ML18053A725

Duke Energy Wsl III Units 1 & 2 COL (Updated Final Safety Analysis Report) Rev.1 - UFSAR Chapter 04 - Reactor

ML18053A725
Person / Time
Site:	Lee
Issue date:	12/19/2017
From:	Donahue J Duke Energy Carolinas
To:	Office of New Reactors
Hughes B
References
DUKE, DUKE.SUBMISSION.15, LEE.NP, LEE.NP.1
Download: ML18053A725 (192)

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UFSAR Table of Contents 1 Introduction and General Description of the Plant 2 Site Characteristics 3 Design of Structures, Components, Equipment and Systems 4 Reactor 5 Reactor Coolant System and Connected Systems 6 Engineered Safety Features 7 Instrumentation and Controls 8 Electric Power 9 Auxiliary Systems 10 Steam and Power Conversion 11 Radioactive Waste Management 12 Radiation Protection 13 Conduct of Operation 14 Initial Test Program 15 Accident Analyses 16 Technical Specifications 17 Quality Assurance 18 Human Factors Engineering 19 Probabilistic Risk Assessment UFSAR Formatting Legend Description Original Westinghouse AP1000 DCD Revision 19 content Departures from AP1000 DCD Revision 19 content Standard FSAR content Site-specific FSAR content Linked cross-references (chapters, appendices, sections, subsections, tables, figures, and references)

4.1 Summary Description .................................................................................... 4.1-1 4.1.1 Principal Design Requirements .................................................... 4.1-3 4.1.2 Combined License Information ..................................................... 4.1-4 4.1.3 References ................................................................................... 4.1-4 4.2 Fuel System Design ...................................................................................... 4.2-1 4.2.1 Design Basis ................................................................................. 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 Structural Design ................................ 4.2-4 4.2.1.6 In-core Control Components ....................................... 4.2-6 4.2.1.7 Surveillance Program .................................................. 4.2-7 4.2.2 Description and Design Drawings ................................................ 4.2-8 4.2.2.1 Fuel Rods .................................................................... 4.2-8 4.2.2.2 Fuel Assembly Structure ............................................. 4.2-9 4.2.2.3 In-core Control Components ..................................... 4.2-13 4.2.3 Design Evaluation ....................................................................... 4.2-17 4.2.3.1 Cladding .................................................................... 4.2-17 4.2.3.2 Fuel Materials Considerations .................................. 4.2-19 4.2.3.3 Fuel Rod Performance .............................................. 4.2-20 4.2.3.4 Spacer Grids ............................................................. 4.2-23 4.2.3.5 Fuel Assembly .......................................................... 4.2-23 4.2.3.6 Reactivity Control Assemblies and Burnable Absorber Rods .......................................................... 4.2-25 4.2.4 Testing and Inspection Plan ....................................................... 4.2-27 4.2.4.1 Quality Assurance Program ...................................... 4.2-27 4.2.4.2 Quality Control .......................................................... 4.2-27 4.2.4.3 Letdown Radiation Monitoring .................................. 4.2-29 4.2.4.4 In-core Control Component Testing and Inspection .................................................................. 4.2-29 4.2.4.5 Tests and Inspections by Others .............................. 4.2-30 4.2.4.6 Inservice Surveillance ............................................... 4.2-30 4.2.4.7 Onsite Inspection ...................................................... 4.2-30 4.2.5 Combined License Information ................................................... 4.2-31 4.2.6 References ................................................................................. 4.2-31 4.3 Nuclear Design .............................................................................................. 4.3-1 4.3.1 Design Basis ................................................................................. 4.3-1 4.3.1.1 Fuel Burnup ................................................................ 4.3-2 4.3.1.2 Negative Reactivity Feedbacks (Reactivity Coefficients) ................................................................ 4.3-2 4.3.1.3 Control of Power Distribution ...................................... 4.3-3 4.3.1.4 Maximum Controlled Reactivity Insertion Rate ........... 4.3-3 4.3.1.5 Shutdown Margins ...................................................... 4.3-4 4.3.1.6 Stability ....................................................................... 4.3-5 4.3.1.7 Anticipated Transients Without Scram (ATWS) .......... 4.3-6 4-i Revision 1

4.3.2.1 Nuclear Design Description ........................................ 4.3-6 4.3.2.2 Power Distribution ....................................................... 4.3-7 4.3.2.3 Reactivity Coefficients ............................................... 4.3-16 4.3.2.4 Control Requirements ............................................... 4.3-19 4.3.2.5 Control Rod Patterns and Reactivity Worth .............. 4.3-24 4.3.2.6 Criticality of the Reactor During Refueling ................ 4.3-25 4.3.2.7 Stability ..................................................................... 4.3-26 4.3.2.8 Vessel Irradiation ...................................................... 4.3-30 4.3.3 Analytical Methods ..................................................................... 4.3-30 4.3.3.1 Fuel Temperature (Doppler) Calculations ................. 4.3-31 4.3.3.2 Macroscopic Group Constants .................................. 4.3-31 4.3.3.3 Spatial Few-Group Diffusion Calculations ................ 4.3-32 4.3.4 Combined License Information ................................................... 4.3-33 4.3.5 References ................................................................................ 4.3-33 4.4 Thermal and Hydraulic Design ...................................................................... 4.4-1 4.4.1 Design Basis ................................................................................. 4.4-1 4.4.1.1 Departure from Nucleate Boiling Design Basis ........... 4.4-1 4.4.1.2 Fuel Temperature Design Basis ................................. 4.4-2 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 of Thermal and Hydraulic Design of the Reactor Core .............................................................................................. 4.4-3 4.4.2.1 Summary Comparison ................................................ 4.4-3 4.4.2.2 Critical Heat Flux Ratio or DNBR and Mixing Technology ................................................................. 4.4-4 4.4.2.3 Linear Heat Generation Rate ..................................... 4.4-8 4.4.2.4 Void Fraction Distribution ............................................ 4.4-8 4.4.2.5 Core Coolant Flow Distribution ................................... 4.4-9 4.4.2.6 Core Pressure Drops and Hydraulic Loads ................ 4.4-9 4.4.2.7 Correlation and Physical Data .................................... 4.4-9 4.4.2.8 Thermal Effects of Operational Transients ............... 4.4-11 4.4.2.9 Uncertainties in Estimates ........................................ 4.4-12 4.4.2.10 Flux Tilt Considerations ............................................ 4.4-13 4.4.2.11 Fuel and Cladding Temperatures ............................. 4.4-14 4.4.3 Description of the Thermal and Hydraulic Design of the Reactor Coolant System .......................................................................... 4.4-16 4.4.3.1 Plant Configuration Data ........................................... 4.4-16 4.4.3.2 Operating Restrictions on Pumps ............................. 4.4-17 4.4.3.3 Power-Flow Operating Map (Boiling Water Reactor BWR) ........................................................................ 4.4-17 4.4.3.4 Temperature-Power Operating Map (PWR) ............. 4.4-17 4.4.3.5 Load Following Characteristics ................................. 4.4-17 4.4.3.6 Thermal and Hydraulic Characteristics Summary Table ......................................................................... 4.4-17 4-ii Revision 1

4.4.4.1 Critical Heat Flux ....................................................... 4.4-17 4.4.4.2 Core Hydraulics ........................................................ 4.4-17 4.4.4.3 Influence of Power Distribution ................................. 4.4-18 4.4.4.4 Core Thermal Response ........................................... 4.4-20 4.4.4.5 Analytical Methods .................................................... 4.4-20 4.4.4.6 Hydrodynamic and Flow Power Coupled Instability ................................................................... 4.4-21 4.4.4.7 Fuel Rod Behavior Effects from Coolant Flow Blockage ................................................................... 4.4-23 4.4.5 Testing and Verification .............................................................. 4.4-24 4.4.5.1 Tests Prior to Initial Criticality .................................... 4.4-24 4.4.5.2 Initial Power and Plant Operation ............................. 4.4-24 4.4.5.3 Component and Fuel Inspections ............................. 4.4-24 4.4.6 Instrumentation Requirements ................................................... 4.4-24 4.4.6.1 Incore Instrumentation .............................................. 4.4-24 4.4.6.2 Overtemperature and Overpower T Instrumentation .................................................... 4.4-25 4.4.6.3 Instrumentation to Limit Maximum Power Output ..... 4.4-26 4.4.6.4 Digital Metal Impact Monitoring System .................... 4.4-26 4.4.7 Combined License Information ................................................... 4.4-27 4.4.8 References ................................................................................. 4.4-27 4.5 Reactor Materials .......................................................................................... 4.5-1 4.5.1 Control Rod and Drive 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 Other Materials ........................................................... 4.5-2 4.5.1.4 Contamination Protection and Cleaning of Austenitic Stainless Steel ............................................................ 4.5-2 4.5.2 Reactor Internal and Core Support Materials ............................... 4.5-3 4.5.2.1 Materials Specifications .............................................. 4.5-3 4.5.2.2 Controls on Welding .................................................... 4.5-3 4.5.2.3 Nondestructive Examination of 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 ............................................................ 4.5-4 4.5.3 Combined License Information ..................................................... 4.5-4 4.6 Functional Design of Reactivity Control Systems ......................................... 4.6-1 4.6.1 Information for Control Rod Drive System .................................... 4.6-1 4.6.2 Evaluations of the Control Rod Drive System ............................... 4.6-1 4.6.3 Testing and Verification of the Control Rod Drive System ............ 4.6-2 4.6.4 Information for Combined Performance of Reactivity Systems .... 4.6-2 4.6.5 Evaluation of Combined Performance .......................................... 4.6-3 4.6.6 Combined License Information ..................................................... 4.6-3 4-iii Revision 1

2 Analytical Techniques in Core Design ............................................................ 4.1-9 3 Design Loading Conditions for Reactor Core Components ......................... 4.1-10 1 [Reactor Core Description (First Cycle)]* ..................................................... 4.3-38 2 [Nuclear Design Parameters (First Cycle)]* ................................................. 4.3-41 3 [Reactivity Requirements for Rod Cluster Control Assemblies]* .................. 4.3-43 4 Not Used ...................................................................................................... 4.3-44 5 Stability Index for Pressurized Water Reactor Cores with a 12-Foot Height ........................................................................................................... 4.3-45 6 Typical Neutron Flux Levels (n/cm2/s) at Full Power ................................... 4.3-46 7 Comparison of Measured and Calculated Doppler Defects ......................... 4.3-47 8 Comparison of Measured and Calculated AG-in-CD Rod Worth ................. 4.3-48 9 Comparison of Measured and Calculated Moderator Coefficients at HZP, BOL ..................................................................................................... 4.3-49 1 Thermal and Hydraulic Comparison Table (AP1000, AP600 and a Typical Westinghouse XL Plant) ............................................................................... 4.4-34 2 Void Fractions At Nominal Reactor Conditions With Design Hot Channel Factors ......................................................................................................... 4.4-36 4-iv Revision 1

2 Fuel Assembly Outline ................................................................................. 4.2-34 3 Fuel Rod Schematic ..................................................................................... 4.2-35 4 Top Grid Sleeve Detail ................................................................................. 4.2-36 5 Intermediate Grid to Thimble Attachment Joint ............................................ 4.2-37 6 Intermediate Flow Mixer Grid to Thimble Attachment .................................. 4.2-38 7 Grid Thimble to Bottom Nozzle Joint............................................................ 4.2-39 8 Rod Cluster Control and Drive Rod Assembly With Interfacing Components ................................................................................................. 4.2-40 9 Rod Cluster Control Assembly ..................................................................... 4.2-41 10 Absorber Rod Detail ..................................................................................... 4.2-42 11 Gray Rod Cluster Assembly ......................................................................... 4.2-43 12 Discrete Burnable Absorber Assembly......................................................... 4.2-44 13 Burnable Absorber Rod Assembly (Pyrex) Borosilicate Glass..................... 4.2-45 14 Primary Source Assembly ............................................................................ 4.2-46 15 Secondary Source Assembly ....................................................................... 4.2-47 1 Fuel Loading Arrangement........................................................................... 4.3-50 2 Typical Production and Consumption of Higher Isotopes ............................ 4.3-51 3 Cycle 1 Soluble Boron Concentration Versus Burnup.................................. 4.3-52 4a Cycle 1 Assembly Burnable Absorber Patterns ........................................... 4.3-53 4b Cycle 1 Assembly Burnable Absorber Patterns (Sheet 1 of 2)..................... 4.3-54 4b Cycle 1 Assembly Burnable Absorber Patterns (Sheet 2 of 2)..................... 4.3-55 5 Burnable Absorber, Primary, and Secondary Source Assembly Locations ...................................................................................................... 4.3-56 6 Normalized Power Density Distribution Near Beginning of Life, Unrodded Core, Hot Full Power, No Xenon .................................................................. 4.3-57 7 Normalized Power Density Distribution Near Beginning of Life, Unrodded Core, Hot Full Power, Equilibrium Xenon ..................................................... 4.3-58 8 Normalized Power Density Distribution Near Beginning of Life, Gray Bank MA+MB Inserted, Hot Full Power, Equilibrium Xenon ........................ 4.3-59 9 Normalized Power Density Distribution Near Middle of Life, Unrodded Core, Hot Full Power, Equilibrium Xenon ..................................................... 4.3-60 10 Normalized Power Density Distribution Near End of Life, Unrodded Core, Hot Full Power, Equilibrium Xenon ............................................................... 4.3-61 11 Normalized Power Density Distribution Near End of Life, Gray Bank MA+MB Inserted, Hot Full Power, Equilibrium Xenon ....................... 4.3-62 12 Rodwise Power Distribution in a Typical Assembly (G-9) Near Beginning of Life Hot Full Power, Equilibrium Xenon, Unrodded Core ......................... 4.3-63 13 Rodwise Power Distribution in a Typical Assembly (G-9) Near End of Life Hot Full Power, Equilibrium Xenon, Unrodded Core .................................... 4.3-64 14 Maximum FQ x Power Versus Axial Height During Normal Operation ........ 4.3-65 15 Typical Comparison Between Calculated and Measured Relative Fuel Assembly Power Distribution ........................................................................ 4.3-66 16 Typical Calculated Versus Measured Axial Power Distribution.................... 4.3-67 17 Measured FQ Values Versus Axial Offset for Full Power Rod Configurations .............................................................................................. 4.3-68 18 Typical Doppler Temperature Coefficient at BOL and EOL ......................... 4.3-69 19 Typical Doppler-Only Power Coefficient at BOL and EOL ........................... 4.3-70 4-v Revision 1

21 Typical Moderator Temperature Coefficient at BOL, Unrodded ................... 4.3-72 22 Typical Moderator Temperature Coefficient at EOL..................................... 4.3-73 23 Typical Moderator Temperature Coefficient as a Function of Boron Concentration at BOL, Unrodded ................................................................. 4.3-74 24 Typical Hot Full Power Temperature Coefficient Versus Cycle Burnup ....... 4.3-75 25 Typical Total Power Coefficient at BOL and EOL ........................................ 4.3-76 26 Typical Total Power Defect at BOL and EOL ............................................... 4.3-77 27 Rod Cluster Control Assembly Pattern......................................................... 4.3-78 28 Typical Accidental Simultaneous Withdrawal of Two Control Banks at EOL, HZP, Moving in the Same Plane ......................................................... 4.3-79 29 Typical Design Trip Curve ............................................................................ 4.3-80 30 Typical Normalized Rod Worth Versus Percent Insertion All Rods Inserting Less Most Reactive Stuck Rod .................................................................... 4.3-81 31 X-Y Xenon Test Thermocouple Response Quadrant Tilt Difference Versus Time ............................................................................................................. 4.3-82 32 Calculated and Measured Doppler Defect and Coefficients at BOL, 2-Loop Plant, 121 Assemblies, 12-foot Core ............................................................ 4.3-83 1 Thermal Diffusion Coefficient (TDC) as a Function of Reynolds Number .... 4.4-37 2 Thermal Conductivity of Uranium Dioxide (Data Corrected to 95%

Theoretical Density) ..................................................................................... 4.4-38 4-vi Revision 1

chapter describes the mechanical components of the reactor and reactor core, including the fuel and fuel assemblies, the nuclear design, and the thermal-hydraulic design.

reactor contains a matrix of fuel rods assembled into mechanically identical fuel assemblies g with control and structural elements. The assemblies, containing various fuel enrichments, are figured into the core arrangement located and supported by the reactor internals. The reactor rnals also direct the flow of the coolant past the fuel rods. The coolant and moderator are light er at a normal operating pressure of 2250 psia. The fuel, internals, and coolant are contained in a heavy walled reactor pressure vessel. An AP1000 fuel assembly consists of 264 fuel rods in x17 square array. The center position in the fuel assembly has a guide thimble that is reserved n-core instrumentation. The remaining 24 positions in the fuel assembly have guide thimbles. The e thimbles are joined to the top and bottom nozzles of the fuel assembly and provide the porting structure for the fuel grids.

fuel grids consist of an egg-crate arrangement of interlocked straps that maintain lateral spacing ween the rods. The grid straps have spring fingers and dimples that grip and support the fuel rods.

intermediate mixing vane grids also have coolant mixing vanes. In addition, there are four rmediate flow mixing (IFM) grids. The IFM grid straps contain support dimples and coolant mixing es only. The top and bottom grids and protective grid do not contain mixing vanes.

AP1000 fuel assemblies are similar to the 17x17 Robust and 17x17 XL Robust fuel assemblies.

17x17 Robust fuel assemblies have an active fuel length of 12 feet and three intermediate flow ng grids in the top mixing vane grid spans. The 17x17 XL Robust fuel assemblies have an active length of 14 feet with no intermediate flow mixing grids. The AP1000 fuel assemblies are the e as the 17x17 XL Robust fuel assemblies except that they have four intermediate flow mixing s in the top mixing vane grid spans.

re is substantial operating experience with the 17x17 Robust and 17x17 XL Robust fuel emblies. The 17x17 Robust fuel assemblies are described in References 1, 2 and 3. The 17 XL Robust fuel assemblies are described in References 4 and 5.

XL Robust fuel assembly evolved from the previous VANTAGE+, VANTAGE 5 and VANTAGE 5 BRID designs. The XL Robust fuel assembly is based on the substantial design and operating erience with those designs. The design is described and evaluated in References 2, 3, 6 ugh 10.

umber of proven design features have been incorporated in the AP1000 fuel assembly design.

AP1000 fuel assembly design includes: low pressure drop intermediate grids, four intermediate mixing (IFM) grids, a reconstitutable Westinghouse integral nozzle (WIN), and extended burnup ability. The bottom nozzle is a debris filter bottom nozzle (DFBN) that minimizes the potential for damage due to debris in the reactor coolant. The AP1000 fuel assembly design also includes a ective grid for enhanced debris resistance.

fuel rods consist of enriched uranium, in the form of cylindrical pellets of uranium dioxide, tained in ZIRLOTM (Reference 8) tubing. The tubing is plugged and seal welded at the ends to apsulate the fuel. An axial blanket comprised of fuel pellets with reduced enrichment may be ed at each end of the enriched fuel pellet stack to reduce the neutron leakage and to improve fuel zation.

4.1-1 Revision 1

de-coated fuel pellets and gadolinium oxide/uranium oxide fuel pellets provide a burnable orber integral to the fuel.

l rods are pressurized internally with helium during fabrication to reduce clad creepdown during ration and thereby prevent clad flattening. The fuel rods in the AP1000 fuel assemblies contain itional gas space below the fuel pellets, compared to the 17x17 Robust, 17x17 XL Robust and r previous fuel assembly designs to allow for increased fission gas production due to high fuel nups.

ending on the position of the assembly in the core, the guide thimbles are used for rod cluster trol assemblies (RCCAs), gray rod cluster assemblies (GRCAs), neutron source assemblies, non-gral discrete burnable absorber (BA) assemblies, or thimble plugs.

the initial core design, discrete burnable absorbers (BAs) and integral fuel burnable absorbers used. Discrete burnable absorber designs, integral fuel burnable absorber designs (including IFBAs and gadolinium oxide/uranium oxide BAs) or combinations may be used in subsequent ads.

bottom nozzle is a box-like structure that serves as the lower structural element of the fuel embly and directs the coolant flow distribution to the assembly. The size of flow passages through bottom nozzle limits the size of debris that can enter the fuel assembly. The top nozzle assembly es as the upper structural element of the fuel assembly and provides a partial protective housing he rod cluster control assembly or other components.

rod cluster control assemblies consist of 24 absorber rods fastened at the top end to a common

, or spider assembly. Each absorber rod consists of an alloy of silver-indium-cadmium, which is in stainless steel. The rod cluster control assemblies are used to control relatively rapid changes activity and to control the axial power distribution.

gray rod cluster assemblies consist of 24 rodlets fastened at the top end to a common hub or er. Geometrically, the gray rod cluster assembly is the same as a rod cluster control assembly ept that 12 of the 24 rodlets are fabricated of stainless steel, while the remaining 12 are silver-um-cadmium (of a reduced diameter as compared to the RCCA absorber) with stainless steel gray rod cluster assemblies are used in load follow maneuvering. The assemblies provide a hanical shim reactivity mechanism to minimize the need for changes to the concentration of ble boron.

reactor core is cooled and moderated by light water at a pressure of 2250 psia. Soluble boron in moderator/coolant serves as a neutron absorber. The concentration of boron is varied to control tivity changes that occur relatively slowly, including the effects of fuel burnup. Burnable orbers are also employed in the initial cycle to limit the amount of soluble boron required and, eby maintain the desired negative reactivity coefficients.

nuclear design analyses establish the core locations for control rods and burnable absorbers.

analyses define design parameters, such as fuel enrichments and boron concentration in the lant.

4.1-2 Revision 1

core has inherent stability against diametral and azimuthal power oscillations. Axial power llations, which may be induced by load changes, and resultant transient xenon may be pressed by the use of the rod cluster control assemblies.

control rod drive mechanisms used to withdraw and insert the rod cluster control assemblies and gray rod cluster assemblies are described in Subsection 3.9.4.

thermal-hydraulic design analyses establish that adequate heat transfer is provided between the clad and the reactor coolant. The thermal design takes into account local variations in ensions, power generation, flow distribution, and mixing. The mixing vanes incorporated in the assembly spacer grid design and the fuel assembly intermediate flow mixers induce additional mixing between the various flow channels within a fuel assembly, as well as between adjacent emblies.

reactor internals direct the flow of coolant to and from the fuel assemblies and are described in section 3.9.5.

performance of the core is monitored by fixed neutron detectors outside the core, fixed neutron ctors within the core, and thermocouples at the outlet of selected fuel assemblies. The ex-core lear instrumentation provides input to automatic control functions.

le 4.1-1 presents a summary of the principal nuclear, thermal-hydraulic, and mechanical design ameters of the AP1000 fuel. A comparison is provided to the fuel design used in AP1000, AP600 in a licensed Westinghouse-designed plant using XL Robust fuel. For the comparison with a t containing XL Robust fuel, a 193 fuel assembly plant is used, since no domestic, stinghouse-designed 157 fuel assembly plants use 17x17 XL Robust fuel.

le 4.1-2 tabulates the analytical techniques employed in the core design. The design basis must met using these analytical techniques. Enhancements may be made to these techniques provided the changes are bounded by NRC-approved methods, models, or criteria. In addition, application e process described in WCAP-12488-A, (Reference 9) allows the Combined License holder to e fuel mechanical changes. Table 4.1-3 tabulates the mechanical loading conditions considered he core internals and components. Specific or limiting loads considered for design purposes of various components are listed as follows: fuel assemblies in Subsection 4.2.1.5; control rods CAs and GRCAs), burnable absorber rods, and neutron source rods, in Subsection 4.2.1.6. The amic analyses, input forcing functions, and response loadings for the control rod drive system reactor vessel internals are presented in Subsections 3.9.4 and 3.9.5.

1 Principal Design Requirements fuel and rod control rod mechanism are designed so the performance and safety criteria cribed in Chapter 4 and Chapter 15 are met. [The mechanical design and physical arrangement e reactor components, together with the corrective actions of the reactor control, protection, and rgency cooling systems (when applicable) are designed to achieve these criteria, referred to as cipal Design Requirements:

z Fuel damage, defined as penetration of the fuel cladding, is predicted not to occur during normal operation and anticipated operational transients.

Staff approval is required prior to implementing a change in this information.

4.1-3 Revision 1

z For normal operation and anticipated transient conditions, the minimum DNBR calculated using the WRB-2M correlation is greater than or equal to 1.14.

z Fuel melting will not occur at the overpower limit for Condition I or II events.

z The maximum fuel rod cladding temperature following a loss-of-coolant accident is calculated to be less than 2200°F.

z For normal operation and anticipated transient conditions, the calculated core average linear power, including densification effects, is less than or equal to 5.718 kw/ft for the initial fuel cycle.

z For normal operation and anticipated transient conditions, the calculated total heat flux hot channel factor, FQ, is less than or equal to 2.60 for the initial fuel cycle.

z Calculated rod worths provide sufficient reactivity to account for the power defect from full power to zero power and provide the required shutdown margin, with allowance for the worst stuck rod.

z Calculations of the accidental withdrawal of two control banks using the maximum reactivity change rate predict that the peak linear heat rate and DNBR limits are met.

z The maximum rod control cluster assembly and gray rod speed (or travel rate) is 45 inches per minute.

z The control rod drive mechanisms are hydrotested after manufacture at a minimum of 125 percent of system design pressure.

z For the initial fuel cycle, the fuel rod temperature coefficient is calculated to be negative for power operating conditions.

z For the initial fuel cycle, the moderator temperature coefficient is calculated to be negative for power operating conditions.]*

2 Combined License Information section contained no requirement for additional information.

3 References Letter from N. J. Liparulo (Westinghouse) to J. E. Lyons (NRC), Transmittal of Response to NRC Request for Information on Wolf Creek Fuel Design Modifications, NSD-NRC-97-5189, June 30, 1997.

Letter from N. J. Liparulo (Westinghouse) to R. C. Jones (NRC), Transmittal of Presentation Material for NRC/Westinghouse Fuel Design Change Meeting on April 15, 1996, NSD-NRC-96-4964, April 22, 1996.

Letter from Westinghouse to NRC, Fuel Criteria Evaluation Process Notification for the 17x17 Robust Fuel Assembly with IFM Grid Design, NSD-NRC-98-5796, October 13, 1998.

Staff approval is required prior to implementing a change in this information.

4.1-4 Revision 1

Letter from H. A. Sepp (Westinghouse) to T. E. Collins (NRC), Fuel Criteria Evaluation Process Notification for the Revised Guide Thimble Dashpot Design for the 17x17 XL Robust Fuel Assembly Design, NSD-NRC-98-5722, June 23, 1998.

Davidson, S. L., and Kramer, W. R., (Ed.), Reference Core Report Vantage 5 Fuel Assembly, WCAP-10444-P-A (Proprietary), September 1985 and WCAP-10445-A (Non-Proprietary), December 1983.

Davidson, S. L., (Ed.), VANTAGE 5H Fuel Assembly, Addendum 2-A, WCAP-10444-P-A (Proprietary) and WCAP-10445-NP-A (Non-Proprietary), February 1989.

Davidson, S. L., and Nuhfer, D. L., (Ed.), VANTAGE+ Fuel Assembly Reference Core Report, WCAP-12610-P-A (Proprietary) and WCAP-14342-A (Non-Proprietary),

April 1995.

[Davidson, S. L. (Ed.), Fuel Criteria Evaluation Process, WCAP-12488-A (Proprietary) and WCAP-14204-A (Non-Proprietary), October 1994.]*

Letter, Peralta, J. D. (NRC) to Maurer, B. F. (Westinghouse), Approval for Increase in Licensing Burnup Limit to 62,000 MWD/MTU (TAC No. MD1486), May 25, 2006.

Staff approval is required prior to implementing a change in this information.

4.1-5 Revision 1

Typical hermal and Hydraulic Design Parameters AP1000 AP600 XL Plant ctor core heat output (MWt) 3400 1933 3800 ctor core heat output (106 Btu/hr) 11,601 6596 12,969 t generated in fuel (%) 97.4 97.4 97.4 em pressure, nominal (psia) 2250 2250 2250 em pressure, minimum steady-state (psia) 2190 2200 2204 mum departure from nuclear boiling (DNBR) for gn transients ypical flow channel >1.25(d), >1.22(d) >1.23 >1.26 himble (cold wall) flow channel >1.25(d), >1.21(d) >1.22 >1.24 arture from nucleate boiling (DNB) correlation(b) WRB-2M(b) WRB-2 WRB-1(a) lant Flow(c) l vessel thermal design flow rate (106 lbm/hr) 113.5 72.9 145.0 ctive flow rate for heat transfer (106 lbm/hr) 106.8 66.3 132.7 ctive flow area for heat transfer (ft2) 41.8 38.5 51.1 rage velocity along fuel rods (ft/s) 15.8 10.6 16.6 rage mass velocity (106 lbm/hr-ft2) 2.55 1.72 2.60 lant Temperature(c)(e) inal inlet (°F) 535.0 532.8 561.2 rage rise in vessel (°F) 77.2 69.6 63.6 rage rise in core (°F) 81.4 75.8 68.7 rage in core (°F) 578.1 572.6 597.8 rage in vessel (°F) 573.6 567.6 593.0 t Transfer ve heat transfer surface area (ft2) 56,700 44,884 69,700 heat flux (BTU/hr-ft2) 199,300 143,000 181,200 imum heat flux for normal operation 518,200 372,226 489,200 BTU/hr-ft2)(f) rage linear power (kW/ft)(g) 5.72 4.11 5.20 k linear power for normal operation (kW/ft)(f)(g) 14.9 10.7 14.0 k linear power (kW/ft)(f)(h) 22.45 22.5 22.45 sulting from overpower transients/operator rs, assuming a maximum overpower of 118%)

4.1-6 Revision 1

Typical hermal and Hydraulic Design Parameters AP1000 AP600 XL Plant t flux hot channel factor (FQ) 2.60 2.60 2.70 k fuel center line temperature (°F) 4700 4700 4700 prevention of center-line melt) assembly design 17x17 XL 17x17 17x17 XL Robust Fuel Robust Fuel/ No IFM ber of fuel assemblies 157 145 193 nium dioxide rods per assembly 264 264 264 pitch (in.) 0.496 0.496 0.496 rall dimensions (in.) 8.426 x 8.426 8.426 x 8.426 8.426 x 8.426 weight, as uranium dioxide (lb) 211,588 167,360 261,000 weight (lb) 43,105 35,555 63,200 ber of grids per assembly op and bottom - (Ni-Cr-Fe Alloy 718) 2(i) 2(i) 2 termediate 8 ZIRLO' 7 Zircaloy-4 or 8 ZIRLO' 7 ZIRLO' termediate flow mixing 4 ZIRLO' 4 Zircaloy-4 or 0 5 ZIRLO' rotective Grid - (Ni-Cr-Fe Alloy 718) 1 1 1 ding technique, first cycle 3 region 3 region 3 region nonuniform nonuniform nonuniform l Rods ber 41,448 38,280 50,952 side diameter (in.) 0.374 0.374 0.374 metral gap (non-IFBA) (in.) 0.0065 0.0065 0.0065 thickness (in.) 0.0225 0.0225 0.0225 material ZIRLO' Zircaloy-4 or Zircaloy-4/ ZIRLO' ZIRLO' l Pellets erial UO2 sintered UO2 sintered UO2 sintered sity (% of theoretical) 95.5 95 95 meter (in.) 0.3225 0.3225 0.3225 gth (in.) 0.387 0.387 0.387 4.1-7 Revision 1

Typical Rod Cluster Control Assemblies AP1000 AP600 XL Plant tron Absorber CCA 24 Ag-In-Cd rodlets 24 Ag-In-Cd 24 Hafnium or RCA 12 304 SS rodlets rodlets Ag-In-Cd 12 Ag-In-Cd rodlets 20 304 SS rodlets 4 Ag-In-Cd rodlets ladding material Type 304 Type 304 SS, Type 304 SS, SS, cold-worked cold-worked cold-worked lad thickness, (Ag-In-Cd) 0.0185 0.0185 0.0185 umber of clusters 53 RCCAs 45 RCCAs 57 RCCAs 16 GRCAs 16 GRCAs 0 GRCAs e Structure ore barrel, ID/OD (in.) 133.75/137.75 133.75/137.75 148.0/152.5 hermal shield Neutron Panel None Neutron Panel affle thickness (in.) Core Shroud Radial reflector 0.875 cture Characteristics ore diameter, equivalent (in.) 119.7 115.0 132.7 ore height, cold, active fuel (in.) 168.0 144.0 168.0 l Enrichment First Cycle (Weight Percent) egion 1 2.35 1.90 Typical egion 2 3.40 2.80 3.8 to 4.4 egion 3 4.45 3.70 (5.0 Max) s:

WRB-2M will be used in future reloads.

See Subsection 4.4.2.2.1 for the use of the W-3, WRB-2 and WRB-2M correlations.

Flow rates and temperatures are based on 10 percent steam generator tube plugging for the AP600 and AP1000 designs.

1.25 applies to core and axial offset limits; 1.22 and 1.21 apply to all other RTDP transients.

Coolant temperatures based on thermal design flow (for AP600 and AP1000).

Based on FQ of 2.60 for AP600 and AP1000.

Based on densified active fuel length. The value for AP1000 is rounded to 5.72 kW/ft.

See Subsection 4.3.2.2.6.

The top grid will be fabricated of nickel-chromium-iron Alloy 718.

4.1-8 Revision 1

Subsection Analysis Technique Computer Code Referenced chanical design of core Static and dynamic modeling BLOWDOWN code, 3.7.2.1 ernals loads, deflections, FORCE, finite element 3.9.2 d stress analysis structural analysis 3.9.3 code, and others el rod design Semi-empirical thermal model of Westinghouse fuel rod 4.2.1.1 Fuel performance fuel rod with considerations such design model 4.2.3.2 characteristics (such as, as fuel density changes, heat 4.2.3.3 temperature, internal transfer, and fission gas release. 4.3.3.1 pressure, and clad stress) 4.4.2.11 clear design Microscopic data; macroscopic Modified ENDF/B 4.3.3.2 Cross-sections and group constants for homogenized core library with constants regions PHOENIX-P X-Y and X-Y-Z power 2-group diffusion theory, 2-group ANC (2-D or 3-D) 4.3.3.3 distributions, fuel nodal theory depletion, critical boron concentrations, X-Y and X-Y-Z xenon distributions, reactivity coefficients Axial power distributions, 1-D, 2-group diffusion theory APOLLO 4.3.3.3 control rod worths, and axial xenon distribution Fuel rod power Integral transport theory LASER 4.3.3.1 Effective resonance Monte Carlo weighing function REPAD 4.3.3.1 temperature Criticality of reactor and 3-D, Monte Carlo theory AMPX system of 4.3.2.6 fuel assemblies codes, KENO-Va ssel irradiation Multigroup spatial dependent DOT 4.3.2.8 transport theory ermal-hydraulic design Subchannel analysis of local fluid VIPRE-01 4.4.4.5.2 ady state conditions in rod bundles, including inertial and cross-flow resistance terms; solution progresses from core-wide to hot assembly to hot channel.

nsient departure from Subchannel analysis of local fluid VIPRE-01 4.4.4.5.4 cleate boiling conditions in rod bundles during transients by including accumulation terms in conservation equations; solution progresses from core-wide to hot assembly to hot channel.

4.1-9 Revision 1

Fuel assembly weight and core component weights (burnable absorbers, sources, RCCA, GRCA)

Fuel assembly spring forces and core component spring forces Internals weight Control rod trip (equivalent static load)

Differential pressure Spring preloads Coolant flow forces (static)

Temperature gradients Thermal expansion Interference between components Vibration (mechanically or hydraulically induced)

Operational transients listed in Table 3.9-1 Pump overspeed Seismic loads (safe shutdown earthquake)

Blowdown forces (due to pipe rupture) 4.1-10 Revision 1

Condition I - normal operation and operational transients Condition II - events of moderate frequency Condition III - infrequent incidents Condition IV - limiting faults pter 15 describes bases and plant operation and events involving each condition.

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

The mechanical design and physical arrangement of the reactor core components, together with corrective actions of the reactor control, protection, and emergency cooling systems (when applicable) provide that:

- Fuel damage, that is, breach of fuel rod clad pressure boundary, is not expected during Condition I and Condition II events. A very small amount of fuel damage may occur. This is within the capability of the plant cleanup system and is consistent with the plant design bases.

- The reactor can be brought to a safe state following a Condition III event with only a small fraction of fuel rods damaged. The fraction of fuel rods damaged must be limited to meet the dose guidelines identified in Chapter 15 although sufficient fuel damage might occur to preclude immediate resumption of operation.

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

The fuel assemblies are designed to withstand non-operational loads induced during shipping, handling, and core loading without exceeding the criteria of Subsection 4.2.1.5.1.

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

The fuel assemblies have provisions for the insertion of in-core instrumentation.

The reactor vessel and internals, in conjunction with the fuel assembly structure, directs reactor coolant through the core. Because of the resulting flow distribution and bypass flow, the heat transfer performance requirements are met for the modes of operation.

following subsection provides the fuel system design bases and design limits. It is consistent the criteria of the Standard Review Plan, Section 4.2.

sistent with the growth in technology, Westinghouse modifies fuel system designs. These ifications utilize NRC approved methods. [A set of design fuel criteria to be satisfied by new fuel igns was issued to the NRC in WCAP-12488-A and its addendums (Reference 1)]* and also ented below in Subsection 4.2.1.

Staff approval is required prior to implementing a change in this information.

4.2-1 Revision 1

ty criteria presented in Section 4.2 of the Standard Review Plan. [The design bases and eptance limits used by Westinghouse are also described in the Westinghouse Fuel Criteria luation Process, WCAP-12488-A and its addendums (Reference 1).]*

fuel rods are designed to satisfy the fuel rod design criteria for rod burnup levels up to the design harge burnup using the extended burnup design methods described in the Extended Burnup luation report, WCAP-10125-P-A (Reference 2).

AP1000 fuel rod design considers effects such as fuel density changes, fission gas release, clad p, and other physical properties which vary with burnup. The integrity of the fuel rods is provided esigning to prevent excessive fuel temperatures as discussed in Subsection 4.2.1.2.1; excessive rnal rod gas pressures due to fission gas releases as discussed in Subsections 4.2.1.3.1 4.2.1.3.2; and excessive cladding stresses, strains, and strain fatigue, as discussed in sections 4.2.1.1.2 and 4.2.1.1.3. The fuel rods are designed so that the conservative design es of the following events envelope the lifetime operating conditions of the fuel. For each design is, the performance of the limiting fuel rod, with appropriate consideration for uncertainties, does exceed the limits specified by the design basis. The detailed fuel rod design also establishes such ameters as pellet size and density, clad/pellet diametral gap, gas plenum size, and helium pre-surization level.

grity of the fuel assembly structure is provided by setting limits on stresses and deformations due arious loads and by preventing the assembly structure from interfering with the functioning of r components. Three types of loads are considered:

Non-operational loads, such as those due to shipping and handling Normal and abnormal loads, which are defined for Conditions I and II Abnormal loads, which are defined for Conditions III and IV design bases for the in-core control components are described in Subsection 4.2.1.6.

1.1 Cladding 1.1.1 Mechanical Properties ZIRLO' cladding material combines neutron economy (low absorption cross-section); high osion resistance to coolant, fuel, and fission products; and high strength and ductility at operating peratures. ZIRLO' is an advanced zirconium based alloy that has the same or similar properties advantages as Zircaloy-4 and was developed to support extended fuel burnup.

AP-12610-P-A (Reference 5) provides a discussion of chemical and mechanical properties of the LO' cladding material and a comparison to Zircaloy-4.

1.1.2 Stress-Strain Limits d Stress e volume average effective stress calculated with the Von Mises equation (considering rference due to uniform cylindrical pellet-clad contact, caused by pellet thermal expansion, pellet lling and uniform clad creep, and pressure differences) is less than the 0.2 percent offset yield ss with due consideration to temperature and irradiation effects for Condition I and II events, AP-12488-A (Reference 1).]* While the clad has some capability for accommodating plastic Staff approval is required prior to implementing a change in this information.

4.2-2 Revision 1

d Strain e total plastic tensile creep strain due to uniform clad creep, and uniform cylindrical fuel pellet ansion associated with fuel swelling and thermal expansion is less than one percent from the radiated condition, WCAP-12488-A (Reference 1).]* The acceptance limit for fuel rod clad strain ng Condition II events is that the total tensile strain due to uniform cylindrical pellet thermal ansion is less than one percent from the pre-transient value. These limits are consistent with en practice.

1.1.3 Fatigue and Vibration gue e usage factor due to cycle fatigue is less than 1.0, WCAP-12488-A (Reference 1).]* That is, for a n strain range, the number of strain fatigue cycles are less than those required for failure. The ue curve is based on a safety factor of two on the stress amplitude or a safety factor of 20 on the ber of cycles, whichever is more conservative.

ration ential fretting wear due to vibration is prevented, giving confidence that the stress-strain limits are exceeded during design life. Fretting of the clad surface can occur due to flow-induced vibration ween the fuel rods and fuel assembly grid springs. Vibration and fretting forces may vary during fuel life due to clad diameter creep down combined with grid spring relaxation.

1.1.4 Chemical Properties mical properties of the ZIRLO' cladding are discussed in WCAP-12610 (Reference 5).

1.2 Fuel Material 1.2.1 Thermal-Physical Properties center temperature of the hottest pellet is below the melting temperature of the uranium dioxide.

melting temperature of unirradiated uranium dioxide, 5080°F, decreases by 58°F per 00 megawatt days per metric ton of uranium, as discussed in WCAP-9179 (Reference 4). Fuel ting will not occur at the overpower limit for Condition I or II events. This provides sufficient margin uncertainties as described in Subsection 4.4.2.9.

nominal design density of the fuel is approximately 95.5 percent of the theoretical density.

itional information on fuel properties is provided in WCAP-9179 (Reference 4).

1.2.2 Fuel Densification and Fission Product Swelling design bases and models used for fuel densification and swelling are provided in AP-10851-P-A (Reference 7), and WCAP-15063-P-A, Revision 1 (Reference 21).

1.2.3 Chemical Properties AP-9179 (Reference 4) and WCAP-12610 (Reference 5) provide the basis for justifying that no erse chemical interactions occur between the fuel and its adjacent material.

Staff approval is required prior to implementing a change in this information.

4.2-3 Revision 1

basic fuel rod models and the ability to predict fuel rod operating characteristics are given in AP-15063-P-A, Revision 1 (Reference 21) and Subsection 4.2.3.

1.3.2 Mechanical Design Limits dding collapse is precluded during the fuel rod design lifetime. Current generation Westinghouse is sufficiently stable with respect to fuel densification. Significant axial gaps in the pellet stack essary for clad flattening do not occur and therefore, clad flattening will not occur. Clad flattening hodologies are described in WCAP-13589-A, (Reference 8) and WCAP-8377 (Reference 22).

rod internal gas pressure remains below the value which causes the fuel/clad diametral gap to ease due to outward cladding creep during steady-state operation. Rod pressure is also limited h that extensive departure from nucleate boiling propagation does not occur as discussed in AP-8963-P-A (Reference 9).

1.4 Spacer Grids 1.4.1 Mechanical Limits and Materials Properties grid component strength criteria are based on experimental tests. The limit is established at the percent confidence level on the true mean crush strength at operating temperature. This limit is cient to provide that, under worst-case combined seismic and pipe rupture event, the core will ntain a geometry amenable to cooling. As an integral part of the fuel assembly structure, the grids sfy the applicable fuel assembly design bases and limits defined in Subsection 4.2.1.5.

grid material and chemical properties are given in WCAP-9179 (Reference 4) and WCAP-12610 ference 5).

1.4.2 Vibration and Fatigue grids provide sufficient fuel rod support to limit fuel rod vibration and maintain clad fretting wear in acceptable limits (defined in Subsection 4.2.1.1).

1.5 Fuel Assembly Structural Design iscussed in Subsection 4.2.1, the structural integrity of the fuel assemblies is provided by setting ign limits on stresses and deformations due to various non-operational, operational, and accident

s. These limits are applied to the design and evaluation of the top and bottom nozzles, guide bles, grids, and thimble joints. [Design changes to the fuel assembly structure qualify for luation in WCAP-12488-A (Reference 1).]*

design bases for evaluating the structural integrity of the fuel assemblies are discussed in sections 4.2.1.5.1 through 4.2.1.5.3.

1.5.1 Non-Operational non-operational load is a loading of 4 g axial (longitudinal) and 6 g lateral (transverse) with ensional stability.

Staff approval is required prior to implementing a change in this information.

4.2-4 Revision 1

the normal operation (Condition I) and upset (Condition II) conditions, the fuel assembly ponent structural design criteria are established for the two primary material categories, tenitic steels and zirconium alloys. The stress categories and strength theory presented in the ME Code,Section III, are used as a general guide. The maximum shear theory (Tresca criterion) ombined stresses is used to determine the stress intensities for the austenitic steel components.

stress intensity is defined as the largest numerical difference between the various principal sses in a three-dimensional field. The design stress intensity value, Sm, for austenitic steels and onium alloys is given by the lowest of the following:

One-third of the specified minimum tensile strength or two-thirds of the specified minimum yield strength at room temperature One-third of the tensile strength or 90 percent of the yield strength at temperature, but not to exceed two-thirds of the specified minimum yield strength at room temperature stress limits for the austenitic steel components are given below. Stress nomenclature follows ASME Code,Section III.

Stress Intensity Limits tegories Limit neral primary membrane Sm ss intensity al primary membrane 1.5 Sm ss intensity mary membrane plus 1.5 Sm ding stress intensity al primary plus 3.0 Sm ondary stress intensity zirconium alloy structural components, which consist of guide thimbles and fuel tubes, are in turn divided into two categories because of material difference and functional requirements. The fuel design criteria are covered separately in Subsection 4.2.1.1. The maximum shear theory is used valuate the guide thimble design. For conservative purposes, the zirconium alloy unirradiated perties are used to define the stress limits.

1.5.3 Infrequent Incidents (Condition III) and Limiting Faults (Condition IV) ical worse case abnormal loads during Conditions III and IV are represented by seismic and pipe ure loadings. The design criteria for this category of loadings are as follows:

Deflections or excessive deformation of components cannot interfere with capability of insertion of the control rods or emergency cooling of the fuel rods.

The fuel assembly structural components stresses under faulted conditions are evaluated primarily using the methods outlined in Appendix F of the ASME Code,Section III. Since the current analytical methods use linear 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 4.2-5 Revision 1

primary membrane and 2.4 Sy or 1.05 Su for primary membrane plus bending. For conservative purposes, the zirconium alloy unirradiated properties are used to define the stress limits.

material and chemical properties of the fuel assembly components are given in WCAP-9179 ference 4) and WCAP-12610 (Reference 5). Subsection 4.2.3.4 discusses the spacer grid crush ing.

rmal-hydraulic design is discussed in Section 4.4.

1.6 In-core Control Components in-core control components are subdivided into permanent and temporary devices. The manent components are the rod cluster control assemblies, gray rod cluster assemblies, and ondary neutron source assemblies. The temporary components are the primary neutron source emblies (which are normally used only in the initial core), the burnable absorber assemblies, and thimble plugs. For some reloads, the use of burnable absorbers may be necessary for power ribution control and/or to achieve an acceptable moderator temperature coefficient throughout life (See Subsection 4.3.1.2.2). [Design changes to the in-core control components qualify for luation using the criteria defined in WCAP-12488-A (Reference 1).]*

erials are selected for:

Compatibility in a pressurized water reactor environment Adequate mechanical properties at room and operating temperatures Resistance to adverse property changes in a radioactive environment Compatibility with interfacing components erial properties are given in WCAP-9179 (Reference 4).

design bases for the in-core control components are given in Subsections 4.2.1.6.1 through 1.6.3.

1.6.1 Control Rods Conditions I and II, the stress categories and strength theory presented in the ASME Code, tion III, are used as a general guide in the design of the RCCA and GRCA structural parts in ition to absorber cladding.

ign conditions considered under the ASME Code,Section III, are as follows:

External pressure equal to the reactor coolant system operating pressure with appropriate allowance for overpressure transients Wear allowance equivalent to 1000 reactor trips Bending of the rod due to a misalignment in the guide thimble Staff approval is required prior to implementing a change in this information.

4.2-6 Revision 1

Radiation exposure during maximum core life. The absorber material temperature does not exceed its melting temperature (1454°F for silver-indium-cadmium [Ag-In-Cd]), (see WCAP-9179, Reference 4).

Temperature effects at operating conditions 1.6.2 Burnable Absorber Rods Conditions I and II, the stress categories and strength theory presented in the ASME Code, tion III, are used as a general guide in the design of the burnable absorber cladding. For ormal loads during Conditions III and IV, code stresses are not considered limiting. Failures of the nable absorber rods during these conditions must not interfere with reactor shutdown or rgency cooling of the fuel rods. The burnable absorber material is nonstructural. The structural ments of the burnable absorber rod are designed to maintain the absorber geometry even if the orber material is fractured.

educe the dissolved boron requirement for control of excess reactivity, burnable absorber rods e been incorporated in the core design. In the first core, the burnable absorber rods (Pyrex) sist of borosilicate glass tubes contained within Type 304 stainless steel tubular cladding, which is ged and seal welded at the ends to encapsulate the glass. The absorber material temperature s not exceed its design limit of 1220°F. Mechanical and thermal design and nuclear evaluation of burnable absorber rods are described in WCAP-7113 (Reference 23).

alternative discrete burnable absorber is the wet annular burnable absorber (WABA). The nable absorber material is boron carbide contained in an alumina matrix. Thermal-physical and release properties of alumina-boron carbide are described in WCAP-9179 (Reference 4) and AP-10021-P-A (Reference 10). Discrete burnable absorber rods are designed so that the orber temperature does not exceed 1200°F during normal operation or an overpower transient.

1200°F maximum temperature helium gas release in a discrete burnable absorber rod will not eed 30 percent of theoretical. See WCAP-10021-P-A (Reference 10).

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

The external pressure equal to reactor coolant system operating pressure with appropriate allowance for overpressure transients An internal pressure equal to the pressure generated by released gases over the source rod life 1.7 Surveillance Program section 4.2.4.6 discusses the testing and fuel surveillance operation experience program that has n and is being conducted to verify the adequacy of the fuel performance and design bases. Fuel eillance and testing results, as they become available, are used to improve fuel rod design and ufacturing processes and to confirm that the design bases and safety criteria are satisfied.

4.2-7 Revision 1

h fuel assembly consists of 264 fuel rods, 24 guide thimbles, and 1 instrumentation tube arranged in a supporting structure. The instrumentation thimble is located in the center position and ides a channel for insertion of an in-core neutron detector, if the fuel assembly is located in an rumented core position. The guide thimbles provide channels for insertion of either a rod cluster trol assembly, a gray rod cluster assembly, a neutron source assembly, a burnable absorber embly, or a thimble plug, depending on the position of the particular fuel assembly in the core.

re 4.2-1 shows a cross-section of the fuel assembly array, and Figure 4.2-2 shows a fuel embly full-length view.

fuel rods are loaded into the fuel assembly structure so that there is clearance between the fuel ends and the top and bottom nozzles. The fuel rods are supported within the fuel assembly cture by fourteen structural grids (top grid (1), bottom grid (1), intermediate grids (8) and rmediate flow mixer (IFM) grids (4)), plus one protective grid. The top grid is fabricated from el-chromium-iron Alloy 718. The bottom grid is fabricated from nickel-chromium-iron Alloy 718.

intermediate grids and the IFM grids are fabricated from ZIRLO' (see WCAP-12610-P-A, erence 5). Top, bottom, and intermediate grids provide axial and lateral support to the fuel rods. In ition, the four IFM grids located near the center of the fuel assembly and between the rmediate grids provide additional fuel rod restraint. The protective grid, in combination with the ris filter bottom nozzle (DFBN), the protective zirconium oxide coated fuel cladding, and the long, d fuel rod bottom end plug, provide debris failure mitigation.

l assemblies are installed vertically in the reactor vessel and stand upright on the lower core e, which is fitted with alignment pins to locate and orient each assembly. After the fuel assemblies set in place, the upper support structure is installed. Alignment pins, built into the upper core e, engage and locate the upper ends of the fuel assemblies. The upper core plate then bears n against the hold-down springs on the top nozzle of each fuel assembly to hold the fuel emblies in place.

roper orientation of fuel assemblies within the core is prevented by the use of an indexing hole in corner of the top nozzle top plate. The assembly is oriented with respect to the handling tool and core by means of a pin inserted into this indexing hole. Visual confirmation of proper orientation is provided by an engraved identification number on the opposite corner clamp.

2.1 Fuel Rods fuel rods consist of uranium dioxide ceramic pellets contained in cold-worked and stress relieved LO' tubing, which is plugged and seal-welded at the ends to encapsulate the fuel. ZIRLO' is an anced zirconium based alloy selected for its mechanical properties and low neutron absorption s-section (see WCAP-12610-P-A, Reference 5). Figure 4.2-3 shows a schematic of the fuel rod.

fuel pellets are right circular cylinders consisting of slightly enriched uranium dioxide powder ch has been compacted by cold pressing and then sintered to the required density. The ends of h pellet are dished slightly, to allow greater axial expansion at the pellet centerline and to increase void volume for fission gas release. The ends of each pellet also have a small chamfer at the r cylindrical surface which improves manufacturability, and mitigates potential pellet damage due el rod handling.

volume and clearances are provided within the rods to accommodate fission gases released the fuel, differential thermal expansion between the clad and the fuel, and fuel density changes ng irradiation. To facilitate the extended burnup capability necessitated by longer operating es, the fuel rod is designed with two plenums (upper and lower) to accommodate the additional 4.2-8 Revision 1

ting of the fuel within the clad during handling or shipping, prior to core loading, is prevented by a nless steel helical spring which bears on top of the fuel pellet stack. Assembly consists of ging and welding the bottom of the cladding, installing the bottom plenum spacer assembly, fuel ets and top plenum spring, and then plugging and welding the top of the rod. The solid bottom plug has an internal grip feature and tapered end to facilitate fuel rod loading during fuel embly fabrication and reconstitution. Additionally, the bottom end plug is designed to be ciently long to extend through the protective grid. The bottom section of the fuel rod has a ective zirconium oxide coated surface feature. Use of the protective grid with a longer end plug the debris filter bottom nozzle, in addition to the coated cladding surface, constitutes a three-l debris protection package, which enhances the fuel reliability performance against trapped ris. This precludes any breach in the fuel rod pressure boundary due to clad fretting wear induced ebris trapped at the bottom section of the fuel assembly.

fuel rods are internally pressurized with helium during the welding process to minimize pressive clad stresses and prevent clad flattening under reactor coolant operating pressures.

fuel rods are pre-pressurized and designed so that:

The internal gas pressure mechanical design limit referred to in Subsection 4.2.1.3 is not exceeded The cladding stress-strain limits (Subsection 4.2.1.1) are not exceeded for Condition I and II events Clad flattening will not occur during the fuel core life AP1000 fuel rod design may also include axial blankets. The axial blankets consist of fuel pellets reduced enrichment at each end of the fuel rod pellet stack. Axial blankets reduce neutron age axially and improve fuel utilization. The axial blankets use chamfered pellets that are longer the enriched pellets to help prevent accidental mixing during manufacturing. Furthermore, axial kets have no impact on the source range detector response, since the reduction in power from axial blanket is limited to the top and bottom 0.67 feet of the core, while the source range ctors are centered typically about three feet from the bottom of the core.

AP1000 fuel rod design may also include annular fuel pellets in the top and bottom 8 inches of fuel stack. These pellets can be either fully enriched or partially enriched. The annular fuel pellets ide additional void volume in the fuel rod to accommodate fission gas release.

AP1000 fuel rods include integral fuel burnable absorbers. The integral fuel burnable absorbers be boride-coated fuel pellets or fuel pellets containing gadolinium oxide mixed with uranium

e. The boride-coated fuel pellets are identical to the enriched uranium dioxide pellets except for addition of a thin boride coating less than 0.001 inch in thickness on the pellet cylindrical surface.

ted pellets occupy the central portion of the fuel column. The number and pattern of integral fuel nable absorber rods within an assembly may vary depending on specific application. See AP-12610-P-A (Reference 5).

2.2 Fuel Assembly Structure hown in Figure 4.2-2, the fuel assembly structure consists of a bottom nozzle, top nozzle, fuel

, guide thimbles, and grids.

4.2-9 Revision 1

lant flow distribution to the assembly. The nozzle is fabricated from Type 304 stainless steel and sists of a perforated plate, and casting which incorporates a skirt and four angle legs with bearing

s. Figure 4.2-2 illustrates this concept. The legs and skirt form a plenum to direct the inlet coolant to the fuel assembly. The perforated plate also prevents accidental downward ejection of the fuel from the fuel assembly. The bottom nozzle is fastened to the fuel assembly guide thimbles by ed thimble screws, which penetrate through the nozzle and engage with a threaded plug in each e thimble.

lant flows from the plenum in the bottom nozzle, upward through the penetrations in the plate, to channels between the fuel rods. The penetrations in the plate are positioned between the rows of fuel rods.

ddition to serving as the bottom structural element of the fuel assembly, the bottom nozzle also tions as a debris filter. The bottom nozzle perforated plate contains a multiplicity of flow holes ch are sized to minimize passage of detrimental debris particles into the active fuel region of the while maintaining sufficient hydraulic and structural margins. Furthermore, the skirt provides roved bottom nozzle structural stability and increased design margins to reduce damage due to ormal handling.

l loads (from top nozzle hold-down springs) imposed on the fuel assembly and the weight of the assembly are transmitted through the bottom nozzle to the lower core plate. Indexing and itioning of the fuel assembly is controlled by alignment holes in two diagonally opposite bearing s that mate with locating pins in the lower core plate. Lateral loads on the fuel assembly are smitted to the lower core plate through the locating pins.

AP1000 bottom nozzle also has a reconstitution design feature which facilitates the easy oval of the nozzle from the fuel assembly. This design incorporates a thimble screw with a circular ing cup located around the screw head. The locking cup is crimped into a local spherical radius f on the bottom nozzle. To remove the bottom nozzle, a counterclockwise torque is applied to the ble screw until the locking cup (detents) is relaxed and the thimble screw is removed. This nstitutable design permits the remote unlocking, the removal, and the relocking of the thimble ws, as the same or a new bottom nozzle is reattached to the fuel assembly.

2.2.2 Top Nozzle reconstitutable top nozzle functions as the upper structural component of the fuel assembly and, ddition, provides a partial protective housing for the rod cluster control assembly, discrete nable absorber, or other core components. As shown in Figure 4.2-2, the top nozzle assembly udes four sets of hold-down springs, which are secured to the top nozzle top plate. The springs made of nickel-chromium-iron Alloy 718. The other top nozzle components are made of Type 304 nless steel.

adapter plate is provided with round penetrations and slots (with semicircular ends) to permit the of coolant upward through the top nozzle. Other round holes are provided in the adapter plate to ept (guide thimble) inserts which are mechanically locked to the adapter plate using a lock tube.

unique design of the insert joint and lock tube are the key design features of the reconstitutable nozzle.

ligaments in the adapter plate cover the top of the fuel rods precluding any upward ejection of fuel rods from the fuel assembly. The enclosure is a box-like structure which establishes the ance between the adapter plate and the top plate. The top plate has a large square hole in the 4.2-10 Revision 1

top plate also contains integral pads located on the two remaining top nozzle corners. The pads ude alignment holes which, when fully engaged with the reactor internals upper core plate guide

, provide proper alignment to the fuel assembly, reactor internals, and rod control cluster embly.

hown in Figure 4.2-4, to remove the top nozzle assembly a tool is first inserted through a lock and expanded radially to engage the bottom edge of the tube. An axial force is then exerted on tool which overrides local lock tube deformations and withdraws the lock tubes from the inserts.

r the lock tubes have been removed, the nozzle assembly is removed by raising it off the upper ed ends of the nozzle inserts, which deflect inwardly under the axial lift load.

h the top nozzle assembly removed, direct access is provided for fuel rod examination or acement. Reconstitution is completed by the remounting of the nozzle assembly and the insertion ck tubes. Details of this design feature, the design bases and evaluation of the reconstitutable nozzle are given in WCAP-10444-P-A (Reference 11).

2.2.3 Guide Thimbles and Instrument Tube guide thimbles are structural members that provide channels for the neutron absorber rods, nable absorber rods, neutron source rods, or other assemblies. Each guide thimble is fabricated Zircaloy-4 or ZIRLO' with constant OD and ID over the entire length. Separate dashpot tubes, ch are made from Zircaloy-4 or ZIRLO' tubing, are inserted into the bottom portion of the guide ble tubes. The larger tube diameter at the top section provides a relatively large annular area essary to permit rapid control rod insertion during a reactor trip, as well as to accommodate the of coolant during normal operation. Holes are provided on the guide thimble above the dashpot educe the rod drop time. The lower portion of the guide thimble with the dashpot tube results in a hpot action near the end of the control rod travel during normal trip operation. The dashpot is ed at the bottom by means of an end plug, which is provided with a small flow port to avoid fluid nation in the dashpot volume during normal operation.

tated previously, the AP1000 fuel assembly includes a reconstitutable top nozzle as a standard ure. To accommodate the reconstitutable feature, the top end of the zirconium alloy guide thimble stened to a tubular sleeve, or insert, by a three tier expansion bulge joint. An expansion tool is rted inside the nozzle insert and guide thimble to the proper elevation. The four lobes on the ansion tool force the guide thimble and insert outward locally to a predetermined diameter, efore joining the two components.

n installation of the top nozzle assembly, the bulge near the top of the nozzle insert is captured in rresponding groove in the hole of the top nozzle adapter plate. As shown in Figure 4.2-4, the hanical connection between the nozzle insert-guide thimble and top nozzle is made by insertion lock tube into the insert. The design of the top grid sleeve-guide thimble and top nozzle insert-e thimble bulge joint connections have been mechanically tested and found to meet applicable ign criteria.

fuel rod support grids, with exception noted for the bottom nickel-chromium-iron Alloy 718 grid, secured to the guide thimbles using a similar bulge joint connection to create an integral cture. Attachment of the intermediate mixing vane and intermediate flow mixer (IFM) zirconium y grids to the guide thimbles is performed using the fastening technique depicted in Figures 4.2-5 4.2-6.

4.2-11 Revision 1

ed and meets the design requirements.

bottom nickel-chromium-iron Alloy 718 grid is secured to the guide thimble assembly by a double bulge connection between the grid sleeve and guide thimble. The design of the double tier bulge connection has also been mechanically tested and meets the design requirements.

lower end of the guide thimble is fitted with a welded end plug. The nickel-chromium-iron y 718 protective grid is secured to the guide thimble assembly by nickel-chromium-iron Alloy 718 cers that are spot-welded to the grid. As shown in Figure 4.2-7, the spacer is captured between guide thimble end plug and the bottom nozzle by means of a (thimble) locking screw.

described methods of grid fastening are standard and have been used successfully since the duction of zirconium alloy guide thimbles in 1969.

central instrumentation tube in each fuel assembly is constrained by seating in counterbores ted in both top and bottom nozzles. The instrumentation tube has a constant diameter and ides an unrestricted passageway for the in-core neutron detector which enters the fuel assembly the top nozzle. Furthermore, the instrumentation tube is secured to the top and mid-grids with e joint connections similar to those previously discussed for securing the grids to the guide bles.

2.2.4 Grid Assemblies hown in Figure 4.2-2, the fuel rods are supported at intervals along their lengths by grid emblies which maintain the lateral spacing between the rods throughout the design life of the embly. Each fuel rod is given support at six contact points within each grid by the combination of port dimples and springs. The grid assembly consists of individual slotted straps assembled and rlocked into an egg-crate type arrangement with the straps permanently joined at their points of rsection. The straps may contain springs, support dimples, and mixing vanes; or any such bination.

types of structural grid assemblies are used on the AP1000 fuel assembly. One type, with mixing es projecting from the edges of the straps into the coolant stream, is used in the high heat flux on of the fuel assemblies to promote mixing of the coolant. The other type, located at the top and om of the assembly, does not contain mixing vanes on the internal straps. The outside straps on grids contain mixing vanes that, in addition to their mixing function, aid in guiding the grids and assemblies past projecting surfaces during handling or during loading and unloading of the core.

ause of its corrosion resistance and high strength properties, the bottom grid material chosen for AP1000 fuel assembly design is nickel-chromium-iron Alloy 718. The top grid is fabricated from el-chromium-iron Alloy 718. The magnitude of the grid restraining force on the fuel rod is set high ugh to minimize possible fretting, without overstressing the cladding at the points of contact ween the grids and fuel rods. The grid assemblies are designed to allow axial thermal expansion e fuel rods without imposing restraint sufficient to develop buckling or distortion of the fuel rods.

eight intermediate (mixing vane), or structural grids on the AP1000 fuel assembly are made of LO'. This material was selected to take advantage of the materials inherent low neutron capture s-section. The zirconium alloy grids have thicker straps than the nickel-chromium-iron alloy grids.

zirconium alloy grid incorporates the same grid cell support configuration as the nickel-mium-iron alloy grid. The zirconium alloy interlocking strap joints and grid/sleeve joints are icated by laser welding, whereas the nickel-chromium-iron alloy grid joints (except the protective 4.2-12 Revision 1

emblies. This additional flow mixing enhances thermal performance.

hown in Figure 4.2-2, the intermediate flow mixer grids are located at selected spans between zirconium alloy mixing vane structural grids and incorporate a similar mixing vane array. Their e function is mid-span flow mixing in the hotter fuel assembly spans. Each intermediate flow er grid cell contains four dimples that are designed to prevent mid-span channel closure in the ns containing intermediate flow mixers and fuel rod contact with the mixing vanes. This simplified arrangement allows short grid cells so that the intermediate flow mixer grid can accomplish its mixing objective with minimal pressure drop.

intermediate flow mixer (IFM) grids, like the mixing vane grids, are fabricated from ZIRLO'. The rmediate flow mixer grids are manufactured using the same basic techniques as the zirconium y structural grid assemblies and are joined to the guide thimbles via sleeves which are welded at bottom of appropriate grid cells.

impact testing has been performed on zirconium alloy structural grids and the intermediate flow er grids indicative of the AP1000 design. The purpose of the testing was to determine the amic buckling, or crush, strength of the grids. The grid impact testing was performed at an ated temperature of 600°F. This temperature is a conservative value representing the core rage temperature at the mid-grid locations.

intermediate flow mixer grids are not intended to be structural members. The intermediate flow er grids do, however, share the loads of the structural grids during faulted loading and, as such, tribute to enhance the load carrying capability of the AP1000 fuel assembly.

dynamic crush strength of the AP1000 structural grids and intermediate flow mixer grids elope the calculated grid impact loading during combined seismic and pipe rupture events. A lable geometry is, therefore, provided at the intermediate flow mixer grid elevations, as well as at structural grid elevations.

2.3 In-core Control Components ctivity control is provided by neutron absorbing rods, gray rods, burnable absorber rods, and a ble chemical neutron absorber (boric acid). The boric acid concentration is varied to control long-reactivity changes such as:

Fuel depletion and fission product buildup Cold to hot, zero power reactivity changes Reactivity change produced by intermediate-term fission products such as xenon and samarium Burnable absorber depletion chemical and volume control system, which is used to adjust the level of boron in the coolant, is ussed in Section 9.3.

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Reactivity changes due to coolant temperature changes in the power range Reactivity changes associated with the power coefficient of reactivity Reactivity changes due to void formation egative power coefficient is maintained at hot, full-power conditions throughout the entire cycle to uce possible deleterious effects caused by a positive coefficient during pipe rupture or loss-of-flow dents. The first fuel cycle needs more excess reactivity than subsequent cycles due to the ing of fresh (unburned) fuel. Since soluble boron alone is insufficient to provide a negative erator coefficient, burnable absorber assemblies are also used. Use of burnable absorber emblies during reloads is discussed in Subsection 4.3.1.2.2.

most effective reactivity control components are the rod cluster control assemblies and the esponding drive rod assemblies, which along with the gray rod cluster assemblies, are the only tic parts in the reactor. Figure 4.2-8 identifies the rod cluster control and drive rod assembly, in ition to the arrangement of these components in the reactor relative to the interfacing fuel embly, guide thimbles, and control rod drive mechanism. The arrangement for the gray rod cluster emblies is the same.

hown in Figure 4.2-8, the guidance system for the rod cluster control assembly is provided by the e thimbles. The guide thimbles provide two regimes of guidance: first, in the lower section, a tinuous guidance system provides support immediately above the core, which protects the rod inst excessive deformation and wear caused by hydraulic loading. Second, the region above the tinuous section provides support and guidance at uniformly spaced intervals.

hown in Figure 4.2-9, the envelope of support is determined by the pattern of the control rod ter. The guide thimbles provide alignment and support of the control rods, spider body, and drive while maintaining trip times at or below required limits.

sections 4.2.2.3.1 through 4.2.2.3.4 describe each reactivity control component in detail. The trol rod drive mechanism assembly is described in Subsection 9.3.4. The neutron source emblies provide a means of monitoring the core during periods of low neutron activity.

2.3.1 Rod Cluster Control Assemblies rod cluster control assemblies are divided into two categories: control and shutdown. The control ups compensate for reactivity changes due to variations in operating conditions of the reactor, that ower and temperature variations. Two nuclear design criteria have been employed for selection e control group. First, the total reactivity worth must be adequate to meet the nuclear uirements of the reactor. Second, in view of the fact that these rods may be partially inserted at er operation, the total power peaking factor should be low enough to confirm that the power ability is met. The control and shutdown groups provide adequate shutdown margin.

llustrated in Figure 4.2-9, a rod cluster control assembly is comprised of a group of individual tron absorber rods fastened at the top end to a common spider assembly.

absorber material used in the control rods is silver-indium-cadmium alloy, which is essentially ck to thermal neutrons and has sufficient additional resonance absorption to significantly ease worth. As shown in Figure 4.2-10, the absorber material is in the form of solid bars sealed in

-worked stainless steel tubes. Sufficient diametral and end clearance is provided to ommodate relative thermal expansions.

4.2-14 Revision 1

material used in the absorber rod end plugs is Type 308 stainless steel. The design stresses d for the Type 308 material are the same as those defined in the ASME Code,Section III, for e 304 stainless steel. At room temperature, the yield and ultimate stresses per ASTM 580 ference 12) are exactly the same for the two alloys. In view of the similarity of composition of the ys, the temperature dependence of strength for the two materials is expected to be the same.

allowable stresses used as a function of temperature are listed in Table I-1.2 of the ASME Code, tion III. The fatigue strength for the Type 308 material is based on the S-N curve for austenitic nless steels in Figure I-9.2 of the ASME Code,Section III.

spider assembly is in the form of a central hub with radial vanes containing cylindrical fingers which the absorber rods are suspended. Internal groove-like profiles to facilitate handling tool drive rod assembly connection are machined into the upper end of the hub. Coil springs inside spider body absorb the impact energy at the end of a trip insertion. The radial vanes are joined to hub by welding and brazing, and the fingers are joined to the vanes by brazing. A bolt, which s the springs and retainer, is threaded into the hub within the skirt and welded to prevent ening while in service.

components of the spider assembly are made from Types 304 and 308 stainless steel except for retainer, which is of Type 630 material, and the springs, which are nickel-chromium-iron y 718.

absorber rods are fastened securely to the spider. The rods are first threaded into the spider ers and then pinned to maintain joint tightness. The pins are then welded in place. The end plug w the pin position is designed with a reduced section to permit flexing of the rods to correct for ll operating or assembly misalignments.

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

2.3.2 Gray Rod Cluster Assemblies mechanical design of the gray rod cluster assemblies plus the control rod drive mechanism and interface with the fuel assemblies and guide thimbles are identical to the rod cluster control embly.

hown in Figure 4.2-11, the gray rod cluster assemblies consist of 24 rodlets fastened at the top to a common hub or spider. Geometrically, the gray rod cluster assembly is the same as a rod ter control assembly except that 12 of the 24 rodlets are stainless steel while the remaining 12 tain the reduced diameter silver-indium-cadmium absorber material clad with stainless steel as rod cluster control assemblies.

gray rod cluster assemblies are used in load follow maneuvering and provide a mechanical shim place the use of changes in the concentration of soluble boron, that is, a chemical shim, normally d for this purpose. The AP1000 uses 53 rod cluster control assemblies and 16 gray rod cluster emblies.

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n assembly. Figure 4.2-12 shows the burnable absorber assemblies. When needed for nuclear siderations, burnable absorber assemblies are inserted into selected thimbles within fuel emblies.

burnable absorber rods (Pyrex) consist of borosilicate glass tubes contained within Type 304 nless steel tubular cladding, which is plugged and seal welded at the ends to encapsulate the

s. The burnable absorber assembly is shown in Figure 4.2-13.

typical discrete burnable absorber rods (WABA) consist of pellets of alumina-boron carbide erial contained within zirconium alloy tubes. These zirconium alloy tubes, which form the outer for the burnable absorber rod, are plugged, pressurized with helium, and seal-welded at each to encapsulate the stack of absorber material. The absorber stack length, shown in re 4.2-12, is positioned axially within the burnable absorber rod by the use of a zirconium alloy om-end spacer.

burnable absorber rods in each fuel assembly are grouped and attached together at the top end e rods to a hold-down assembly by a flat, perforated retaining plate, which fits within the fuel embly top nozzle and rests on the adapter plate.

retaining plate and the burnable absorber rods are held down and restrained against vertical ion through a spring pack which is attached to the plate and is compressed by the upper core e when the reactor upper internals assembly is lowered into the reactor. With this arrangement, burnable absorber rods cannot be ejected from the core by flow forces. Each rod is attached to baseplate by a nut that is crimped into place.

2.3.4 Neutron Source Assemblies purpose of a neutron source assembly is to provide a base neutron level to give confidence that detectors are operational and responding to core multiplication neutrons. For the first core, a tron source is placed in the reactor to provide a positive neutron count of at least two counts per ond on the source range detectors attributable to core neutrons. The detectors, called source ge detectors, are used primarily during subcritical modes of core operation.

source assembly also permits detection of changes in the core multiplication factor during core ing, refueling, and approach to criticality. This can be done since the multiplication factor is ted to an inverse function of the detector count rate. Changes in the multiplication factor can be cted during addition of fuel assemblies while loading the core, changes in control rod positions, changes in boron concentration.

h primary and secondary neutron source rods are used. The primary source rod, containing a oactive material, spontaneously emits neutrons during initial core loading, reactor startup, and al operation of the first core. After the primary source rod decays beyond the desired neutron flux l, neutrons are then supplied by the secondary source rod. The secondary source rod contains a le material, which is activated during reactor operation. The activation results in the subsequent ase of neutrons.

r source assemblies are typically installed in initial load of the reactor core: two primary source emblies and two secondary source assemblies. Each primary source assembly contains one ary source rod and a number of burnable absorber rods. Each secondary source assembly tains a symmetrical grouping of secondary source rodlets. Figure 4.2-14 shows the primary rce assembly. Figure 4.2-15 shows the secondary source assembly.

4.2-16 Revision 1

hown in Figures 4.2-14 and 4.2-15, the source assemblies contain a hold-down assembly tical to that of the burnable absorber assembly.

primary and secondary source rods both use the same cladding material as the absorber rods.

secondary source rods contain antimony-beryllium pellets stacked to a height of approximately nches. The primary source rods contain capsules of californium (plutonium-beryllium possible rnate) source material and alumina spacers to position the source material within the cladding.

rods in each assembly are fastened at the top end to a hold-down assembly.

other structural members, except for the springs, are constructed of Type 304 stainless steel.

springs exposed to the reactor coolant are nickel-chromium-iron Alloy 718.

3 Design Evaluation e fuel assemblies, fuel rods, and in-core control components are designed to satisfy the ormance and safety criteria of]* Section 4.2 of the Standard Review Plan, the mechanical design es of Subsection 4.2.1 and [the Fuel Criteria Evaluation Process per WCAP-12488-A ference 1)]*, and other interfacing nuclear and thermal and hydraulic design bases specified in tions 4.3 and 4.4.

cts of Conditions II, III, IV or anticipated transients without trip on fuel integrity are presented in pter 15.

initial step in fuel rod design evaluation for a region of fuel is to determine the limiting rod(s).

iting rods are defined as those rods whose predicted performance provides the minimum margin ach of the design criteria. For a number of design criteria, the limiting rod is the lead burnup rod of el region. In other instances, it may be the maximum power or the minimum burnup rod. For the t part, no single rod is limiting with respect to all the design criteria.

r identifying the limiting rod(s), an analysis is performed to consider the effects of rod operating ory, model uncertainties, and dimensional variations. To verify adherence to the design criteria, evaluation considers the effects of postulated transient power changes during operation sistent with Conditions I and II. These transient power increases can affect both rod average and l power levels. Parameters considered include rod internal pressure, fuel temperature, clad ss, and clad strain. In fuel rod design analyses, these performance parameters provide the basis omparison between expected fuel rod behavior and the corresponding design criteria limits.

l rod and assembly models used for the performance evaluations are documented and ntained under an appropriate control system. Material properties used in the design evaluations given in WCAP-12610 (Reference 5).

3.1 Cladding 3.1.1 Vibration and Wear l rod vibrations are flow induced. The effect of vibration on the fuel assembly and individual fuel is minimal. The cyclic stress range associated with deflections of such small magnitude is gnificant and has no effect on the structural integrity of the fuel rod.

reaction force on the grid supports, due to rod vibration motions, is also small and is much less the spring preload. Adequate fuel clad spring contact is maintained. No significant wear of the Staff approval is required prior to implementing a change in this information.

4.2-17 Revision 1

d fretting and fuel vibration has been experimentally investigated, as shown in WCAP-8278 ference 13).

3.1.2 Fuel Rod Internal Pressure and Cladding Stresses urnup-dependent fission gas release model (WCAP-15063-P-A, Revision 1 [Reference 21]) is d in determining the internal gas pressure as a function of irradiation time. The plenum volume of fuel rod has been designed to provide that the maximum internal pressure of the fuel rod will not eed the value which would cause:

The fuel/clad diametral gap to increase during steady-state operation Extensive departure from nucleate boiling propagation to occur clad stresses at a constant local fuel rod power are low. Compressive stresses are created by pressure differential between the coolant pressure and the rod internal gas pressure. Because of pre-pressurization with helium, the volume average effective stresses are always less than roximately 14,000 psi at the pressurization level used in the AP1000 fuel rod design. Stresses to the temperature gradient are not included in this average effective stress because thermal sses are, in general, negative at the clad inside diameter and positive at the clad outside meter, and their contribution to the clad volume average stress is small. Furthermore, the thermal ss decreases with time during steady-state operation due to stress relaxation. The stress due to sure differential is highest in the minimum power rod at beginning-of-life due to low internal gas sure and decreases as rod power increases. Thermal stresses are maximum in the maximum er rod due to the larger temperature gradient and decrease as the rod power is decreased.

internal gas pressure at beginning-of-life ranges from approximately 200 to 750 psi for typical burnup fuel rods. The total tangential stress at the clad inside diameter at beginning-of-life is roximately 19,500 psi compressive (approximately 18,500 psi due to P and approximately 0 due to T) for a low-power rod operating at four kilowatts/foot. Total tangential stress is roximately 20,500 psi compressive (approximately 18,000 psi due to P and approximately 0 psi due to T) for a high-power rod operating at 10 kilowatts/foot. However, the volume rage effective stress at beginning-of-life is between approximately 13,500 psi (high-power rod) approximately 14,000 psi (low-power rod). These stresses are substantially below even the radiated clad yield strength (approximately 55,500 psi) at a typical clad mean operating perature of 700°F.

sile stresses could be created once the clad has come in contact with the pellet. These stresses ld be induced by the fuel pellet swelling during irradiation. Swelling of the fuel pellet can result in ll clad strains (less than one percent) for expected discharge burnups, but the associated clad sses are very low because of clad creep (thermal- and irradiation-induced creep). The one ent strain criterion is extremely conservative for fuel-swelling driven clad strain because the in rate associated with solid fission products swelling is very slow. A detailed discussion of fuel performance is given in Subsection 4.2.3.3.

3.1.3 Material and Chemical Evaluation LO' clad has a high corrosion resistance to the coolant, fuel, and fission products. As shown in AP-8183 (Reference 3), there is considerable pressurized water reactor operating experience on capability of Zircaloy-4 as a clad material. ZIRLO', an advanced zirconium based alloy, has al or better corrosion resistance than Zircaloy-4 (see WCAP-12610-P-A, [Reference 5]). Controls uel fabrication specify maximum moisture levels to preclude clad hydriding.

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rface layer remains very thin even at high burnup. Thus, there is no indication of propagation of layer and eventual clad penetration.

ss corrosion cracking is another postulated phenomenon related to fuel/clad chemical interaction.

-of-pile tests have shown that in the presence of high clad tensile stresses, large concentrations dine can chemically attack the zirconium alloy tubing and may lead to eventual clad cracking.

ensive post-irradiation examination has produced no evidence that this mechanism has been rative in Westinghouse commercial pressurized water reactor fuel.

3.1.4 Rod Bowing AP-8691 (Reference 14) presents the model used for evaluation of AP1000 fuel rod bowing. This el has been used for bow assessment in 14x14, 15x15, and 17x17 type cores.

3.1.5 Consequences of Power Coolant Mismatch sequences of power coolant mismatch are discussed in Chapter 15.

3.1.6 Creep Collapse and Creepdown subject and the associated irradiation stability of cladding have been evaluated. In AP-13589-A (Reference 8), it is shown that current generation Westinghouse fuel is sufficiently le with respect to fuel densification. Significant axial gaps do not form in the pellet stack, enting clad collapse from occurring. The design basis of no clad collapse during planned core life erefore satisfied. Cladding collapse analyses, if required, would be performed using the methods cribed in WCAP-8377 (Reference 22).

3.2 Fuel Materials Considerations ered, high-density uranium dioxide fuel reacts only slightly with the clad at core operating peratures and pressures. In the event of clad defects, the high resistance of uranium dioxide to ck by water protects against fuel deterioration, although limited fuel erosion can occur. The sequences of defects in the clad are greatly reduced by the ability of uranium dioxide to retain on products, including those which are gaseous or highly volatile.

ervations from several early Westinghouse pressurized water reactors as discussed in AP-8218-P-A (Reference 6) have shown that fuel pellets can densify under irradiation to a sity higher than the manufactured values. Fuel densification and subsequent settling of the fuel ets can result in local and distributed gaps in the fuel rods. The densification process is related to elimination of very small as-fabricated porosity in the fuel during irradiation. Early fuels were ntionally manufactured to low initial density and were undersintered, which resulted in a large tion of very small pores. Densification behavior in current fuel is controlled by improved ufacturing process controls and by specifying a nominal 95.5 percent initial fuel density, which lts in reduced levels of small, densifying porosity.

evaluation of fuel densification effects and the treatment of fuel swelling and fission gas release described in WCAP-13589-A (Reference 8) and WCAP-15063-P-A, Revision 1 (Reference 21).

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ors are considered:

Clad creep and elastic deflection Pellet density changes, thermal expansion, gas release, and thermal properties as a function of temperature and fuel burnup Internal pressure as a function of fission gas release, rod geometry, and temperature distribution se effects are evaluated using fuel rod design models, as discussed in WCAP-15063-P-A, ision 1 (Reference 21), that include appropriate models for time dependent fuel densification.

h these interacting factors considered, the model determines the fuel rod performance racteristics for a given rod geometry, power history, and axial power shape. In particular, internal pressure, fuel and clad temperatures, and clad deflections are calculated. The fuel rod is divided several axial sections and radially into a number of annular zones. Fuel density changes are ulated separately for each segment. The effects are integrated to obtain the internal rod sure.

initial rod internal pressure is selected to delay fuel/clad mechanical interaction and to avoid the ntial for clad flattening. It is limited, however, by the design criteria for the rod internal pressure, iscussed in Subsection 4.2.1.3.

gap conductance between the pellet surface and the clad inner diameter is calculated as a tion of the composition, temperature and pressure of the gas mixture, and the gap size or contact sure between the clad and pellet. After computing the fuel temperature for each pellet zone, the tional fission gas release is assessed using an empirical model derived from experimental data, etailed in WCAP-15063-P-A, Revision 1 (Reference 21). The total amount of gas released is ed on the average fractional release within each axial and radial zone and the gas generation

, which, in turn, is a function of burnup. Finally, the gas released is summed over the zones, and pressure is calculated.

model shows close agreement in fit for a variety of published and proprietary data on fission gas ase, fuel temperatures, and clad deflections, as detailed in WCAP-15063-P-A, Revision 1 ference 21). These data include variations in power, time, fuel density, and geometry.

3.3.1 Fuel/Cladding Mechanical Interaction factor in fuel element duty is potential mechanical interaction of the fuel and clad. This fuel/clad raction produces cyclic stresses and strains in the clad, and these, in turn, reduce clad life. The uction of fuel/clad interaction is therefore a goal of design. The technology for using pre-surized fuel rods in Westinghouse pressurized water reactors has been developed to further this ctive.

gap between the fuel and clad is initially sufficient to prevent hard contact between the two.

ever, during power operation a gradual compressive creep of the clad onto the fuel pellet occurs to the external pressure exerted on the rod by the coolant. Clad compressive creep eventually lts in fuel/clad contact. Once fuel/clad contact occurs, changes in power level result in changes in stresses and strains. By using pre-pressurized fuel rods to partially offset the effect of the lant external pressure, the rate of clad creep toward the surface of the fuel is reduced. Fuel rod pressurization delays the time at which fuel/clad contact occurs and, hence, significantly reduces 4.2-20 Revision 1

o-dimensional (r,) finite element model has been established to investigate the effects of radial et cracks on stress concentrations in the clad. Stress concentration herein is defined as the rence between the maximum clad stress in the direction and the mean clad stress. The first e has the fuel and clad in mechanical equilibrium; and, as a result, the stress in the clad is close ero. In subsequent cases the pellet power is increased in steps and the resultant fuel thermal ansion imposes tensile stress in the clad.

ddition to uniform clad stresses, stress concentrations develop in the clad adjacent to radial ks in the pellet. These radial cracks have a tendency to open during a power increase, but the ional forces between fuel and clad oppose the opening of these cracks and result in localized eases in clad stress. As the power is further increased, large tensile stresses exceed the ultimate ile strength of uranium dioxide and additional cracks in the fuel pellet are created, limiting the nitude of the stress concentration in the clad.

part of the standard fuel rod design analysis, the maximum stress concentration evaluated from e element calculations is added to the volume-averaged effective stress in the clad as determined one-dimensional stress/strain calculations. The resultant clad stress is then compared to the perature-dependent cladding yield stress to confirm that the stress/strain criteria are satisfied.

transient evaluation method is described in the following paragraphs.

et thermal expansion due to power increases is considered the only mechanism by which ificant stresses and strains can be imposed on the clad.

er increases in commercial reactors can result from fuel shuffling (for example, region 3 itioned near the core center for cycle 2 operation after operating near the periphery during e 1), reactor power escalation following extended reduced power operation, and full-length trol rod movement. In the mechanical design model, lead rods are depleted using best-estimate er histories as determined by core physics calculations. During burnup, the amount of diametral closure is evaluated based upon the pellet expansion cracking model, clad creep model, and fuel lling model. At various times during the depletion, the power is increased locally in the rod to the nup-dependent attainable power density as determined by core physics calculations. The radial, ential, and axial clad stresses resulting from the power increase are combined into a volume rage effective clad stress.

von Mises criterion is used to determine whether the clad yield stress has been exceeded. This rion states that an isotropic material in multi-axial stress will begin to yield plastically when the ctive stress exceeds the yield stress as determined by an axial tensile test. The yield stress elation is that for irradiated cladding, since fuel/clad interaction occurs at high burnup. In applying criterion, the effective stress is increased by an allowance which accounts for stress centrations in the clad adjacent to radial cracks in the pellet, prior to the comparison with the yield ss. This allowance was evaluated using a two-dimensional (r,) finite element model.

w transient power increases can result in large clad strains without exceeding the clad yield stress ause of clad creep and stress relaxation. Therefore, in addition to the yield stress criterion, a rion on allowable clad strain is necessary. Based upon high strain rate burst and tensile test data rradiated tubing, one percent strain was determined to be a conservative lower limit on irradiated ductility and that was adopted as a design criterion.

4.2-21 Revision 1

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

assessment of the fatigue life of the fuel rod cladding is subjected to considerable uncertainty ause of the difficulty of evaluating the strain range which results from the cyclic interaction of the pellets and cladding. This difficulty arises, for example, from such highly unpredictable nomena as pellet cracking, fragmentation, and relocation. Westinghouse investigated this icular phenomenon both analytically and experimentally. Strain fatigue tests on irradiated and irradiated hydrided Zircaloy-4 cladding were performed. These tests permitted the definition of a servative fatigue-life limit and recommendation of a methodology to treat the strain fatigue luation of the Westinghouse-referenced fuel rod designs. (See WCAP-9500-P-A, Reference 15.)

cessful load follow operation has been performed on several reactors. There was no significant lant activity increase that could be associated with the load follow mode of operation.

Westinghouse analytical approach to strain fatigue is based on a comprehensive review of the ilable strain fatigue models. The review included the Langer-ODonnell model (Reference 16) the

-Munse model, and the Manson-Halford model. Upon completion of this review, and using the lts of the Westinghouse experimental programs as documented in WCAP-9500-P-A ference 15), it was concluded that the approach defined by Langer-ODonnell would be retained the empirical factors of their correlation modified to conservatively bound the results of the stinghouse testing program.

design equations followed the concept for the fatigue design criterion according to the ASME e, Section III:

The calculated pseudo stress amplitude (Sa) has to be multiplied by a factor of two to obtain the allowable number of cycles (Nf).

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

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

k n

Nk 1 1 fk re:

= number of diurnal cycles of mode k.

= number of allowable cycles.

3.3.2 Irradiation Experience stinghouse fuel operational experience is presented in WCAP-8183 (Reference 3). Additional test embly and test rod experience is given in WCAP-10125-P-A (Reference 2).

4.2-22 Revision 1

3.3.4 Potentially Damaging Temperature Effects During Transients fuel rod experiences many operational transients (intentional maneuvers) during its residence in core. A number of thermal effects must be considered when analyzing the fuel rod performance.

clad can be in contact with the fuel pellet at some time in the fuel lifetime. Clad/pellet interaction urs if the fuel pellet temperature is increased after the clad is in contact with the pellet. Clad/pellet raction is discussed in Subsection 4.2.3.3.1.

d flattening has been observed in some operating power reactors. This is no longer a concern ause clad flattening is precluded during the fuel residence in the core (Subsection 4.2.3.1) by the of stable fuel.

ential differential thermal expansion between the fuel rods and the guide thimbles during a sient is considered in the design. Excessive bowing of the fuel rods is precluded because the grid emblies allow axial movement of the fuel rods relative to the grids. Specifically, thermal expansion e fuel rods is considered in the grid design so that axial loads imposed on the fuel rods during a mal transient will not result in excessively bowed fuel rods.

3.3.5 Fuel Element Burnout and Potential Energy Release discussed in Subsection 4.4.2.2, the core is protected from departure from nucleate boiling over full range of possible operating conditions. In the extremely unlikely event that departure from leate boiling should occur, the clad temperature will rise due to the steam blanketing at the rod ace and the consequent degradation in heat transfer. During this time there is a potential for mical reaction between the cladding and the coolant. However, because of the relatively good film ng heat transfer following departure from nucleate boiling, the energy release resulting from this tion is insignificant compared to the power produced by the fuel.

3.3.6 Coolant Flow Blockage Effects on Fuel Rods coolant flow blockage effects on fuel rods are presented in Subsection 4.4.4.7.

3.4 Spacer Grids coolant flow channels are established and maintained by the structure composed of grids and e thimbles. The lateral spacing between fuel rods is provided and controlled by the support ples of adjacent grid cells. Contact of the fuel rods on the dimples is maintained through the ping force of the grid springs. Lateral motion of the fuel rods is opposed by the spring force and internal moments generated between the spring and the support dimples. Grid testing is ussed in WCAP-8236 (Reference 17) and WCAP-10444-P-A (Reference 11).

3.5 Fuel Assembly 3.5.1 Stresses and Deflections fuel assembly component stress levels are limited by the design. For example, stresses in the rod due to thermal expansion and zirconium alloy irradiation growth are limited by the relative ion of the rod as it slips over the grid spring and dimple surfaces. Clearances between the fuel ends and nozzles are provided so that zirconium alloy irradiation growth does not result in rod 4.2-23 Revision 1

n subjected to structural tests to verify that the design bases requirements are met.

fuel assembly design loads for shipping have been established at 4 g axial and 6 g lateral.

elerometers are permanently placed in the shipping cask to monitor and detect fuel assembly elerations that would exceed the criteria. Experience indicates that loads that exceed the wable limits rarely occur. Exceeding the limits requires reinspection of the fuel assembly for age. Tests on various fuel assembly components, such as the grid assembly, sleeves, inserts, structure joints, have been performed to confirm that the shipping design limits do not result in airment of fuel assembly function. Seismic analysis methodology of the fuel assembly is ented in WCAP-8236 (Reference 17), WCAP-9401-P-A (Reference 18), and WCAP-10444-P-A ference 11).

emonstrate that the fuel assemblies will maintain a geometry that is capable of being cooled er the worst-case accident Condition IV event, a plant specific or bounding seismic analysis is ormed.

fuel assembly response resulting from safe shutdown earthquake condition is analyzed using

-history numerical techniques. The vessel motion for this type of event primarily causes lateral s on the reactor core. Consequently, the methodology and analytical procedures as described in AP-8236 (Reference 17) and WCAP-9401-P-A (Reference 18) are used to assess the fuel embly deflections and impact forces.

motions of the reactor internals upper and lower core plates and the core barrel at the upper core e elevation, which are simultaneously applied to simulate the reactor core input motion, are ined from the time-history analysis of the reactor vessel and internals. The fuel assembly onse, namely the displacements and impact forces, is obtained with the reactor core model.

ilar dynamic analyses of the core were performed using reactor internals motions indicative of the tulated pipe rupture. Scenarios regarding breaches in the pressure boundary are investigated to rmine the most limiting structural loads for the fuel assembly. The application of leak-before-ak limits the size of the pipe rupture loads for which the fuel assemblies must be analyzed. The rupture used in the fuel assembly analysis is the largest pipe connected to the reactor coolant em which does not satisfy the leak-before-break criteria. Subsection 3.6.3 discusses mechanistic break.

3.5.1.1 Grid Analyses maximum grid impact force obtained from seismic analyses is less than the allowable grid ngth. With respect to the guidelines of Appendix A of the Standard Review Plan, Section 4.2, stinghouse has demonstrated that a simultaneous safe shutdown earthquake and pipe rupture nt is highly unlikely. The fatigue cycles, crack initiation, and crack growth due to normal operating seismic events will not realistically lead to a pipe rupture. More information is available in AP-9283 (Reference 19).

ed on the deterministic fracture mechanics evaluation of small flaws in piping components, stinghouse has demonstrated that the dynamic affects of a large pipe rupture in the primary lant piping system for the AP1000 design does not have to be considered.

sign basis for the piping design in the AP1000 is that the reactor coolant loop and surge lines will sfy the leak-before-break criteria for mechanistic pipe break. In addition, the piping connected to reactor coolant system that is six inch nominal diameter or larger is evaluated for leak-before-ak. The result of a pipe leakage event consistent with the mechanistic pipe break evaluation 4.2-24 Revision 1

pressure boundary integrity for numerous branch lines is analyzed to determine the most limiting ak of a line not qualified for leak-before-break for the dynamic loading of the reactor core. Grid s resulting from a combined seismic and pipe rupture event do not cause unacceptable grid rmation as to preclude a core coolable geometry.

3.5.1.2 Nongrid Analyses stresses induced in the various fuel assembly nongrid components are assessed based on the t limiting seismic condition. The fuel assembly axial forces resulting from the hold-down spring together with its own weight distribution are the primary sources of the stresses in the guide bles and fuel assembly nozzles. The fuel rod accident induced stresses, which are generally very ll, are caused by bending due to the fuel assembly deflections during a seismic event. The mic-induced stresses are compared with the allowable stress limits for the fuel assembly major ponents. The component stresses, which include normal operating stresses, are below the blished allowable limits. Consequently, the structural designs of the fuel assembly components acceptable for the design basis accident conditions for the AP1000.

3.5.2 Dimensional Stability alized yielding and slight deformation in some fuel assembly components are allowed to occur ng a Condition III or IV event. The maximum permanent deflection, or deformations, do not result ny violation of the functional requirements of the fuel assembly.

3.6 Reactivity Control Assemblies and Burnable Absorber Rods 3.6.1 Internal Pressure and Cladding Stresses during Normal, Transient, and Accident Conditions designs of the burnable absorber and source rods provide a sufficient cold void volume to ommodate the internal pressure increase during operation. This is not a concern for the rod ter control assembly absorber rod or gray rod cluster assembly rodlets because no gas is ased by the silver-indium-cadmium absorber material.

the discrete burnable absorber rod, there is sufficient cold void volume to limit the internal ssure to a value, which satisfies the design criteria. For the source rods, a void volume is vided within the rod to limit the maximum internal pressure increase at end-of-life.

ures 4.2-14 and 4.2-15 detail the primary and secondary source assemblies.

ing normal transient and accident conditions, the void volume limits the internal pressures to es that satisfy the criteria in Subsection 4.2.1.6. These limits are established not only to prevent peak stresses from reaching unacceptable values, but also to limit the amplitude of the oscillatory ss component in consideration of the fatigue characteristics of the materials.

, guide thimble, and dashpot flow analyses indicate that the flow is sufficient to prevent coolant ng within the guide thimble. Therefore, clad temperatures at which the clad material has quate strength to resist coolant operating pressures and rod internal pressures are maintained.

4.2-25 Revision 1

radial and axial temperature profiles within the source and absorber rods are determined by sidering gap conductance, thermal expansion, neutron or gamma heating of the contained erial as well as gamma heating of the clad.

maximum temperatures of the silver-indium-cadmium control rod absorber material are ulated and found to be significantly less than the material melting point and found to occur axially nly the highest flux region. The mechanical and thermal expansion properties of the silver-indium-mium absorber material are discussed in WCAP-9179 (Reference 4).

e first core, the burnable absorber rods (Pyrex) consist of borosilicate glass tubes contained in Type 304 stainless steel tubular cladding, which is plugged and seal welded at the ends to apsulate the glass. The absorber material temperature does not exceed its design limit of 1220°F.

hanical and thermal design and nuclear evaluation of the burnable absorber rods are described CAP-7113 (Reference 23).

maximum temperature of the alumina-boron carbide burnable absorber pellet (WABA) is ected to be less than 1200°F which takes place following the initial power ascent. As the rating cycle proceeds, the burnable absorber pellet temperature decreases due to a reduction in t generation due to boron depletion and better gap conduction as the helium produced diffuses the gap.

icient diametral and end clearances have been provided in the neutron absorber, burnable orber, and source rods to accommodate the relative thermal expansions between the enclosed erial and the surrounding clad and end plug.

3.6.3 Irradiation Stability of the Absorber Material, Taking into Consideration Gas Release and Swelling irradiation stability of the silver-indium-cadmium absorber material is discussed in WCAP-9179 ference 4). Irradiation produces no deleterious effects in the absorber material.

mentioned in Subsection 4.2.3.6.1, gas release is not a concern for the control rod material ause no gas is produced by the absorber material. Sufficient diametral and end clearances are ided to accommodate any potential expansion and/or swelling of the absorber material.

alumina-boron carbide burnable absorber pellets are designed such that gross swelling or mbling of the pellets is not predicted to occur during reactor operation. Some minor cracking of the ets may occur, but this cracking should not affect the overall absorber and stack integrity.

3.6.4 Potential for Chemical Interaction, Including Possible Waterlogging Rupture structural materials selected have good resistance to irradiation damage and are compatible with reactor environment.

rosion of the materials exposed to the coolant is quite low, and proper control of chloride and gen in the coolant minimizes potential for the occurrence of stress corrosion. The potential for the rference with rod cluster control assembly movement due to possible corrosion phenomena is low.

4.2-26 Revision 1

silver-indium-cadmium absorber material is relatively inert and will remain inert even when jected to high coolant velocity regions. Rapid loss of reactivity control material will not occur. Test lts detailed in WCAP-9179 (Reference 4) concluded that additions of indium and cadmium to er, in the amounts to form the silver-indium-cadmium absorber material composition, result in ll corrosion rates.

the discrete burnable absorber, in the unlikely event that the zirconium alloy clad is breached, the on carbide in the affected rod(s) could be leached out by the coolant water. If this occurred early, ore instruments could detect large peaking factor changes, and corrective action would be taken, arranted. A postulated clad breach after substantial irradiation would have no significant effect on king factors since the boron will have been depleted. Breaching of the zirconium alloy clad by rnal hydriding is not expected due to moisture controls employed during fabrication. Rods of this ign have performed very well with no failures observed.

4 Testing and Inspection Plan 4.1 Quality Assurance Program Quality Assurance Program Plan of the Westinghouse Commercial Nuclear Fuel Division for the 000 is summarized in Chapter 17.

program provides for control over activities affecting product quality, commencing with design development and continuing through procurement, materials handling, fabrication, testing and ection, storage, and transportation. The program also provides for the indoctrination and training ersonnel and for the auditing of activities affecting product quality through a formal auditing gram.

stinghouse drawings and product, process, and material specifications identify the inspections to erformed.

4.2 Quality Control lity control philosophy is generally based on the following inspections being performed to a ercent confidence that at least 95 percent of the product meets specification, unless otherwise d.

4.2.1 Fuel System Components and Parts characteristics inspected depend on the component parts. The quality control program includes ensional and visual examinations, check audits of test reports, material certification, and destructive examination, such as X-ray and ultrasonic.

material used in the AP1000 core is accepted and released by Quality Control.

4.2.2 Pellets ection is performed for dimensional characteristics such as diameter, density, length, and areness of ends. Additional visual inspections are performed for cracks, chips, and surface ditions according to approved standards.

4.2-27 Revision 1

4.2.3 Rod Inspection l rod, rod cluster control rod, discrete burnable absorber rod, and source rod inspections consists e following nondestructive examination techniques and methods, as applicable:

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

Rod welds are inspected by ultrasonic test or X-ray in accordance with a qualified technique and Westinghouse specifications meeting the requirements of ASTM-E-142-86 (Reference 20).

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

Fuel rods are inspected by gamma scanning or other approved methods, as discussed in Subsection 4.2.4.5, to confirm proper plenum dimensions.

Fuel rods are inspected by gamma scanning, or other approved methods, as discussed in Subsection 4.2.4.5, to confirm that no significant gaps exist between pellets.

Fuel rods are actively and/or passively gamma scanned to verify enrichment control prior to acceptance for assembly loading.

Traceability of rods and associated rod components is established by quality control.

4.2.4 Assemblies h fuel rod, control rod, burnable absorber rod and source rod assembly is inspected for pliance with drawing and/or specification requirements. Other in-core control component ection and specification requirements are given in Subsection 4.2.4.4.

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

Tool and gauge inspection and control, including standardization to primary and/or secondary working standards. Tool inspection is performed at prescribed intervals on serialized tools.

Complete records are kept of calibration and conditions of tools.

Audits are performed of inspection activities and records to confirm that prescribed methods are followed and that records are correct and properly maintained.

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

4.2.6 Process Control revent the possibility of mixing enrichments during fuel manufacture and assembly, strict chment segregation and other process controls are exercised.

4.2-28 Revision 1

onfirmed by analysis.

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

shed pellets are placed on trays identified with the same color code as the powder containers and sferred to segregated storage racks within the confines of the pelleting area. Samples from each et lot are tested for isotopic content and impurity levels prior to acceptance by quality control.

sical barriers are used to prevent mixing of pellets of different nominal densities and enrichments e pellet storage area. Unused powder and substandard pellets are returned to storage in the inal color-coded containers.

ding of pellets into the clad is performed in isolated production lines; only one density and chment (with possible exception for top and bottom (axial blanket) zones) are loaded on a line at me.

rialized traceability code is placed on each fuel tube, which identifies the contract and chment. The end plugs are inserted and then welded (in an inert gas atmosphere) to seal the

. The fuel tube remains coded and traceability identified until just prior to installation in the fuel embly.

ilar traceability is provided for wet annular burnable absorber, source, and control rods, as uired.

4.3 Letdown Radiation Monitoring iation monitoring of the reactor coolant is made by grab samples and laboratory analysis of the ary coolant. Refer to information presented in Subsections 9.3.3 and 9.3.6, and Table 9.3.3-1.

4.4 In-core Control Component Testing and Inspection s and inspections are performed on each reactivity control component to verify the mechanical racteristics. In the case of the rod cluster control assembly, prototype testing has been ducted. Manufacturing test/inspections and functional testing at the plant site are both performed.

ing the component manufacturing phase, the following requirements apply to the reactivity control ponents to provide the proper functioning during reactor operation:

Materials are procured to specifications to attain the desired standard of quality.

Spider assemblies are proof-tested by applying a 5000-pound load to the spider body, so that approximately 310 pounds is applied to each vane. This proof load provides a bending moment at the spider body approximately equivalent to 1.4 times the load caused by the acceleration imposed by the control rod drive mechanism.

Rods are checked for integrity by the applicable nondestructive methods described in Subsection 4.2.4.2.3.

4.2-29 Revision 1

rod cluster control assemblies and gray rod cluster assemblies are also functionally tested, wing core loading but prior to criticality, to demonstrate reliable operation of the assemblies. Each embly is operated (and tripped) one time at full-flow/hot conditions. In addition, any assembly that a drop time greater than a two sigma limit from the average rod drop time is subjected to itional rod drops to confirm drop time. Thus, each assembly is sufficiently tested to confirm proper tioning and operation.

emonstrate continuous free movement of the rod cluster control assemblies, and gray rod cluster emblies, and to provide acceptable core power distributions during operations, partial movement cks are performed on every assembly as required by the technical specifications. In addition, odic drop tests of the assemblies are performed at each refueling shutdown to demonstrate tinued ability to meet trip time requirements.

rod cluster control assembly and/or gray rod cluster assembly cannot be moved by its hanism, adjustments in the boron concentration of the coolant provide that adequate shutdown gin will be achieved following a trip. Thus, inability to move one assembly can be tolerated until reactor can be safely taken to Mode 3. More than one inoperable assembly could be tolerated but ld impose additional demands on the plant operator. Therefore, the number of inoperable emblies has been limited to one.

4.5 Tests and Inspections by Others tests and inspections performed by others, Westinghouse reviews and approves the quality trol procedures, and inspection plans to be utilized to confirm that they are equivalent to the cription provided in Subsections 4.2.4.1 through 4.2.4.4 and are performed properly to meet stinghouse requirements.

4.6 Inservice Surveillance detailed in WCAP-8183 (Reference 3), significant 17x17 fuel assembly operating experience has n obtained. A surveillance program is expected to be established for the AP1000 for inspection of t-irradiated fuel assemblies. This surveillance program will establish the schedule, guidelines, and ection criteria for conducting visual inspection of post-irradiated fuel assemblies and/or insert ponents. The surveillance program includes a quantitative visual examination of some harged fuel assemblies from each refueling. This program also includes criteria for additional ection requirements for post-irradiated fuel assemblies if unusual characteristics are noticed in visual inspection or if plant instrumentation and subsequent laboratory analysis indicates gross d fuel. The post-irradiated fuel surveillance program will address disposition of fuel assemblies

/or insert components receiving an unsatisfactory visual inspection. Those post-irradiated fuel emblies receiving an unsatisfactory visual inspection are not reinserted into the core until a more iled inspection and/or evaluation can be performed. Normally the fuel assemblies are taken to spent fuel inspection station.

4.7 Onsite Inspection ten procedures are used for the post-shipment inspection of the new fuel assemblies in addition activity control and source components. Fuel handling procedures specify the sequence in which dling and inspection take place.

4.2-30 Revision 1

sit exceeded design limitations.

owing removal of the fuel assembly from the container in accordance with detailed procedures, fuel assembly plastic wrapper is examined for evidence of damage. The polyethylene wrapper is removed, and a visual inspection of the entire fuel assembly is performed.

trol rod, gray rod, secondary source rod and discrete burnable absorber rod assemblies are ally shipped in fuel assemblies. They are inspected prior to removal of the fuel assembly from the tainer. The control rod assembly is withdrawn a few inches from the fuel assembly to confirm free unrestricted movement, and the exposed section is visually inspected for mechanical integrity, aced in the fuel assembly, and stored with the fuel assembly. Control rod, secondary source or rete burnable absorber assemblies may be stored separately or within fuel assemblies in the new storage area.

5 Combined License Information nges to the reference design of the fuel, burnable absorber rods, rod cluster control assemblies, itial core design from that presented in the DCD are addressed in APP-GW-GLR-059 ference 24).

6 References

[Davidson, S. L. (Ed.), Fuel Criteria Evaluation Process, WCAP-12488-A (Proprietary) and WCAP-14204-A (Non-Proprietary), October 1994.]*

Davidson, S. L. (Ed.) et al., Extended Burnup Evaluation of Westinghouse Fuel, WCAP-10125-P-A (Proprietary) and WCAP-10126-NP-A (Non-Proprietary), December 1985.

Operational Experience with Westinghouse Cores, WCAP-8183, (revised annually).

Beaumont, M. D., et al., Properties of Fuel and Core Component Materials, WCAP-9179, Revision 1 (Proprietary) and WCAP-9224 (Non-Proprietary), July 1978.

Davidson, S. L., and Nuhfer, D. L. (Ed.), VANTAGE+ Fuel Assembly Reference Core Report, WCAP-12610-P-A (Proprietary), June 1990 and WCAP-14342-A (Non-Proprietary), April 1995.

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

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 (Nonproprietary), August, 1988.

Davidson, S. L. (Ed) et al., Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel, WCAP-13589-A (Proprietary) and WCAP-14297-A (Non-Proprietary), March 1995.

Staff approval is required prior to implementing a change in this information.

4.2-31 Revision 1

Skarita, J., et al., Westinghouse Wet Annular Burnable Absorber Evaluation Report, WCAP-10021-P-A, Revision 1 (Proprietary) and WCAP-10377-NP-A, Revision 2 (Non-Proprietary), October 1983.

Davidson, S. L. (Ed.) et al., Reference Core Report VANTAGE 5 Fuel Assembly, WCAP-10444-P-A (Proprietary) and WCAP-10445-NP-A (Nonproprietary), September 1985.

ASTM-A-580-90, Specification for Stainless and Heat-resisting Steel Wire.

Demario, E. E., Hydraulic Flow Test of the 17x17 Fuel Assembly, WCAP-8278 (Proprietary) and WCAP-8279 (Non-Proprietary), February 1974.

Skaritka, J. (Ed.), Fuel Rod Bow Evaluation, WCAP-8691, Revision 1 (Proprietary) and WCAP-8692, Revision 1 (Non-Proprietary), July 1979.

Davidson, S. L. and Iorii, J. A., Reference Core Report 17x17 Optimized Fuel Assembly, WCAP-9500-P-A (Proprietary) and WCAP-9500-A (Nonproprietary), May 1982.

ODonnell, W. J., and Langer, B. F., Fatigue Design Basis for Zircaloy Components, Nuclear Science and Engineering 20, pp 1-12, 1964.

Gesinski, L., and Chiang, D., Safety Analysis of the 17x17 Fuel Assembly for Combined Seismic and Loss-of-Coolant Accident, WCAP-8236 (Proprietary) and WCAP-8288 (Nonproprietary), December 1973.

Davidson, S. L., et al., Verification, Testing, and Analysis of the 17x17 Optimized Fuel Assembly, WCAP-9401-P-A (Proprietary) and WCAP-9402-A (Nonproprietary), August 1981.

Witt, F. J., Bamford, W. H., and Esselman, T. C., Integrity of the Primary Piping Systems of Westinghouse Nuclear Power Plants During Postulated Seismic Events, WCAP-9283 (Nonproprietary), March 1978.

ASTM-E-142-86, Methods for Controlling Quality of Radiographic Testing.

Foster, J. P., et al., Westinghouse Improved Performance Analysis and Design Model (PAD 4.0), WCAP-15063-P-A, Revision 1 (Proprietary) and WCAP-15064-NP-A, Revision 1 (Non-Proprietary), July 2000.

George, R. A., et al., Revised Clad Flattening Model, WCAP-8377 (Proprietary), July 1974.

WCAP-7113, Use of Burnable Poison Rods in Westinghouse Pressurized Water Reactors, October 1967.

APP-GW-GLR-059/WCAP-16652-NP, AP1000 Core & Fuel Design Technical Report, Revision 0.

4.2-32 Revision 1

Figure 4.2-1 Fuel Assembly Cross-Section 4.2-33 Revision 1

WLS 1&2 - UFSAR Figure 4.2-2 Fuel Assembly Outline 4.2-34 Revision 1

Figure 4.2-3 Fuel Rod Schematic 4.2-35 Revision 1

Figure 4.2-4 Top Grid Sleeve Detail 4.2-36 Revision 1

Figure 4.2-5 Intermediate Grid to Thimble Attachment Joint 4.2-37 Revision 1

Figure 4.2-6 Intermediate Flow Mixer Grid to Thimble Attachment 4.2-38 Revision 1

Figure 4.2-7 Grid Thimble to Bottom Nozzle Joint 4.2-39 Revision 1

Figure 4.2-8 Rod Cluster Control and Drive Rod Assembly With Interfacing Components 4.2-40 Revision 1

Figure 4.2-9 Rod Cluster Control Assembly 4.2-41 Revision 1

Figure 4.2-10 Absorber Rod Detail 4.2-42 Revision 1

Figure 4.2-11 Gray Rod Cluster Assembly 4.2-43 Revision 1

Figure 4.2-12 Discrete Burnable Absorber Assembly 4.2-44 Revision 1

145.0

[3683.0]

Figure 4.2-13 Burnable Absorber Rod Assembly (Pyrex) Borosilicate Glass 4.2-45 Revision 1

Figure 4.2-14 Primary Source Assembly 4.2-46 Revision 1

Figure 4.2-15 Secondary Source Assembly 4.2-47 Revision 1

section describes the design bases and functional requirements used in the nuclear design of fuel and reactivity control system and relates these design bases to the General Design Criteria C). The design bases are the fundamental criteria that must be met using approved analytical niques. [Enhancements to these techniques may be made provided that the changes are ded by NRC approved methodologies as discussed in]* WCAP-9272-P-A (Reference 1) and AP-12488-P-A (Reference 2).]*

plant conditions for design are divided into four categories:

Condition I Normal operation and operational transients Condition II Events of moderate frequency Condition III Infrequent incidents Condition IV Limiting faults reactor is designed so that its components meet the following performance and safety criteria:

In general, 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 occurrences are accommodated with, at most, a shutdown of the reactor with the plant capable of returning to operation after corrective action.

Fuel damage, that is, breach of fuel rod clad pressure boundary, is not expected during Condition I and Condition II occurrences. A very small amount of fuel damage may occur.

This is within the capability of the chemical and volume control system (CVS) and is consistent with the plant design basis.

Condition III occurrences 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 occurrences is not sufficient to interrupt or restrict public use of those areas beyond the exclusion area boundary.

A Condition III occurrence does not by itself generate a Condition IV occurrence or result in a consequential loss of function of the reactor coolant or reactor containment barriers.

Condition IV faults do not cause a release of radioactive material that results in exceeding the dose limits identified in Chapter 15. Condition IV occurrences are faults that are not expected to occur but are defined as limiting faults which are included in the design.

core design power distribution limits related to fuel integrity are met for Condition I occurrences ugh conservative design and are maintained by the action of the control system.

requirements for Condition II occurrences are met by providing an adequate protection system ch monitors reactor parameters.

Staff approval is required prior to implementing a change in this information.

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1.1 Fuel Burnup 1.1.1 Basis mitation on initial installed excess reactivity or average discharge burnup is not required other as is quantified in terms of other design bases, such as overall negative power reactivity back discussed below. [The NRC has approved, in WCAP-12488-P-A (Reference 2), maximum rod average burnup of 60,000 MWD/MTU. Extended burnup to 62,000 MWD/MTU has been blished in Reference 61.]*

1.1.2 Discussion l burnup is a measure of fuel depletion which represents the integrated energy output of the fuel egawatt-days per metric ton of uranium (MWD/MTU) and is a useful means for quantifying fuel osure criteria.

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

al excess reactivity installed in the fuel, although not a design basis, must be sufficient to maintain criticality at full-power operating conditions throughout cycle life with equilibrium xenon, arium, and other fission products present. Burnable absorbers, control rod insertion, and/or mical shim are used to compensate for the excess reactivity. The end of design cycle life is ned to occur when the chemical shim concentration is essentially zero with control rods present to degree necessary for operational requirements. In terms of soluble boron concentration, this esponds to approximately 10 ppm with the control and gray rods essentially withdrawn.

1.2 Negative Reactivity Feedbacks (Reactivity Coefficients) 1.2.1 Basis the initial fuel cycle, the fuel temperature coefficient will be negative, and the moderator perature coefficient of reactivity will be negative for power operating conditions, thereby providing ative reactivity feedback characteristics. The design basis meets General Design Criterion 11.

1.2.2 Discussion en compensation for a rapid increase in reactivity is considered, there are two major effects.

se are the resonance absorption (Doppler) effects associated with changing fuel temperature and neutron spectrum and reactor composition change effects resulting from changing moderator sity. These basic physics characteristics are often identified by reactivity coefficients. The use of htly enriched uranium results in a Doppler coefficient of reactivity that is negative. This coefficient ides the most rapid reactivity compensation. The initial core is also designed to have an overall ative moderator temperature coefficient of reactivity during power operation so that average lant temperature changes or void content provides another, slower compensatory effect. For e core designs, if the compensation for excess reactivity is provided only by chemical shim, the erator temperature coefficient could become positive. Nominal power operation is permitted only range of overall negative moderator temperature coefficient. The negative moderator perature coefficient can be achieved through the use of discrete burnable absorbers (BAs) and/or Staff approval is required prior to implementing a change in this information.

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nable absorber content (quantity and distribution) is not stated as a design basis. However, for e reloads, the use of burnable absorbers may be necessary for power distribution control and/or chieve an acceptable moderator temperature coefficient throughout core life. The required nable absorber loading is that which is required to meet design criteria.

1.3 Control of Power Distribution 1.3.1 Basis nuclear design basis is that, with at least a 95 percent confidence level:

The fuel will not operate with a power distribution that would result in exceeding the departure from nucleate boiling (DNB) design basis (i.e., the departure from nucleate boiling ratio (DNBR) shall be greater than the design limit departure from nucleate boiling ratio as discussed in Subsection 4.4.1) under Condition I and II occurrences, including the maximum overpower condition.

Under abnormal conditions, including the maximum overpower condition, the peak linear heat rate (PLHR) will not cause fuel melting, as defined in Subsection 4.4.1.2.

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 fuel will not be operated at Peak Linear Heat Rate (PLHR) values greater than those found to be acceptable within the body of the safety analysis under normal operating conditions, including an allowance of one percent for calorimetric error (calorimetric uncertainty calculation will be provided per Subsection 15.0.15.1).

above basis meets General Design Criterion 10.

1.3.2 Discussion culation of extreme power shapes which affect fuel design limits are performed with proven hods. The conditions under which limiting power shapes are assumed to occur are chosen servatively with regard to any permissible operating state. Even though there is close agreement ween calculated peak power and measurements, a nuclear uncertainty is applied bsection 4.3.2.2.1) to calculated power distribution. Such margins are provided both for the lysis for normal operating states and for anticipated transients.

1.4 Maximum Controlled Reactivity Insertion Rate 1.4.1 Basis maximum reactivity insertion rate due to withdrawal of rod cluster control assemblies (RCCAs) or rod cluster assemblies (GRCAs) or by boron dilution is limited by plant design, hardware, and ic physics. During normal power operation, the maximum controlled reactivity insertion rate is ed. The maximum reactivity change rate for accidental withdrawal of two control banks is set h that PLHR and the departure from nucleate boiling ratio limitations are not challenged. This sfies General Design Criterion 25.

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jection accident. (See Chapter 15).

owing any Condition IV occurrence, such as rod ejection or steam line break, the reactor can be ught to the shutdown condition, and the core maintains acceptable heat transfer geometry. This sfies General Design Criterion 28.

1.4.2 Discussion ctivity addition associated with an accidental withdrawal of a control bank (or banks) is limited by maximum rod speed (or travel rate) and by the worth of the bank(s). For this reactor, the imum control and gray rod speed is 45 inches per minute.

reactivity change rates are conservatively calculated, assuming unfavorable axial power and on distributions. The typical peak xenon burnout rate is significantly lower than the maximum tivity addition rate for normal operation and for accidental withdrawal of two banks.

1.5 Shutdown Margins 1.5.1 Basis imum shutdown margin as specified in the technical specifications is required in all operating es.

nalyses involving reactor trip, the single, highest worth rod cluster control assembly is postulated emain untripped in its full-out position (stuck rod criterion). This satisfies General Design erion 26.

1.5.2 Discussion independent reactivity control systems are provided: control rods and soluble boron in the lant. The control rods provide reactivity changes which compensate for the reactivity effects of the and water density changes accompanying power level changes over the range from full load to oad. The control rods provide the minimum shutdown margin under Condition I occurrences and capable of making the core subcritical rapidly enough to prevent exceeding acceptable fuel age limits (very small number of rod failures), assuming that the highest worth control rod is k out upon trip.

boron system can compensate for xenon burnout reactivity changes and maintain the reactor in cold shutdown condition. Thus, backup and emergency shutdown provisions are provided by hanical and chemical shim control systems which satisfy General Design Criterion 26. Reactivity nges due to fuel depletion are accommodated with the boron system.

1.5.3 Basis en fuel assemblies are in the pressure vessel and the vessel head is not in place, keff will be ntained at or below 0.95 with control rods and soluble boron. Further, the fuel will be maintained ciently subcritical that removal of the rod cluster control assemblies will not result in criticality.

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ipment flooded with pure water and a keff not to exceed 0.98 in normally dry new fuel storage s, assuming optimum moderation. No criterion is given for the refueling operation. However, a percent margin, which is consistent with spent fuel storage and transfer and the new fuel storage, dequate for the controlled and continuously monitored operations involved.

boron concentration required to meet the refueling shutdown criteria is specified in the Core rating Limits Report (COLR). Verification that these shutdown criteria are met, including ertainties, is achieved using standard design methods. The subcriticality of the core is tinuously monitored as described in the technical specifications.

1.6 Stability 1.6.1 Basis core will be inherently stable to power oscillations at the fundamental mode. This satisfies eral Design Criterion 12.

tial power oscillations within the core with a constant core power output, should they occur, can eliably and readily detected and suppressed.

1.6.2 Discussion illations of the total power output of the core, from whatever cause, are readily detected by the temperature sensors and by the nuclear instrumentation. The core is protected by these ems; a reactor trip occurs if power increases unacceptably, thereby preserving the design gins to fuel design limits. The combined stability of the turbine, steam generator and the reactor er control systems are such that total core power oscillations are not normally possible. The undancy of the protection circuits results in a low probability of exceeding design power levels.

core is designed so that diametral and azimuthal oscillations due to spatial xenon effects are

-damping; no operator action or control action is required to suppress them. The stability to metral oscillations is so great that this excitation is highly improbable. Convergent azimuthal llations can be excited by prohibited motion of individual control rods.

cations of power distribution anomalies are continuously available from an online core monitoring em. The online monitoring system processes information provided by the fixed in-core detectors, ore thermocouples, and loop temperature measurements. Radial power distributions are efore continuously monitored, thus power oscillations are readily observable and alarmed. The ore long ion chambers also provide surveillance and alarms of anomalous power distributions. In posed core designs, these horizontal plane oscillations are self-damping by virtue of reactivity back effects inherent to the basic core physics.

l xenon spatial power oscillations may occur during core life, especially late in the cycle. The ne core monitoring system provides continuous surveillance of the axial power distributions. The trol rod system provides both manual and automatic control systems for controlling the axial er distributions.

fidence that fuel design limits are not exceeded is provided by reactor protection system rpower T (OPT) and overtemperature T (OTT) trip functions, which use the loop perature sensors, pressurizer pressure indication, and measured axial offset as an input.

ection and suppression of xenon oscillations are discussed in Subsection 4.3.2.7.

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protection monitoring system. The diverse reactor trip reduces the probability and consequences postulated ATWS. The effects of anticipated transients with failure to trip are not considered in design bases of the plant. Analysis has shown that the likelihood of such a hypothetical event is ligibly small. Furthermore, analysis of the consequences of a hypothetical failure to trip following cipated transients has shown that no significant core damage would result, system peak sures should be limited to acceptable values, and no failure of the reactor coolant system would lt. (See WCAP-8330, Reference 5). The process used to evaluate the ATWS risk in compliance 10 CFR 50.62 is described in Section 15.8.

2 Description 2.1 Nuclear Design Description reactor core consists of a specified number of fuel rods held in bundles by spacer grids and top bottom fittings. The fuel rods are fabricated from cylindrical tubes made of zirconium based y(s) containing uranium dioxide fuel pellets. The bundles, known as fuel assemblies, are arranged pattern which approximates a right circular cylinder.

h fuel assembly contains a 17 x 17 rod array composed nominally of 264 fuel rods, 24 rod cluster trol thimbles, and an in-core instrumentation thimble. Figure 4.2-1 shows a cross-sectional view 17 x 17 fuel assembly and the related rod cluster control guide thimble locations. Detailed criptions of the AP1000 fuel assembly design features are given in Section 4.2.

initial core loading, the fuel rods within a given assembly have the same uranium enrichment in the radial and axial planes. Fuel assemblies of three different enrichments are used in the initial loading to establish a favorable radial power distribution. Figure 4.3-1 shows the fuel loading ern used in the initial cycle. Two regions consisting of the two lower enrichments are interspersed rm a checkerboard pattern in the central portion of the core. The third region is arranged around periphery of the core and contains the highest enrichment. The enrichments for the initial cycle shown in Table 4.3-1. Axial blankets consisting of fuel pellets of reduced enrichment placed at the s of the enriched pellet stack have been considered and may be used in reload cycles. Axial kets are included in the design basis to reduce neutron leakage and to improve fuel utilization.

oad core loading patterns can employ various fuel management techniques including low-age designs where the feed fuel is interspersed checkerboard-style in the core interior and leted fuel is placed on the periphery. Reload core designs, as well as the initial cycle design, are cipated to operate approximately 18 months between refueling, accumulating a cycle burnup of roximately 21,000 MWD/MTU. The exact reloading pattern, the initial and final positions of emblies, and the number of fresh assemblies and their placement are dependent on the energy uirement for the reload cycle and burnup and power histories of the previous cycles.

core average enrichment is determined by the amount of fissionable material required to provide desired energy requirements. The physics of the burnout process is such that operation of the tor depletes the amount of fuel available due to the absorption of neutrons by the U-235 atoms their subsequent fission. In addition, the fission process results in the formation of fission ducts, some of which readily absorb neutrons. These effects, the depletion and the buildup of on products, are partially offset by the buildup of plutonium shown in Figure 4.3-2 for a typical 17 fuel assembly, which occurs due to the parasitic absorption of neutrons in U-238. Therefore, e beginning of any cycle a reactivity reserve equal to the depletion of the fissionable fuel and the dup of fission product poisons less the buildup of fissile fuel over the specified cycle life is built the reactor. This excess reactivity is controlled by removable neutron-absorbing material in the 4.3-6 Revision 1

rmined on a design specific basis. Figure 4.3-3 is a plot of the initial core soluble boron centration versus core depletion.

concentration of the soluble neutron absorber is varied to compensate for reactivity changes due el burnup, fission product poisoning including xenon and samarium, burnable absorber letion, and the cold-to-operating moderator temperature change. Throughout the operating ge, the CVS is designed to provide changes in reactor coolant system (RCS) boron concentration ompensate for the reactivity effects of fuel depletion, peak xenon burnout and decay, and cold tdown boration requirements.

nable absorbers are strategically located to provide a favorable radial power distribution and ide for negative reactivity feedback. Figures 4.3-4a and 4.3-4b show the burnable absorber ributions within a fuel assembly for the several patterns used in a 17 x 17 array. The initial core nable absorber loading pattern is shown in Figure 4.3-5.

les 4.3-1 through 4.3-3 contain summaries of reactor core design parameters including reactivity fficients, delayed neutron fraction, and neutron lifetimes. Sufficient information is included to mit an independent calculation of the nuclear performance characteristics of the core.

2.2 Power Distribution accuracy of power distribution calculations has been confirmed through approximately 1000 flux s under conditions very similar to those expected. Details of this confirmation are given in AP-7308-L-P-A (Reference 7) and in Subsection 4.3.2.2.7.

2.2.1 Definitions ative power distributions within the reactor are quantified in terms of hot channel factors. These channel factors are normalized ratios of maximal absolute power generation rates and are a sure of the peak pellet power within the reactor core relative to the average pellet (FQ) and the rgy produced in a coolant channel relative to the core average channel (FH). Absolute power eration rates are expressed in terms of quantities related to the nuclear or thermal design; more cifically, volumetric power density (qvol) is the thermal power produced per unit volume of the (kW/liter).

ear heat rate (LHR) is the thermal power produced per unit length of active fuel (kW/ft). Since fuel embly geometry is standardized, LHR is the unit of absolute power density most commonly used.

practical purposes, LHR differs from qvol by a constant factor which includes geometry effects the heat flux deposition fraction. The peak linear heat rate (PLHR) is defined as the maximum ar heat rate occurring throughout the reactor. PLHR directly impacts fuel temperatures and decay er levels thus being a significant safety analysis parameter.

rage linear heat rate (ALHR) is the total thermal power produced in the fuel rods expressed as t flux divided by the total active fuel length of the rods in the core.

al heat flux is the heat flux at the surface of the cladding (Btu/hr-ft2). For nominal rod ameters, this differs from linear heat rate by a constant factor.

power is the total power generated in one rod (kW).

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hot channel factors used in the discussion of power distributions in this section are defined as ws:

heat flux hot channel factor, is defined as the maximum local heat flux on the surface of a fuel divided by the average fuel rod heat flux, allowing for manufacturing tolerances on fuel pellets rods.

, nuclear heat flux hot channel factor, is defined as the maximum local fuel rod linear heat rate ded by the average fuel rod linear heat rate, assuming nominal fuel pellet and rod parameters.

, engineering heat flux hot channel factor, is the allowance on heat flux required for ufacturing tolerances. The engineering factor allows for local variations in enrichment, pellet sity and diameter, burnable absorber content, surface area of the fuel rod, and eccentricity of the between pellet and clad. Combined statistically, the net effect is a factor of 1.03 to be applied to fuel rod surface heat flux.

H , nuclear enthalpy rise hot channel factor, is defined as the ratio of the maximum integrated power within the core to the average rod power.

ufacturing tolerances, hot channel power distribution, and surrounding channel power ributions are treated explicitly in the calculation of the departure from nucleate boiling ratio cribed in Section 4.4.

convenient for the purposes of discussion to define subfactors of FQ . However, design limits are in terms of the total peaking factor.

PLHR

= total peaking factor or heat flux hot channel factor = ALHR hout densification effects:

FQ = FQN x FQ E= N x Nx Nx E FXY FZ FU FQ N E re F Q and F Q are defined above and:

= factor for calculational uncertainty, assumed to be 1.05.

= ratio of peak power density to average power density in the horizontal plane of peak local power.

= ratio of the power per unit core height in the horizontal plane of peak local power to the average value of power per unit core height. If the plane of peak local power coincides with the plane of maximum power per unit core height, then FNZ is the core average axial peaking factor.

2.2.2 Radial Power Distributions power shape in horizontal sections of the core at full power is a function of the fuel assembly and nable absorber loading patterns, the control rod pattern, and the fuel burnup distribution. Thus, at time in the cycle, a horizontal section of the core can be characterized as unrodded or with trol rods. These two situations combined with burnup effects determine the radial power shapes 4.3-8 Revision 1

considered, but these are quite small. The effect of nonuniform flow distribution is negligible.

le radial power distributions in various planes of the core are often illustrated, since the moderator sity is directly proportional to enthalpy, the core radial enthalpy rise distribution, as determined by integral of power up each channel, is of greater interest. Figures 4.3-6 through 4.3-11 show cal normalized power density distributions for one-eighth of the core for representative operating ditions. These conditions are as follows:

Hot full power (HFP) near beginning of life, unrodded, no xenon Hot full power near beginning of life, unrodded, equilibrium xenon Hot full power near beginning of life, gray bank MA+MB in, equilibrium xenon Hot full power near middle of life (MOL), unrodded equilibrium xenon Hot full power near end of life, unrodded, equilibrium xenon Hot full power near end of life, gray bank MA+MB in, equilibrium xenon e the position of the hot channel varies from time to time, a single-reference radial design power ribution is selected for departure from nucleate boiling calculations. This reference power ribution is chosen conservatively to concentrate power in one area of the core, minimizing the efits of flow redistribution. Assembly powers are normalized to core average power. The radial er distribution within a fuel rod and its variation with burnup as utilized in thermal calculations and rod design are discussed in Section 4.4.

2.2.3 Assembly Power Distributions the purpose of illustration, typical rodwise power distributions from the beginning of life and end e conditions corresponding to Figures 4.3-7 and 4.3-10, respectively, are given for the same embly in Figures 4.3-12 and 4.3-13, respectively.

e the detailed power distribution surrounding the hot channel varies from time to time, a servatively flat radial assembly power distribution is assumed in the departure from nucleate ng analysis, described in Section 4.4, with the rod of maximum integrated power artificially raised e design value of FNH . Care is taken in the nuclear design of the fuel cycles and operating ditions to confirm that a flatter assembly power distribution does not occur with limiting values of 2.2.4 Axial Power Distributions distribution of power in the axial or vertical direction is largely under the control of the operator ugh either the manual operation of the control rods or the automatic motion of control rods in junction with manual operation of the chemical and volume control system. The automated mode peration is referred to as mechanical shim (MSHIM) and is discussed in Subsection 4.3.2.4.16.

rod control system automatically modulates the insertion of the axial offset (AO) control bank trolling the axial power distribution simultaneous with the MSHIM gray and control rod banks to ntain programmed coolant temperature. Operation of the chemical and volume control system is ated manually by the operator to compensate for fuel burnup and maintain the desired MSHIM k insertion. Nuclear effects which cause variations in the axial power shape include moderator sity, Doppler effect on resonance absorption, spatial distribution of xenon, burnup, and axial 4.3-9 Revision 1

online core monitoring system provides the operator with detailed power distribution information oth the radial and axial sense continuously using signals from the fixed in-core detectors. Signals also available to the operator from the ex-core ion chambers, which are long ion chambers ide the reactor vessel running parallel to the axis of the core. Separate signals are taken from the h ion chamber. The ion chamber signals are processed and calibrated against in-core surements such that an indication of the power in the top of the core less the power in the bottom e core is derived. The calibrated difference in power between the core top and bottom halves, ed the flux difference ( I ) , is derived for each of the four channels of ex-core detectors and is layed on the control panel. The principal use of the flux difference is to provide the shape penalty tion to the OTT DNB protection and the OPT overpower protection.

2.2.5 Local Power Peaking l densification occurred early in the evolution of pressurized water reactor fuel manufacture under diation in several operating reactors. This caused the fuel pellets to shrink both axially and ally. The pellet shrinkage combined with random hang-up of fuel pellets can result in gaps in the column when the pellets below the hung-up pellet settle in the fuel rod. The gaps vary in length location in the fuel rod. Because of decreased neutron absorption in the vicinity of the gap, power king occurs in the adjacent fuel rods, resulting in an increased power peaking factor. A ntitative measure of this local peaking is given by the power spike factor S(Z), where Z is the axial tion in the core. The power spike factor S(z) is discussed in References 8, 9, and 10.

ern PWR fuel manufacturing practices have essentially eliminated significant fuel densification acts on reactor design and operation. It has since been concluded and accepted that a sification power spike factor of 1.0 is appropriate for Westinghouse fuel as described in AP-13589-A (Reference 59).

2.2.6 Limiting Power Distributions ording to the ANSI classification of plant conditions (Chapter 15), Condition I occurrences are e expected frequently or regularly in the course of power operation, maintenance, or euvering of the plant. As such, Condition I occurrences are accommodated with margin between plant parameter and the value of that parameter which would require either automatic or manual ective action. Condition I occurrences are considered from the point of view of affecting the sequences of fault conditions (Conditions II, III, and IV). Analysis of each fault condition described ased on a conservative set of corresponding initial conditions.

list of steady-state and shutdown conditions, permissible deviations, and operational transients ven in Chapter 15. Implicit in the definition of normal operation is proper and timely action by the tor operator; that is, the operator follows recommended operating procedures for maintaining ropriate power distributions and takes any necessary remedial actions when alerted to do so by plant instrumentation.

online monitoring system evaluates the consequences of limiting power distributions based upon conditions prevalent in the reactor at the current time. Operating space evaluations performed by online monitoring system include the most limiting power distributions that can be generated by propriate operator or control system actions given the current core power level, xenon ribution, MSHIM or AO bank insertion and core burnup. Thus, as stated, the worst or limiting er distribution which can occur during normal operation is considered as the starting point for lysis of Conditions II, III, and IV occurrences.

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e power distributions which deviate from the normal operating condition within the allowable rating space as defined in the core operating limits; e.g., due to lack of proper action by the rator during a xenon transient following a change in power level brought about by control rod ion. Power distributions which fall in this category are used for determination of the reactor ection system setpoints to maintain margin to overpower or departure from nucleate boiling limits.

means for maintaining power distributions within the required absolute power generation limits described in the technical specifications. The online core monitoring system provides the rator with the current allowable operating space, detailed current power distribution information, mal margin assessment and operational recommendations to manage and maintain required mal margins. As such, the online monitoring system provides the primary means of managing maintaining required operating thermal margins during normal operation.

e unlikely event that the online monitoring system is out of service, power distribution controls ed on bounding, precalculated analysis are also provided to the operator such that the online itoring system is not a required element for short term reactor operation. Limits are placed on the l flux difference so that the heat flux hot channel factor FQ is maintained within acceptable limits.

scussion of precalculated power distribution control in Westinghouse pressurized water reactors Rs) is included in WCAP-7811 (Reference 11). Detailed background information on the design straints on local power density in a Westinghouse PWR, on the defined operating procedures, on the measures taken to preclude exceeding design limits is presented in the Westinghouse cal report on power distribution control and load following procedures WCAP-8385 ference 12). The following paragraphs summarize these reports and describe the calculations d to establish the upper bound on peaking factors.

N calculations used to establish the upper bound on peaking factors, FQ and F H , include the lear effects which influence the radial and axial power distributions throughout core life for various es of operation, including load follow, reduced power operation, and axial xenon transients.

er distributions are calculated for the full-power condition. Fuel and moderator temperature back effects are included within these calculations in each spatial dimension. The steady-state lear design calculations are done for normal flow with the same mass flow in each channel and redistribution effects neglected. The effect of flow redistribution is calculated explicitly where it is ortant in the departure from nucleate boiling analysis of accidents. The effect of xenon on radial er distribution is small (compare Figures 4.3-6 and 4.3-7) but is included as part of the normal ign process.

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

Core power level Core height Coolant temperature and flow Coolant temperature program as a function of reactor power 4.3-11 Revision 1

Rod bank overlaps mal operation of the plant assumes compliance with the following conditions:

Control rods in a single bank move together with no individual rod insertion differing from the bank demand position by more than the number of steps identified in the technical specifications.

Control banks are sequenced with overlapping banks.

The control bank insertion limits are not violated.

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

axial power distribution procedures referred to above are part of the required operating edures followed in normal operation with the online monitoring system out of service. In service, online core monitoring system provides continuous indication of power distribution, shutdown gin, and margin to design limits.

relaxed axial offset control (RAOC) procedures described in WCAP-10216-P-A (Reference 13) e developed to provide wide control band widths and consequently, more operating flexibility.

se wide operating limits, particularly at lower power levels, increase plant availability by allowing ker plant startup and increased maneuvering flexibility without trip. This procedure has been ified to accommodate AP1000 MSHIM operation. It is applied to analysis of axial power ributions under MSHIM control for the purpose of defining the allowed normal operating space h that Condition I thermal margin limits are maintained and Condition II occurrences are quately protected by the reactor protection system when the online monitoring system is out of ice.

purpose of this analysis is to find the widest permissible I versus power operating space by lyzing a wide range of achievable xenon distributions, MSHIM/AO bank insertion, and power l.

bounding analyses performed off line in anticipation of the online monitoring system being out of ice is similar to that based on the relaxed axial offset control analysis, which uses a xenon nstruction model described in WCAP-10216-P-A (Reference 13). This is a practical method ch is used to define the power operating space allowed with AP1000 MSHIM operation. Each lting power shape is analyzed to determine if loss-of-coolant accident constraints are met or eeded.

online monitoring system evaluates the effects of radial xenon distribution changes due to rational parameter changes continuously and therefore eliminates the need for overly servative bounding evaluations when the online monitoring system is available. A detailed ussion of this effect may be found in WCAP-8385 (Reference 12). The calculated values have n increased by a factor of 1.05 for method uncertainty and a factor of 1.03 for the engineering or FEQ .

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online monitoring system measures the core condition continuously and evaluates the thermal gin condition directly in terms of peak linear heat rate and margin to departure from nucleate ng limitations directly.

wing for fuel densification effects, the average linear power at 3400 MW is 5.72 kW/ft. From re 4.3-14, the conservative upper bound value of normalized local power density, including ertainty allowances, is 2.60 corresponding to a peak linear heat rate of 15.0 kW/ft at each core ation at 101 percent power.

etermine reactor protection system setpoints with respect to power distributions, three categories vents are considered: rod control equipment malfunctions and operator errors of commission or ssion. In evaluating these three categories of events, the core is assumed to be operating within four constraints described above.

first category comprises uncontrolled rod withdrawal (with rods moving in the normal bank uence) for both AO and MSHIM banks. Also included are motions of the AO and MSHIM banks w their insertion limits, which could be caused, for example, by uncontrolled dilution or primary lant cooldown. Power distributions are calculated throughout these occurrences, assuming short-corrective action; that is, no transient xenon effects are considered to result from the function. The event is assumed to occur from typical normal operating situations, which include mal xenon transients. It is further assumed in determining the power distributions that total core er level would be limited by reactor trip to below the overpower protection setpoint of nominally percent rated thermal power. Since the study is to determine protection limits with respect to er and axial offset, no credit is taken for OTT or OPT trip setpoint reduction due to flux rence. The peak power density which can occur in such events, assuming reactor trip at or below percent, is less than that required for fuel centerline melt, including uncertainties and sification effects.

second category assumes that the operator mispositions the AO and/or MSHIM rod banks in ation of the insertion limits and creates short-term conditions not included in normal operating ditions.

third category assumes that the operator fails to take action to correct a power distribution limit ation (such as boration/dilution transient) assuming automatic operation of the rod control system ch will maintain constant reactor power.

each of the above categories, the trip setpoints are designed so as not to exceed fuel centerline t criteria as well as fuel mechanical design criteria.

appropriate hot channel factors FQ and FNH for peak local power density and for DNB analysis ll power are based on analyses of possible operating power shapes and are addressed in the nical specifications.

maximum allowable FQ can be increased with decreasing power, as shown in the technical cifications. Increasing F N with decreasing power is permitted by the DNB protection setpoints H

allows radial power shape changes with rod insertion to the insertion limits, as described in section 4.4.4.3. The allowance for increased FNH permitted is addressed in the technical cifications.

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urring in normal operation are used in verifying that this criterion is met. The worst values erally occur when the rods are assumed to be at their insertion limits. Operation with rod positions ve the allowed rod insertion limits provides increased margin to the FNH criterion. As discussed in tion 3.2 of WCAP-7912-P-A (Reference 14), it has been determined that the technical cifications limits are met, provided the above conditions are observed. These limits are taken as t to the thermal-hydraulic design basis, as described in Subsection 4.4.4.3.1.

en a situation is possible in normal operation which could result in local power densities in excess ose assumed as the precondition for a subsequent hypothetical accident, but which would not f cause fuel failure, administrative controls and alarms are provided for returning the core to a condition. These alarms are described in Chapter 7.

independence of the various individual uncertainties constituting the uncertainty factor on FQ bles the uncertainty ( FUQ ) to be calculated by statistically combining the individual uncertainties he limiting rod. The standard deviation of the resultant distribution of FUQ is determined by taking square root of the sum of the variances of each of the contributing distributions AP-7308-L-P-A (Reference 7). The values for FEQ and FNU are 1.03 and 1.05, respectively. The e for the rod bow factor, FBQ , is 1.056, which accounts for the maximum FQ penalty as a function urnup due to rod bow effects.

2.2.7 Experimental Verification of Power Distribution Analysis subject is discussed in WCAP-7308-L-P-A (Reference 7) and WCAP-12472-P-A (Reference 4).

mmary of these reports and the extension to include the fixed in-core instrumentation system is n below. Power distribution related measurements are incorporated into the evaluation of ulated power distribution information using the in-core instrumentation processing algorithms tained within the online monitoring system. The processing algorithms contained within the online itoring system are functionally identical to those historically used for the evaluation of power ribution measurements in Westinghouse PWRs. Advances in technology allow a complete tional integration of reaction rate measurement algorithms and the expected reaction rate dictive capability within the same software package. The predictive software integrated within the ne monitoring system supplies accurate, detailed information of current reactor conditions. The orical algorithms are described in detail in WCAP-12472-P-A (Reference 4).

measured versus calculational comparison is performed continuously by the online monitoring em throughout the core life. The online monitoring system operability requirements are specified e technical specifications.

measurement of the reactor power distribution and the associated thermal margin limiting ameters, with the in-core instrumentation system described in Subsections 7.7.1 and 4.4.6, the wing uncertainties must be considered:

Reproducibility of the measured signal Errors in the calculated relationship between detector current and local power generation within the fuel bundle Errors in the detector current associated with the depletion of the emitter material, manufacturing tolerances and measured detector depletion 4.3-14 Revision 1

appropriate allowance for category A has been accounted for through the imposition of strict ufacturing tolerances for the individual detectors. This approach is accepted industry practice has been used in PWRs with fixed in-core instrumentation worldwide. Errors in category B above quantified by calculation and evaluation of critical experiment data on arrays of rods with ulated guide thimbles, control rods, burnable absorbers, etc. These critical experiments provide quantification of errors of categories A and D above. Errors in category C have been quantified ugh direct experimental measurement of the depletion characteristics of the detectors being used uding the precision of the in-core instrumentation systems measurement of the current detector letion. The description of the experimental measurement of detector depletion can be found in I-NP-3814 (Reference 16).

AP-7308-L-P-A (Reference 7) describes critical experiments performed at the Westinghouse ctor Evaluation Center and measurements taken on two Westinghouse plants with movable on chamber in-core instrumentation systems. The measurement aspects of the movable fission mber share the previous uncertainty categories less category C which is independent of the other rces of uncertainty. WCAP-7308-L-P-A (Reference 7) concludes that the uncertainty associated peak linear heat rate (FQ*P) is less than five percent at the 95 percent confidence level with only percent of the measurements greater than the inferred value.

omparing measured power distributions (or detector currents) with calculations for the same rating conditions, it is not possible to isolate the detector reproducibility. Thus, a comparison ween measured and predicted power distributions includes some measurement error. Such a parison is given in Figure 4.3-15 for one of the maps used in WCAP-7308-L-P-A (Reference 7).

e the first publication of WCAP-7308-L-P-A, hundreds of measurements have been taken on tors all over the world. These results confirm the adequacy of the five percent uncertainty wance on the calculated peak linear heat rate (ALHR*FQ*P).

milar analysis for the uncertainty in hot rod integrated power FH*P measurements results in an wance of four percent at the equivalent of a 95 percent confidence level.

easurement in the fourth cycle of a 157-assembly, 12-foot core is compared with a simplified one-ensional core average axial calculation in Figure 4.3-16. This calculation does not give explicit esentation to the fuel grids.

accumulated data on power distributions in actual operation are basically of three types:

Much of the data is obtained in steady-state operation at constant power in the normal operating configuration.

Data with unusual values of axial offset are obtained as part of the ex-core detector calibration exercise performed monthly.

Special tests have been performed in load follow and other transient xenon conditions which have yielded useful information on power distributions.

se data are presented in detail in WCAP-7912-P-A (Reference 14). Figure 4.3-17 contains a mary of measured values of FQ as a function of axial offset for five plants from that report.

4.3-15 Revision 1

sfactory results are described in Chapter 14. Since not all limiting situations can be created at inning of life, the main purpose of the tests is to provide a check on the calculational methods d in the predictions for the conditions of the test. Tests performed at the beginning of each reload e are limited to verification of the selected safety-related parameters of the reload design.

2.2.9 Monitoring Instrumentation adequacy of instrument numbers, spatial deployment, required correlations between readings peaking factors, calibration, and errors are described in WCAP-12472-P (Reference 4). The vant conclusions are summarized in Subsection 4.3.2.2.7 and Subsection 4.4.6.

vided the limitations given in Subsection 4.3.2.2.6 on rod insertion and flux difference are erved, the in-core and ex-core detector systems provide adequate monitoring of power ributions when the online monitoring system is out of service. Further details of specific limits on observed rod positions and flux difference are given in the technical specifications, together with scussion of their bases.

its for alarms and reactor trip are given in the technical specifications. Descriptions of the systems ided are given in Section 7.7.

2.3 Reactivity Coefficients kinetic characteristics of the reactor core determine the response of the core to changing plant ditions or to operator adjustments made during normal operation, as well as the core response ng abnormal or accidental transients. These kinetic characteristics are quantified in reactivity fficients. The reactivity coefficients reflect the changes in the neutron multiplication due to varying t conditions, such as thermal power, moderator and fuel temperatures, coolant pressure, or void ditions, although the latter are relatively unimportant. Since reactivity coefficients change during life of the core, ranges of coefficients are employed in transient analysis to determine the onse of the plant throughout life. The results of such simulations and the reactivity coefficients d are presented in Chapter 15.

reactivity coefficients are calculated with approved nuclear methods. The effect of radial and l power distribution on core average reactivity coefficients is implicit in those calculations and is significant under normal operating conditions. For example, a skewed xenon distribution which lts in changing axial offset by five percent typically changes the moderator and Doppler perature coefficients by less than 0.01 pcm/°F. An artificially skewed xenon distribution which lts in changing the radial FNH by three percent typically changes the moderator and Doppler perature coefficients by less than 0.03 pcm/°F and 0.001 pcm/°F, respectively. The spatial effects accentuated in some transient conditions, for example, in postulated rupture of the main steam 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.

analytical methods and calculational models used in calculating the reactivity coefficients are n in Subsection 4.3.3. These models have been confirmed through extensive qualification efforts ormed for core and lattice designs.

ntitative information for calculated reactivity coefficients including fuel-Doppler coefficient, erator coefficients (density, temperature, pressure, and void), and power coefficient, is given in following sections.

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ctive fuel temperature and is primarily a measure of the Doppler broadening of U-238 and Pu-240 nance absorption peaks. Doppler broadening of other isotopes is also considered, but their tribution to the Doppler effect is small. An increase in fuel temperature increases the effective nance absorption cross sections of the fuel and produces a corresponding reduction in reactivity.

fuel temperature coefficient is calculated using approved nuclear methods. Moderator perature is held constant, and the power level is varied. Spatial variation of fuel temperature is n into account by calculating the effective fuel temperature as a function of power density, as ussed in Subsection 4.3.3.1.

pical Doppler temperature coefficient is shown in Figure 4.3-18 as a function of the effective fuel perature (at beginning of life and end of life conditions). The effective fuel temperature is lower the volume-averaged fuel temperature, since the neutron flux distribution is non-uniform through pellet and gives preferential weight to the surface temperature. A typical Doppler-only tribution to the power coefficient, defined later, is shown in Figure 4.3-19 as a function of relative power. The integral of the differential curve in Figure 4.3-19 is the Doppler contribution to the er defect and is shown in Figure 4.3-20 as a function of relative power. The Doppler temperature fficient becomes more negative as a function of life as the Pu-240 content increases, thus easing the Pu-240 resonance absorption. The upper and lower limits of Doppler coefficient used ccident analyses are given in Chapter 15.

2.3.2 Moderator Coefficients moderator coefficient is a measure of the change in reactivity due to a change in specific coolant ameters, such as density/temperature, pressure, or void. The coefficients obtained are moderator sity/temperature, pressure, and void coefficients.

2.3.2.1 Moderator Density and Temperature Coefficients moderator temperature (density) coefficient is defined as the change in reactivity per degree nge in the moderator temperature. Generally, the effects of the changes in moderator density and temperature are considered together.

soluble boron used in the reactor as a means of reactivity control also has an effect on the erator density coefficient, since the soluble boron density and the water density are decreased n the coolant temperature rises. A decrease in the soluble boron density introduces a positive ponent in the moderator coefficient. If the concentration of soluble boron is large enough, the net e of the coefficient may be positive.

initial core hot boron concentration is sufficiently low that the moderator temperature coefficient egative at operating temperatures with the burnable absorber loading specified. Discrete or gral fuel burnable absorbers can be used in reload cores to confirm the moderator temperature fficient is negative over the range of power operation. The effect of control rods is to make the erator coefficient more negative, since the thermal neutron mean free path, and hence the me affected by the control rods, increase with an increase in temperature.

h burnup, the moderator coefficient becomes more negative, primarily as a result of boric acid ion, but also to a significant extent from the effects of the buildup of plutonium and fission ducts.

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nge. The moderator temperature coefficient is shown as a function of core temperature and boron centration for the core in Figures 4.3-21 through 4.3-23. The temperature range covered is from

, about 70°F, to about 550°F. The contribution due to Doppler coefficient (because of change in erator temperature) has been subtracted from these results. Figure 4.3-24 shows the unrodded full-power moderator temperature coefficient plotted as a function of burnup for the initial cycle.

temperature coefficient corresponds to the unrodded critical boron concentration present at hot power operating conditions.

moderator coefficients presented here are calculated to describe the core behavior in normal accident situations when the moderator temperature changes can be considered to affect the re core.

2.3.2.2 Moderator Pressure Coefficient moderator pressure coefficient relates the change in moderator density, resulting from a reactor lant pressure change, to the corresponding effect on neutron production. This coefficient is of h less significance than the moderator temperature coefficient. A change of 50 psi in pressure approximately the same effect on reactivity as a one half degree change in moderator perature. This coefficient can be determined from the moderator temperature coefficient by ting change in pressure to the corresponding change in density. The typical moderator pressure fficient may be negative over a portion of the moderator temperature range at beginning of life L) (-0.004 pcm/psi) but is always positive at operating conditions and becomes more positive ng life (+0.3 pcm/psi, at end of life).

2.3.2.3 Moderator Void Coefficient moderator void coefficient relates the change in neutron multiplication to the presence of voids in moderator. In a PWR, this coefficient is not very significant because of the low void content in the lant. The core void content is less than one-half of one percent and is due to local or statistical ng. The typical void coefficient varies from 50 pcm/percent void at BOL and at low temperatures inus 250 pcm/percent void at EOL and at operating temperatures. The void coefficient at rating temperature becomes more negative with fuel burnup.

2.3.3 Power Coefficient combined effect of moderator temperature and fuel temperature change as the core power level nges is called the total power coefficient and is expressed in terms of reactivity change per ent power change. Since a three-dimensional calculation is performed in determining total power fficients and total power defects, the axial redistribution reactivity component described in section 4.3.2.4.3 is implicitly included. A typical power coefficient at beginning of life (BOL) and of life (EOL) conditions is given in Figure 4.3-25.

total power coefficient becomes more negative with burnup, reflecting the combined effect of erator and fuel temperature coefficients with burnup. The power defect (integral reactivity effect)

OL and EOL is given in Figure 4.3-26.

2.3.4 Comparison of Calculated and Experimental Reactivity Coefficients section 4.3.3 describes the comparison of calculated and experimental reactivity coefficients in il.

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2.3.5 Reactivity Coefficients Used in Transient Analysis le 4.3-2 gives the limiting values as well as the best-estimate values for the reactivity coefficients he initial cycle. The limiting values are used as design limits in the transient analysis. The exact es of the coefficient used in the analysis depend on whether the transient of interest is examined e BOL or EOL, whether the most negative or the most positive (least negative) coefficients are ropriate, and whether spatial non-uniformity must be considered in the analysis. Conservative es of coefficients, considering various aspects of analysis, are used in the transient analysis. This escribed in Chapter 15.

reactivity coefficients shown in Figures 4.3-18 through 4.3-26 are typical best-estimate values ulated for the initial cycle. Limiting values are chosen to encompass the best-estimate reactivity fficients, including the uncertainties given in Subsection 4.3.3.3 over appropriate operating ditions. The most positive, as well as the most negative, values are selected to form the design is range used in the transient analysis. A direct comparison of the best-estimate and design limit es for the initial cycle is shown in Table 4.3-2. In many instances the most conservative bination of reactivity coefficients is used in the transient analysis even though the extreme fficients assumed may not simultaneously occur at the conditions assumed in the analysis. The d for a reevaluation of any accident in a subsequent cycle is contingent upon whether the fficients for that cycle fall within the identified range used in the analysis presented in Chapter 15 due allowance for the calculational uncertainties given in Subsection 4.3.3.3. Control rod uirements are given in Table 4.3-3 for the initial cycle and for a hypothetical equilibrium cycle, e these are markedly different. These latter numbers are provided for information only.

2.4 Control Requirements stablish the required shutdown margin stated in the COLR under conditions where a cooldown to ient temperature is required, concentrated soluble boron is added to the coolant. Boron centrations for several core conditions are listed in Table 4.3-2 for the initial cycle. For core ditions including refueling, the boron concentration is well below the solubility limit. The rod ter control assemblies are employed to bring the reactor to the shutdown condition. The minimum uired shutdown margin is given in the COLR.

ability to accomplish the shutdown for hot conditions is demonstrated in Table 4.3-3 by paring the difference between the rod cluster control assembly reactivity available with an wance for the worst stuck rod with that required for control and protection purposes. The tdown margin includes an allowance of seven percent for analytic uncertainties which assumes use of silver-indium-cadmium rod cluster control assemblies. Use of a seven percent uncertainty wance on rod cluster control assembly worth is discussed and shown to be acceptable in AP-9217 (Reference 17). The largest reactivity control requirement appears at the EOL when the erator temperature coefficient reaches its peak negative value as reflected in the larger power ct.

control rods are required to provide sufficient reactivity to account for the power defect from full er to zero power and to provide the required shutdown margin. The reactivity addition resulting power reduction consists of contributions from Doppler effect, moderator temperature, flux stribution, and reduction in void content as discussed below.

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ncrease in effective pellet temperature. This effect is most noticeable over the range of zero er to full power due to the large pellet temperature increase with power generation.

2.4.2 Variable Average Moderator Temperature en the core is shut down to the hot zero-power condition, the average moderator temperature nges from the equilibrium full-load value determined by the steam generator and turbine racteristics (such as steam pressure, heat transfer, tube fouling) to the equilibrium no-load value, ch is based on the steam generator shell side design pressure. The design change in temperature onservatively increased to account for the control system dead band and measurement errors.

en the moderator coefficient is negative, there is a reactivity addition with power reduction. The erator coefficient becomes more negative as the fuel depletes because the boron concentration duced. This effect is the major contributor to the increased requirement at EOL.

2.4.3 Redistribution ing full-power operation, the coolant density decreases with core height. This, together with ial insertion of control rods, results in less fuel depletion near the top of the core. Under steady-e conditions, the relative power distribution will be slightly asymmetric toward the bottom of the

. On the other hand, at hot zero-power conditions, the coolant density is uniform up the core, and e is no flattening due to Doppler effect. The result will be a flux distribution which at zero power be skewed toward the top of the core. Since a three-dimensional calculation is performed in rmining total power defect, flux redistribution is implicitly included in this calculation. An itional redistribution allowance for adversely skewed xenon distributions is included in the rmination of the total control requirement specified in Table 4.3-3.

2.4.4 Void Content mall void content in the core is due to nucleate boiling at full power. The void collapse coincident power reduction makes a small positive reactivity contribution.

2.4.5 Rod Insertion Allowance ull power, the MSHIM and AO banks are operated within a prescribed band of travel to pensate for small changes in boron concentration, changes in temperature, and very small nges in the xenon concentration not compensated for by a change in boron concentration. When MSHIM banks reach a predetermined insertion or withdrawal, a change in boron concentration ld be required to compensate for additional reactivity changes. Use of soluble boron is limited to depletion and shutdown considerations. Since the insertion limit is set by rod travel limit, a servatively high calculation of the inserted worth is made, which exceeds the normally inserted tivity.

2.4.6 Installed Excess Reactivity for Depletion ess reactivity is installed at the beginning of each cycle to provide sufficient reactivity to pensate for fuel depletion and fission product buildup throughout the cycle. This reactivity is trolled by the addition of soluble boron to the coolant and by burnable absorbers when necessary.

soluble boron concentration for several core configurations and the unit boron worth are given in les 4.3-1 and 4.3-2 for the initial cycle. Since the excess reactivity for burnup is controlled by ble boron and/or burnable absorbers, it is not included in control rod requirements.

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wing rapid power level changes, that the resulting reactivity change can be controlled by nging the gray and/or control rod insertion. (Also see Subsection 4.3.2.4.16).

2.4.8 pH Effects nges in reactivity due to a change in coolant pH, if any, are sufficiently small in magnitude and ur slowly enough to be controlled by the boron system WCAP-3896-8 (Reference 18).

2.4.9 Experimental Confirmation owing a normal shutdown, the total core reactivity change during cooldown with a stuck rod has n measured on a 121-assembly, 10-foot-high core and a 121-assembly, 12-foot-high core. In h case, the core was allowed to cool down until it reached criticality simulating the steam line ak accident. For the 10-foot core, the total reactivity change associated with the cooldown is over dicted by about 0.3-percent with respect to the measured result. This represents an error of ut five percent in the total reactivity change and is about half the uncertainty allowance for this ntity. For the 12-foot core, the difference between the measured and predicted reactivity change n even smaller 0.2 percent . These measurements and others demonstrate the capability of the hods described in Subsection 4.3.3.

2.4.10 Control e reactivity is controlled by means of a chemical poison dissolved in the coolant, rod cluster trol assemblies, gray rod cluster assemblies and burnable absorbers as described below.

2.4.11 Chemical Shim on in solution as boric acid is used to control relatively slow reactivity changes associated with:

The moderator temperature defect in going from cold shutdown at ambient temperature to the hot operating temperature at zero power The transient xenon and samarium poisoning, such as that following power changes to levels below 30 percent rated thermal power The reactivity effects of fissile inventory depletion and buildup of long-life fission products The depletion of the burnable absorbers boron concentrations for various core conditions are presented in Table 4.3-2 for the initial cycle.

2.4.12 Rod Cluster Control Assemblies number of rod cluster control assemblies is shown in Table 4.3-1. The rod cluster control emblies are used for shutdown and control purposes to offset fast reactivity changes associated The required shutdown margin in the hot zero power, stuck rod condition The reactivity compensation as a result of an increase in power above hot zero power (power defect, including Doppler and moderator reactivity changes) 4.3-21 Revision 1

Reactivity changes resulting from load changes allowed control bank reactivity insertion is limited at full power to maintain shutdown capability.

he power level is reduced, control rod reactivity requirements are also reduced, and more rod rtion is allowed. The control bank position is monitored, and the operator is notified by an alarm if limit is approached. The determination of the insertion limit uses conservative xenon distributions axial power shapes. In addition, the rod cluster control assembly withdrawal pattern determined the analyses is used in determining power distribution factors and in determining the maximum th of an inserted rod cluster control assembly ejection accident. For further discussion, refer to the nical specifications on rod insertion limits.

er distribution, rod ejection, and rod misalignment analyses are based on the arrangement of the tdown and control groups of the rod cluster control assemblies shown in Figure 4.3-27. Shutdown cluster control assemblies are withdrawn before withdrawal of the control and AO banks is ated. The approach to critical is initiated by using the chemical and volume control system to blish an appropriate boron concentration based upon the estimated critical condition then drawing the AO bank above the zero power insertion limit and finally withdrawing the control ks sequentially. The limits of rod insertion and further discussion on the basis for rod insertion s are provided in the COLR and technical specifications.

2.4.13 Gray Rod Cluster Assemblies rod cluster control assembly control banks include four gray rod banks consisting of gray rod ter assemblies (GRCAs). Gray rod cluster assemblies consist of 24 rodlets fastened at the top to a common hub or spider. Geometrically, it is the same as a rod cluster control assembly ept that 12 of the 24 rodlets are comprised of stainless steel while the remaining 12 rodlets are uced diameter silver-indium-cadmium clad with stainless steel. The term gray rod refers to the uced reactivity worth relative to that of a rod cluster control assembly consisting of 24 silver-um-cadmium rodlets. The gray rod cluster assemblies are used in load follow maneuvering and ide a mechanical shim reactivity mechanism to eliminate the need for changes to the centration of soluble boron (that is, chemical shim).

2.4.14 Burnable Absorbers rete burnable absorber rods or integral fuel burnable absorber rods or both may be used to ide partial control of the excess reactivity available during the fuel cycle. In doing so, the nable absorber loading controls peaking factors and prevents the moderator temperature fficient from being positive at normal operating conditions. The burnable absorbers perform this tion by reducing the requirement for soluble boron in the moderator at the beginning of the fuel e, as described previously. For purposes of illustration, the initial cycle burnable absorber pattern hown in Figure 4.3-5. Figures 4.3-4a and 4.3-4b show the burnable absorber distribution within a assembly for several burnable absorber patterns used in the 17 x 17 array. The boron in the rods epleted with burnup but at a slow rate so that the peaking factor limits are not exceeded and the lting critical concentration of soluble boron is such that the moderator temperature coefficient ains within the limits stated above for power operating conditions.

2.4.15 Peak Xenon Startup pensation for the peak xenon buildup may be accomplished using the boron control system.

tup from the peak xenon condition is accomplished with a combination of rod motion and boron 4.3-22 Revision 1

2.4.16 Load Follow Control and Xenon Control ing load follow maneuvers, power changes are primarily accomplished using control rod motion e, as required. Control rod motion is limited by the control rod insertion limits as provided in the LR and discussed in Subsections 4.3.2.4.12 and 4.3.2.4.13. The power distribution is maintained in acceptable limits through limitations on control rod insertion. Reactivity changes due to the nging xenon concentration are also controlled by rod motion.

id power increases (five percent/min) from part power during load follow operation are omplished with rod motion.

rod control system is designed to automatically provide the power and temperature control cribed above 30 percent rated power for most of the cycle length without the need to change on concentration as a result of the load maneuver. The automated mode of operation is referred s mechanical shim (MSHIM) because of the usage of mechanical means to control reactivity and er distribution simultaneously. MSHIM operation allows load maneuvering without boron change ause of the degree of allowed insertion of the control banks in conjunction with the independent er distribution control of the axial offset (AO) control bank. The worth and overlap of the MA, MB, MD, M1, and M2 control banks are designed such that the AO control bank insertion will always lt in a monotonically decreasing axial offset. MSHIM operation uses the MA, MB, MC, MD, M1, M2 control banks to maintain the programmed coolant average temperature throughout the rating power range. The AO control bank is independently modulated by the rod control system to ntain a nearly constant axial offset throughout the operating power range. The degree of control insertion under MSHIM operation allows rapid return to power without the need to change boron centration.

target axial offset used during MSHIM load follow and base load operation is established at a e negative value than the axial offset associated with the all rods out of condition. The negative is necessary to maintain both positive and negative axial offset control effectiveness by the control bank. Extended base load operation is performed by controlling axial offset to the target e using the AO control bank, and by controlling the coolant average temperature to the grammed value with the M-banks. Boron concentration changes are made periodically as the fuel letes to reposition the M-banks and allow for a periodic exchange of the gray rod bank insertion uence. MSHIM load follow and base load operations (including the gray rod bank insertion uence exchanges) are considered Condition I normal operations.

2.4.17 Burnup trol of the excess reactivity for burnup is accomplished using soluble boron and/or burnable orbers. The boron concentration is limited during operating conditions to maintain the moderator perature coefficient within its specified limits. A sufficient burnable absorber loading is installed at beginning of a cycle to give the desired cycle lifetime, without exceeding the boron concentration

. The end of a fuel cycle is reached when the soluble boron concentration approaches the tical minimum boron concentration in the range of 0 to 10 ppm.

2.4.18 Rapid Power Reduction System reactor power control system is designed with the capability of responding to full load rejection out initiating a reactor trip using the normal rod control system, reactor control system, and the d power reduction system. Load rejections requiring greater than a fifty percent reduction of rated mal power initiate the rapid power reduction system. The rapid power reduction system utilizes 4.3-23 Revision 1

uction system is included in the cycle specific safety analysis and licensing process.

2.5 Control Rod Patterns and Reactivity Worth rod cluster control assemblies are designated by function as the control groups and the tdown groups. The terms group and bank are used synonymously to describe a particular uping of control assemblies. The rod cluster control assembly patterns are displayed in re 4.3-27. The control banks are labeled MA, MB, MC, MD, M1, M2, and AO with the MA, MB, and MD banks comprised of gray rod cluster assemblies; and the shutdown banks are labeled

, SD2, SD3, and SD4. Each bank of more than four rod cluster control assemblies, although rated and controlled as a unit, is composed of two or more subgroups. The axial position of the cluster control assemblies may be controlled manually or automatically. The rod cluster control emblies are dropped into the core following actuation of reactor trip signals.

criteria have been employed for selection of the control groups. First, the total reactivity worth t be adequate to meet the requirements specified in Table 4.3-3. Second, in view of the fact that e rods may be partially inserted at power operation, the total power peaking factor should be low ugh to meet the power capability requirements. Analyses indicate that the first requirement can met either by a single group or by two or more banks whose total worth equals at least the uired amount. The axial power shape is more peaked following movement of a single group of worth three to four percent . Therefore, control bank rod cluster control assemblies have n separated into several bank groupings. Typical control bank worth for the initial cycle are shown able 4.3-2.

position of control banks for criticality under any reactor condition is determined by the centration of boron in the coolant. On an approach to criticality, boron is adjusted so that criticality be achieved with control rods above the insertion limit set by shutdown and other considerations.

e the technical specifications and COLR). Early in the cycle, there may also be a withdrawal limit w power to maintain the moderator temperature coefficient within the specified limits for that er level.

ted rod worths for several different conditions are given in Subsection 15.4.8.

wable deviations due to misaligned control rods are discussed in the technical specifications.

presentative differential rod worth calculation for two banks of control rods withdrawn ultaneously (rod withdrawal accident) is given in Figure 4.3-28.

culation of control rod reactivity worth versus time following reactor trip involves both control rod city and differential reactivity worth. The rod position versus time of travel after rod release umed is given in Figure 4.3-29. For nuclear design purposes, the reactivity worth versus rod ition is calculated by a series of steady-state calculations at various control positions, assuming rods out of the core as the initial position in order to minimize the initial reactivity insertion rate.

, to be conservative, the rod of highest worth is assumed stuck out of the core, and the flux ribution (and thus reactivity importance) is assumed to be skewed to the bottom of the core. The lt of these calculations is shown in Figure 4.3-30.

shutdown groups provide additional negative reactivity to establish adequate shutdown margin.

tdown margin is the amount by which the core would be subcritical at hot shutdown if the rod ter control assemblies were tripped, but assuming that the highest worth assembly remained fully 4.3-24 Revision 1

values given in Table 4.3-3 show that the available reactivity in withdrawn rod cluster control emblies provides the design bases minimum shutdown margin, allowing for the highest worth ter to be at its fully withdrawn position. An allowance for the uncertainty in the calculated worth of rods is made before determination of the shutdown margin.

2.6 Criticality of the Reactor During Refueling basis for maintaining the reactor subcritical during refueling is presented in Subsection 4.3.1.5, a discussion of how control requirements are met is given in Subsections 4.3.2.4 and 4.3.2.5.

2.6.1 Criticality Design Method Outside the Reactor cality of fuel assemblies outside the reactor is precluded by adequate design of fuel transfer, ping, and storage facilities and by administrative control procedures. The two principal methods reventing criticality are limiting the fuel assembly array size and limiting assembly interaction by g the minimum separation between assemblies and/or inserting neutron poisons between emblies. The details of the methodology used for the new fuel rack and spent fuel rack criticality lysis are included in the Chapter 9.1 references.

design criteria are consistent with General Design Criterion (GDC) 62, Reference 19, and NRC ance given in Reference 20. The applicable 10 CFR Part 50.68 requirements are as follows:

The maximum K-effective value, including all biases and uncertainties, must be less than 0.95 with soluble boron credit and less than 1.0 with full density unborated water. Note this design criterion is provided in 10 CFR Part 50.68, Item 4 of Paragraph b. Note that the specific terminology is:

If no credit for soluble boron is taken, the k-effective of the spent fuel storage racks loaded with fuel of the maximum fuel assembly reactivity must not exceed 0.95, at a 95 percent probability, 95 percent confidence level, if flooded with unborated water. If credit is taken for soluble boron, the k-effective of the spent fuel storage racks loaded with fuel of the maximum fuel assembly reactivity must not exceed 0.95, at a 95 percent probability, 95 percent confidence level, if flooded with borated water, and the k-effective must remain below 1.0 (subcritical), at a 95 percent probability, 95 percent confidence level, if flooded with unborated water.

The maximum enrichment of fresh fuel assemblies must be less than or equal to 5.0 weight-percent U-235. Note this design criterion is provided in 10 CFR Part 50.68, Item 7 of Paragraph b. Note that the specific terminology is:

The maximum nominal U-235 enrichment of the fresh fuel assemblies is limited to five (5.0) percent by weight.

following conditions are assumed in meeting this design bases:

The fuel assembly contains the highest enrichment authorized without any control rods or non-integral burnable absorber(s) and is at its most reactive point in life.

For flooded conditions, the moderator is pure water at the temperature within the design limits which yields the largest reactivity.

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Mechanical uncertainties are treated by combining both the worst-case bounding value and sensitivity study approaches.

Credit is taken for the neutron absorption in structural materials and in solid materials added specifically for neutron absorption.

l depletion analyses during core operation were performed with CASMO-4 (using the 70-group s-section library), a two-dimensional multigroup transport theory code based on capture babilities (Reference 53). CASMO-4 is used to determine the isotopic composition of the spent

. In addition, the CASMO-4 calculations are restarted in the storage rack geometry, yielding the

-dimensional infinite multiplication factor (kinf) for the storage rack to determine the reactivity ct of fuel and rack tolerances, temperature variation, and to perform various studies.

design method which determines the criticality safety of fuel assemblies outside the reactor uses MCNP4a code (Reference 21), with continuous energy cross-sections based on ENDF/B-V and DF/B-VI.

t of 62 critical experiments has been analyzed using the above method to demonstrate its licability to criticality analysis and to establish the method bias and uncertainty. The benchmark eriments cover a wide range of geometries, materials, and enrichments, all of them adequate for lifying methods to analyze light water reactor lattices (References 22 to 28, and 65 to 68).

analysis of the 62 critical experiments results in an average Keff of 0.9991. Comparison with the sured values results in a method bias of 0.0009. The standard deviation of the set of reactivities 0011. The 95/95 tolerance factor is conservatively set to 2.0.

analytical methods employed herein conform with ANSI N18.2 (Reference 3), Section 5.7, Fuel dling System; ANSI N16.9 (Reference 29), NRC Standard Review Plan, Subsection 9.1.2, the C guidance, OT Position for Review and Acceptance of Spent Fuel Storage and Handling lications (Reference 30).

2.6.2 Soluble Boron Credit Methodology minimum soluble boron requirement under normal and accident conditions must be determined how that the reactivity of the spent fuel racks remains below 0.95. This is achieved by crediting a rete amount of soluble boron and then determining by linear interpolation the appropriate amount oluble boron necessary to reduce the maximum Keff to 0.95 with all uncertainties and biases uded.

2.7 Stability 2.7.1 Introduction stability of the PWR cores against xenon-induced spatial oscillations and the control of such sients are discussed extensively in References 11, 31, 32, and 33. A summary of these reports is n in the following discussion, and the design bases are given in Subsection 4.3.1.6.

large reactor core, xenon-induced oscillations can take place with no corresponding change in total power of the core. The oscillation may be caused by a power shift in the core which occurs dly by comparison with the xenon-iodine time constants. Such a power shift occurs in the axial ction when a plant load change is made by control rod motion and results in a change in the 4.3-26 Revision 1

to the negative power coefficient of reactivity, PWR cores are inherently stable to oscillations in l power. Protection against total power instabilities is provided by the control and protection em, as described in Section 7.7. Hence, the discussion on the core stability will be limited to on-induced spatial oscillations.

2.7.2 Stability Index er distributions, either in the axial direction or in the X-Y plane, can undergo oscillations due to urbations introduced in the equilibrium distributions without changing the total core power. The monics and the stability of the core against xenon-induced oscillations can be determined in terms e eigenvalue of the first flux harmonics. Writing the eigenvalue of the first flux harmonic as:

= b + ic (1) n b is defined as the stability index and T = 2 /c as the oscillation period of the first harmonic.

time dependence of the first harmonic in the power distribution can now be represented as:

(t ) = A e t = a e bt cos ct (2) re A and a are constants. The stability index can also be obtained approximately by:

1 A b= ln n+1 (3)

T An re A n and A n+1 are the successive peak amplitudes of the oscillation and T is the time period ween the successive peaks.

2.7.3 Prediction of the Core Stability core described in this report has an active fuel length that is 24 inches longer (nominal) than that previous Westinghouse PWRs licensed in the U.S. with 157 fuel assemblies. For this reason, it is ected that this core will be as stable as the 12-foot designs with respect to radial and diametral on oscillations since the radial core dimensions have not changed. This core will be slightly less le than the 12-foot, 157 assembly cores with respect to axial xenon oscillations because the ve core height has been increased by 24 inches. The effect of this increase will be to decrease burnup at which the axial stability index becomes zero (Subsection 4.3.2.7.4 below). The erator temperature coefficients and the Doppler temperature coefficients of reactivity will be lar to those of previous designs. Control banks included in the core design are sufficient to pen any xenon oscillations that may occur. Free axial xenon oscillations are not allowed to occur a core of any height, except during special tests as described in Subsection 4.3.2.7.4.

2.7.4 Stability Measurements 2.7.4.1 Axial Measurements axial xenon transient tests conducted in a PWR with a core height of 12 feet and 121 fuel emblies are reported in WCAP-7964 (Reference 34) and are discussed here. The tests were ormed at approximately 10 percent and 50 percent of cycle life.

4.3-27 Revision 1

ulse motion of the lead control bank and the subsequent oscillation period was monitored. In the trolled test conducted early in the cycle, the part-length rods were used to follow the oscillations aintain an axial offset within the prescribed limits. The axial offset of power was obtained from the ore ion chamber readings (which had been calibrated against the in-core flux maps) as a function me for both free-running tests, as shown in Figure 12 of WCAP-7964 (Reference 34).

total core power was maintained constant during these spatial xenon tests, and the stability x and the oscillation period were obtained from a least-square fit of the axial offset data in the of equation 2. The axial offset of power is the quantity that properly represents the axial stability e sense that it essentially eliminates any contribution from even-order harmonics, including the amental mode. The conclusions of the tests follow:

The core was stable against induced axial xenon transients, at the core average burnups of both 1550 MWD/MTU and 7700 MWD/MTU. The measured stability indices are -0.041 h-1 for the first test and - 0.014 h-1 for the second test. The corresponding oscillation periods are 32.4 and 27.2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />, respectively.

The reactor core becomes less stable as fuel burnup progresses, and the axial stability index is essentially zero at 12,000 MWD/MTU. However, the movable control rod systems can control axial oscillations, as described in Subsection 4.3.2.7.

2.7.4.2 Measurements in the X-Y Plane X-Y xenon oscillation tests were performed at a PWR plant with a core height of 12 feet and fuel assemblies. The first test was conducted at a core average burnup of 1540 MWD/MTU and second at a core average burnup of 12,900 MWD/MTU. Both of the X-Y xenon tests show that core was stable in the X-Y plane at both burnups. The second test shows that the core became e stable as the fuel burnup increased, and Westinghouse PWRs with 121 and 157 assemblies stable throughout their burnup cycles. The results of these tests are applicable to the

-assembly AP1000 core, as discussed in Subsection 4.3.2.7.3.

ach of the two X-Y tests, a perturbation was introduced to the equilibrium power distribution ugh an impulse motion of one rod cluster control unit located along the diagonal axis. Following perturbation, the uncontrolled oscillation was monitored, using the movable detector and mocouple system and the ex-core power range detectors. The quadrant tilt difference (QTD) is quantity that properly represents the diametral oscillation in the X-Y plane of the reactor core in the differences of the quadrant average powers over two symmetrically opposite quadrants entially eliminates the contribution to the oscillation from the azimuthal mode. The quadrant tilt rence data were fitted in the form of equation 2 of Subsection 4.3.2.7.2 through a least-square hod. A stability index of - 0.076 hr-1 (per hour) with a period of 29.6 hr was obtained from the mocouple data shown in Figure 4.3-31.

as observed in the second X-Y xenon test that the PWR core with 157 fuel assemblies had ome more stable due to an increased fuel depletion, and the stability index was not determined.

2.7.5 Comparison of Calculations with Measurements direct simulation of axial offset data was carried out using a licensed one-dimensional code AP-7084-P-A Reference 35). The analysis of the X-Y xenon transient tests was performed in an geometry, using a licensed few group two-dimensional code (WCAP-7213-A Reference 36).

h of these codes solve the two-group, time-dependent neutron diffusion equation with 4.3-28 Revision 1

detailed experimental data during the tests, including the reactor power level, the enthalpy rise, the impulse motion of the control rod assembly, as well as the plant follow burnup data, were ely simulated in the study.

results of the stability calculation for the axial tests are compared with the experimental data in le 4.3-5. The calculations show conservative results for both of the axial tests with a margin of roximately 0.01 hr-1 in the stability index.

analytical simulation of the first X-Y xenon oscillation test shows a calculated stability index of 081 hr-1, in good agreement with the measured value of - 0.076 hr-1. As indicated earlier, the ond X-Y xenon test showed that the core had become more stable compared to the first test, and valuation of the stability index was attempted. This increase in the core stability in the X-Y plane to increased fuel burnup is due mainly to the increased magnitude of the negative moderator perature coefficient.

vious studies of the physics of xenon oscillations, including three-dimensional analysis, are orted in a series of topical reports (References 31, 32, and 33). A more detailed description of the erimental results and analysis of the axial and X-Y xenon transient tests is presented in AP-7964 (Reference 34) and Section 1 of WCAP-8768 (Reference 37).

2.7.6 Stability Control and Protection online monitoring system provides continuous indication of current power distributions and ides guidance to the plant operator as to the timing and most appropriate action(s) to maintain le axial power distributions. In the event the online monitoring system is out of service, the ore detector system is utilized to provide indications of xenon-induced spatial oscillations. The dings from the ex-core detectors are available to the operator and also form part of the protection em.

2.7.6.1 Axial Power Distribution rod control system automatically maintains axial power distribution within very tight axial offset ds as part of normal operation. The AO control bank is specifically designed with sufficient worth e capable of maintaining essentially constant axial offset over the power operating range. The rod trol system is also allowed to be operated in manual control in which case the operator is ructed to maintain an axial offset within a prescribed operating band, based on the ex-core ctor readings. Should the axial offset be permitted to move far enough outside this band, the ection limit is encroached, and the turbine power is automatically reduced or a reactor trip signal erated, or both.

uel burnup progresses, PWR cores become less stable to axial xenon oscillations. However, free on oscillations are not allowed to occur, except for special tests. The AO control bank is sufficient ampen and control any axial xenon oscillations present. Should the axial offset be inadvertently mitted to move far enough outside the allowed band due to an axial xenon oscillation or for any r reason, the OTT and/or OPT protection setpoint including the axial offset compensation is hed and the turbine power is automatically reduced and/or a reactor trip signal is generated.

4.3-29 Revision 1

core life.

X-Y stability of large PWRs has been further verified as part of the startup physics test program PWR cores with 193 fuel assemblies. The measured X-Y stability of the cores with 157 and assemblies was in close 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 possible and would be implemented, if necessary, to increase the natural stability of the core.

is based on the fact that several actions could be taken to make the moderator temperature fficient more negative, which would increase the stability of the core in the X-Y plane.

visions for protection against non-symmetric perturbations in the X-Y power distribution that could lt from equipment malfunctions are made in the protection system design. This includes control drop, rod misalignment, and asymmetric loss of coolant flow.

ore detailed discussion of the power distribution control in PWR cores is presented in AP-7811 (Reference 11) and WCAP-8385 (Reference 12).

2.8 Vessel Irradiation view of the methods and analyses used in the determination of neutron and gamma ray flux nuation between the core and the pressure vessel is provided below. A more complete ussion on the pressure vessel irradiation and surveillance program is given in Section 5.3.

materials that serve to attenuate neutrons originating in the core and gamma rays from both the and structural components consist of the core shroud, core barrel and associated water annuli.

se are within the region between the core and the pressure vessel.

eneral, few group neutron diffusion theory codes are used to determine fission power density ributions within the active core, and the accuracy of these analyses is verified by in-core surements on operating reactors. Region and rodwise power-sharing information from the core ulations is then used as source information in two-dimensional transport calculations which pute the flux distributions throughout the reactor.

neutron flux distribution and spectrum in the various structural components vary significantly the core to the pressure vessel. Representative values of the neutron flux distribution and ctrum are presented in Table 4.3-6.

iscussed in Section 5.3, the irradiation surveillance program utilizes actual test samples to verify accuracy of the calculated fluxes at the vessel.

3 Analytical Methods culations required in nuclear design consist of three distinct types, which are performed in uence:

Determination of effective fuel temperatures Generation of microscopic few-group parameters Space-dependent, few-group diffusion calculations 4.3-30 Revision 1

3.1 Fuel Temperature (Doppler) Calculations peratures vary radially within the fuel rod, depending on the heat generation rate in the pellet; the ductivity of the materials in the pellet, gap, and clad; and the temperature of the coolant.

fuel temperatures for use in most nuclear design Doppler calculations are obtained from a plified version of the Westinghouse fuel rod design model described in Subsection 4.2.1.3, which siders the effect of radial variation of pellet conductivity, expansion coefficient and heat eration rate, elastic deflection of the clad, and a gap conductance which depends on the initial fill

, the hot open gap dimension, and the fraction of the pellet over which the gap is closed. The tion of the gap assumed closed represents an empirical adjustment used to produce close eement with observed reactivity data at beginning of life. Further gap closure occurs with burnup accounts for the decrease in Doppler defect with burnup which has been observed in operating ts. For detailed calculations of the Doppler coefficient, such as for use in xenon stability ulations, a more sophisticated temperature model is used, which accounts for the effects of fuel lling, fission gas release, and plastic clad deformation.

ial power distributions in the pellet as a function of burnup are obtained from LASER AP-6073, Reference 38) calculations.

effective U-238 temperature for resonance absorption is obtained from the radial temperature ribution by applying a radially dependent weighing function. The weighing function was rmined from REPAD (WCAP-2048, Reference 39) Monte Carlo calculations of resonance ape probabilities in several steady-state and transient temperature distributions. In each case, a pellet temperature was determined which produced the same resonance escape probability as actual distribution. The weighing function was empirically determined from these results.

effective Pu-240 temperature for resonance absorption is determined by a convolution of the al distribution of Pu-240 densities from LASER burnup calculations and the radial weighing tion. The resulting temperature is burnup dependent, but the difference between U-238 and 240 temperatures, in terms of reactivity effects, is small.

effective pellet temperature for pellet dimensional change is that value which produces the same r pellet radius in a virgin pellet as that obtained from the temperature model. The effective clad perature for dimensional change is its average value.

temperature calculational model has been validated by plant Doppler defect data, as shown in le 4.3-7, and Doppler coefficient data, as shown in Figure 4.3-32. Stability index measurements provide a sensitive measure of the Doppler coefficient near full power (Subsection 4.3.2.7).

3.2 Macroscopic Group Constants OENIX-P (WCAP-11596-P-A, Reference 40) has been used for generating the macroscopic cross ions needed for the spatial few group codes. PHOENIX-P or other NRC approved lattice codes be used for reload designs.

OENIX-P has been approved by the NRC as a lattice code for the generation of macroscopic and roscopic few group cross sections for PWR analysis. (See WCAP-11596-P-A, Reference 40).

OENIX-P is a two-dimensional, multigroup, transport-based lattice code capable of providing essary data for PWR analysis. Since it is a dimensional lattice code, PHOENIX-P does not rely on 4.3-31 Revision 1

solution for the detailed spatial flux and energy distribution is divided into two major steps in OENIX-P (See References 40 and 41). First, a two-dimensional fine energy group nodal solution btained, coupling individual subcell regions (e.g., pellet, clad and moderator) as well as ounding pins, using a method based on Carlviks collision probability approach and rogeneous response fluxes which preserve the heterogeneous nature of the pin cells and their oundings. The nodal solution provides an accurate and detailed local flux distribution, which is used to homogenize the pin cells spatially to few groups.

n, a standard S4 discrete ordinates calculation solves for the angular distribution, based on the up-collapsed and homogenized cross sections from the first step. These S4 fluxes normalize the iled spatial and energy nodal fluxes, which are then used to compute reaction rates, power ributions and to deplete the fuel and burnable absorbers. A standard B1 calculation evaluates the amental mode critical spectrum, providing an improved fast diffusion coefficient for the core tial codes.

OENIX-P employs either a 42 or 70 energy group library derived mainly from the ENDF/B-V files ference 21). This library was designed to capture the integral properties of the multigroup data perly during group collapse and to model important resonance parameters properly. It contains tronics data necessary for modelling fuel, fission products, cladding and structural materials, lant, and control and burnable absorber materials present in PWRs.

up constants for burnable absorber cells, control rod cells, guide thimbles and instrumentation bles, or other non-fuel cells, can be obtained directly from PHOENIX-P without any adjustments h as those required in the cell or 1D lattice codes.

OENIX-P has been validated through an extensive qualification effort which includes calculation-surement comparison of the Strawbridge-Barry critical experiments (See References 42 and 43),

KRITZ high temperature criticals (Reference 44), the AEC sponsored B&W criticals ferences 45 through 47) and measured actinide isotopic data from fuel pins irradiated in the ton and Yankee Rowe cores (References 48 through 52). In addition, calculation-measurement parisons have been made to operating reactor data measured during startup tests and during mal power operation.

dation of the cross section method is based on analysis of critical experiments, isotopic data, t critical boron concentration data, and control rod worth measurement data such as that shown able 4.3-8.

firmatory critical experiments on burnable absorber rods are described in WCAP-7806 ference 42).

3.3 Spatial Few-Group Diffusion Calculations 3D ANC code (see WCAP-10965-P-A, Reference 57) permits the introduction of advanced fuel igns with axial heterogeneities, such as axial blankets and part-length burnable absorbers, and ws such features to be modeled explicitly. The three dimensional nature of this code provides radial and axial power distribution. For some applications, the updated version APOLLO (see AP-13524 Reference 60) of the PANDA code (see WCAP-7084-P-A Reference 35) will continue e used for axial calculations, and a two-dimensional collapse of 3D ANC that properly accounts he three-dimensional features of the fuel is used for X-Y calculations.

4.3-32 Revision 1

dation of the reactivity calculations is associated with validation of the group constants mselves, as discussed in Subsection 4.3.3.2. Validation of the Doppler calculations is associated the fuel temperature validation discussed in Subsection 4.3.3.1. Validation of the moderator fficient calculations is obtained by comparison with plant measurements at hot zero power ditions, similar to that shown in Table 4.3-9.

l calculations are used to determine differential control rod worth curves (reactivity versus rod rtion) and to demonstrate load follow capability. Group constants are obtained from the three-ensional nodal model by flux-volume weighing on an axial slicewise basis. Radial bucklings are rmined by varying parameters in the buckling model while forcing the one-dimensional model to oduce the axial characteristics (axial offset, midplane power) of the three-dimensional model.

dation of the spatial codes for calculating power distributions involves the use of in-core and ore detectors and is discussed in Subsection 4.3.2.2.7.

discussed in Subsection 4.3.3.2, calculation-measurement comparisons have been made to rating reactor data measured during startup tests and during normal power operation. These parisons include a variety of core geometries and fuel loading patterns, and incorporate a onable extreme range of fuel enrichment, burnable absorber loading, and cycle burnup.

lification data identified in Reference 40 indicate small mean and standard deviations relative to surement which are equal to or less than those found in previous reviews of similar or parallel roved methodologies. For the reload designs the spatial codes described above, other NRC roved codes, or both are used.

4 Combined License Information nges to the reference design of the fuel, burnable absorber rods, rod cluster assemblies, or initial design from that presented in the DCD are addressed in APP-GW-GLR-059 (Reference 64).

5 References Bordelon, F. M, et al., Westinghouse Reload Safety Evaluation Methodology, WCAP-9272-P-A (Proprietary) and WCAP-9273-NP-A (Nonproprietary), July 1985.

[Davidson, S. L. (Ed.), Fuel Criteria Evaluation Process, WCAP-12488-P-A (Proprietary) and WCAP-14204-A - (Nonproprietary), October 1994.]*

ANSI N18.2-1973, Nuclear Safety Criteria for the Design of Stationary Pressurized Water Reactor Plants.

Beard, C. L. and Morita, T., BEACON: Core Monitoring and Operations Support System, WCAP-12472-P-A (Proprietary) and WCAP-12473-A (Nonproprietary), August 1994; Addendum 1, May 1996; and Addendum 2, March 2001.

Gangloff, W. C. and Loftus, W. D., Westinghouse Anticipated Transients Without Reactor Trip Analysis, WCAP-8330, August 1974.

Not used.

Staff approval is required prior to implementing a change in this information.

4.3-33 Revision 1

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

Meyer, R. O., The Analysis of Fuel Densification, Division of Systems Safety, U.S.

Nuclear Regulatory Commission, NUREG-0085, July 1976.

Hellman, J. M., Olson, C. A., and Yang, J. W., Effects of Fuel Densification Power Spikes on Clad Thermal Transients, WCAP-8359; July 1974.

Moore, J. S., Power Distribution Control of Westinghouse Pressurized Water Reactors, WCAP-7811, December 1971.

Morita, T., et al., Power Distribution Control and Load Following Procedures, WCAP-8385 (Proprietary) and WCAP-8403 (Nonproprietary), September 1974.

Miller, R. W., et al., Relaxation of Constant Axial Offset Control, FQ Surveillance Technical Specification, WCAP-10216-P-A, (Proprietary) and WCAP-10217-A, (Nonproprietary) Revision 1A, February 1994.

McFarlane, A. F., Power Peaking Factors, WCAP-7912-P-A (Proprietary) and WCAP-7912-A (Nonproprietary), January 1975.

Meyer, C. E., and Stover, R. L., Incore Power Distribution Determination in Westinghouse Pressurized Water Reactors, WCAP-8498, July 1975.

Warren, H. D., Rhodium In-Core Detector Sensitivity Depletion, Cycles 2-6, EPRI-NP-3814, December 1984.

Henderson, W. B., Results of the Control Rod Worth Program, WCAP-9217 (Proprietary) and WCAP-9218 (Nonproprietary), October 1977.

Cermak, J. O., et al., Pressurized Water Reactor pH - Reactivity Effect Final Report, WCAP-3696-8 (EURAEC-2074), October 1968.

USNRC Code of Federal Regulations, Title 10, Part 50, Appendix A, Criterion 62, Prevention of Criticality in Fuel Storage and Handling.

Kopp, L. (NRC), Guidance on the Regulatory Requirements for Criticality Analysis of Fuel Storage at Light-Water Reactor Power Plants, February 1998.

Briesmeister, J. F., Editor, MCNP - A General Monte Carlo N-Particle Transport Code, Version 4A, LA-12625, Los Alamos National Laboratory (1993).

Baldwin, M. N., et al., Critical Experiments Supporting Close Proximity Water Storage of Power Reactor Fuel, BAW-1484-7, Babcock & Wilcox Company, July 1979.

Hoovier, G. S., et al., Critical Experiments Supporting Underwater Storage of Tightly Packed Configurations of Spent Fuel Pins, BAW-1645-4, Babcock & Wilcox Company, November 1991.

4.3-34 Revision 1

Manaranche, J. C., et al., Dissolution and Storage Experimental Program with 4.75 w/o Enriched Uranium-Oxide Rods, Trnas. Am. Nucl. Soc. 33:362-364 (1979).

Bierman, S. R. and Clayton, E. D., Criticality Experiments with Subcritical Clusters of 2.35 w/o and 4.31 w/o 235U Enriched UO2 Rods in Water with Steel Reflecting Walls, PNL-3602, Batelle Pacific Northwest Laboratory, April 1981.

Bierman, S. R., et al., Criticality Experiments with Subcritical Clusters of 2.35 w/o and 4.31 w/o 235U Enriched UO2 Rods in Water with Uranium or Lead Reflecting Walls, PNL-3926, Batelle Pacific Northwest Laboratory, December 1981.

Bierman, S. R., et al., Criticality Experiments with Subcritical Clusters of 2.35 w/o and 4.31 w/o 235U Enriched UO2 Rods in Water with Fixed Neutron Poisons, PNL-2615, Batelle Pacific Northwest Laboratory, October 1977.

ANSI N16.9-1975, Validation of Calculational Methods for Nuclear Criticality Safety.

NRC Letter OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications, from Grimes, B. K., to all power reactor licenses, April 14, 1978.

Poncelet, C. G., and Christie, A. M., Xenon-Induced Spatial Instabilities in Large Pressurized Water Reactors, WCAP-3680-20 (EURAEC-1974), March 1968.

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.

Skogen, F. B., and McFarlane, A. F., Xenon-Induced Spatial Instabilities in Three Dimensions, WCAP-3680-22 (EURAEC-2116), September 1969.

Lee, J. C., et al., Axial Xenon Transient Tests at the Rochester Gas and Electric Reactor, WCAP-7964, June 1971.

Barry, R. F., and Minton, G., The PANDA Code, WCAP-7048-P-A (Proprietary) and WCAP-7757-A (Nonproprietary), February 1975.

Barry, R. F., and Altomare, S., The TURTLE 24.0 Diffusion Depletion Code, WCAP-7213-A (Proprietary) and WCAP-7758-A (Non-Proprietary), February 1975.

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.

Poncelet, C. G., LASER - A Depletion Program for Lattice Calculations Based on MUFT and THERMOS, WCAP-6073, April 1966.

Olhoeft, J. E., The Doppler Effect for a Non-Uniform Temperature Distribution in Reactor Fuel Elements, WCAP-2048, July 1962.

4.3-35 Revision 1

Mildrum, C. M., Mayhue, L. T., Baker, M. M., and Isaac, P. G., Qualification of the PHOENIX/POLCA Nuclear Design and Analysis Program for Boiling Water Reactors, WCAP-10841 (Proprietary), and WCAP-10842 (Nonproprietary), June 1985.

Barry, R. F., Nuclear Design of Westinghouse Pressurized Water Reactors with Burnable Poison Rods, WCAP-7806, December 1971.

Strawbridge, L. E., and Barry, R. F., Criticality Calculation for Uniform Water-Moderated Lattices, Nuclear Science and Engineering 23, p. 58, 1965.

Persson, R., Blomsjo, E., and Edenius, M., High Temperature Critical Experiments with H2O Moderated Fuel Assemblies in KRITZ, Technical Meeting No. 2/11, NUCLEX 72, 1972.

Baldwin, M. N., and Stern, M. E., Physics Verification Program Part III, Task 4: Summary Report, BAW-3647-20, March 1971.

Baldwin, M. N., Physics Verification Program Part III, Task 11: Quarterly Technical Report January-March 1974, BAW-3647-30, July 1974.

Baldwin, M. N., Physics Verification Program Part III, Task 11: Quarterly Technical Report July-September 1974, BAW-3647-31, February 1975.

Nodvik, R. J., Saxton Core II Fuel Performance Evaluation Part II: Evaluation of Mass Spectrometric and Radiochemical Analyses of Irradiated Saxton Plutonium Fuel, WCAP-3385-56 Part II, July 1970.

Smalley, W. R., Saxton Core II - Fuel Performance Evaluation Part I: Materials, WCAP-3386-56 Part I, September 1971.

Goodspeed, R. C., Saxton Plutonium Project - Quarterly Progress Report for the Period Ending June 20, 1973, WCAP-3385-36, July 1973.

Crain, H. H., Saxton Plutonium Project - Quarterly Progress Report for the Period Ending September 30, 1973, WCAP-3385-37, December 1973.

Melehan, J. B., Yankee Core Evaluation Program Final Report, WCAP-3017-6094, January 1971.

APP-GW-GLR-029P, Revision 3, AP1000 Spent Fuel Storage Racks Criticality Analysis, Westinghouse Electric Company LLC (Westinghouse Proprietary).

Not used.

Not used.

Not used.

Davidson, S. L., (Ed.), et al., ANC: A Westinghouse Advanced Nodal Computer Code, WCAP-10965-P-A (Proprietary) and WCAP-10966-A (Nonproprietary), September 1986.

4.3-36 Revision 1

Davidson, S. L., et al., Assessment of Clad Flattening and Densification Power Spike Factor Elimination in Westinghouse Nuclear Fuel, WCAP-13589-A (Proprietary) and WCAP-14297-A (Nonproprietary), March 1995.

Yarbrough, M. B., Liu, Y. S., Paterline, D. L., Hone, M. J., APOLLO - A One Dimensional Neutron Theory Program, WCAP-13524, Revision 1 (Proprietary), August 1994 and WCAP-14952-NP-A, Revision 1A (Nonproprietary), September 1977.

Letter, Peralta, J. D. (NRC) to Maurer, B. F. (Westinghouse), Approval for Increase in Licensing Burnup Limit to 62,000 MWD/MTU (TAC No. MD1486), May 25, 2006.

Not used.

Not used.

APP-GW-GLR-059/WCAP-16652-NP, AP1000 Core & Fuel Design Technical Report, Revision 0.

Bierman, S. R., Criticality Experiments with Neutron Flux Traps Containing Voids, PNL-7167, Battelle Pacific Northwest Laboratory, April 1990.

Durst, B. M., et al., Critical Experiments with 4.32 wt% 235U Enriched UO2 Rods in Highly Borated Water Lattices, PNL-4267, Battelle Pacific Northwest Laboratory, August 1982.

Bierman, S. R., Criticality Experiments with Fast Test Reactor Fuel Pins in Organic Moderator, PNL-5803, Battelle Pacific Northwest Laboratory, December 1981.

Taylor, E. G., et al., Saxton Plutonium Program Critical Experiments for the Saxton Partial Plutonium Core, WCAP-3385-54, Westinghouse Electric Corp., Atomic Power Division, December 1965.

4.3-37 Revision 1

(FIRST CYCLE)]*

ve core Equivalent diameter (in.) ............................................................................................................................... 119.7 Active fuel height first core (in.), cold ............................................................................................................... 168 Height-to-diameter ratio .................................................................................................................................. 1.40 Total cross section area (ft2) .......................................................................................................................... 78.14 H2O/U molecular ratio, cell, cold ...................................................................................................................... 2.40 lector thickness and composition Top - water plus steel (in.) ................................................................................................................................ ~10 Bottom - water plus steel (in.) .......................................................................................................................... ~10 Side - water plus steel (in.)................................................................................................................................ ~15 l assemblies Number ............................................................................................................................................................ 157 Rod array ................................................................................................................................................... 17 x 17 Rods per assembly .......................................................................................................................................... 264 Rod pitch (in.)................................................................................................................................................ 0.496 Overall transverse dimensions (in.).................................................................................................. 8.426 x 8.426 Fuel weight, as UO2 (lb) ............................................................................................................................ 211,588 Zircaloy clad weight (lb) .............................................................................................................................. 43,105 Number of grids per assembly Top and bottom - (Ni-Cr-Fe Alloy 718) .......................................................................................................2(a)

Intermediate ...................................................................................................................................8 ZIRLO' Intermediate flow mixing (IFM).......................................................................................................4 ZIRLO' Number of guide thimbles per assembly............................................................................................................. 24 Composition of guide thimbles ..................................................................................................................ZIRLO' meter of guide thimbles, upper part (in.) .................................................................................. 0.442 ID x 0.482 OD meter of guide thimbles, lower part (in.)................................................................................... 0.397 ID x 0.482 OD meter of instrument guide thimbles (in.)................................................................................... 0.442 ID x 0.482 OD e:

The top grid will be fabricated of nickel-chromium-iron Alloy 718.

Staff approval is required prior to implementing a change in this information.

4.3-38 Revision 1

l rods Number ....................................................................................................................................................... 41,448 Outside diameter (in.).................................................................................................................................... 0.374 Diameter gap (in.) ....................................................................................................................................... 0.0065 Clad thickness (in.)...................................................................................................................................... 0.0225 Clad material .............................................................................................................................................ZIRLO' l pellets Material ............................................................................................................................................. UO2 sintered Density (% of theoretical) (nominal) ................................................................................................................ 95.5 Fuel enrichments (weight %)

Region 1................................................................................................................................................... 2.35 Region 2................................................................................................................................................... 3.40 Region 3................................................................................................................................................... 4.45 Diameter (in.) .............................................................................................................................................. 0.3225 Length (in.) .................................................................................................................................................... 0.387 Mass of UO2 per ft of fuel rod (lb/ft) ............................................................................................................... 0.366 Cluster Control Assemblies Neutron absorber .....................................................................................................................................Ag-In-Cd Diameter (in.) ......................................................................................................................................... 0.341 Density (lb/in.3) .......................................................................................................................Ag-In-Cd 0.367 Cladding material ......................................................................................................... Type 304, cold-worked SS Clad thickness (in.)...................................................................................................................................... 0.0185 Number of clusters, full-length ........................................................................................................................... 53 Number of absorber rods per cluster .................................................................................................................. 24 y Rod Cluster Assemblies Neutron absorber ..........................................................................................................................Ag-In-Cd/304SS Diameter (in.) .......................................................................................................................................... 0.160 Density (lb/in.3) ................................................................................................Ag-In-Cd 0.367 / 304SS 0.285 Cladding material .......................................................................................................... Type 304, cold-worked SS Clad thickness (in.)....................................................................................................................................... 0.0185 Number of clusters, full-length ............................................................................................................................ 16 mber of absorber rods per cluster........................................................................................ 12 Ag-In-Cd / 12 304SS Staff approval is required prior to implementing a change in this information.

4.3-39 Revision 1

crete Burnable absorber rods (first core)

Number ........................................................................................................................................................... 1558 Material ...................................................................................................................................... Borosilicate Glass OD (in.)........................................................................................................................................................... 0.381 Inner tube, OD (in.) ...................................................................................................................................... 0.1815 Clad material ................................................................................................................................... Stainless Steel Inner tube material .......................................................................................................................... Stainless Steel B10 content (Mg/cm)......................................................................................................................................... 6.24 Absorber length (in.).......................................................................................................................................... 145 gral Fuel Burnable Absorbers (first core)

Number .......................................................................................................................................................... 8832 Type................................................................................................................................................................ IFBA Material .......................................................................................................................................... Boride Coating B10 Content (Mg/cm)..................................................................................................................................... 0.772 Absorber length (in.).......................................................................................................................................... 152 ess reactivity Maximum fuel assembly K (cold, clean, ..................................................................................................... 1.328 unborated water)

Maximum core reactivity Keff (cold, zero power, ........................................................................................... 1.205 beginning of cycle, zero soluble boron)

Staff approval is required prior to implementing a change in this information.

4.3-40 Revision 1

(FIRST CYCLE)]*

e average linear power, including densification effects (kW/ft) ..........................................................................5.72 l heat flux hot channel factor, FQ.......................................................................................................................2.60 lear enthalpy rise hot channel factor, F N H .....................................................................................................1.65 ctivity coefficients (a) Design Limits Best Estimate Doppler-only power coefficients (see Figure 15.0.4-1) (pcm/% power)(b)

Upper curve............................................................................................ -19.4 to -12.6 .............. -13.3 to -8.7 Lower curve............................................................................................ -10.2 to -6.7 .................-11.3 to -8.4 Doppler temperature coefficient (pcm/°F)(b) .................................................. -3.5 to -1.0 .................... -2.1 to -1.3 Moderator temperature coefficient (pcm/°F)(b) .............................................. 0 to -40................................0 to -35 Boron coefficient (pcm/ppm)(b) ...................................................................... -13.5 to -5.0 ................ -10.5 to -6.9 ded moderator density (pcm/g/cm3)(b) ........................................................... 0.47x105 ..................... 0.45x105 ayed neutron fraction and lifetime, eff ....................................................................................... 0.0075(0.0044)(c) mpt Neutron Lifetime, l* , s ...........................................................................................................................19.8 trol rods Rod requirements..........................................................................................................................See Table 4.3-3 Maximum ejected rod worth ...........................................................................................................See Chapter 15 k worth HZP no overlap (pcm)(b) BOL, Xe Free EOL, Eq. Xe MA Bank ........................................................................................................ 299............................................205 MB Bank........................................................................................................ 195............................................250 MC Bank........................................................................................................ 139............................................218 MD Bank........................................................................................................ 312............................................198 M1 Bank ........................................................................................................ 856............................................632 M2 Bank ........................................................................................................ 933..........................................1405 AO Bank ........................................................................................................ 2027.........................................1571 Staff approval is required prior to implementing a change in this information.

4.3-41 Revision 1

cal Hot Channel Factors F N H ..................................................................... BOL..........................................EOL Unrodded....................................................................................................... 1.40..........................................1.33 MA bank ........................................................................................................ 1.46..........................................1.38 MA + MB banks ............................................................................................. 1.46..........................................1.33 MA + MB + MC banks ................................................................................... 1.50..........................................1.31 MA + MB + MC + MD banks.......................................................................... 1.50..........................................1.37 MA + MB + MC + MD + M1 banks................................................................. 1.52..........................................1.45 AO bank ........................................................................................................ 1.60...........................................1.52 on concentrations (ppm)

Zero power, keff = 0.99, cold(d) RCCAs out ....................................................................................................1574 Zero power, keff = 0.99, hot(e) RCCAs out.......................................................................................................1502 Design basis refueling boron concentration ....................................................................................................2700 Zero power, keff 0.95, cold(d) RCCAs in....................................................................................................... 1179 Zero power, keff = 1.00, hot(e) RCCAs out......................................................................................................1382 Full power, no xenon, keff = 1.0, hot RCCAs out ............................................................................................ 1184 Full power, equilibrium xenon, k = 1.0, hot RCCAs out ....................................................................................827 Reduction with fuel burnup First cycle (ppm/(GWD/MTU))(f) ........................................................................................... See Figure 4.3-3 Reload cycle (ppm/(GWD/MTU)) ...............................................................................................................~40 es:

Uncertainties are given in Subsection 4.3.3.3.

1 pcm = 10-5 where is calculated form two statepoint values of keff by ln (k1/k2).

Bounding lower value used for safety analysis.

Cold means 68°F, 1 atm.

Hot means 557°F, 2250 psia.

1 GWD = 1000 MWD. During the first cycle, a large complement of burnable absorbers is present which significantly reduce the boron depletion rate compared to reload cycles.

Staff approval is required prior to implementing a change in this information.

4.3-42 Revision 1

EOL Reactivity Effects BOL EOL Representative (Percent) (First Cycle) (First Cycle) (Equilibrium Cycle)

Control requirements Total power defect (%)(a) 1.89 2.54 3.02 Redistribution (adverse xenon only) (%) 0.27 0.40 0.32 Rod insertion allowance (%) 2.00 2.00 2.00 Total control (%) 4.16 4.94 5.34 Estimated RCCA worth (69 rods)

a. All full-length assemblies inserted (%) 12.69 10.89 10.64

b. All assemblies but one (highest worth inserted 10.49 9.27 9.35

(%)

Estimated RCCA credit with 7 percent adjustment to 9.76 8.62 8.70 accommodate uncertainties, item 3b minus 7 percent (%)

Shutdown margin available, item 4 minus item 2 5.60 3.68 3.36

(%)(b) es:

Includes void effects.

The design basis minimum shutdown is 1.60 percent.

Staff approval is required prior to implementing a change in this information.

4.3-43 Revision 1

Not Used 4.3-44 Revision 1

Reactor Cores with a 12-Foot Height Burnup CB Axial Stability Index (h-1)

(MWD/MTU) FZ (ppm) Experiment Calculated 1550 1.34 1065 -0.0410 -0.0320 7700 1.27 700 -0.0140 -0.0060 (a) -0.0325 -0.0255 5090 Radial Stability Index (h-1)

Experiment Calculated (b) -0.0680 -0.0700 2250 s:

Four-loop plant, 12-foot core in cycle 1, axial stability test Four-loop plant, 12-foot core in cycle 1, radial (X-Y) stability test 4.3-45 Revision 1

1.00 MeV > E 5.53 KeV > E E 1.0 MeV 5.53 KeV 0.625 eV E < 0.625 eV e center 1.12x1014 1.76x1014 1.28x1014 5.47x1013 e outer radius at midheight 3.86x1013 6.08x1013 4.42x1013 1.83x1013 e top, on axis 3.02x1013 4.75x1013 3.46x1013 2.17x1013 e bottom, on axis 2.92x1013 4.59x1013 3.34x1013 2.40x1013 ssure vessel ID azimuthal peak 4.71x1010 8.4x1010 5.56x1010 5.32x1010 4.3-46 Revision 1

Core Burnup Measured Calculated Plant Fuel (MWD/MTU) (pcm)(a) (pcm) 1 Air filled 1800 1700 1710 2 Air filled 7700 1300 1440 3 Air and helium filled 8460 1200 1210 pcm = 105 x ln (k2/k1) 4.3-47 Revision 1

2-Loop Plant, 121 Assemblies, 10-ft Core Measured (pcm) Calculated (pcm) up B 1885 1893 up A 1530 1649 tdown group 3050 2917 (a)

DA critical, 0.69-in. pitch

/o PuO2, 8% Pu-240, 9 control rods

-in. rod separation 2250 2250

-in. rod separation 4220 4160

-in. rod separation 4100 4019 Benchmark Critical Experiment Hafnium Control Rod Worth Control No. of Measured(b) Calculated(b)

Rod Fuel Worth Worth Configuration Rods (ppm B-10) (ppm B-10) 9 hafnium rods 1192 138.3 141.0 s:

Report in WCAP-3726-1 (Reference 58).

Calculated and measured worth are given in terms of an equivalent charge in B-10 concentration.

4.3-48 Revision 1

Coefficients at HZP, BOL Plant Type/ Measured iso(a) Calculated iso Control Bank Configuration (pcm/°F) (pcm/°F) op, 157-assembly, 12-ft core at 160 steps -0.50 -0.50 in, C at 190 steps -3.01 -2.75 in, C at 28 steps -7.67 -7.02

, C, and D in -5.16 -4.45 op, 121-assembly, 12-ft core at 180 steps +0.85 +1.02 in, C at 180 steps -2.40 -1.90 and D in, B at 165 steps -4.40 -5.58

, C, and D in, A at 174 steps -8.70 -8.12 op, 193-assembly, 12-ft core RO -0.52 -1.2 in -4.35 -5.7 and C in -8.59 -10.0

, C, and B in -10.14 -10.55

, C, B, and A in -14.63 -14.45 Isothermal coefficients, which include the Doppler effect in the fuel k2 iso = 105 ln / T °F k1 4.3-49 Revision 1

Figure 4.3-1 Fuel Loading Arrangement 4.3-50 Revision 1

Figure 4.3-2 Typical Production and Consumption of Higher Isotopes 4.3-51 Revision 1

Figure 4.3-3 Cycle 1 Soluble Boron Concentration Versus Burnup 4.3-52 Revision 1

Figure 4.3-4a Cycle 1 Assembly Burnable Absorber Patterns 4.3-53 Revision 1

Figure 4.3-4b (Sheet 1 of 2)

Cycle 1 Assembly Burnable Absorber Patterns 4.3-54 Revision 1

Figure 4.3-4b (Sheet 2 of 2)

Cycle 1 Assembly Burnable Absorber Patterns 4.3-55 Revision 1

Figure 4.3-5 Burnable Absorber, Primary, and Secondary Source Assembly Locations 4.3-56 Revision 1

Figure 4.3-6 Normalized Power Density Distribution Near Beginning of Life, Unrodded Core, Hot Full Power, No Xenon 4.3-57 Revision 1

Figure 4.3-7 Normalized Power Density Distribution Near Beginning of Life, Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-58 Revision 1

Figure 4.3-8 Normalized Power Density Distribution Near Beginning of Life, Gray Bank MA+MB Inserted, Hot Full Power, Equilibrium Xenon 4.3-59 Revision 1

Figure 4.3-9 Normalized Power Density Distribution Near Middle of Life, Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-60 Revision 1

Figure 4.3-10 Normalized Power Density Distribution Near End of Life, Unrodded Core, Hot Full Power, Equilibrium Xenon 4.3-61 Revision 1

Figure 4.3-11 Normalized Power Density Distribution Near End of Life, Gray Bank MA+MB Inserted, Hot Full Power, Equilibrium Xenon 4.3-62 Revision 1

Figure 4.3-12 Rodwise Power Distribution in a Typical Assembly (G-9)

Near Beginning of Life Hot Full Power, Equilibrium Xenon, Unrodded Core 4.3-63 Revision 1

Figure 4.3-13 Rodwise Power Distribution in a Typical Assembly (G-9)

Near End of Life Hot Full Power, Equilibrium Xenon, Unrodded Core 4.3-64 Revision 1

Figure 4.3-14 Maximum FQ x Power Versus Axial Height During Normal Operation 4.3-65 Revision 1

Figure 4.3-15 Typical Comparison Between Calculated and Measured Relative Fuel Assembly Power Distribution 4.3-66 Revision 1

Figure 4.3-16 Typical Calculated Versus Measured Axial Power Distribution 4.3-67 Revision 1

Figure 4.3-17 Measured FQ Values Versus Axial Offset for Full Power Rod Configurations 4.3-68 Revision 1

Figure 4.3-18 Typical Doppler Temperature Coefficient at BOL and EOL 4.3-69 Revision 1

Figure 4.3-19 Typical Doppler-Only Power Coefficient at BOL and EOL 4.3-70 Revision 1

Figure 4.3-20 Typical Doppler-Only Power Defect at BOL and EOL 4.3-71 Revision 1

Figure 4.3-21 Typical Moderator Temperature Coefficient at BOL, Unrodded 4.3-72 Revision 1

Figure 4.3-22 Typical Moderator Temperature Coefficient at EOL 4.3-73 Revision 1

Figure 4.3-23 Typical Moderator Temperature Coefficient as a Function of Boron Concentration at BOL, Unrodded 4.3-74 Revision 1

Figure 4.3-24 Typical Hot Full Power Temperature Coefficient Versus Cycle Burnup 4.3-75 Revision 1

Figure 4.3-25 Typical Total Power Coefficient at BOL and EOL 4.3-76 Revision 1

Figure 4.3-26 Typical Total Power Defect at BOL and EOL 4.3-77 Revision 1

Figure 4.3-27 Rod Cluster Control Assembly Pattern 4.3-78 Revision 1

Figure 4.3-28 Typical Accidental Simultaneous Withdrawal of Two Control Banks at EOL, HZP, Moving in the Same Plane 4.3-79 Revision 1

Figure 4.3-29 Typical Design Trip Curve 4.3-80 Revision 1

Figure 4.3-30 Typical Normalized Rod Worth Versus Percent Insertion All Rods Inserting Less Most Reactive Stuck Rod 4.3-81 Revision 1

Figure 4.3-31 X-Y Xenon Test Thermocouple Response Quadrant Tilt Difference Versus Time 4.3-82 Revision 1

Figure 4.3-32 Calculated and Measured Doppler Defect and Coefficients at BOL, 2-Loop Plant, 121 Assemblies, 12-foot Core 4.3-83 Revision 1

heat generation distribution in the core. This provides adequate heat removal by the reactor lant system, the normal residual heat removal system, or the passive core cooling system.

1 Design Basis following performance and safety criteria requirements are established for the thermal and raulic design of the fuel. Condition I, II, III, and IV transients and events through out this section as defined in ANSI N18.2a-75 (Reference 1).

Fuel damage (defined as penetration of the fission product barrier; that is, 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 are within the capability of the plant cleanup system and are consistent with the plant design bases.

The reactor can be brought to a safe state following a Condition III event with only a small fraction of fuel rods damaged (as defined in the above definition), although sufficient fuel damage might occur to preclude resumption of operation without considerable outage time.

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.

atisfy these requirements, the following design bases have been established for the thermal and raulic design of the reactor core.

1.1 Departure from Nucleate Boiling Design Basis 1.1.1 Design Basis re is at least a 95-percent probability at a 95-percent confidence level that departure from leate boiling (DNB) does not occur on the limiting fuel rods during normal operation and rational transients and any transient conditions arising from faults of moderate frequency ndition I and II events).

1.1.2 Discussion design method employed to meet the DNB design basis for the AP1000 fuel assemblies is the ised Thermal Design Procedure, WCAP-11397-P-A (Reference 2). With the Revised Thermal ign Procedure methodology, uncertainties in plant operating parameters, nuclear and thermal ameters, fuel fabrication parameters, computer codes, and DNB correlation predictions are sidered statistically to obtain DNB uncertainty factors. Based on the DNB uncertainty factors, ised Thermal Design Procedure design limits departure from nucleate boiling ratio (DNBR) es are determined such that there is at least a 95-percent probability at a 95-percent confidence l that DNB will not occur on the most limiting fuel rod during normal operation and operational sients and during transient conditions arising from faults of moderate frequency (Condition I II events).

umed uncertainties in the plant operating parameters (pressurizer pressure, primary coolant perature, reactor power, and reactor coolant system flow) are evaluated. Only the random portion e plant operating parameter uncertainties is included in the statistical combination.

rumentation bias is treated as a direct DNBR penalty. Since the parameter uncertainties are 4.4-1 Revision 1

those transients that use the VIPRE-01 computer program (Subsection 4.4.4.5.2) and the B-2M correlation (Subsection 4.4.2.2.1), the Revised Thermal Design Procedure design limits are for the typical cell and 1.25 for the thimble cell for Core and Axial Offset Limits and 1.22 for the cal cell and 1.21 for the thimble cell for all other RTDP transients. These values may be revised htly) when plant specific uncertainties are available.

maintain DNBR margin to offset DNB penalties such as those due to fuel rod bow (as described in section 4.4.2.2.5), the safety analyses are performed to DNBR limits higher than the design limit BR values. The difference between the design limit DNBRs and the safety analysis limit DNBRs lts in DNBR margin. A portion of this margin is used to offset rod bow and unanticipated DNBR alties.

Standard Thermal Design Procedure is used for those analyses where the Revised Thermal ign Procedure is not applicable. In the Standard Thermal Design Procedure method the ameters used in analysis are treated in a conservative way from a DNBR standpoint. The ameter uncertainties are applied directly to the plant safety analyses input values to give the est minimum DNBR. The DNBR limit for Standard Thermal Design Procedure is the appropriate B correlation limits increased to give sufficient margins to cover any DNBR penalties associated the analysis.

preventing DNB, adequate heat transfer is provided from the fuel clad to the reactor coolant, eby preventing clad damage as a result of inadequate cooling. Maximum fuel rod surface perature is not a design basis, since it is within a few degrees of coolant temperature during ration in the nucleate boiling region. Limits provided by the nuclear control and protection ems are such that this design basis is met for transients associated with Condition II events uding overpower transients. There is an additional large DNBR margin at rated power operation during normal operating transients.

1.2 Fuel Temperature Design Basis 1.2.1 Design Basis ing modes of operation associated with Condition I and Condition II events, there is at least a percent probability at a 95-percent confidence level that the peak kW/ft fuel rods will not exceed uranium dioxide melting temperature. The melting temperature of uranium dioxide is 5080°F ference 3) unirradiated and decreasing 58°F per 10,000 MWD/MTU. By precluding uranium ide melting, the fuel geometry is preserved and possible adverse effects of molten uranium ide on the cladding are eliminated. Design evaluations for Condition I and II events have shown fuel melting will not occur for achievable local burnups up to 75,000 MWD/MTU (Reference 81).

NRC has approved design evaluations up to 60,000 MWD/MTU in Reference 81 and up to 00 MWD/MTU in References 9 and 88.

1.2.2 Discussion l rod thermal evaluations are performed at rated power, at maximum overpower, and during sients at various burnups. These analyses confirm that this design basis and the fuel integrity ign bases given in Section 4.2 are met. They also provide input for the evaluation of Condition III IV events given in Chapter 15.

center-line temperature limit has been applied to reload cores with a lead rod average burnup of o 60,000 MWD/MTU. For higher burnups, the peak kilowatt-per-foot experienced during 4.4-2 Revision 1

1.3 Core Flow Design Basis 1.3.1 Design Basis ical minimum value of 94.1 percent of the thermal flow rate is assumed to pass through the fuel region of the core and is effective for fuel rod cooling. Coolant flow through the thimble and rumentation tubes and the leakage between the core barrel and core shroud, head cooling flow, leakage to the vessel outlet nozzles are not considered effective for heat removal.

1.3.2 Discussion e cooling evaluations are based on the thermal flow rate (minimum flow) entering the reactor sel. A typical maximum value of 5.9 percent of this value is allotted as bypass flow. This includes cluster control guide thimble and instrumentation tube cooling flow, leakage between the core el and the core shroud, head cooling flow, and leakage to the vessel outlet nozzles. The shroud cavity flow is considered as active flow that is effective for fuel rod cooling.

maximum bypass flow fraction of 5.9 percent assumes the use of thimble plugging devices in the cluster control guide thimble tubes that do not contain any other core components.

1.4 Hydrodynamic Stability Design Basis es of operation associated with Condition I and II events do not lead to hydrodynamic instability.

1.5 Other Considerations design bases described in Subsections 4.4.1 through 4.4.1.4 together with the fuel clad and fuel embly design bases given in Subsection 4.2.1 are sufficiently comprehensive that additional limits not required.

l rod diametral gap characteristics, moderator coolant flow velocity and distribution, and erator void are not inherently limiting. Each of these parameters is incorporated into the thermal hydraulic models used to confirm that the above-mentioned design criteria are met. For instance, fuel rod diametral gap characteristics change with time, as described in Subsection 4.2.3, and the rod integrity is evaluated on that basis. The effect of the moderator flow velocity and distribution cribed in Subsection 4.4.2.2 and the moderator void distribution described in Subsection 4.4.2.4 included in the core thermal evaluation and thus affect the design basis.

ting the fuel clad integrity criteria covers the possible effects of clad temperature limitations. Clad ace temperature limits are imposed on Condition I and Condition II operation to preclude ditions of accelerated oxidation. A clad temperature limit is applied to the loss-of-coolant accident cribed in Subsection 15.6.5; control rod ejection accident described in Subsection 15.4.8; and ed rotor accident described in Subsection 15.3.3.

2 Description of Thermal and Hydraulic Design of the Reactor Core 2.1 Summary Comparison le 4.4-1 provides a comparison of the design parameters for the AP1000, the AP600, and a nsed Westinghouse-designed plant using XL Robust fuel. For the comparison with a plant 4.4-3 Revision 1

2.2 Critical Heat Flux Ratio or DNBR and Mixing Technology minimum DNBRs for the rated power and anticipated transient conditions are given in le 4.4-1. The minimum DNBR in the limiting flow channel is typically downstream of the peak heat location (hotspot) due to the increased downstream enthalpy rise.

BRs are calculated by using the correlation and definitions described in Subsections 4.4.2.2.1 4.4.2.2.2. The VIPRE-01 computer code described in Subsection 4.4.4.5, is used to determine flow distribution in the core and the local conditions in the hot channel for use in the DNB elation. The use of hot channel factors is described in Subsections 4.4.4.3.1 (nuclear hot channel ors) and 4.4.2.2.4 (engineering hot channel factors).

2.2.1 DNB Technology primary DNB correlation used for the analysis of the AP1000 fuel is the WRB-2M correlation ferences 82 and 82a). The WRB-2M correlation applies to the Robust Fuel Assemblies, which are ned to be used in the AP1000 core. This correlation applies to most AP1000 conditions.

rrelation limit of 1.14 is applicable for the WRB-2M correlation.

applicable range of parameters for the WRB-2M correlation is:

ssure 1495 P 2425 psia al mass velocity 0.97 Gloc/106 3.1 lb/ft2-hr al quality -0.1 Xloc 0.29 ated length, inlet to CHF location LH 14 feet d spacing 10 gsp 20.6 inches uivalent hydraulic diameter 0.37 De 0.46 inches uivalent heated hydraulic diameter 0.46 Dh 0.54 inches WRB-2 (Reference 4) or W-3 (References 5 and 6) correlation is used wherever the B-2M correlation is not applicable. The WRB-2 correlation limit is 1.17.

applicable range of parameters for the WRB-2 correlation is:

ssure 1440 P 2490 psia al mass velocity 0.9 Gloc/106 3.7 lb/ft2-hr al quality -0.1 Xloc 0.3 at length, inlet to DNB location Lh 14 feet d spacing 10 gsp < 26 inches uivalent hydraulic diameter 0.37 De 0.51 inches uivalent heated hydraulic diameter 0.46 Dh 0.59 inches WRB-2 correlation was developed based on mixing vane data and, therefore, is only applicable e heated rod spans above the first mixing vane grid.

4.4-4 Revision 1

range of the primary correlation. For system pressures in the range of 500 to 1000 psia, the correlation limit is 1.45 (Reference 7). For system pressures greater than 1000 psia, the correlation limit is 1.30. The pressures associated with some of the steam line break statepoints in the range of 300 to 500 psia. Using additional information, the W-3 correlation is shown to be licable with these pressures and a correlation limit of 1.45.

ld wall factor, described in WCAP-7695-L (Reference 8), is applied to the W-3 DNB correlation to servatively account for the presence of the unheated thimble surfaces.

2.2.2 Definition of DNBR DNB heat flux ratio, DNBR, as applied to typical cells (flow cells with all walls heated) and ble cells (flow cells with heated and unheated walls) is defined as:

q "DNB, predicted DNBR =

q"actual re:

q q "DNB, predicted = "WRB 2 q "WRB 2 M or NB, predicted =

F F RB-2M = the uniform DNB heat flux as predicted by the WRB-2M DNB correlation RB-2 = the uniform DNB heat flux as predicted by the WRB-2 DNB correlation

= the flux shape factor to account for nonuniform axial heat flux distributions (Reference 10) with the term C modified as in Reference 5 ctual = the actual local heat flux DNBR as applied to the W-3 DNB correlation is:

q "predicted DNBR =

q "actual re:

q x CWF q "predicted = "EU W 3 F

U-W-3 = the uniform DNB heat flux as predicted by the W-3 DNB correlation (Reference 5)

CWF = 1.0-Ru [T]

re:

G T = 13.76 - 1.372e1.78x -4.732 ( 6

)0.0535 -0.0619 ( P )0.14 8.509 D0.017 10 h 1000

= 1-De/Dh 4.4-5 Revision 1

2.2.3 Mixing Technology rate of heat exchange by mixing between flow channels is proportional to the difference in the l mean fluid enthalpy of the respective channels, the local fluid density, and the flow velocity. The portionality is expressed by the dimensionless thermal diffusion coefficient (TDC) which is defined w

TDC =

Va re:

= flow exchange rate per unit length (lbm/ft-s)

= fluid density (lbm/ft3)

= fluid velocity (ft/s)

= lateral flow area between channels per unit length (ft2/ft) application of the thermal diffusion coefficient in the VIPRE-01 analysis for determining the rall mixing effect or heat exchange rate is presented in Reference 83.

iscussed in WCAP-7941-P-A (Reference 12) those series of tests, using the R mixing vane grid ign on 13-, 26-, and 32-inch grid spacing, were conducted in pressurized water loops at Reynolds bers similar to that of a pressurized water reactor core under the following single- and two-phase cooled boiling) flow conditions:

Pressure 1500 to 2400 psia Inlet temperature 332 to 642°F Mass velocity 1.0 to 3.5 x 106 lbm/hr-ft2 Reynolds number 1.34 to 7.45 x 105 Bulk outlet quality -52.1 to -13.5 percent thermal diffusion coefficient is determined by comparing the THINC code predictions with the sured subchannel exit temperatures. Data for 26-inch axial grid spacing are presented in re 4.4-1, where the thermal diffusion coefficient is plotted versus the Reynolds number. The mal diffusion coefficient is found to be independent of the Reynolds number, mass velocity, sure, and quality over the ranges tested. The two-phase data (local, subcooled boiling) falls in the scatter of the single-phase data. The effect of two-phase flow on the value of the thermal sion coefficient is demonstrated in WCAP-7941-P-A (Reference 12), by Rowe and Angle ferences 13 and 14), and Gonzalez-Santalo and Griffith (Reference 15). In the subcooled boiling on, the values of the thermal diffusion coefficient are indistinguishable from the single-phase es. In the quality region, Rowe and Angle show that in the case with rod spacing similar to that in surized water reactor core geometry, the value of the thermal diffusion coefficient increased with lity to a point and then decreased, but never below the single-phase value. Gonzalez-Santalo and fith show that the mixing coefficient increased as the void fraction increased.

4.4-6 Revision 1

ference 83). A mixing test program similar to the one just described was conducted for the current 17 geometry and mixing vane grids on 26-inch spacing, as described in WCAP-8298-P-A ference 16). The mean value of the thermal diffusion coefficient obtained from these tests is 9.

inclusion of intermediate flow mixer grids in the upper spans of the fuel assembly results in a grid cing of approximately 10 inches giving higher values of the thermal diffusion coefficient. A servative value of the thermal diffusion coefficient, .038, is used to determine the effect of coolant ng in the core thermal performance analysis.

2.2.4 Hot Channel Factors total hot channel factors for heat flux and enthalpy rise are defined as the maximum-to-core-rage ratios of these quantities. The heat flux hot channel factor considers the local maximum ar heat generation rate at a point (the hotspot), and the enthalpy rise hot channel factor involves maximum integrated value along a channel (the hot channel).

h of the total hot channel factors is composed of a nuclear hot channel factor, Subsection 4.4.4.3, cribing the neutron power distribution and an engineering hot channel factor, which allows for ations in flow conditions and fabrication tolerances. The engineering hot channel factors are e up of subfactors which account for the influence of the variations of fuel pellet diameter, sity, enrichment, and eccentricity; inlet flow distribution; flow redistribution; and flow mixing.

E t Flux Engineering Hot Channel Factor, FQ heat flux engineering hot channel factor is used to evaluate the maximum linear heat generation in the core. This subfactor is determined by statistically combining the fabrication variations for pellet diameter, density, and enrichment. As shown in WCAP-8174 (Reference 17), no DNB alty needs be taken for the short, relatively low-intensity heat flux spikes caused by variations in above parameters, as well as fuel pellet eccentricity and fuel rod diameter variation.

E halpy Rise Engineering Hot Channel Factor, F H effect of variations in flow conditions and fabrication tolerances on the hot channel enthalpy rise rectly considered in the VIPRE-01 core thermal subchannel analysis, described in section 4.4.4.5.1 under any reactor opening condition. The following items are considered as tributors to the enthalpy rise engineering hot channel factor:

Pellet diameter, density, and enrichment Variations in pellet diameter, density, and enrichment are considered statistically in establishing the limit DNBRs, described in Subsection 4.4.1.1.2, for the Revised Thermal Design Procedure (Reference 2). Uncertainties in these variables are determined from sampling of manufacturing data.

Inlet flow maldistribution The consideration of inlet flow maldistribution in core thermal performances is described in Subsection 4.4.4.2.2. A design basis of five-percent reduction in coolant flow to the hot assembly is used in the VIPRE-01 analyses.

4.4-7 Revision 1

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-01 analyses for every operating condition evaluated.

Flow mixing The subchannel mixing model incorporated in the VIPRE-01 code and used in reactor design is based on experimental data, as detailed in WCAP-7667-P-A (Reference 18) and discussed in Subsections 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. The VIPRE-01 mixing model is discussed in Reference 83.

2.2.5 Effects of Rod Bow on DNBR phenomenon of fuel rod bowing, as described in WCAP-8691 (Reference 19), is accounted for in DNBR safety analysis of Condition I and Condition II events for each plant application. Applicable eric credits for margin resulting from retained conservatism in the evaluation of DNBR and/or gin obtained form measured plant operating parameters (such as FNH or core flow), which are limiting than those required by the plant safety analysis, can be used to offset the effect of rod the safety analysis of the AP1000, sufficient DNBR margin was maintained, as described in section 4.4.1.1.2, to accommodate the full and low flow rod bow DNBR penalties identified in erence 20. The referenced penalties are applicable to the analyses using the WRB-2M or WRB-2 B correlations.

maximum rod bow penalties (less than about 2 percent DNBR) accounted for in the design ty analysis are based on an assembly average burnup of 24,000 MWD/MTU. At burnups greater 24,000 MWD/MTU, credit is taken for the effect of FNH burndown, due to the decrease in onable isotopes and the buildup of fission product inventory, and no additional rod bow penalty is uired (Reference 21).

e upper spans of the fuel assembly, additional restraint is provided with the intermediate flow er grids such that the grid-to-grid spacing in those spans with intermediate flow mixer grids is roximately 10 inches compared to approximately 20 inches in the other spans. Using the NRC roved scaling factor [see WCAP 8691 (Reference 19) and Reference 21], results in predicted nnel closure in the limiting 10 inch spans of less than 50 percent closure. Therefore, no rod bow BR penalty is required in the 10 inch spans in the safety analyses.

2.3 Linear Heat Generation Rate core average and maximum linear heat generation rates are given in Table 4.4-1. The method of rmining the maximum linear heat generation rate is given in Subsection 4.3.2.2.

2.4 Void Fraction Distribution calculated core average and the hot subchannel maximum and average void fractions are ented in Table 4.4-2 for operation at full power. The void models used in the VIPRE-W code are cribed in Subsection 4.4.2.7.3.

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lysis. Extensive experimental verification of VIPRE-01 is presented in Reference 84.

2.6 Core Pressure Drops and Hydraulic Loads 2.6.1 Core Pressure Drops analytical model and experimental data used to calculate the pressure drops shown in le 4.4-1 are described in Subsection 4.4.2.7. The core pressure drop includes the fuel assembly, er core plate, and upper core plate pressure drops. The full-power operation pressure drop values wn in Table 4.4-1 are the unrecoverable pressure drops across the vessel, including the inlet and et nozzles, and across the core. These pressure drops are based on the best-estimate flow for al plant operating conditions as described in Subsection 5.1.4. This subsection also defines and cribes the thermal design flow (minimum flow) that is the basis for reactor core thermal ormance and the mechanical design flow (maximum flow) that is used in the mechanical design e reactor vessel internals and fuel assemblies. Since the best-estimate flow is that flow which is t likely to exist in an operating plant, the calculated core pressure drops in Table 4.4-1 are based his best-estimate flow rather than the thermal design flow.

uncertainties associated with the core pressure drop values are presented in section 4.4.2.9.2.

2.6.2 Hydraulic Loads re 4.2-2 shows the fuel assembly hold-down springs. These springs are designed to keep the assemblies in contact with the lower core plate under Condition I and II events, except for the ine overspeed transient associated with a loss of external load. The hold-down springs are igned to tolerate the possibility of an overdeflection associated with fuel assembly lift-off for this e and to provide contact between the fuel assembly and the lower core plate following this sient. More adverse flow conditions occur during a loss-of-coolant accident. These conditions are ented in Subsection 15.6.5.

raulic loads at normal operating conditions are calculated considering the best-estimate flow, cribed in Section 5.1, and accounting for the minimum core bypass flow based on manufacturing rances. Core hydraulic loads at cold plant startup conditions are based on the cold best-estimate

, but are adjusted to account for the coolant density difference. Conservative core hydraulic loads a pump overspeed transient, which could possibly create a flow rate 18-percent greater than the t estimate flow, are evaluated to be approximately twice the fuel assembly weight.

raulic verification tests for the fuel assembly are described in Reference 86.

2.7 Correlation and Physical Data 2.7.1 Surface Heat Transfer Coefficients ced convection heat transfer coefficients are obtained from the Dittus-Boelter correlation ference 24), with the properties evaluated at bulk fluid conditions:

0.8 0.4 hD e D G Cp

= 0.023 e K K 4.4-9 Revision 1

= equivalent diameter (ft)

= thermal conductivity (Btu/h-ft-°F)

= mass velocity (lbm/h-ft2)

= dynamic viscosity (lbm/ft-h)

= heat capacity (Btu/lb-°F) correlation has been shown to be conservative (Reference 25) for rod bundle geometries with h-to-diameter ratios in the range used by pressurized water reactors.

onset of nucleate boiling occurs when the clad wall temperature reaches the amount of erheat predicted by Thoms correlation (Reference 26). After this occurrence, the outer clad wall perature is determined by:

Tsat = [0.072exp(-P/1260)](q)0.5 re:

at = wall superheat, TW - Tsat (°F)

= wall heat flux (Btu/h-ft2)

= pressure (psia)

= outer clad wall temperature (°F)

= saturation temperature of coolant at pressure P (°F) 2.7.2 Total Core and Vessel Pressure Drop ecoverable pressure losses occur as a result of viscous drag (friction) and/or geometry changes m) in the fluid flow path. The flow field is assumed to be incompressible, turbulent, single-phase er. Those assumptions apply to the core and vessel pressure drop calculations for the purpose of blishing the primary loop flow rate. Two-phase considerations are neglected in the vessel sure drop evaluation because the core average void is negligible, as shown in Table 4.4-2. Two-se flow considerations in the core thermal subchannel analysis are considered and the models described in Subsection 4.4.4.2.3. Core and vessel pressure losses are calculated by equations e form:

L V2 PL = (K + f )

De 2 gc (144) 4.4-10 Revision 1

= fluid density (lbm/ft3)

= length (ft)

= equivalent diameter (ft)

= fluid velocity (ft/s)

= 32.174 (lbm-ft/lbf p-s2)

= form loss coefficient (dimensionless)

= friction loss coefficient (dimensionless) d density is assumed to be constant at the appropriate value for each component in the core and sel. Because of the complex core and vessel flow geometry, precise analytical values for the form friction loss coefficients are not available. Therefore, experimental values for these coefficients obtained from geometrically similar models.

es are quoted in Table 4.4-1 for unrecoverable pressure loss across the reactor vessel, including inlet and outlet nozzles, and across the core. The results of full-scale tests of core components fuel assemblies are used in developing the core pressure loss characteristic.

s of the primary coolant loop flow rates are made prior to initial criticality as described in section 4.4.5.1, to verify that the flow rates used in the design, which are determined in part from pressure losses calculated by the method described here, are conservative. See Section 14.2 for operational testing.

2.7.3 Void Fraction Correlation RE-01 considers two-phase flow in two steps. First, a quality model is used to compute the ing vapor mass fraction (true quality) including the effects of subcooled boiling. Then, given the void quality, a bulk void model is applied to compute the vapor volume fraction (void fraction).

RE-01 uses a profile fit model (Reference 83) for determining subcooled quality. It calculates the l vapor volumetric fraction in forced convection boiling by: 1) predicting the point of bubble arture from the heated surface and 2) postulating a relationship between the true local vapor tion and the corresponding thermal equilibrium value.

void fraction in the bulk boiling region is predicted by using homogeneous flow theory and uming no slip. The void fraction in this region is therefore a function only of the thermodynamic lity.

2.8 Thermal Effects of Operational Transients B core safety limits are generated as a function of coolant temperature, pressure, core power, and l power imbalance. Steady-state operation within these safety limits provides that the DNB ign basis is met. Subsection 15.0.6 discusses the overtemperature T trip (based on DNBR limit) us Tavg. This system provides protection against anticipated operational transients that are slow respect to fluid transport delays in the primary system. In addition, for fast transients (such as 4.4-11 Revision 1

2.9 Uncertainties in Estimates 2.9.1 Uncertainties in Fuel and Clad Temperatures described in Subsection 4.4.2.11, the fuel temperature is a function of crud, oxide, clad, pellet-gap, and pellet conductances. Uncertainties in the fuel temperature calculation are essentially of types: fabrication uncertainties, such as variations in the pellet and clad dimensions and the et density; and model uncertainties, such as variations in the pellet conductivity and the gap ductance. These uncertainties have been quantified by comparison of the thermal model to the ile thermocouple measurements (References 30 through 36), by out-of-pile measurements of the and clad properties (References 37 through 48), and by measurements of the fuel and clad ensions during fabrication. The resulting uncertainties are then used in the evaluations involving fuel temperature. The effect of densification on fuel temperature uncertainties is also included in calculation of the total uncertainty.

ddition to the temperature uncertainty described above, the measurement uncertainty in rmining the local power and the effect of density and enrichment variations on the local power considered in establishing the heat flux hot channel factor. These uncertainties are described in section 4.3.2.2.1.

ctor trip setpoints, as specified in the technical specifications, include allowance for instrument measurement uncertainties such as calorimetric error, instrument drift and channel oducibility, temperature measurement uncertainties, noise, and heat capacity variations.

ertainty in determining the cladding temperature results from uncertainties in the crud and oxide knesses. Because of the excellent heat transfer between the surface of the rod and the coolant, film temperature drop does not appreciably contribute to the uncertainty.

2.9.2 Uncertainties in Pressure Drops e and vessel pressure drops based on the best-estimate flow, as described in Section 5.1, are ted in Table 4.4-1. The uncertainties quoted are based on the uncertainties in both the test results the analytical extension of these values to the reactor application.

ajor use of the core and vessel pressure drops is to determine the primary system coolant flow s, as described in Section 5.1. In addition, as described in Subsection 4.4.5.1, tests on primary em prior to initial criticality, are conducted to verify that a conservative primary system coolant rate has been used in the design and analysis of the plant.

2.9.3 Uncertainties Due to Inlet Flow Maldistribution effects of uncertainties in the inlet flow maldistribution criteria used in the core thermal analyses described in Subsection 4.4.4.2.2.

2.9.4 Uncertainty in DNB Correlation uncertainty in the DNB correlation described in Subsection 4.4.2.2, is written as a statement on probability of not being in DNB based on the statistics of the DNB data. This is described in section 4.4.2.2.2.

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section 4.4.4.5.1, due to uncertainties in the nuclear peaking factors are accounted for by lying conservatively high values of the nuclear peaking factors. Measurement error allowances included in the statistical evaluation of the limit DNBR described in Subsection 4.4.1.1 using the ised Thermal Design Procedure. More information is provided in WCAP-11397-P-A ference 2). In addition, conservative values for the engineering hot channel factors are used as ented in Subsection 4.4.2.2.4. The results of a sensitivity study, WCAP-8054-P-A ference 22), with THINC-IV, a VIPRE-01 equivalent code, show that the minimum DNBR in the channel is relatively insensitive to variations in the core-wide radial power distribution (for the e value of FNH ).

ability of the VIPRE-01 computer code to accurately predict flow and enthalpy distributions in rod dles is discussed in Subsection 4.4.4.5.1 and in Reference 83. Studies (Reference 84) have n performed to determine the sensitivity of the minimum DNBR to the void fraction correlation also Subsection 4.4.2.7.3) and the inlet flow distributions. The results of these studies show that minimum DNBR is relatively insensitive to variation in these parameters. Furthermore, the RE-01 flow field model for predicting conditions in the hot channels is consistent with that used in derivation of the DNB correlation limits including void/quality modeling, turbulent mixing and sflow and two phase flow (Reference 83).

2.9.6 Uncertainties in Flow Rates uncertainties associated with reactor coolant loop flow rates are discussed in Section 5.1. A mal design flow is defined for use in core thermal performance evaluations accounting for both diction and measurement uncertainties. In addition, another 5.9 percent of the thermal design flow ssumed to be ineffective for core heat removal capability because it bypasses the core through various available vessel flow paths described in Subsection 4.4.4.2.1.

2.9.7 Uncertainties in Hydraulic Loads described in Subsection 4.4.2.6.2, hydraulic loads on the fuel assembly are evaluated for a pump rspeed transient which creates flow rates 18 percent greater than the best estimate flow. The best mate flow is the most likely flow rate value for the actual plant operating condition.

2.9.8 Uncertainty in Mixing Coefficient nservative value of the mixing coefficient, that is, the thermal diffusion coefficient, is used in the RE-01 analyses.

2.10 Flux Tilt Considerations nificant quadrant power tilts are not anticipated during normal operation since this phenomenon is sed by some asymmetric perturbation. A dropped or misaligned rod cluster control assembly ld cause changes in hot channel factors. These events are analyzed separately in Chapter 15.

er possible causes for quadrant power tilts include X-Y xenon transients, inlet temperature matches, enrichment variations within tolerances, and so forth.

ddition to unanticipated quadrant power tilts as described above, other readily explainable mmetries may be observed during calibration of the ex-core detector quadrant power tilt alarm.

ing operation, in-core maps are taken at least one per month and additional maps are obtained 4.4-13 Revision 1

mmetry in the core, from quadrant to quadrant, is frequently a consequence of the design when embly and/or component shuffling and rotation requirements do not allow exact symmetry ervation. In each case, the acceptability of an observed asymmetry, planned or otherwise, ends solely on meeting the required accident analyses assumptions. In practice, once eptability has been established by review of the incore maps, the quadrant power tilt alarms and ted instrumentation are adjusted to indicate zero quadrant power tilt ratio as the final step in the bration process. This action confirms that the instrumentation is correctly calibrated to alarm in event an unexplained or unanticipated change occurs in the quadrant-to-quadrant relationships ween calibration intervals.

per functioning of the quadrant power tilt alarm is significant. No allowances are made in the ign for increased hot channel factors due to unexpected developing flux tilts, since likely causes presented by design or procedures or are specifically analyzed.

lly, in the event that unexplained flux tilts do occur, the Technical Specifications provide ropriate corrective actions to provide continued safe operation of the reactor.

2.11 Fuel and Cladding Temperatures sistent with the thermal-hydraulic design bases described in Subsection 4.4.1, the following ussion pertains mainly to fuel pellet temperature evaluation. A description of fuel clad integrity is ented in Subsection 4.2.3.1.

thermal-hydraulic design provides that the maximum fuel temperature is below the melting point ranium dioxide, Subsection 4.4.1.2. To preclude center melting and to serve as a basis for rpower protection system setpoints, a calculated center-line fuel temperature of 4700°F is cted as the overpower limit. This provides sufficient margin for uncertainties in the thermal luations, as described in Subsection 4.4.2.9.1. The temperature distribution within the fuel pellet redominantly a function of the local power density and the uranium dioxide thermal conductivity.

ever, the computation of radial fuel temperature distributions combines crud, oxide, clad gap, pellet conductances. The factors which influence these conductances, such as gap size (or tact pressure), internal gas pressure, gas composition, pellet density, and radial power ribution within the pellet, have been combined into a semi-empirical thermal model, discussed in section 4.2.3.3, that includes a model for time-dependent fuel densification, as given in AP-10851-P-A (Reference 49) and WCAP-15063-P-A, Revision 1 (Reference 85). This thermal el enables the determination of these factors and their net effects on temperature profiles. The perature predictions have been compared to in-pile fuel temperature measurements ferences 30 through 36, 50 and 85) and melt radius data (References 51 and 52) with good lts.

l rod thermal evaluations (fuel centerline, average and surface temperatures) are performed at eral times in the fuel rod lifetime (with consideration of time-dependent densification) to determine maximum fuel temperatures.

principal factors employed in the determination of the fuel temperature follow.

2.11.1 Uranium Dioxide Thermal Conductivity thermal conductivity of uranium dioxide was evaluated from data reported in References 37 ugh 48 and 53. At the higher temperatures, thermal conductivity is best obtained by using the 4.4-14 Revision 1

Kdt = 93 W/cm o

conclusion is based on the integral values reported in References 51 and 53 through 57.

design curve for the thermal conductivity is shown in Figure 4.4-2. The section of the curve at peratures between 0° and 1300°C is in agreement with the recommendation of the International mic Energy Agency (IAEA) panel (Reference 58). The section of the curve above 1300°C is ved for an integral value of 93 W/cm. (References 51, 53, and 57).

rmal conductivity for uranium dioxide at 95-percent theoretical density can be represented by the wing equation:

1 K= + 8.775 x 1013 T3 11.8 + 0.0238 T re:

= W/cm-°C

= °C.

2.11.2 Radial Power Distribution in Uranium Dioxide Fuel Rods accurate description of the radial power distribution as a function of burnup is needed for rmining the power level for incipient fuel melting and other important performance parameters, h as pellet thermal expansion, fuel swelling, and fission gas release rates. Radial power ribution in uranium dioxide fuel rods is determined with the neutron transport theory code, ER. The LASER code has been validated by comparing the code predictions on radial burnup isotopic distributions with measured radial microdrill data, as detailed in WCAP-6069 ference 59) and WCAP-3385-56 (Reference 60). A radial power depression factor, f, is rmined using radial power distributions predicted by LASER. The factor, f, enters into the rmination of the pellet centerline temperature, Tc, relative to the pellet surface temperature, Tg, ugh the expression:

Tc q" f K(T) dT = 4 Ti re:

) = the thermal conductivity for uranium dioxide with a uniform density distribution

= the linear power generation rate 2.11.3 Gap Conductance temperature drop across the pellet-clad gap is a function of the gap size and the thermal ductivity of the gas in the gap. The gap conductance model is selected so that when combined the uranium dioxide thermal conductivity model, the calculated fuel center-line temperature 4.4-15 Revision 1

2.11.4 Surface Heat Transfer Coefficients fuel rod surface heat transfer coefficients during subcooled forced convection and nucleate ng are presented in Subsection 4.4.2.7.1.

2.11.5 Fuel Clad Temperatures outer surface of the fuel rod at the hotspot operates at a temperature a few degrees above fluid perature for steady-state operation at rated power throughout core life due to the onset of leate boiling. At beginning of life this temperature is the same as the clad metal outer surface.

ing operation over the life of the core, the buildup of oxides and crud on the fuel rod surface ses the clad surface temperature to increase. Allowance is made in the fuel center melt luation for this temperature rise. Since the thermal-hydraulic design basis limits DNB, adequate t transfer is provided between the fuel clad and the reactor coolant so that the core thermal output ot limited by considerations of clad temperature.

2.11.6 Treatment of Peaking Factors total heat flux hot channel factor, FQ, is defined by the ratio of the maximum-to-core-average t flux. The design value of FQ, as presented in Table 4.3-2 and described in Subsection 4.3.2.2.6, 6 for normal operation.

described in Subsection 4.3.2.2.6, the peak linear power resulting from overpower transients/

rator errors (assuming a maximum overpower of 118 percent) is less than or equal to 22.45 kW/ft.

centerline fuel temperature must be below the uranium dioxide melt temperature over the me of the rod, including allowances for uncertainties. The fuel temperature design basis is cribed in Subsection 4.4.1.2 and results in a maximum allowable calculated center-line perature of 4700°F. The peak linear power for prevention of center-line melt is 22.5 kW/ft. The ter-line temperature at the peak linear power resulting from overpower transients/operator errors uming a maximum overpower of 118 percent) is below that required to produce melting.

3 Description of the Thermal and Hydraulic Design of the Reactor Coolant System 3.1 Plant Configuration Data nt configuration data for the thermal-hydraulic and fluid systems external to the core are provided ppropriate in Chapters 5, 6, and 9. Areas of interest are as follows:

Total coolant flow rates for the reactor coolant system and each loop are provided in Table 5.1-3. Flow rates employed in the evaluation of the core are presented throughout Section 4.4.

Total reactor coolant system volume including pressurizer and surge line and reactor coolant system liquid volume, including pressurizer water at steady-state power conditions, are given in Table 5.1-2.

The flow path length through each volume may be calculated from physical data provided in Table 5.1-2.

4.4-16 Revision 1

The parameters for components of the reactor coolant system are presented in Section 5.4.

The steady-state pressure drops and temperature distributions through the reactor coolant system are presented in Table 5.1-1.

3.2 Operating Restrictions on Pumps minimum net positive suction head is established before operating the reactor coolant pumps.

operator verifies that the system pressure satisfies net positive suction head requirements prior perating the pumps.

3.3 Power-Flow Operating Map (Boiling Water Reactor BWR) subsection is not applicable to AP1000.

3.4 Temperature-Power Operating Map (PWR) relationship between reactor coolant system temperature and power is a linear relationship ween zero and 100-percent power.

effects of reduced core flow due to inoperative pumps is described in Subsections 5.4.1 15.2.6 and Section 15.3. The AP1000 does not include power operation with one pump out of ice. Natural circulation capability of the system is described in Subsection 5.4.2.3.2.

3.5 Load Following Characteristics d follow using control rod and gray rod motion is described in Subsection 4.3.2.4.16. The reactor er is controlled to maintain average coolant temperature at a value which is a linear function of

, as described in Section 7.7.

3.6 Thermal and Hydraulic Characteristics Summary Table thermal and hydraulic characteristics are given in Tables 4.1-1, 4.4-1, and 4.4-2.

4 Evaluation 4.1 Critical Heat Flux critical heat flux correlations used in the core thermal analysis are explained in Subsection 4.4.2.

4.2 Core Hydraulics 4.2.1 Flow Paths Considered in Core Pressure Drop and Thermal Design following flow paths for core bypass are considered:

Flow through the spray nozzles into the upper head for head cooling purposes Flow entering into the rod cluster control and gray rod cluster guide thimbles 4.4-17 Revision 1

Flow introduced through the core shroud for the purpose of cooling and not considered available for core cooling above contributions are evaluated to confirm that the design value of the core bypass flow is he total allowance, one part is associated with the core and the remainder is associated with the rnals (items A, C, and D above). Calculations have been performed using drawing tolerances in worst direction and accounting for uncertainties in pressure losses. Based on these calculations, core bypass is no greater than the 5.9 percent design value.

w model test results for the flow path through the reactor are described in Subsection 4.4.2.7.2.

4.2.2 Inlet Flow Distributions re inlet flow distribution reduction of five percent to the hot assembly inlet is used in the RE-01 analyses of DNBR in the AP1000 core. Studies shown in WCAP-8054-P-A ference 22), made with THINC-IV, a VIPRE-01 equivalent code, show that flow distributions ificantly more nonuniform than five percent have a very small effect on DNBR, which is ounted for in the DNB analysis.

4.2.3 Empirical Friction Factor Correlations friction factor for VIPRE-01 in the axial direction, parallel to the fuel rod axis, is evaluated using a elation for a smooth tube (Reference 83). The effect of two-phase flow on the friction loss is ressed in terms of the single-phase friction pressure drop and a two-phase friction multiplier. The tiplier is calculated using the homogenous equilibrium flow model.

flow in the lateral directions, normal to the fuel rod axis, views the reactor core as a large tube

k. Thus, the lateral friction factor proposed by Idel'chik (Reference 64) is applicable. This elation is of the form:

FL = A Re L0.2 re:

= a function of the rod pitch and diameter as given in Idel'chik (Reference 64)

= the lateral Reynolds number based on the rod diameter comparisons of predictions to data given in Reference 83 verify the applicability of the RE-01 correlations in PWR design.

4.3 Influence of Power Distribution core power distribution, which is largely established at beginning of life by fuel enrichment, ing pattern, and core power level, is also a function of variables such as control rod worth and ition, and fuel depletion through lifetime. Radial power distributions in various planes of the core often illustrated for general interest. However, the core radial enthalpy rise distribution, as rmined by the integral of power up each channel, is of greater importance for DNBR analyses.

4.4-18 Revision 1

4.3.1 Nuclear Enthalpy Rise Hot Channel Factor, FNH en the local power density q (kW/ft) at a point x, y, z in a core with N fuel rods and height H, then:

H Max q x o yo , zo) dz hot rod power o FNH = = H average rod power 1 q (x, y, z) dz N all rods o way in which FNH is used in the DNBR calculation is important. The location of minimum DNBR ends on the axial profile, and the value of DNBR depends on the enthalpy rise to that point.

ically, the maximum value of the rod integral power is used to identify the most likely rod for imum DNBR. An axial power profile is obtained that, when normalized to the design value of

, recreates the axial heat flux along the limiting rod. The surrounding rods are assumed to have same axial profile with rod average powers which are typical distributions found in hot emblies. In this manner, worst-case axial profiles can be combined with worst-case radial ributions for reference DNBR calculations.

ould be noted again that FNH is an integral and is used as such in DNBR calculations. Local heat es are obtained by using hot channel and adjacent channel explicit power shapes which take into ount variations in horizontal power shapes throughout the core.

operation at a fraction of full power, the design FNH used is given by:

H [1 + 0.3(1 P)]

FNH = FRTP re:

is the limit at rated thermal power (RTP):

the fraction of rated thermal power and FRTP H = 1.59.

permitted relaxation of FNH is included in the DNB protection setpoints and allows radial power pe changes with rod insertion to the insertion limits, as detailed in WCAP-7912-P-A ference 65). This allows greater flexibility in the nuclear design.

4.3.2 Axial Heat Flux Distributions described in Subsection 4.3.2.2, the axial heat flux distribution can vary as a result of rod motion ower change or as a result of a spatial xenon transient which may occur in the axial direction. The ore nuclear detectors, as described in Subsection 4.3.2.2.7, are used to measure the axial power alance. The information from the ex-core detectors is used to protect the core from excessive l power imbalance. The reference axial shape used in establishing core DNB limits (that is, rtemperature T protection system setpoints) is a chopped cosine with a peak-to-average value

.61. The reactor trip system provides automatic reduction of the trip setpoints on excessive axial er imbalance. To determine the magnitude of the setpoint reduction, the reference shape is plemented by other axial shapes skewed to the bottom and top of the core.

4.4-19 Revision 1

lant temperature, and power level changes.

initial conditions for the accidents for which DNB protection is required are assumed to be those missible within the specified axial offset control limits described in Subsection 4.3.2.2. In the case e loss-of-flow accident, the hot channel heat flux profile is very similar to the power density profile ormal operation preceding the accident. It is therefore possible to illustrate the calculated imum DNBR for conditions representative of the loss-of-flow accident as a function of the flux rence initially in the core. The power shapes are evaluated with a full-power radial peaking factor H ) of 1.59. The radial contribution to the hot rod power shape is conservative both for the initial dition and for the condition at the time of minimum DNBR during the loss-of-flow transient. The imum DNBR is calculated for the design power shape for non-overpower/overtemperature DNB nts. This design shape results in calculated DNBR that bounds the normal operation shapes.

4.4 Core Thermal Response neral summary of the steady-state thermal-hydraulic design parameters including thermal output flow rates is provided in Table 4.4-1.

tated in Subsection 4.4.1, the design bases of the application are to prevent DNB and to prevent melting for Condition I and II events. The protective systems described in Chapter 7 are igned to meet these bases. The response of the core to Condition II transients is given in pter 15.

4.5 Analytical Methods 4.5.1 Core Analysis objective of reactor core thermal design is to determine the maximum heat removal capability in ow subchannels and to show that the core safety limits, as presented in the technical cifications, are not exceeded while combining engineering and nuclear effects. The thermal ign takes into account local variations in dimensions, power generation, flow redistribution, and ng. The Westinghouse version of VIPRE-01, a three-dimensional subchannel code that has been eloped to account for hydraulic and nuclear effects on the enthalpy rise in the core and hot nnels, is described in Reference 83, VIPRE-01 modeling of a PWR core is based on a one-pass eling approach (Reference 83). In the one-pass modeling, hot channels and their adjacent nnels are modeled in detail, while the rest of the core is modeled simultaneously on a relatively rse mesh. The behavior of the hot assembly is determined by superimposing the power ribution upon the inlet flow distribution while allowing for flow mixing and flow distribution between channels. Local variations in fuel rod power, fuel rod and pellet fabrication, and turbulent mixing also considered in determining conditions in the hot channels. Conservation equations of mass, l and lateral momentum, and energy are solved for the fluid enthalpy, axial flow rate, lateral flow, pressure drop.

4.5.2 Steady State Analysis VIPRE-01 core model as approved by the NRC (Reference 83) is used with the applicable DNB elations to determine DNBR distributions along the hot channels of the reactor core under all ected operating conditions. The VIPRE-01 code is described in detail in Reference 84, including ussions on code validation with experimental data. The VIPRE-01 modeling method is described eference 83, including empirical models and correlations used. The effect of crud on the flow and alpy distribution in the core is not directly accounted for in the VIPRE-01 evaluations. However, 4.4-20 Revision 1

mates of uncertainties are discussed in Subsection 4.4.2.9.

4.5.3 Experimental Verification ensive additional experimental verification of VIPRE-01 is presented in Reference 84.

VIPRE-01 analysis is based on a knowledge and understanding of the heat transfer and rodynamic behavior of the coolant flow and the mechanical characteristics of the fuel elements.

use of the VIPRE-01 analysis provides a realistic evaluation of the core performance and is used e thermal hydraulic analyses as described above.

4.5.4 Transient Analysis RE-01 is capable of transient DNB analysis. The conservation equations in the VIPRE-01 code tain the necessary accumulation terms for transient calculations. The input description can ude one or more of the following time dependent arrays:

Inlet flow variation Core heat flux variation Core pressure variation Inlet temperature or enthalpy variation he beginning of the transient, the calculation procedure is carried out as in the steady state lysis. The time is incremented by an amount determined either by the user of by the time step trol options in the code itself. At each new time step the calculations are carried out with the ition of the accumulation terms which are evaluated using the information from the previous time

. This procedure is continued until a preset maximum time is reached.

me intervals selected by the user, a complete description of the coolant parameter distributions ell as DNBR is printed out. In this manner the variation of any parameter with time can be readily rmined.

4.6 Hydrodynamic and Flow Power Coupled Instability ing flow may be susceptible to thermohydrodynamic instabilities (Reference 68). These abilities are undesirable in reactors, since they may cause a change in thermohydraulic ditions that may lead to a reduction in the DNB heat flux relative to that observed during a steady condition or to undesired forced vibrations of core components. Therefore, a thermo-hydraulic ign criterion was developed which states that modes of operation under Condition I and II events ll not lead to thermohydrodynamic instabilities.

specific types of flow instabilities are considered for AP1000 operation. These are the Ledinegg low excursion) type of static instability and the density wave type of dynamic instability.

4.4-21 Revision 1

P l internal G

omes algebraically smaller than the loop supply (pump head) pressure drop-flow rate curve:

P l external G

criterion for stability is thus:

P P internal external G G reactor coolant pump head curve has a negative slope (P/G external less than zero),

reas the reactor coolant system pressure drop-flow curve has a positive slope (P/G internal ater than zero) over the Condition I and Condition II operational ranges. Thus, the Ledinegg ability does not occur.

mechanism of density wave oscillations in a heated channel has been described by R. T. Lahey F. J. Moody (Reference 69). Briefly, an inlet flow fluctuation produces an enthalpy perturbation.

perturbs the length and the pressure drop of the single-phase region and causes quality or void urbations in the two-phase regions that travel up the channel with the flow. The quality and length urbations in the two-phase region create two-phase pressure drop perturbations. However, since total pressure drop across the core is maintained by the characteristics of the fluid system rnal to the core, then the two-phase pressure drop perturbation feeds back to the single-phase on. These resulting perturbations can be either attenuated or self-sustained.

mple method has been developed by M. Ishii (Reference 70) for parallel closed-channel systems valuate whether a given condition is stable with respect to the density wave type of dynamic ability. This method had been used to assess the stability of typical Westinghouse reactor igns, including the design outlined in References 71, 72, and 73, under Condition I and II ration. The results indicate that a large margin-to-density wave instability exists. Increases on the er of 150 percent of rated reactor power would be required for the predicted inception of this type stability.

application of the Ishii method (Reference 70) to Westinghouse reactor designs is conservative to the parallel open-channel feature of Westinghouse pressurized water reactor cores. For such s, there is little resistance to lateral flow leaving the flow channels of high-power density. There is energy transfer from channels of high-power density to lower power density channels. This pling with cooler channels leads to the conclusion that an open-channel configuration is more le than the above closed-channel analysis under the same boundary conditions.

w stability tests (Reference 74) have been conducted where the closed channel systems were wn to be less stable than when the same channels were cross-connected at several locations.

cross-connections were such that the resistance to channel cross-flow and enthalpy urbations would be greater than would exist in a pressurized water reactor core which has a tively low resistance to cross-flow.

4.4-22 Revision 1

nnel stability experiments simulating a reactor core flow. These experiments were conducted at sures up to 2200 psia. The results showed that, for flow and power levels typical of power reactor ditions, no flow oscillations could be induced above 1200 psia.

itional evidence that flow instabilities do not adversely affect thermal margin is provided by the from the rod bundle DNB tests. Many Westinghouse rod bundles have been tested over wide ges of operating conditions with no evidence of premature DNB or inconsistent data which might ndicative of flow instabilities in the rod bundle.

ummary, it is concluded that thermohydrodynamic instabilities will not occur under Condition I II for Westinghouse pressurized water reactor designs. A large power margin, greater than percent of rated power, exists to predicted inception of such instabilities. Analysis has been ormed which shows that minor plant-to-plant differences in Westinghouse reactor designs such uel assembly arrays, power-to-flow ratios, and fuel assembly length do not result in gross rioration of the above power margins.

4.7 Fuel Rod Behavior Effects from Coolant Flow Blockage lant flow blockages can occur within the coolant channels of a fuel assembly or external to the tor core. The effects of fuel assembly blockage within the assembly on fuel rod behavior are e pronounced than external blockages of the same magnitude. In both cases, the flow blockages se local reductions in coolant flow. The amount of local flow reduction, where the reduction occurs e reactor, and how far along the flow stream the reduction persists are considerations which will ence the fuel rod behavior. The effects of coolant flow blockages in terms of maintaining rated performance are determined both by analytical and experimental methods. The experimental are usually used to augment analytical tools such as computer programs similar to the RE-01 program. Inspection of the DNB correlation (Subsection 4.4.2.2 and References 4, 5,

6) shows that the predicted DNBR is dependent upon the local values of quality and mass city.

VIPRE-01 code is capable of predicting the effects of local flow blockages on DNBR within the assembly on a subchannel basis, regardless of where the flow blockage occurs. Reference 84 ws that, for a fuel assembly similar to the Westinghouse design, VIPRE-01 accurately predicts the distribution within the fuel assembly when the inlet nozzle is completely blocked. Full recovery of flow was found to occur about 30 inches downstream of the blockage. With the reactor operating e nominal full-power conditions specified in Table 4.4-1, the effects of an increase in enthalpy decrease in mass velocity in the lower portion of the fuel assembly would not result in the fuel reaching the DNBR limit.

open literature supports the conclusion that flow blockage in open-lattice cores, similar to the stinghouse cores, causes flow perturbations which are local to the blockage. For example, hstubo and S. Uruwashi (Reference 76) show that the mean bundle velocity is approached mptomatically about four inches downstream from the flow blockage in a single flow cell. Similar lts were also found for two and three cells completely blocked. P. Basmer, et al., (Reference 77) ed an open-lattice fuel assembly in which 41 percent of the subchannels were completely blocked e center of the test bundle between spacer grids. Their results show that the stagnant zone ind the flow blockage essentially disappears after 1.65 L/De or about five inches for their test dle. They also found that leakage flow through the blockage tended to shorten the stagnant zone n essence, the complete recovery length. Thus, local flow blockages within a fuel assembly have effect on subchannel enthalpy rise. In reality, a local flow blockage would be expected to mote turbulence and, therefore would not likely affect DNBR at all.

4.4-23 Revision 1

s-flow velocity exceeds the limit established for fluid elastic stability, large amplitude whirling lts. The limits for a controlled vibration mechanism are established from studies of vortex dding and turbulent pressure fluctuations. The cross-flow velocity required to exceed fluid elastic ility limits is dependent on the axial location of the blockage and the characterization of the cross-(jet flow or not). These limits are greater than those for vibratory fuel rod wear. Cross-flow city above the established limits can lead to mechanical wear of the fuel rods at the grid support tions. Fuel rod wear due to flow-induced vibration is considered in the fuel rod fretting evaluation iscussed in Section 4.2.

5 Testing and Verification 5.1 Tests Prior to Initial Criticality actor coolant flow test is performed, as discussed in Chapter 14, following fuel loading but prior itial criticality. Coolant loop pressure data is obtained in this test. This data allows determination e coolant flow rates at reactor operating conditions. This test verifies that proper coolant flow s have been used in the core thermal and hydraulic analysis.

5.2 Initial Power and Plant Operation e power distribution measurements are made at several core power levels, as discussed in pter 14. These tests are used to confirm that conservative peaking factors are used in the core mal and hydraulic analysis.

itional demonstration of the overall conservatism of the THINC analysis was obtained by paring THINC predictions to in-core thermocouple measurements, as detailed WCAP-8453-A ference 78). VIPRE-01 has been confirmed to be as conservative as the THINC code in erence 83.

5.3 Component and Fuel Inspections ections performed on the manufactured fuel are described in Subsection 4.2.4. Fabrication surements critical to thermal and hydraulic analysis are obtained to verify that the engineering channel factors in the design analyses (Subsection 4.4.2.2.4) are met.

6 Instrumentation Requirements 6.1 Incore Instrumentation primary function of the incore instrumentation system is to provide a three-dimensional flux map e reactor core. This map is used to calibrate neutron detectors used by the protection and safety itoring system as well as to optimize core performance. A secondary function of the incore rumentation system is to provide the protection and safety monitoring system with the signals essary for monitoring core exit temperatures. This secondary function is the result of the hanical design that groups the detectors used for generating the flux map in the same thimble as core exit thermocouples.

incore instrumentation system consists of incore instrument thimble assemblies, which house d incore detectors, core exit thermocouple assemblies contained within an inner and outer sheath embly, and associated signal processing and data processing equipment. There are 42 incore 4.4-24 Revision 1

thimbles are inserted into the active core through the upper head and internals of the reactor sel. The signals output from the fixed incore detectors are digitized inside containment and tiplexed out of the containment. The signal processing software integral to the incore rumentation system allows the fixed incore detector signals to be used to calculate an accurate e-dimensional core power distribution suitable for developing calibration information for the ore nuclear instrumentation input to the overtemperature and overpower T reactor trip setpoints.

system is also capable of accurately determining whether the reactor power distribution is ently within the operating limits defined in the technical specifications while the reactor is rating above approximately 20 percent of rated thermal power.

incore instrument system data processor receives the transmitted digitized fixed incore detector als from the signal processor and combines the measured data with analytically-derived stants, and certain other plant instrumentation sensor signals, to generate a full three-ensional indication of nuclear power distribution in the reactor core. It also edits the three-ensional indication of power distribution to extract pertinent power distribution parameters outputs use by the plant operators and engineers. The data processor also generates hardcopy esentations of the detailed three-dimensional nuclear power indications.

hardware and software which performs the three-dimensional power distribution calculation are able of executing the calculation algorithms and constructing graphical and tabular displays of conditions at intervals of less than one minute. The software provides information to enable the tor operator to ascertain how the measured peaking factor performance agrees with the peaking or performance predicted by the design model used to determine the acceptability of the fuel ing pattern. The analysis software provides information required to activate a visual alarm display lert the reactor operator about the current existence of, or the potential for, reactor operating limit ations. The calculation algorithms are capable of determining the core average axial offset using a imum set of the total 42 incore monitor assemblies. A minimum set of incore monitor assemblies least 30 operating assemblies, with at least two operating assemblies in each quadrant, prior to lear model calibration; and at least 21 operating assemblies, with at least two operating emblies in each quadrant, after nuclear model calibration. The nuclear model calibration is ormed after each new core load. The hardware which performs the online power distribution itoring is configured such that a single hardware failure will not necessitate a reactor maximum er reduction or restrict normal reactor operations.

ing plant operation, the incore instrument thimble assembly is positioned within the fuel assembly exits through the top of the reactor vessel QuickLoc seal connection. The fixed incore detector core exit thermocouple signal exit the detector through a multipin connector to the incore rument thimble tube cables. The fixed incore detector and core exit thermocouple cables are then ed to different data conditioning and processing stations. The data is processed and the results available for display in the main control room.

6.2 Overtemperature and Overpower T Instrumentation overtemperature T trip protects the core against low DNBR. The overpower T trip protects inst excessive power (fuel rod rating protection).

described in Subsection 7.2.1.1.3, factors included in establishing the overtemperature T and rpower T trip setpoints include the reactor coolant temperature in each loop and the axial ribution of core power as seen by excore neutron detectors.

4.4-25 Revision 1

d to limit the maximum power output of the reactor within their respective ranges.

re are eight radial locations containing a total of twelve neutron flux detectors installed around the tor between the vessel and the primary shield. Four proportional counters for the source range located at the highest fluence portions of the core containing the primary startup sources at an ation approximately one-fourth of the core height. Four pulse fission chambers for the rmediate range, located in the same instrument wells as the source range detectors, are itioned at an elevation corresponding to one-half of the core height. Four uncompensated zation chamber assemblies for the power range are installed vertically at the four corners of the

. These assemblies are located equidistant from the reactor vessel along the length and, to imize neutron flux pattern distortions, within approximately one foot of the reactor vessel. Each er range detector provides two signals corresponding to the neutron flux in the upper and in the er sections of a core quadrant. The three ranges of detectors are used as inputs to monitor tron flux from a completely shutdown condition to 120 percent of full power, with the capability of rding overpower excursions up to 200 percent of full power.

output of the power range channels is used for:

Protecting the core against the consequences of rod ejection accidents Protecting the core against the consequences of adverse power distributions resulting from dropped rods The rod speed control function Alerting the operator to an excessive power imbalance between the quadrants intermediate range detectors also provide signals for the post-accident monitoring system.

ails of the neutron detectors and nuclear instrumentation design and the control and trip logic are n in Chapter 7. The limits on neutron flux operation and trip setpoints are given in the technical cifications.

6.4 Digital Metal Impact Monitoring System digital metal impact monitoring system is a nonsafety-related system that monitors the reactor lant system for metallic loose parts. It consists of several active instrumentation channels, each prising a piezoelectric accelerometer (sensor), signal conditioning, and diagnostic equipment.

digital impact monitoring system conforms with Regulatory Guide 1.133.

digital metal impact monitoring system is designed to detect a loose parts that weigh from 0.25 to ounds, and can also detect impact with a kinetic energy of 0.5 foot-pounds on the inside surface e reactor coolant system pressure boundary within three feet of a sensor.

digital impact monitoring system consists of several redundant instrumentation channels, each prised of a piezoelectric accelerometer (sensor), preamplifier, and signal conditioning equipment.

output signal from each accelerometer is amplified by the preamplifier and signal conditioning ipment before it is processed by a discriminator to eliminate noise and signals which are not cative of loose part impacts. The system starts up and operates automatically.

4.4-26 Revision 1

ormed using special test modules. These modules simulate impacts and test performance of the al processing equipment. Hardware integrity tests are also performed to verify equipment ration.

impact detect algorithm, used by the signal processing equipment, is designed to minimize the ber of false alarms. False impact detection, attributable to normal hydraulic, mechanical and trical noise, is minimized by a number of techniques including:

Utilizing a floating level within the impact detection algorithm. The floating level is based on signal levels not characteristic of an impact, and is generally a function of the background noise level.

Comparing the impact event with the times and type of normally occurring plant operation events received from plant control system such as a control rod stepping.

Comparing the number of events detected within a given time interval.

sensors of the impact monitoring system are fastened mechanically to the reactor coolant em at potential loose part collection regions including the upper and lower head region of the tor pressure vessel, and the reactor coolant inlet region of each steam generator.

equipment inside the containment is designed to remain functional through an earthquake of a nitude equal to 50 percent of the calculated safe shutdown earthquake and normal environments iation, vibration, temperature, humidity) anticipated during the operating lifetime. The instrument nnels associated with the sensors at each reactor coolant system location are physically arated from each other starting at the sensor locations to a point in the plant that is always essible for maintenance during full-power operation.

digital metal impact monitoring system is calibrated prior to plant startup. Capabilities exist for sequent periodic online channel checks and channel functional tests and for offline channel brations at refueling outages.

7 Combined License Information 7.1 Changes to the reference design of the fuel, burnable absorber rods, rod cluster control assemblies, or initial core design from that presented in the DCD are addressed in APP-GW-GLR-059 (Reference 87).

7.2 Following selection of the actual plant operating instrumentation and calculation of the instrumentation uncertainties of the operating plant parameters as discussed in Subsection 7.1.6, the design limit DNBR values will be calculated. The calculations will be completed using the RTDP with these instrumentation uncertainties and confirm that either the design limit DNBR values as described in this section remain valid or that the safety analysis minimum DNBR bounds the new design limit DNBR values plus DNBR penalties, such as rod bow penalty. This will be completed prior to fuel load.

8 References ANSI N18.2a-75, Nuclear Safety Criteria for the Design of Stationary Pressurized Water Reactor Plants.

4.4-27 Revision 1

Christensen, J. A., Allio, R. J., and Biancheria, A., Melting Point of Irradiated UO2, WCAP-6065, February 1965.

Davidson, S. L. and Kramer, W. R. (Ed.), Reference Core Report VANTAGE 5 Fuel Assembly, WCAP-10444-P-A (Proprietary) and WCAP-10445-NP-A (Non-Proprietary),

September 1985.

Tong, L. S., Boiling Crisis and Critical Heat Flux, AEC Critical Review Series, TID-25887, 1972.

Tong, L. S., Critical Heat Fluxes in Rod Bundles, Two Phase Flow and Heat Transfer in Rod Bundles, Annual Winter Meeting ASME, November 1968, p. 3146.

Letter from A. C. Thadani (NRC) to W. J. Johnson (Westinghouse), January 31, 1989,

Subject:

Acceptance for Referencing of Licensing Topical Report, WCAP-9226-P/

9227-NP, Reactor Core Response to Excessive Secondary Steam Releases.

Motley, F. E., Cadek, F. F., DNB Test Results for R-Grid Thimble Cold Wall Cells, WCAP-7695-L Addendum 1, October 1972.

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

Tong, L. S., Prediction of Departure from Nucleate Boiling for an Axially Nonuniform Heat Flux Distribution, Journal of Nuclear Energy 21, pp 241-248, 1967.

Not used.

Cadek, F. F., Motley, F. E., and Dominicis, D. P., Effect of Axial Spacing on Interchannel Thermal Mixing with the R Mixing Vane Grid, WCAP-7941-P-A (Proprietary) and WCAP-7959-A (Non-Proprietary), January 1975.

Rowe, D. S., and Angle, C. W., Crossflow Mixing Between Parallel Flow Channels During Boiling, Part II Measurements of Flow and Enthalpy in Two Parallel Channels, BNWL-371, Part 2, December 1967.

Rowe, D. S., and Angle, C. W., Crossflow Mixing Between Parallel Flow Channels During Boiling, Part III Effect of Spacers on Mixing Between Two Channels, BNWL-371, Part 3, January 1969.

Gonzalez-Santalo, J. M., and Griffith, P., Two-Phase Flow Mixing in Rod Bundle Subchannels, ASME Paper 72-WA/NE-19.

Motley, F. E., Wenzel, A. H., and Cadek, F. F., The Effect of 17 x 17 Fuel Assembly Geometry on Interchannel Thermal Mixing, WCAP-8298-P-A (Proprietary) and WCAP-8290A (Non-Proprietary), January 1975.

Hill, K. W., Motley, F. E., and Cadek, F. F., Effect of Local Heat Flux Spikes on DNB in Non Uniform Heated Rod Bundles, WCAP-8174 (Proprietary), August 1973, and WCAP-8202 (Non-Proprietary), August 1973.

Staff approval is required prior to implementing a change in this information.

4.4-28 Revision 1

Skaritka, J., Ed, Fuel Rod Bow Evaluation, WCAP-8691, Revision 1 (Proprietary) and WCAP-8692, Revision 1 (Non-Proprietary), July 1979.

Partial Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1, Letter from E. P. Rahe, Jr. (Westinghouse) to J. R. Miller (NRC),

NS-EPR-2515, October 9, 1981; Remaining Response to Request Number 1 for Additional Information on WCAP-8691, Revision 1, Letter from E. P. Rahe, Jr.

(Westinghouse) to R. J. Miller (NRC), NS-EPR-2572, March 16, 1982.

Letter from C. Berlinger (NRC) to E. P. Rahe, Jr. (Westinghouse),

Subject:

Request for Reduction in Fuel Assembly Burnup Limit for Calculations of Maximum Rod Bow Penalty, June 18, 1986.

Hochreiter, L. E., Applications of the THINC-IV Program to PWR Design, WCAP-8054-P-A (Proprietary), February 1989 and WCAP-8195 (Non-Proprietary),

October 1973.

Hochreiter, L. E., Chelemer, H., and Chu, P. T., THINC-IV, An Improved Program for Thermal-Hydraulic Analysis of Rod Bundle Cores, WCAP-7956-P-A, February 1989.

Dittus, F. W., and Boelter, L. M. K., Heat Transfer in Automobile Radiators of the Tubular Type, California University Publication in Engineering 2, No. 13, 443461, 1930.

Weisman, J., Heat Transfer to Water Flowing Parallel to Tube Bundles, Nuclear Science Engineering 6, pp 78-79, 1959.

Thom, J. R. S., et al., Boiling in Subcooled Water During Flowup Heated Tubes or Annuli, Proceedings of the Institution of Mechanical Engineers 180, Part C, pp 226-246, 1955-1966.

Not used.

Not used.

Not used.

Kjaerheim, G., and Rolstad, E., In-Pile Determination of UO2, Thermal Conductivity, Density Effects, and Gap Conductance, HPR-80, December 1967.

Kjaerheim, G., In-Pile Measurements of Center Fuel Temperatures and Thermal Conductivity Determination of Oxide Fuels, Paper IFA-175 Presented at the European Atomic Energy Society Symposium on Performance Experience of Water-Cooled Power Reactor Fuel, Stockholm, Sweden, October 1969.

Cohen, I., Lustman, B., and Eichenberg, D., Measurement of the Thermal Conductivity of Metal-Glad Uranium Oxide Rods During Irradiation, WAPD-228, 1960.

Clough, D. J., and Sayers, J. B., The Measurement of the Thermal Conductivity of UO2, under Irradiation in the Temperature Range 150 to 1600°C, AERE-4690, UKAEA Research Group, Harwell, December 1964.

4.4-29 Revision 1

Devold, I., A Study of the Temperature Distribution in UO2, Reactor Fuel Elements, AE-318, Aktiebolaget Atomenergi, Stockholm, Sweden, 1968.

Balfour, M. G., Christensen, J. A., and Ferrari, H. M., In-Pile Measurement of UO2 Thermal Conductivity, WCAP-2923, 1966.

Howard, V. C., and Gulvin, T. G., Thermal Conductivity Determinations on Uranium Dioxide by a Radial Flow Method, UKAEA IG-Report 51, November 1960.

Lucks, C. F., and Deem, H. W., Thermal Conductivity and Electrical Conductivity of UO2, in Progress Reports Relating to Civilian Applications, BMI-1448 (Revised) for June 1960, BMI-1489 (Revised) for December 1960, and BMI-1518 (Revised) for May 1961.

Daniel, J. L., Matolich, J. Jr., and Deem, H. W., Thermal Conductivity of UO2, HW-69945, September 1962.

Feith, A. D., Thermal Conductivity of UO2 by a Radial Heat Flow Method, TID-21668, 1962.

Vogt, J., Grandel, L., and Runfors, U., Determination of the Thermal Conductivity of Unirradiated Uranium Dioxide, AB Atomenergi Report RMB-527, 1964, Quoted by IAEA Technical Report Series No. 59, Thermal Conductivity of Uranium Dioxide.

Nishijima, T., Kawada, T., and Ishihata, A., Thermal Conductivity of Sintered UO2 and 4Al2O3 at High Temperatures, Journal of the American Ceramic Society 48, pp 31-44, 1965.

Ainscough, J. B., and Wheeler, M. J., Thermal Diffusivity and Thermal Conductivity of Sintered Uranium Dioxide, Proceedings of the Seventh Conference of Thermal Conductivity, National Bureau of Standards, Washington, p 467, 1968.

Godfrey, T. G., et al., Thermal Conductivity of Uranium Dioxide and Armco Iron by an Improved Radial Heat Flow Technique, ORNL-3556, June 1964.

Stora, J. P., et al., Thermal Conductivity of Sintered Uranium Oxide Under In-Pile Conditions, EURAEC-1095, August 1964.

Bush, A. J., Apparatus for Measuring Thermal Conductivity to 2500°C, Reporting 64-1P6-401-43 (Proprietary), Westinghouse Research Laboratories, February 1965.

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.

Kruger, O. L., Heat Transfer Properties of Uranium and Plutonium Dioxide, Paper 11-N-68F, presented at the Fall Meeting of Nuclear Division of the American Ceramic Society, Pittsburgh, September 1968.

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.

4.4-30 Revision 1

Duncan, R. N., Rabbit Capsule Irradiation of UO2, CVTR Project, CVNA Project, CVNA-142, June 1962.

Nelson, R. G., et al., Fission Gas Release from UO2 Fuel Rods with Gross Central Melting, GEAP-4572, July 1964.

Gyllander, J. A., In-Pile Determination of the Thermal Conductivity of UO2 in the Range 500 to 2500°C, AE-411, January 1971.

Lyons, M. F., et al., UO2 Powder and Pellet Thermal Conductivity During Irradiation, GEAP-5100-1, March 1966.

Coplin, D. H., et al., The Thermal Conductivity of UO2 by Direct In-Reactor Measurements, GEAP-5100-6, March 1968.

Bain, A. S., The Heat Rating Required to Produce Center Melting in Various UO2 Fuels, ASTM Special Technical Publication No. 306, Philadelphia, pp 30-46, 1962.

Stora, J. P., In-Reactor Measurements of the Integrated Thermal Conductivity of UO2 -

Effect of Porosity, Transactions of the American Nuclear Society 13, pp 137-138, 1970.

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, 1966.

Poncelet, C. G., Burnup Physics of Heterogeneous Reactor Lattices, WCAP-6069, June 1965.

Nodvick, R. J., Saxton Core II Fuel Performance Evaluation, WCAP-3385-56, Part II, Evaluation of Mass Spectrometric and Radiochemical Analyses of Irradiated Saxton Plutonium Fuel, July 1970.

Not used.

Not used.

Not used.

Idel'chik, I. E., Handbook of Hydraulic Resistance, 2nd Edition, Hemisphere Publishing Corp., 1986.

McFarlane, A. F., Power Peaking Factors, WCAP-7912-P-A (Proprietary) and WCAP-7912-A (Non-Proprietary), January 1975.

Not used.

Not used.

Boure, J. A., Bergles, A. E., and Tong, L. S., Review of Two-Phase Flow Instability, Nuclear Engineering Design 25, pp 165-192, 1973.

4.4-31 Revision 1

Saha, P., Ishii, M., and Zuber, N., An Experimental Investigation of the Thermally Induced Flow Oscillations in Two-Phase Systems, Journal of Heat Transfer, pp 616-622, November 1976.

Virgil C. Summer Nuclear Station FSAR, Chapter 4, South Carolina Electric & Gas Company, Docket No. 50-395.

Byron/Braidwood Stations FSAR, Chapter 4, Commonwealth Edison Company, Docket No. 50-456.

South Texas Project Electric Generating Station FSAR, Chapter 4, Houston Lighting and Power Company, Docket No. 50-498.

Kakac, S., et al., Sustained and Transient Boiling Flow Instabilities in a Cross-Connected Four-Parallel-Channel Upflow System, Proceedings of Fifth International Heat Transfer Conference, Tokyo, September 1974.

Kao, H. S., Morgan, T. D., and Parker, W. B., Prediction of Flow Oscillation in Reactor Core Channel, Transactions of the American Nuclear Society 16, pp 212-213, 1973.

Ohtsubo, A., and Uruwashi, S., Stagnant Fluid Due to Local Flow Blockage, Journal of Nuclear Science Technology, No. 7, pp 433-434, 1972.

Basmer, P., Kirsh, D., and Schultheiss, G. F., Investigation of the Flow Pattern in the Recirculation Zone Downstream of Local Coolant Blockages in Pin Bundles, Atomwirtschaft 17, No. 8, pp 416-417, 1972 (in German).

Burke, T. M., Meyer, G. E., and Shefcheck, J., Analysis of Data from the Zion (Unit 1)

THINC Verification Test, WCAP-8453-A, May 1976.

Not used.

Not used.

Davidson, S. L., and Ryan, T. L., VANTAGE+ Fuel Assembly Reference Core Report, WCAP-12610-P-A (Proprietary) and WCAP-14342-A (Non-Proprietary), April 1995.

Smith, L. D., et al., Modified WRB-2 Correlation, WRB-2M, for Predicting Critical Heat Flux in 17x17 Rod Bundles with Modified LPD Mixing Vane Grids, WCAP-15025-P-A (Proprietary) and WCAP-15026-NP (Non-Proprietary), April 1999.

. Letter from D. S. Collins (USNRC) to J. A. Gresham (Westinghouse), Modified WRB-2 Correlation WRB-2M for Predicting Critical Heat Flux in 17x17 Rod Bundles with Modified LPD Mixing Vane Grids, February 3, 2006.

Sung, Y. X., et al., VIPRE-01 Modeling and Qualification for Pressurized Water Reactor Non-LOCA Thermal-Hydraulic Safety Analysis, WCAP-14565-P-A and WCAP-15306-NP-A, October 1999.

4.4-32 Revision 1

Slagle, W. H. (ed.) et al., Westinghouse Improved Performance Analysis and Design Model (PAD 4.0), WCAP-15063-P-A, Revision 1 (Proprietary) and WCAP-15064-NP-A, Revision 1 (Non-Proprietary), July 2000.

Kitchen, T. J., Generic Safety Evaluation for 17x17 Standard Robust Fuel Assembly (17x17 STD RFA), SECL-98-056, Revision 0, September 30, 1998.

APP-GW-GLR-059/WCAP-16652-NP, AP1000 Core & Fuel Design Technical Report, Revision 0.

Letter, Peralta, J. D. (NRC) to Maurer, B. F. (Westinghouse), Approval for Increase in Licensing Burnup Limit to 62,000 MWD/MTU (TAC No. MD1486), May 25, 2006.

4.4-33 Revision 1

(AP1000, AP600 and a Typical Westinghouse XL Plant)

Typical Design Parameters AP1000(a) AP600 XL Plant ctor core heat output (MWt) 3400 1933 3800 ctor core heat output (106 BTU/hr) 11601 6596 12,969 t generated in fuel (%) 97.4 97.4 97.4 tem pressure, nominal (psia) 2250 2250 2250 tem pressure, minimal (psia) 2190 2200 2204 imum DNBR at nominal conditions ypical flow channel 2.80 3.48 2.20 himble (cold wall) flow channel) 2.74 3.33 2.12 imum DNBR for design transients ypical flow channel >1.25b >1.22b >1.23 >1.26 himble (cold wall)flow channel >1.25b >1.21b >1.22 >1.24 B correlation(c) WRB-2M WRB-2 WRB-1 lant conditions(d) essel minimum measured flow rate (MMF) 106 lbm/hr 115.55 74.4 148.9 pm 301,670 193,200 403,000 essel thermal design flow rate (TDF) 106 lbm/hr 113.5 72.9 145.0 pm 296,000 189,600 392,000 ffective flow rate for heat transfer(e) 106 lbm/hr 106.8 66.3 132.7 pm 278,500 172,500 358,700 ffective flow area for heat transfer (ft2) 41.8 38.5 51.1 Average velocity along fuel rods (ft/s)(e) 15.8 10.6 16.6 Average mass velocity, 106 lbm/hr-ft2(e) 2.55 1.72 2.60 lant temperature(d)(e) ominal inlet (°F) 535.0 532.8 561.2 verage rise in vessel (°F) 77.2 69.6 63.6 verage rise in core (°F) 81.4 75.8 68.7 verage in core (°F) 578.1 572.6 597.8 verage in vessel (°F) 573.6 567.6 593.0 4.4-34 Revision 1

Typical Design Parameters AP1000(a) AP600 XL Plant t transfer ctive heat transfer surface area (ft2)(f) 56,700 44,884 69,700 verage heat flux (BTU/hr-ft2) 199,300 143,000 181,200 Maximum heat flux for normal operation (BTU/hr-ft2)(g) 518,200 372,226 498,200 verage linear power (kW/ft)(f)(m) 5.72 4.11 5.20 eak linear power for normal operation (kW/ft)(g,h) 14.9 10.7 14.0 eak linear power resulting from overpower transients/operator errors, assuming a maximum overpower of 118% (kW/ft)(h) 22.45 22.5 22.45 eak Linear power for prevention of center-line melt (kW/ft)(i) 22.5 22.5 22.45 ower density (kW/liter of core)(j) 109.7 78.82 98.8 pecific power (kW/kg uranium)(j) 40.2 28.89 36.6 l central temperature eak at peak linear power for prevention of 4700 4,700 4700 centerline melt (°F) ssure drop(k) cross core (psi) 39.9 +/- 4.0(l) 17.5 +/- 1.7 38.8 +/- 3.9 cross vessel, including nozzle (psi) 62.3 +/- 6.2(l) 45.3 +/- 4.5 59.7 +/- 6.0 s:

Robust Fuel Assembly.

1.25 applies to Core and Axial Offset limits; 1.22 and 1.21 apply to all other RTDP transients.

WRB-2M is used for AP1000. WRB-2 or W-3 is used for AP1000 where WRB-2M is not applicable. See Subsection 4.4.2.2.1 for use of W-3, WRB-2 and WRB-2M correlations.

Based on vessel average temperature equal to 573.6°F. Flow rates and temperatures based on 10 percent steam generator tube plugging.

Based on thermal design flow and 5.9 percent bypass flow.

Based on densified active fuel length. The value for AP1000 is rounded to 5.72 kW/ft.

Based on 2.60 FQ peaking factor.

See Subsection 4.3.2.2.6.

See Subsection 4.4.2.11.6.

Based on cold dimensions and 95.5 percent of theoretical density fuel for AP1000; 95 percent for others.

These are typical values based on best-estimate reactor flow rate as discussed in Section 5.1.

Inlet temperature = 536.8°F.

The value for AP1000 is rounded to 5.72 kW/ft.

4.4-35 Revision 1

With Design Hot Channel Factors (Based On VIPRE-01)

Average Maximum re, % 0.0 -

t Subchannel, % 0.1 0.9 4.4-36 Revision 1

Figure 4.4-1 Thermal Diffusion Coefficient (TDC) as a Function of Reynolds Number 4.4-37 Revision 1

Figure 4.4-2 Thermal Conductivity of Uranium Dioxide (Data Corrected to 95% Theoretical Density) 4.4-38 Revision 1

1.1 Materials Specifications parts of the control rod drive mechanisms and control rod drive line exposed to reactor coolant made of metals that resist the corrosive action of the coolant. Three types of metals are used usively: stainless steels, nickel-chromium-iron alloys, and, to a limited extent, cobalt-based ys. These materials have provided many years of successful operation in similar control rod drive hanisms. In the case of stainless steels, only austenitic and martensitic stainless steels are used.

ere low or zero cobalt alloys are substituted for cobalt-based alloy pins, bars, or hard facing, the stitute material is qualified by evaluation or test.

ssure-containing materials comply with the ASME Code,Section III. The material specifications portions of the control rod drive mechanism that are reactor coolant pressure boundary are uded in Table 5.2-1. These parts are fabricated from austenitic (Type 316, 316L, 316LN and e 304, 304L, 304LN) stainless steel. Nickel-chromium-iron alloy (Alloy 690) is used for the reactor sel head penetration. For pressure boundary parts, austenitic stainless steels are not used in the t-treated conditions which can cause susceptibility to stress-corrosion cracking or accelerated osion in pressurized water reactor coolant chemistry and temperature environments. Pressure ndary parts and components made of stainless steel do not have specified minimum yield ngth greater than 90,000 psi.

material selection is based in part on the duty cycle specified for the control rod drive hanisms and control rods. The materials are specified so that the components do not suffer erse effects, such as excessive wear or galling, as a result of a minimum 300 trips from full power 60 coupling and decoupling cycles of the drive rod coupling assembly. The material for the trol rod drive mechanisms and the control rod assemblies are selected for acceptable ormance. That is, the design goal is to achieve a service life of 9 x 106 full-step cycles. Inspection hanges in operation indicate the need for replacement or refurbishment. The worst case result of etected wear of a control rod drive mechanism or drive rod is a rod assembly drop or a failure to p an assembly during a trip. Both events are accounted for in safety analyses. The pressure ndary components are not subject to significant wear due to stepping cycles.

rnal latch assembly parts are fabricated of heat-treated martensitic and austenitic stainless steel.

t treatment is such that stress-corrosion cracking is not initiated. Components and parts made of nless steel do not have specified minimum yield strength greater than 90,000 psi. Magnetic pole es are immersed in the reactor coolant and are fabricated from Type 410 stainless steel.

magnetic parts, except shims, pins, and springs, are fabricated from Type 304 stainless steel. A alt alloy or qualified substitute is used to fabricate the latch, link, and link pins. Springs and shims made from nickel-chromium-iron alloy (Alloy X-750 and Alloy 625). Lock screws are fabricated of e 316 stainless steel. Latch arm tips fabricated of stainless steel may be surfaced with a suitable d facing material to provide improved resistance to wear. Hard chrome plate is used selectively for ring and wear surfaces.

drive rod assembly is also immersed in the reactor coolant and uses a Type 410 stainless steel e rod. The drive rod coupling is machined from Type 403 or 410 stainless steel. The protective ve and disconnect button are also Type 410 stainless steel. The remaining parts are Type 304 or e 304L stainless steel with the exception of the springs, button retainer, and locking button, which fabricated of nickel-chromium-iron alloy.

absorber rodlets in the rod control cluster assemblies and the gray rod cluster assemblies are ed stainless steel tubes (cladding) containing absorber material. The other rodlets in the gray rod 4.5-1 Revision 1

tron flux in the control rod materials are addressed in Section 4.2. The outside surface of the orber and other rodlets is chromium plated to enhance resistance to wear due to the stepping ion and vibration of the rods. The rods included in one rod control cluster assembly or gray rod ter assembly are attached at the top to a common hub which connects with the drive rod of the trol rod drive mechanism. The hub is fabricated of type 316 stainless steel.

coil housing is exposed to containment atmosphere and requires a magnetic material. Low on cast steel and ductile iron are qualified by tests or other evaluations for this application. The hed housings are electroless nickel plated to provide resistance against general corrosion.

s are wound on composite bobbins, with double glass-insulated copper wire. Coils are vacuum regnated with silicone varnish. A wrapping of mica sheet is secured to the coil outside diameter.

result is a well-insulated coil capable of sustained operation at 392°F (200°C).

1.2 Fabrication and Processing of Austenitic Stainless Steel Components discussions provided in Subsection 5.2.3.4 concerning the processes, inspections, and tests on tenitic stainless steel components to prevent increased susceptibility to intergranular corrosion sed by sensitization are applicable to the austenitic stainless steel pressure-housing components e control rod drive mechanism. The discussions provided in Subsection 5.2.3.4, concerning the trol of welding of austenitic stainless steels especially control of delta ferrite are also applicable.

section 5.2.3.4 discusses the compliance with the guidelines of Regulatory Guides 1.31, 1.34, 1.44. The welded control rod drive mechanism austenitic stainless steels that come into contact the primary reactor coolant meet the guidance of Regulatory Guide 1.44.

1.3 Other Materials the cobalt alloy used to fabricate the latch, link, and link pins in latch assemblies, stress-corrosion king has not been observed in this application. Where hardfacing material is used in the latch embly, a cobalt base alloy equivalent to Stellite-6 or qualified low or zero cobalt substitute is used.

or zero cobalt alloys used for hardfacing or other applications where cobalt alloys have been iously used are qualified using wear and corrosion tests. The corrosion tests qualify the corrosion stance of the alloy in reactor coolant. Low cobalt or cobalt free wear resistant alloys considered his application include those developed and qualified in industry programs.

springs in the control rod drive mechanism are made from nickel-chromium-iron alloy oy 750), ordered to Aerospace Material Specification (AMS) 5698 or AMS 5699 with additional rictions on prohibited materials. Operating experience has shown that springs made of this erial are not subject to stress-corrosion cracking in pressurized water reactor primary water ironments. Alloy 750 is not used for bolting applications in the control rod drive mechanisms.

1.4 Contamination Protection and Cleaning of Austenitic Stainless Steel control rod drive mechanisms are cleaned prior to delivery in accordance with the guidance ided in NQA-1 (see Chapter 17). Process specifications in packaging and shipment are ussed in Subsection 5.2.3. Westinghouse personnel conduct surveillance of these operations to fy that manufacturers and installers adhere to appropriate requirements as described in section 5.2.3.

4.5-2 Revision 1

2 Reactor Internal and Core Support Materials 2.1 Materials Specifications major core support material for the reactor internals is SA-182, SA-336, SA-376, SA-479, or 240 Types 304, 304L, 304LN, or 304H stainless steels. Fabricators performing welding of any of e materials are required to qualify the welding procedures for maximum carbon content and heat t for each welding process in accordance with Regulatory Guide 1.44. For threaded structural eners the material used is strain hardened Type 316 stainless steel and for the clevis insert-to-sel bolts either UNS N07718 or N07750. Remaining internals parts not fabricated from Types 304, L, 304LN, or 304H stainless steels typically include wear surfaces such as hardfacing on the al keys, clevis inserts, alignment pins (Stellite' 6 or 156 or low cobalt hardfaces); dowel pins e 316); hold down spring (Type 403 stainless steel (modified)); clevis inserts (UNS N06690); and diation specimen springs (UNS N07750).Instrument guide assembly materials that are not es 304, 304L, 304LN, or 304H stainless steel are the guide bushings and guide stud tip S S21800) and the instrument guide tube spring (UNS N07718). Core support structure and aded structural fastener materials are specified in the ASME Code,Section III, Appendix I as plemented by Code Cases N-60 and N-4. The qualification of cobalt free wear resistant alloys for in reactor coolant is addressed in Subsection 4.5.1.3.

use of cast austenitic stainless steel (CASS) is minimized in the AP1000 reactor internals. If d, CASS will be limited in carbon (low carbon grade: L grade) and ferrite contents and will be luated in terms of thermal aging effects.

estimated peak neutron fluence for the AP1000 reactor internals has been considered in the ign. Susceptibility to irradiation-assisted stress corrosion cracking or void swelling in reactor rnals identified in the current pressurized water reactor fleet are being addressed in reactor rnals material reliability programs. The selection of materials for the AP1000 reactor internals siders information developed by these programs. Ni-Cr-Fe Alloy 600 is not used in the AP1000 tor internals.

2.2 Controls on Welding discussions provided in Subsection 5.2.3.4 are applicable to the welding of reactor internals and support components.

2.3 Nondestructive Examination of Tubular Products and Fittings nondestructive examination of wrought seamless tubular products and fittings is in accordance ASME Code,Section III, Article NG-2500. The acceptance standards are in accordance with the uirements of ASME Code,Section III, Article NG-5300.

2.4 Fabrication and Processing of Austenitic Stainless Steel Components discussions provided in Subsection 5.2.3.4 and Section 1.9 describes the conformance of tor internals and core support structures with Regulatory Guides 1.31 and 1.44.

discussion provided in Section 1.9 describes the conformance of reactor internals with ulatory Guides 1.34 and 1.71.

4.5-3 Revision 1

core support structures describe the conformance of the process specifications with Regulatory de 1.37. The process specifications follow the guidance of NQA-1 (Reference 1).

3 Combined License Information section contained no requirement for additional information.

4.5-4 Revision 1

control rod drive mechanism (CRDM) and operation of the control rod drive system are cribed in Subsection 3.9.4. Figure 3.9-4 provides the details of the control rod drive mechanisms.

re 4.2-8 provides the configuration of the driveline, including the control rod drive mechanism. No raulic system is associated with the functioning of the control rod drive system. The rumentation and controls for the reactor trip system are described in Section 7.2. The reactor trol system is described in Section 7.7.

control rod drive mechanisms are contained within an integrated head package located on top of reactor vessel head as described in Subsection 3.9.7. This assembly provides the support uired for seismic restraint in conjunction with the attachment of the control rod drive mechanisms e reactor vessel head. An outer shroud and the seismic restraint structure isolate the control rod e mechanisms from the effects of ruptures of high-energy lines outside the shroud, and from siles. The shroud also is used to direct air from the cooling fans past the control rod drive hanisms. The cooling system maintains the temperatures of the coils in the control rod drive hanisms below the design operating temperature. The integrated head package provides the per support and required separation for electrical lines providing power to the control rod drive hanisms and signals from the rod position sensors.

line for the reactor head vent system is located among the control rod drive mechanisms and is ported by the integrated head package. This line is pressurized to reactor coolant system sure and considered to be a high-energy line. This line is constructed to the appropriate uirements of the ASME Code. Figure 3.9-7 shows elements of the integrated head package ounding the control rod drive mechanisms.

2 Evaluations of the Control Rod Drive System control systems of the type used in the AP1000 have been analyzed in detailed reliability ies. These studies include fault tree analysis and failure mode and effects analyses. These ies, and the analyses presented in Chapter 15, demonstrate that the control rod drive system orms its intended safety-related function - a reactor trip. The control rod drive system puts the tor in a subcritical condition when a safety-related system setting is reached with an assumed ible failure of a single active component.

essential elements of the control rod drive system (those required to provide reactor trip) are ated from nonessential portions of the rod control system by the reactor trip switchgear, as cribed in Section 7.2. The essential portion of the control rod drive system is shielded from the ct effects of postulated moderate- and high-energy line breaks by the integrated head package.

dynamic effects of pipe ruptures do not have to be considered for those pipes that satisfy the uirements for mechanistic pipe break, as outlined in Subsection 3.6.3.

reactor vessel head vent lines and instrumentation conduits are one inch nominal diameter or ller. Breaks in lines of this size do not have to be postulated for dynamic effects, pressurization, spray wetting. The pressure boundary housing of the control rod drive mechanisms is structed to the requirements of the ASME Code and a break in this pressure boundary is not ible.

only instrumentation required of the control rod drive mechanism and supporting systems to rate safely is the rod position indicator. A break in the cables connected to the rod position cators would neither preclude a reactor trip, nor would it result in an unplanned withdrawal of a assembly. A break in the power cable to the control rod drive mechanism coils results in a drop of 4.6-1 Revision 1

afety valves. Overheating of the control rod drive mechanism coils due to a failure of the cooling em would in the worst case result in a drop of one or more rod assemblies. The reactor and tor protection system is designed to accommodate and protect against rod drop events.

itional information is provided in Subsection 3.9.1, and Sections 7.2, and 15.4.

3 Testing and Verification of the Control Rod Drive System control rod drive system is extensively tested prior to its operation. These tests may be divided into five categories:

Prototype tests of components Prototype control rod drive system tests Production tests of components following manufacture and prior to installation Onsite pre-operational and initial startup tests Periodic in-service tests se tests, which are described in Subsection 3.9.4.4 and Sections 4.2 and 14.2, are conducted to fy the operability of the control rod drive system when called upon to function.

4 Information for Combined Performance of Reactivity Systems ndicated in Chapter 15, there are only three postulated events that assume credit for reactivity trol systems, other than a reactor trip to render the plant subcritical. These events are the steam-break, feedwater line break, and small break loss of coolant accident. The reactivity control ems in these accidents are the reactor trip system and the passive core cooling system (PXS).

itional information on the control rod drive system is presented in Subsection 3.9.4. The passive cooling system is discussed further in Section 6.3.

credit is taken for the boration capabilities of the chemical and volume control system (CVS) as a em in the analysis of transients presented in Chapter 15. Information on the capabilities of the mical and volume control system is provided in Subsection 9.3.6. The adverse boron dilution sibilities due to the operation of the chemical and volume control system are investigated in section 15.4.6. Prior proper operation of the chemical and volume control system has been umed as an initial condition to evaluate transients. Appropriate technical specifications promote correct operation or remedial action.

AP1000 instrumentation and control system includes a diverse actuation system (DAS). This em provides for automatic control rod insertion, turbine trip, passive residual heat removal heat hanger start, core makeup tank start, isolation of critical containment penetrations, and start of the sive containment cooling system as appropriate upon conditions indicative of an anticipated sient without scram or other failure of the plant control and reactor protection system. This em is diverse and independent from the reactor trip system from the sensor through actuation ices.

ddition to the above, the AP1000 plant systems provide for operator response to an anticipated sient without scram (ATWS) event that includes core reactivity control followed by core decay t removal. Core reactivity control is provided by a manual trip of the control rods, insertion of the 4.6-2 Revision 1

5 Evaluation of Combined Performance evaluations of the steam-line break, the feedwater line break, and the small break loss of coolant dent, which presume the combined actuation of the reactor trip system and the control rod drive em and the passive safety injection, are presented in Subsections 15.1.5 and 15.2.8 and tion 15.6. Reactor trip signals and signals to actuate passive safety features for these events are erated from functionally diverse sensors. These signals actuate diverse means of reactivity trol; that is, control rod insertion and injection of soluble neutron absorber.

-diverse but redundant types of equipment are used only in the processing of the incoming sor signals into appropriate logic which initiates the protective action. This equipment is described ections 7.2 and 7.3. In particular, protection from equipment failures is provided by redundant ipment and periodic testing. Effects of failures of this equipment have been extensively stigated. Reliability studies, including failure mode and effects analysis for this type of equipment fy that a single failure does not have an adverse effect upon the engineered safety features ation system. Adequacy of the passive core cooling system performance under faulted ditions is verified in Section 6.3.

ddition to the automatic actuations provided for by the diverse actuation system, that system also ides for manual actuation of the reactor trip.

probability of a common mode failure impairing the ability of the reactor trip system to perform its ty-related function is extremely low. However, analyses are performed to demonstrate pliance with the requirements of 10 CFR 50.62. These analyses demonstrate that safety criteria ld not be exceeded even if the control rod drive system were rendered incapable of functioning ng anticipated transients for which its function would normally be expected. The evaluation onstrates that borated water from the core makeup tank shuts down the reactor with no rods uired, and the passive residual heat removal system provides sufficient core heat removal.

6 Combined License Information section contained no requirement for additional information.

4.6-3 Revision 1