. The temperature distribution within the fuel

pellet is primarily a function of the local power density and the UO 2 thermal conductivity. However, the computation of radial fuel rod 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 semiempirical thermal model (see Section 4.2.3.3) with the model modifications

for time dependent fuel densification given in Reference 3. This thermal model

enables the determination of these factors and their net effects on temperature

profiles. The temperature predictions have been compared to inpile fuel

temperature measurements (Ref. 30 through 36) and melt radius data (Ref. 50 and

51) with good results. 4.4-19 Rev. 11 WOLF CREEK As described in Reference 3, fuel rod thermal evaluations (fuel centerline, average and surface temperatures) are determined throughout the fuel rod

lifetime with consideration of time dependent densification. To determine the

maximum fuel temperatures, various burnup rods, including the highest burnup rod, are analyzed over the rod linear power range of interest.

The principal factors which are employed in the determination of the fuel

temperature are discussed below. 4.4.2.11.1 UO 2 Thermal Conductivity The thermal conductivity of uranium dioxide was evaluated from data reported by Howard, et al. (Ref. 37); Lucks et al. (Ref. 38); Daniel, et al. (Ref. 39);

Feith (Ref. 40); Vogt, et al. (Ref. 41); Nishijima, et al. (Ref. 42); Wheeler, et al. (Ref. 43); Godfrey, et al. (Ref. 44); Stora, et al. (Ref.45); Bush (Ref.

46); Asamoto, et al. (Ref. 47); Kruger (Ref. 48); and Gyllander (Ref. 52).

At 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 0 2800°C Kdt is 93 watts/cm. This conclusion is based on the integral values reported by Gyllander (Ref. 52), Lyons, et al. (Ref. 53), Coplin, et al. (Ref. 54), Duncan (Ref. 50), Bain (Ref. 55), and Stora (Ref.

56). The design curve for the thermal conductivity is shown in Figure 4.4-9. The

section of the curve at temperatures between 0° and 1,300°C is in excellent

agreement with the recommendation of the IAEA panel (Ref. 57). The section of

the curve above 1,300°C is derived for an integral value of 93 watts/cm (Ref.

50, 52 and 56).

Thermal conductivity of UO2 at 95-percent theoretical density can be

represented best by the following equation:

K = 1 11.8 + 0.0238T

 + 8.775 x 10

-13 T 3 [4.4-13] where: K = watts/cm-°C

T = °C

4.4.2.11.2 Radial Power Distribution in UO2 Fuel Rods

An accurate description of the fuel rod 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. Radial power distributions in

UO 2 fuel rods are determined with the neutron transport theory code, LASER. The LASER Code has been

4.4-20 Rev. 23 WOLF CREEK validated by comparing the code predictions on radial burnup and isotopic distributions with measured radial microdrill data (Ref. 58 and 59). 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 c , relative to the pellet surface temperature, T s , through the expression: T S T C K(T) dT = g'f 4p [4.4-14] where: K(T) = the thermal conductivity for UO 2 with a uniform density distribution q' = the linear power generation rate 4.4.2.11.3 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 the UO 2 thermal conductivity model, the calculated fuel centerline temperatures predict the inpile temperature

measurements. A more detailed discussion of the gap conductance model is presented in Reference 88. 4.4.2.11.4 Surface Heat Transfer Coefficients

The fuel rod surface heat transfer coefficients during subcooled forced convection and nucleate boiling are presented in Section 4.4.2.7.1. 4.4.2.11.5 Fuel Clad Temperatures The outer surface of the fuel rod at the hot spot operates at a temperature of approximately 660 F for steady state operation at rated power throughout core

life due to the presence 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 clad temperature. 4.4.2.11.6 Treatment of Peaking Factors The total heat flux hot channel factor, F Q , is defined as the ratio of the maximum to core average heat flux. The design value of F Q as presented in Table 4.3-2 and discussed in Section 4.3.2.2.6, is 2.50 for normal operation. This results in a peak linear power at full power conditions of 14.48 for 3565 MWt operation. 4.4-21 Rev.9 WOLF CREEK As described in Section 4.3.2.2.6, the peak linear power resulting from overpower transients/operator errors (assuming a maximum overpower of 118

percent) is limited such that the centerline fuel melt kW/ft limit is never exceeded. The centerline temperature kW/ft must be below the UO 2 melt temperature over the lifetime of the rod, including allowances for uncertainties. The fuel temperature design basis is discussed in Section 4.4.1.2 and results in a maximum allowable calculated centerline temperature of 4,700°F. The centerline temperature at the peak linear power resulting from

overpower transients/operator errors (assuming a maximum overpower of 118 percent) is below that required to produce melting. 4.4.2.12 Revised Thermal Design Procedure (RTDP) WCGS utilizes the Revised Thermal Design Procedure (RTDP), Reference 91. to determine a design limit DNBR value used as a basis in thermal-hydraulic

analyses. 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). 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. The RTDP design limit DNBR value for the WCGS is 1.23. The design limit DNBR is used as a basis for the technical specifications and for consideration in evaluations completed in accordance with 10 CFR 50.59. To maintain DNBR margin to offset DNBR penalties such as those due to rod bow, the safety analyses are performed to DNBR limits higher than the design limit

DNBR value. The difference between the design limit DNBR and the safety

analysis limit DNBR results in available DNBR margin. The net DNBR margin, after consideration of all applicable penalties, is available for operating and design flexibility. The safety analysis limit DNBR is 1.76. The Standard Thermal Design Procedure (STDP is used for those analyses where RTDP is not applicable. In the STDP method, 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 analysis input

values to give the lowest minimum DNBR. The DNBR limit for STDP is the

appropriate DNB correlation limit after consideration of applicable penalties is made. 4.4-22 Rev. 16 WOLF CREEK 4.

4.3 DESCRIPTION

OF THE THERMAL AND HYDRAULIC DESIGN OF THE REACTOR COOLANT SYSTEM 4.4.3.1 Plant Configuration Data Plant configuration data for the thermal hydraulic and fluid systems external to the core are provided as appropriate in Chapters 5.0, 6.0, and 9.0. Implementation of the emergency core cooling system (ECCS) is discussed in Chapter 15.0. Some specific areas of interest are the following:

a. Total coolant flow rates for the reactor coolant system (RCS) and each loop are provided in Table 5.1-1. Flow rates employed in the evaluation of the core are presented throughout Section 4.4.

b. 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.

c. The flow path length through each volume may be calculated from physical data provided in the above

referenced tables.

d. The height of fluid in each component of the RCS may be determined from the physical data presented in Section 5.4. The components of the RCS are water filled during

power operation with the pressurizer being approximately

60 percent water filled.

e. Components of the ECCS are to be located so as to meet the criteria for net positive suction head described in

Section 6.3.

f. 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 Chapter 15.0.

g. The parameters for components of the RCS are presented in Section 5.4.

h. The steady state pressure drops and temperature distributions through the RCS are presented in Table 5.1-

1.

4.4.3.2 Operating Restrictions on Pumps The minimum net positive suction head and minimum seal injection flow rate must be established before operating the reactor coolant pumps. With the minimum 6-gpm labyrinth seal injection flow rate established before each pump operation, the operator will have to verify that the system pressure satisfies net positive suction head requirements. 4.4-23 Rev.10 WOLF CREEK 4.4.3.3 Power-Flow Operating Map (BWR) Not applicable to WCGS. 4.4.3.4 Temperature-Power Operating Map The relationship between RCS temperature and power is shown in Figure 4.4-10.

The effects of reduced core flow due to inoperative pumps are discussed in Sections 5.4.1, 15.2.5, and 15.3.4. Natural circulation capability of the

system is shown in Table 15.2-2. 4.4.3.5 Load Following Characteristics Load follow using control rod motion and dilution or boration by the boron system is discussed in Section 4.3.2.4.16. The RCS is designed on the basis of steady state operation at full power heat load. The reactor coolant pumps utilize constant speed drives as described in Section 5.4, and the reactor power is controlled to maintain average coolant temperature at a value which is a linear function of load, as described in Section 7.7. 4.4.3.6 Thermal and Hydraulic Characteristics Summary Table The thermal and hydraulic characteristics are given in Tables 4.3-1, 4.4-1, and

4.4-1.4.4.4 EVALUATION 4.4.4.1 Critical Heat Flux The critical heat flux correlation utilized in the core thermal analysis is explained in detail in Section 4.4.2. 4.4.4.2 Core Hydraulics 4.4.4.2.1 Flow Paths Considered in Core Pressure Drop and Thermal Design The following flow paths for core bypass flow are considered:

a. Flow through the spray nozzles into the upper head for head cooling purposes

b. Flow entering into the rod cluster control guide thimbles to cool the control rods

c. Leakage flow from the vessel inlet nozzle directly to the vessel outlet nozzle through the gap between the vessel

and the barrel 4.4-24 Rev. 11 WOLF CREEK

d. Flow introduced between the baffle and the barrel for the purpose of cooling these components

e. Flow in the gaps between the fuel assemblies on the core periphery and the adjacent baffle wall The above contributions are evaluated to confirm that the design value of the core bypass flow is met. The design value of core bypass flow for the standard plant is equal to 8.4 percent of the total vessel flow.

Of the total allowance, 3.01 percent is associated with the internals (items a, c, d, and e above), 3.6 percent for the core, and a 1.79 percent flow measurement uncertainty. Calculations have been performed using drawing tolerances on a worst-case basis and accounting for uncertainties in pressure losses. Based on these calculations, the core bypass flow is <8.4 percent. Flow model test results for the flow path through the reactor are discussed in Section 4.4.2.7.2. 4.4.4.2.2 Inlet Flow Distributions Data from several 1/7 scale hydraulic reactor model tests (Ref. 23, 24, and 62) have been utilized in arriving at the core inlet flow maldistribution criteria

to be used in the VIPRE analyses (see Section 4.4.4.5.1). 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 (Ref. 63). The same design basis of 5-percent reduction to the hot assembly inlet is used in VIPRE analyses. The experimental error estimated in the inlet velocity distribution has been considered as outlined in Reference 18 where the sensitivity of changes in inlet velocity distributions to hot channel thermal performance is shown to be small. Studies (Ref. 18) 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

assembly for a loop out of service based on the experimental data in References

23 and 24. The effect of the total flow rate on the inlet velocity distribution was studied in the experiments of Reference 23. As was expected, on the basis of

the theoretical analysis, no significant variation could be found in inlet

velocity distribution with reduced flow rate. 4.4.4.2.3 Empirical Friction Factor Correlations

Two empirical friction factor correlations are used in the VIPRE-01 Code (described in Section 4.4.4.5.1). The friction factor in the axial direction, parallel to the fuel rod axis, is evaluated using the Darcy formulation of the friction pressure drop in one-

dimensional flow (Ref. 9). The frictional pressure loss in the axial direction

is given by: dP fG 2 v' __ = _____ [4.4-15] dX 2D h g c 4.4-25 Rev.10 WOLF CREEK where G = mass velocity (lbm/sec ft

2) Dh = hydraulic diameter based on wetted perimeter (ft) f = friction factor

g c = force-to-mass conversion constant v' = specific volume for momentum (ft 3/lbm)The pressure loss in lateral flow, for either gaps or leakage paths, is treated as a form drag loss rather than a wall friction loss. This permits the formulation of the pressure loss in terms of the known geometric quantities: lwl w v' P =K G____________ [4.4-16] 2S 2 g c where v' = specific volume for momentum (ft 3/lbm) w = crossflow through the gap (lbm/sec ft) K G = form loss coefficient g c = force-to-mass conversion constant S = gap width (ft.) 4.4.4.3 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, characterized by F H N (defined in Section 4.3.2.2.1) as well as axial heat flux profiles are discussed in the following two sections. 4.4.4.3.1 Nuclear Enthalpy Rise Hot Channel Factor, F H NGiven the local linear power density q'(kW/ft) at a point x, y, z in a core with N fuel rods and height H, F H N = hot rod power average rod power

= Max0 H q'(x o ,y o ,z o)dz 1 Nall rods0 H q'(x,y,z)dz         [4.4-18]

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 F H N recreates the axial heat 4.4-26 Rev. 10 WOLF CREEK flux along the limiting rod. The surrounding rods are assumed to have the same axial profile with rod average powers which are typical distributions found in

hot assemblies. In this manner, worst-case axial profiles can be combined with worst-case radial distributions for reference DNB calculations. It should be noted again that F H N is an integral and is used as such in 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 VIPRE analysis to

radial power shapes is discussed in Reference 93. For operation at a fraction P of full power, the design F H N used is given by: F H N = 1.65 [1 + 0.3 (1 - P)] [4.4-19] The permitted relaxation of F H N with power level is included in the DNB protection setpoints and allows radial power shape changes with rod insertion to the insertion limits (Ref. 66), thus allowing greater flexibility in the nuclear design. 4.4.4.3.2 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 or power change or due to a spatial xenon transients which may occur in the axial direction. Consequently, it is necessary to measure the

axial power imbalance by means of the excore 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 T channels on excessive axial power imbalance; that is, when a 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 is a chopped cosine shape with a peak to average value of 1.55. To determine the penalty to be taken in protection setpoints for extreme values of flux difference, this reference shape is supplemented by other axial shapes

skewed to the bottom and top of the core. The course of those accidents in

which DNB is a concern is analyzed in Chapter 15.0, assuming that the

protection setpoints have been set on the basis of these shapes. In many

cases, the axial power distribution in the hot channel changes throughout the

course of the accident due to rod motion, coolant temperature, and power level

changes.The initial conditions for the accidents for which DNB protection is required are assumed to be those permissible within the relaxed axial offset control strategy for the load maneuvers described in Reference 67. In the case of the loss-of-flow accident, the hot channel heat 4.4-27 Rev. 11 WOLF CREEK flux profile is very similar to the power density profile in normal operation preceding the accident. It is, therefore, possible to illustrate the

calculated minimum DNBR for conditions representative of the loss-of-flow accident as a function of the flux difference initially in the core. A plot of this type is provided in Figure 4.4-11 for first core initial conditions. As

noted on this figure, all power shapes were evaluated with a full power radial peaking factor (F H N) of 1.55. The radial contribution to the hot rod power shape is conservative both for the initial condition and for the condition at the time of minimum DNBR during the loss of flow transient. Also shown is the minimum DNBR calculated for the reference power shape at the same conditions. 4.4.4.4 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. 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 Chapter 7.0 are designed to meet these bases. The

response of the core to Condition II transients is given in Chapter 15.0. 4.4.4.5 Analytical Techniques 4.4.4.5.1 Subchannel Analysis Method

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

are not exceeded, using the most conservative power distribution. The thermal

design takes into account local variations in dimensions, power generation, flow redistribution, and mixing. The following sections describe the use of the VIPRE-01 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. The VIPRE-01 computer program uses the subchannel analysis concept where the reactor core or fuel bundle is divided into a number of quasi-one-dimensional

channels that communicate laterally by diversion crossflow and turbulent

mixing.The VIPRE-01 subchannel modeling allows a region of fluid flow to be described by a number of channels of various sizes and shapes. Channel size and shape

may be small and regular or relatively large and irregular simply by inputting

flow area and wetted perimeters. Hence, 4.4-28 Rev. 11 WOLF CREEK the hottest location in the core can be modeled in detail (such as a subchannel in a bundle array) and cooler locations in the core, which may include several

bundles, can be lumped together into a single channel. In any analysis where channel sizes differ, it is desirable to input model dependent cross flow variables consisting of the gaps between the channels and the distance between

connecting channel centroids. For each axial segment, conservation equations of mass, axial and lateral momentum, and energy are solved for the fluid enthalpy, axial flow rate, lateral flow per unit length, and momentum pressure drop. The flow field is

assumed to be incompressible and homogeneous, although models are added to

reflect subcooled boiling and co-current liquid/vapor slip. Fluid properties

are functions of the local enthalpy and a uniform but time-varying system

pressure with an option to add the effects of local pressure. The governing equations for the conservation of mass, energy, and momentum, and the two phase flow models of friction factors, two phase friction multipliers, subcooled and bulk quality/void models have been reviewed by comparison with

similar equations in the other widely used thermal hydraulic codes such as

COBRA-IIIC and COBRA-IV (Reference 9). In addition, VIPRE-01 contains a

recirculation solution option and a fuel rod heat conduction model with a

dynamic gap model not found in these codes. Estimates of uncertainties are discussed in Section 4.4.2.9.

Experimental Verification VIPRE-01 has been compared with a wide range of flow and heat transfer data to verify the accuracy and versatility of the code, and to demonstrate capabilities and limitations. Sensitivity studies on models, solution parameters and various input variables have been performed to provide users

with guidance in applying the code to practical problems (Reference 9). Comparisons were performed with VIPRE-01 for five main categories of data. These were single and two-phase flow field data, void/quality relation data, rod temperature measurements, heat transfer tests and experimental CHF data. Models of FSAR Chapter 15 transients were compared for different operating

plants, and in general the VIPRE-01 results were quite consistent with the

FSAR results. The flow distribution and two-phase pressure drop data comparisons (described in Reference 9, Volume 4) show that the VIPRE code does an excellent job of

predicting single phase flow distributions, even with severe blockages. The

code also does a good job of predicting pressure drop, in both single and two-

phase flows. The VIPRE-01 code is not capable of correctly predicting the flow distribution on a subchannel basis in two phase flow, however. This is 4.4-29 Rev. 10 WOLF CREEK a generic limitation of the homogeneous model. The problem is further illustrated in the code inability to predict the void distribution in

subchannels with two phase flow. It does, however, do a good job in predicting

the axial distribution of void and the overall two-phase pressure drop. It is

recommended that the most appropriate way to model two-phase flow is on a one-

dimensional, bundle average basis, rather than subchannel analysis. Evaluation of the fuel and gap conductance models in VIPRE against in-pile fuel centerline temperature data shows that the code predicts reasonable results for a large range of operating powers. There are however, significant

simplifications in the model. It is recommended that a user use a fuel

performance code for detailed fuel performance results, and use the VIPRE results only as a general guideline for expected behavior. The comparisons with boiling heat transfer data show that in nucleate boiling the Thom-plus-Dittus-Boelter heat transfer correlation option does the best

overall, of the available options in VIPRE. The range of data is limited, however, and it is recommended that the user evaluate the results carefully in

light of the intended application. The results presented for post-CHF heat

transfer illustrate the limitations of the homogeneous model in VIPRE in

calculating film boiling. Use of the VIPRE code for post-CHF analysis is not

recommended. Transient Analysis The VIPRE-01 thermal-hydraulic computer code has a transient evaluation capability. Operating transient data can be input for system pressure, core inlet temperature or enthalpy, core power level, and core inlet mass flux. The VIPRE-01 Code also has the capability for evaluating fuel rod thermal response. This is treated by the methods described in Section 15.0.11. 4.4.4.6 Hydrodynamic and Flow Power Coupled Instability Boiling flows may be susceptible to thermohydrodynamic instabilities (Ref. 72). These instabilities are undesirable in reactors, since they may cause a change in thermohydraulic 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 thermohydraulic design criterion

was developed which states that mode of operation under Condition I and II events shall not lead to thermohydrodynamic 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 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 (see Section 4.4.2.2 and Ref. 8) shows that the predicted DNBR is dependent upon the local values of quality and mass velocity. 4.4-30 Rev. 10 WOLF CREEK Operating experience to date has indicated that a flow resistance-allowance for possible crud deposition is not required. There has been no detectable long-

term flow reduction reported at any Westinghouse plant. Inspection of the inside surfaces of steam generator tubes removed from operating plants has confirmed that there is no significant surface deposition that would affect

system flow. Although all of the coolant piping surfaces have not been inspected, the small piping friction contribution to the total system resistance and the lack of significant deposition on piping near steam

generator nozzles support the conclusion that an allowance for piping deposition is not necessary. The effect of crud enters into the calculation of

core pressure drop through the fuel rod frictional component by use of a surface roughness factor. Present analyses utilize a surface roughness value which is a factor of three greater than the best estimate obtained from crud sampling from several operating Westinghouse reactors. The operator has at his disposal several methods of detecting significant RCS flow reduction, these are:

a. Flow meter on each RCS loop.

b. If operating in an automatic control rod mode (T held constant) a reduction in reactor power would be present

for significant reductions in RCS flow.

c. If operating in a manual control rod mode (power held constant) an increase in T across the core would be present for significant reductions in flow.

d. Local changes in flow could be indicated by incore flux maps (assuming significant changes in local power), and

e. Core exit thermocouple readings.

The operator will verify flow, perform calorimetric power checks, and perform incore flux maps as required by the Technical Specifications. Tests were performed at Batelle, Pacific Northwest Laboratories to investigate postulated flow blockages in fuel rod bundles caused by clad ballooning. VIPRE

predictions are in very good agreement with the test data. The code correctly

predicts the velocity decrease just before the blockage, the expected

acceleration in the blockage throat, the expansion loss at the blockage exit, and the subsequent downstream recovery for various sizes of blockage. From a review of the open literature, it is concluded that flow perturbations caused by flow blockage in "open lattice cores" similar to the Westinghouse cores are confined to the vicinity of the blockage. For a flow blockage in a

single flow cell, Ohtsubo, et al. (Ref. 78) show that the mean bundle velocity is approached asymptotically about 4 inches downstream from the blockage. Similar results were also found 4.4-31 Rev. 10 WOLF CREEK for two and three cells completely blocked. Basmer, et al. (Ref. 79) 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 plants were operating at full power and nominal steady state

conditions, as specified in Table 4.4-1, a reduction in local mass velocity greater than 88 percent would be required to reduce the DNBR to the safety analysis 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 effect DNBR at all. Coolant flow blockages induce local crossflows 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 cross-flow

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. The crossflow velocity required to exceed fluid elastic stability limits is dependent on the axial location of the blockage and the characterization of the crossflow (jet flow or not). These limits are greater

than those for vibratory fuel rod wear. Cross-flow 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 (see Section 4.2.3.1). 4.4.5 TESTING AND VERIFICATION

4.4.5.1 Tests Prior to Initial Criticality A reactor coolant flow test is performed following fuel loading but prior to initial criticality. Coolant loop elbow differential pressure readings are

obtained in this test. This data allows 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. Chapter

14.0 describes the initial test programs. 4.4.5.2 Initial Power and Plant Operation Core power distribution measurements are made at several core power levels (see Chapter 14.0). These tests are used to ensure that conservative peaking

factors are used in the core thermal and hydraulic analysis. 4.4-32 Rev. 10 WOLF CREEK Additional demonstration of the overall conservatism of the THINC analysis was obtained by comparing THINC predictions to incore thermocouple measurements (Ref. 80). These measurements were performed on the Zion reactor. No further in-reactor testing is planned. 4.4.5.3 Component and Fuel Inspections Inspections performed on the manufactured fuel are described in Section 4.2.4. Fabrication measurements critical to thermal and hydraulic analysis are obtained to verify that the engineering hot channel factors in the design analyses (see Section 4.4.2.2.4) are met. 4.4.6 INSTRUMENTATION REQUIREMENTS

4.4.6.1 Incore Instrumentation Instrumentation is located in the core so that by correlating movable neutron detector information with fixed thermocouple information radial, axial, and

azimuthal core characteristics may be obtained for all core quadrants. The incore instrumentation system is comprised of thermocouples, positioned to measure fuel assembly coolant outlet temperatures at preselected positions, and fission chamber detectors positioned in guide thimbles which run the length of selected fuel assemblies to measure the neutron flux distribution. Figure 4.4-21 shows the number and location of instrumented assemblies in the core. The core-exit thermocouples can provide a backup to the flux monitoring

instrumentation for monitoring power distribution. The movable incore neutron detector system would be used for more detailed mapping if the thermocouple system were to indicate an abnormality. These two

complementary systems are more useful when taken together than either system

alone would be. The incore instrumentation system is described in more detail in Section 7.7.1.9. The incore instrumentation is provided to obtain data from which fission power density distribution in the core, and fuel burnup distribution may be determined. 4.4.6.2 Overtemperature and Overpower DT Instrumentation The Overtemperature D T trip protects the core against low DNBR. The Overpower D T trip protects against excessive power (fuel rod rating protection). As discussed in Section 7.2.1.1.2, factors included in establishing the Overtemperature D T and Overpower D T trip setpoints includes the reactor coolant temperature in each loop and the axial distribution of core power through the use of the two section excore neutron detectors. 4.4.6.3 Instrumentation to Limit Maximum Power Output The output of the three ranges (source, intermediate, and power) of detectors, with the electronics of the nuclear instruments, are used to limit the maximum

power output of the reactor within their respective ranges. 4.4-33 Rev. 10 WOLF CREEK There are six radial locations containing a total of eight neutron flux

detectors installed around the reactor in the primary shield, two proportional

counters for the source range installed on opposite "flat" portions of the core

containing the primary startup sources at an elevation approximately 1/4 of the

core height. Two compensated ionization chambers for the intermediate range, located in the same instrument wells and detector assemblies as the source

range detectors, are positioned at an elevation corresponding to 1/2 of the

core height. Four dual section uncompensated ionization chamber assemblies for

the power range are installed vertically at the four corners of the core and

are located equidistant from the reactor vessel at all points and, to minimize

neutron flux pattern distortions, within 1 foot of the reactor vessel. Each

power range detector provides two signals corresponding to the neutron flux in

the upper and in the lower sections of a core quadrant. The three ranges of

detectors are used as inputs to monitor neutron flux from a completely shutdown

condition to 120 percent of full power with the capability of recording

overpower excursions up to 200 percent of full power.

The output of the power range channels is used for:

a. The rod speed control function

b. Alerting the operator to an excessive power unbalance

between the quadrants

c. Protecting the core against the consequences of rod

ejection accidents, and

d. Protecting the core against the consequences of adverse

power distributions resulting from dropped rods

Details of the neutron detectors and nuclear instrumentation design and the

control and trip logic are given in Chapter 7.0. The limits on neutron flux

operation and trip setpoints are given in the Technical Specifications.

4.4.6.4 Digital Metal Impact Monitoring System (DMIMS-DX TM) General System Description

The metal impact monitoring system (DMIMS-DX TM) at Wolf Creek is designed to detect loose parts in the reactor coolant system. The system consists of

sensors preamplifiers, signal conditioners, signal processors and a display.

It contains 12 active instrument channels, each comprised of a piezoelectric

accelerometer (sensor), signal conditioning and diagnostic equipment.

Redundant sensors are fastened mechanically to the reactor coolant system at

each of the following potential loose parts collection regions:

Reactor pressure vessel - upper head region

Reactor pressure vessel - lower head region

Each steam generator - reactor coolant inlet region

4.4-34 Rev. 29 WOLF CREEK The output signal from each accelerometer is passed through a preamplifier and an amplifier. The amplified signal is processed through a discriminator to eliminate noises and signals that are not indicative of loose parts. The processed signal is compared to a preset alarm setpoint. Loose parts detection is accomplished at a frequency range of 1 kHz to 20 kHz, where background signals from the RCS are acceptable. Spurious alarming from control rod stepping is prevented by a module that detects CRDM motion commands and automatically inhibits alarms during control rod stepping. If a measured signal exceeds the preset alarm level, audible and visible alarms in the control room are activated. Digital signal processors record the times that the first and subsequent impact signals reach various sensors. This timing information provides a basis for locating the loose part. The DMIMS-DX TM also has a provision for audio monitoring of any channel. The audio signal can be compared to a previously recorded audio signal, if desired. The on-line sensitivity of the LPMS is such that the system will detect a loose part that weighs from 0.25 to 30 pounds and impacts with a kinetic energy of 0.5 feet pounds on the inside surface of the RCS pressure boundary within 3 feet of a sensor. The DMIMS-DX TM audio and visual alarm capability will remain functional after an Operating Basis Earthquake (OBE). All of the DMIMS-DX TM components are qualified for structural integrity during a Safe Shutdown Earthquake (SSE) and will not mechanically impact any safety-related equipment. The components of the loose parts monitoring system are designed for the environmental conditions specified in Table 4.4-5. The DMIMS-DX TM components outside containment are located in a mild environment. In addition, the equipment inside containment is designed to remain functional through normal radiation exposures anticipated during a 40-year operating lifetime. Physical separation of the two instrument channels, associated with the redundant sensors at each reactor coolant system location, exists from each sensor to a location accessible during power operation. Capabilities exist for subsequent periodic online channel checks and channel functional tests and for offline channel calibrations at refueling outages. The loose parts monitoring system complies with NRC Regulatory Guide 1.133, except as noted in USAR Chapter 3, Appendix A, Conformance to NRC Regulatory Guides.Operators were trained in the operation and maintenance of the LPMS prior to Refuel 14 criticality. This consisted of a formal training session onsite by either the onsite training staff or by the vendor using the DMIMS-DX TM that was set up for training in a training lab prior to installation. The vendor can also provide service personnel on short notice to assist the operating staff in operation or maintenance of the equipment and analysis of loose parts signals, as may be required. 4.4-35 Rev. 19 WOLF CREEK 4.

4.7 REFERENCES

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32. Cohen, I., Lustman, B. and Eichenberg, D., "Measurement of the Thermal Conductivity of Metal-Clad Uranium Oxide Rods During Irradiation," WAPD-228, 1960.

33. Clough, D. J. and Sayers, J. B., "The Measurement of the Thermal Conductivity of UO 2 under Irradiation in the Temperature Range 150 C," AERE-R-4690, UKAEA Research Group, Harwell, December, 1964.

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36. Balfour, M. G., Christensen, J. A. and Ferrari, H. M., "In-Pile Measurement of UO 2 Thermal Conductivity," WCAP-2923, March, 1966. 4.4-38 Rev. 10 WOLF CREEK

37. Howard, V. C. and Gulvin, T. G., "Thermal Conductivity Determinations on Uranium Dioxide by a Radial Flow

Method," UKAEA IG-Report 51, November, 1960.

38. Lucks, C. F. and Deem, H. W., "Thermal Conductivity and Electrical Conductivity of UO 2 " in Progress Reports Relating to Civilian Applications, BMI-1448 (Rev.) for

June, 1960; BMI-1489 (Rev.) for December, 1960 and BMI-

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39. Daniel, J. L., Matolich, Jr., J. and Deem, H. W.

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40. Feith, A. D., "Thermal Conductivity of UO 2 by a Radial Heat Flow Method," TID-21668, 1962.

41. 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

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43. Ainscough, J. B. and Wheeler, M. J., "Thermal Diffusivity and Thermal Conductivity of Sintered Uranium Dioxide," in Proceedings of the Seventh Conference of

Thermal Conductivity, p. 467, National Bureau of

Standards, Washington, 1968.

44. Godfrey, T. G., Fulkerson, W., Killie, T. G., 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.

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46. Bush, A. J., "Apparatus for Measuring Thermal Conductivity to 2500 C," Westinghouse Research

Laboratories Report 64-lP6-401-43 (Proprietary), February, 1965.

47. 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. 4.4-39 Rev.10 WOLF CREEK

48. 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, September, 1968, Pittsburgh.

49. 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.

50. Duncan, R. N., "Rabbit Capsule Irradiation of U0 2 ," CVTR Project, CVNA-142, June, 1962.

51. Nelson, R. C., Coplin, D. H., Lyons, M. F. and Weidenbaum, B., "Fission Gas Release from U0 2 Fuel Rods with Gross Central Melting," GEAP-4572, July, 1964.

52. Gyllander, J. A., "In-Pile Determination of the Thermal Conductivity of U0 2 in the Range 500-2500 C," AE-411, January, 1971.

53. Lyons, M. F., et al., "UO 2 Powder and Pellet Thermal Conductivity During Irradiation," GEAP-5100-1, March, 1966.

54. Coplin, D. H., et al., "The Thermal Conductivity of U0 2 by Direct In-Reactor Measurements," GEAP-5100-6, March, 1968.

55. 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.

56. Stora, J. P., "In-Reactor Measurements of the Integrated Thermal Conductivity of UO 2 - Effect of Porosity," Trans. ANS , 13 , p.p. 137-138 (1970).

57. 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.

58. Poncelet, C. G., "Burnup Physics of Heterogeneous Reactor Lattices," WCAP-6069, June, 1965.

59. Nodvick, 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.

60. Dean, R. A., "Thermal Contact Conductance Between UO 2 and Zircaloy-2," CVNA-127, May, 1962.

61. Ross, A. M. and Stoute, R. L., "Heat Transfer Coefficient Between UO 2 and Zircaloy-2," AECL-1552, June, 1962. 4.4-40 Rev. 10 WOLF CREEK

62. Carter, F. D., "Inlet Orificing of Open PWR Cores," WCAP-7836, January, 1972.

63. Shefcheck, J., "Application of the THINC Program to PWR Design," WCAP-7359-L (Proprietary), August, 1969 and WCAP-7838 (Non-Proprietary), January, 1972.

64. Novendstern, E. H. and Sandberg, R. O., "Single Phase Local Boiling and Bulk Boiling Pressure Drop Correlations," WCAP-2850-L (Proprietary) and WCAP-7916

(Non-Proprietary), April, 1966.

65. Owens, Jr., W. L., "Two-Phase Pressure Gradient," in International Developments in Heat Transfer , Part II , pp. 363-368, ASME, New York, 1961.

66. McFarlane, A. F., "Power Peaking Factors," WCAP-7912-P-A (Proprietary) and WCAP-7912-A (Non-Proprietary),

January, 1975. 67. Miller, R. W., et al., Relaxation of Constant Axial offset Control - FQ Surveillance Technical Specification, WCAP-10216-P-A Rev. 1A, approved version dated February, 1994

68. Vallentine., H. R., "Applied Hydrodynamics," Buttersworth Publishers, London, 1969.

69. Kays, W. M. and London, A. L., "Compact Heat Exchangers," National Press, Palo Alto, 1955.

70. Rowe, D. S., "COBRA-III, a Digital Computer Program for Steady State and Transient Thermal-Hydraulic Analysis of

Rod Bundle Nuclear Fuel Elements," BNWL-B-82, 1971.

71. Weisman, 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," Chem. Eng. Prog. Symp. Ser. 64 , No. 82 , pp. 114-125 (1968).

72. Boure, J. A., Bergles, A. E., and Tong, L. S, "Review of Two-Phase Flow Instability," Nucl. Engr. Design 25, pgs.

165-192, 1973.

73. Lahey, R. T. and Moody, F. J., "The Thermal Hydraulics of a Boiling Water Reactor," American Nuclear Society, 1977.

74. Ishii, M., Saha, P., and Zuber, N., "An Experimental Investigation of the Thermally Induced Flow Oscillations

in Two-Phase Systems," Journal of Heat Transfer, No.

1976, pgs. 612-622.

75. Byron/Braidwood Stations Final Safety Analysis Report, Docket Nos. 50-454, 50-455, 50-456, and 50-457. 4.4-41 Rev. 11 WOLF CREEK

76. South Texas Project Units l and 2 Final Safety Analysis Report, Docket Nos. 50-498 and 50-499.

77. Virgil C. Summer Nuclear Station Final Safety Analysis Report, Docket No. 50-395.

78. Ohtsubo, A., and Uruwashi, S., "Stagnant Fluid due to Local Flow Blockage," J. Nucl. Sci. Technol., 9 , No. 7 , pp. 433-434, (1972).

79. 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," Automwirtschaft , 17 , No. 8 , pp. 416-417, (1972). (In German).

80. Burke, T. M., Meyer, C. E. and Shefcheck J., "Analysis of Data from the Zion (Unit 1) THINC Verification Test," WCAP-8453-A, May, 1976.

81. Kakac, S., Vexiroglu, T. N., Akyuzlu, K., and Berkol, O, "Sustained and Transient Boiling Flow Instabilities in a Cross-Connected Four-Parallel-Channel Upflow System," Proc of 5th International Heat Transfer Conference, Tokyo, September 3-7, 1974.

82. Kao, H. S., Morgan, C. D., and Parker, W. B., "Prediction of Flow Oscillation in Reactor Core Channel," Trans. ANS Vol. 16, pgs. 212-213, 1973.

83. Skaritka, J. (Ed.), "Fuel Rod Bow Evaluation," WCAP-8691, Rev. l (Proprietary) and WCAP-8692, Rev. 1 (Non-

Proprietary), July 1979.

84. "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 R. J. Miller (NRC), NS-EPR-2572, dated March 16, 1982.

85. Letter from C Berlinger, NRC, to E. P. Rahe, Jr., Westinghouse, "Request for Reduction in Fuel Assembly

Burnup Limit for Calculation of Maximum Rod Bow

Penalty", June 18, 1986.

86. (DELETED) 4.4-42 Rev. 11 WOLF CREEK

87. Letter from A.C. Thadani (NRC) to W.J. Johnson (Westinghouse), Jan. 31, 1989,

Subject:

Acceptance for

Referencing of Licensing Topical Report, WCAP-9226-P/9227-NP, "Reactor Core Response to Excessive Secondary Steam Releases." 88. Leech, W.J., et. al., "Revised PAD Code Thermal Safety Model," WCAP-8720, Addendum 2, October 1982.

89. Neises, G.J, (WCNOC), "WCGS Rerating Program Monthly Status Report for January/February 1992 and the Phase 3 Limiting Events Analysis Report", attachment to NS 92-0278, 2/28/92.

90. Letter from C.E.Rossi, NRC, to J.A.Blaisdell, UGRA, "Acceptance for Referencing of Licensing Topical

Report, EPRI NP-2511-CCM, 'VIPRE-01 : A Thermal

Hydraulic Code for Reactor Cores,' Volumes 1,2,3 and

4" May, 1986.

91. Friedland, A. J., Ray, S., "Revised Thermal Design Procedure," WCAP-11397-P-A, April, 1989.

92. Letter from Stone, J. C., USNRC, to Carns, N. S., WCNOC, "Wolf Creek Generating Station - Amendment No. 92 to Facility Operationg License No. NPF-42," dated December 8, 1995.

93. Kennamore, William, et. al., "Core Thermal-Hydraulic Analysis Methodology for the Wolf Creek Generating Station," TR-90-0025 W01, July, 1990. 4.4-43 Rev. 10 WOLF CR EE K TABL E 4.4-1 TH E RMAL AND HYDRAULIC COMPARISON TABL E Comanche Peak WCGS Design Parameters Units 1 and 2 UnitReactor core heat output, MWt 3,4113,565 Reactor core heat output, 10 6Btu/hr 11,63912,164Heat generated in fuel, % 97.497.4System pressure, nominal psia 2,2502,250 System pressure, minimum steady state, psia 2,2202,220 Minimum DNBR at nominal design conditions Typical flow channel 2.082.50 Thimble (cold wall) flow channel 1.742.44Minimum DNBR for design transients >1.30>1.76DNB Correlation "R"WRB-2

(W-3 with modified

spacer factor) Coolant FlowTotal thermal flow rate, gpm 390,214361,296 E ffective flow rate for heat transfer, gpm 367,740337,414

E ffective flow area for heat transfer, ft 2 51.151.3 Average velocity along fuel rods, ft/sec 16.614.7

Average mass velocity, 10 6 lb m/hr ft 2 2.622.31 Comanche PeakWCGS Coolant Temperature Units 1 and 2 Unit Nominal Inlet, oF 558.8553.7 Average rise in vessel, oF 59.465.6 Average rise in core, oF 62.668.6 Average in core, oF 591.8588.0 Average in vessel, oF 588.5586.5 Rev. 13 WOLF CR EE K TABL E 4.4-1 (Sheet 2) Comanche PeakWCGS Design Parameters Units 1 and 2 Unit Heat Transfer Active heat transfer surface area, ft 2 59,70059,742 Average Heat Flux, Btu/hrft 2 189,800198,340 Maximum Heat Flux for normal operation, Btu/hrft 2 440,300(a)460,100(b)Average Linear Power, kW/ft 5.445.68 Peak Linear Power for normal operation, kW/ft 12.6(a)14.48(b)

Peak linear power resulting from overpower transients, operator errors, assuming a maximum overpower of 118%, kW/ft (c) 18.0<22.5 Peak linear power for prevention of centerline melt, kW/ft (d) >18.022.5 Power density, kW per liter of core (e) 104.5109.2 Specific power, kW per kg Uranium (e) 38.440.1 Fuel Central Temperature Peak at peak linear power for

prevention of centrline melt, oF 47004700 Pressure drop (f) Across core, psi 26.1+2.628.0+ 2.6 Across vessel, including nozzle, psi 46.2+4.649.6+ 4.7 (a) This limit is associated with the value of Fq = 2.32 (b) This limit is associated with the value of F q = 2.50 (c) See Section 4.3.2.2.6. (d) See Section 4.4.2.11.6. (e) Based on cold dimensions and 95% of theoretical density fuel.(f) Based on best estimate reactor flow rate, at discussed in Section 5.1. Rev. 13 WOLF CR EE K Table 4.4-2 This Table has been Deleted Rev. 11 WOLF CR EE K TABL E 4.4-3 VOID FRACTIONS AT NOMINAL R E ACTOR CONDITIONS WITH D E SIGN HOT CHANN E L FACTORS Average (%) Maximum (%)Core0.1-Hot Subchannel-3.6 Rev. 11 WOLF CR EE K TABL E 4.4-4 COMPARISON OF THINC-IV AND THINC-I PR E DICTIONS WITH DATA FROM R E PR E S E NTATIV E W E STINGHOUS E TWO AND THR EE LOOP R E ACTORS Improvement (F) Power Percent Full Measured Inletrms (F) (F) for THINC-IV Reactor (MWt) Power Temperature (F) THINC-I THINC-IV over THINC-I 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 98.8 540.2 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 2207.3 100.7 543.0 2.41 2.41 0.00

2213.9 101.0 533.8 2.52 2.44 0.08 Rev. 0 WOLF CR EE K TABL E 4.4-5 LOOS E PARTS MONITORING SYST E M E nvironmental Conditions A. Accelerometers Temperature 40-650°F Humidity 0-100% Radiation 10 18 nvt and 10 8 rad Pressure 69 psig Vibration OB E Atmosphere Air B. Preamplifiers and Cables (inside containment) Temperature-electronics 40-150°F Hardline Cable 40-650°F

Cable inside 40-150°F

containment

Humidity 0-100% Radiation 10 12 nvt and 6x10 6 rad Pressure 69 psig Shock and Vibration OB E Atmosphere Air C. Signal Conditioning Amplifier, Signal Processor, and Associated E quipment (outside of containment) Temperature 40-120°F Radiation 103 rad

Pressure Atmospheric

Humidity 0-95%

Shock and Vibration In accordance with good engineering practice Atmosphere Air Rev. 0 WOLF CREEK 0.12 0.10 f-.-0.08 -LEGEND: 0 26" SPACING 1-PtiASE 0 26" SPACING 2-PHASE c.> 0.06 1--c 1-CDO 0 8 <6o I:Po 0 O.OIJ. I--Q9f/R 0 0 0 0 C2> c9 0.02 I--0 I I I I I I I 0 1.0 2.0 3.0 IJ..O 5.0 6.0 7.0 8.0 Rev. 0 G De WOLF CREEK Re = f:L ( 10-5) UPDATED SAFETY AHALYSIS REPORT FIGURE 4.4-4 TDC VERSUS REYNOLDS NUMBER FOR 26 INCH GRID SPACING I ct. I 1.096 1.001 1.029 1.00 I I. 153 1.000 I. 166 1.000 I I. 223 1.000 I. 126 I. 00 I I I 1.025 1.002 I 0.717 0.999 I -I-I. 120 1.001 1.074 I. 185 1.001 1.000 I. 209 I. 162 1.000 1.001 I. 170 1.188 1.00 I 1.000 I. 161 1.093 1.00 I 1.002 1.025 0.990 1.002 1.002 0.780 0.664 0.999 0.998 f-WOLF CREEK I. 185 1.000 1.065 1.002 1.086 1.002 0.975 I .002 0.563 0.997 KEY: +* + 6h/6h -G/G I. 238 1.000 0.916 0.967 1.001 1.001 0.823 0.466 1.000 0.997 FOR RADIAL POWER DISTRIBUTION NEAR BEGINNI"G OF LIFE, HOT FULL POWER, EQUILIBRIUM XENON CALCULATED =I. 34 Rev. 0 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.4-5 NORMALIZED RADIAL FLOW AND ENTHALPY DISTRIBUTION AT 4 FOOT ELEVATION I I .096 --0.996 I I .026 I. 120 I .002 0.994-I. 155 I .072 I. 188 0.991 0.998 0.988 I. 169 I .212 I. 165 0.990 0.986 0.990 I. 228 I. 173 I. 191 0.985 0.989 0.988 I I. 129 I. 165 I .095 0.992 0.990 0.995 I .024-I .025 0.989 I .000 1.000 1.003 I I 0.715 0.777 0.663 1.022 1.019 1.021 I -WOLF CREEK --I. 189 0.998 I .065 0.998 1.086 0.996 0.973 I. 004-0.562 1.019 KEY: -II--t::. h It::. h G/G I. 24-2 0.981 0.916 0.963 1.009 1.006 0.819 0.4-68 1.018 1.018 FOR RADIAL POWER DISTRIBUTION NEAR BEGINNING OF LIFE, HOT FULL POWER, EQUILIBRIUM XENON CALCULATED F N = I. 34-LiH WOLF CREEK Rev. 0 UPDATED SAFETY ANALYSIS REPORT FIGURE 4.4-6 NORMALIZED RADIAL FLOW AND ENTHALPY DISTRIBUTION AT 8 FOOT ELEVATION t-I t I 1.097 0.995 1.026 0.999 I I. 157 0.991 I 1.170 0.990 I. 231 0.987 I. 130 0.993 1.023 1.000 0. 711 1.016 I I. 121 0.993 1.073 I. 189 0.996 0.989 I. 215 I. 166 0.980 0.991 I. 175 I. 193 0.990 0.989 I. 165 1.095 0.991 0.995 I. 024-0.987 1.000 1.002 0.774 0.660 I. 013 1.019 WOLF CREEK 1.190 0.990 1.066 0.997 1.087 0.996 0.971 1.003 0.560 1.025 KEY: -6h/6h G/G I. 24-3 0.987 0.914-0.961 1.005 1.003 0.817 0.4-69 1.011 1.030 FOR RADIAL POWER DISTRIBUTION NEAR BEGINNING OF LIFE, HOT FULL POWER. EQUILIBRIUM XENON CALCULATED F N = I. 34 . 6H Rev. 0 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.4-7 NORMALIZED RADIAL FLOW AND ENTHALPY DISTRIBUTION AT 12 FOOT ELEVATION -CORE EXIT z:: 0 (..) < 0::: u.. 0 0 > REGION I NO BUBBLE DETACHMENT WOLF CREEK LOCAL BOILING >-0... --' <C ::r: I-:z: w :z: 0 I->-<C 0... 0::: --' ::::> <( I-::r: <C I-(I) :z: w (I) --' ""' <C --' ::::> ::::> C>' en w REG I ON II BUBBLE DETACHMENT 0 w 0::: :z: =>0 I--<CI-0::: <C we::: C...::::>W

r-c

::

W<C::::;, 1-(1)1-<C 0 (I) 0::: ---' w :::><C... C>'::::;,:::;: -C>'W __lWI-BULK BOILING VOID FRACTION PREDICTED FROM THERMODYNAMIC QUALITY WITH NO SLIP Rev. 0 WOLF CREEK THERMODYNAMIC QUALITY, X = H-H sAT f Hg -H sAT FIGURE 4.4-8 VOID FRACTION VERSUS THERMODYNAMIC QUALITY WOLF CREEK 0.07 LEGEND: PERCENT PERCENT 0 u THEORETICAL THEORETICAL 0.06 \* DENSITY DENSITY u 0 GYLLANDER 96.4 0 HOWARD & GULVIN 96.0 0 I 2

• ASAMOTO 95.0
• LUCKS & DEEM 96.5 u \o 0 GODFREY 93.4 0 DANIEL ET AL. 96.5 "'-., $ 0.05
• STORA 92.2
• FEITH 97.9 E 0 OA.D. BUSH 94.4 0 VOGT ET AL. 97.0 ::.::: o,t
• ASAMOTO 91.0 6 NISHIJIMA 95.0 > \l KRUGER 93.7 .. WHEELER & 97.0 1-0' --! ::J: (}rrl Q 0;:;::) :::03: "C :::0 )> 0 ,., r :J:II ("") 1-i -i() 1::1\t rrl 0 0 oz m 0 ., -i c H oo ("") G) 1::1\tt:"" rrl -i c zc.DH :::0 I< CI)(J1< rrl n :J:II" -i -i ..c -< --!-< . Zt"'l :J:IIt"'t

-:::c -+- rrl 0 I I< 0..,., c.D m :::0 rrl c 1-4 -i C) m > g'l AINSCHOUGH 1-0.04 u o't SEE REFERENCES FOR THERMAL-HYDRAULIC SECTION :J c 6 z 0 u o ...J 0.03 <t

• o' o*
• t
• 2 6 0 a: w 0 ::I: 'Q \l 0 1-* \l
• 0.02 6 .. 0.01 0 1200 1600 2400 800 400 2000 2800 3200 1--!1'0 " ("") )>-1::1\t ro "C 0 )> " -i )> 1-i TEMPERATURE (T), °C 0 WOLF CREEK 586.5 553.7 619.3 557 520 540 560 580 600 620020406080100PERCENT POWERTEMPERATURE (o F)TCOLD T AVG THOT Rev. 13WOLF CREEKUPDATED SAFETY ANALYSIS REPORTFIGURE 4.4-10REACTOR COOLANT SYSTEMTEMPERATURE - PERCENT POWER MAP 0 -t--<( a:: co z: 0 ::i: :=:> ::::E z: ::i: WOLF CREEK 1.8 I. 7 1...: -5 BOL I +5 ,.., j .... -5 EOL I +5 ,..I I. 6 0 0 0 g 0 0 0 I, 5 0 6 o Oo 0 0 0 eo o 00 0 o 0 00 0 I . lf coo 1.3

-20.0 -10.0 0 AXIAL OFFSET (PERCENT) 10.0 Rev. 0 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4.4-11 100% POWER SHAPES EVALUATED AT CONDITIONS REPRESENTATIVE OF LOSS OF FLOW, ALL EVALUATED. WITH F 1-1 = 1. 55 R p 1 2 3 4

• 5 6 t!J 1 [!] 9
• 10 0 11
• 12 13 14 15 Wolf Creek N M L K J H G F E D C 8 A
• 0 [!] 0
• 0
• 0 * [!] 0 0
• D * * [!] 0
• D
• 0 * * [!] [!] I!J D
• 0 (!] [!] 0 *
• 0 0
• 0 [!] * * * [!] D [!] 0 * [!] 0 THERMOCOUPLE e MOVEABLE DETECTOR
• 0 0 * *
• D D *
• D D
• 0 * *
• 0 * (!]
• 0 *
• 0
• 0 0 *
• 0
• D **
• 0
• 0 REV. 2 WOLF CREEK UPDATED SAFETY ANALYSIS REPORT FIGURE 4. 4-21 MOVEABLE DETECTOR AND THERMOCOUP LOCA f IONS

<.11 n::s .... I.Cl .... 0'1

• n::s <li
• ..:::: u 0 0 0 0 0'1 0 N 1-L.l... 0 n::s 3:"'0 0 I co 0:::. :I: 0 ..... -0 0 -c 0'1 . 0 0 cc 0 0 ....... 0 0 <..C 0 0 1.!) 0 0 o::T 0 0 ("I") 0 I.Cl.--=-=

0'1 n::s ..... 0 -1< ..... u 0 N 0 0 0 WOLF CREEK 0 0 ....... 0 0 <..C 0 0 1.!) 0 0 o::T . 0 0 ("I") 0 0 N 0 0 r-0 0 0 0 0 0 0 UPDATED SAFETY ANALYSiS REPORT FIGURE 4.4-22 -:::::> 1-c:c. >< :::::> _J L.l... 1-<. l.U. :I: -1. < u -0:::. u 0 I..W 1-u -0. I..W 0:::. 0.. 0 I..W 0:::. 0.. <..C 0 --L.l... :.U 0 MEASURED VERSUS PREDICTED CRITICAl HEAT FlUX WRB-2 CORRELATION WOLF CREEK 4.5 REACTOR MATERIALS 4.5.1 CONTROL ROD SYSTEM STRUCTURAL MATERIALS 4.5.1.1 Materials Specifications 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, nickel-chromium-iron, and cobalt based alloys. In the case of stainless steels, only austenitic and martensitic stainless steels are used. The martensitic stainless steels are not used in the heat treated conditions which cause susceptibility to stress corrosion cracking or accelerated corrosion in the Westinghouse pressurized water reactor water chemistry.

a. Pressure vessel

All pressure containing materials comply with Section III of the ASME Boiler and Pressure Vessel Code, and are fabricated from austenitic (Type 304) stainless steel.

b. Coil stack assembly

The coil housings require a magnetic material. Both low carbon cast steel and ductile iron have been successfully

tested for this application. The choice, made on the

basis of cost, indicates that ductile iron will be

specified on the control rod drive mechanism (CRDM). The

finished housings are zinc plated or flame sprayed 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 outside

diameter. The result is a well insulated coil capable of

sustained operation at 200°C.

c. Latch assembly

Magnetic pole pieces are fabricated from Type 410 stainless steel. All 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 nickel-chromium-iron alloy (Inconel-X). Latch

arm tips are clad with Stellite-6 to provide improved

wearability. Hard chrome plate and Stellite-6 are used

selectively for bearing and wear surfaces. 4.5-1 Rev. 0 WOLF CREEK

d. Drive rod assembly The drive rod assembly utilizes a Type 410 stainless steel drive rod. The coupling is machined from Type 403 stainless steel. Other parts are Type 304 stainless steel with the exception of the springs, which are nickel-chromium-iron alloy, and the locking button, which is Haynes 25.

4.5.1.2 Fabrication and Processing of Austenitic Stainless Steel Components The discussions provided in Section 5.2.3 concerning the processes, inspections, and tests on austenitic stainless steel components to ensure

freedom from increased susceptibility to intergranular corrosion caused by

sensitization, and the discussions provided in Section 5.2.3 on the control of

welding of austenitic stainless steels, especially control of delta ferrite, are applicable to the austenitic stainless steel pressure housing components of the CRDM. 4.5.1.3 Contamination Protection and Cleaning of Austenitic Stainless Steel The CRDMs are cleaned prior to delivery in accordance with the guidance of ANSI N45.2.1. Process specifications in packaging and shipment are discussed in

Section 5.2.3. Westinghouse personnel do conduct surveillance to ensure that

manufacturers and installers adhere to appropriate requirements, as discussed

in Section 5.2.3. 4.5.2 REACTOR INTERNALS MATERIALS 4.5.2.1 Materials Specifications 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 radial support

key bolts, which are fabricated of Inconel-750. These materials are listed in

Table 5.2-4. There are no other materials used in the reactor internals or

core support structures which are not otherwise included in ASME Code, Section

III, Appendix I. 4.5.2.2 Controls on Welding The discussions provided in Section 5.2.3 are applicable to the welding of reactor internals and core support components. 4.5-2 Rev. 0 WOLF CREEK 4.5.2.3 Nondestructive Examination of Wrought Seamless Tubular Products and Fittings The nondestructive examination of wrought seamless tubular products and fittings is in accordance with Section III of the ASME Code. 4.5.2.4 Fabrication and Processing of Austenitic Stainless Steel Components The discussions provided in Section 5.2.3 and Appendix 3A verify conformance of reactor internals and core support structures with Regulatory Guide 1.44. The discussions provided in Section 5.2.3 and Appendix 3A verify conformance of reactor internals and core support structures with Regulatory Guide 1.31. The discussion provided in Appendix 3A verifies conformance of reactor internals with Regulatory Guide 1.34. The discussion provided in Appendix 3A verifies conformance of reactor internals and core support structures with Regulatory Guide 1.71. 4.5.2.5 Contamination Protection and Cleaning of Austenitic Stainless Steel The discussions provided in Section 5.2.3 and Appendix 3A are applicable to the reactor internals and core support structures and verify conformance with ANSI N45 specifications and Regulatory Guide 1.37. 4.5-3 Rev. 0 WOLF CREEK 4.6 FUNCTIONAL DESIGN OF REACTIVITY CONTROL SYSTEMS 4.6.1 INFORMATION FOR CONTROL ROD DRIVE SYSTEM (CRDS)

The CRDS is described in Section 3.9(N).4.1. Figures 3.9(N)-5 and 3.9(N)-6 provide the details of the control rod drive mechanisms, and Figure 4.2-8 provides the layout of the CRDS. No hydraulic system is associated with its functioning. The instrumentation and controls for the reactor trip system are

described in Section 7.2, and the reactor control system is described in Section 7.7. 4.6.2 EVALUATION OF THE CRDS The CRDS has been analyzed in detail in a failure mode and effects analysis (Ref. 1). This study, and the analyses presented in Chapter 15.0, demonstrates that the CRDS performs its intended safety function, reactor trip, by putting

the reactor in a subcritical condition when a safety system setting is

approached, with any assumed credible failure of a single active component.

The essential elements of the CRDS (those required to ensure reactor trip) are

isolated from nonessential portions of the CRDS (the rod control system) as

described in Section 7.2. Despite the extremely low probability of a common mode failure impairing the ability of the reactor trip system to perform its safety function, analyses have been performed in accordance with the requirements of WASH-1270. These analyses, documented in References 2 and 3, have demonstrated that acceptable safety criteria would not be exceeded even if the CRDS were rendered incapable

of functioning during a reactor transient for which their function would

normally be expected. The design of the control rod drive mechanism is such that failure of the control rod drive mechanism cooling system will, in the worst case, result in an individual control rod trip or a full reactor trip (see Section 9.2). 4.6.3 TESTING AND VERIFICATION OF THE CRDS

The CRDS is extensively tested prior to its operation. These tests may be subdivided into five categories: 1) prototype tests of components, 2) prototype

CRDS tests, 3) production tests of components following manufacture and prior

to installation, 4) onsite preoperational and initial startup tests, and 5) periodic inservice tests. These tests, which are described in Sections 3.9(N).4.4, 4.2, 14.2, and the Technical Specifications , are conducted to

verify the operability of the CRDS when called upon to function. 4.6-1 Rev. 0 WOLF CREEK 4.6.4 INFORMATION FOR COMBINED PERFORMANCE OF REACTIVITY SYSTEMS As is indicated in Chapter 15.0, the only postulated events which assume credit for reactivity control systems other than a reactor trip to render the plant subcritical are the steam line break, feedwater line break, and loss-of-coolant accident. The reactivity control systems for which credit is taken in these accidents are the reactor trip system and the safety injection system (SIS).

Additional information on the CRDS is presented in Section 3.9(N).4 and on the SIS in Section 6.3. Note that no credit is taken for the boration capabilities

of the chemical and volume control system (CVCS) as a system in the analysis of transients presented in Chapter 15.0. Information on the capabilities of the CVCS is provided in Section 9.3.4. The adverse boron dilution possibilities due to the operation of the CVCS are investigated in Section 15.4.6. Prior proper operation of the CVCS has been presumed as an initial condition to

evaluate transients, and appropriate Technical Specifications have been

prepared to ensure the correct operation or remedial action. 4.6.5 EVALUATION OF COMBINED PERFORMANCE The evaluations of the steam line break, feedwater line break, and the loss-of-coolant accident, which presume the combined actuation of the reactor trip

system to the CRDS and the SIS, are presented in Sections 15.1.5, 15.2.8, and

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

poison.Nondiverse but redundant types of equipment are utilized only in the processing of the incoming sensor signals into appropriate logic, which initiates the protective action. This equipment is described in detail in Sections 7.2 and 7.3. In particular, note that protection from equipment failures is provided

by redundant equipment and periodic testing. Effects of failures of this

equipment have been extensively investigated as reported in Reference 4. The

failure mode and effects analysis described in this reference verifies that any

single failure will not have a deleterious effect on the engineered safety

features actuation system. Adequacy of the emergency core cooling system and

SIS performance under faulted conditions is verified in Section 6.3. 4.6-2 Rev. 0 WOLF CREEK 4.

6.6 REFERENCES

1. Shopsky, W. E., "Failure Mode and Effects Analysis (FMEA) of the Solid State Full Length Rod Control System," WCAP-8976, August 1977.

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

3. Gangloff, W. C. and Loftus, W. D., "An Evaluation of Solid State Logic Reactor Protection in Anticipated

Transients," WCAP-7706-L (Proprietary) and WCAP-7706 (Non-Proprietary), July 1971.

4. Mesmeringer, J.C., "Failure Mode and Effects Analysis (FMEA) of the Engineered Safety Features Actuation

System," WCAP-8584, Revision 1 (Proprietary) and WCAP-

8760, Revision 1 (Non-Proprietary), February 1980. 4.6-3 Rev. 0}}