ML20098F120

Reload Design Methodology

ML20098F120
Person / Time
Site:	Oconee
Issue date:	04/30/1984
From:	DUKE POWER CO.
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ML16152A361	List: ML20098F120
References
NFS-1001A, NUDOCS 8410020354
Download: ML20098F120 (558)

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4 4 DUKE POWER COMPANY OCONEE NUCLEAR STATION I

l RELOAD DESIGN METHODOLOGY t NFS-1001 A

APRIL 1984 I

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DUKE POWER COMPANY OCONEE NUCLEAR STATION RELOAD DESIGN METHODOLOGY 4

f Technical Report NFS-1001 l-i April 23, 1979 i

i '- Revision 4 June 1981 l

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• i & Sdst Mr. William O. Parker, Jr.

l Vice President - Steam Production Duke Power Company P. O. Box 33189 422 South Church Street Charlotte, North Carolina 28242

Dear Mr. Parker:

The staff has completed the review of Technical Report NFS-1001, "0conee Nuclear Station Reload Design Methodology" which was submitted by letter dated April 23, 1979 and revised by letters dated May 20,1980, January 28, April 22 and June 16, 1981. The results of our review are contained in the enclosed Safety Evaluation.

We have found the revised report to be an acceptable method of performing reload design calculations for future Oconee Nuclear Station, Units 1, 2 and 3 reloads.

If you have any questions on this subject, please contact me.

Sincerely,

b. h%M Philip C. Wagner, Projeck Manager Operating Reactors Branch #4 Division of Licensing

Enclosure:

Safety Evaluation cc w/ enclosure:

See next page

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l ABSTRACT This Technical Report describes Duke Power Company's Reload Design Methodology for the Oconee Nuclear Station. Included in this report are descriptions of Fuel Design, Fuel Cycle Design, Fuel Mechanical Performance, Maneuvering Analysis, Thermal Hydraulic Design, Technical Specifications Review and Development, t Accident Analysis Review, and the Development of Core Physics Parameters.

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l TABLE OF CONTENTS Page

1. Introduction 1-1

2. Fuel Design 2.1 Fuel Pellet 2-1 2.2 Fuel Rod 2-1 2.3 Fuel Assembly Design 2-2 2.4 Core Component Data 2-3

3. Fuel Cycle Design 3.1 Preliminary Fuel Cycle Design 3-1 3.1.1 Overview of Nuclear Calculational System 3-1 3.1.2 Calculations and Results of PFCD 3-3 3.2 Final Fuel Cycle Design 3-3 3.2.1 Fuel Shuffle Optimization and Cycle Depletion 3-4 3.2.2 Rod Worth Calculations 3-4 3.2.3 Pcuer Distribution Calculations 3-8 3.2.4 Fuel Burnup Calculations 3-8 3.2.5 Reactivity Coefficients and Deficits 3-9 3.2.6 Boron Related Parameters 3-13 3.2.7 Xenon Worth 3-13 3.2.8 Kinetics Parameters 3-13

4. Fuel Mechanical Performance 4.1 Introduction 4-1 l

4.2 Cladding Collapse 4-2 4.3 Cladding Strain Analysis 4-3 l 4.4 Cladding Stress Analysis 4-5 2 l 4.5 Fuel Pin Pressure Analysis 4-7 4.6 Linear Heat Rate Capability 4 l l

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5. Maneuvering Analysis 5.1 Fuel Cycle Depletion 5-1 5.2 Integral Rod Worth 5-2 5.3 Power Maneuver 5-2 1 5.4 Control Rod Scans Off the Power Maneuver 5-3 5.5 Control Rod Scans Off Fuel Cycle Depletion 5-3

6. Thermal Hydraulic Design 6.1 Introduction 6-1 6.2 Thermal-Hydraulic Design Criteria 6-1 6.3 Analysis Methodology 6-2 6.4 Core Inlet conditions 6-3 6.5 Reference Design DNBR Analysis 6-7 6.6- Equivalent Two Channel Model 6-7 1 6.7 Determination of Pressure-Temperature Core Protection Safety Limits 6-7 6.8 Determination of Power Distribution Limits 6-8 6.9 Transient Analysis - Determination of the Flux-Flow Ratio 6-10 6.10 Application of the Rod Bow Penalty 6-10

7. Technical Specifications 1 7.1 Technical Specifications Review 7-1 7.2 Development of Core Safety Limits 7-1 7.3 Development of Limiting Safety System Settings 7-8 7.4 Development of Limiting Conditions for Operation 7-10 l 8. Accident Analysis Review 8.1 Introduction 8-1 8.2 Overview of Accident Analysis Review 8-2 1 8.3 Discussion of Individual Accidents 8-3

9. Development of Core Physics Parameters 7.1 Startup Test Predictions 9-1 9.1.1 Critical Boron Concentration and Boron Worths 9-1 9.1.2 Xenon Worths 9-2 9.1.3 Rod Worths 9-3 9.1.4 Reactivity Coefficients 9-4 9.1.5 Power Distributions 9-5 9.1.6 Kinetics Parameters 9-5 iv i

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, 9.2 Core Physics Report 9-6

10. References 10-1 Appendix A - Code Summary A-1 Amendments - NRC Questions and Responses Amend 1-1 Supplement 1 Physics Test Comparisons S1-1 Supplement 2 Nuclear Reliability Factors for EPRI-NODE-P S2-1 i ,.

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• LIST OF TABLES Page 2-1 'Oconee System and Component Data 2-4 3 3-1" Shutdown Margin Calculation 3-15 3-2 Ejected Rod Worths 3-16 3-3 Radial Pin Peak 3-17 3-4 Boron-Parameters 3-Ic 4-1 , Fuel Mechanical Performance Assessment Criteria 4-9 2 4-2 Axici Flux Shapes Used for Thermal Analyses 4-10 2 7-1. Reactor Protection System Trip Functions 7-16 8-l' Accident Analysis Review Key Safety Parameter Checklist 8-18 l 9-1 Critical Boron Concentracion (PPM) 9-7 9-2 Boron Worth (PPMB/%Ap) 9-8 3 Radial and Total Peaking Power Maps. 9-9 9-4 Core Physics Data 9-10 i

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LIST OF FIGURES Page

'l-1 Elements of Reload Design 1-1 3-1 Nuclear Flow Chart for EPRI-ARMP 3-19 4-1 Pin Power Versus Burnup Envelope For Thermal Analysis Assessments 4-11 4-2 Radial Assembly Power Versus Burnup For Creep Collapse Analysis Assessments 4-12 4-3 Thermal Analysis Flow Diagram 4-13 2 4-4' Hechanical. Analysis Flow Diagram 4-14 4-5 Fuel Pin Pressure Versus Burnup 4-15 4 Fuel Linear Heat Rate to Melt Versus Burnup 4-16 6-1 Thermal Hydraulic Analysis Methodology 6-12 6-2 Steady State Pressure-Temperature Core Protective Safety Limits 6-13 6-3 Generic DNBR Curves 6 '7-1 Core Safety Pressure Temperature Limits 7-17 7-2 Margin to Center Fuel Melt LHR Versus Core Offset 7-18 7-3 Core Safety Power-Power Imbalance Limits 7-19 7-4 Determination of RPS P-T Trip Setpoints 7-20 7-5 Determination of RPS Power-Flow-Imbalance Trip Setpoints 7-21 7-6 Operating Limits for Full Length Control Rod Position (0-200 EFPD) 7-22 7-7 Operating Limits for Axial Imbalance (0-200 EFPD) 7-23

'b '7-8 Operating Limits for Part Length Rod Position (0-200 EFPD) 7-24 8-1 Accident Analysis Review Process 8-20 1

,9-1 Boron Letdown Curves 9-11 9-2 Differential Boron Worth vs. Burnup 9-12 9-3 Differential Boron Worth 9-13 9-4 Integral Boron Worth 9-14 9-5 Equilibrium Xenon Worth Vs. Burnup at HFP 9-15 vii 1:

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1. INTRODUCTION The design of a commercial light water reactor is such that the reactor core is loaded with a specified number of fuel assemblies which are generally iden-tical in design but different in the amount of fissile material content. In the initial core the fuel assemblies differ in the initial enrichment of the fuel, and in subsequent fuel cycles they differ in the amount of the burnup of the fuel as well. The reactor is refueled at intervals varying from 6 to 18 months. The refueling of a reactor consists of removing part of the core (a certain number of irradiated fuel assemblies, the number and identity of which are determined by a fuel management scheme) and loading an equal number of fresh and possibly previously burned icel assemblies called the " reload batch." In general, after refueling, the neutronic, thermal-hydraulic, safety, and operating parameters of the core would be different from the previous fuel cycle. The design analyses required to determine the mechanical design, enrichment and number of assemblies of the reload batch as well as the core loading pattern, the nuclear and thermal-hydraulic characteristics of the reloaded core, and the safety analyses demonstrating the safety of operation of the reloaded reactor is called reload design.

This report describes the various aspects of the reload design. In the following paragraphs, a brief overview of the major elements of the reload design process and the reload design criteria are provided. Subsequent sections provide detailed discussion including descriptions of design methods, analytical formulation, and calculational' procedures of the major reload ( sign tasks used for Oconee reload design.

The reload design is essentially a series )f analytical exercises with the objective to design the reload core in such a manner that the reactor can be operated up to a specified power level for a specified number of days with acceptable safety criteria. It consists o. the development of the basic speci-fications of the reload batch (mechanical characteristics of the fuel assembly, fuel roo and associated structures, fuel enrichment, pellet dimensions; shape and enrichment, fuel stack length, fill gas pressure, number of assemblies, uranium loading, etc.); it sets forth the number and identity of each residual fuel assembly, selects the location of each fuel assembly and control rod in 1-1

the core for the new fuel cycle, establishes the core characteristics and operating-limits and protection system setpoints and demonstrates that the operation of the reactor during the new fuel cycle will be within safety con-siderations already evaluated and approved or provides new safety analyses to demonstrate conformance to applicable safety criteria. -

In arriving at the final reload design, the designer tries to meet the require-ments imposed by the operational considerations, fuel economics considerations and safety considerations. These requirements are called reload design criteria and are as follows:

1. Initial core excess reactivity rill be sufficient to enable full power operation for the der!. J 1ength of the cycle.

2. Technical Specification limits of specified core parameters (quadrant power tilt, power imbalance, control rod positions, xenon conditions, coolant flow).and on core protection system trip setpoints after allow-ance for appropriate measurement tolerances should have adequate margin from nominal values of these parameters during operatiqual corditions throughout the cycle so that sufficient operating flexibility is retained for.the fuel cycle.

3. 'The fuel assemblies to be discharged at the end of the fuel cycle will attain maximum permissible burnup so that maximum energy extraction con-sistent with the fuel mechanical integrity criteria is achieved.

4. Values of important core parameters (moderator temperature coefficient, Doppler coefficient, ejected rod worth, boron worth, total control rod group worth, maximum linear heat rate of the fuel pin at various eleva-tions in the core, and shutdown margin) predicted for the cycle are con-servative with respect to their values assumed in the safety analysis of various postulated accidents, and if they are not conservative, acceptable reanalysis of applicable accidents is performed.

5. The power distributions within the reactor core for all possible (or permissible) core conditions that could exist during the operation of the l-2

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l cycle will not lead to exceeding the thermal design criteria of the fuel and exceeding the LOCA-limited peak linear heat rates.  ;

6. Fuel management will produce fuel rod power and burnup consistent with the  ;

mechanical integrity analysis of the fuel rod--that is, the clad tensile  !

strain is less than 1%, the effective clad stress is less that the yield strength, and clad collapse will not occur during the life of the fuel. l l

7. The mechanical design of the reload fuel, including its vibration, flow, i

structural characteristics, and seismic and LOCA response, is compatible with the residual fuel.

The reload design process is comprised of the coordinated effort of designers and analysts from many areas, each of which generates specified information in a sequential and sometimes iterative manner to develop the final reload design, meeting the design criteria. The major elements of the reload design process

. are (1) fuel design, (2) fuel cycle design , (3) fuel mechanical performance analysis, (4) maneuvering analysis, (5) thermal-hydraulic analysis, (6) safety analysis and Technical Specification development, and (7) reload report, and (8) development of core physics parameters.

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The fuel design consists of the fuel assembly design (material selection, fuel rod lattice and fuel rod number specification, spacer grid design--number of spacer grids, material selection, mixing vanes, etc.- , and fuel assembly end fittings design) and fuel rod design (rod dimensions, cladding type and dimensions, pellet density and dimensions, design of fuel stack spacers, fuel stack length, fuel rod fill gas pressure and composition, and specified toler-l

-ances,on fuel rod design parameters). lua physical properties of the fuel assembly, fuel pin, and non-fuel regions established by the fuel design I ~ process are necessary input into other phases of reload design.

The fuel cycle design establishes the number and enrichment of the reload batch fuel assemblies, specifies _the number and identity of residual fuel assemblies, and determines the arrangement (location and orientation) of the fuel assemblies and the locations of control rods and their grouping in such a manner that the 1-3 1

specified criterion on energy production and certain specified criteria on fuel burnup, power distribution and control rod worth requirements are satisfied.

The fuel mechanical performance analysis consists of the evaluations to confirm that the fuel assembly mechanical performance with regard to vibrations, hy-draulic loading (fuel assembly lift-off), and seismic and LOCA reponse are acceptable. It also~ includes the evaluation of the extent of fuel densifica-tion and its effect; the evaluation of the fuel rod mechanical performance with regard to clad stress, strain, and clad collapse; and the evaluation of the extent of fuel rod bowing and its effects. In the absence of any changes in the mechanical design of the fuel assembly, no reevaluation of the mechanical performance of the fuel assembly is needed. The extent of fuel densification and its effects depends on the fuel fabrication process, the initial density of the fabricated fuel pellet, and the analytical model utilized for the evaluation; if any of these factors changes for a particular reload, a reanaly-sis of fuel densification and its effect is required. The fuel rod mechanical performance is influenced by the fuel density, operating RCS pressure, initial fill gas pressure, fuel rod design dimensions, fuel pellet density, and the predicted power history of the fuel rod; and if any of these factors changes for a particular reload, a reevaluation of the fuel mechanical performance is required. Fuel rod bowing is recognized to increase with fuel burnup. Since fuel rod bowing is considered to have the potential to decrease the thermal-hydraulic performance in certain flow channels in the core, an evaluation of the magnitude of rod bowing and its effect on DNBR is required for each reload considering the maximum expected fuel assembly burnup during that cycle. The thermal analysis establishes the maximum permissible power density of a fuel rod to preclude center fuel melting during steady-state and anticipated transient conditions. The thermal analysis needs to be repeated for a particular reload only if there is a change in the fuel design or there is a change in the regu-latory requirements.

The maneuvering analysis involves detailed power distribution evaluation in three dimensions by sinulating various anticipated and postulated' design con-ditions and is performed to confirm that the fuel cycle design is acceptable from the point of view of safety requirements. The data generated in the ma-neuvering analysis are used to define the core safety limits pertaining to 1-4

l the thermal design limits of the fuel and the limiting conditions on control rod position, axial-imbalance, and xenon distribution.

The thermal-hydraulic analyses establish the maximum permissible power distri-bution for various flow conditions, the permissible combination of core pressure and core coolant temperature and the minimum permissible core pressure to ensure that a minimum DNBR of 1.3 or greater can be maintained during anticipated tran-sients. The need to perform the thermal-hydraulic analysis in conjunction with a reload arises when there is a change in the fuel design, a change in the input assumptions of the original analysis, or a change in the regulatory criteria.

The results of the maneuvering analysis in conjunction with the results of thermal and thermal-hydraulic analyses, as appropriate, are used either to confirm that the existing Technical Specifications continue to be valid for the reload cycle or to generate new Technical Specifications limits. The accident analyses are reviewed to ensure that important core safety parameters predicted for the reload cycle are conservative compared to their values used

, in the existing accident analysis, and where necessary, appropriate accidents are reanalyzed.

1 A number of physics parameters pertinent to the reload cycle should be cal-culated to confirm that important core parameters for the reload cycle are conservative compared to their values used in the accident analyses. Other physics parameters are calculated to enable an orderly and safe startup of the cycle, to perform startup testing, and to perform core follow calculations.

These parameters include the critical boron concentration as a function of core i

burnup and for various conditions, ejected and stuck rod worths, control rod group and total group worths, reactivity worths of xenon and samarium, core excess reactivity, moderator and Doppler reactivity coefficients, reactivity worth of boton, effective delayed neutron fractions and decay constants.

The' final phase of the reload design is the integration and documentatior. of the results of all other phases into a report called the reload report. This report includes a description of the reload core, che fuel design, results of nuclear, thermal, thermal-hydraulic, and fuel mechanic.) saalysis, accident analysis review, and proposed Technis.13pectrications, if any.

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Considerable design and engineering effort is needed to incorporate a new fuel design into a reload design; and unless there is sufficient economic or engineering incentive or regulatory requirements, a new fuel design is 6 not usually attempted for a reload cycle. Therefore, a typical reload de-sign involves only a change in the fuel cycle design and as such is an ex-trapolation from the original core design, with many of the design variables being constrained by the original core design. In the following sections each of the major phases of the reload design process is discussed in more detail, and Figure 1-1 shows a flow chart of the various phases.

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ELEMENTS OF RELOAD DESIGN Fuel Design if Fuel Cycle Design i I _

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Fuel Mechanical Maneuvering Thermal Hy-I

! Performance Analysis draulic Design I

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9 P Technical Specification Development

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Accident Analysis Review 1P lf lf l f Reload Report and Development of Core Technical Specifications Physics Parameters 1

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2. FUEL DESIGN The fuel design consists of the fuel assembly design (material selection, fuel rod lattice and fuel rod number specification, spacer grid design--number of spacer grids, material selection, and fuel assembly end fittings design) and fuel rod design (rod dimensions, cladding type and dimensions, pellet density and dimensions, design of fuel stack spacers, fuel stack length, fuel rod fill gas pressure and composition, and specified tolerances on fuel rod design paramenters).

2.1 Fuel Pellet The fuel is in the form of UO cylindrical pellets which are sintered, ground, 2

and dried to obtain the desired density, dimensions, and moisture content.

The ends of the pellets cre chamfered to minimize the effects of differential thermal expansion. The pellet radius is constrained to be less than the clad inner radius and is set to accommodate radial growth resulting from a peak pellet burnup of 55,000 MWD /NTU without the plastic clad strain exceeding 1%. The pellet radius is also determined from economic and nuclear calculations performed to minimize fuel cycle costs.

2.2 Fuel Rod The fuel rod is composed of a cylindrical metal cladding, top and bottom fuel j spscers, and a column of fuel pellets. Zircaloy -4 has proven to be an accept-l able clad material because of its nuclear and material properties. The thick-ness of the clad is determined from heat transfer and clad strain and stress

! _ requirements. Fuel spacers hold the column of fuel pellets in position within the fuel rod and are designed to permit axial growth and thermal expansion of the fuel column. The void spaces at the top and bottom of the fuel rod are of

, sufficient volume to accommodate the fission gas release during the fuel burnup.

l The volume is designed to maintain the internal pin pressure less than the primary system pressure at coolant temperatures greater than 425*F for Condition I and II occurences. All fuel rods are pre pressurized with helium gas to 2-1

aid heat transfer, to prevent cladding collapse due to fuel densification effects,-and to avoid hydrogen contamination. The-fuel rods when loaded are

-sealed by Zr-4 endcaps welded at each end.

h.3 Fuel Assembly Design Fuel assembly design consists of specifying the number, location, material and fabrication techniques for fuel rods and non-fuel components such as guide tubes, instrument tubes, spacer grids, end fittings, and spacer sleeves.

Generally the number of fuel rods per assembly and the number of assemblies per plant-is determined by trying to limit the linear heat rate of the fuel. The fuel' pellet radius and rod pitch are determined by both neutronics and economic calculations aimed at minimizing the fuel cycle costs through optimizing the fuel-to-moderator ratio. The overall assembly pitch la constrained by the size and number of assemblies in the core.

Instrument tubes are typically locate 1 in the center of the fuel assembly and guide tubes.are synetrically dispersed throughout the assembly, and provide continuous guidance for the control rod assembly. both the control rod guide tubes and the instrument tubes are made from Zr-4.

The lower end fitting positions the assembly in the lower core plate. The lower ends of the fuel rod rest on the grid of the lower end fitting. The guide tubes penetrate the end fitting and are attached to it.

9 The upper end fitting positions the upper end of the fuel assembly in the upper core. grid plate and provides means for handling and identification of the assem-b!f ha internal hollow post in the center of the end fitting provides a means for supporting control rod assemblies and for retention of either an orifice rod assembly or a burnable poison assembly. Attached to the upper end fitting i is a holdcwn spring. This spring provides a holdown force to oppose hydraulic forces resulting from the primary coolant flow. These end fittings are cast stainless steel.

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r Spacer grids are constructed from strips of metal which are slotted end fitted together in an " egg crate" fashion. These spacer grids provide support for the fuel rods in the X-Y direction (keeps them at a fixed distance apart) and are located axially along the fuel rods to decrease the amount of fuel assembly and fuel rod bow. Oconee uses spacer grids constructed from Inconel-718 arranged in a 15 x 15 lattice. Each assembly contain.s eight of these grids. The spacer grids are held in position axially by the frictional grip force exerted on the fuel rods and guide tubes by the spacer grids.

Because of the considerable design, engineering and testing needed to incor-porate a new fuel design into a reload core, it is usually not attempted unless there is sufficient economic, engineering, or regulatory justification. If however sufficient justification exists, the new fuel design is typically documented in a generic topical report and the reload report would reference this topical report.

2.4 Core Component Data The basic physical dimensions and materials of the fuel pellet, fuel rod, fuel assembly, control rods, and orifice rods are used in the fuel cycle design, thermal-hydraulic design, and fuel mechanical performance. Table 2-1 contains a summary of this data for the B&W Mark-B2 assembly and is intended as an ex-ample.

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TABLE 2-1 OCONEE SYSTEM AND COMPONENT DATA MARK B2 CORE COMPONENT DESIGN DATA Fuel Assembly Designation Mark B2 Calculated Fuel Assembly Total Weight, lbs. 1550 Material Between Active Core Limits per Assembly, lbs.

Zircaloy 274 Inconel . 9.2 UO 1172 2

Fuel Assembly Cell Flow Areas, in. 39.76 Assembly Pitch, in. 8.587-Spacer Grid Assemblies Spacer Grid Material. Inconel 718 3

Spacer Grid Material Density, lbs/in 0.296 Number of Spacer. Grids per Assembly Total- 8 Between Active Core Limits 6 Spacer Grid Weight, lbs.

Intermediate, each 1.5 End, each 1.6 Total per Fuel Assembly 12.4 2-4

Fuel Rods Fuel Rod Pitch, in. 0.568 Fuel Rod Array 15 x 15 Number of Fuel Rods per Assembly 208 Fuel Rod Weight, lbs.

Per-Fuel Rod 7.0

' Clad- 1.15 Fuel Rod Clad OD, in. 0.430 Fuel Rod Clad ID, in. 0.377

' Fuel Rod Clad Wall Thickness, in. minimum 0.024

. Fuel Rod Length, in. 153.6875 Fuel Rod Clad Material Zircaloy-4 Fuel Rod Clad Material Density, lbs/in. 0.237 2-5

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Fuel Column Starts, From Bottom of Fuel Rods, in. 4-1/8 Fuel Pellet Diameter, in. 0.370 Fuel Pellet Length, in. 0 700 Pellet Density, g/cc (93.5% of Theoretical) 10.248 Pellet Dish Depth, in. 0.018 Pellet Dish Factor 0.983014 Diametrical Gap, in. 0.007 Fuel Column Length, in. 144 Unit Fuel Cell Flow Area, in. 0.1774 Material UO 2

U per Fuel Rod, kg 2.2530 U per Assembly, kg 468.6240 UO2 Per Fuel Rod, kg 2.5559 U02 Per Fuel Assembly, kg 531.6272 U/UO2 Ratio Used for Calculation 0.8815 Metal / Water Ratio 0.828227 Control Rod Guide Tubes Control Rod Guide Tubes Attachment to Spacer Grids None No. of Control Rod Guide Tubec per Assembly 16 Guide Tube Material Zircaloy-4 Guide Tube Material Density, lbs/in.3 0.237 Weight of control rod guide tubes, lbs.

Per Control Rod Guide Tube 1.06

. Total per Assembly 16.9 Between Active Core Limits 14.1 Guide. Tube OD, in. 0.530 Guide Tube ID, in. 0.498 Guide Tube Wall Thickness, in. 0.016 Guide Tube Length, in. 155.625 2-6 1

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Instrumentation Tube Number of Instrumentation Tubes per Assembly 1 l Instrumentation Tube Attachment'to Spacer Grids None Instrumentation Tube Material. Zircaloy-4 Instrumentation Tube Material Density, lbs/ft.3 0.237 Instrumentation Tube Weight, lbs.

Per Tube 1.40 Between Active Core Limits 1.30 Instrumentation Tube OD, in. 0.493 Instrumentation Tube ID, in. 0.441 Instrumentation Tube Wall Thickness,.in. 0.026

. Instrumentation Tube Cell Flow Area, in.2 0.0867 Length of Instrumentation Tube, in. 154.9375 Spacer Sleeves Spacer Sleeve, OD, in. 0.550

' Spacer Sleeve, ID, in.' O.498 Spacer Sleeve Material Zircaloy-4 Spacer Sleeve Weight, lbs.

' Per Assembly 1.39 m

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Spacer Sleeve Length, in.

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6 Sleeves at 19.593 1 Sleeve at 18.781 Length of 7 Spacer Sleeves, in. 136.339 Control Rod

~ Cladding OD, in. 0.440 Cladding ID, in. 0.398 Cladding Wall Thickness 0.021 Cladding Length, in. 149.500 Cladding Material SS304 Absorber OD,'in. 0.392 Absorber Length,' in. 134.0 Absorber Material Ag-In-Cd Axial' Power Shaping Rod Cladding OD, in. 0.440 Cladding ID, in. 0.398 Cladding Wall Thickness, in. 0.021 Cladding Length, in. 30.500 Cladding Material SS304 Absorber OD, in. 0.375 -

Absorber Length, in. 36 Absorber Material Ag-In-Cd Follower Tube OD, in. 0.440 Follower Tube ID, in. 0.376 Follower Tube Wall Thickness, in. 0.032

. Follower. Tube Length, in. 113.750 l

Follower Tube Material Zircaloy-4 1

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. Burnable Poison Rod ~

Cladding OD,'in. 0.430 Cladding ID, in.- 0.360 C1' adding Wall Thickness, in. 0.035

. Cladding Length, in. . 157.250 Cladding Materir.1 ..,

Zircaloy-4

- Burnable' Poison. Material OD, in~. 0.340 Burnable Poison.Lengh, in. 126.000

' Burnable Poisoa Material A1 0 ~

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' Orifice Rod

-Orifice'. Rod OD, in. 0.480

! Orifice Rod Length, in. 10.000 LOrificeRodLength,in. (0.489 d'ia'.' section) 7.000 iOrifice Rod Material SS304

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Orifice Rod Spring OD, in. '

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O.381 Orifice Rof Spring Wire Dia., in. 0.063 )

Orifice Rod Sprin's Material Inconel X-750 t

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3. FUEL CYCLE DCSIGN' 3.1 Preliminary Fuel Cycle Design The purpose of the preliminary fuel cycle design (PFCD) is to determine the number and enrichment of the fresh and possibly burned assemblies to be in-serted during the next refueling. A prelimiaary fuel shuffling scheme is

~ developed and check calculations on certain key parameters are performed.

The input required for the PFCD consists of general ground rules and design bases developed from cycle energy, contract, and operating requirements. The output of the PFCD is the number and enrichment of the feed assemblies.

The majority of this section will discuss the calculations necessary to deter-mine the above information.

3.1.1 Overview of Nuclear Calculational System The nuclear calculational system enables the nuclear designer to numerically mod-el and simulate the nuclear reactor core. The current system in use at Duke Power is outlined in Figure 3-1. A brief description of each code is included as Appendix A to this report.

3.1.1.1 Cross Section Generation t

In order to model the neutronics of a reload core, it is necessary to generate nuclear cross sections for subsequent use in a diffusion theory code. The cross section generator sometimes referred to as a sttice code is EPRI-CELL. Basic input consists of the geonietry and materials constituting the fuel assembly, I specific power, soluble boron concentration, temperature for the pellet, clad,

-and moderator, effective resonance temperature, icel enrichment, number of deple-tion steps and the length of each step, etc.

3-1

The output of the-EPRI-CELL run is a set of nuclear cross sections which charac-

.terize that particular assembly. These are then formatted into table set structure through the use of an auxiliary code called NUPUNCHER for input to PDQi7.

l This procedure is repeated for the different fuel. assemblies comprising the re-l

actor core as well as for a spread of enrichments covering the typical range of re' load enrichments.

Non-fuel cross sections with the exception of burnable poison assemblies and con-76 trol rods are.also generated using EPRI-CELL. Cross sections for burnable poison assemblies and control rods for use in diffusion theory calculations are generated by matching reaction rates between the diffusion theory code PDQ97 and CPM (a' collision probability code).

3.1.1.2 Diffusion Theory Calculations Once the cross sections have been generated they are input to PDQ97 which solves the diffusion depletion equations in one, two, or three dimensions.

Two types of PDQ97 calculations are routinely run; the color set calculation and the quarter core calculation. The color set calculation is a single assembly or four quart r assemblies pin mesh x-y geometry PDQ97. This is used to calculate assembly parameters necesscry for input to EPRI-NODE.

The quarter core PDQ97 is a pin mesh x y geometry calculation which yields radial power distributions, local pin peaking, reactivity versus burnup, and

<other nuclear parameters of interest.

h b 3.1.1.3' Nodal Calculations The output from the color set PDQW7 calculations discussed in Section 3.1.1.2 is processed by.the auxiliary codes EPRI-FIT, and SUPERLINK and input to the nodal code EPRI-NODE. EPRI-N0DE is a three dimensional nodal code that is currently in use to provide three dimensional information such as three dimen-sional power distributions and integral and differential rod worth, etc.

l 3-2

l I

3.1.2 Calculations and Results of PFCD Once the calculational models are prepared for the cycle of interest (steps dis-cussed in 3.1.1 are complete), the nuclear designer chooses a feed enrichment, number of assemblies, and preliminary loading pattern for the reload core.

Calculations using either or both EPRI-NODE and PDQ97 are performed to veri-fy cycle lifetime and power peaking. The process is iterated until the num-ber and enrichment of feed assemblies as well as a preliminary shuffle scheme has been determined which yield the desired cycle lifetime and a reasonable power distribution.

The preliminary number and enrichment of the feed assemblies must typically be determined eighteen months prior to reactor shutdown for refueling to assure that an adequate quantity of separative wurk is available. Changes to these preliminary estimates are normally possible up to twelve months prior.to reactor shutdown. Therefore, it is necessary that the results of the PFCD be complete at that time.

3.2 Final Fuel Cycle Design Having determined the number and enrichment of the fuel assemblies during the PFCD, the final fuel cyt;2 design (FFCD) concentrates on optimizing the placement of fresh and burned assemblies, control rod groupings, and burnable poison assemblies (if any) to result in an acceptable fuel cycle design. It must meet the following design criteria with appropriat.e reductions to account for calculational uncer-tainties:

1. Radial Pin Peak < 1.714 (inner flow zone)

< 1.629 (outer flow zone)

2. Moderator Temperature Coefficient 1 0.0 at >95% hot full power.

3. Maximum pellet burnup 5 55,000 MWD /MTU.

4. Shutdown Margin at HZP > 1.0 % Ap.

3-3

5. Ejected rod worth at HZP $ 1.0% Ao.

6. Ejected rod worth at HFP $ 0.65% ap.

During the FFCD, nuclear calculations are performed to generate these parameters for input to fuel mechanical performance,-thermal and thermal-hydraulic perfor-mance, maneuvering analysis, and accidents and trap-tects analyzed during the safety analysis.

3.2.1 Fuel Shuffle Optimization and Cycle Depletion ,

Beginning of cycle (BOC) power distribution calculations are performed using com-binations of EPRI-SHUFFLE and PDQ97. Initial runs start with the fuel shuffle scheme developed in the PFCD and modify the shuffle scheme in an attempt to minimize the power peaking. This is accomplished by a trial and error type search until an acceptable BOC power distribution results. .The cycle is then 1 depleted using point depletion in steps corresponding to 0, 4, 12, 25, 50, 100, 150...EFPD to verify that power peaking versus burnup remains acceptable. The shuffling variations include re-arranging the location of the burned or fresh

. fuel' assemblies, location of control rods (groups 5, 6, 7) and rotation of

- the spent fuel assemblies. These calculations are typically performed assuming quarter core symmetry.

The shuffle pattern determiped by optimizing power distribution may later need to be' modified based upon results obtained in the ra map'.ing nuclear calculations.

d 3.2.2 Rod Worth Calculations Control rods serve several functions in the Oconee reactor. The primary func-tion is'to provide adequate shutdown capability during normal and accident con-

-ditions. They are also used in the " rods in" mode to maintain criticality during power maneuvers and to compensate for reactivity loss due to fuel 4

3-4 m - , - -w +- v: , -- r,-g er, m , ,-m4--< - - -

l depletion. Since the presence of control rods influences both power distri-butions and criticality, it is necessary in many calculations to evaluate not only the reactivity effect but also the perturbation that a given rod config-uration has on the power distribution.

Oconee may be designed and operated in either a " rods out" (feed and bleed) or, a " rods in" mode. A typical " rods in" design allows for the first regulating bank (7) to be almost fully inserted into the reactor core during HFP operation.

This regulating bank is typically withdrawn two months prior to the end of the fuel cycle to provide the additional reactivity to operate to E0C. Shutdown margin for a " rods in" design is typically less before the regulating bank is removed than at E0C and is normally calculated for that point in the cycle instead of EOC.

Most calculations of control rod worth are used in the safety analysis of the re-load. core. The calculations discussed in subsequent sections include the follow-ing:

1. Choice of Control Rod Groupings and Worths

2. Shutdown Margin

3. Ejected Rod Worth

4. Dropped Rod Worth 3.2.2.1 Control Rod Groupings and Worths Control rod locations in Oconee are fixed, however, the rods in a particular group may vary from cycle to cycle. The contrcl and rod groupings are determined by nuclear calculations to evaluate the effects that a particular rod grouping has on power distribution, group worth, and ejected rod worth. The worth of each regulating bank (5, 6, 7) is calculated in quarter core geometry using either PDQ97 or EPRI-N0DE at BOC, EOC, and before and after control rod inter-change at HFP and HZP. .The total rod worth (all rods in) is calculated at B0C, EOC, and any limiting burnup at HZP only for use in the shutdown margin calcu-lation.

The locations chosen during the FFCD are confirmed during the maneuvering analy-sis.

3-5

3.2.2.2 Shutdown Margin Searches for the highest worth stuck rod are performed at BOC, EOC, or any limit-ing burnup for H2P conditions using full core EPRI-N0DE and/or PDQS7 calculations.

Table 3-1 summarizes the results of a shutdown margin calculation. The total rod worth described in section 3.2.2.1 is shown as Item 1. This value is reduced by an estimate of the worth lost by the control rod poison due to burnup (Item 2).

' tem 3 is the worth of the highest stuck rod. The total worth reduced by the con-trol rod burnup penalty and the stuck rod worth is shown as the net worth (Item 4). A calculational uncertainty of 10% is subtracted off in step 5 and step 6 shows the available rod worth.

The required rod worth is calculated next in steps 7-10. The power deficit ob-tained by running an EPRI-NODE or PDQ97 case at HFP and a second case at HZP and subtracting the reactivities is shown as item 7. The maximum allowable inserted rod worth, item 8, is obtained from the allowable rod insertion and the integral rod worth curve for that insertion (generated by EPRI-NODE). This accounts for any rod insertion allowed at HFP. An axial flux redistribution occurs when the power level is reduced from HFP to HZP. This redistribution causes an increase in reactivity.

If Item 7 is calculated using a 3-D code such as EPRI-N0DE no additional penalty is required here. However, if Item 7 was calculated using 2-D PDQ97 an additional reactivity penalty is assessed as Item 9. The sum of these required worths (Item

10) is the total required worth.

The shutouwt :rgin is shown as Item 11 and is defined as the total available worth minue the total required worth. For Oconee this is required to be > 1.0%

Ap.

l l 3.2.2.3 Ejected Rod Worth I

The maximum allowable ejected rod worth is limited by the Technical Speci " eations.

For Oconee these limits are shown in Table 3-2. A calculated limit for setting rod withdrawal positions has been established by applying a 15% uncertainty to the Technical Specification limits.

3-6

To verify that the ejected rod worths are within this calculational limit, ejected rod worth calculations are performed at BOC and EOC at both HFP and HZP.

The calculation of ' ejected rod worths is accomplished using full core two di-mensional pin mesh PDQ97 or EPRI-NODE calculations. The HZP ejected rod worth calculations are performed with control groups 5 through 7 fully inserted in the core and with group 8 centered.

Single rods in groups 5, 6, and 7 are removed in subsequent cases and the worth of the ejected rod is calculated by subtracting the reactivities of the cases before and after the rod was removed. The fuel and moderator temperature is held constant and equal to the HZP moderator temperature for these calculations. The highest worth calculated by the above procedure is the worst ejected rod at HZP.

• If the ejected rod worth exceeds the calculational limit, rod position limits are

' imposed or a new control rod grouping is developed.

' The HFP ejected rod worths are performed in a similar manner to the HZP calcu-lations with the exceptions that for a " rods out" design only groups 7 and 8 are inserted into the core and that the fuel temperature and moderator tempera- 1 tures correspond to those of HFP conditions. The HFP ejected rod worths are performed without thermal feedback to be conservative. If the ejected rod worth exceeds the calculational limit, rod position limits are imposed or a new control rod grouping is developed.

3.2.2.4 Dropped Rod The analysis of the rod drop worth is required to determine the reactivity insertion resulting from the rod drop. Full core calculations using EPRI-NODE are performed with thermal-hydraulic feedback.

Search cases are run dropping full length control rods until the highest worth rod has been identified. This value of dropped rod worth is used as input to the accident analysis evaluation.

3-7

,= -

1 l

l l3.2.3 Power Distribution Calculations

, For Oconee emphasis in the FFCD is on radial power distributions both cn an as-sembly and local rod basis. Thermal and thermal-hydraulic analyses have been per-formed on the Oconee reactors which indicate that radial pin peaks shown in l

Table'3-3 will result in acceptable DNB and Center Fuel Melt (CFM) margins.

- These margins are calculated and confirmed during the maneuvering analyses.

Power distributions are calculated both with the 2-D PDQ97 model and the 3-D EPRI-  !

NODE model. The local radial peaking factor, maximum pin to assembly power ra-tio, calculated by PDQ97 is multiplied by the three dimensional peak calculated by EPRI-NODE to produce maximum three dimensional power peaks in a fuel rod.

)

3.2.4. Fuel Burnup Calculations 4

One of the current design criterion is that maximum pellet burnup is < 55,000 MWD /MTU. This criterion _is confirmed during the final fuel cycle design.

Depletion calculations from 2-D quarter core pin mesh PDQ$7 models yield core ,

and assembly average burnup as well as single fuel rod burnups. From these

' values a maximum ratio of-single rod to assembly average burnup can be cal-culated for each assembly. This data is then used in conjunction with 3-D EPRI-NODE depletion cal.culations (where'the axial burnup distribution is calculated) to arrive at a local burnup limit which can be compared to the design limit of 55,000 MWD /MTU.

Generally, the assembly average burnups are in the 30,000 to 33,000 MWD /MTU range

and sufficient margin to the 55,0p0 MWD /MTU limit exists to make the detailed cal-

- culation described above unnecessary.

3-8 w , e , - - - ~ - - - , ,e- ,-w-,--,-,,,-.,-,r -- - - n,--- , , , , , , - - - , , , - - ,,--

3.2.5 Reactivity Coefficients and Deficits

~ Reactivity coefficients define the reactivity insertion for small changes in re-actor parameters such as moderator temperature, fuel temperature, and power level.

These parameters are input to safety analysis and used in modeling the reactor re-sponse during accidents an.d transients. Whereas reactivity coefficients represent reactivity effects over small changes in reactor parameters, reactivity deficits usually apply to reactivity inserted from larger changes typical of HFP to HZP.

An example of a reactivity deficit is the power deficit from HFP to HZP used in the shutdown margin calculation. A different way of looking at the terms is that.

the coefficient when integrated over a given range yields the deficit, or the co-efficient is the parcial derivative of reactivity with respect to one specific parameter.

Coefficients of reactivity are calculated using PDQ97 or EPRI-N0DE. First a nominal case is established at some reference conditions. Then one parameter of interest is varied up and/or down by a fixed amount in another calculation and the resulting change in core. reactivity divided by the parameter change is the reactivity coefficient.

3.2.5.1 Doppler Coefficient The Doppler Coefficient or Fuel Temperature Coefficient (FTC) is the change in core reactivity produced by a small change in fuel temperature. The major com-ponent of the Doppler coefficient arises from the behavior of the Uranium-238 and Plutonium-240 resonance absorption cross sections. As the fuel temperature in-creases, the resonances broaden increasing the chance that a neutron will be ab-sorbed and thus decreasing the core reactivity.

If Case 1 represents the reference case with an effective fuel temperature T 3 (and KI effective) and Case 2 represents a second case where the fuel temperature has been increased or decreased by approximately 50*F and is2 T , (and 2K effective) the Doppler coefficient is mathematica, ty calculated from the following equation:

~

eff ff eff

• ff = ap/ F Ti-T2 3-9

In the final fuel cycle design both HFP and HIP. Doppler coefficients are calcu-lated.

k

~3.2.5.2 Moderator Temperature Coefficient The Moderator Temperature Coefficient (NTC) is the change in reactivity produced by a

'small change in' moderator temperature. .In Oconee the average core moderator temp-erature is increased as power is escalated from 0 to 15% HFP. At and above 15% HFP the. average moderator temperature is held constant at 580 F. However, for acci-dent and transient analyses it is necessary to know the moderator temperature co-efficient at HFP and also at HZP.

-These analyses can be performed with either EPRI-NODE and/or PDQ97 by effecting a

change in the. core average moderator temperature. Two cases are run with the moderator temperature at +5"F and -5 F around the HZP (12.5 F at HFP) average 1 temperature.- If these cases and the resulting K effectives are identified as.

2 Case-l'(TMOD,Kjgg)andCase2(TMOD,K ff), the moderator temperature co-i 2

. efficient is' calculated from the following equation:

K,ff - K,f f .

HTC =

K,f f . x Keff =. Ap/*F (TMOD2 - TMOD2 )

i.

l 3.2.5.3 Temperature Coefficient-

-The' fractional change in reactivity due to a small change in core temperatures is defined as the co2e temperature coefficient of reactivity. This is equal to the sum of the moderator and Doppler temperature coefficients and may be explicitly

calculated at HZP for isothermal conditions (TFUEL=TMOD) by varying both the fuel

' and moderator temperatures approximately +5 F about the average moderator temp- 1 ferature at HZP. Similarly the temperature coefficient at HFP may be explicitly calculated by varying the moderator and fuel temperatures by 12.5 F. This calcu-

,. lation may be performed with PDQ$7 and/or EPRI-NODE.

a

i. 3-10

- , - . - , .--,=.-..:.:,.. ..- -.,. - - - , , - - - - - . . ~ . . - - .

3.2.5.4 Power Coefficient and Power Deficit The power coefficient of reactivity is the core reactivity change resulting from an incremental change in core power level. The power deficit is usually the to-tal reactivity change associated with a power level change from HZP to HFP.

The power coefficient is defined by the following equation:

K1 eff

-K2eff GP = Ki gg xK2 gf Pt -P2 where: K1 gg is K effective for the core at power P1 (%)

K2 fg is K effective for the core at power P2 (%).

Neglecting second order effects this equation is equivalent to the following:

ATMOD + FTC ATFUEL ap = MTC AP AP where: NTC is the moderator temperature coefficient and FTC is the fuel tempera-ture coefficient (Doppler coefficient).

. In Oconee the core average moderator temperature is constant at approximately 580 F above 15% IIFP. Therefore, for power levels above 15% HFP the power coef-ficient can be reduced to just the fuel temperature contribution or ATFUEL ap - FTC AP Since the power coefficient should include flux redistribution effects resulting from axial variations in burnup and isotopics as well as non-uniform fuel temp-erature distributions it should be performed using a 3-D simulator with thermal hydraulic feedback. If the calculation is performed using a 2-D model then it should be corrected for the 3-D effects.

3-11

A typical power coefficient calculation for HFP would proceed in the following manner. The HFP case is run using EPRI-NODE and the core Keffective is calcula-1 ted (K gg g ). Then a second EPRI-NODE case is run with the core power level reduced 5% while holding everything else constant. The Keffective from this case, Keffective, is used along with the results from the reference case to calculate the power coefficient:

K1 eff

-K2eff eff

• eff op = .P t -P2 M

% POWER The power deficit is calculated for use in the shutdown margin calculation (see Section 3.2.2.2) and is the reactivity change from HZP to HFP. This calculation should be performed in three dimensions to satisfactorily model the axial flux redistribution, however, a two dimensional calculation may be performed and cor-rected for this flux redistribution phenomenon. Two EPRI-NODE or PDQ97 runs are made to calculate the power deficit. The first is made at 100% HFP and the sec-ond at HZP. These calculations are usually performed at least two times during the cycle burnup.

The HFP and the HZP case should have the equilibrium xenon concentration cor-responding to HFP. The power deficit is calculated from the following equation:

~

eff eff Power Deficit = Kjgg x K,ff2 x 100 = W HZP - HFP where: K,gg,ctgy, is core Keffective at HZP and K effective is core Keffective at HFP.

3.2.5.5 Miscellaneous Coefficients For reload design, certain coefficients of reactivity are not routinely calculated.

These include moderator density coefficient, moderator pressure coefficient, and moderator void coefficient.

3-12

3.2.6 Boron Related Parameters ,

1 Critical boron concentrations at BOC and EOC for HFP and HZP and various rod posi-tions, are calculated using PDQ97 and/or EPRI-NODE. Table 3-4 lists conditions that critical boron concentrations and boron worths are calculated. In addition to these, an all rods out (ARO) critical bo'ron letdown curve, and a critical boron letdown curve at nominal rod index are generated for HFP equilibrium conditions.

3.2.7 Xenon Worth The HFP equilibrium xenon worth is calculated at BOC (4 EFPD) and at EOC. These values are compared to previous cycle values in the reload report.

Calculations using either PDQ$7 or EPRI-NODE are performed for HFP equilibrium xenon conditions. If PDQ97 is being used, a second no xenon case is run by either zeroing out the number density for xenon or zeroing out the xenon cross section.

If EPRI-NODE is being used the power level on the xenon card can be set to zero and the time in hours set to 40.0 to obtain a no xenon concentration.

The difference in reactivities between the equilibrium and no xenon cases is the xenon worth.

3.2.8 Kinetics Parameters The kinetics behavior of the nuclear reactor is often described in terms of solu-tions to the Inhour equation for six effective groups of delayed neutrons. ,

Transient and accident analyses often involve kinetic modeling af the reactor core. The rate of change in power from a given reactivity insertion can be cal-

\ .

l culated by solving the kinetics equations if the six group effective delayed neu-l tron fractions, the six group precursor decay constants, and the prompt neutron lifetime are known.

The computer codes used to calculate these parameters are PDQ97 and DELAY.

PDQ97 is used to obtain spatially averaged isotope fission fractions as a function l

of burnup, and DELAY calculates kinetics parameters and then uses these parameters to solve the Inhour equation and thereb- relate the stable reactor period to the 3-13 I

reactivity insertion. This information is also needed for startup physics test-i ing. -The sum of the six group p effective, p :.ffective, for the new reload cycle is compared to that of the previous cycle in the reload report.

a s

e 4

e 3-14 m , -~ -- , , ,, , - , - -

---,1- -,--- - . - - - , , -

. , - - - - - , - + - - . - , - + - - - , - - - - - -

Table 3-1 Shutdown Margin Calculation BOC, % ap EOC, % ap Available Rod Worth

1. Total rod worth, HZP 8.91 8.79

2. Worth reduction due to burnup of poison material .36 .42

3. Maximum stuck rod, HZP -2.17 -2.01

4. Net Worth 6.38 6.36 -

5. Less 10% uncertainty -0.64 -0.64

6. Total available worth 5.74 5.72 Required Rod Worth

7. Power deficit, HFP to HZP 1.31 2.12

8. Max allowable inserted rod worth 0.40 0.60

9. Flux redistribution 0.59 1.20

10. Total required worth 2.30 3.92

11. Shutdown Margin (total avail.

worth minus total required worth) 3.44 1.80 Note: Required shutdown margin is 1.00% ap.

3-15

b %

Table 3 Ejected Rod Worths Condition Technical Specification Calculational Limit Limit HZP (Banks 5-8 in.)- 1.00% ap 0.85% Ap 1

. MFP (Banks 7-8 in.) 0.65% Ap 0.55% Ap SP l

l 3-16 I

Table 3-3 Radial Pin Peak FLOW ZONE Maximum Allowable Calculational Radial Pin Peak Limit

• Inner 1.714 1.587 Outer 1.629 1.508

• An 8% reduction has been applied to the radial peaks to account for calculational uncertainty and to provide a margin for the 3-D calculations performed during the maneuvering analysis.

3-17

Table 3-4 Boron Parameters Critical Boron - BOC, ppe (no Xenon)

HZP, group 8 inserted HZP, Eroup 7 and 8 inserted HFP, group 7 and 8 inserted Critical Boron - EOC, ppe (equilibrium Xenon)

HZP, group 8 37.5% WD HFP, group 8 37.5% WD Boron Worth - HFP, pps/%6p BOC (xx ppe)

EOC (17 ppe) f 1'

O e

1 3-18

Figure 3-1 NUCLEAR FLOW CHART FOR EPRI-ARMP

. EPRI-CELL 1V NUPUNCHER > 4 4 CPM or CASMO p 1 f

I - 1F 1F

- PDQW7COLORSETf 2-D 1/4 CORE PDQ97 n

ir EPRI-SHUFFLE EPRI-FIT 1F SUPERLINK m v

k 1V EPRI-NODE 4 1r Desired 3-D Information J l 4

3-19 i

t '1 e-- -"t- ---e-'--- ** ++w=-* --e+-=-w-- wee-e--'-- -

ee---- ~~-a-- ----e--- -----

4. FUEL MECHANICAL AND THERMAL PERFORMANCE 4.1 Introduction Each fuel cycle design requires that thorough fuel mechanical and thermal assessments be performed. A reload design utilizes fuel designs that are bound by previous fuel assembly design analyses.

Occasionally, however, minor differences in the design will occur (such as a change from 94% TD fuel to 95% TD fuel). These changes must then be assessed in regard to the following:

Cladding creep collapse,

• Cladding strain,

+ Cladding stress, Fuel pin temperature, and

• Fuel pin pressure Design analyses that envelope the operation of all current fuel designs have been completed by the fuel vendor, and reanalysis is normally not required for a new fuel cycle design. Rather, a specific fuel cycle design is compared against the enveloping design analyses. The assessment must compare cladding and pellet designs against the pellet and cladding geometries and densities, etc., that have been considered in the enveloping design analyses.

Further, the individual radial power histories during the fuel cycle (current and previous batches)-must be compared against the generic radial power envelopes that hsve been used in the design analyses.

In most cases, the design analyses will envelope the fuel cycle design being considered and no reanalysis is required. However, in some cases, either the radial power history or fuel geometry may lie outside of the enveloping design analyses, thus requiring partial or full reanalysis. The following subsections describe the types of comparisons that must be made to justify a fuel cycle de-sign without reanalysis and provides some detail concerning the types of analyses that must be performed if required by either the fuel cycle design or by changes in the fuel design itself.

4-1 Rev. 2

Table 4-1 presents a summary of all types of fuel thermal and me-chanica) performance assessment criteria that are used to determine whether a fuel cycle design, the cladding, and the pellets are enve-loped by existing analyses. As shown in Table 4-1, several of these analyses require either a comparison against a fuel pin power ver-sus burnup envelope or a comparison against an assembly radial Power versus burnup envelope. Examples of these power history en-velopes are presented in Figures 4-1 and 4-2. These envelopes change, as reanalysis is occasionally required, resulting in an 4

- expanded power history envelope. Figure 4-3 presents a flow chart

- for the fuel pin pressure and linear heat rate to melt analyses.

Figure 4-4 is a mechanical analysis flow diagram.

f 4.2 Claddina Collapse

-Cladding creepdown under the influence of external (system) pres-sure is a phenomenon that must be evaluated during each reload fuel cycle design to ensure that the most limiting fuel rod does not exceed the cladding collapse exposure limit. Cladding creep is a function of neutron flux, cladding temperature, applied stress, cladding thickness, and-initial ovality. . Acceptability of a fuel cycle design is demonstrated by comparing the power histories of all l the fuel assemblies against the generic assembly power history used in existing design analyses, similar to Figure 4-2. The ge-neric power history must be completely enveloping.to avoid reanaly-sis. Duke Power Company uses its own PDQ edit code to automatically perform this comparison for all' fuel assemblies at each depletion step. Changes in pellet or cladding design are also assessed in a

similar manner: direct comparison with the fuel rod geometries of Table 4-1 and reanalysis, if necessary. 'Four separate fuel designs have been analyzed to form the generic cladding creep collapse analysis.

The CROV1 computer code calculates ovality changes in the fuel rod cladding due to thermal and irradiation creep and is used to perform the fuel rod creep collapse analysis when required. CROV predicts

. 4-2 Rev. 2

,e,-we m- ---- , - - ,e, ew- -

wem- - - , e,rr- - ,-r--. ,-r-- -->e,em,, -

-wav v ,,me--- - , - . -s ep -wne,*e r- y -v m e - e -- ,,m ,vv-- -, -n ~ g v- r r- ww -

the conditions necessary for collapse and the resultant time to collapse. Conservative inputs to the CROV cladding collapse analy-sis include the use of minimum cladding wall thickness and maximum initial ovality (conservatively assumed to be uniformly oval), all as allowed by manufacturing specifications. Other conservatisms included'are minimum prepressurization pressure and zero fission

- gas' release. Internal-pin pressure and cladding temperatures, Linput to CROV, are calculated by TAC 0 s1 (or TACO 2s when approved) l Rev. 4 <

using a radial power history similar to that of Figure 4-2, a

. generic pin to assembly local peak, and a standard axial flux shape.

The conservative fuel rod geometry and conservative power history are used to predict the number of EFPH required for complete clad-ding collapse. To demonstrate acceptability, the maximum expected residence time of the cycle is compared against the EFPH required for complete collapse.

4.3 Cladding Strain Analysis t

' The limit on cladding strain is that uniform strain of the clad-ding should not exceed 1.0%.

A generic strain analysis has been completed by the fuel vendor using TACO (or TACO 2 when approved) to ensure that the strain l Rev. 4

_ criterion above is not exceeded. To determine whether the fuel '

and fuel cycle designs are enveloped by existing analyses, the criteria of Table 4-1 are reviewed.

E Should reanalysis be required, TACO (or TACO 2 when approved) will l Rev. 4 be used to determine the fuel rod dimensional changes that occur

-between the two power levels considered by the conservative desigh

. power ramp used in the strain analysis. Then, the maximum tensile L (elastic and plastic) strain, which occurs at the cladding I.D., is determined from the following equation:  !

j. 4-3 Rev. 4

, . , - ..--,-,...n-.- ,,,--,-,-,--e-- , -.-- , ,-,- .--- --,,- - ,v , ~ - - - - -

.. - - . _ - _ _ . . .~

[Jd:

JStrain = (Pellet 0.D.) peak - (Pellet 0.D.) x 100

<g

~

(Pellet 0.D.)

o where (Pellet 0.D.) peak = the maximum pellet 0.D. at the local pcuar peak, and (Pellet 0.D.), = pellet 0.D. prior to and after a local power

- ramp.

Pellet 0.D. dimensions are used to calculate cladding strain be-

)

cause the strain itself is caused by pellet thermal expansion. l The strain analysis is completed in two parts:

Part 1 employs TACO (or TACO 2 when approved) to determine when Rev.

pellet contact occurs. A conservative fuel rod geometry is used in conjunction with a < l.5 axial flux shape, and the core average linear heat rate at 100% power to characterize gap closure. If contact occurs prior to 30,000 MWD /NTU, then Part 2 will use a ramp from 2 KW/FT to a final linear heat rate that is consistent with centerline fuel melt. Whereas, if contact occurs after 30,000 MWD /NTU, then the ramp's peak linear power is reduced to a lower value that is consistent with maximum local powers that could occur at burnups greater than 30,000 MWD /MTU.

Part 2 of the strain analysis is the power ramp calculation, also performed on TACO (or TACO 2 when approved), which calcu- Rev.

lates the change in fuel pellet 0.D. that occurs from the change in power level induced by the power ramp. The change in pellet 0.D. is then used to perform the hand calculation of cladding strain using the equatica above. The cladding and pellet are assumed to be in hard contact at the initiation of this ramp.

! Thus, there are two major conditions in this scenario that make i it conservative. The first is the extreme power change that is used to simulate the worst case peaking. The second is tha't the pellet is assumed to be in hard contact at inititation of the ramp. This is a conservative assumption since the power ramp is 4-4 4

Rev. 4

-initiated from 2 KW/FT, and pellet / cladding contact is not expected to occur at this low linear heat rate.

4.4 ' Cladding Stress Analysis The cladding stress analysis for a new fuel cycle design is sini-larly bounded by a conservative design analysis that uses Section III of the ASME Boiler and Pressure Vessel Code as a guide in classifying the stresses into various categories, assigning appro-priate limits to these categories, and combining these stresses to determine stress intensity. Each new fuel cycle design is assessed against tae criteria stated in Table 4-1 to determine if reanalysis is required. The stress analysis is very conservative, and reana-lysis should not be required for standard Mark B reloads. However, an assessment is made for each reload design using the criteria of Table 4-1.

The fuel rod stress analysis considers those stresses that are not relaxed by small material deformation, and this anaysis complies with the following design critera:

- All fuel cladding stresses (primary and secondary) shall not exceed minimum unirradiated yield strength for condition I and II occurrences.

• The stress intensity value of the primary membrane stresses in the fuel rod cladding, which are not relieved by small material deformation of the cladding, shall not exceed 2/3 of the minimum unirradiated yield strength.

The above criteria keep the primary loads well belon material allowable.

In performing the stress analysis, all the loads were selected to represent the worst case loads and were then combined. This repre-4-5 Rev. 2

sents a conservative approach since they cannot occur simultaneously.

'This insures that the worst case conditions for condition I and II events are satisfied. In addition, these input parameters were chosen so that they conservatively envelope all Mk-B design condi-tions.

'The primary membrane stresses result from the compressive pressure loading. . Stresses resulting from creep ovalization are addressed in.the creep collapse analysis.

Since the internal fuel rod pressure cannot exceed system pressure for condition I and II occurrences (at coolant temperatures greater than 425'F), the need to address tensile stresses at hot zero power ,

(NZP) conditions and higher is eliminated. The tensile stresses were addressed at cold conditions. The minimum internal fuel rod pressure at HZP conditions is combined with the maximum design

, system pressure during a transient to simulate the maximum pres-sure differential across the cladding. The tensile stress analyzed occurs at cold (room temperature) conditions at BOL. This is the worst case since the grid loads will be maximum at that point.

The worst case compressive pressure loads were combined with the other worst case loads. These are described below:

The maximum grid loads will occur at BOL. During operation, the contact force will relax with time due to fuel rod creep-down and ovalization as well as grid spring relaxation.

The maximum radial thermal stress will occur at the maximum rated power (power level corresponding to centerline fuel melt).

This stress cannot physically occur at the same time the maximum pressure loading occurs, but is assumed to do so for conserva-  ;

tism. (Maximum cladding temperature gradient is combined with l s minimum pin pressure.) l l

1 I

l I

4-6 Rev. 2

. _ , . _ _- - . - _ _ _ . . . . . _ , _ _ . . _ , _ _ . _ _ _ _ _ _ , _ . ~ . _ _ _ . _ _ _ _ . _ _ _ . . _ . _

4

- The ovality bending stresses are calculated at BOL conditions.

A. linear stress distribution is assumed. The creep collapse analysis calculates the stress increase with time and ovali-zation.

• Flow induced vibration and differential fuel rod growth stresses are also addressed.

- 4.5 Fuel Pin Pressure Analysis 4 'The pin pressure analysis is assessed against the design basis analysis criteria and envelopes *as indicated in Table 4-1. If any of the parameters of this table are violated, then a resnalysis is performed.

Pin' pressure analysis is performed using TACO (or TACO 2 when Rev.'4 approved). The rod is assumed to have a 1.5 symmetric axial flux shape, with a pin power history similar to that presented in Figure 4-1. Incore fuel densification is minimized in this analysis to yield a smaller plenum volume and a maximum pin pressure.

. Figure 4-5 presents the result of an analysis of pin pressure versus burnup, performed by Duke Power Company, using TACO (or TACO 2 when Rev. 4 approved). This analysis was performed for an extended burnup fuel cycle design, using the pin power history indicated in Figure 4-1,

. but with lower, more realistic axial flux shapes than the 1.5 cosine shape that is used for Reload Design purposes. (Refer to Table 4-2 for the axial flux shapes used in this extended burnup analysis.) ,

To satisfy mechanical design criteria, pin pressure must be less than system pressure (2200 psia).

4.6 Linear Heat Rate Capability Linear heat rate capability of all fuel rods in a relcad batch is assessed by comparison against the criteria and envelopes of Table 4-7 Rev. 4

p 4-1. Any rod whose geometry or power history falls outside of those criteria must be reanalyzed.

The Linear Heat Rate to Melt (LHRTM) analysis is performed using TACO (or TACO 2 when approved), assuming naximum incore pellet Rev. 44 4

densification. This analysis assumes a conservative pin power history, similar to that of Figure 4-1, and a 1.5 cosine axial flux shape. In this analysis, very small axial segments of the fuel rod are spiked to high linear heat rates at each burnup step until centerline fuel melt occurs. The resulting heat rate required to reach centerline fuel melt at each burnup is then plotted versus burnup.

Figure 4-6 is a plot of fuel LHRTM versus burnup for an extended burnup fuel cycle design. This TACO (or tat 02 when approved) Rev. 4 analysis, performed by Duke Power Company, represents the pin power history of Figure 4-1, but with more realistic axial flux shapes than the 1.5 cosine that is used for reload fuel cycles. (Refer to Table 4-2 for the axial flux shapes used in this analysis.) The t

minimum LHRTM occurs early in life due to fuel densification, but quickly increases due to the offsetting effects of cladding creepdown, apellet swelling, and fuel relocation. (No credit is taken for fuel 4

relocation in LHRTM analyses).

3 i

i i

I l

I 4-8 Rev. 4 1

TABLE 4-1 FUEL MECHANICAL PERFORMANCE ASSESSMENT CRITERIA Analysis Catesory Linear Heat-Item No. Parameter Reviewedt Claddina Collapsea Claddina Strain Claddina Stress Pin Pressure Rate Capability ,

1 Pin Power History vs Burnup NA NA NA. Figure 4-1 Figure 4-1 2 Radial Assembly Power History i

vs Burnup Figure 4-2 NA NA NA NA 3 Clad O.D. Yes Yes ,

Yes Yes- Yes I 4 Clad I.D. Yes Yes Yes Yes Yes i 5 Clad Thickness Yes Yes Yes Yes Yes 6 Clad Initial Ovality Yes NA NA NA NA 7 Pellet Diameter Yes Yes Yes Yes' Yes 8 Pellet Density Yes Yes Yes Yes Yes  ;

9 Initial Prepressure Yes Yes Yes Yes Yes ,

~

f

. NOTES: 1. These criteria are the more significant items reviewed for a reload fuel cycle design, and do not i include minor assumptions that sre part of the bases.

2. The cladding collapse review actually is performed separately for each type of Mark B fuel design j (four sets of parametere exist, corresponding to four, separate fuel designs).

1 l

i l

! 4-9 Rev. 2

. TABLE 4-2 Axial. Flux Shapes Used for Thermal. Analysis (Reference, Figures 4-5, 4-6)

-Burnup Ranae Peak of Axial Co ine Shapes 0 -

15,100 1.28 15,100 - 35,000 1.22

> 35,000 1.16 NOTE: Standard reload desig:2 analyses always employ a 1.5 P/P axial flux shape for pin pressure and LHRTH analysis.

4-10 Rev. 2

0 I

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i 1 i 1 1

^a.t_ mw32 z i >P5 r

$- mms "

f FIGURE 4-2 RADIAL ASSEMBLY POWER VERSUS BURNUP FOR CREEP COLLAPSE ANAlySgS ASSESSMENTS (EXAMPLE ONLY) 1.6 -

1.5 -

z 1,4 -

iF-6 O

1.3 -

Z 3

e 1.2 -

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t 2 0.9 -

w

$ 0.s -

1 O

4 0.7 -

m 0.6 -

Q g g e a ' ' 1 I ' '

o 100 , a 1100 IM TIME (EFPD)

FIGURE 4-3 THERMAL ANALYSIS FLOW DIAGRAM COMPARE COMPARE CLADDING ER ENVELOPE W YES ENVELOPED? YES

-> D PmET O NO CHARACTER.

(F tG. 41)

ISTICS REVISE STANDARD INPUTS TO TACD(2)

CREATE A

NEW ENVELOPE 1r THAT INCLUDES  : PERFORM -

REPORT LIST, MATCH OUTLIERS REANALYSIS RESULTS

$ CYCLE _

w DESIGN PARAMETERS FUELTEMPERATURE ANALYSIS (LHRTM)

COMPARE COMPARE CLADDING WER EWELOPED? YES D? YES

-> AND PELLET ,,

Opy NO NO CHARACTER.

(FIG. 411 ISTICS REVISE STANDARD INPUTS TO TAC 0(2) 1r CREATE A NEW ENVELOPE 1r y THAT RERDRT

< INCLUDES - PERFORM RESULTS i OUTLIERS REANALYSIS FUEL PIN PRESSURE ANALYSIS

. i. _ ._ _ _

FIGURE 4 4 MECHANICAL ANALYSES FLOW DIAGRAM YES RERUN TACU(2)

CALCULATE PIN PRESSURES DETERMINE WORST ' '

CASE POWER HISTORY ENVELOPED? , ,NO, DETERMINE CREEP O , . CALCU' LATE AND CHECK CLADDING '  : COLLAPSE TIME  ;

CLAD TEMPS AND PELLET -CROV-CHARACTERISTICS DETERMINE FAST CREEP COLLAPSE ANALYSIS FLUX 1 MEV LIST MATCil COMPARE CLADDING

. DESIGN  : AND PELL ET ,  ;

• PARAMETERS CHARACTE RISTICS NO a t STRAIN ANALYSIS RECALCU-LATE CLAD STRAIN COMPARE WiTH GENERIC STRESS REPORT ASSESSMENT ENVELOPED? YES RESul.TS CRITERIA NO

< L STRESS ANALYSIS 1 r

,< RECALCU l

• LATE CLAD STRESS

0 0

0, 0

8 0

0 I 0, 0

7 0

0 I 0, 0 _

6 _

, i8 l l l 1 l I I 0

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0 2

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3 Figure 4-6 Fuel Linear Heat Rate to Melt Vs Burnup 26.0 -

25.0 -

i 24.0 -

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I 3

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I I i l i I i l

[

w 0 5000. 10000 15000 20000 25000 30000 35000 Burnup (MWD /MTU)

5. MANEUVERING ANALYSIS The purpose of a maneuvering analysis is to generate three dimensional power distributions and imbalances for a variety of rod positions, xenon distributions, and power levels.

The' maneuvering analysis can be divided into five discrete phases. The first is

-the fuel cycle depletion performed to establish a nominal rod index and fuel de-pletion history. The second is the generation of integral rod worth curves at several burnup steps. The third step is the power maneuver performed in the 1 nominal manner at BOC (4EFPD) and usually after a rod pull or at EOC. The fourth phase is to perform control rod and APSR scans at the most severe times of the power transient. The fifth step is to perform selected control rod and APSR ,

scans at the various depletion steps. Each of these phases involves the running

.of'aultiple EPRI-NODE cases and generation of three dimensional power distribu-tions, rod positions, and imbalance for each case.

This dataLis processed by the Node Utility Code (NUC) to calculate CFM, DNBR, and LOCA margins and to produce " fly speck" graphs to be used to set Technical Speci-fication (see Section 7.1) limits on rod position, offset versus power level, and 1 reactor protective system setpoints.

5.1 Fuel Cycle Depletion If the three dimensional EPRI-NODE model has not been previously depleted during the FFCD with rod positions which limit imbalance to within 1 5%, then the fuel cycle depletion is performed as the first step of the maneuvering analysis. The depletion is typically performed in steps of 0, 4, 12, 25, 50, 100, 150 ... EFPD adjusting rod positi6ns and. soluble boron where necessary to hold imbalance with-in 1 5% and the reactor near critical. The xenon, power, and exposure files for these cases are saved for.use in later analyses.

f L

l l

l 5-1 r.

5.2 Integral Rod Worth EPRI-NODE cases are run at- BOC (4EFPD) and after a rod pull or at EOC stepping the control rods into the core to calculate an integral rod worth curve for groups 7,-6, and'5 'with a 25% overlap. These cases input the exposure history calculated from the nominal depletion discussed in Section 5.1. The resulting integral rod worth curve is used in determining rod positions for maintaining criticality dur-Eing the power maneuver and to define rod insertion limits based on shutdown mar- 1 gin and ejected rod worth.

5.3 Power Maneuver 5.3.1 " Rods In" Design 1 The burnup and xenon distribution from the BOC (4EFPD) depletion case at 100% HFP forms the input starting point for the 100-30-110 power maneuver. First an EPRI-NODE case is run reducing power to 30% HFP and holding everything else constant.

The difference in reactivity between the 100% and 30% cases represents the power deficit. Using the integral rod worth curve generated in Section 5.2, the change in rod position necessary to compensate this power deficit is calculated.

The next step is to run a third case reducing the power level to 30% with the above estimated control rod positions. Subsequent cases may be necessary to correct the rod positions to maintain the problem near critical. APSR positions are also varied to maintain imbalance within +5% where possible.

The transient is performed using one hour time steps for updating the xenon con-centrations, rod positions, and power levels (soluble boron concentration is held constant). The power is held at 30% until peak xenon occurs. When peak xenon has occurred (approximately six hours) the power is raised back to 100% and the control rods are withdrawn enough to compensate for the power deficit and the increased xenon worth. The-transient is followed using one hour steps inserting the control rods as xenon burns out and maintaining imbalance by varying APSR position until xenon concentration approaches its new equilibrium concentration.

5-2

The power and xenon distributions, rod positions, and imbalance from this cominal 100-30-100 power maneuver are saved and used in' the next phar e of the analysis.

5.'3.2 " Rods Out"' Design For. a " Rods Cat" fuel cycle design, the power maneuver is simulated by inserting

- group 7 to 50% and allowing EPRI-NODE to calculate the reduced power level which results in criticality. This reduced power level is used in place of the 30% )

power level of the " Rods In" Design. The remainder of the calculations proceed 1 I

.s imilarly to those previously discussed for the " Rods In" Design.

5.4 Control Rod Scans Off the Power Maneuver

' Control rod scans are performed at the most severe times during the power 1 maneuver intentionally trying to produce large offsets and high power peaking to evaluate the effect of mispositioned control rods on power peaking. During these scans control groups 6 and 7 with appropriate overlap are inserted and/or. removed to produce large offse.ts and high power peaks. APSR scans are I also perforne'd to determine the offset and power peaking that results as these rods are moved.

The resulting power distributions and offset from these rod scans are used to

- evaluate CFM and DNBR margins which lead to the core safety limits on power and imbalance (see Section 7.2).

5.5 Control Rod' Scans Off Fuel Cycle Depletion Control rod scans on groups 6 and 7 with overlap are also performed from the fuel f

cycle depletion cases usually at 4 EFPD,100 EFPD, after any rod pull, and at EOC.

APSR scans are also performed at these times during the fuel cycle.

l The resulting power distributions, rod positions, and offset from these rod scans are also evaluated for CFM, DNBR, and LOCA margins and used in setting Technical Specification limits.

i 5-3

. . - . . - . . , . _ . . - - - --- - ---. n. - - - -. - . , , . , - . . - _ _ . - - - -

6. THERMAL HYDRAULIC DESIGN

.6.1 Introduction The thermal-hydraulic analyses establish the maximum permissible core power and power distribution for various operating conditions and the permissible combination of core outlet pressure and rc&ctor outlet temperature to ensure that-a minimum DNBR of 1.3 or greater can be maintained during the steady state overpower condition or during anticipated transients. The need to perform the thermal-hydraulic analyses in conjunction with a reload arises when there is a change in the fuel design, a change in the input assumptions of the original analysis, or a change in the regulatory criteria.

6.2 Thermal-Hydraulic Design Criterion The general criterion for thermal-hydraulic performance is that no core damage due to critical heat flux take place during steady state operations or during anticipated transiants.

The critical heat flux criterion is expressed as a departure from nucleate boiling ratio, or DNBR. Nucleate boiling refers to the heat transfer regime where 3 mall steam bubbles are forming on the clad surface and efficiently transferring heat (small temperature difference between clad and water). As fuel rod power is increased, the bubble generation increases to a point where the bubbles form an insulating blanket over the heating surface, causing a large rise in clad temperature. This point is the critical heat flux, or 0

departure from nucleate boiling. The DNB ratio is the ratio of this critical heat flux at a given point on a fuel rod to the actual heat flux at that same L point, or locatien.

LDNBR is calculated asing the Babcoch O VU cox BAW-2 correlation. This correla-tion, approved by NRC, has replaced the older and more conservative Westinghouse W-3 correlation used initially for the Oconee cores. The BAW-2 correlation, like the W-3, is na empirical correlation of hundreds of experimental data points covering the range of temperatures, pressures, mass velocities, and coolant 6-1 Rev. 1

qualities typical of PWRs. Validity limits on the BAW-2 which are observed during all thermal-hydraulic analyses follow:

Coolant Quality: -3% to +26% (analysis conservatively limited to +22%)

Mass Velocity : 0.95 x 10e to 4.0 x 108 Lba/hr-ft*

• Pressure  : 1750 to 2457 psia The minimum DNBR, during normal operation and anticipated transients is limit-ed to 1.30. A DNBR of 1.30 corresponds to a 95 percent probability at a 95 per-cent confidence level that DNB will not occur. Existing analyses conservatively 4

use a DNBR criterion of 1.4326 to accommodate rod bow (Refer to Section 6.10).

6.3 Analysis Methodology L

In order to show that the design criterion of 1.4326 minimum DNBR is met, analysis of core conditions with respect to coolant flow, core pressure, core inlet temperature, power level, and power distribution must be per-formed. The objective of the analysis is to define permissible values of these parameters such that the criterion is met. Assuring that the many pos-sible variations of power distribution in three dimensions meet the DNBR cri-terion, for example, requires a systematic analysis of possible power distri-butions and comparison with predicted conditions meeting the DNBR criterion.

A flow chart of this analysis methodology is provided in Figure 6.1.

The steady state thermal-hydraulic analysis covers the overrower (112% full power) condition as a reference point, then determines the allawable pressure-temperature operating limits, and finally determines power distribution limits i called generic DNBR curves that conservatively envelope the allowable core operating conditions. Subsequently, the transient thermal-hydraulic analysis of the two pump coastdown is completed, and results of this analysis are used to determine a flux / flow reactor trip setpoint. This setpoint ensures that the DNBR crite-ion is met upon loss of one or more primary coolant pumps.

l 6-2 Rev. 1 i

6.4 Core Inlet Conditions The first block of the Figure 6.1, Core Inlet Conditions, represents the hand calculations required to determine the core mass flow rate and the core inlet temperature (enthalpy) for each plant operating condition to be analyzed.

The reactor coolant pumps are constant volumetric flow pumps; thus the RCS nass flow rate varies with cold leg temperature. Further, the integrated control system reintains a constant average temperature over the power range

< of 15-100 percent, which requires that cold leg temperature decrease with

.reasing core power. These two factors when combined require that an 1 arative calculation be used to determine core inlet temperature and mass flow rate over the power range analyzed. Additional density changes (flow corrections) are made to. account for parametric variations in the core inlet temperature and outlet pressure as well as for the temperature and pressure errors which are applied during the analysis.

6.5 Reference Design DNBR Analysis This section represents block number 2 of Figure 6.1 and discusses the method used to determine the reference design DNBR, which is reported in Table 6-1,

-Thermal-Hydraulic Design Conditions, of each Reload Report.

A two stage analysis is used to determine this reference DNBR: 1) a core-wide analysis and 2) a hot assembly / hot channel analysis. These two stages are described in subsections 6.5.1 and 6.5.2, which follow.

l 6.5.1 Core-Wide Analysis i .

l The CHATA10 (Core Hydraulic And Thermal Analysis) program is used to determine the core-wide flow distribution. CHATA is a closed channel model (no energy _

j or mass interchange among assemblies) that varies flow to each assembly until 3

) each one has the same pressure drop and the sum of the assembly flows is equal .

to the input core flow.

l l

l i

6-3 Rev. 3 l-

Total core flow effective for heat transfer is input into CHATA, which models single fuel assenblies on an eighth-core symmetric basis to deter-mine the core flow distribution. The calculated result of particular impor-tance that is derived from this CHATA core flow distribution is the hot assem-

-bly flow, which is subsequently input into the hot assembly analysis described

.ia Subsection 6.5.2.

Primary inputs to this core-wide analysis are core flow effective for heat transfer, individual fuel assembly geometries, form loss ;oefficients, the

~

generic radial peaking distribution, the 1.5 design cosine axial flux shape, and core operating conditions.

Core flow rate is one of the most important inputs to the thermal-hydraulic analyses, and the possibility exists that this input flow can change depend-  ;

ing on measured flow. Reloads are now being designed based on 106.5% of the original design reactor coolant system flow rate of 88,000 gpa per pump. The 106.5% figure was selected based on the lowest reasured flow rate less measure-ment uncertainties. I Core flow is equal to total loop flow less the bypass flow, which is defined as that part of the flow that does not contact the active heat transfer surface area. These bypass flow paths are 1) core shroud, 2) core barrel annulus,  ;

j 3) control rod guide tubes and instrument tubes, and 4) all interfaces separat-ing the inlet and outlet regions. A typical value of the design bypass flow is 8.10%; however, this flow rate is dependent on the number of orifice rod and burnable poison rod assemblies. Removal of orifice rod assemblies and/or changes in the number of burnable poison rod assemblies and retainers cause both changes in the core bypass flow rate and in the core flow distribution.

Such changes will either be ref1'ected appropriately in the core-wide flow distribution or will be conservatively enveloped.

The basic assumption for the core inlet flow distribution. which is based on vessel model flow tests and Oconee 1 core pressure drop measurements, is that the isothermal flow distribution is relatively flat with a maximum deviation of 6-4 Rev. 1

, r ,- - - - _ w . . . - . . - - - - - - . . - --,% , , --. - , - - - . - - - - .- - - - -.--.-. , , - , , - -

less than 5% for 4 pump flow conditions. Therefore, the hot assembly is as-sumed to receive only 95% of the nominal assembly flow for this isothermal, four pump condition. ,

These isothermal flow maldistribution factors are considered during the core-wide CHATA analysis by the use of an additional form loss coefficient located at the entrance of the " hot" assembly. However, it is important~to note that the numerical value of this form loss coefficient is determined in an iso-thermal flow distribution analysis to be consistent with the fact that the vessel model flow test is an isothermal model. Subsequently, when the CHATA core-wide model is run st power conditions, a significantly larger hot assembly flow maldistribution results due to the radial peaking factor of the hot assembly. Further conservatisms imposed on the hot assembly during the core-wide analysis are minimum spacing effects on the flow area and on the form loss coefficients.

6.5.2 Hot Assembly / Hot Channel Analysis The second. step toward determining the reference design DNBR is the hot assembly /

hot channel analysis, which it.also represented by block number 2 of Figure 6.1.

The term " hot" asserbly refers to that fuel assembly with the highest radial peaking factor (actually the intersection of four 1/4 assemblies). The term hot channel refers to the subchannel with the highest single pin peaking factor.

This subchannel occurs within the hot assembly and is generally formed by the intersection of four fuel pins, although the hot channel could occur in a pin-control rod subchannel. (Hot channel factors are always included in all subchannel types within the hot bundle to permit this possibility.)

i-l The conservative hot assembly flow rate fro's Section 6.5.1 is input into the TEMP 14 (Thermal Energy Mixing Program) code for detailed analysis of the single 3

" hot" assembly. The Oconee hot ssembly pin by pin peaking distribution con-servatively models the intersection of the pin arrays of four 1/4 assemblies, with a relatively flat local peaking gradient, to conservatively minimize bene-ficial energy mixing effects. This generic pin-by pin power distribution in-l cludes the design radial peaking factor of 1.714 at the hot channel. All hot

! 6-5 Rev. 3 l

I l

channel factors are applied, so that the resulting DNBR calculated represents the' worst case (lowest) DNBR for the reactor core for the specified input condi-tions.

The assembly-wide modeling utilizes an open channel calculation; that is, it calculates energy interchange between channels at each calculational increment up the channel. This energy interchange reduces the enthalpy rise in the hot channel, thereby raising the minimum DNBR and permitting a higher allowable peak- 1 ing factor for the reactor core for conditions when DNBR is limiting. However, .

~the model does not include mass interchange between subchannels.

The outputs of primary importance from tae hot assembly / hot channel analysis are the hot channel minimum DNBR and ;0e hot channel flow rate. The hot '

channel minimum DNBR from the 112% overpower analysis is the reference design DNBR. In general, these outputs of mit1 mum DNBR and hot channel flow are used to establish the equivalent hot channel model of Section 6.6, which itself is used for parametric studies.

-6.5.3 Hot Channel Factors The following hot channel factors are utilized in the thermal-hydraulic analyses of Sections 6.5, 6.6, 6.7, and 6.9. Additional hot channel factors are included in the analyses of Section 6.8 and are described therein.

The flow area reduction factor, is 0.98 on the hot unit and the hot pin-rod cells, and is 0.97 on the instrument guide tube, wall and corner cells. .This factor, a statistical computation from measured as-built rod gaps, is applied to the channel flow area at each increment.

The hot channel factor on average pin power, is 1.0107. j

- This factor accounts for variations in average pin power caused by differences in the absolute number of grams of U-235 per rod. The loading tolerance on U-235 per fuel stack and variation on the powder lot mean enrichment are considered in determining the factor.

1 The hot channel factor due to manufacturing tolerances, is 1.014.

Variations in pellet density, pellet cross-sectional area, weight per unit length, local enrichment, and local outer clad diameter are all i accounted for in this factor.

6-6 Rev. 1

~6.6 Equivalent Two Channel Model Results of the hot channel analysis performed using the TEMP code, described in Subsection 6.5.2, are used to build a closed, dual channel model that is used for all subsequent parametric analyses. This two channel model contains an average channel and a hot channel and is modeled using the CHATA. code.

The hot channel model in CHATA contains all of the conservatisms used in TEMP and is forced'to duplicate the hot channel minimum DNBR calculated by TEMP.

This matching of the CHATA hot channel with that of TEMl' is accomplished by use of an engineering hot channel factor on enthalpy rise, FLah, that is applied during the CHATA analysis. This factor is varied until the CHATA hot channel minimus DNBR equals that in TEMP.

t In parametric analyses, the average channel represents a pseudo-core average channel and acts as the." driver" of the hot channel. Thus, an accurate, yet efficient representation of the TEMP hot channel is created for use in para-metric analyses, such as the pressure-temperature core protection safety

- limits and the generic DNBR curves, described in Subsections 6.7 and 6.8.

6.7' Determination of Pressure-Temperature Core Protection Safety Limits The curves shown in Figure 6.2, Core Protection Safety Limits, represent values of reactor outlet temperatura and core outlet pressure for which a minimum DNBR of 1.4326 (or 22% quality) is obtained for various pump combinations. The thermal-nydraulic analysis considers all conservatisms discussed in Section 6.5.

To determine the allowable range of pressure-temperature combinations for each pump combination, a series of DNBR calculations is done using the equivalent I two channel model for a ' range of core outlet pressures and reactor outlet tem-peratures. The results of these calculations are used to generate the locus

[

l of pressure-temperature points corresponding to the minimum DNBR of 1.4326.

1 i- For.the 3 pump and 2 pump cases, the minimum core flows as a fraction of 4-pump flow are standard, previously verified numbers; their corresponding power ,

l levels are calculated using the flux-flow ratio determined for each reload. ,

I I

6-7 Rev. 1 1

Section 7.3.1 discusses calculation of Pressure T ,erature Trip Setpoints.

J 6.8 Determination of Power Distribution Limits j 6.8.1' Summary

- Calculation of power power imbalance limits based on the DNBR criterion in-volves a synthesis of thermal-hydraulic analysis and the results of the maneu-vering analysis. Margin to DNBR is calculated in the maneuvering analysis for design power transients at various burnups. The method used to generate these DNBR margins efficiently is to precalculate a set of generic curves plotting the total peaking factor producing the minimum allowable DNBR (1.4326) for.the overpower condition associated with each pump combination. These i - curves of allowable total peaking are plotted versus distance up the channel for a range of axial peaking factors as shown in Figure 6-3. The fact that the curves are generic means that they can be generated once and used for all

, maneuvering analyses until fuel design or core flow rate changes impacting the thermal-hydraulic analyses are made. The following two sections. describe che thermal-hydraulic analyses involved in obtaining the generic DNBR curves.

Since the curves are plotted in terms of maximum allowable total peaking factors that envelope all core operating power distributions, their comparison to actual peaking during the maneuvering analysis becomes a relatively simple numeric exercise rather than a thermal-hydraulic analysis. Conversion of these calcu-F . lated DNBR margins into Power-Power Imbalance limits is described in Section 7.2.2.2.

l 6.8.2 Generic DNBR Curves  !

Sections 6.3, 6.4, 6.5, and 6.6 have described the methods used to arrive at a dual channel model which can be used for performing parametric studies.

This subsection will summarize how this dual channel model is used in deter-mining the generic DNBR; curves.

These parametric variations on the reference hot channel analysis are based on i the concept that for specified reactor core operating conditions - power level, l

i. 1 flow rate, temperature and pressure - all channels in the core have the same l 1

l 6-8 Rev. 1 l I

i

pressure drop regardless of variations in local peaking and axial power shape.

In other words, hot channel flow rate will be adjusted by the code to satisfy core-wide pressure drop as local conditions are varied. The axial power shapes input to'these parametric hot channel runs are smooth cosine curves whose 9' peak can be specified at various distances up the channel for each series of axial peaking factors. To obtain the maximum allowable peaking factor for each ,

' data point, power input to the channel is increased until the limiting DNBR of 1.4326 is reached. This process determines a maximum allowable total peak for a specified axial peak and its location.

After completion of these parametric analyses, two sets of generic DNBR curves

~

or Maximum Allowable Peaking (MAP) curves are determined. One set is used for DNB operational offset 1Luits, and the second set is used for RPS DNb offset limits. The generic DNBR curves used as operational limits are a conservative overley of 1) the generic DNBR curves used for RPS offset limits, and 2) another set of MAP curves which have the reference design DNBR as their basis. Both sets of limits consider the extremities of the P-T core protection envelope (619'F and 1800 psig) as potential core operating conditions. Thus both the operational DNB offset limits and the RPS DNB offset limits have considered the worst case temperature and pressure envelope permitted by the RPS.

The last step in the thermal-hydraulic analysis is to take actual power shapes that gave the lowest DNBRs during the maneuvering analysis and input these irregularly shaped axial curves into the hot channel code to verify conserva-tism of the corresponding cosine curves used to develop the generic DNBR curves.

A typical set of generic DNB curves is provided in Figure 6.3.

6.8.3 Hot Channel Factors l

The following additional hot channel factors on local heat flux are utilized in the thermal-hydraulic analyses for developing the generic DNBR curves:

f~

1.026 = penalty incurred to increase calculated axial powers since l flux depressions at the spacer grids are ignored.

l f

- 1.024 = the ratio of the total nuclear uncertainty of 1.075 to the i '

radial nuclear uncertainty of 1.05.

6-9 Rev. 1 l

'Thus, in determining the generic DNB curves, the normal value of Fq" is -

increased from 1.014 to 1.065.

6.9 Transient Analysis - Determination of the Flux - Flow Ratio During a loss of one or more reactor coolant pumps, the core is prevented from violating the 1.4326 minumum DNBR criterion by a reactor trip that is initiated by exceeding the allowable reactor power to reactor coolant flow ratio setpoint.

Loss of one or more reactor coolant (RC) pumps is also detected by the RC pump monitors. That is, in' dependently of the power to flow trip, loss of one RC pump will result.in an automatic reactor runback. Similarly, loss of two or more RC pumps from above 55% full power will cause a reactor trip.

The thermal-hydraulic analysis that is used to set the power to flow trip set-(point for.coastdown protection conservatively assumes the loss of two RC pumps.

The transient is analyzed using the RADAR 15 code to assure that the 1.4326 mini- 3 mum DNBR criterion is not violated at anytime during the loss of one or more RC pumps.

-The steady state thermal-hydraulic analysis provides the starting point for the transient analysis. The power to flow setpoint itself is derived from

.this analysis by varying the time of reactor trip following the loss of two RC pumps (that is by considering various trip setpoints) until the minimum ratio required to maintain the minimum DNBR of 1.4326 has been determined.

' Calculation of the actual (error corrected) power to flow setpoint used at the nuclear station is described in Section 7.3.2.

6.10 Application of the Rod Bow Penalty In existing thermal-hydraulic analyses, a very conservative DNBR penalty is included to account for rod bowing effects. This pc'alty (11.2%), however, has been reduced _by 1% because of the flow area (rod pitch) reduction factor already included in the thermal hydraulic analysis.

2 For some reloads, additional credit can be applied based on the fact that

. primary coolant flow can be proven to be higher than the 106.5% design flow.

6-10 Rev. 3 i I

I

l l

1 l

The resulting net penalty is applied directly to the final DNBR margins or  ;

by increasing the 1.3 DNBR criteria by the percent penalty, resulting in a DNBR criterion of 1.4326. 1 2

In future fuel cycle designs, this penalty will be revised to reflect the true effect of measured rod bowing on minimum DNBR (if any additional penalty J is required). References 12 and 13 document the methods to be used for deter-mining the true r,d bow penalty. Then, a determination will be made to I

either maintain the current margin which exists or to eliminate part or all of this margin.

l I

l l

!~

l l

{

6-11 Rev. 2 l

Fl!UIE 6.1 THERMAL HYDRAULIC ANALYSIS METHODOLOGY 1

CORE INLET CONDITIONS )

(REFER TO SECTION 6.4) l REFERENCE DESIGN DN8R ANALYSIS l

.COREWIDE ANALYSIS

. HOT ASSEMBLY / MOT CHANNEL (REF5R TO SECTION 6.5) l HOT CHANNEL FLOW RATE, REFERENCE DESIGN DNBR REFE ENCE DESIGN (REFER TO SECTION 6.6)

DNBR EOUlVALENT HOT AND AVERAGE CHANNEL MODELS PRESSURE-TEMPERATURE CORE POWER DISTRIBUTION LIMITS J CORE PROTECTION SAFETY LIMITS INPUT VARIOUS AXlAL FLUX SHAPES INPUT P-T COM8INATIONS (COMBINATIONS OF TOTAL PEAK, TO DETERMINE LIMITING AXlAL PEAK, AND AXtAL PEAK COMBINATIONS FOR CORE ,

LOCATION) TO DETERMINE MAX.

PROTECTION ALLOWABLE PEAKING.

(REFER TO SECTION 6.7) (REFER TO SECTION 6.8) 1 1f ALLOWABLE P-T COMBINATIONS ALLOWA8LE PO ER DISTRIBUTIONS TRANSIENT ANALYSIS OF

7. PUMP COASTDOWN (REFER TO SECTION 6.9) u FLUX / FLOW SETPOINT &

PARTIAL PUMP POWER LEVELS l

1 1

f

, 1 l

l 6-12 Rev.1

0 7

6 N N N O O O I I I T T T A A AR R R E 0 E E P

, 6 P P 6 O O OP P P M MM U U U P P P 0

, 5 4 3 2 6

0 4

6 E F R O U 0 E T . 3 R A 6 U R T E

P T

I A M R MI E EL 0 P TY ET

. 2 6

M E

RE T UF T SA E SS L E

R E .

0 T P IV 1 6 U ET O TC L AE E S

2.

6 SO TT S 0 E EY R . 0 V RD P 6 R

UA E O GETOR I T FSC C .

0 9

A 5 E R

s 0

, 8 5

0

_ 7 5

0 6

5 I m - - - - - - ~ - -

0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0

_ 5 4 3 2 1 0 8 7 2 1 2 2 2 2 2 1 f Ea$w s 8 = 0D .

!w F* ~

A

, FIGURE 63 gg . MAXIMUM ALLOWABLE PEAKING 2.8 -

u _

2.7 -

l 2.6 U- O C 1.70 l

2.s -

l la O - . .

y - - -

x 2.4 -

E

( 1.50 23 -

1.40 2.2 -

i 130 2.1 .

:  : g 2.0 -

33 i

1.9 -

0 0 0 0 0 - ppex=1.10 1.8 - 5 >

t, I f I t t 0 0.4 0.5 0.6 0.7 0.8 0.9 1.0

'7NAL DISTANCE CHANNEL X/L l 6-14 Rev.1

7. TECHNICAL SPECIFICATIONS REVIEW AND DEVELOPMENT 7.1 ~ Technical Specifications Review One'of the license' conditions applicable to the operation of a power reactor is that the reactor facility should be operated in accordance with the " Technical Specifications". . Technical Specifications are criteria for safe operation of

^

the reactor and are established from applicable design evaluations, safety analyses, and other considerations. Included in the Technical Specifications are safety limits, limiting system settings, limiting conditions for operation, surveillance requirements, identification of design features, and identification

'of administrative controls.

The. Technical Specifications on core safety limits, certain limiting safety.

system settings, and certain limiting conditions for operation are established on the basis of, among other things, the nuclear and thermal-hydraulic charac-teristics of the core and applicable accident analyses. Since the nuclear and

. thermal-hydraulic behavior of the core and accident analyses may be affected by the reload design, the Technical Specifications (and their bases), parti-

-cularly the sections pertaining to core safety limits, limiting safety system

- settings,' limiting conditions for operation, surveillance requirements, and reactor' design features are reviewed to confirm their continued validity for the reload cycle, and modifications are made as necessary to ensure safety of o -ation and/or to improve flexibility in operation. Technical Specifications ally affected by a typical reload design are (i) core safety limits, limiting safety system settings based on core safety limits and fuel

.gn limits, and (iii) limiting conditions for operation based on LOCA-power

.tribution limits and shutdown margin and ejected rod worth limits. The following subsections-describe the manner in which these Technical Specification l Ilimits'are developed.

7.2 Development of Core Safety Limits

! The core safety limits define limits on the values of pertinent core parameters y such that if core operation is within these limits, the integrity of the fuel cladding is maintained. Fuel cladding integrity can be assured (within per-l 7-1 i

A

missible tolerances) by maintaining the minimum DNBR in the core at or above tne design minimum value of 1.3 and by limiting the maximum linear heat rate in the core to less than or equal to the center fuel melt LHR limit. In order to achieve this condition, values of pertinent core parameters which correspond to a minimum DNBR of 1.3 and/or the center fuel melt LHR limit are calculated, and there values form the core safety limits. Core safety limits l l

are specified on core pressure-core outlet temperature combinations (P-T limits) and on-reactor power power imbalance combinations. In calculating these limits it is assumed that all other pertinent variables are at their design limits (maximum or minimum, as appropriate).

7.2.1 Determination of Core Safety P-T Limits The P-T limits are based entirely on the DNBR criterion, and they represent the values of core outlet pressure--vessel outlet temperature combinations for which a minimum DNBR of 1.3 is predicted when other pertinent parameters are at their respective design limits. The thermal-hydraulic analysis of Section 6.7 defines the values of core outlet pressure as a function of vessel outlet temperature for which a minimum DNBR of 1.3 is predicted for the maximum

' design conditions during 4 pump, 3 pump, and 2-pump modes of operation. (The design conditions during 4-pump operation consist of a reactor power of 112%

power, a combination of radial peak of 1.71 with an axial peak of 1.5, and a minimum reactor coolant flow of 374,880 gpm). The core safety limit is obtained by superimposing the P-T curves corresponding to 4-pump, 3-pump, and 2 pump modes of operation and by-drawing the enveloping curve as shown in Figure 7-1.

7.2.2 Detenaination of Core Safety Power-Power Imbalance Limits The core safety power-power imbalance limits define the values of reactor power as a function of axial imbalance such that for these values a minimum DNBR of 1.3 and/or a linear heat rate equal to the center fuel melt limit is predicted when other pertinent parameters (RCS flow, pressure and temperature, and hot channel _ factors) are at their design limits. These limits indirectly represent the limits on the DNBR criterion-limited power peaks and the center fuel melt criterion-limited power peaks. Since power peaking is not directly measurable 7-2

(

by the RPS, the DNBR criterion-limited power peaks and the center fuel melt criterion-limited power peaks are separately correlated to RPS measurable reactor power and power imbalance, and limits are then established on reactor '

power-power imbalance combinations to satisfy the DNBR and center fuel melt criteria. The power-power imbalance limits separately established for the DNBR and center fuel melt criteria are then superimposed, and the resulting -

most limiting power-power imbalance envelope forms the core safety limit.

7.2.2.1 Calculation of Power-Power Imbalance Limits for Center Fuel Melt Criterion

{'

The power-power imbalance limits based on the center fuel melt criterion are determined by a synthesis of the results of the fuel thermal analysis and the .

results of the maneuvering analysis.

The fuel thermal analysis (Section 4.6) establishes the maximum permissible linear heat rate in the core to prevent center fuel melting (center fuel melt linear heat rate limit). Using the center fuel melt linear heat rate limit (CFMLHR), the allowable total peaking factor is established by the relation:

MAPF = CFMLHR LHR X FOP -

where LHR is the average linear heat rate in the core (including densification effects) and FOP is the power level expressed as a fraction of rated power.

The maneuvering analysis (Section 5) establishes the maximum calculated total i f peaking factors for various core conditions (power levels, xenon conditions, control rod positions and burnups). These calculated maximum total peaking factors are increased by several conservative factors to obtain the worst c

case expected total peaking factor corresponding to each condition. The in-7 dividual conservative factors and their values are as follows:

1. Nuclear uncertainty factor = 1.075 .

+ .

7-3

.a

=

. 4_ .

-  ?'

.._f-'m -

2. Spacer grid effect factor = 1.026

3. Radial-local factor = varies with location of the assembly in the core (typical value is 1.10)

4. Engineering hot channel factor = 1.014

5. Densification power spike factor: varies with axial location of the peak in the core.

The nuclear uncertainty factor accounts for the uncertainty in the calculated i peak due to the limitations of the analytical models; the spacer grid effect factor accounts for the flux distortion caused by the spacer grids; and the radial-local factor is applied to account for the fact that the calculations are

.I performed using an assembly-by-assembly model rather than by using a pin-by pin model. The engineering hot channel factor accounts for the manufacturing tolerances of critical fuel rod design parameters (pellet enrichment, pellet density, pellet diameter, etc.). The densification power spike factor accounts for the local flux enhancement resulting from gaps in the fuel column induced by fuel densification. Although fuel rod bowing is considered to have the potential for enhancing the power peaks, no explicit allowance is provided for the rod bow power spike factor on the basis that the other conservatism factors (nuclear uncertainty factor and engineering het channel factor) are adequate to offset the effect of the rod bow power spike factor without an additional allowance.

l The worst case expected maximum total peaking factors calculated in this manner for different power levels are compared to the respective allowable total peaking factors, and the central fuel melt mar' gin for each condition can be determined.

The central fuel melt margin at a particular power level is given by: , ..

, . :, ,A l CFM Margin (%) =

{ (

allowable total peak - worst case expected maximum total peak ..? f x 100 allowable total peak AN ri y 0; '

MJDC :

j-lw' w.

7-4

• k

.a ; -

? .

Core conditions which correspond to non-negative margins are acceptable conditions, and core conditions which correspond to negative margins cannot be permitted. In order to preclude core :onditions with negative margins, limits should be estab-lished on acceptable values of power peaking conditions for each power level, and corresponding reactor trip setpoints should be established so as to trip the re-actor when conditions approach unacceptable values. Since power peaking cannot directly be measured by the RPS, power peaks are first correlated with the RPS-measurable axial offset for each power level. The outputs of the maneuvering calculations include the maximum total peaking factor in the core, its location '

and the corresponding core axial offset. In order to determine the axial offset limits that correspond to an acceptable center fuel melt margin for a particular power level, the center fuel melt margin for each calculated maximum total peak for that power level is plotted against the corresponding axial offset. These plots include the data for the entire cycle. For each plot two straight lines are drawn, one conservatively enveloping the d. a corresponding to positive of fsets and the other conservatively enveloping the data corresponding to negative offsets. The maxiuum allowable positive and negative offsets are found by extrapolating these straight lines to zero margin. Figure 7-2 illustrates the analysis for the 100% FP case.

In practice, detailed calculations are performed only for the 100% FP case, and the limits for other power levels are determined by conservatively extrapolating the 100% FP limits to other power levels by using the power feedback effect on peaking factors and by validating these limits by comparison with results ot a limited number of maneuvering calculations at these power levels.

Offset limits are typically established for power levels of 112% FP, 100% FP, 80% FP, and 50% FP.

7.2.2.2 Calculation of Power-Power Imbalance Limits for DNBR Criterion The power-power imbalance limits based on the DNBR criterion are determined by a synthesis of the results of the thermal-hydraulic analysis and the results of the maneuvering analysis.

The thermal-hydraulic analysis establishes the maximum allowable total peaking

=

factors as a function of core elevation for various axial flux shapes to prevent 7-5

, ~

4 1

violation'of the DNBR criterion (Section 6.8). The maneuvering analysis gener- 1 ates the power distribution in the core (including the maximum :;tal peaking  !

factor and the associated axial pe.sking factor for each fuel assembly in a l k-core representation and the core axial offset) for various design conditions and for various times in the cycle. For each power distribution, the calculated maximum total peaking factors of each of the assemblies is increased by a

, radial uncertainty factor of 1.05, a radial-local factor, and the resulting 1 adjusted peak is compared to the allowable peaking factor for that avial peak-ing factor and axial peak location. The DNBR reargin is then obtained as:

l DNBR margin (%) = (all wable total peek - adjusted maximum total peak) x 100 -

allowable total peak

' For each calculated power distribution, the DNBR margin is calculated for each assembly in the k-core, and then the mininium DNBR margin in the core for each power distribution.is determined.

In order to' determine the axial offset limits that correspond to the acceptable DNBR margin, the minimum DNBR margins are plotted for each calculated power distribution against the corresponding axial offset, and the maximum allowable positive and negative offset limits are determined in a manner similar to that used to establish the center ruel melt limited offset limits. In this case also, l offset limits are established typically for power levels of 112% FP,100% FP, 80% FP and 50% FP at full flow conditions.

7.2.2.3 Calculation of the Core Safety Limits on Power-Power Imbalance  !

I

- The core safety limits on power-power imbalance are the most limiting values 1 of the cent're fuel melt power imbalance limits and the DNBR power imbalance limit for each power level. To determine the core safety limits, first the

, limiting offsets at 112% FP, 100% FP, 80% FP, and 50% FP are determined by

, superimposing the DNBR and center fuel melt offset limits at each power level.

1 The following example illustrates the procedure:

l 7-6 l

i

l 1

(CFM Offset Limits (%) DNBR Offset Limiting Offset

' Power Level (% FP)

Limits (%) (%) .

-ve +ve -ve +ve -ve +ve 112 31 29 35 33 31 29 1 100 49 47 55 50 49 47 The limiting offsets at each power level are converted to imbalance limits using the relation:  ;

Power imbalance = axial offset x fraction of full power.  ;

l The resulting imbalance limits are plotted on a power power imbelance graph, as shown'in Figure 7-3. The following additional ateps are required to com-plete the procedure of determining the core safety limits on power power imbalance:

1. Draw a horizontal straight line corresponding to 112% FP.

2. From points where this line intersects the imbalance limit envelope, draw two straight lines, one on the positive imbalance and one on the negative imbalance side, that conservatively envelope the imbalance points.

These three straight lines define the power-power imbalance limits for 4-pur, operation.

l I The power-power imbalance limits for 3 pump and 2-pump modes of operation can be determined by reducing the thermal power associated with each break point of the 4 pump curve to the values of the maximum allowable core thermal power l

for 3 pump and 2 pump modes and by drawing straight lines parallel to the 4-l pump envelope through the points defined by the 3-pump and 2 pump thermal power and the 4-pump imbalance limits. The maximum thermal power for the 3-l i

7-7

I l

pump mode is obtained by multiplying the 3 pump flow (74.7% of the full flow) by the flux-flow trip setpoint and adding the allowance for calibration and instrumentation error for power measurement (6.5%) to the product. The maxi-num allowabie core thermal power for the 2 pump case (one-pump for each loop) is determined by a similar manner.

7.3 -Development of Limiting Safety Settings The reactor protection system contains several trip functions designed to pre-vent the process variables from exceeding the safety limits, te ensure that the fuel design limits (minimum DNBR and center fuel melt LHR limit) are not exceeded l l

during conditions of normal operation and anticipated transients, and to enable '

reactor shutdown during accident conditions. These trip functions, their intended purpose, and their setpoints are shown in Table 7-1. The trip setpoints are established by reducing the safety limits or other design analysis limits I by appropriate error adjustment factors, which account for any uncertainty in the measurement of that variable and the calibration and instrumentation errors.

It re aral, the trip setpointe requiring modification for a reload cycle are the P-T trip setpoints and the power-flow-imbalance trip setpoints as a result of a 1 change.in the core safety limits and/or a change in the flux / flow trip setpoints.

7.3.1 Determination of RPS P-T Trip Setpoints The P-T trip _ function defines values of RCS pressure as a function of RC out-let temperature at which the RPS should trip and provides protection of the P-T core safety limits.

l The P-T trip setpoints are derived by errot-adjusting the P-T core safety limits and by considering the high RCS pressure, low RCS pressure, and high RC outlet temperature trip setpoints. Error adjustment is performed on the RCS pressure

-(to account for the difference in pressure between the core outlet and the point of measurement and to account for the error in the measuremey, of pressure by the RPS) and the RC outlet temperature (to account for the error l in temperature measurement by the RPS). The P-T trip setpoints are to be modi-1 l

7-8 l 1

l

.-,1

fied whenever the P-T core safety limits are changed, P-T error adjustment factors are changed, or the high RC outlet temperature trip setpoints are changed, or the low RCS pressure trip setpoint is changed.

c In' order to determine the P-T trip setpoints; first the locus of pressure-temperature points constrained by the high RCS pressure trip setpoint (2300 psig), 4 the high RCS temperatur~e trip setpoint (619*F), and the low RCS pressure trip setpoint (1800 psig) are identified on the Core Safety P-T Limit curve, as shown in Figure.7-4. Referring to Figure 7-6, the straight lines AB, BC, and DE respectively represent.tne locus of P-T points constrained by the high RCS pressure trip, the high RCS temperature trip, and the low RCS pressure trip setpoints. Next, the pressure-temperature points C and D are adjusted for the difference between the core pressure and the RCS pressure at the measurement

. location and for the errors in the temperature and pressure measurements by the RPS. Referring to Figure 7-4, C' and D' are the error adjusted points, and the straight line C'D' joining these points defines the locus of RPS P-T trip setpoints.

7.3.2 Determination of RPS Power-Flow-Imbalance Trip Setpoints The power-flow-imbalance trip setpoints define the values of reactor power at which RPS trip should occur whenever the combinations of power, flow, and their uncertainties produce limiting values of power and flow which result in the design minimum DNBR during a flow transient and whenever the ecmbination

'of power, imbalance, and their uncertainties correspond to the core safety limits on power-imbalance. This trip function is established by considering maximum allowable power-to-flow ratio and by considering the maximum allowable values of power as a function of imbalance. The maximum allowable power-to-flow ratio is constrained by the requirement that the minimum DNBR, in the event of a limiting flow transient, is equal to or greater than the design limit of 1.3.

Thus the power-flow-imbalance trip setpoints ensure core protection during transients involving a flow reduction (by the power-to-flow trip portion of the trip function) and during conditions involving adverse power distributions (by ,

the power-imbalance trip portions of the trip function). .

7-9

. = . -

1 In order to determine the power-flow-imbalance trip setpoints, first the maximum allowable power-to-flow ratio is to be obtained. The maximum allow-

' able power-to-flow ratio (also called the flux / flow trip setpoint) is obta'ined by reducing the calculated flux / flow retio (Section 6.9) by an error adjustment 4

_ factor, which takes into account the noise in the RPS flow signal and other electronic errors in the RPS flow instrumentation. Next, the core safety power-imbalance limits are error-adjusted both on the power level limit and tne in- 1 j

balance limit. The error adjustment factor for power level is 6.5% FP, which includes 4% FP allowance for the neutron flux error (uncertainty in correlating the RPS measured neutron flux to reactor power), 2% FP allowance for the calori-metric error, and % r? allowance for any setpoint error. The error adjustment factor for imbalance accounts for the uncertainty in the measurement of axial i imbalance by the out-of-core detector system, and it is a function of the in-i balance 1Dait and the power level. To establish the RPS power-flow-imbalance  !

trip setpoints,'the error adjusted power and imbalance are plotted on a figure {

with imbalance as the horizontal axis and power as the vertical axis. The

. envelope obtained by the straight lines passing through pairs of these points and the horizontal straight line drawn passing through the point representing 112% power for the 4 pump case or the maximum power allowed by the flux / flow I trip setpoint, as illustrated in Figure 7-5.

7.4 Development of Limitina Conditions for Operation The limiting conditions of operation generally requiring modification in con-junction with a reload cycle are the LOCA-limited power distribution limits, shutdown margin-limited control rod insertion limits, and the ejected rod worth , limited control rod insertion limits.

l The LOCA-limited power distribution limits are limits on pertinent core para- l meters (such as control rod positions, axial imbalance, quadrant power tilt, and xenon conditions which influence the power distribution in the core) such that the power distributions in the core during normal operation are within the values assumed in the safety analysis of the loss of coolant accident. l The shutdown margin-limited control rod insertion limits are limits on the maximum allowable cont.ol rod insertions satisfying the shutdown margin 7-10

criterion, and the ejected rod worth-limited control rod insertion limits are limits on the maximum allowable control rod insertions satisfying the ejected _ rod worth criterion.

7.4.1 Determination of LOCA-Limited Power Distribution Limits

.The ECCS analysis establishes acceptable values of the linear heat rate in the core ruch that the performance of the Emergency Core Cooling System conforms to the requirements of the Final Acceptance Criteria. The values of the allowable linear heat rates established by the currently applicable ECCS analysis for Oconee class reactors are 15.5 kw/ft. at the 2 ft. core elevation, 16.6 kw/ft. at the 4 ft. elevation, 18 kw/ft. at the 6 ft. elevation, 17 kw/ft.

at the 8 ft. elevation, and 16 kw/ft. at the 10 ft. elevation. The maximum operating linear heat rates at the designated core elevations should be main-tained at or below the allowable values. The maximum operating linear heat rate is a function of the power level and the maximum operating peaking factor.

Thus, for a given power level the maximum operating linear heat rate varies E

with the maximum operating peaking factor. Therefore, for a given power level the maximum operating linear heat rates can be maintained within the allowable linear heat rates by maintaining the maximum operating power peaks at the designated axial locations within the allowable peaking factor. The allowable

. peaking fa.ctor at axial location z for the power level F0P is given by:

APF (F0P,z) = (*)

L LHR x FOP, l

[

where APF (FOP,z) is the allowable peaking factor at elevation z for power levels equal to or less than FOP, ALHR (z) is the allowable linear heat rate at axial

! location z, and LHR is the densified average linear heat rate at 100% FP.

l l

t The power peaking factor in the core changes with fuel burnup, axial imbalance, full length control rod position, and part length control rod position. In addition, the peaking factor is influenced by the existence of any quadrant power tilt and non-equilibrium xenon conditions. Therefore, allowable ranges of

these core operation parameters would have to be established in order that the f-maximum operating peaking factors at the designated axial' locations be within 7-11 i

~

the allowable values. Although the fuel densification phenomenon has the po-tential for enhancing power peaks, no explicit allowance is provided for power l

-spikes associated with this phenomenon in the LOCA power distribution limits on the basis that the densification power spikes do not enhance the local heat flux.

The'effect of a positive quadrant power tilt on the maximum peaking f.=ctor has been conservatively established to be an increase of 1.5% in peaking factor per percent positive quadrant tilt. The current Technical Specifications permit re-actor operation with a positive quadrant tilt of 5.00%, which amounts to a pos-

-sible 7.5% increase in peaking factor. Therefore, the allowable peaking factor would have to be reduced by 7.5% to account for the permitted quadrant tilt con-dition.

The effect of non-equilibrium xenon conditions on peaking factors is quantified by the analysis of the power peaking factors occurring during various power man-euvers. Based on the results of this analysis, it has been determined that for

" rods in" operation the effect of non-equilibrium xenon on power peaking can be accounted for by reducing the maximum allowable peaking factor by 8% for power levels equal or greater than 90% full power. For " rods out" operation the non-equilibrium xenon peaking factor is 5%.

The allowable peaking factor for " rods out" operation after accounting for quad- -

~

rant tilt and xenon becomes:

1 APF (F0P,z) = ()

LHRx F0P x 1.075 x 1.05 The remaining core parameters which influence the maximum operating power peaks are the full length control rod position, part length control rod position, I

axial imbalance, and core burnup. The permissible values of these quantities are h

to be determined such that resulting power peaks, after accounting for any uncer-tainties, would be within the maximum allowable power peaks.

I

( The maneuvering analysis establishes the relationship of operating peaking fac-

-tors at various axial locations with the core imbalance and control rod posi-

'tions. The maneuvering analysis calculations include part length cont I rod scans inducing a range of values of core axial offset for different full 7-12 g - -- - -

length control rod positions. The calculations are performed for various power levels and for the full range of core burnups. The calculations yield the values of the maximum peaking factor at the different axial planes corres-ponding to various full-length control rod positions, various axial offsets, and for different part length rod positions, and these calculations also yield the variations of the maximum peaking factor nith axial offset. The calculated maximum peaks at each axial plane are increased by the nuclear uncertainty factor (1.075), the spacer grid effect factor (1.026), the radial-local factor (its value varies with the radial lociticg of the assembly containing the maximum peak), the power level uncertainty factor (1.02) and the engineering hot channel factor (1.014) to obtain the worst case operating peaking factors.

To determine the allowable values of full-length and part-length control rod positions and the axial offsets, first an operating range for the full-length control rod position is chosen and then the ranges of axial offsets and part-length control rod positions for which the worst case operating peaking factors at the designated axial planes are less than or equal to their respective allowable values are determined. If the resulting ranges of axial offset and part-length control rod position are acceptable from the standpoint of operational flexibility, the assumed full-length control rod position ranges and the calculated range of axial offset and part-length control rod position are taken as the operating limit for the full-length control rod positions, axial offsets and part-length control rod positions. If, however, the resulting ranges of axial offsets and part-length control positions are unacceptable from the standpoint of operational flexibility, a more restrictive full-length control rod position bank is selected and the corresponding axial offset and part-length control rod position limits established.

Since the core peaking factors ds not remain constant throughout the entire fuel cycle, the operating limits on control rod positions and axial offsets should be based on the composite results of calculations for representative times in the cycle. In order to provide maximum operating flexibility, the operating limits on control rod positions and axial offsets are established for different cycle burnup intervals (e.g., BOC - 100EFPD, 100 EFPD - 250 EFPD and 250 EFPD-EOC).

The operating limits applicable to each burnup interval are generated on the 7-13

J - -

' basis of.the results of maneuvering calculations corresponding to the beginning and end of each burnup interval'. (For each burnup interval, the control rod grouping and the nominal position of the regulating control rod groups are the same).

Calculations of axial offset limits, part length control rod position limits, and full length control rod position limits are performed for various power levels (typically for 102% FP, 92% FP, 80% FP, and 50% FP). The offset limits 1 ar each power level are converted to imbalance limits by multiplying the offset

' limits by the applicable power fraction. Typical operating limits established in this manner are shown in Figures 7 7-8.

7.4.2- Determination of Control Rod position Limits Based on Shutdown Margin Criterion The criterion on shutdown margin is that a minimum of 1% Ak/k shutdown margin l should be available at all times. The shutdown margin decreases with increasing power and also with increasing inserted rod worth. Therefore, associated with each power level, there is a maximum allowable full length control rod insertion l limit which corresponds to a minimum shutdown margin of 1% Ak/k. Shutdown margin limited rod insertion 2imits are determined by evaluating the shutdown margins at different power levels (typically at 100% FP, 507. FP, and 15% FP) )

and by using the integral group worth results. Since shutdown margins change with' cycle burnup, shutdown margin limited rod insertion limits are calculated for different burnup intervals of the fuel cycle. The rhutdown margin limited rod insertion limits are identified in Figure 7-6.

7,4.3 Determination of Control Rod Position Limits Based on Ejected Rod Worth Criterion The criterion on the ejected rod worth is that its value shall not exceed 1%

-Ak/k at hot zero power (HZP) conditions and 0.65% Ak/k at hot full power (HFP) conditions. The limits at intermediate power levels are assumed to vary linearly with power--that is, the ejected rod worth at 80% FP is 0.72, the limit at 20% FP is 0.93, etc. The ejected rod worth is a function of, among other things, the inserted control rod group worth and the cycle burnup 7-14

(through changes in power distribution). For a fixed burnup the ejected rod 1

worth changes with control rod insertions; therefore limits on the allowable control rod insertion should be placed at various power levels so that the ejected rod worth criterion is satisfied. In order to determine the ejected rod worth limited control rod position limits, the ejected rod worths are cal-culated corresponding to the most limiting of the shutdown margin and LOCA-limited full length rou insertion limits for different power levels. The cal-culated ejected rod worths are increased by 15% and compared to the allowable values. If the adjusted calculated ejected rod worths are within the allowable values, no further calculations are needed; otherwise, the control rod insertion limit is changed to the value that corresponds to acceptable ejected rod worths.

When the ECCS-limited and ejected rod worth limited rc 3 insertion limits are more limiting than the shutdown margin limited insertion limits, the ECCS and ejected rod worth limited rod insertion limits are combined by superposition into a single rod insertion limit.

7-15

i I Tebh ?-1

{ Reactor Protection Systee Trip Functions Reactor Trip Nealtered Foremeter Trip Setroint During Purpose of 1 rip '

4-Pump Operation

1. Overpower Neutree flus 105.51 FP To provide core protection during transients <

trip involving uncontrollei power increase.

1

2. Power-flow- Neutros flex, AC f t:;. Flux / Flow e 1.08 To prodoe cora protection during tressie-nts I

imbalesce and power ishalance .imvelring e flow reduction and during core trip coed'.tions involving excessive power peaking .

3. RCS pressure- RCS pressure end RC Function of RC outlet- To yrovide core protecties during transients temperature outlet temperet4re temperature le<olving a reduction in pressure or a 1

trip reduction in core heat removal and to ensure reactor shut down during e IACA.

J

4. Low RCS RCS pressure 1800 pois To provide core protection during

]

g pressure treestents involving a pressure

triy reduction I

5. RC Fump Neutroe flus med pump Loss of two pumps To provide core protection during

);

Nonitor trip costect monitor voltage above 551 FP loss of RC pumps i

6. Righ RCS R'1 pressure 2300 pelg To provide protection of RCS pressure 4 f prtseere boundary from excessive pressures trip

, 7. Righ ?tCS RC outlet temp. 619'r To pri ont excessive temperature in the

, touperature RCS trip I

j S. Eigh RC RS pressure 4 psig To ensure reactor shutdown during a LOCA pressure and SLR inside containment.

trip .

0 3 7 6

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Figure 7-2 Margin To Center Fuel Melt LHR Versus Core Offset 60, ,,

X X X' ~X X y XX X- K X X X X X

X x

x x x xtxx X X x X X X X X X - .

X XX X X- 10 .

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Offset Limits j "

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-60 -40 -20 20 40 Core Offset, % )

7-18 l I

f Tigure 7-3 Core Protection Safety Limits Thermal P0wer Level. 5 UNACCEPTAsLE OPERATION -

1 120

(-31.112) (2s.sl2)

ACCEPTAsLE e Puwp OPERATION

(-se.ioo) - 100 (s7.800) 2

(-31.sr.2) (2s.s7 2)

~

- 80

(.,,,7$) ACCEPTAsLE (s7.75) 3As PUMP OPERATION

(-31.ss.s) 3 60 (2s.ss.s)

ACCEPTAsLE

(-us.s7) hP!RffGN

- 40 i

l CURVE RC FLOW (GPM) i 37s.sse

- 20 2 2:0.03s i

3 is3.sso

-60 ,

-40 -20 0 20 40 66 Reactor P0wer imbalance. 5 .

l l

l i

i~

l l 7-19 Rev. 1

Figura 7-4 Determinatien of RPS P-T Trip Setpoints 2400 --

A B 2200 - -

C' C RPS Trip Setpoints N

2 *

- 2000 - -

O g- .

n.

h Core Safety Limit D'

-O 1800 E <D o

t

=

1600 l l l l 560 580 600 620 640 Reactor Outlet Temperature, F 7-20 Rev. 4 i _ .

l Figura 7--5 Protectiva System Maximum Allowzble ~

Setpoints .

Thermal P0wer Level, 5

. 120 U N ACCE PT A st.E OPERATION

(-Is,10s) . .110 (is, 0s) i I

' .100 I N, = 0.ss0 l "2 = .72s s PuwP I I (as,ss.5)

(-s7.ss.s) I . 90 l OPERATION I I

(-is,s0.s)  ! - (es s0.s) 8 i . . so i l i I I (35,66.I)

(-37,6s.1) 3&s PUMP I

l OPERATION l .!. 60 l

3(15,52.9)

(-is 52.s) !

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i I 8 I i - - 40 g 2,3ss PuwP 1 (35.as.s)

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OPERATION -- 30 I I

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-20 20 40 .

-40 g Reactor P0wer imbalance, 5 1

^

7-21 Rev. 1

i l

~ Figure '7-6 Rod Position Limits for Four-Pump Operation. l (0 to 200 10 EFPD) '

I POWER LEVEL l

110

{CUTOFFe I'

\_I I

~

(180.102) (274.102)

~ ~

OPERATION IN THl3 REGION IS NOT ALLOWE0

, (270.92) 80 -

(284.4,80) l RESTRICTE0 g 70 -

REGION SHUTDOWN ,

'3 IRARGIN 2 60 -

LIMIT PERNIS$18LE I OPERATING

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4 Figure 7-7 Power Imbalance Limits, - ~.

(0 to 200 10 EFPD)

Power, 5 of 2588 N t 110

_ (20.4,102)

(-24.4.102) _

-100

(-24.8,92) _ , go a i (13.4,92)

(-33.8,80) i - 80

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- - 20

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- - 10 g l l l l 3 g g g j

-50, -40 -30 10 O 10 20 30 40 50 Axial Power Imbalance, 5 l

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7-23 Rev. 1

Figure 7-8 APSR Position Limits, (From 0 to 200 t 10 EFPD)

(8.0.102) (32,102)

RESTRICTED

.(6.0,92) 4 (32,92) 80 4 40.0,80) (38.6,80) 60 , _

E 3 '

E (100.'50)

. 40 -

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= PERMISSIBLE OPERATING E REC 10h 20 -

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O i , ,

0 20 40 60 80 100 Bank 8 Position, 5 Withdrawn 7-24 Rev. 1

8. ACCIDENT ANALYSIS REVIEW 8.1 Introduction A major aspect of the safei.y consideration of a reactor is the safety analysis of postulated accidents. These safety analyses enable one to confirm that the reactor system is designed to mitigate such events and that the resulting con-sequences of such events are acceptable. The most important considerations af-fecting the calculated consequences of the various postulated accidents are (a) the values of plant parameters assumed in the analysis, (b) the performance characteristics of the mitigating systems assumed in the analysis, and (c) the analytical models used. In general, the accident analyses documented in the E are based on values of plant FSAR( } and Fuel Densification Reports parameters that correspond to the mos,t adverre conditions expected to exist throughout the life of the plant, are based on conservative performance char-acteristics of the mitigating systems, and were performed utilizing generally accepted analytical methodology. Therefo e, t.he reference safety analyses (FSAR and Fuel Densification Analysis Reports) are intended to be valid for the entire life of the plant.

The primary goal of safety analysis during the reload design process is to ensure the continued safe operation of the facility with the refueled core.

The reference safety analyses and facility Technical Specifications establish the bases and condition.s for safe operation of the initial core. An equivalent level of safety for the refueled core is established when it is determined that the reload design satisfies the original bases and conditions. In particular, the accident analyses contained in the licensing basis safety analyses ramain valid if a reload design predicts steady-state and transient parameters that lie within the ranges of the values assumed in the original analyses. Thus, reload safety analysis may consist of verifying that the core physics, fuel performance, thermal-hydraulic, and mechanical design parameters for the reload design are bounded by the licensing basis analysis values.

8-1 Rev. 1

8.2 Overview of Accident Analysis Review The role of accident analysis review in typical Oconee reload design consists of a systematic review of the reference analysis of all postulated accidents.

In this review each accident is examined by comparing the values of important plant parameters and RPS trip functions and trip setpoints assumed in the re-ference accident analysis to the corresponding values predicted for the fuel '

cycle under consideration. The safety parameters of interest for the reload cycle are obtai.aed from appropriate nuclear design, thermal-hydraulic design, and fuel performance. analyses. If the safety analysis review confirms that all pertinent plant parameters and RPS trip functions and trip setpoints for the reload cycle are conservative with respect to their values assumed in the acci-

-dent analyses, it is concluded that the reference accident analyses continue to be valid for.the fuel cycle, and therefore in these situations no reanalyses of l

accidents are performed. If, however, one or more plant parameters or RPS trip functions or trip setpoints assumed in the reference accident analyses are found to be non-conservative for the fuel cycle, a reanalysis of affected accidents is performed. This process is shown schematically in Figure 8-1.

The safety parameters of interest for the reload cycle are obtained from appro-priate nuclear design, thermal-hydraulic design, and fuel performance analyses.

Table 8-1 presents a list of the key safety parameters that are reviewed for each reload design. The table indicates the conservative direction that each parameter value'should take relative to the reference analysis value.

In addition to the safety parameters addressed above, the reference analyses also incorporate the RPS trip functions and setpoints. (The role of the func-l tions and the determination of the setpoint values is discussed in Chapter 7 i of this report.) If a particular reload design results in revised setpoint values, a review of the reference accident analyses is performed and the l effect of the revision on the analyses is evaluated.

-In some cases, the reference analysis calculations explicitly include the  ;

various conservative engineering factors, densification and rod bow factors, and fuel pin design parameters. If a reload design results in changes in these l vslues, their impact on the reference analyses is evaluated.

8-2 Rev. 1 l

. , - - - . _ . ~ , - . _ - , . _ . _ _ _ _ _ _ _ . _ . _ _ . - - , _ _,_

8.3 Discussion of Individual Accidents A discussion of each of the accidents addressed in the reference analyses follows. For each event, a brief description of the accident is followed by a listing and discussion of the key safaty parameters associated with the accident.

8.3.1 Uncompensated Operating Reactivity Changes 8.3.1.1 Accident Description During the normal operation of the reactor, the overall reactivity of the core changes because of fuel depletion and variations in fission product poison concentrations. These reactivity changes, if not compensated for, could produce an increase or decrease in reactor power (depending on the direction of the reactivity change) and consequently change the fuel and moderator temperatures. Ultimately, core operating limits could be exceeded.

Normal functioning of the Integrated Control System would compensate for the reactivity changes. In the absence of automatic or manual compensatory re-sponses, the reactor coolant system average temperature will change to com-pensate for the reactivity disturbances.

The reference analyses demonstrate that the reactor protective system prevents safety limits from being exceeded.

8.3.1.2 Key Safety Parameters The reference analyses are based on the following parameter values:

Doppler Coefficient, Ak/k/*F -1.17 x 10 5 Moderator Temperature Coefficient, Ak/k/*F +0.5 x 10 ~4 These values are representative of beginning of core life for the first cycle and provide least negative bounds for the expected range of values. That is, mire negative (less positive) valat.:., would lessen the severity of the accident by emplifying the compensating moaerator an.d Derpler feedback effects.

6-3 Rev. 1

8.3.2 Start-Up Accident 8.3.2.1 Accident Description l l

During reactor startup, an uncontrolled positive reactivity insertion by mal-operation of control rods could result in a nuclear excursion. In addition to the Reactor Protection System trip functions, several design features have been

~

utilized to minimize the possibility of inadvertent rod withdr'awal. In the ab-sence of all other protection actions, the excursion is terminated by the neg-ative Doppler coefficient.

The core protection criteria for this accident specify that the reactor thermal power shall not exceed 112% FP and that the RCS pressure not exceed code allow-able limits.

The reference analyses demonstrate that the reactor is completely protected against any startup accident involving the withdrawal of any or all control rods, since in no case does the thermal power exceed 112% and peak pressure never exceeds code allowable limits.

~8.3.2.2 Key Safety Parameters The reference analyses are based on the following parameter values:

Doppler Coefficient, Ak/k/*F -1.17 x 10~5 Moderator Temperature Coefficient, Ak/k/*F +0.5 x 104 Total Rod Worth, Ak/k 10.0 The reactivity coefficient values are representative of beginning of core life for the first core nd provide a least negative bound for the expected range of values. That is, more negative (less positive) values would lessen the severity of the accident by amplifying the compenssting moderator and

-Doppler feedback effects. The total rod worth value is a maximum worth that provides for the largest positive reactivity insertion.

l 8-4 Rev. 1

A 8.3.3 Rod Withdrawal Accident at Rated Power Operation 8.3.3.1. Accident Description A rod withdrawal accident presupposes an operator error or equipment failure which results-in accidental withdrawal of a control rod group while the reactor is at rated power. As a result, the power level increases, the coolant and fuel rod temperatures increase, and core damage would eventually occur if the withdrawal were not terminated by operator or protection system action.

The reference analysis of this accident utilizes Reactor Protection System action to mitigate the effects of the rod withdrawal and demonstrates that thermal power and system pressure remain below acceptable limits. These results satisfy the core protection criteria for this accident.

8.3.3.2 Key Safety Parameters The reference analyses are based on the following parameter values:

Doppler Coefficient, Ak/k/*F -1.17 x 10 5 Moderator Temperature Coefficient, Ak/k/*F +0.5 x 10 ~4 Total Rod Worth, % Ak/k 10.0 The reactivity coefficient values are representative of beginning of core li'e for the first cycle and provide least negative bounds for the expected range of values. That is, more negative (less positive) values would lessen the severity of the accident by amplifying the compensating moderator and Doppler feedback effects. The total rod worth value is a maximum worth that provides for the largest positive reactivity insertion.

1 8-5 Rev. 1

-8.3.4 Moderator Dilution Accident 8.3.4.1 Accident Description Moderator dilution, a periodic operational procedure, occurs when the soluble boron concentration of the coolant make-up flow is less than the average con-centration of the coolant in the primary system. An uncontrolled moderator dilution accident occurs when the process continues for long periods of time at excessive make-up flow rates. The postive reactivity insertion caused by the decreasing soluble boron concentration would cause an increase in reactor power and hence increased coolant and fuel rod temperatures.

J. The_ automatic dilution process incorporates several design interlocks and alarms to prevent improper operation. However, if a dilution accident were to occur, the Reactor Protection System would function to safely mitigate the event.

'The criteria for reactor protection for this accident are:

1. Reactor thermal power shall be less than 112% FP.

l

2. RCS pressure shall be less than the code allowable limits.

3. The reactor minimum shutdown margin of 1% Ak/k suberitical shall be maintained.

The reference analyses evaluate plant responses to dilution rates ranging from -

70 gym to 500 gpm of unborated make-up water. In all caces, thermal power and system pressure remain below the specified limits and the shutdown margin is maintained. . Additional analyses demonstrate complete protection during refueling operations.

l 8.3.4.2 Key Safety Parameters l l

Doppler Coefficient, Ak/k/*F -1.17 x 10~5 l Moderator Temperature Coefficient, Ak/k/*F +0.94 x 10~4 Boron Worth, pps/% Ak/k 75 l

, 8-6 Rev. I

i The reactivity coefficient values are representative of beginning of core life conditions and provide least negative bounds on the range of expected values. More negative (less positive) values would lessen the severity of l the accident by amplifying the compensating moderator and Doppler feedback l effects. The boron worth value is a conservatively low value (high reac- I tivity worth per ppa)-that produces a high positive reactivity insertion rate. l 8.3.5 Cold Water Accident 8.3.5.1 Accident Description i - A cold ~ water accident involves the introduction of a slug of coolant water 1

into the reactor core with a temperature lower than that of the coolant in the core. Or, a cold water accident may involve a sudden increase in reactor coolant flow rate (idle pump startup) which would reduce the average coolant temperature in the core. In tne presence of negative reactivity coefficients,

a reduction in coolant and fuel temperatures would yield a positive reactivity 2 insertion and thus increase the power level.

The power increase response to this type of accident is inherently self-limiting due to the compensating reactivity feedback effects. Furthermore, the Reactor Protection System provides a high neutron flux trip function. The protection criteria for this accident are that the minimum DNBR be greater than 1.3 and that sy.tes pressure limits not be exceeded.

s.

The reference analysis for this event demonstrates that thermal power and system pressure remain below the specified limits and DNBR remains above 1.3.

8.3.5.2 Key Safety Parameters The reference analysis is based on the following parameter valu-s

Doppler Coefficient Ak/k/*F -1.3 x 106 Moderator Temperature Coefficient, Ak/k/*F -3.0 x 10~4 8-7 Rev. 1 i

l I

l l

The MTC value is representative of end of core life conditions and provides  !

a most-negative bound on the range of expected values. A less negative value would lessen the severity of the accident by decreasing the positive reactivity

~ insertion. The Doppler value assumed is a conservative, least negative bound.

Values.in the conservative direction (more negative) would also lessen the severity of the accident through enhanced feedback effects during a power increase.

8.3.6 Loss of Coolant Flow 8.3.6.1 Accident Description A reduction in reactor coolant flow rate occurs if one or more of the reactor coolant pumps should fail. A pumping failure can occur from mechanical failures.

or from a loss of electrical power. The effects of loss or reduction in coolant flow are an increase in coolant temperature and system pressure which could result in exceeding the core thermal limits if the reactor is not tripped promptly.

The core protection criteria of concern in this event is the minimum DNBR, which must be greater than 1.3 for electrical malfunction events and greater than 1.0 for mechanical malfunction events. Reactor protection is provided by three RPS ,

l trip functions: power - number of RC pumps, power - flow - imbalance, and RC j pressure - temperature.

The reference analyses demonstrate that the reactor can sustain a loss-of-coolant-flow accident without damage to the fuel. The minimum DNBR reached for the loss of flow due to electrical failure analyses was greater than 1.3.

The analysis for the loss of flow 'due to mechanical failure -(lockeil rotor) demonstrated a minimua DNBR greater than 1.0.

8.3.6.2 Key Safety Parameters l

l The reference analyses were based on the following parameter values:

Doppler Coefficient, Ak/k/*F -1.2 x 105 Moderator Temperature Coefficient, Ak/k/*F +0.5 x 10~4 l 8-8 Rev. 1 l

-Reactor Coolant Flow, gpa 352,000 Core Power Peaking Factors

-radial-local 1.783 axial-(cosine) 1.5 These reactivity coefficient values provide least negative bounds _for the expected range of values. More negative values would amplify the compensating

. feedback effects due to increasing temperatures. The assumed reactor coolant flow is extremely conservative compared to the available flow (greater than 108.5%). The core power peaking factors considered in this analysis are those corresponding to the maximum design condition. The combination of 1.783 (radial-local) x 1.5 (axial cosine) is more conservative with respect to DNBR criteria than any other power shape that exists in typical reload cores.

In addition, the effects of a loss of coolant flow accident are strongly '

influenced by the flow coastdown characteristics, fuel densification and rod bow effects, and hot channel power peak augmentation factors. These parameters are not expected to change during the normal reload design process. If changes do occur, their impact on the reference analyses will be evaluated.

8.3.7 Stuck-Out, Stuck-In, or Dropped-In Control Rod Accident 8.3.7.1 Accident Description In the event that a control rod beccnes significantly misaligned from the other control rods in its group, the effect of such a condition on localized power peaking (flux distortioa) and on available shutdown margin must be con-sidered. A stuck-out control rod reduces the available shutdown worth and o

hence-reduces the shutdown margin. The effects of this accident are mitigated by requiring a shutdown margin of 1% Ak/k, with the control rod of greatest worth fully withdrawn from the core. A stuck-in or dropped-in control rod causes neutron flux distortions that could result in localized power densities and heat' fluxes in excess of the design limits if the reactor is allowed to return to full power. The effects of this type of accident are mitigated by 8-9 Rev. 1

providing sufficient margin between the expected,' distorted power peaks and the design limits to prevent the limits from being exceeded. The core protection criteria of concern for these events are that the minimum DNBR shall be greater than~1.3'and that the system pressure shall not exceed code allowable limits.

The reference analyses demonstrate that, even in the absence of ICS action (to accomplish a power runback to 60L FP) or a reactor trip, thermal power does not exceed the original velues nor does system pressure exceed allowable limits.

8.3.7.2 Key Safety Parameters The reference analyses are based upon the following parameter values:

Doppler Coefficient, Ak/k/'F -

-1.3 x 10 s Moderator Temperature Coefficient, Ak/k/*F -3.0 x 104 Maximum Dropped Control Rod Worth, % Ak/k HFP, No Xe 0.46 HFP, P'Xe 0.36 The MTC value is representative of end of core life conditions and maximizes the positive reactivity insertion during the initial temperature decrease.

The Doppler value is a conservative, least negative value that minimizes the compensating feedback effects during a return to power.

1 8.3.8 Loss of Electric Power 1

8.3.8.1 Loss of Load Transient 8.3.8.1.1 Accident Description The effect of of a loss-of-load condition on a unit would be that the unit generator breakers would open and thus disconnect the unit from the trans-mission system. When this occurs, a runback signal causes an automatic power reduction to 15 percent power. Depending on the initial power level at the time of the loss of load, the Reactor Protection System may initiate a reactor trip on high recctor coolant temperature or pressure.

8-10 Rev. 1

1 E The loss-of-los'd accident does not result in any fuel damage or excessive apressures on the reactor coolant system. There is no resultant radiological hazard to station operating personnel or to the public as only secondary system steam is discharged to the atmosphere. Unit operation with I percent defective fuel and 1 gpa steam generator tube leakage is demonstrated to be safe by the reference analyses. For these conditions, the steam relief accompanying a loss-of-load accident Oould not change the whole body dose

because the primary contributors are normally released thrcugh the condenser i air ejector.

8.3.8.1.2 Key Safety Parameters 4

'The course and consequences of this accident are independent of the parameters affected in the reload design.

8.3.8.2 Complete Loss of All Station Power 8.3.8.2.1 Accident Description T

! The hypothetical. initiator of this accident is the complete loss of all station power except the station batteries. The loss of power results in gravity insertion of the control rods and trip of the turbine stop valves.

'The main steam safety valves prevent excessive temperatures and pressures

. in the reactor coolant system. The reactor coolant system flow decays without fuel damage occurring, and decay heat removal is provided by natural circulation. The turbine-driven emergency feedwater pump, taking i its suction from the condenser hotwell and upper surgetank, provides feed-water to the steam generators. Condenser cooling is maintained through a gravity feed line from Lake Keowee. The station batteries provide power for the necessary control and auxiliary systems.

The reference analyses demonstrate that neither fuel damage nor excessive

- pressures occur.

i 8-11 Rev. 1

8.3.8.2.2 Key Safety Parameters The course and consequences of this accident are independent of the parameters affected in the reload design.

I 8.3.9 Steam Line Failure 8.3.9.1 Accident Description The steam line failure accident assumes a break in the secondary system press-ure boundary that results in inadequate secondary pressure control. The worst case steam line failure involves the maximum break size (34 inch diameter) at rated power and end of core life. Under these conditions, the rapidly decreasing secondary pressure results in an excessive primary system cooldown which, under the influence of a negative moderator temperature coefficient, produces a ;csi-tive reactivity insertion. If feedwater flow. continues to the affected steam generator the excessive heat removal and concurrent primary cooldown will conti-nue and the reactor may experience a return to low power levels if the positive reactivity inserted exceeds the shutdown margin. The reactor coolant contraction accompanying the primary cooldown may result in ECCS actuation.

i 1

The criteria for unit protection and the release of fission products to the environment are:

l

1. That the core will remain intact for effective core cooling, l assuming minimum tripped rod worth with a stuck rod.

2. That no steam generator tube loss of primary boundary integrity will occur due to the loss of secondary pressure and resulta:it temperature gradients.

3. That doses will be within 10CFR100 limits.

l The reference analyses consider three major accident scenarioes: (1) the base case that assumes proper ICS and operator actions; (2) a case that assumes ICS action but no operator action; and (3) a case that assumes neither ICS nor 8-12 Rev. 1

l

. operator action. The reference analyses demonstrate that the protection criteria ]

are satisfied.

8.3.S.2 Key Safety Parameters The reference analyses were based upon the following parameter values:

Doppler Coefficient, Ak/k/*F -1.2 x 105 4

Moderator Temperature Coefficient, Ak/k/*F -3.0 x 10 Available Scram Worth, % Ak/k 3.46 The MTC value used in the analyses provides a most negative bound for the expected range of values. Less negative values would decrease the positive reactivity insertion and thus lessen the severity of the accident. The Doppler Coefficient is a least negative value that minimizes the compen-sating feedback effects during a return to power. A minimum rod worth value yields the most adverse effects.

8.3.10 Steam Generator Tube Failures 8.3.10.1 Accident Description The occurrence of a double-ended rupture of one steam generator tube would result in the release of the activity contained in the reactor. coolant to the secondary system. The initial leak rate is in excess of the normal

- makeup flow and hence would result in a low reactor coolant system pressure or pressure-temperature trip. Continued primary to secondary flow would result in the automatic initiation of the high pressure injection system which would provide sufficient makeup to compensate for the tube leakage and thus terminate the depressurization.

The volume of primary coolant released to the atmosphere through steam relief would' produce acceptable consequences.

i i

)

l l

l 8-13 Rev. 1

1 l

1

~8.3.10.2 EKey Safety Parameters The course and consequences of this accident are independent of the parameters  ;

affected in the reload design.

l

\

8.3.11 Fuel Handling Accidents -

The reference analyses for the fuel handling accident are not affected by the reload design process.

8.3.12 Rod Ejection Accident 8.3.12.1 Accident Description For reactivity to be added to the core at a rapid rate, physical failure of a pressure barrier component in the control rod drive assembly must occur.

Such a failure could cause a pressure differential to act on.a control rod assembly and rapidly eject the assembly from the core region. The power excursion due to the rapid lacrease in reactivity is limited by the Doppler effect and terminated by the Reactor Protection System.

The criterion for reactot protection in this accident is that the reactor will be operated in such a manner that a control rod ejection accident will not further damage the reactor coolant system.

The consequences of the rod ejection accident are largely dependent upon the rate at which thermal energy is released to the coolant. In turn, the amount of thermal energy is a function of the vorth of the ejected rod and the initial power level. -The reference analyses include calculations for a range of ejected rod worths at rated power and hot zero power and at beginning and end of core life. The effects of varying the Doppler and moderator coefficients and rod worths are also calculated. The analyses demonstrate that the reactivity l transient resulting from this accident will be limited by the Doppler effect and terminated by the RPS with no serious core damage or additional loss of the coolant system integrity.

l l

8-14 Rev. 1 i

i

8-3.12.2 Key Safety Parameters

.The reference' analyses are based upon the following parameter values:

BOL EOL Doppler Coefficient, Ak/k/*F -1.17 x 10 ~5 -1.33 x 10 s Moderator Temperature Coefficient, Ak/k/*F +3.5 x 10 ~4 -3.0 x 10 ~4 Delayed Neutron Fraction 0.0071 0.0053 Neutron Lifetime, p, 24 8 23.0 Ejected Rod Worth, *a ak/k

.HZP 1.0 1.0 HFP 0.65 0.65 The' analyses calculate the effects of an ejected rod using a spectrum of reactivity coefficients between-the values shown. The MTC bounds define the range of allowable values based upon Technical Specification limits. The results.of the Doppler sensitivity study show the highest neutron power for

-the least negative coefficients. Thus, more negative values would lessen the severity of the accident. The kinetics parameters are nominal values that are representative of the range of values expected in the reload design work.

The rod worth values provide upper limits for the calculations. Lower rod worths would lessen the severity of the accident.

8.3.13 loss of Coolant Accident 8.3.13.1 Accident Description A loss of coolant accident (LOCA) occurs when a break in the reactor coolant

' pressure boundary results in coolant expulsion in excess of the normal make-up flow rate. The blowdown rate, the time period before reactor trip and ECCS actuation and the amount of stored energy initially removed from the core are all dependent upon the break size. In order to evaluate this accident, a range of rupture sizes from small leaks up to the complete severance of a 36 inch ID primary coolant pipe have been evaluated.

8-15 Rev. 1

i 4 1 \

l Should a break occur, depressurization of the RCS causes coolant to flow out of the pressurizer into the primary loop resulting in a pressure and level decrease in the pressurizer. A reactor trip occurs when the low pressure l or pressure-temperature trip setpoints are reached. The Engineered Safeguards system is actuated when the appropriate setpoints are reached. The consequences '

of the accident are limited in two ways:

1. The reactor trip and borated water injection complement void

{

formation in causing a rapid decrease in nuclear power to a residual level corresponding to the delayed fission and fission product decay.

2. Injection of borated water ensures sufficient flooding of the core to prevent excessive clad temperatures.

The core protection criteria for a LOCA are specified in the regulatory requirements of 10CFR50.46. Briefly, the five criteria are:

1. The peak cladding temperature shall not exceed 2200'F.

2. The percentage of local cladding oxidation shall not exceed 17%.

3. The percentage of hydrogen generation resulting from whole-core cladding oxidation shall not exceed 1%.

4. Calculated changes in the core geometry shall be such that i the core remains amenable to cooling. l

5. The mode of long term cooling shall be established.

The reference analyses 11 demonstrate that these criteria are satisfied at all times.

l l

8-16 Rev. 1 l l

I

-l 1

8.3.13.2 Key Safety Parameters The reference analyses are based upon two major input parameters that are affected by the reload desi,s process. They are:

Average Fuel Temperature, 'F (@ 18 kw/ft) 3120 Peak Linear Heat Rate, kw/ft Core Elevation, ft 2 15.5 4 16.6 6 18.0 8 17.0 10 16.0 These parameter values are the limiting values applicable to the generic ECCS

analysis, s

8-17 Rev. I

.. - - - - -. . .=. . ._

I

. 1 l

1 Table 8-1 ACCIDENT ANALYSIS REVIEW XEY SAFETY PARAMETER CHECKLIST l

Reference Analysis Conservative Reload Cycle )

Parameter (Units) -

Value(s) Direction Value(s) 1

' Doppler Coefficient, -1.17 x 10~8 more cegative ak/k/*F s

Moderator Temperature BOL +0.5 x 10~4 less positive

-Coefficient, Ak/k/*F EOL -3.0 x 10~4 less negative Delayed Neutros.- BOL 0.0071 Fraction, (nominal) EOL 0.0053 Neutron Lifetime BOL 24.8 (nominal), micro- EOL 23.0 seconds Total Rod Worth, 10.0 smaller

% Ak/k Maximum Ejected Rod HFP 0.65 smaller Worth, 1 Ak/k HZP 1.0 smaller ,

Maximum Dropped Rod HFP, no Xe 0.46 smaller l Worth, % Ak/k HFP, w/Xe 0.36 smaller Minimum Tripped Rod 3.46 larger Worth (for SLB),

% Ak/k Minimum Shutdown 1.0 larger Margin, % Ak/k l

8-18 Rev. 1  !

Table 8-1 (cont'd)

ACCIDENT ANALYSIS REVIEW KEY SAFETY PARAMETER CHECKT,IST Reference Analysis Conservative Reload Cycle Parameter (Units) Value(s) Direction Value(s)

Boron Worth, PPM / 75 larger

% Ak/k Average Fuel Tem- 3120 smaller perature at 18 kW/

ft,*F Peak Linear Heat Rate,

, kA/ft Core Elevation, ft 2 15.5 smaller 4 16.6 smaller 6 18.0 smaller 8 17.0 smaller 10 16.0 smaller Reactor Coolant Flow, gpm 352,000 larger Core Power Peaking Factors radial-local 1.783 smaller axial (cosine) 1.5 smaller 8-19 Rev. 1

e ele Reload Cycle s Reload All Values Within Range Design Parameters

/\

Some Values Current Out of Safety Analysis Range Valid sr Detailed Review Effects Iterate on f Analyses ,

Reload Design Revised Effects Licensing Basis

( / Uncertain or Safety Analysis

\ Unacceptable

\/

Consequences Consequences Acceptable Unacceptable Re-analysis of Event a-Figure 8-1 Accident Analysis Review Process 8-20 Rev. 1 y ,w - n-- w -. ,, mr.g e, y -

w-- -w--- ,- +---- , , - - - --w - - ,.- c ,-- 9 -g

9. -DEVELOPMENT OF CORE PHYSICS PARAMETERS

- Upon completion of the Final Fuel Cycle Design and the Maneuvering Analysis both PDQ97 and EPRI-NODE depletions, rod scans, boron concentrations and worths, power {

distributions, etc. have been generated primarily for HFP and some HZP conditions.

The purpose of this stage of developing core physics parameters is to provide l additional, calculations to supplement those already performed. The results of these calculations are used for startup test predictions and core physics para-meters throughout the cycle.

9.1 Startup Test Predictions After each refueling, the reactor undergoes a startup test program aimed at verifying that the reactor core is correctly loaded, control rods are in the cor-rect locations and are functioning properly, and to verify reactor behavior is as predicted by the nuclear simulators which were used in generating the data 4 used in the plant's safety analysis.

7 9.1.1 Critical Boron Concentrations and Boron Worths EPRI-NODE and/or PDQ97 may be used to calculate critical boron concentrations and boron worths at a variety of rod configurations, at HZP and HFP, as a function boron concentration, at different xenon concentrations, and at different times in the fuel cycle. EPRI-NODE and PDQ97 both are capable of critical boron searches and when critical boron concentrations are desired are usually run in this mode.

An acceptable alternative, however, is to not search on critical boron but to correct the input boron concentration to the critical boron concentration using a calculated boron worth and the calculated reactivity.

Table 9-1 shows some of the critical boron calculations normally performed for startup physics tests. Table 9-2 shows the soluble boron worths usually per- '

formed for startup physics tests. The boron worths are usually calculated by running two identical cases except that the soluble boron concentration is

[ different. The differential boron worth is calculated by subtracting the re-activities and, dividing by the boron difference. Differential boron worths are usually quoted in %p/100 PPM or in PPM /%p (the latter is sometimes referred to as the inverse boron worth).

9-1

- - .-. - - - . - - - _ _ - - - - . ~ , - - . - . . _ , . - . - . - -

l-(

Critical boron concentration at 68'F, 532*F, and KFP'with all rods out except i

APSR's 4.s calculated as a function of cycle burnup. Figure 9-1 illustrates the form in which these results are displayed.

Differential boron worth is calculated as a function of boron concentration and also as a function of cycle burnup. Figures 9-2 and 9-3 show the results of these types of calculations. Integral boron worth calculation is performed at

'BOC (4EFPD) as a function of boron concentration. The results of this are dis-played in the format illustrated by Figure 9-4.

9.1.2 Xenon Worths l

Xenon worth is calculated as a function of cycle burnup using either PDQ97 or EPRI-NODE. The nominal HFP depletion cases with equilibrium xenon are used as input to a second set of cases where the xenon concentration is set to zero (or the xenon cross sections are set to zero). The difference in reactivities between the equilibrium xenon and no xenon cases equals the equilibrium xenon worth at HFP. The results are displayed in a format similar to Figure 9-5.

1 l

l i

l l

9-2

9.1.3 Rod Worths 9.1.3.1 Group Worths The worth of groups 1 to 8 and the integral rod worth curves for groups 5-7 are calculated at BOC HZP for use in the zero power physics testing. The rod groups are sequentially inserted or withdrawn from the EPRI-NODE calculation assuming no control rod overlap. The group worth is the difference in reactivity between the fully inserted case and the fully withdrawn case.

At HFP equilibrium xenon BOC (4EFPD), the above rod worths are calculated in a

. similar manner except that when calculating the intergral rod worth curves a control rod overlap of 25% is used.

At'HFP and HZP group 8 rod scans are performed where group 8 is stepped in small increments into or out of the core. The HZP results are used to provide tables of rod worth versus position and plots of relative rod worth versus position.

The HFP results are used to provide the same information plus a table of imbalance as a function of rod index. Rod scans on group 7 are performed at BOC HFP to provide a table of imbalance versus rod position.

9.1.3.2 Stuck Rod Worth The maximum worth of a single control rod stuck out of the reactor core at HZP is calculated during the final fuel cycle design. The worth of the stuck rod is used by the site engineers in the reactivity balance procedures to guarantee shutdown margin. If the stuck rod worth is to be measured during the startup test program, then a recalculation of the worth is performed simulating the test conditions. This worth would then be provided as a startup test prediction.

9.1.3.3 Dropped Rod Worth The maximum worth of a single control rod dropped into the reactor core is cal-culated during the final fuel cycle design. If this parameter is to be measured during the startup test program, then a recalculation of the worth is performed simulating the test conditions. This worth would then be provided as a startup test prediction.

9-3

q

. 9.1.3.4 EjectedIRod Worth' 1 l

During startup physics testing the maximum ejected control rod worth at HZP is measured and compared to the predicted worth. The maximum ejected rod worth is calculated during the final fuel cycle design (Section 3.2.2.3) and a re-calculation of this parameter is not usually necessary since the calculation is l Y

performed at conditions similar to those used in the testing.

l

- 9.1.4 ' Reactivity Coefficients 9.1.4.1 HZP Coefficients

- At HZP'the isothermal temperature coefficient is measured by reducing the average moderator temperature 5 F to 527 F taking data once equilibrium is reached then increasing the temperature 10 F'to 537 F taking data and establishing equilibrium.

The temperature is then reduced 5 F back to the original 532 F value. The cal-culations used for predicting the isothermal temperature coefficient should be run 'at 527*F and 537 F using either EPRI-NODE or PDQW7. The resulting reactivity change is then divided by the 10 F temperature change to yield the NZP isothermal temperature coefficient.

l The Doppler or fuel temperature coefficient at HZP can be calculated by varying the fuel temperature while maintaining the moderator temperature constant at 532*F.

The-resulting reactivity change divided by the change in fuel temperature is.

the Doppler coefficient at HZP.

l The predicted moderator coefficient is calculated by subtracting the Doppler co-efficient from the isotbermal coefficient and is compared to the measured mod-

. erator coefficient obtained by subtracting the predicted Doppler coefficient from-the measured isothermal coefficient. Alternately, the moderator temperature coefficient can also be explicitly calculated.

9.1.4.2' HFP Coefficients L Both a temperature coefficient of reactivity and a power doppler coefficient of I

l l

9-4 l

reactivity are measured at HFP. Changes in temperature or power are compensated for by control rod insertion or withdrawal. A calculated Doppler coefficient is subtracted from the temperature coefficient to obtain the moderator coefficient.

A moderator coefficient is calculated by running one equilibrium HFP case at BOC (4EFPD EPRI-N0DE or PDQ97) and a second case which has lowered the moderator temperature 5"F. The difference in reactivity divided by the temperature change 1 is the moderator coefficient.

A third case is run to determine the power doppler. In this case the power level is reduced to 95% HFP. The difference in reactivity between the HFP and the 1 95% HFP cases divided by 5% FP is the power doppler coefficient.

9.1.5 Power Distributions Power distributions, both assembly radial and total peaking factors, are sea-sured at 40 and 10C% HFP for Oconee reload startups. Calculations using 1 EPRI-NODE are run at these power levels and nominal conditions to provide predicted power distributions to compare to measured. Typical power distri-butions generated are shown in Table 9-3.

9.1.6 Kinetics Parameters Kinetics parameters are calculated using the methodology and codes as discussed i I in section 3.2.8. These parsneters include the six group S effective and A ,

total $ effective and A, and reactivity versus positive and negative doubling times. These kinetics parameters are generated for two sets of HZP conditions.

The first is with group 8 inserted and the second is with groups 5 through 8 1 inserted. In addition to the BOC HZP parameters, one set of BOC NfP parameters are generated with groups 5-7 at 100% WD and group 8 at 37.5% WD.

9-5

i 9.2 Core Physics Report The purpose of the core physics report is to document the predicted behavior of l the reactor core as a function of burnup and power level. It is intended to be used for operator guidance and the site engineer. Portions of the information

~

included will reiterate data found in the final fuel cycle design report and the startup test prediction report, however, much data not needed for these reports is useful to the operator and site engineers. t This report will include sufficient information to calculate reactivity

. balance throughout the cycle. Table 9-4 lists items typical of what will be calculated for this report. Any additional c.'lculations will be performed using either EPRI-NODE or PDQW7.

l i

9-6 l

l

Table 9-1 CRITICAL BORON CONCENTRATIONS (PP.'f)

HZP,.N0XE, WEFPD ARO CRGP 1-7 out CRGP8=37.5%WD CRGP 7 in CRGP8=37.5%WD CRGP 6,7' in CRGP8:37.5%WD CRGP 5-7 in CRGP8=37.5%WD CRGP 4-7 ,in CRGP8=37.5%WD CRGP 3-7 in CRGP8=37.5%WD CRGP 2-7 in CRGP8=37.5%WD CRGP 1-7 in CRGP8=37.5%WD HFP, NOXE, WEFPD CRGP 1-6 out, CRGP 7 in, CRGP 8=37.5%WD HFP, EQXE, 4EFPD ARO CRGP 1-7 out CRGP8=37.5%WD CRGP 7 in CRGP8=37.5%WD CRGP 6,7 in CRGP8=37.5%WD.

CRGP 5-7 in CRGP8=37.5%WD HFP,EQXE,EOC

. CRGP 1-7 out CRGP8=37.5%WD 9-7

l Table 9-2  !

i BORON WORTH (PPMB/%Ap) i HZP, NOXE, CRGP 7 and 8 in p EFPD, XX PPMB '

1. EFPD, XXXX PPMB Rod Patch, XXX PPMB l EOC, XXX PPMB l HFP, EQXE, CRGP 7 and 8 in 4 EFPD, XXXX PPMB Rod Patch, 17 PPMB EOC, 17 PPMB I

I I

9-8 l

l

-a n- n - ,- ,--1 --l- ,e----w,~ww ,--,,g-,- -

-e-r- - p~~w-- -----,-- ,v-- --

~-m-~ - --,------mg+-we-

trF1-Y Table 9-3 RADIAL AND TOTAL PEAKING POWER MAPS

i CONDITIONS POWER LEVEL BURNUP

-% HFP EFPD-40 2 100 4 1 100 12 100 25 1

f

, 4 e

k gf f

9-9

Table 9-4 Core Physics Data A. Critical Boron Concentrations

1. ARO HFP Versus Burnup

2. ARO HZP Versus Burnup

3. .ARO 68 F Versus Burnup B. Critical Boron Concentrations required for 1% shutdown with highest worth rod stuck out (NoXe)

1. HZP Versus Burnup

2. 68 F Versus Burnup C. Differential Boron Worth HFP, HZP versus burnup.

D. Power Distributions from the Cycle Depletion ~

E. Rod Worths BOC, EOC, HFP and NZP F. Imbalance versus APSR position BOC, EOC at HFP G. . Imbalance versus Group 7 position BOC, EOC at HFP H. Xenon worth versus Power Level l

I. Xenon Worth versus Burnup i

I I

l 4

' 9-10 l I

1 i i i

.1

t-FIGURE 9-1 l BORON LETDOWN CURVES CRGP 1-7=100%WD l 1

CRGR 8-37.5%WD I

\

i l

l l

l t i 1

1 A

_--- __ =_____ _ _ = = ____ . = = _ . _ _ _ _ _ _ =

e __ = =_ = = - - , .__- - - -

12% ======-_____ __ __

= = _ -

-=__ =_=_._ _-__

= -

===-.-==_=--_-

_ -_==_=__-- --

1

_=. - __ =_ - =_-__-=_ _ _ =_ , ___=_.=_ =._ =--- _ . _ _ = =_ _ _=_= _ ._=_- _- ,_ - -- _- - ,

_ -= _=__ _ __ _ =- -_---_ _ _ = = _-- -

< = - _ _ - . __

_- ==ssrrrr__ --= srr-s=-_ _ _ _ _-

' x _ _= - -- -

1 1

=- --_

1000___._-.=___==__-_____---_=======-=__==_==___==_____

=_ = ___

= = = -

=

_ = _ _ _ __ - - _

=_=_= _ - - _- _-___.- - - - -

_=

s == s

_=________=_-- c__ _ __ _-_ _= = .r .

__ _ =a_ s__a_ r, =_= -

__ =__ =_ s_

gg =. = gs= == =- = =e 1

3= == =- _ _ === s==.

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800 =

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~9 2 - =__ =,P__

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t

_ - - _ = _

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. = = = - = = _ . - - - _ = = - - - _ = = = _ = _ - - - - = - - - - - - - - = - __.

=_ =_ = =_---- - __

= = = = = = = = -

_ =__= _ = = = ==_=__ = = =_ _=-_ ____--;_====--

_ _ _ = = = = = = = = = = = = _ =

_ - =__=_ _ _=_=_ =___=_

7

=

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t 400 =_ r a r r r -

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===-mm====--*

l ======-- = ==========:==

- & & & & &=-~ = = ==

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======--_===--=-=--=-====--.=_==:==============.

--==~

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= - - -

=========.-===_=_=:===================_===-

=

E EZ- EEE==;=;=;;EM

=

i

===__==_

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= = = = _ =.--

-- - = ==

= = =_= .== = = .=_ 3-[ s= 5 5 35 h[ ' k[ ] ~ ~-

~~~~*-----

?_[-@~,_- _--- ;-

---..-~~===========_=-=_==_=.;=_==r=._=-=.===

._.'~"---"~

== -

====

0=--==-~~===.--=====-----==------.===_ ==g---- =

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__7-----------.-;__- .

= = = = =

= =- -

_ _ _ _w

===._.=_..=.__.-=.-_.._=.=_=__=====__====-_=====-==_===;___=_== -

=

_==___=.______...____=___=_=._=....___=_=__5s==_=-.==_=.___==_____._..._______.____= . _

i 0 100 200 300 BURNUP. EFPD 1

9-11 .

J

-. ..~ . ._..__... .. _..._ ___. _

! FIGWtg 9-2 i i

.DIFFERINTIAL BORON WORTH

{

L  !

! VERSUS BURNUP EFP,54XE i

i f

i I

~~-5

= = , _555EE15

= - - - - EE===E . _~=Ei-i=5 = = ~EE= = . 6-E- _ = =E = E- EEE EEE EEE E59=' 555 -=

E=45J-, <

1.04 =F = = ="E - Z EE=EE E~~-=

g E E E=gg= g=guss gg=EEE=E EEi - FRCN 0-100 RI=193 2EHE" :-

EiEE=h = =E MEE=T= EEE EEE ===EE -

FRCH 100-235 RI=203 l _==_'_"=

GMM=23

- - _ _ _ _ . 55E*d5EMM,:5MM5 = - - _= =m = _ . _ _ = -

. FROM 235-270 RI=276 - M M =- -

1.02 - - - - -- g== g = g . =x = g= - g .

,- = - =;= . .

= = - - ===:=====_==-:: - :: =- --_-===_==:---=-

ETE - -"=E =

= =E' A =g E s

=55

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- L-=="- M =-

RD

- == ======== ==== _

=

= -.

== =-

.== :

== =:E= =z wE= - == = 25% :+: -L_

- .w:m_ - -.:. n =_m t = _

c m =E  :

_ a= mm E- z 5 = _ E= - ia _ ; - u_= _+u - u = w =_=:m == m .=z =

1.00 -

l l'=r=-J=fE=EM=

- 5= f E f M .7,5@ 5 ::f heli 5"b 42=?=:n-O?11Lc=E90?i==fQ,--

z M- ~- 5: E- [ 'j- 5=E = E ?-5 -$-' ~ .d E N .

z

mmi E = t i_a == 7 2 z

6 8 =: = = = == - . : : == = = = = : = = : = -:=:'===-----x=:=-=

=: = =mi 5 m: = mr E=a m m - -

O.

a *. E _ _ gEs!

, _ _ = = =M"E

= 5 545 z3h-E EWE == ~ kf= .h=i='E=E =::- =W-~.4 = = =M,dMP === . .--

= = = = = = = =. =-~:-.~ ==:^'

.; u~~_. ^ ~. :=: . : ~ = = L^:: '-~;* =L9.=

o.98 - ** _

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= == _== z= :;c+ _  : c n' = ~ E = f.

.- =

n  := _=_- - == =_ - = 5 = = _-; g =_ -

===-==-:.::==-==- -.==:- - :- = ^ -

z ; q = z : _ . :5 2=;2.-

= = =: -- : = n

_= = r ,

=_u-=:

- E=2 u + =c = + s = u x =- - -

=-: . s = c = : 5+ ++ l i

-- ._=_a_._ - s_==_ =+ =_=. c.__ =_n _=_ m_ _e . _e ___ _ _ _3 x .- _ :_=: n.

- =_ =. -.

g-.= g - =9:- _ ._-- .- :- _= +- w- -: -_

y =_= 3 =-- = - ._.= = = - - m . ; = _=:

= = O . - a u + - g: - - - -

= -

=

=.n.

+ ==--

-- = =-

_ 5w~ _ _

0.96 4,1;EE g =a - C _ 3 =_: M :E =] E = 3 W + -~~ J-:--

~

^ ::-=== 2- - : = g=  : : 5.=Z g -

_ = = = = = = . = - - . i = ~.. =. -.=' = :'~~. -.-. =~-~'. :-----

: ' = = = :0 - -_, L _ _ = 1. ~_

i - ---- '

- = -

7 ~~--

-'.b

. b _- .. b: 1 -~. -

^--- '

- 2- ..

~n %

: += 8:2n: = = e E- -E a= = -

=-.:---===.-=-':==:== . - - - = = - - .=.-

-= .5 - .= :- :==-  : =--=+: = :=-- =.

L

=EEE=-i==: E~ EEE *= b - G-d=I= L i ? = 4 W 4 i=ci_. LT;jf l 0.94 s.+ =.=i=== IE ATE =ris J= = F - 4 1_: n i L M;cl J.13 ===== --

g-a-- = = =s; E+= r _:- - n = =7 n - -

--- - :: - - . . .. 7 7-t 0 100 200 t

300 A

e l i 1

1 l

l I

l 9-12 i

l l

En v

)

1.

~

l l

FIGURE 9-3 f

I DI g 1AL 30 ROM WORTE l

, 4 mm, , u-m  !

f.

i

- I

-  ?

i -

l' . .

i 1

k l

1 I

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[- .

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g f l

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f l l i.

~

- +

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l '3 t

1 . . . . _ _ . - _ ,

l t'

l g

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= -. ...

.. s1 - .

e

! 8 8 8 E E E R f

. . . e e e s '

j l . . Widd 001/8 Y I l . . . .. 9-13

FIGURE 944

.IlfrECRAL BORON WORTH HFP, 4 EFPD, EQXE, 11=XXX i . ,. _ .; -

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=== =:

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=: a -

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~ .

2 }..: , 2*

3: s . y- n*r; &_ ~%m t_. n9 .:; r.

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.; n m -11 : na wn 8

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pwAy n

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e '3 ~ C,L: .29 'o b

"* GD to y N C$

07I l

9-14

( m h. n. " ' ' ' '

FIGURE 9-5 EQUILIBRIUM IENON WORTH VERSUS BURNUP AT HFP

• ~

FROM kOORI193 _

, FROM 100-C5 E=203 FR m 235-270 E=276 __

l 2.9 a

4 2.8 n 2.7 2.6 BURNUP. EFPD 9-15

10. REFERENCES

1. Program to Determine In-Reactor Performance of B&W Fuels - Cladding Creep 2

Collapse, BAW-10084, Rev. 2, Babcock & Wilcox, Lynchburg, Virginia, October, 1978.

2. R. A. Turner, Fuel Densification Report, BAW-10054, Rev. 2, Babcock &

Wilcox, Lynchburg, Virginia, May 1973.

3. Oconee 1 Fuel Densification Report, Revision 1, BAW-1387, Rev. 1, Babcock & Wilcox, Lynchburg, Virginia, April 1973.

4. A. J. Eckert, H. W. Wilson, and K. E. Yoon, Oconee 2 Fuel Densification 3 Report, BAW-1395, Babcock & Wilcox, Lynchburg, Virginia, June 1973.

5. Oconee 3 Fuel Densification Report, BAW-1399, Babcock & Wilcox, Lynchburg, Virginia, November 1973.

6. TACO 2 - Fuel Pin Performance Analysis, BAW-10141, Babcock & Wilcox, 2 Lynchburg, Virginia, August 1979.

7. Correlation of Critical Heat Flux in a Bundle Cooled by Pressurized Water, BAW-10000A, Babcock & Wilcox, Lynchburg, Virginia, June 1976.

8. Oconee Nuclear _ Station, Units 1, 2, and 3, Final Safety Analysis Report, Docket Nos. 50-269, -270, and -287.

9. Electric Power Research Institute (EPRI), Advanced Recycle Methodology Program (ARMP) System Documentation, September 1977.

10. J. M. Alcorn and R. H. Wilson, CHATA - Core Hydraulics and Thermal Analysis, BAW-10110, Rev. 1, Babcock & Wilcox, Lynchburg, Virginia, 3 May 1977.

11. R. C. Jones, J. R. Biller, and B. M. Dunn, ECCS Analysis of B&W's 177-FA Lowered Loop NSS, BAW-10103A, Rev. 3, Babcock & Wilcox, Lynchburg, 1 Virginia, July 1977.

10-1 Rev. 3

W 12.

J. H. Taylor (B&W) to D. B. Vassallo (NRC), Letter, " Determination of the 2 Fuel Rod Bow DNB Penalty," December 13, 1978.

13.

D. 3. Vassallo (USNRC) to J. H. Taylor (B&W), Letter " Calculation of the 2 Effect of Fuel Rod Bowing on the Critical Heat Flux for Pressurized Water Reactors," June 12, 1978.

14.

J. M. Alcorn, H. C. Cheatwood, C. D. Morgan, and R. H. Wilson, TEMP - 3 Thermal Energy Mixing Program, BAW-10021, Babcock & Wilcox, Lynchburg, Virginia, April 1970.

i

15. J. R. Gloudemans and H. C. Cheatwood, RADAR - Reactor Thermal and 3 1

Hydraulic Analysis During Reactor Flow Coastdown, BAW-10069A, Rev. 1, 1 Babcock & Wilcox, Lynchburg, Virginia, October 1974.

1 l

16. TACO - Fuel Performance Analysis, BAW - 10087A, Rev. 1, Babcock & Wilcox, 4 Lynchburg, Virginia, August 1977.

10-2 Rev. 4

(.,

t 5

. APPENDIX A - CODE

SUMMARY

3

,o -_ ._

l

}

}.

?

a i

E 9 Appendix A 1

I Code Summary i

l 5

f l

I

!~

I I

A-1 1

CASMO CASNO is a multigroup two-dimensional transport tnecry code for burnup calcula- 1 tions.on BWR and PWR assembliet. This code has been developed by Studsvick Energiteknik AB and supported by EPRI.

CHATA CHATA is a steady state closed channel thermal-hydraulic code which can be used in a multichannel or single channel configuration. It calculates flow, pressure drop, coolant properties, and DNBR. .It has several different options that give it the capability to iterate on an input parameter, such as finding the maximum power for a specified DNBR and pressure drop. It can be used to calculate assembly-by-assembly core flow distribution and a hot channel analysis, and is suitable for parametric studies because of its short running time.

COMETHE-III-J The COMETHE code calculates fuel pin thermal and mechanical behavior as a function of burnup. This code was developed by Belgo Nucleaire and licensed in this country by the S. M.' Stoller Corp. EPRI is sponsoring the distribution and further l development of this code for the utilities. The code does all the calculations described for TAFY and includes a relocation and cracking model to determine

. fuel-clad interaction forces.

1 CROV The Creep Ovalization Analysis Program for Fuel Cladding (CROV) calculates ovality changes in fuel rod cladding due to thermal and irradiation induced creep. CROV conservatively predicts the ovality time history and time to collapse'under a prescribed pressure, inside and outside temperature, and flux level time history loading.

The creep rate calculation utilizes a modified von Mises flow rule and includes a strain-hardening model. Empirical constants used in the creep rate equation

'are conservatively representative of B&W zircaloy-4 cladding at a temperature range between 600 F and 750 F.

A-2

l DELAY F . DELAY calculates core averaged delayed neutron fractions for six energy groups, core averaged decay constants for six energy groups, core averaged delayed neu-

~

" tron fraction with and without importance factor, estimated prompt neutron lifetime,;and reactivity versus period. -Input consists primarily of isotopic fission' fractions versus burnup and enrichment from PDQ97 calculations.

EPRI-CELL EPRI-CELL computes the space, energy and burnup dependence of the neutron spectrum within cylindrical cells of Light Water Reactor fuel rods. Its primary output

-consists of broad group, microscopic, exposure dependent cross-sections for subsequent use in multidimensional diffusion theory depletion analysis. EPRI-

CELL ~ utilizes'three inaustry accepted subcodes; GAM-1, THERMOS, and CINDER.

EPRI-CPM EPRI-CPM is a multigroup two-dimensional collision probability code for burnup calculations on BWR and PWR assemblies. The code handles a geometry consisting

-of cylindrical fuel rods of varying composition in a square pitch array with

~ allowance-for fuel rods loaded with gadolinium, burnable absorber rods, cluster control rods,.in-core instrument channels, water gaps, boron steel curtains and cruciform control rods in the regions separating fuel assemblies.

o

!- EPRI-FIT

'EPRI-FIT is a program which processes the PDQ97 integral file and calculates and edits values needed by the EPRI-NODE code. EPRI-FIT greatly reduces the hand calculation time needed to extract these values from the PDQ97 printout and improves the quality assurance. A data file under the local.name of COLOR

=is written which contains the EPRI-FIT edited data and is used as input to the SUPERLINK program.  :

4 A-3 if

- ~ - . .. .

EPRI-NODE EPRI-NODE is a multidimensional model code similar in theory to FLARE. The EPRI-N0DE program computes the core effective multiplication factor, the three-dimen- l

-sional core power distribution, core coolant flow and temperature distribution, and fuel exposure distribut' ion. The program includes-the effects of partially inserted full-length control rods, part-length rods, and up to 13 different fuel assembly types with different enrichments and burnable absorber shim loadings. I l

EPRI-NODE has a capacity to represent the core with 32 axial nodes for each fuel assembly and 30x30 nodes in the XY plane.

The program iterates to account for.the interaction between power distribution and core nuclear properties which depend on coolant flow and coolant temperature distributions, fuel temperature distribution and xenon distribution. The pro-gram computes the time dependence of xenon following changes in power level and/or changes in power distribution. The program permits fuel shuffling from one location.to another and fresh fuel insertion for burnup cycle calculations.

Individual steps can by stacked for either xenon transient or fuel cycle burnup

^

calculations.

1 EPRI-NUPUNCHER l l

NUPUNCHER prepares cross section tables in HARMONY format from' cross section l

data produced by EPRI-CELL and placed on the ECDATA file. NUPUNCHER reduces j significantly the tedious task of hand transferring values from the EPRI-CELL- l printout to macroscopic and microscopic tables in card image HARMONY format.

Two, three and four group cross section data may be obtained with one dimen-sional HARMONY interpolating tables.

EPRI-PDQ97 MODIFICATIONS PDQ#7_is the' industry accepted multigroup one, two, or three dimensional dif-

fusion depletion code. EPRI-ARMP uses PDQ97/ Version II with minor modifications I to. allow options for mixed number density, improved removal treatment, peak power editing, and re-editing.

A-4

t

. EPRI-SHUFFLE

^

The EPRI-SHUFFLE program will read a PDQ97 concentration file, make certain modifications to this file, and write a new updated concentration file. This procedure is accomplished by defining " assembly regions" in the program input.

Assembly regions are square arrays of mesh points containing depletable nuclide concentrations and superimposed on the original PDQ97 geometry. These assembly regions are then used to describe the movement of existing nuclide concentrations by translation, reflection and/or rotation. In addition, new fuel' concentrations can replace spent fuel concentrations in selected assembly regions described in the program's input.

EPRI-SUPERLINK SUPERLINK accesses data on the files produced by EPRI-FIT and together with

. relevant input information for file management and for data processing control produces polynomial coefficients for use in EPRI-NODE.

PDQ97 See EPRI-PDQ97 Modifications.

NODE UTILITY CODE (NUC)

The NUC program is a package of subroutines that performs any necessary utility function to EPRI-NODE files. The major subroutines are:

I. FILE - this mode lists, merges, purges, adds, rearranges, edits, etc. the NODE cases on one or more history files.

II. FLEX - this mode takes an existing file, expands or collapses it to a new problem size, and then stores it on a new disk.

III. COPY - this mode copies a given history file from disk storage (working file) to magnetic tape storage (permanent backup file) and vice versa.

-IV. MARGINS - this mode performs those operations which are necessary to cal-culate CFM, DNB, and LOCA margins from an input history file (s).  ;

It also plots the results in the form of a " fly speck" graph.

A-5

l TACO 2-l t

TACO 2 conservatively predicts fuel pin temperature and fuel pin pressure. It 3 <

includes models for fuel densification, fuel swelling, fuel restructuring, gas I

release, cladding creep, and gap closure. i l

TEMP l

' TEMP is a steady state open channel thermal hydraulic code that considers energy mixing between channels and is used to calculate flow distribution among individ-ual channels in an assembly or a cluster of fuel pins. It calculates flow, pressure drop, coolant parameters up the channel, and DNBR. 1 RADAR l

l 1

RADAR performs a thermal analysis of a slow reactor transient such as the loss of a primary pump, computing as a function of time fuel pin and clad surface i

. temperatures, DNBR, and coolant thermodynamic conditions when given pin power and either channel flow or pressure drop as a function of time.

l l TACO TACO conservatively predicts fuel pin temperature and fuel pin pressure. It 4 i

includes models for fuel densification, fuel swelling, fuel restructuring, gas release, cladding creep, and gap closure.

l l

A-6 Rev. 4

1 QUESTIONS AND ANSWERS

m .

)

.(-

\q . - \

~

ATTACHMENT DUKE PCWER COMPANY.

.OCONEE NUCLEAR STATION OCONEE RELOAD DESIGN METHODOLOGY TEC3NICAL REPORT RESPONSES TO NRC QUESTIONS OF-OCIORER 16, 1980 s.

O e

h a

?

bgg

• _ ~ ~ - .r~ w n . -- - _ ,_, __.__

y- . . . - . +=. --%.. . . - - .. - = . - . %. ..

,_s Pegn 1 of 24 m )

(7 Q. 1. Paragraph 3.2.5. Reactivity Coefficients and Daficies.

3 The described procedure for the calculation of the reactivity deficits involves PDQ07 or EPRI-N0DE. However, it is not clear whether for widely different str.tes the reac'tivity difference due I to the spectral component is also included. The same comment ,

applies to the differential boron worth calculation.

A..-l. The lattice code-EPRI-CELL does change cross section libraries an a function of moderator temperature. These cross sections

. are then used in PDQ07 Version 2 for both color set calculations, which lead to input for EPRI-NODE, and for quarter core calculations.

Therefore, the spectral component is included in the calculations of reactivity coefficients and reactivity oeficits.

The effects of soluble boron on the flux spectrum is accounced for in two ways. First the soluble boron concentration inp,uc to the EPRI-CELL fuel depletion is varied from 1200 ppa at 30L to 400 ppa et 6000 MWD /MTU and is held constant at this con-centration for the rest of the depletion. Second, the non-fuel cross sections (eg. control rod guide tubes, reflector, etc.) are generated as a function of soluble boron concentration.

Q. 2. Table 3-1, Shutdown Margin Calculation.

Give a description of the manner in which the " Worth reduction i,, due to burnup of poison material" has been calculated. -

k.'

A. 2. CPM has been used to generate a curve of control rod reactivity reduction (% ao) as a function of fuel burnup at HFP Nominal conditions. This is changed to a % reduction in control rod worth versus burnup. For rodded fuel cycles the control rod bank that is inserted is con'servatively assumed to have been inserted for the whole cycle. For unrodded (feed & bleed) eycles the lead regulating bank is conservatively assumed to have been inserted 20% for the whole cycle. Knowing the worth 4 of the rod groups, the integral red worth curve, and the accu-mulated burnup that. each has seen, the burnup penalty can be calculated.

e g O" 1

p, .

e..

c- -

^

, ,c . , , .

3 Paga 2 of 24

). .

l

.~'

Q. 3. Paragraph 3.2.8, Kinetics Parameters.

Present a more detailed description of the DELAY code. Provide the source of the' code, e.g., Duke Power Company.

l A. 3. The DELAY code has been written by Duke Power Company. The, following four pages have been extracted' froar the DELAY code manual and describe the theory, equations, and data sources for the code.

{

l

~.

9 i

1 l

l l

,v

.s_..

_, _ ,, e ' - " '

~

y Pega 3 of 24 1.0 ~ INTRODUCTION s DELAY is a utility type code which calculaties the six group delayed neutron s's, A's and also reduces them by a group independent effectiveness value.

In addition to this, DELAY calculates the prompt neutron lifetime and then solves the In-hour equation to correlate reactivity insertion and doubling time.

. . - o In;:ut for DELAY is available from two dimensional quarter core FC0 calcu'a-

-ions and EPRI-CELL fuel depletion calculations.

2.0 THEORY f# Calculation 2.' 1.1 s.,1; ands?

i, is defined as the fraction of fission neutrons produced that appear 2s d& layed neutrons of delayed group 1. A is defined as the effective decay constant for the precursors that produce, delayed neutrons in delayed group i.

These quantities are defined by the following equations: - e F F (I} i jg"j9 j9'j9

• ij9"j9 j9'j9 and

'52) f A

gjg C$ )g Af/Cgjg g =,

wnere

-(vb) is the neutron production rate, C denotes the concentration of delayed neutron ' precursors, and the subscripts i, j, g refer to the delayed neutron grouc, fissioning isotope, and incident neutron energy group respectively.

The concentration of delayed neutron precursors is related to the fission rate by F

W i j g C,3g =ks ijgugjgt ,jgt ijg Using ecuation 3, the solution to equations (1) and (2' becemes:

'aa) I li 3 E g

• 29 tjg jg -

'CIIV' =

(ab) S SgEFFECTIVENESS FACTOR m -

i24 p, .;/(E;g?)glijjg) 3

~

s.. 49 G

es e oe s pgu*e==='a -w we4 gen om emme w~-w-- dim m es==mpe w owe- w m pame e- --- -- = _ _ . -

, , wh;re m

- Pags 4 of 24

('6 ) P)g a vg t F

f- $$"jgjg'jg E

i

f. is the fraction ~of the total neutron production rate arising from fissions of isotope j by incident neutrons of group g. Equation (5) is solved using integrated fission rate data from P0Q calculations. Suggested effectiveness factors are 0.961 for Oconee and 0.97 for McGuire.

2.1.2 Delayed Neutron Data .

Tomlinson's values of delayed neutron parameters have been chosen for DELAY.

The values have been reproduced here as Table I for documentation purposes and have been used in DELAY.

~.

2.2 Promet Neutrcn Lifetime The' prompt neutron lifetime, 1* is defined 1 k (7) I* = + 2 YE VE 1 T1 2 T2 wnere

\

(8) :T1 * "UF1 N

1 "E

(9) 5 R1 F2

~

72 = Ei1 N .

(10) V $=eg-10 at 2200 m/sec x220000yc/sec eg710 The parameters and their units are defined in Table 2.

2.3

' Reactivity Calculation The In-hour ecuation has been simplified to include only the as;mototic re-actor period. The form programmed into DELAY is the following:

=p+i 1*

s

, effective

=9 i=1 1+\gi ,

i anere  ? = asymptotic + sactor period

(-,.  : = reactivity e e eo

• Pzg2 5 of 24 3

^

TABLE 1-(

(_ s Delayed Neutr6n Data i From Tomlinson AERE-R-6993  ;

Fast Fission A Relative Absol. Gp.

Group - (see-1)- m S.0. Abundance 5.0. Yield (n/100F) = 5.0. -

'soteoe

.0127 .0003 .038 . 004 .063 .007 U235 1 2 .0317 .0012 .213 . 007 .351 .016 3 .115 .004 .188 . 024 .310 .042 4 .311 .012 .407 . 010 .672 .034 5 1.40 .012 .128 . 012 .211 .022 6- 3.87 .548 .026 . 004 .043 '..007 U238 1 .0132 .0004 .013 . 001 .058 .007 2 .0321 .0009 .137 . 003 .602 .037 3 . .139 .007 .162 . 030 .712 .129 4 .358. .021 .388 . 018 1.708 .120 5 1.41 .099 .225 . 019 .989 .089 6 4.02 .317 .075 . 007 .330 .036

'?u239 1 .0129 .0003 .038 . 004 .024 .003 2 .0311 .0007 .280 . 006 .179 .013 3 .134 .004 .216 . 027 .138 .019 4 .331 .018 .328 . 015 .210 .018 5 1.26. .171 .103 . 013 .066 .010 6 3.21 .378 ,

.035 . 007 .022 .004 Pu240 1. .0129 .0006 .028 . 004 .022 .004 2 .0313 .0007 .273 . 006 .238 .024 3 .135 .016 .192 . 079 .162 .065 4 .333 .046 .350 . 030 .315 .040 5 1.36 .304 .128 . 027 .119 .027 6 - 4.04 1.16 .029 . 009 .024 .007

u242+ 1 .0129 .004 .006 2 .0295 .195 .31 3 .131 .162 .26

.1 .338 .411 .56 5 1.39 .218 .35 5 3.55 .010 .016 .

r. .

n ,

ye- se es e ge,,

'****b=- . . . . . . .. . e

- .-. . . . - - +<~.-%. - - . . - . ...

~ Peg 2 6 of 24 TABLE 2 Parameters for Prompt Neutron Lifetime Calculation r

Parameter Descriotion Units Source l

kg k effective, fast group none PDQ k none 2 effective, thennal group PDQ Ig Removal cross section to thennal group- cm-I r PM' flux weighted edit j fuel only 1

.: 7

• Neutron production cross section in fast cm'1 90Q i group flux weighted edi; j fuel only ,

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Neutron velocity, thennal group cm/sec equation 10 2200m) Thermal cross section at 2200 m/sec for barns Chart of the  !

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I -Q. 4. Paragraph 8.3.2 Start-up Accident s

Give the variation of the total (and its components) reactivity for the lh' l

start-up accident for the first 10 seconds after the accident initiation,

' ' (these would complement Fig. 14-1 and 14-2 of the Oconee FSAR Rev. 16).

I A' . 4. The approach taken in the review of FSAR transient analyses as an inte-gral part of the reload design methodology is discussed in Section 8 of NFS-1001. For each FSAR analysis the main parameters of interest have been identified and-documented in the FSAR. In order to assure that a reload core is in conformance with the assumptions in the analysis, it i' is necessary to determine that the parameters associated with the re-load core. are bounded by the parameters assumed in the FSAR. If this criterion is met, it can be concluded that the existing FSAR analysis remains valid for the reload core.

Question 4 requests additional information for the start-up accident concerning the variation of the components of the reactivity response.

These parameters are an intermediate output of the analysis whose re-sponse is indicated by other documented parameters such as power level, but are not normally included in the analysis documentation. However, the components of the reactivity response are determined by the para-meters which are reviewed and shown to be within the bounds of the FSAR analysis. The reactivity response determined by those parameters re-maina valid until the value of a parameter is no longer bounded for a -

reload core. The safety review methodology of Section 8 assures the identification of all pertinent reload core parameters affecting the reference safety analysis, confirmation of the validity of the re-farence safety analysis for the reload core, and the resolution of any t - non-conservative parameter, i

In order to respond to the question, the variation of the total re-activity and its components were calculated froeit the results presented in FSAR Figures 14-1 and 14-2, utilizing the analysis assumptions specified in the FSAR. The variatit'n of the total reactivity during a startup accident is the sum of three reactivity effects. The with-drawal of the control rod banks adds positive reactivity which causes the neutron power level to increase and raise the average core tenpara-ture. The fncrease in fuel temperature causes a negative reactivity feedback due to the negative Doppler coefficient. The increase in power level increases heat transfer from the fuel to the coolant, re-sulttng in an increase in moderator temperature. This causes a positive reactivity feedback since a positive beginning of cycle moderator coefficient is assumed. The transient response is primarily decsratined by the rate of positive reactivity addition from the with-2 dravel of rods, and the Doppler feedback which slows or terminates the nuclear excursion. The moderator feedback has a smaller effect.

Figures 4-1 and 4-2 show the variation of the reactivity consistent -

with FSAR Figures 14-1 and 14-2 respectively. It should be noted that these figures do not represent the first 10 seconds of the transients, considering that the initial conditions are 10E-9 rated power and 1%

k/k suberitical. Figures 4-1 and 4-2 illustrate the time interval of

- greatest interest during the transient, Figure 4-1 is the same scale'

{'

- as Figure 14-1, and Figure 4-2 is the first one second of the response

~

in Fi.ure 14-2. For both transients the reactivity addition for the '

first 10 seconds following initiation of rod withdrawal vould only cause a reduction in the suberiticality margin.

I, Revised 3/18/81

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Q. 5. Paragraph 8.3.3. Rod Withdrawal Accident at Rated Power Operation

' Give th& variation of the reactivity as in 4. above.

A. 5.

i- The reactivity response of the rod withdrawal accident at rated power simulation performed by B&W and used in the o'riginal FSAR analysis 1 l

is not available. In order to . respond to the question a similar ana-lysis was performed by Duke Power Company using the RETRAN code and matching as accurately as possible the modeling assumptions of the o riginal ' analysis . Figure"5-1, a revised FSAR Fighre 14-9, shows the comparison between the original analy:is (solid lines) and the new analysis (dashed lines). No attempt was made to match the re-sults of the original analysis, the intent being to match the assump-tions and initial conditions. The similarity between the results of the two analyses supports the conclusion that the reactivity response of the new analysis shown in Figure 5-2 is representative of the ori-ginal analysis.

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Prg3 B3 og 86 Q. 6. Paragraph 8.3. Discussion of Individual- Accidents e

s Have the computer codes used in accident analysis (summarized in Ap-pendix A) been updated and revised since the Oconee FSAR was issued?

If so, would the general conclusions of the accident analysis change if the analysis was to be performed with the updated codes? Justify your conclusion.

A. 6. The computer codes summarized in Appendix A of NFS-1001 are primarily .

the ' nuclear, chermal, and thermal-hydraulic analysis' codes intended -

for the reload core design. All the codes necessary for accident analyses are not included in that appendix.

The analysis of the loss of coolant accident was revised since the issuance of the Oconee FSAR using updated codes. BAW-10103 represents this revised analysis. Although many of the other accidents have not been reanalyzed utilizing updated codes, it is believed that the gen-eral codelusions of the existing analyses would not change if the analysis was repeated utilizing state-of-the-art computer codes. This conclusion is based on the premise that the earlier computer codes employed generally conservative modeling compared to the more accurate modeling utilized in current computer codes. Furthermore, the input

. parameters and assumptions employed in establishing the plant models have the dominant influence on accident consequences.

As discussed in the report, the safety analysis review performed dur-ing reload design involves a thorough review of the input data and

- assumptions used in the accident analyses and a comparison to the values generated by the reload design. The goal of the review is to

- verify that the reload design values remain bounded by the accident values and thus confirm that the safety analyses remain valid.

!( c N*

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• • Peg 2 14 og 24

,_ Q. 7. Paragraph 3.3.4. Moderator Dilution Accident ks "AdditiSnal Analysis" is claimed to demonstrate complete protection during refueling operations. Give more information of this analysis.

-E. 7.

The " Additional Analysis" referred to is summarized in FSAR Section 14.1.2.4.2,' the last paragraph on page 14-9. This paragraph is re-produced below.

During refueling or maintenance operations when the reactor closure head has been removed, the sources of dilution water makeup to the letdown storage tank--and therefore to the reactor coolant system--are i locked closed, and the high pressure injection pumps are not operating.

At the beginning of core life when the boron concentration is highest, the reactor is about 9.5 per cent ak/k suberitical with the maximum worth rod stuck out. To demonstrate the ability of the reactor to accept m.oderator dilution during shutdown, the consequences of acci-dentally filling the letdown storage tank with dilution water and starting the high pressure injection pumps have been evaluated. The entire ' water volume from the letdown storage tank could be pumped into the reactor coolant system (assuming only the coolant in the reactor vessel is diluted), and the reactor would still be 4.9 per cent ak/k suberitical.

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, _ , Q. 8. Paragraph 8.3.6. Loss of Coolant Flow It is s&ated that the hot channel power peak augmentation f actors, fuel densification, and rod bow effects are not expected to change for the reloads; however, it is.not stated how this conclusion has been arrived at.

A. 8. Not channel power peak augmentation factors are associated with the 1

mechanical design of the fuel assembly. The mechanical design is not ,

normally modified in the reload design process. The fuel assembly ,

design for Oconee has a history of very few modifications, none signi-ficantly affecting mechanical or nuclear performance. For example, the het channel factors which account for the ef fect of statistical uncertainty in parameters such as enrichment, fuel rod loading, and geometry on the fuel rod heat flux and heat generation rate, remain valid for all fuel manufa:tured wis*ain the specified tolerances in these parameters.

The presently accepted treatment of the fuel densification eff et on minimum DNBR analysis is the use of densified fuel stack length for calculating the heat flux. The original analysis was based on an initial fuel density of 92.5", which produced the maximum stack length reduction compared to the subsequent reload fuel batches consisting of higher density fuel. For each reload, values of the densified heat flux are evaluated in the thermal hydraulic design analysis section of the reload report.

The effect of fuel rod bowing, dependent on the fuel assembly mechan-

. ical design and burnup, is explicitly factorad into the thermal-hydraulic design of the reload core. The reactor protection system setpoints necessary fur DNBR protection are established to provide the necessary margin to account for the ef fect of fuel rod bowing, as discussed in Sections 4.8 and 6.10 of NFS-1001.

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P:33 16 of 24 Q. 9. Paragraph 8.3.9, Steam Line Failure  !

• k" .

It is s6ated in the accident description that continued feedwater flow in the affected steam generator, combined with excessive heat removal and primary cool down the reactor may experience "a return to low power levels." There is not quantification of this power level, its potential consequences, or measures and actions for the return of the reactor to suberttical. Under what conditions is '

there a minimum of rod worth which could have the most adverse effects? '

A. 9.

The answers to these questions may be found in the Oconee FSAR, Chap-ter 14 and Supplement 3. However, a brief response summarizi.ng the FSAR material follows.

A number of cases involving a variety of secondary system behavior during a steam line break are evaluated in the FSAR. Cases involving i

failure 'to isolate the affected steam generator, excessive feedwater addition due to malfunction in the feedwater control function, or of l

the auxiliary feedwater in additon to the continuing feedwater to the affected steam generator predict a return to power (1% TP, 8% TP.

35% FP, respectively) for a brief period of time. In each case, the

. reactor is returned to a subcritical condition by the action of the ECCS (high pressure injection, core flood tank and low pressure in-jection) within 350 seconds. The return to power situations are calculated to occur with the conservative assumption of the minimum tripped rod work associated with the minimum shutdown margin speci-fied in the Technical Specifications and considering the highest- '

s worth rod to be stuck out.

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- , Q. 10. Supplement 2, Figure 4-1 and Paragraph 3.1.1.1.

s Figure 4-1, Supplement 2 appears :o contradict the scacement in paragraph 3.1.1.1 that reads:

"i'ON-fuel cross sections with the excepcion of burnable poison

' assemblies and control rods are also generated using EPRI-CELL.

Cross seccions for burnable poison assemblies and control, rods for use in diffusion theory ~ calculations are generated by macching reaction rates between the diffusion theory code .

- PDQ07 and C.PM (a collision probability code) ."

Give a more decailed description of the procedure for control rod and burnable poison cross section generation and the use of burnable poison cross seccions in PDQ07-HARMONY depletion calculations.

~.

A. 10. While there appears to be a contradiction both statements have ms.t:. The ARMP procedure for generation of burnable poison cross sections was developed from CPM and PDQ07 calculacions.. The procedure however needs only EPRI-CELL .

and PDQ07 calculations to use it. Detailed description of the procedure can be found in the " Advanced Recycle Methodology Program System Documentacion, September 1977." Part I Chapter 6 Section -4.2 cascribes the development of the procedure using CPM and PDQ07 while Section 4.3 describes the procedure using

,, EPRI-CELL and PDQ07.

~ The procedure for developing control rod cross sections is described in Part I Chapter 6 Section 3.4 of the " Advanced Recycle Machodology Program System Documentation, September 1977."

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Pags 18 of 24 I

- Q. 11'. Supplement 2,' Paragraph 3.2, Comparison of ARMP PDQ07 to Cold Criticals. '

~k The'two-dimensional simulation of'the criticals has not been' performed )

at Duke nor with PDQ07, yet it was concluded thac the results would l have been identical sith the PDQ07 results. Justify the above l conclusion. l A. .11. The cold criticals have been simulated with PDQ07. The results have j been published in Part I Chapter 2, Rev.1 of the AILT System l Documentation. This work was performed under EPRI Research Project '

118-1. l l

These benchmark calculations use standard ARMP methodology, standard ,

ARMP codes (EPRI-CELL, NUPUNCHER, PDQ07) and Duke Power also uses these codes and methodology. Duke Power Company has been actively involved in developing in-core fuel management capability since 1969.  ;

Currently in the. Nuclear Fuel Services Section, there are a total of nine employees with a en=n1stive thirty-two (32) man-years of PDQ experience. The level of individual experience ranges from one to j nine years, and includes experience with Combustion Engineering,  ;

Westinghouse, and Babcock & Wilcox core design calculations. There-fore, Duke Power considers that if it had performed these benchmark j) calculations, the results would have been identical.

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( Q. 12. Supp'iement 2, paragraph 3.4, Conclusions.

The conclusions for the calculated results of the peak power are not tenable. There is no reason why the diffusion theory estimation by PDQ07 of the local radial peaking should be more conservative than those. calculated with transport theory codes,

- or.the measured values. This result. must be regarded as .for-

- tuitous. For example (Fig. 3-4), many fuel assembly maxima ,

were underpredicted by PDQ07. Justify the conclusion that.

PDQ07 will always be conservative in peak power predictions and present physical arguments for this justification.

A. 12. In Section 3, PDQ07'_s ability to conservatively predict the assembly local radial is addressed. In Figures 3-2, 3-3, and 3-4,e it was shown that the maximum local radial as calculated by PDQ07 was conservative with respect to the measured or '-

transport theory calcalated values for three completely dif-ferent latrica conditions. Each of these figures show the pin-wise power distributions within a single. fuel assembly.

In Figures 3-2, 3, 4, the eight highese measured' (or EPRI-CPM calculated) pin powers were selected. The means and standard deviations of the (calculated-measured) difference were cab-ulated for all three groups together, and by each group (by Figure) individually.

t. In these samples, the mean was taken as the sample mean with the true standard deviation unknown. Then 95 confidence limits of the true mean were detectined by:

c(.025, n-1)

• S(D)

Dg,g = D + _ f Table 1 d.i. splays the results of this analysis.

Table 1 95: Confidence f.evel Estimates of the C-M Radial Local Means Figure n D S (D) Eu 5 3-2 8 . 0070 .01739 .0215 -

.0075 3-3 3 . 02225 .01268 .0329 .0116 l

3-4 8 . 0105 .008767 .0178 .0032 3-2,3,4 24 . 01325 .01445 .0194 .0071

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-e-Since D>0.0 for all' four sample groups, it is concluded that 4 PDQ07_ vould overpredict the mean radial local of the highest power pins within an assembly., Furthermore, using 95: con- '

• fidence limits estimates, PDQ07 over-predicts the mean radial local'in the lower 2.5% interval (Dt >0.0) for three of the four cases considered.

Besides the observations in Chapter 3 of Supplement 2 _the Oconee V fual assembly employs a uniform lattice with a small interassembly water gap. . A water hole's area is only as large as that of a fuel rod so that thermal flux peaking is minimized. Likewise, even at cold conditions, the nominal water' gap between assemblies is only 12*. of a pin pitch. .

Thermal physics constants are standardly calculated using the Mixed Number , Density (MND) procedure. Thermal absorption and fission constants are products of their respective 2200 m/see cross s. actions and the call average velocity (relative to 2200 m/sec). Thermal diffusion constants are-created in a similar fashion.

"*her-al reaction races in PDQ07 are proportional to the magnitude of.the charmal flux. When excess thermalization occurs, e.g., near a water hole, MND cross. sections conservatively yield higher thermal

{

re. action races than conventional cross sections.

This conservatism of the MND method is shown in Figure l. Here a 1 7 comparison was made of MND and conventional PDQ07 pin powers  !

('. relative to EPRI-CPM. The data source for the MND PDQ07 and EPRI '

CPM assembly simulation was Figure 3 4 of Supplement 2. It was shown that for the.eight maximum pin powers, MND cross sections i

yielded a mean percent difference of .99%; while the conventional cross section PDQ07 had a nonconservative mean of .31%. I The statistics presented in Supplement -2 justify use of a radial

, ONRF of 1.03 for unrodded fuel cycles. We have suggested use of 1.05 which allows approximately two percent conservatism for any local pin peak uncertainties.

The above statistics, physical geometry, and modeling procedures support the. conclusion that no additional uncertainty is needed on the radial local peak. However, a 2 conservatism is built into the 1.05 radial ONRF we propose using, h-p- -

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Q. 13. - Supplemen:; 2, paragraph 4.2, Oconee 7uel Cycle Simulation.

It appears that the EPRI-NODE-P 'almost consistently under-predicts the assembly peak power for cycles 2 and 3. Justify the conclusion in paragraph 4.3 that the EPRI-NODE-P "yieIdad consistently good power distributions..."

A. 13. Conclusions about power distributions are reached in view of the global behavior of EPRI-NODE-P. *he Cycle 3 data was-shown in Sc4 tion 4 of Supplement 2 only for illustrative purposes since the measured data was not considered benchmark quality as the other four cycles. .

The derived total ONRF from chapter 5 was 1.10 for redded '

cycles. Only 6% of the products of the ONRF and calculated peak exceeded the cycle 2 measured peaks. Therefore, based on a 95/95 criterion, the agreement was judged good.

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Prg3 23 of 24

~ Q. 14. Suppiament 2, Figure 4-2 through 4 '127.

b The h?RI-NODE-P calculated power distributions for the first four cycles of operation of Oconee 1 consistently underpre-dicted the relative power in assembly H-8, often by more chan 10%. Is the reason for this anomaly known?

A. 14. Yes. It is current Duk's design practice to perform only one ,

radial power nocmalization at approximately 25 EFPD. The normalization is referenced to a two-dimensional discrete '

pin model PDQ07 power distribution.

The normalization is performed such that there is good radial power agreement in both the central nine (H-8 included) and the peripheral assembly regions. Since only the internal leakage factor, gh, was adjusted for the central nine, agree-ment'of the central nine as a whole was addressed rather U-8 specifically. This method yielded radial differences of 5';

or less early in each cycle for H-8 asshown by Figures 4-4, 4-41, and 4-87. Assembly K-9 in Cycle 3 had a 20% larger radial at BOC than H-8, therefore the central nine normalization

. gave a more accurate agreement with a more limiting assembly.

Cycles 1, 2, and 3 were all rodded cycles, and therefore rod interchanges severely changed the radial power shape. A radial power renormalization to PDQ07 after the rod interchange would have significantly improved radial and peak agreement.

!(.'- The reactors at Oconee will soon all be operated in the unrodded mode and so only the statistics for Cycles 4 and 5 are repre-sentative of future design calculations.

' In Cycia 4, the largest radial power difference for H-8 was 3.3%. In Cycle 5, differences of up to 10% were seen. However,

. B-8 was a low power assembly, and K-9 was the assembly of concern.

Good agreement was shown between assemblies K-9 and also H-9 throughout this cycle.

The only other method of assuring less chan 5 power difference to H-3 would have been to apply a K- multiplier. Such an ad hoc method of normalization is contrary to Duke design practice.

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15..-Supplement 2, paragraph 5.2, Normalitiy Test Results.

, Q.

All, data sets have been used with the assumption of normal distribution, yet some have failed the normality test. Justify the use of the data sets as normal. i A. 15. The D' test for normality is a very rigorous test, and in

,. Table 5-1 of Supplement 2 it was shown that. nine of 16 individual and grouped data sets passed the normality criteria outright - with a 5% level of significance.

Table 1 below presents the percent differences by which the other seven data sets missed the D' percentage point' cutoff values for normality. Of these seven, four data sets were combinations of individual nonnormal datasets which in turn, carried inherent near-normality into the larger sets.

Table I Nearly Normal Data Sets Percent Difference M h N from Cutoff. Figute

^

1 Radial 308 -2.16% 5-11 1,2 Radial 455 -1.75% 5-21 k 1,2,4,5 1

Radial Peak 730 377

-1.56%

.26%

5-23 5-16 3 Peak 211 -3.67% 5-19 1,2 Peak 612 -1.38% 5-24~ l 1,2,4,5 Peak 1027 -1.72% 5-26 i The argusent presented in paragraph 5.2 was that although certain distri-butions did not pass the no6mality test criteria, an ocular inspection of the histograms indicated that, for engineering purposes, normality would be a reasonable approximation of these distributions. This is further supported by Table I above.

It should also be noted that cycle 4, cycle 5, and cycle 4 & 5 radial and peak power data sets passed the normality test. These unrodded I cycles are typical of future Oconee reload designs. l I

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/'

N Re Q. 11 NRC reviewer would like a copy of the work performed under EPRI Research Project 118-1 Re A. 11 < Enclosed is a copy of the EPRI-CELL Criticals Benchmarking

. Portion of Project 118-1. Figures 6-5 and 6-6 of 118-1 correspond to Figures 3-2 and 3-3 in Supplement 2 of NFS-1001. .

h o

e f

0 e

8 I

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~~  : ~~~- --- - - - . -

016/701/@104D Rev. 1

. October, 1978 Part I ,

-- Chapter 2 Advanced Recycle Methodology Proaram p

system DocumentatAon assearch Project 118-1 l

EPRI-CELL Criticals Leachmarking october 30, 1978 Prepared fort Electric Power Research Institute 3412 Millview Avenue Palo Alto, CA 94304 Principal Investigators

(

W. J. Eich (NAI)*

M. L. Kennedy (NAI)

R. D. Moste11er (Science Applications, Inc.)

EPRI Project Manager

3. A. Eclotar Prepared by:

Nuclear Associates International corporation 6003 Essecutive Boulevard md:kville, MD 20852 ~

With revisions by:

Electric Power Research Institute

(-

-

• Current affiliation: Electric Power Research Institute Palo Alto, CA 94303

-- - - - - - - - - ,-.4------em- -,'e we-e enw m----ew------s- --% -qa.m-=w --------m--e-+ -

- - - - r -----

- 5 -

..WBO/AW4/tstwfe TABLE OF CONTENTS

-Section Title Page

sJ 1 07ERvZEW 2-1 2 ANALYTIC PRDCEDJRE 2-2 3 PRINCIPAL 002 RESULTS 2-5 4 PRINCIPAL NO2 IEEULT8 2*I

- l 5 SUPPLEMENTARY INFORMATION AND OTHER RESULTS 2-11 1 6 LARGE-SCALE NOCE-UP RESULTS 2-10 E EFERENCES E-1 M

9 4

e 4

>- 111

" ** **O4O &m-D-- Oweg_ _ . . ,

w- -v-r v--r y-T

<w.-, , .-w -- w- -- - - - -w . ._

u qta;itu w/3

. , . offect, however, adds several tenths of a percent in reactivity to very

(.

(- watery lattices but such are so far from reactor conditions that their ,

analysis lacks most practical relevance.) Finally, the two items of input required for the simulation of grain heterogeneities have been entered for the M02 cases.

Box 3 of Figure 2-1 sipifies the non-depletion EPRI-CELL (GAM / THERMOS) .

run which produces printed output (Box 4) and, by option, the fow PDQ-7 input cards containing the macroscopic few group EPRI-CELL output in Table set Format (Box 5). These cards are part of the input to a "one-dimensional" radial plane FDQ Box 7 (one mesh in the 2-direction with zero current boundaries). Another item of input is the axial 2

buckling, 3g, (Box 6) which has generally been measured. If this buckling was not available in the literature, then it has been accurately estimated from measured critical water heights and reflector savings measured in similar lattices. Since the criticals analyzed in the

-murse of this Program have been restricted to arrays having relatively high moderator heights, dependence of the final value of k,gf (Box 8) i is quite minimal on axial buckling uncertainty. Another item of input to these FDQs is a set of (four fast group) reflector constants which were developed .o match the results of multirroup transpert (F3 )

calculatioas#. These critical analyses could as validly have been conducted with 3 fast groups mutatis mutandis but the effort had been initiated before the installation of the collapsed broad group edits. The Mixed Husher Density model is implicit in the core and reflector thermal group constants used in these FM calculations.

The approach used in analyzing large-scale mock-up experiments differs in some respects from the procedure discussed above. That approach is described in more detail in section 6 of this chapter.

e

- 2-4 I

. _ _ . . . -- ~ _ . . . _ . _ . _ _ - ._ _ . _ _ _ . . _ _ . .

016/A04/01g/3-1 SECTION 6 LARGE-SCALE MOCK-UP REStLTS 6.1 Introduction Figure 4-1 schematically illustrates the calculational process followed in the analysis of five large-scale mock-ups. The procedure .

'is basically similar to the approach described in Section 2 of this chapter for critical lattices. There are three principal differences-between the two methodelogies:

(1) the large-scale sock-ups were analysed for the verification of existing ARMP libraries and procedures rather than to aid in the development of the system (2) the mock 1mps were sufficiently heterogeneous that two-dimen-sional rectangular diffusion theory calculations were required in place of one-dimensional radial calculations (3) separate EPRI-CELL calculations were required fer different parts of lattices--fuel pins, water holes, and burnable 4 poison pins These sock-ups are of special interest because they permit accurate determination of the worth of burnable poison reds (SPR's) . Stretofore, SPR contributions to reactivity in PWR's have been subs med ints core analyses which integrate a number of additional effects, such as control rod worth, Eenon worc , Doppler defect, and soluble boron worth. These mock-ups, however, determine the EPR worth up to 9 I percent t.c by means of straightforward soluble bar: . substitution.

Furtherstore, these particular BPR's have a boron loading which is approximately 70 percent heavier than that for PWR assemlilies of any current design. The agreement achieved with the experimental data therefere uniquely validates the ARMP representation of burnable l poisons and, in addition, further. substantiates the benchmarking t

of EPRI-CELL against critical experiments, which is described in the preceding sections of this chapter.

~ ,

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,' 2-18 l

i

. . . . _ . . _. . . - . _ . ~ , _ . __ . . . . . - - . . _ _

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1 4

- M , Description s.

2 Salance of General 1

- 1 of Peel, unter moles, Impets options, burnable Poleons, lattice Resonaaee Parameters, I V i 1f u v _ y 3 1 EPRI - CELL d 5-EPRI - M EPRI - M Puel Pia Water Bole Burnable Poison Pia (No Depletion)

_ No ( Depletion) if l 6

X-Y One-Group *i Fixed Source PDQ-7 Y V V (s

__lf 7

8_

AxialBuckling,3l m X-Y M ouP-Mock-Up Geometry, etc. PDQ-7 V l 9 l

~

hit l FIGURE 6-1 Flow Chart for Large-scale Mock-Up Analysis 2-19 h_.

_ ,_ . ... . .. .e= - - " *~ ~

M,a 4MMM.

, ._e,-.,..-y

,-,,.-------ww.-,w-+.-g-e

016/A06/019/3 6.2 Descriotion of Experiments 21 The experimental configuration employed in these critical mock-ups is shown in Figure 6-2. The subassembly regions indicated there are fictitious in the sense that there is no structural material in the active region of the geometry and that there is no physical significance to the subassembly boundary. A subassembly region, -

however, does correspond to a 15 x 15 assembly in size and config-uration. The outer buffer region was comprised only of fuel pins and horated moderator, but the contents of the subassembly regions were rearranged from case to case and the soluble boron concentration was adjusted until a multiplication factor of 1.0007 was achieved.

The subassembly configurations for the different cases, or " loads,"

are summarized in Figure 6-3. All locations other than those indicated are fuel cells.

The fuel pins and barnable poison rods are described in Table 6-1. Unlike normal fuel pins, these pins are clad with aluminum. The i .BPR's are unclad cylinders of pyrex glass which have a much higher boren content than normal BPR's. Water holes contain nothing but borated water, and moderator characteristics are summarized in Table 6-1, as well. All measurements were performed at room temperature and pressure, with a moderator height of 145 cm.

For the loadings of interest in this study relative power densities l were obtained for one octant of the central subarnambly. These measure-monts were made at the midplane of the active he- st, using a medium iodide (thallium activated) scintillation counter to count collicated l

fission-product gamma rays from activated fuel rods.

The five loadings considered here allow direct determination of BPR worth by the method of soluble boron substitution. In load 1 .

the subassemblies contain a uniform lattice of fuel pins, and the central region is identical to the buffer. In loads 2 and 3, 17

/.

2-20 l

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"" " FUEL SUBASS**P' Y FUEL SUBASSEMBLY

. .... . . .ee.

3OUNDARY (FOR ILLUSTRATION 07.4LY)

FIGURE 6.2 Geometry for Large-Scale Mock-Up Experiments 4

(n4

%W f* I 2-21 4

... . .. . . . ....e

- - . -  %--,- g.-----w-.,. --.w,-+,.- . . - . . -- ,- . - - - - . , , -+.- -,

p' o s.

I 2 2 4 3 3 .

2 2 2 2 I

2 2 l2 2 l i

3 3

' i i

i l4 '

2  !

2 , ,

4 i -

l .

l LOAD 1 LOAD 2 LOAD 3 Loc 5 LcA: 9 l LOCATION 1 FUEL WATER WATER WATER WATER l LOCATIONS 2 FUEL WATER WATER PCISON ICIs04 -

LOCATIONS 3 FUEL WATER FUEL Fo! son Fust

(- LO AT!cNs 4 FUEL FUEL WATER Fust PorsoN FIGURE 6-3 Subassembly Configurations 2-22

--w -

4 -4 e - -+ +.mm.m.+--- _ . , ..w.,, ., , , , , _ , _ , , , _ , , ,, , ,__

1.

\

TABLE 6-1

. ~ PHYSICAL CEk3UCTERISTICS T PINS AND MODERATOR Fuel Pin l

Enriclument, w/o -

2.459, 1 .002 l Pellet Material 002 Pellet Density, g/cmi3 10.24 1 .04 Pellet Diameter, em 1.0297 1 .0013 Active Fuel Ianyth, em 153.34 1 .89 Clad Material 6061 Aluminum Clad Thicknees, cm

.0813 1 .0025 Clad Outer Diameter, em 1.2060 ' 0015 Fuel Pin Pitch, cm 1.636 1 .003 Surnable Poison mod i Poison Material Pyrex Glass Poison Density, g/cm3 2.244 1 .008 i'

Poison Diameter, em 1.170 1 .001 Boron content, w/o-3.919 1 .002 Poison 3mngth, em 188.0 1 .1 Clad Hone 2 l Moderator Water Density, g/cm3 .9973 Water Temperature, OC 2111 soluble Boron content, ppm Imad 1 1511 1 3 Load 2 1335.5 1 3 Imad 3 1335.5 1 3 -

Imad 8 794 1 3 Imad 9 779 1 3 i  ; .

/

\

~

2-23

.-...-r-------.---.,_-,w--,--y,y,,.-...w- y-- - _m.m .-- v-, - . - _ . ~ .4

016/A08.'019/3-1 s

fuel pins have been removed from each of the subassemblies, leaving borated water in their place. In both loads tle subassemblies are y

octant symmetric, but the water hole locations are slightly different.

In loads 8 and 9 the same fuel pins have been removed as in loads 2 and 3, respectively, but SPR's have been inserted in their place everywhere anoopt in the contral location of each =h== - kly. Connerison of results from loads 2 and 8 and from loads 3 and 9 therefore provides a value for the BPR worth in terms of the change in the soluble boron concentration.

6.3 Analytical ?rocedure Imads 1, 2 and 8 first were simulated with the standard ARMP PWR procedures described in Part I, chapter 6 of this documentation, following the process indicated in Figure 6-1. It it to be emphasized that only the standard procedures were used - more detailed treatments normally employed during benchmarking against critical experiments, such as four energy groups and four mesh spaces per pin cell side in the two-dimensional PDQ calculation, were not needed because of the very low leakage of all these configurations.

This approach produces very good agreement with the experimental data for loads 1 and 8 but not as good for load 2. In the ARMP proce-dure, a four group fine-mesh correction is applied to the multiplication factor when water holes are present (see Part I, Chapter 8, Section 3),

but the discrepancy in the result for load 2 is #2newhat beyond the range of the recommended correction factor for c; rating PWR's. On tne other hand, the water density in these mock-up experiments is abwut 50%

froater than under normal operating conditions, and so the higher i soluble boron density can produce a larger reactivity discrepancy.

, i Because the leakage from these mock-up experiments is quite low, a change in group structure would hava very little effect and so only a fine-mesh correction is needed. A finer mesh spacing, two e

i 2-24 l

{

4 mesh spaces per pin cell side rather than one, was selected and the two-dimensional calculations for loads 1, 2, and S were re-run.

As Figures 6-4, 6-5, and 6-6 illustrate, this change produced a significantly better value for the multiplication factor for load 2 and left the multiplication factors for loads 1 and 8, which were already in good agreement with the experimental data, essentially unchanged. (The convention adopted in these Figures is that water holes are represented by an "X" and that EPR's are represented by a  ;

"+" . ) In load 1, no non-fuel locations are present and so no corree- I tion 'is necessary. In load 8, the EPR parameterisation itself, which preserves the reaction rate predicted by EPRI-CELL by adjusting the PDQ thermal MFD absorption cross section for the BPR, produces a EPR worth which is mesh independent.

Calculations also have been performed for loads 3 and 9, and the results are presented in Figures 6-7 and 6-8, respectively.

Once this mesh change was made in PDQ-7, the AMP system produced excellent agreement with the measured data from all if - W.ds. No additional modification of any of the AMP procedures was needed, and it should be emphasized that this one change was necessitated by the high density of the moderator, relative to normal operating conditions.

EPRI-CELI. therefore has been shown to describe acturately the neutronic behavior of BPR's, even when they are as heavily loaded as the ones in these experiments.

i

2-25

~

1.026 1.025 1.022 1.018 1.013 1.005 .997 .986

k. 1.026 1.025 1.022 1.018 1.013 1.005 .996 .986 1.024 1.021 1.017 1.012 1.004 .996 .985

- 1.024 1.021 1.017 1.012 1.004 .996 .985 1.019 1.015 1.009 1.002 .993 .983 1.019 -1.015 1.009 1.002 .993 .983 1.011 1.005 .998 .989 .979

.979 1.011 1.005 .998 .989 1.000 .992 .984 .973 1.000 .991 .984 .973

.985 .976 .966

.985 .976 .966

- .968 .958

.968 .957 s

' ' .947

.947; Relative Pin Power in Central Subassembly K gyp K ,. y2 AR!P, Star:oAan APJP, FINER Mess EXPERIMENT 1.0007 - -

ARMP STANDARD .9999 1.0182 35.64 ARMP, FINER MESH .9998 1.0179 35.64 -

l FIGURE 6-4 Comparison of Results for Load 1 9

I 2-26

- . . w . . . - . - . . . . . . . .-

1.072 .993 .968 . 992 .993 .948

.951

{- 1.085 1.085 1.010 1.004

.985

.979

.978 .978 .957 .934

. 973 .976 .955 .933 1.033 -1.040 1.002 1.013 1.062 .993 .970 1.075 1.074 1.033 1.028 1.050 .998 .940 1.058 1.076 1.016 1.011 1.0% .986 . 9E 1.080 1.082 1.035 .930 {

1.087 1.087 1.036 .947 1.088 1.089 1.043 .94c 1.056 1.108 1.096 .999 .894 1.096 1.122 1.102 1.003 .939 1 1.069 1.117 1.099 .992 .937 .

l 1.073 .974 .942 1.056 .959 .925

_ . _ _ 1.060 .955 .925

.982 .941 .940

.995 .936 .914 ,

.983 .934 .916 l

.939 .919 '

.917 .905

.918 .908

l. i .890

.036

.Ohh Relative Pin Power in Central Subassembly

)

K EFF K. EXPERIMENT 1 APJ4P, STANDARD EXPERIMENT 1.0007 - - ARi4P, FINER llESH ARiiP, STANDARD 1.0052 1.0240 35.24 ART 1P, FINER IiESH 1.0018 1.0205 35.25 FZacRE 6-5 Comparison of Results for Load 2

,n, 1

w 2-27

- -- -~ - ~ ~ ~ ~ ^ ~ ~ ~ ~

,_ . . . . , _ . _ _ . . . . - - . _ . . . _ _ . . _ . . , , - - , . , _ . - , _ _ , _ , _ _,,,,_,._,___..,,,_,,,__......-___,,_,_,,__,_m,___.~ , . _ .

.986 1.003 1.015 1.005 1.045 1.079 1.102

!(' .998 .990 .996 1.005 1.046 1.087 1.114

.996 .991 .998 1.004 1.046 1.08) 1.120 995 .907 .961 .943 .924 1.027 1.061

.993 1.077 999 .899 .924 .931 .914 1.003 1.078 1.011 .890 .938 .945 .904

.961 1.044 l .864 .864

.851 .855 .945 1.070

.844 .845 .932 1,072

.878 .820 .896 1.007 1.045

.839 .816 .855 .994 1.090

.860 .813 .851 1.005 1.092

.931 1.053 1.093

.926 1.067 1.123

.914 1.070 1.121 1.028 1.095 1.147 1.021 1.115 1.153 1.030 1.116 1.150 1.118- l1.151 !

4 1.155 l1.175 '

1.152 i1.178 1.158

, 1.19c 1.191

~

Relative Pin Power in Central Subassembly K " M EFF EXPERIMENT

- ARMP, STANDARD EXPERIMENT 1.0007 -

ARMP, FINER Mess ARMP, STANDARD .9993 1.0234 35.52 35.54 ARMP, FINER MESH .9998 1.0234 FIGURE 6-6 Comparison of Results for Load 8 w

2-28

~. .

~~~

(' 1.061 .985 .979 .988 .977 .950 .948 1.086 1.006 .979 .973 .976 .955 .933 1.030 1.033 .996 1.004 1.036 .993 .940 1.060 1.076 l'.015 1.009 1.052 .985 .938 1.063 1.055 1.043 .937 1.057 1.052 1.022 .927 l

1.014 1.034 1.080 .997 .931 l 1.021 1.017 I.063 .990 .937 )

.997 1.095 .997 .951 i 1.012 1.059 .986 .933 1.023 .953 1.028 .932

.961 .936

[ .963 .917

.912

.901 R.1. sty. ma %r a c.n=rai se.. ety K gyp K. g EXPERIMENT ARMP, FINER iiESH EXPERIMENT 1.0007 - -

ARMP, FINER IlESH 1.0022 1.0209 35.25 nGURE.6 .7 Camparison of R ults for Load 3 2-29

1.003 1.016 1.011 1.011 1.058 1.072 1.107 '

k- 1.016 1.011 1.016 1.019 1.058 1.097 1.144 1.013 .901 .961 .958 .931 1.036 1.070 1.032 .908 .959 .965 .919 1.014 1.088

.882 .900 .954 1.056 ^

.875 .880 .940 1.076

.932 .968 .925 .993 1.055

~

.941 .949 .897 1.003 1.086

.956 .960 1.018 1.082

.956 .

903 1.010 1.094

.996 1.096

.961 1.010 1.045 1.094 1.055 1.127 1.105 1.153 Asiative Pin Power in Central Subassembly 9 XPERI MNT .

K K=

EFF M-

- - ARltP, FINER IIESH l EXPERIMENT 1.0007 - -

i ARfP, FINER FLESH .9997 1.0235 35.54 FIGURE 6-8 Comparison of Resulta fos Load 9

,x _ .

2-30 ,

C16/,A10/019/3-1 1

. l I

17. v.c. Uotinen, et al., *14ttices of Plutonium-Enriched Dods in Livht unter-Part In Drperimental Results," Nucl. Tech., M 257 L (1972).

18. M. Windsor and R. Goldstein, " Analysis of Lattices Containing Mixed-oxide Fuel in Particulate Form," Trans. Am. Nucl. jgc,., i 3 , 107 (1972).

19. Askew, et al., op cit.

20. Nellens, op cit.

21. M.N. Baldwin and M.E. Stern, " Physics verification Program -

Part III, Task 4, Summary Report," BAW-3647-20 (1971).

9 9

e b

0 6

E-2 i

n.- , , , , - - - . . . - . . -

Re Q. 12 Additional justification is required to support the conclusion that PDQ07 conservatively predicts mexism pin powers.

Re A. 12 Nuclear reactor cores are modeled in two dimensions at Duke

- Power Company using the FDQ07 code. A discrete pin geometry and two neutron energy group Mixed Number Density (MND)

EPRI-CELL physics coastants are used.

In the following figures, hot full power (NFF) PDQ07 and CAgMD individual pia powers are presented from quarter-

- - y assembly calculations. These calculations were performed

- at' beginair.g-of-life with no menos: at this time pin power -

/ peaking is most severe. The enrichments used are typical

' of future reloads at 0 cones. A variety of solubla boron j/ L ~

" concentr dions and burnable poison (BF) weight percents (5 4C)

' were used., Also, water filled control rod guide tubes (CRGT) were used. All assemblies contained an instrument tube (IT).

Table 1 identifies the five cases.

P TABLE 1 gag, U-235 w/o M FFM-goron

- 1 3.03 1.0 w/o 3 4C $00

/ 2 3.03 1.0 w/o 3 4C 1000 3 3.33 .2 w/o 3 4C 1000 4 3.34 CRGT 1000 3 3.04 CRCT 0

In evaluating pin powers, the CAgM0 code solves the transport equation is two dimensions and seven neutron energy groupel.

~

7 FDQ07 used only two energy groupe in evaluating the diffusion

/ e'7 ,

, equation. Therefore, the Duke FDQ07 model was tested not only by a higher order neutronics method, but aise by more neutron energy groups.

, - , in all five cases it is shown that FDQ07 predicts accurately

/ and conservatively each scoemt,1y's maximum pin power. FDQ07 also predicted the sese location of the maximum pin for each e case as CAgMO.

For pin powers equal to or greater than 1.000, pinwise powers  !

usually agree within 12. The CROT cases, however, show FDQ07 j to be up to 2% were conservative. ,

Therefore it is cont.luded from these esaparisons, as well as these in NFs-1001 Supplement 2, that the two group MND PDQ07 accusately sad conservatively predicts the monimum pin power . <

within as assembly over a wide range of moderator and fuel temper.tures, enrichments, soluble boron concentrations, and SF loadings.

1. These CAgM0 ealculations were ren using 69 energy groups in the k, .

niuroregion calculation.

e6 -ee ns e eem e et __ _a en a +. a* e aume ~ = + +eem e r4 g =wooe, e e . - . e - ee t e, &

s

FIGURE 1 QUARTER ASSEMBLY PINWISE POWERS - CASE 1 PDQf7 CASMO CODE E-INT 1.1419 1.1421 ,

1 SA n.91% w/o 1.0A 500 500 PPMB 1.0 1.0 B-4-C w/o IT 1.044 1.009 CASMO 1.050 1.012 PDQ97

(' .

0.992 0.970 yp 0.993 0.973 0.982 0.974 0.952 0.956 0.977 0.966 0.952 0.948 0.981 0.973 0.952 0.946 BP 0.975 0.965 0.95; 0.947 0.086 0.967 0.959 0.971 1.000 SP 0.980 0.964 0.960 0.973 0.994 1.003 0.997 0.981 0.997 1.010 1.024 1.041 1.000 0.992 0.186 0.994 1.007 1.022 1.043 ,

s.. 1.035 1.033 1.031 1.035 1.043 1.054 1.068 1.096 1.036 1.034 1.033 1.038 1.046 1.058 1.076 1.109 l

. se esmuunnegne e mm . , , , , , 4 ,

9

a .

s s

, FIGURE 2 QUARTER ASSDfBLY PINWISE POWERS - CASE 2

. 1 l

CO.3E PDQ97 CASMO K-INF 1.0938 1.0930 IT-99% w/n 9.0A q og

,- PPMB _ 1000 1000

'T B-4-C w/o 1.0 1.0 I

- 1 s

u

---~nm IT '

1.043 1.009 CASHO 1.049 1.012 PDQ97 L

m ~

0.992 0.971 BP

-0.993 0.974 l

0.983' O.975 0.953- 0.958 0.978 0.967 0.354 0.950 0.982 0.975 0.953 0.948 BP 0.976 0.966 0.954 0.949 0.986 0.968 0.960 0.972 1.000 BP~

0.981 0.969 0.961 0.974 0.994 m_

1.003 0.997 0.982 0.997 1.010 1.023 1.039 0.999 0.993 0.987 0.994 1.006 1.020 1.040 m-

. 1.034 1.032 1.030 1.034 1.042 1.052 1.065 1.092 k's-. 1.035 1.033 1.032 1.036 1.045 1.056 1.073 1.105

._ M--w w w * --==e-,,..- - - - - . . =

e g

FIGURE 3

^

QUARTER ASSEMBLY PINWISE POWERS - CASE 3 PDQ97 CASMO CODE K-INF 1.1891 1.1876

~

11-? q q w / n 9MR 9 9R 1000 1000 PPMB _

B-4-C w/o 0.2 0.2 IT 1.037 1.013 CASMO

.,- 1.043 1.023 PDQ97

( ,

0.994 1.002 3p 0.993 1.004 0.984 0.988 0.999 0.995 0.976 0.988 1.002 1.007-0.982 0.986 0.999 1.003 BP 0.972 0.986 1.003 1.014 0.984 0.995 1.002 0.999 0.990 BP 0.975 0.994 1.011 1.001 0.988 l

i 0.988 0.992 1.001 0.994 0.991 0.990 0.9.15 l 0.980 0.991 1.002 0.996 0.986 .

0.983 0.98.1 i

l

1.007 1.008 1.009 1.008 1.008 1.009 1.015 1.037

{'s 1.002 1.005 1.008 1.007 1.004 1.006 1.015 1.042

._ FIGURE 4 QUARTER ASSEMBLY PINWISE POWERS - CASE 4 I

l l

CODE PDQ97 CASMO I K-INF 1.2210 1.2170 l

U-29% w/o 1.1A 1_qR PPMB 1000 1000 B-4-C w/o _

CRGT CRGT l

l l

l IT i

1.024 1.011 CA3MG 1.028 1.029 PDQ97 0.990 1.028 0.989 1.031 UE i

1 0.980 0.996 1.041 1.030 0.971 1.006 1.050 1.063 0.977 0.994 1.042 1.058 gg 0.966 1.004 1.052 1.080 0.978 1.019 1.046 1.028 0.983 0.968 1.017 1.062 1.030 0.985 0.971 0.985 1.020' O.992 0.974 0.962 0.958 0.960 0.988 1.017 0.999 0.968 0.949 0.942

( "

0.980 0.967 0.983 0.975 0.988 0.983 0.984 0.977 0.977 0.966 0.972 0.959 0.974 0.961 0.993 0.984

l FIGURE 5 QUARTER ASSEMBLY PINWISE POWERS - CASE 5 PDQ97 CASMO CODE K-INF 1.3272 1.3267 U-235 w/o 3.08 3.08 .

0 0 PPMB CRGT CRGT B-4-C w/o IT 1.026 1.013 CASMO 1.030 1.031 PDQ97

(~

0.992 1.030 CRGY 0.993 1.032 0.981 0.998 1.043 1.033 0.974 1.009 1.051 1.065 0.978 0.996 1.044 1.060 0 0.969 1.006 1.053 1.080 l

0.979 1.019 1.047 1.028 0.982 U# 1.061 1.029 0.985 l~

0.970 1.016

, 0.971 0.985 1.020 0.992 0.974 0.960 0.956 0.900 0.988 1.016 0.999 0'.968 0.948 0.941

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0.979 0.982 0.987 0.982 0.975 0.969 0.9 J 0.989 0.966 0.974 0.982 0.976 0.964 0.956 0.1. 8 0.979 k_. '^

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DUKE POWER COMPANY OCONEE NUCLEAR STATION Attachment 1 l

Reload Design Methodology Technical Report.

NFS - 1001 ,

Revision 2 hK**.

Title Page 4- 9 (New) 11 4-10 (New) 111 4-11 (New) vi 4-12 (New) v11 4-13 (New) 4-1 4-14 (New)

. 4-2 4-15 (New) ,

4-3 4-16 (New) 4-4 6- 9 4-5 . ' 6-10 4-6 6-11

- 4-7 10-1 4-8 (New) 10-2 (New)

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'D N POWER COMPANT I 'OCONEE NUCLEAR STATION i .. . RELOAD DESIGN ETHODOLOGY i-p.

y e 'c-Technical. Report

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April 23, 1979

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4 FUEL $!ECHANICAL AND THER.'!Al PERFOR.'!ANCE k.

4.1 Introduction Each fuel cycle design requires that thorough fuel mechanical and thermal assessments be performed. A reload design utilizes fuel designs that are bound by previous fuel assembly design analyses.

Occasionally, however, minor differences in the design will occur (such as a change from 94% TD fuel to 95% TD fuel). These changes must then be assessed in regard to the following:

- Cladding creep collapse,

- Cladding strain,

- Cladding stress,

- Fuel pin temperature, and

- Fuel pin pressure ~

Design analyses 'that envelope the opefation of all current fuel designs have been completed by the fuel vendor, and reanalysis

[ is nonsally not required for a new fuel cycle design. Rather, a specific fuel cycle design is compared against the enveloping

(

w design analyses. The assessment must compara cladding and pellet designs against the pellet and cladding geometries and densities, etc., that have been considered in the enveloping design analyses.

Further, the individual radial power histories during the fuel cycle (current and previous batches) must be compared against the generic

( radial power envelopes that have been used in the design analyses.

l In most cases, the design analyses will envelope the fuel cycle design being considered and no reanalysis is required. However, in some cases, either the radial power history or fuel geometry may lie outside of the enveloping design analyses, thus requiring, partial or full resnalysis. The following subsections describe the  ;

types of comparisons that must be made to justify a fuel cycle de-l sign without reanalysis and provides some detail concerning the t .

types of analyses that must be performed if required by either the l

fuel cycle design or by changes in the fuel design itself.

l

[.

7,,j l ( 4-1 Rev. 2 r

1 l

I Table 4-1 presents a summary of all types of fuel thermal and me-chanical performance assessment crita;ia that are used to deter: sine whether a fuel cycle' design, the cladding, and the pellets are enve-loped by existing analyses. As shown in Table 4-1, several :f these analyses require either a comparison against a fuel pin power ver-sua burnup envelope or a comparisen against an assembly radial power versua burnup envelope. Examples of these power history en-velopes are presented in Figures 4-1 and 4-2. These envelopes I

change, as reanalysis is occasionally required, resulting in an expanded power history envelope. Figure 4-3 presents a flow chart for the fuel pin pressure and linear heat rate to melt analyses. l Figure 4-4 f.s a mechanical analysis ' flow diagram.

4.2 Claddint Collapse

• Cladding creepdown under the influence of external (system) pres-sure is a phenomenon that must be evaluated during each reload fuel cycle design to ensure that the most limiting fuel rod does N

.not exceed the cladding collapse exposure limit. Cladding creep '

is a function of neutron flux, cladding temperature, applied stress, cladding thickness, and initial ovality. Acceptability of a fuel cycle design is demonstrated by comparing the power histories of all

, - the fuel assemblies against the generic assembly power history used in existing design analyses, similar to Figure 4-2. The ge-neric power-history must be completely enveloping to avoid.reanaly-sis. Duke Power Company uses its own PDQ edit code to automatically perform this comparison for all fuel assemblies at each depletion step. Changes in pellet or cladding design are also assessed in a similar manner: direct comparison with the fuel rod geometries of

. Table 4-1 and reanalysis, if necessary. Four separste fuel designs have been analyzed to form the generic cladding creep collapse analysis.

The CROV1 computer code calculates ovality changes in the fuel rod cladding due to thermal and irradiation creep and is used to perform the fuel rod creep collapse analysis when required. CROV predicts 4-t Rev. 2

j-..

the conditions necessary for collapse and the resultant time to

.c

'b collapse. Conservative inputs to che CROV cladding collapse analy-sis include the use of minimum cladding wall thickness and maximum

- initial ovality (conservatively assumed to be uniformly oval), all as allowed by manufacturing specifications. Other conservatisms included are minimum prepressurization pressure and zero fission gas release. Internal pin pressure and cladding temperatures, ,

input to CR07, are calculated by TACO 2e using a radial power his-tory similar to that of Figure 4-2, a generic pin to assembly.

local peak, and.a standard axial flux shape.

The conservative fuel rcd geometry and conservative power history are used to predict the number of ETPH required for complete clad-ding collapse. To' demonstrate acceptability, the maximum expected residence time of the cycle is compare'd against the ETPH required for complete collapse.

4.3 Claddina Strain Analysis 7--

t .

The limit on cladding strain is that uniform strain of the clad-ding shot _3 not exceed 1.0"..

A generic strain analysis has been completed by the fuel vendor using TACO 2 to ensure that the sumin criterion above is not ex-ceeded. To determine whether the fuel and fuel cycle designs are enveloped by existing analyses, the criteria of Table 4-1 are re-viewed.

Should reanalysis be required, TACO 2 will be used to determine the fuel rod dimensional changes that occur between the two power levels considered by the conservative design power ramp used in the strain analysis. Then, the maximum tensile (elastic and plastic) strain, which occurs at the cladding I.D., is determined from the .

following equation:

n. /

,p. .

G 4-3 Rev. 2

@ 4- e de ++>*-- MM h-e nemureM = 1 O - em .e., $-

Str n = (Pellet 0.D. ) peak - (Pellet 0lD.) x 100 N

~

(Pellet 0.D.)

-o where (Pellet 0.D.) peak = the maximum pellet 0.D. at the local power peak, and (Pellet 0.D.), a pellet 0.D. prior to tad after a local power ramp.

Pellet 0.D. dimensions are used to calculate cladding strain be-cause the strain itself is caused by pellet thermal expansion.

1 i

The strain analysis is completed in two parts: l i

• - Part 1 employs TACO 2 to determine when pellet contact occurs, i l

A conservative ' fuel rod geometry is used in conjunction with q a, < l.5 axial flux shape, and the core average linear heat rate at 100T, power to characterize gap closure. If contact occurs prior to 30,000 MWD /MTU, then Part 2 will use a ramp from 2 4

KW/FT to a final linear heat rate that is consistent with centerline fuel melt. Whereas, if contact occurs after 30,000 )

!!WD/!!TU, then the ramp's peak linear power is reduced to a s I

lower value that is consistent wirJ2 maximum local powers that -

could occur at burnups greater than 30,000 MWD /MTU.

l Part 2 of the strain analysis is the power ramp calculation, l also performed on TACO 2, which calculates the change in fuel l

- pellet 0.D. that occurs from the change in power level induced by the power ramp. The change in pellet 0.D. is then used to perform the hand calculation of cladding strain using the equa-

, tion above. The cladding and pellet are assumed to be in hard

- contact at the initiation of t'ais ramp.

Thua, there are two major conditions in this scenario tnat make ,

1 it conservative. The first is the extrese power change that is used to simulate the worst case peaking. The second is that the pellet is assumed to be in hard contact at inititation of the ramp. This is a conservative assumption since the power ramp is 4-4 Rev. 2 e-e-- e m --.e- .m .-.e- - - . . . - - - _.e - - . . --ye-- - , , , - . - - - - -m- w-- w ,,w, t- ---w-- ---

---

+--w--'-wr-- er

• s O

. ~

initiated from 2 KV/FT, and pellet / cladding contact is not expected i

to occur at this low linear heat rate. e 4.4 Claddine Stress Analvsis The cladding stress analysis for a new fuel cycle design is siai-larly bounded by a conservative design anal.ysis that uses Section III of the ASME Boiler and Pressure Vessel Code as a guide in classifying the stressas into various categories, assigning appro-priate limits to these categories, and combining these stresses to determine stress intensity. Each new fuel cycle design is assessed against the criteria stated in Table 4-1 to determine if reanalysis is required. The stress analysis is very conservative, and reana-lysis should not be required for standard Mark B reloads. However, an assessment is made for each reload design using the criteria of .

Table 4-1.

g The fuel rod stress analysis considers those stresses that are not relaxed by small meterial deformation, and this anaysis complies with

(

the following design critera:

• All fuel cladding stresses (primary and secondary) shall not exceed minimus unieradiated yield strength for condition I and II occurrences.

• The stress intensity value of the primary membrane stresses in the fuel rod cladding, which are not relieved by small seterial deformation of the cladding, shall not exceed 2/3 of the minimus unirradiated yield strength.

The above criteria keep the primary loads well below material allowable. .

In performing the stress analysis, all the loads were selected to represent the worst case loads and were then combined. This repre-

.e

+

q '

4-5 Rev. 2

s sents a conservative approach since they cannot occur simultaneously.

This insures that the worst case conditions for condition I and II

- events are satisfied. In addition, these input parameters were chosen so that they conservatively envelope all Mk-B design condi-tions.

The primary membrane stresses result from the compressive pressure l loading. Stresses resulting from creep ovalization are addressed in the creep collapse analysis.

Since the internal fuel rod pressure cannot exceed systes pressure for condition I and II occurrences (at coolant temperatures greater '

than 425'*F), the need to address tensile stresses at hot zero power (HZP) conditions and higher is eliminated. The tensile stresses were addressed at cold conditions. The mini == internal fuel rod pressure at HZP conditions is combined with the maximum design systes pressure during a transient to simulate the maximum pres-sure differential across the cladding. The tensile stress analyzed '

occurs at cold (room temperature) conditions at BOL. This is the worst case since the grid loads will be maximum at that point.

The worst case compressive pressure loads were combined with the other worst case loads. These are described below:

The maximum grid loads will occur at BOL. During operation, the contact force will relax with time due to fuel rod creep-down and ovalization as well as grid spring relaxation.

The maximum radial thermal stress will occur at the maximum rated power (power level corresponding to centerline fuel melt).

This stress cannot physically occur at the same time the maximum pressure loading occurs, but is assumed to do so for conserva-tism. (Maximum cladding temperature gradient is combined with minimum pin pressure.)

4-6 Rev. 2

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The ovality bending stresses are calculated at BOL conditions.

A linear stress distribution is assumed. The creep collapse analysis calculates the stress increase with time and ovali-zation.

- Flov induced vibration and differential fuel rod growth stresses are also addressed. .

14.5 Fuel Pin Pressure Analysis The pin pressure analysis is assessed against the design basis analysis criteria and envelopes as indicated in Table 4-1. If any of the parameters of this table are vic h ted, then a reanalysis is performed.

The rod is assumed

~

Pin pressure analysis is performed using TACO 2.

to have a 1.5 symmetric axial flux shape, with a pin power history similar to that presented in Figure 4-1. Incore fuel densification is minimized in this analysis to yield a smaller plenum volume and

[

( a maximum pin pressure.

Figure 4-5 presents the result of an analysis of pin pressure versus burnup, performed by Duke Power Company, using TACO 2. This analy-sis was performed for an extended burnup fuel cycle design, using the pin power history indicated in Figure 4-1, but with lover, more realistic axial flux shapes than the 1.5 cosine shape that is i

used for Reload Design purposes. (Refer to Table 4-2 for the axial flux shapes used in'this extended burnup analysis.) To satisfy acchanical design criteria, pin pressure must be less than system pressure (2200 psia).

4.6 ' Linear Heat Rare Capability .

i l Linear heat rate capability of all fuel rods in a reload batch is l j assessed by comparison against the criteria and envelopes of Table l- . .-

IY 4-7 Rev. 2 l

l

4 4 '1. Any rod whose geometry or power history falls outside of those criteria sust be reanalyzed.

l The Linear Heat Rate to Melt (LERTM) analysis is performed using  !

I TACO 2, assuming maximum incore pellet densification. This analysis i

. assumes a conservative pin power history, similar to that of Figure

~4-1, and a 1.5 cosine axial flux shape. In this analysis, very small axial segments of the fuel rod are spiked to high linear heat rates at each burnup step until centerline fuel melt occurs. The resulting heat rate required to reach centerline fuel melt at each burnup is

, then plotted versus burnup.

Figure 4-6 is a plot of fuel LHRTM versus burnup for s extended burnup fuel cycle design. This TACO 2 analysis, performed by Duke Power Company, represents the pin power history of Figure 4-1, but with more realistic axial flux shapes than the 1.5 cosine that is used for reload fuel cycles. (Refer to Table 4-2 for the axial flux shapes used in this analysis.) The minimum LHRTM occurs early -

in life due to fuel densification, but quickly increases due to the offsetting affects of cladding creepdown, pellet swelfing. and fuel relocation. (No credit is taken for fuel relocation in LHRTM

. analyses).

l l

4-s Rev. 2 l 1

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---.e -- - - - - - . ,.

s v-y x y (n.( x)

TAul.E 4-1 ,

FUEf. HECNANICAL PERFORMANCE ASSESSMENT CRITERIA j

j Assalysis Category .

l.inea r Heat l Item No. Parameter Nevieweds Claddina Collapse 2 Claddlina Strain Cladding Stress Pin Pressure Rate Capability 1 Pin rower History vs Burnup NA NA NA Figure 4-l' Figure 4-1 2 Radial Assembly Power History vs Burnup Figure 4-2 NA NA NA NA 3

• Clad 0.11. Yes Yes Yes Yes Yes 4 Clad I.D. Yes Yes Yes Yes Yes 5 Clad Thickness Yes Yes Yes Yes Yes l

6 Clad Initial Ovality Yes NA NA NA NA 7 Pellet Diameter Yes Yes Yes Yes Yes i

8 Pellet Density Yes Yes Yes Yes Yes 9 Initial Prepressure Yes Yes Yes Yes Yes aN)TES: 1. These criteria are the more significant items reviewed for a reload fuel cycle design, and do not t

Isacinde minor assiseptiossa Liiat are part of the bases.

l

2. The cladding collapse review actually is performed separately for each type of Mark B fuel designi

(loser sets of parameters exist, correspondieng to four separate inel designis).

4 4

1 i e l I 1

l 4-9 Rev. 2

TABLE 4-2 Axial Flux Shapes Used for Ther:nal Analysis (Reference, Figures 4-5, 4-6)

Burnus Ranas Peak of Axial Cosine Shapes 0 - 15,100 1.28 15,100 - 35,000 1.22

> 35,000 1.16 NOTE: Standard reload de' sign analyse.s always employ a 1.5 P/P axial flux shape for pin pressure and LHRTM analysis.

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1' I FlGURE 4-3 illEHMAL ANALYSIS Fl OW tslAGHAM COMPAHE COMPARE CLADDING YES MME H ENVELOPE, P YES E NVE LOPEDP

, ET HISIOHY NO NO CHAHAClEH.

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' HEVISE STANDAHO IS TO 162 t, CHEA1E A NEN ENVELOPE THAT ELuiMS PERFOHM REPOltI LIST. MATCH OUTLIE RS REANALYSIS RESULTS ,

t CYCLE ~

O DESIGN l PAHAMEIEHS F UEL T EMPERATURE ANALYSIS lLHHT Ml COMPAHE

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IULL PIN PHESSul4E ANAL.YSIS 8 8 e

FIGUHE 4 4 MEC44AN4 CAL ANALYSES FIOW DIAGHAM YES HEHUN TAQa2

CALCULATE PIN PHESSUHES DEIEHMINti WONSI CASE POWCH ills 10HY ENVELOPED? . .NO. , _ CALCULATE IIEP O H T AND CHECK CL ADD 4NG CLAD IEMPS HESULIS AND PELLET .

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pressure drop regardless of variations in l'ocal peaking and axial pcwer shape.

In other words, hot channel flow rate will be adjusted by the code to satisfy core-wide pressure drop as local conditions are varied. The axial power shapes input to these parametric hot channel runs are smooth cosine curves whose peak can be specified at various distances up the channel for each series of I axial peaking factors. To obtain the mad == allowable peaking factor for each l '

data point, power input to the channel is increased until the limiting DNBR of 1.4326 is reached. This process determines a maximum allowable total peak for a specified axial peak and its location.

After completion of these parametric analyses, two sets of generic DNBR curves or hwimus Allowable Peaking (MAP) curves are determined. One set is used for DNB operational offset limits, and the second set is used for RPS DNB offset limits. The generic DNBR curves used as operational limits are a conservative I

overlay of 1) the generic DNER curves used for RPS offset limits, and 2) another j I

set of MAP curves which have the reference design DNBR as their basis. Both sets of limits consider the extremities of the P-T core protection envelope (619'F and 1800 psig) as potential core operating conditions. Thus both the f '

f h operational DNB offset limits and the RPS DNR offset limits have considered the l worst case temperature and pressure envelope permitted by the RPS. l l

The last step in the thermal-hydraulic analysis is to take actual power shapes j that gave the lowest DNERs during the maneuvering analysis and input these  :

I irregularly shaped axial curves into the hot channel code to verify conserva- l j

time of the corresponding cosine curves used to develop the generic DNER curves.

f A typical set of generic DNB curves is provided in Figure 6.3.

I 6.8.3 Hot Channel Factors The following additional hot channel factors on local heat flux are utilized in the thermal-hydraulic analyses for developing the generic DNBR curves:

i 1.026 = penalty incurred to increase calculated axial powers since ,

flux depressions at the spacer grids are ignored.

, - 1.024 = the ratio of the total nuclear uncertainty of 1.075 to the h radial nuclear uncertainty of 1.05. .

6-9 Rev. 1 -

..___.7.___

i -

Thus, in determining the generic DNB curves, the normal value of Fq" is

- increased from 1.014 to 1.065. ,

6.9 Transient Analvs'is - Determination of the Flux - Flow Ratio During a_ loss of one or more reactor coolant pumps, the core is prevented frore violating the 1.4326 minimum DNBR criterion by a reactor trip that is initiated by exceeding the allowable reactor power to reactor coolant flow ratio setpoint.

loss of one or more reactor coolant (RC) pumps is also detected by the RC pump sonitors. That is, independently of the power to flow trip, loss of one RC pump will result in an automatic reactor runback. Similarly, loss of two or more RC pumps fron'above 55% full power will cause a reactor trip.

The thermal-hydraulic analysis that is used to set the power to ficw trip set-point for coastdown protection conservatively assumes the loss of two RC pumps. ,

The transient is analyzed using the RADAR code to assure that the 1.4326 mini-

^

4 sus DNBR critarion is not violated at anytime during the loss of one or more RC pumps. -

The steady state thermal-hydraulic analysis provides the starting point for the transient analysis. The power to flow setpoint itself is derived from this analysis by varying the time of reactor trip following the loss of two 1 RC pumps (that is by considering various trip setpoint.s) until the minimum ratio required to maintain the minimum DNBR of 1.4326 bas been determined.

Calculation of the actual (error corrected) power to flow setpoint used at the nuclear station is described in Section 7.3.2.

6.10 Application of the Rod Bow Penaltv i

In existing thermal-hydraulic analyses, a very conservative DNBR penalty is included to account for rod bowing effects. This penalty (11.2%), however,

' has been reduced by 1% because of the flow area (rod pitch) reduction factor already included in the therssi-hydraulic analysis.

2 For some reloads, additional credit can be applied based on the fact that primary coolant flow can be proven to be higher than the 106.5% design flow.

f 6-10 Rev. 2 y- --------.rw---w. - eme pwww e-www'N' N e y a- T-- - ----v'-'-*=-- v' '--r---- ' - - ' - - ' - - ' ' - - ' - - - ~ ='*-'vvF-'V-w

The resulting net penalty is applied directly to the final DNBR margins or by increasing the 1.3 DNBR criteria by the percent penalty, resulting in a DNBR criterion of 1.4326. 2 In future fuel cycle designa, this penalty will be revised to reflect the l penalty true effect of measured rod bowing on minimus DNBR (if any additiona is required). References 12 and 13 document the methods to be used for det Then, 'a determination will be made to mining the true rod bow penalty.

either maintain the current margin which exists or to eliminate part or all of this margin.

i

(.

w l 6-11 Rev. 2

~ ~ - - -- - - - . ._. . . __ __

4 e

10. REFERENCES

1. Program to Determine In-Reactor Performance of B&W Tuels - Cladding Creep Collapse, BAW-1C084, Rev. 2, Babcock & Wilcox, Lynchburg, Virginia, ,2 October, 1978.

2. R. A. Turner, Tuel Densification Report, BAW-10054, Rev. 02, Babcock &

Wilcox, Lynchburg, Virginia, May 1973.

3. Oconee 1 Fuel Densification Report, Revision 1, BAW-1387, Rev. 1, Babcock & ,Wilcox, Lynchburg, Virginia, April 1973.

4 A. J. Eckert, H. W. Wilson, and Y. E. Yoon, Oconee 2 Tuel Densification Report, BAW-1395, Babcock & Wilcox, Lynchburg, Virginia, June 1973.

5. Oconee 3 Fuel' Densification Report, BAW-1399, Babcock & Wilcox, Lynchburg, Virginia, November 1973.

?

6. TAC 02 - Fuel Pin Performance Analysis, BAW-10141, Babcock & Wilcox, 2

.&m Lynchburg, Virginia, August 1979.

7. Correlation of Critical Heat Flux in a Bundle Cooled by Pressurized Water, BAW-10000A, Babcock & Wilcox, Lynchburg, Virginia, June 1976.

8. Oconee Nuclear Station, Unita 1, 2, and 3, Tinal Safety Analysis Report, Docket Nos. 50-269, -270, and -287.

i I 9. Electric Power Research Institute (EPRI), Advanced Recycle Methodology j Program (ARMP) System Documentation, September 1977.

l 10. J. M. Alcorn and R. H. Wilson, CHATA - Core Hydraulics and Thermal i Analysis, BAW-10110, Babcock & Wilcox Lynchburg, Virginia, January 1976. ,

l

11. R. C. Jones, J. R. Biller, and B. M. Dunn, ECCS Analysis of R&W's 177-FA 1 Lowered Loop NSS, BAW-10103A, Rev. 3, Babcock & Wilcox, Lynchburg, O Virginia, July 1977.

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- 3 y . .,-w._ -.___m._,,-- _,. , , _. _ _

12. J. H. Taylor (B&W) to D. B. Vassallo (NRC), Letter. " Determination of the 2 Tuel Rod low DNB Penalty " December 13, 1978. I i

13. D. B. Vassallo (USNRC) to J. I. Taylor (B&W), Letter " Calculation of the 2 Effect of Tuel Rod Bowing on the Critical Heat Flux for Pressurized Water Reactors," June 12, 1978.

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2 3. . , 3 s c.. =.enwe o-Mr. Harold R. Denton, Director Office of Nuclear Reactor Regulation U. S. Nuclear Regulatory Commission Washhston, D. C. 20$55 Attention: Mr. J. F. Stol:, Chief Operating Reactors Branch No. 4 Re: Oconee Nuclear Station Docket Nos. 50-269, -270, -287

Dear Sir:

( In response to your letter dated June 2, 1981 requesting additional informa-l '. tion regarding Tschnical Report NFS 1001, " Reload Design Methodology," please

\. - find the attached responses in Attachment 1 of this submittal. Attachment 2 transmits Revision 4 of Technical Repore. 1001, " Reload Design Methodology."

V ry truly yours, N d,

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Qw- '

William O. Parker, Jr.

JLJ:ses Attachments been w/o Attachment 2 R. M. Gribble K. S. Canady J. E. Smith N. A. Rutherford R. T. Bond R. L. Gill T. B. Owen R. C. Futrell etion File 05-801.01 P. M. Abraham R. R. Clark .

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i DUKE POWER COMPANY l OCONEE NUCLEAR STATION f'-

Response to NRC Letter of June 2, 1981 O

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. s Question 492.1-(Section'6.7)

Provide a more detailed discussion on how the core outlet pressure - reactor outlet temperature curves are' determined.

Response

The core outlet pressure - reactor outlet temperature curves (P-T Safety Limits, Figure 6.2) are determined by varying core outlet pressure and core inlet temperature using CHATA Command Routines 1 and 2 (CR 1/2). Using the -

equivalent-two channel model, described in Section 6.6, core inlet temperature is varied at a constant pressure (one inlet temperature value per CHATA run) until the inlet temperature that yields a hot channel minimum DNBR of 1.4326 at that pressure has been determined. This single limiting combination of reactor coolant pressure and inlet temperature is then used to calculate the corresponding reactor vessel outlet temperature, using a simple reactor vessel heat balance. . ,

This process is repeated over a range of pressures, typically 1800, 1900, 2000,:2100, 2200, and 2300 psia. For each of these pressures, a limiting in-let temperature is determined and a corresponding reactor outlet temperature is calculated. Finally, the resulting P-T Safety Limits are plotted for each allowable combination of operating reactor coolant pumps.

Question 492.2 (Section 6.8.2) 7 The method used to determine the Maximum Allowable Peaking (MAP) factor was to

~

vary the hot channel power until the limiting DNBR was reached. Babcock and v, Wilcox varies the radial peaking factor rather than the power. Demonstrate that the Duke method is an acceptable and equivalent method when compared to the Babcock and Wilcox method.

! Response -

-The Duke method is identical to the Babcock and Wilcox method; further, the operation of CHATA Command Routine 8 prohibits such a variation in this

-procedure. In addition to-this response, it may also be helpful to review i Reference 10 of NFS-1001, specifically page 10-3 and Appendix H, which i describe the CHATA Command Routines.

l

-The MAP curves are generated using CHATA Command Routines 1 and 8 (CR 1/8) and the equivalent two channel model described in Section 6.6. (This two channel model contains an average channel (Command Routine 1) that represents the overall core and a hot channel (Command Routine 8) that is " driven" by the average channel's pressure drop.

Command Routine 8 (CR8) accepts the average channel (CRI) pressure drop as a

'ooundary ecndition, and varies hot channel flow and percent over power in the -

hot channel until the criteria of dP and minimum DNBR are satisfied in the hot channel.

The hot channel in CR8 is a single rod; therefore, for this single rod, over- '

{ ,. power is functionally equivalent to pin peak. Usually the pin power input g

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data field in the CR8 hot channel model is set equal to the core overpower a

fraction (for example 1.12) such that CR8 will output the allowable pin peak directly.'

Question'492.3 (Section 6.8.2)

More information is needed on the generic DNBR curves or MAP curves.

Item 1: Provide a detailed discussion of how the curves are developed.

Ressonse-

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MAP curves are developed using the equivalent two channel model described in Section 6.6 and further described in Duke's response to question 492.2, t above. CHATA Command Routines 1 and 8 are used for MAP analyses.

l Maximum allowable total peaking (KAP) limits are determined both for RPS DNB l offset' limits and for " operational" DNB offset limits. These two types of MAP l curves are described in the response to Item 2 of this question.

\

CHATA Command Routines 1 and 8 are used to vary (in a series of hundreds of separate computer analyses) the axial flux shape peak and axial peak loca-tion. One computer run is required for each combination of axial peak rnd i

axial peak location, for example, an axial peak of 1.7 at 80% of the active fuel length. CHATA CR 1/8 is run such that the average channel model (CRI) calculates and transmits the dP boundary condition to the hot channel (CR8). ,

The hot channel model then determines the maximum rod power (peak) and the  !

hot channel flow that satisfy the dP r.nd DNBR boundary conditions. '

The inputs to CR 1/8 for MAP analyses are core operating conditions (tempera-  !

ture, pressure, power, and average channel flow), average and hot channel geo-metries, hydraulic characteristics, average channel pin power (pin peak = ]

1.0), and a specific axial flux shape to be assessed. To develop a set of MAP curves, axial flux shape is varied from an axial peak of 1.1 to 2.0, with the location of the axial peak varying from the bottom to top of the active fuel length in increments of 10 percent of active fuel length. Output from the hot channel model (CR-8) is the allowable hot channel overpower fraction (fune-tionally equivalent to pin peak for this single rod model). The output pin  ;

peak is then multiplied by the axial peak to yield the maximum allowable total I peak for the flux shape being analyzed.

i Item 2: State the differences between the RPS DNB offset curves and the DNB operational offset curves. l f

Response

Two types of MAP curves are developed. One type is used for the RPS DNB off-set limits. Multiple. subsets of RPS MAP limits are determined, one subset for

- each allowable combination of cperating reactor coolant pumps. The second type of 1 DUP curves is used for DNB operational offset limits.

RPS MAP curves are determined at two separate operating conditions (tempera-ture and pressure) for each allowable combination of operating reactor coolant (7.-

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... pumps, as shown in Table 492.3-1. These two sets of RPS MAP curves (high temperature and low pressure) are overlayed at each allowable pump combina-(- tion, and the conservative overlay is chosen for RPS DNB offset limits. The result of the RPS MAP analysis is three separate families of curves (similar to Figure 6.3), one for four RC pumps operating, one for three pumps, and one for two pumps.

Operational MAP curves are developed for operation with four reactor coolant pumps and are based on the most conservative overlay of the RPS MAP curves and a new set of MAP curves that are determined at the conditions stated in Table 492.3-2.

Item 3: State how the MAP curves which have the reference design DNBR as their basis are obtained.

Response

The MAP curves referred to in this item are the " operational" MAP curves, pre- I viously described. As stated in the Response to Item 2, above, the opera-tional MAP curves are the conservative overlay of 1) the RPS MAP curves at l four pump conditions and 2) MAP curves determined at 102% power and based on the reference DNBR at 102% power. The purpose of this additional overlay at 102% power is to insure that the operational offset limits preserve the ini-tial DNB ratio assumed for DNB limited accidents.

Item 4: State how the extremities of the P-T core protection envelope are

,- considered in developing the DNB offset 1.imits.

Response

The low pressure and high temperature extremities of the variable P-T envelope are used as operating conditions for the RPS MAP analysis by performing the RPS MAP. analyses at the operating conditions stated in Table 492.3-1. The extremities for the four pump RPS MAP analyses carry-through into the opera-tional MAP limits because the operational MAP curves are an overlay of the RPS MAP curves and the 102% power reference DNB condition.

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i' m s m Table 492.3-1 MAP Analysis Input Operating Conditions 4 Pump Operation High Temperature Low Pressure Core' Power Level = 112% Rated Core Power Level = 112% Rated

- T RV outlet:= 619F

• Teore inlet = 544F (typical)

• Peore z-2063 psia (typical)

Peore = 1800 psia MDNBR = 1.4326 MDNBR = 1.4326

.i 3 Pump Operation High Temperature Low Pressure Core Power Level = 87.2% Rated

• Core Power Level = 87.2% Rated T RV outlet = 619F.

• Teore inlet = 542 (typical) 4

• Peore = 2065 PSIA Peore = 1800 psia

_ MDNBR = 1.4326 MDNBR = 1.4326 2 Pump Operation High Temperature Low Pressure Core Power Level = 59.4% Rated Core Power Level = 59.4% Rated T RV outlet = 619F

• Teore inlet = 552F (typcial)

• Peore = 1870 psia Peore = 1800 psia MDNBR = 1.4326 MDNBR = 1.4326

• Peore is that pressure which

• Teore inlet is that temper-results in a MDNBR = 1.4326 ature that results in a MDNBR with a RV outlet temperature at = 1.432o with a pressure at the high temperature setpoint. the low pressure setpoint.

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r t-Table 492.3-2 2,

s.- Operational MAP Input Operating Conditions The following operating conditions describe the additional set of MAP curves that.are developed at 4 pump conditions and are overlayed with the RPS MAP curves to form the operational MAP curves.

Core Power I,evel = 102% Rated

• Teore inlet E 557.2F (includes +2*F error)

Peore = 2135.0 psia (includes -65 psi error)

MDNBR E 2.38 (B&W-2)*

• NOTE: The maximum' allowable total peak resulting from these constraints is the same as the maximum allowable peak that results from an analysis performed at 112% power and a DNBR of 2.05.

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Question 492.4 (Section 7.3.1)

. f' In determining the reactor protection system P-T set points, tLe applicant

( stated that the RCS high pressure trip set point was 2356 psig. In the Technical Specifications for Oconee Units 1, 2, and 3, the high pressure trip is at 2300 psig.- Correct this discrepancy.

Response

The current value for the high pressure trip set point is indeed 2300 psig.

This discrepancy will be corrected in the next revision of the report on the following pages: i

1) Paragraph 2, page 7-0

-2) Table 7-1, page 7-16

3) Figure 7-4, page 7-20 Question 492.5 (Section 7.3.1)

Provide the values that are used to error adjust the P-T set' point curve. How are these numbers obtained?

Response

- The error-adjustment of the P-T set point curve is the same as for previous Oconee reload designs. The error-adjustment for temperature is +1 F. This con-servatively accounts for the maximum temperature error in the instrumentation string. The pressure measurement error is + 30 psi which is added to the mini-l'. num pressure difference between the core outlet and the pressure tap on the hot leg, AP = +30 psi. The net error-adjustment for pressure 14 0 psi.

l Question 492.6 (Section 7.3.2) .

I On page 7-10 reference is made to the flux / flow ratio ratio calculated in Section 6.8. The flux / flow ratio is calculated in Section 6.9. Correct this discrepancy.

Response-

. This editorial correction will be in the next revision of the report.

Question 492.7 (Section 7.3.2)

Provide a reference for. che 6.5 percent full power error-adjustment factor used in setting the RPS power-flow imbalance.

Response +

The 6.5 percent full power error-adjustment is .the same as for previous Oconee reload designs and is discussed in the B&W Topical Report, "RPS Limits and Set-points", BAW-10121, on page 5-13. Although this report is based on the 205 class plant, this factor is the same for the Oco. nee Units (see Technical Specifications 2.3 and 4.1), l s

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L SUPPLEMENT 1 l

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ABSTRACT Measurement and calculational techniques and comparisons of calculated and measured results for core physics parameters are presented in this supple-ment. The measurements are from Oconee Unit 1, Cycles 1-5 and the calcu-lations are performed with EPRI-N0DE-P. Comparisons of calculated and measured parameters show good agreement and confirm the adequacy of present calculational procedures in predicting core physics parameters.

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TABLE OF CONTENTS Page i

1. Introduction S1 1-1

2. Critical Boron Concentrations 2.1 Measurement Technique S1 2-1 2.2 Calculational Technique S1 2-1 2.3 Comparisons of Calculated and Measured Results S1 2-1 2.3.1 Hot Zero Power Comparison S1 2-1 2.3.2 Hot Full Power Comparison S1 2-1 2.4 Summary S1 2-2

3. Control Rod Worths 3.1- Measurement Techniques S1 3-1

-3.2 Calculational Techniques S1 3-1 3.3 Comparisons of Calculated and Measured Results S1 3-2 3.3.1 Comparison in Terms of Reactivity S1 3-2 3.3.2 Comparison in Terms of Soluble Boron Concentration S1 3-2 '

3.4 Summary S1 3-2

4. Ejected Rod Worth:

4.1 Measurement Techniques S1 4-1 1

-4.1.1 Boron Swap S1 4-1 4.1.2 Rod Swap S1 4-1 4.1.3 Rod Drop S1 4-1 4.2 Calculational Techniques S1 4-2 4.2.1 Boron Swap S1 4-2' 4.2.2 Rod Swap -

S1 4-2 4.2.3 Rod Drop S1 4-2 4.3 Comparison of Calculated and Measured Results S1 4-2 4.4 Summary S1 4-3

5. Isothermal Temperature Coefficients 5.1 Measurement Technique S15-1 5.2 Calculational Technique S15-1 5.3 Comparison of Calculated and Measured Results S1 5-1 5.4 Summary S1 5-2

6. References S1 6-1 S1 iii

LIST.0F TABLES

?^82 Critical Boron Concentrations at Hot Zero Power S1 2-3

.2-1 Hot Full Power Critical Baron Concentrations S1 2-4 2-2 Control Rod Worths - In Terms of Reactivity S1 3-3 3-1 Control Rod Worths - In Terms of Soluble Boron Concentration S1 3-4 3-2 4-1 Ejected Rod Worths S1 4-4 5-l' Isothermal Temperature Coefficients S1 5-3 S1 iv

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! LIST OF FIGURES l

2-1 E'1' Oconee 1, Cycle 1 - Boron Letdown Curves S1 2-5 2-2. Oconee 1, Cycle 2 - Boron Letdown Curves S1 2-6 1 2-3 Oconee 1, Cycle 3 - Boron Letdown Curves  !

S1 2-7 2-4 Oconee 1, Cycle 4 - Boron Letdown Curves S1 2-8 l- 2-5 Oconee 1, Cycle 5 - Boron Letdown curves l S1 2-9 l

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_ _ _ _ . . , . . _ . , . . , _ . _ . _ , _ _ , _ - - _ , . ,. ~_ _ _ _ _ . _ _ _. _. _ _ _ - . , . . _ . - . _ _ . _ . _ . _ _ _ _ _ . _ . _ . , . _ _ _ _ . - _ . . . . _ ,

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1. INTRODUCTION This supplem nt presents measurement and calculational techniques and compar-isons of calculated and measured results for so.ne key core physics parameters.

The physics parameters include hot zero power (HZP) and hot full power (HFP) critical boron concentrations, HZP control rod worths and ejected rod worths, and HZP isothermal temperature coefficients.

The measured data is from the Oconee Nuclear Station Unit 1, Cycles 1-5. The measurement techniques discussed are those currently used at the station. The HZP measurements were taken at beginning-of-cycle (BOC) during the Zero Power Physics Testing. The HFP boron concentration measurements were taken at various time steps throughout the cycles.

All' calculations were performed with EPRI-NODE-P. In contrast to predictions, which are calculated before the measurements are taken, the calculations pre-sented here were performed after the measurements were taken. Therefore, the i

plant conditions at the time of the measurements could be closely modeled with EPRI-NODE-P.

Y,%.,, .

The comparisons of calculated and measured results present the means of the dif-ferences between the measured and calculated data and the corresponding standard deviations. The mean and standard deviation are defined as follows:

Mean = x = I[x

> Standard

• 8*

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Deviation a-I where: x g = value for the iS observation n = number of observations.

Si 1-1

2. CRITICAL BORON CONCENTPJLTIONS

'2.1 Measurement Technique

' Critical boron concentrations are measured at HZP and HFP by an acid-base-titration of a rea.ctor coolant system sample.

=The measurement uncertainty for critical boron concentrations is due to (1) error in the titration method and (2) error due to differences between the sample concentration and the core average concentration. Based on con-servative estimates of these errors, the total uncertainty associated with the critical boron concentration measurements is less than 20 ppab.

2.2 Calculational. Technique Critical boron concentrations are calculated at HZP and HFP using EPRI-NODE-P in the boron search mode. Since the search does not yield an exactly critical value, fixed boron runs using EPRI-NODE-P are also made to calculate a boron worth, which is then used to correct the calculated boron concentration to exactly critical.

2.3 Comparison of Calculated and Measured Results 2.3.1 Hot Zero Power Comparison The calculated and measured critical boron concentrations at HZP and BOC for

- Oconee Unit 1, Cycles 1-5 are compared in Table 2-1. Each entry corresponds to a different control rod position. The mean of the differences for these five cycles was found to be 32 ppab with a standard deviation of 24 ppab.

Excluding cycle 3 data, which does not follow the biasing trend of the other cycles, the near of the differences is 43 ppab with a standard deviation of only 14 ppab, d

2.3.2 _ Hot Full Power Comparison The calculated and measured critical boron concentrations at HFP for the reload cycles 2-5 of Oconee Unit 1, are compared in Table 2-2. The mean of S1 2-1

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the differences for these cycles is 46 ppeb with a standard deviation of 19 l

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PPab.

- The data displayed in Table 2-2 can be visualized better by examining plots of soluble boron concentration as a function of burnup. These boron letdown curves are shown in Figures 2-1 through 2-5.

2.4 Summary The comparison between EPRI-NODE-P and measured critical boron concentrations at HZP and HFP indicate EPRI-NODE-P can adequately predict soluble boron con-centrations over both " rods in and "unrodded" (feed-and-bleed) fuel cycles.

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S1 2-2

Table 2-1 OCONEE 1 CYCLES 1-5 CRITICAL BORON CONCENTRATIONS AT HOT ZERO POWER, BOC Critical Boron Conc., PPM Difference, Cycle Calculated Measured PPM 1 1443 1476 33 1441 1480 39 1458 1467 29 1 1403 1440 37 1333 1364 31 1336 1367 31 1247 1262 15 1246 1253 7 2 1257 1301 44 1250 1296 46 1221 1276 55 1143 1194 51 1059 1119 60 1046 1099 53 992 1043 51 965 1013 48 3 1378 1373 -5 1357 1365 8 1320 1331 11 l 1376 1356 -20 L e-- 1356 1350 -6 1321 1324 3 1241 1226 -15 1019 1018 -1 4 1290 1334 44

.1257 1310 53 1003 1057 54 5 1376 1423 47 1350 1405 55 1349 1399 50 1348 1412 64 1045 1083 38 Hean -- --

31.6 Standard Deviation -- -- 24.1 Hean (w/o Cycle 3 Data) -- --

43.1 Standard Deviation -- --

13.7 D'.fference = Measured - Calculated S1 2-3

TABLE 2-2 OCONEE 1, CYCLES 2-5 HOT FULL POWFR CRITICAL BORON CONCENTRATIONS Critical Boron Con _c., PPf* Difference, Calculated Measured PPM Cycle EFPD 30.6 662 697 35

.2 52.2 592 645 53 83.0 446 511 65 103.5 383 428 45 129.0 298 352 54 156.0 217 265 48 184.0 129 188 59 203.8 66 127 61 222.9 1 75 74 3 25.3 721 717 -4 58.5 606 625 19 91.2 499 504 5 121.9 423 482 59 143.9 361 403 42 179.1 243 279 36 203.3 173 212 39 232.6 79 127 48 1

4 28.3 770 797 27 56.6 672 702 30 83.2 591 640 49 103.4 524 583 59 125.1 452 507 55 150.6 373 430 57 174.8 301 354 53 194.5 234 316 82 234.7 113 187 74 5 24.5 910 930 20 44.4 848 877 29 90.4 708 722 14 118.1 604 670 66 145.4 517 564 47 175.3 417 461 44 207.5 316 377 61 239.7 217 266 49 Mean -- --

45.7 Standard Deviation -- --

19.7 Difference = Measured - Calculated S1 2-4

Figura 2-1 e

.. 5 0-1 OCONEE 1. CYCLE 1 a;

BORON LETDOWN CURVES l

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~ Z L i ;- . . : ' - *l : : : ; ; . . . . : . :.. . . . . . . ; . . : . : . . ~ : : - . . . . . . . . .. . . Z : :'; . . = .: . ' . .. . . . - . . . ::: . : n- n; . ; ; , , , : ,

175 5. . =. iirg:Mi:n n Eng n iF iiE.:.:f m; pip: = r n H ' E. .=

- r E n u m.m;n:mg':i=: =- : nm j ~~E yi:jjij!":'ji R j [. . . . . . . _ . . . ,

I

}5}EE MH:i HiI S-- hhiiz ii:i:Shii [m - ?!:i"=i:.@ = 4= ili 2:i5" fil:jiF Ti':. "fi:i;;;;;O iE2EC:

n'i-. ... [Siim3idi!H s!!N!:jS5 u

500 ':- - = -- -- - =-x - :- -- - - :- - -

- -

=s53

. x=::=

= _g. = . . xW;s E

. = . m x :.m=u.

=.L inM t :..h=NA! a:: x _ : . m. a .n .....

O=: n o:nn ::.- E =win  :...x _x._n m: . .m = = ;

@iWi#a=:s x mx:-: m r :: .: mn: = xm o -.no .. . -

- -.:n -. . n=. =m = ; :-. - - - . . :: m n : n =.2.=x.:=:_::::  :=

n.

. 21. ---------2 . . .. _-

'i ~i -

ii:  : - Ei :i -; b

=.!?a. :. . . = .' _m;!: i$5 -

..:.x= == E = n.mn:auma:

_. . : . 2 2 : = = . = .= amru =.a.. . x = x = == - - -

== =::~-n 5E" g;;g == = : 2. = x= = : x= . = == = = = = rn : .rn = m m . . :xx unm,. .nz n= ==@

-= - ~"----- -'===-

- 400 " iE =E..F.i-

_._.. . ===. G_ E.. .F. -". E. =E. ii. f.. Ei n F' P. . . .E.2.E =~: '~ iiiFii := = E"- = ==R= F:iE = ; ' =EEE  ; = : =~ E. == " " = = ~"

E:i.E_

. : . .x : n ,. ..=u-.

1. m==2....:. m.=c.1: n. n2 = == =r.

= =::

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. : 2: =a: = _ A x= : m =2. n n: . m. m. m u m. . _....m.2 .:- x . . = : . . _n t= x =n c = == =

.=x x .. . : r . = --. - : = n= = ==

ru u===:==='

x: = = - = == 322 :== = = == =:= =r r :=m=ur n.cm:m====

-==W = == :== m := == n: = ,_ m = = = = m : = xx = * = = m = 2= ===

= =. m==x: nu.

- - " - -== - - =@= 3Ei i

== = = =m= == - -EE - = - ==: - :=:== = = - =:r - - -

--"m----u- - .= 1 1

2- ==_=.=.m:=..n.=,.x... m = ; =. .x2.._=

= '325 - === ==.= == = . : . =n : = = = -  ;; = _

===:r==========:====ru==_=- = r=n=:.=.

n ; == = = = = T. 2 : .= ==~ =i;;g

= = =:: -- = = == = = = g= = cc r = = m

= = == EE: = =x = x u m = x= ; -

m a m : r -.ncm.-=  :.n zr == == z=

= = = =me n--- r n nn = - r E = =--: == === =: n n 2. : : m o -

r=r=;- - = = - r r = st:- =r--m

====n. == == :2. _2. E.m=.r ==: m == d ==; :"= ==::===-==: 'x=n == :=r.a n= . - m.-= ..: = = x.:. . ==_;=;=-ru.- =m====== = 2.==x===

- = = = r = x :m m .= .c. 2. n . = =r m - - -- -- .- .. n:  : . 71 =- =

200 "5I  : = Eu555EEi

. .an =s: = = .x x m x = x : . c= = u n . . . : r=

E ~~~~ il5EMb FE  :..= ... .- ..'"..h~*--. .2 ---

i~5"E E 5" E

==

n r. - .nqp.-.

4 m

a

. :.r = =x: = cm =--== =;c-=:.m.=: m n a =. :m x = m:y -o n ------ n o ;_  :

.-~.1 . E . ". -. 2 ;;;= g= iE: 1 7.____--__.

m-u= = == = = mm c.umm = x u-33 :_ _ l

u m = r-- . :m = : -

n.m r.

--- - - - -- - - - - - - g -- . - - . - - .

m.:.. 2 a._. _. .. x = = : . .; . u u r : : . . - . -

. u m = .n= = -= = m = x c = m::r n r :::

, . . : - a n- :1 m - .--,:z=====

n :.

. - g .: n ~ - .x; 5 : :- p -

.. . c = ===x . = . r - 7 7. . . i. .n:- ... . . 5:EF  :: - '=-. :. -rm . 5:E a: :_ a.;

. . _ m n m:: r . x x -. . ..-

1M

= = g=. i ==. =. .i : -: ..&. . a -. . = . . . . - . . .

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.: 1 ;g= n- =-m

.:-=w=rmx: = s=i i{:-ix? x = ==  :=.n :::.=4. ... . =. : . =. . i. . =. . p:.. .:if.r :?c6ia.

- -- _m. . m : n m -= : : = = =

NEE 5p%iiFiG tii Ni E -' ' ' ~

.E E E ' E #2E* "'E" e n. E : = x 2 ndisiii:22 3 = =ii n ; = = =. u n: = .u ==:== =n = xun= nra.m m a :.. = = m a_ o- 5

. . ; = = =

. . :n . y :3 m gu .-i.

r = = == r mn =: m n r : - m yc. j na s:: . . - - -

=

. = - nm = == ::: := E= -=-

.. .z = . = n. -E n =. . . =.:_v; - : : =. . =. . =. . . x. . . . m. .. _=. =. . m- =. r . .n x. .:.. . . n. . n . .:... . . . . Ox m ,. . u- : z E..g_i ==::==::r -

BURNUP (EFPD)

S1 2-5 v

u.- - . , _ . _ . . -. .- .. .. __ _ ~ _ -.. . . ..

. _ _ __-- m.___

Figure 2-2 OCONEE 1 CYCLE 2 BORON LETDOWN CURVES 1000- -. .

__ l he

-: a-. --- -

__ _ z.- =,.1 -~_-  :

--~

m

-- ---- LEGEND: measured data '

---+ -

~'~

- calculated data ~7 900" " ----

=;- ::-

NOTES: control rod interchange at 53 EFFD bank 7 withdrawal at 237 EFPD  :.,_

-~ -

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s -

4 700 _.

E

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y=".

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ta_

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g

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m U.C* '=+-~~ .._

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300- +===,='== - ..-- =

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Z

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n-

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-m 2_8 _' $

= .

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= = = + r== ~n . - . ...- ..t-t : :M-n ,

Q. 2 0 100 200 300 p p S1 2-6

Figura 2 :

r ,

~~

OCONEE 1. CYC12 3

. ~ BORON LETDOWN CURVES 1000 LaGRED: amasured data ,

E

- - - - - i 1. a data l i

NerES: rod patah at 100 EfrD l bank 7 withdrawal at 245.8 EFFD i

.i I

E 1

700 I -600 1

l i

i l

-500

, 1 400 ,

I . ,00 i

t l

. 200 100 M

0 100 2M BURNUP(EFPD)

S1 2-7

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$dd) MIME:.D MOS USW10S

FigurO 2-5 OCONEE 1. CYCLE 5 BORON LETDOWN CURVES

- 1000 .

p ,

__% =,n e' . n .__..__ __

c _ . _ __

_ . _ _ ~ _= _ '_ . _, ._  : -- -

. -- ,t-  :

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900 ,-

M_+_ , . --__ _ _ -__

LEGEND: measured data m__

w -_=_ :E

.m -

. ---- calculated data __i_

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= _ -

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600 -+' -M-A - - - -

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S1 2-9

3. CONTROL ROD IdORTHS 3.1 Heasurement Techniques Individual control rod group worths are measured by the boron swap technique.

This technique involves a continuous decrease in boron concentration together with an insertion of the control rods in small, discrete steps. The change in reactivity due to each insertion is determined from reactimeter readings before and after the insertion. The worth of each rod group is the sum of all the reactivity changes for that group.

The worth of the total regulating banks 5-7 can be measured in two ways. The t

first is to add up the worths of the individual banks as determined from the reactimeter readings. The worth in this case is in terms of reactivity. This nearcrement technique is the approved test procedure method. The second way is to measure rod worth in terms of change in soluble boron concentration.

This worth is defined as the difference between the equilibrium critical boron concentration when all the regulating rods are out and the concentration when all the regulating rods are in.

I 3.2 Calculational Techniques

. Individual and total regulating rod group worths in terms of reactivity are calculated by making two EPRI-NODE-P runs. The first is a boron search run with the rod group (s) out. The boron concentration found in this run is then used in a fixed boron run with the rod group (s) in. The difference in reac-tivity between these two runs with constant boron concentration is the rod group (s) worth.

To calculate the total regulating red groups worth in terms of soluble baron concentration, a boron search using EPRI-NODE-P is performed for both the rods-out case and the rods-in case. The resulting boron concentrations are then corrected to exact criticality and the group worths are determined as i the difference between these critical boron concentrations.

S1 3-1 s

. , _ . - . . - . - - - . - . . - _ . _ - - - - , _ _ ~ . _ - - . _ . - - - - . , . - . - . - - _ . . _ _ _ - _ _ _ _ _ _ _ - - - _ _ - _ . - . - - . _ _ _ -

1

! 3.3 Comparison of Calculated and Measured Results e

3.3.1 Comparison in Terms of Reactivity L

A comparison of cal:ulated and measured control rod worths in terms of reac-tivity is shown in Table 3-1. This table compares the worths of the indivi-dual banks 5, 6 and 7 and the total regulating banks 5-7 at HZP and BOC for l Oconee Unit 1, Cycles 1-5. The differences between measured and calculated worths .for all the banks. are small. . For the total banks 5-7, the mean of the

% difference was 7.98% with a standard deviation of 5.36%.

3.3.2 Comparison in Terms of Soluble Baron Concentration The comparison of calculated and measured regulating bank worths in terms of soluble boron concentration for Oconee Unit 1, Cycles 2-5 at NZP and BOC is given in Table 3-2. The agreement between calculated and measured worths is very good. All the differences are less than the measurement uncertainty associated with the boron measurement. The mean of the differences between measured and calculated values for the total banks 5-7 was found to be -3.3 ppeb with a standard deviation of 13.5 ppeb. The mean of the % differences was

-0.98% with a standard deviation of 4.26%.

3.4 Summa ry l l

The comparisons between the calculated and measured control rod worths at HZP indicate that EPRI-NODE-P can adequately predict control rod worths. This has been verified by comparing calculated worths to two independent measurement techniques for Oconee 1, Cycles 1-5.

1 1

S1 3-2

Table 3-1 OCONEE 1 CYCLES 1-5 CONTROL ROD WORTHS AT HOT ZERO POWER, BOC Rod Worth, %Ap Difference Cycle Bank Calculated Measured %Ap  % Difference 0.973 1.11 0.14 12.6 1(* 7 6 0.842 0.97 0.13 13.4 5 0.642 0.68 0.04 5.9 5-7 2.457 2.76 0.30 10.9 1 7 0.967 1.11 0.14 12.6 6 1.007 1.10 0.09 8.2 5 0.606 0.69 0.08 11.6 5-7 2.580 2.90 0.32 11.0 2(b) 7 0.726 0.898 0.172 19.2 6 0.893 0.912 0.019 2.1 5 1.192 1.242 0.050 4.0 5-7 2.811 3.052 0.241 7.9 3 7 1.235 1.228 -0.007 -0.6 6 1.025 1.013 -0.012 -1.2 5 1.332 1.376 0.044 3.2 5-7 3.592 3.617 0.025 0.7 4 7 0.859 0.86 0.00 0.0 6 0.891 0.89 0.00 0.0 5 1.163 1.24 0.08 6.5 5-7 2.913 2.99 0.08 2.7 5 7 1.155 1.36 0.20 14.7 6 1.030 1.20 0.17 14.2 5 1.012 1.19 0.18 15.1 5-7 3.197 3.75 0.55 14.7 Mean 7 NA NA 0.108 9.75 6 NA NA 0.066 6.12 5 NA NA 0.079 7.72 5-7 NA NA , 0.253 7.98 Standard 7 NA NA 0.089 8.15 Deviation 6 NA NA 0.075 6.78 5 NA NA 0.033 4.66 5-7 NA NA 0.188 5.36 (a) Group 8 100% withdrawn; Group 8 inserted in all other cases.

(b) Group 7 not completely withdrawn; Group 5 not completely inserted.

Difference = Nessured - Calculated i J

easured - Calculated

% Difference = Measured x 100 S1 3-3

Table 3-2 OCONEE 1 CYCLES 2-5 CONTROL ROD WORTHS - REGULATING BANKS, NZP, BOC Rod Worth, PPM Difference, Cycle Banks calculated Measured PPM  % Difference 2 5-7 256 263 +7 + 2.7 3 5-7 357 338 -19 - 5.6

-4 5-7 287 277 -10 - 3.6 5 5-7 331 340 +9 + 2.6 MEAN - -

-3.3 -0.98 STANDARD DEVIATION - -

13.5 4.26 l

Difference = Measured - Calculated

% Difference a Measured - Calculated x 100 Measured 1

i l

i e

S1 3-4

4. EJECTED ROD WORTHS Ejected rod worth is defined here as the measured worth of the worst case ejected rod. No error adjustments have been included.

4.1 Measurement Technique boron swap, Ejected rod worths have previously been measured by three methods:

rod swap, and rod drop.

4.1.1 Boron Swap The boron swap method is similar to the method used to measure control rod worth. It involves maintaining criticality by varying the boron concentration The control rod positions to compensate for the ejection of the worst case rod.

are held constant. As was done for control rod worth, the ejected rod worth is determined from the reactimeter readings.

4.1.2 Rod Swap For the rod swap method, criticality is maintained by varying the position of the controlling rod group (usually Group 5) to compensate for the ejection of the worst case rod. In this method, the boron concentration is held constant.

The ejected rod worth is the reactivity due to the controlling rod group move-ment as determined from the differential rod worth measurement.

4.1.3 Rod Drop The rod drop method starts with the ejected rod fully withdrawn. The rod is then tripped into the core and the reactivity is charted by the reactimeter.

The extrapolation to Rod position and boron concentration are held constant.

zero inverse time of a plot of reactivity vs. inverse time yields the ejected rod worth. The uncertainty associated with this method is much greater than that associated with either of the other two methods.

S1 4-1

4.2 Calculational Techniques Ejected rod wortha are calculated using EPRI-NODE-P to simulate either boron sway, rod swap, or rod drop.

4.2.1 Baron Swap For boron swap, a boron search run is first performed to determine the critical boron concentration at the rod group position. The boron concentration as calculated in the EPRI-NODE-P run should be corrected for exact criticality.

Using this corrected boren concentration and a constant rod group position, the reactivity is determined with the worst case rod first in and then out. The ejected rod worth is the difference in reactivity between the worst case rod in and out. '

'4.2.2 Rod Swap For rod sway, the reactivity is determined with the worst case rod first in and then out, keeping the boron concentration constant and the controlling rod group position constant at the least withdraw position (least withdrawn corresponds to ejected rod out). The ejected rod worth is the difference in reactivity between the worst case rod in and out. .

4.2.3 Rod Drop To calculate ejected rod worths by rod drop, EPRI-NODE-P cases with the worst case rod first out and then in are run, with boron concentration and rod group position held constant. The ejected rod worth is the difference in reactivity between these two cases.

4.3 Comparison of Calculated and Measured Results j A comparison of calculated and measured ejected rod worths for Oconee Unit 1, Cycles 1-5 is given in Table 4-1. Overail, the agreement is good. The mean of all the differences between measured and calculated values is -0.0053%Ap with a standard deviation of 0.0602%Ap.

51 4-2 L

L

- 4.41 Sgumary j' The comparison between measured and calculated ejected rod worths indicate EPRI-NODE-P can adequately predict ejected rod worths.

ft l

f b i

e r

e d.

L h

P t

S1 4-3 c

r- . - , . ~ - n ,.- .,,.-n - --- .- -.--- --,-- - --- - , ,

l' TABLE 4-1 OCONEE 1, CYCLES 1-5 EJECTED ROD WORTHS I i

-Nessurement Worth %Ap Difference

. Cycle Location Technique Calculated Measured %Ap 1 H-2 Rod drop 0.287 0.33 0.04

._ L . Rod drop 0.279 0.27 -0.01 l N-12 Rod drop 0.183 0.20 0.02 l K-S Rod drop 0.122 0.15 0.03

=2 L-14 Rod swap 0.580 0.46 -0.12 L-14 Rod swap 0.232 0.20 -0.03 L-14 Rod swap 0.239 0.15 -0.09

-H-8 Baron swap 0.554 0.639 0.085 l

1 3 K Boron swap '0,537 0.57 0.03 l 4 D-12 Rod swap 0.385 0.39 0 N-12 Rod swap 0.363 , 0.25 -0.11 N-4 Rod swap 0.381 0.34 -0.04 D-4 Rod swap 0.420 0.51 0.09 F-8 Rod swap 0.109 0.12 0.01 H-10 Rod swap 0.109 0.07 -0.04 L-8 Rod swap 0.109 0.09 -0.02 H-6 Rod swap 0.111 0.11 0 5 N-12 Boron swap 0.610 0.67 0.06 Mean - - - -

-0.0053 Standard Deviation - - - - 0.0602 Difference Measured - Calculated S1 4-4

= - .

- 5. ISOTHERMAL TEMPERATURE COEFFICIENTS The isothermal temperature coefficient is defined as the change in reactivity per unit change in moderator tesaperature at hot zero power, i.e. ,

a T

• bE AT

-5.1 Measurement Techniques i The isothermal temperature coefficient is measured by executing average moderator temperature changes of +5'F, -10*F and +5'F from initial equilibrium critical con-ditions. After each change, steady state conditions are established and pertinent data are recorded by the reactimeter or reactivity program at the resulting plateaus. The isothermal temperature coefficient is determined as the change in reactivity between plateaus divided by the change in temperature. Since three different temperature ramps are executed, three coefficients can be determined.

e The reported isothermal temperature coefficient is a temperature-weighted average of these three coefficients.

The uncertainty associated with the isothermal temperature coefficient measure-ment is dependent on the value of the coefficient. For acceptable coefficient i values, the measurement uncertainty is less than 10.06 x 10 4 Ak/k/*F.

5.2 Calculational Technique The isothermal temperature coefficient at HZP is calculated using EPRI-NODE-P.

Two cases with the same boron concentration and rod positions but different moderator temperatures are run. The isothermal temperature coefficient is the difference in reactivity between the two cases divided by the difference in the moderator temperatures.

5.3 -Comparison of Calculated and Measured Results A comparison of calculated and measured isothermal temperature coefficients at HZP and BOC for Oconee Unit 1, Cycles 1-5 is presented in Table 5-1. The S1 5-1

I agreement between these calculated and measured coefficients is very good; all values agree within 0.2 x 104

' Ap/'F. The menn of all the differences was found to be only +0.027 x 104 Ap/'F with a standard deviation of 0.089 x 104 Ap/*F.

5.4 Summary The comparison between calculated and measured isothermal temperature coef-ficients indicates that EPRI-NODE-P is an excellent predictor of isothermal temperature coefficients.

i O

1 S1 5-2

Table 5-1 OCONEE 1 CYCLES 1-5 ISOTHERMAL TEMPERATURE COEFFICIENTS AT HOT ZERO POWER, BOC Soluble Boron Temp. Coeff, 10 ~4 Ap/ F Differente Calculated Measured 10 4 Ap/7 Cycle Cone., PPM 1478 +1.09 +1.07 -0.02 1

1428 +0.999 +1.052 +0.053 7355 +0.736 +0.77 +0.03 1263 +0.509 +0.51 0.00 1253 +0.474 +0.5315 +0.058 952 -0.301 -0.391 -0.09 2 1295 -0.035 +0.15 +0.19 1194 -0.341 -0.17 +0.17 1013 -0.120 -0.14 -0.02 3- 1330 -0.0078 +0.046 +0.054 1018 -0.708 -0.726 +0.018 4 1324 +0.121 +0.090 -0.031 1057 -0.542 -0.701 -0.159 5 1405 +0.122 +0.20 +0.08 1083 -0.719 -0.65 +0.07 Mean -- -- -- +0.027 Standard Deviation -- -- -- -0.089 Difference = Measured - Calculated I

S1 5-3 1

.~

d

6. REFERENCES

1. Comparison of Core Physics Calculations with Measurements, BAW-10120, Babcock & Wilcox, Lynchburg, Virginia, June 1978.

2. Duke Power Company, Oconee Nuclear Station Unit 1, Startup Physics Test Program, February 1980.

3. Duke Power Company, Oconee Nuclear Station Unit 1, Cycle 6, Zero Power Physics Test, TT/1/A/711/06.

I e

S1 6-1

SUPPLEMENT 2

ABSTRACT In-Supplement 2 describes Duke Power Company's benchmarking of EPRI-NODE-P.

c1cled in this supplement are measured Assembly Powers, Local Radial Comparisons, EPRI-N0DE-P Calculations, Statistical Analyses, and Fitting Procedures.

S2-ii

l TABLE OF CONTENTS Page 1

1. Introduction and Summary 1.1 Introduction S2 1-1

'1.2' Structure of Supplement 1 S2 1-1 1.3 Summary S2 1-1

2. Measurement Data 2.1: Measured Assembly Power Data S2 2-1 2.2 Measurement System Description S22-1 2.3 Measured Powers: Cycle 1 and Cycle 2 - 30.6 EFPD S2 2-2 2.4- Measured Powers Beyond 30.6 EFPD - Cycle 2 S2 2-3

-3. Local Radial Analysis 3.1 Local Radial Factor Analysis S2 3-1 3.2 . Comparisons of ARMP PDQ07 to Cold Criticals S2 3-1 3.3 ARMP Local Radials Compared to Simulated Hot Full Power Conditions S2 3-2 3.4 Conclusions S2 3-3

4. EPRI-NOPE-P Power Distribution Comparisons 4.1 'EPRI-NODE-P Model S2 4-1

. 4.2 Oconee Fuel Cycle Simulations S2 4-2

'4. 3 - Conclusions S2 4-5

5. Statistical Analysis 5.1 0bserved Nuclear Reliability Factor (ONRF) Derivation S2 5-1 5.2 Normality Test Results S2 5-4 l 5.3 ONRFs For EPRI-NODE-P S2 5-5 5.4a Quantitative Comparisons of EPRI-NODE-P to Measurements S2 5-6 5.4b Relative Percent Differences S2 5-7

-5.5 Conclusions S2 5-8

6. References S2 6-1 Appendix A - Power Peak Methodology 1

A.1 Objective S2 A-1 l A.2 Method of Series Evaluation S2 A-1 !

l S2-iii 1 1

LIST OF TABLES Page B&W Cold Criticals Used for Local Radial S2 3-4 3-l' Uncertainty Analysis Oconee-1 Core Conditions S2 4-6 4-1 4-2 Oconee Unit 1 State Points Cycle 1 S2 4-7 4-3 Oconee Unit 1 State Points Cycle 2 S2 4-8

~4-4 Oconee Unit 1 State Points Cycle 3 S2 4-9 4-5 Oconee Unit 1 State Points Cycle 4 ,

S2 4-10

'4-6 Oconee Unit 1 State Points Cycle 5 S2 4-11 5-1 Difference Distribution Normali^y Tests (C, M 1 1.0) S2 5-10 5-2 EPRI-NODE-P ONRF Values S2 5-11 5 Maximum State Point Data Used in ONRF Calculation S2 5-12 5-4 Difference Means and Standard Deviations for Peaks S2 5-13 5-5 -Difference Means and Standard Deviations for Radials S2 5-14 5-6 Percent Difference Means - (C, M 1 1.0) - Peaks S2 5-15 5-7 Percent Difference Means - (C, M 1 1.0) - Radials S2 5-16 i

i i

S2-iv 1

r -- ,

LIST OF FIGURES

- gP 2-1 Oconee Instrument String Locations S2 2-4 2-2 Eighth Core Instrument String Map S2 2-5 3-l' Uranium Critical Experiment Geometry S2 3-5 3-2 Relative Pin Powers Critical - ARMP PDQ07 S2 3-6 3-3 -Relative Pin Powers Critical (LBP) - PDQ07 S2 3-7 3-4 Hot Full Power Simulation - PDQ07 Pin Powers S2 3-8 4-1 _ Outline of PWR Data Flow S2 4-12 4-2 Oconee 1 Cycle 1 Control Rod Configuration S2 4-13 4-3 Oconee 1 Cycle 1-Quarter Core Loading Diagram S2 4-14 4-4 Oconee 1 Cycle 1 Calculated - Measured Assembly S2 4-15 (to) 4-19 Radial Powers (to) S2 4-30 4-20' Oconee 1 Cycle 1 Calculated vs. Measured S2 4-31 (to) 4-37 Assembly Peak Powers (to) S2 4-48 4-38 Oconee 1 Cycle 2 Control Rod Configuration S2 4-49 4-39 Oconee 1 Cycle 2 Quarter Core Leading Diagram S2 4-50 4-40 Oconee 1 Cycle 2 Quarter Core Shuffle Pattern S2 4-51 4-41 Oconee 1 Cycle 2 Calculated vs. Measured S2 4-52

-(to) 4-50 Assembly Radial Powers (to) S2 4-61 4-51 Oconee 1 Cycle 2 Calculated vs. Measured S2 4-62

-(to) 4-60 Assembly Peak Powers (to) S2 4-71 4-61 Oconee 1 Cycle 3 Control Rod Configuration S2 4-72 4-62 Oconee 1 Cycle 3 Quarter Core Loading Diagram S2 4-73 4-63 Oconee 1 Cycle 3 Quarter Core Shuffle Pattern S2 4-74 4-64 Oconee 1 Cycle 3 Calculated vs. Measured S2 4-75 (to) 4-73 Assembly Radial Powers (to) S2 4-84 4-74 Oconee 1 Cycle 3 Calculated vs. Measured S2 4-85 (to) 4-83 Assembly Peak Powers (to) S2 4 i S2-v

LIST OF FIGURES P_ ale 4-84 Oconee 1 Cycle 4 Control Rod Configuration S2 4-?5 ,

4-85 Oconee 1 Cycle 4 Quarter Core loading Pattern S2 4-96 '

4-86 Oconee 1 Cycle 4 Quarter Core Shuffle Pattern S2 4-97 4-87 Oconee 1 Cycle 4 Calculated vs. Measured S2 4-98 (to) 4-95 Assembly Radial Powers (to) S2 4-106 4-96 Oconee 1 Cycle 4 Calculated vs. Measured S2 4-107 (to) 4-104 .

Assembly Peak Powers (to) S2 4-115 4-105 Oconee 1 Cycle 5 Control Rod Configuration S2 4-116 4-106 Oconee 1 Cycle 5 Quarter Core Loading Diagram S2 4-117 4-107 Oconee 1 Cycle 5 Quarter Core Shuffle Pattern S2 4-318 4-108 Oconee 1 Cycle 5 Calculated vs. Measured S2 4-119 (to) 4-117 Assembly Radial Powers (to) S2 4-128 4-118 Oconee 1 Cycle 5 Calculated vs. Measured S2 4-129 (to) 4-127 Assembly, Radial Powers (to) S2 4-138 5-1 Oconee 1 Cycles 1,2,3,4,5 Calculated Minus Measured S2 5-17 (to) 5-5 Radial Power Difference Histograms (All C, M) (to) S2 5-21 5-6 Oconee 1 Cycles 1,2,3,4,5 Calculated Minus Measured S2 5-22 (to) 5-10 Peak Power Difference Histograms (All C, M) (to) S2 5-26 5-11 Oconee 1 Cycles 1,2,3,4,5 Calculated Minus Measured S2 5-27 (to) 5-15 Radial Power Differences (C, M 1 1.0) (to) S2 5-31 5-16 Oconee 1 Cycles 1,2,3,4,5 Calculated Minus Measured S2 5-32

( (to) 5-20 Peak Power Difference Histograms (C, M 1 1.0) (to) S2 5-36 i 5-21 Radial Power Difference Histogram Cycles 1,2 S2 5-37 (C, M 1 1.0)

I S2-vi l

LIST OF FIGURES Page 5-22 Radial Power Difference Histogram Cycle 4,5 S2 5-38 (C, M 1 1.0) 5-23 Radial Power Difference Histogram Cycle 1,2,4,5 S2 5-39 (C, M 1 1.0) -

5-24 Peak Power Difference Histogram Cycles 1,2 S2 5-40

)

(C, M 1 1.0) 5-25 Peak Power Difference Histograr Cycles 4,5 S2 5-41 l (C, M 1 1.0) 5-26 Peak Power Difference Histogram Cycles 1,2,4,5 S2 5-42

. (C, M 1 1.0) 5-27 Radial Power Difference Histograms - State Point S2 5-43 Maximum Calculated Minus Maximum Measured 5-28 Peak Power Difference Histograms - State Point S2 5-44 l Maximum Calculated Minus Maximum Heasured i A-1 Assembly Axial Power Plots Oconee 1 Cycle 1 S2 A-3 (to) A-13 (to) S2 A-15 A-14 Assembly Axial Power Plots Oconee 1 Cycle 2 S2 A-16 (to) A-27 (to) S2 A-29 A-28. Assembly Axial Power Plots Oconee 1 Cycle.3 S2 A-30.

(to) A-43 (to) S2 A-45 A-44 Assembly Axial Power Plots Oconee 1 Cycle 4 S2 A-46 (to) A-58 (to) S2 A-60 A-59 Assembly Axial Power Plots Oconee 1 Cycle 5 S2 A-61 (to) A-70 (to) S2 A-72 S2-vii

-e., >

1. INTRODUCTION AND

SUMMARY

1.1 Introduction The current nuclear code employed by Duke Power Company for three dimensional assembly power calculations is EPRI-NODE-P. This code has been benchmarked against two rodded cycles and two unrodded cycles of operation of the Oconee Unit I reactor.

This work encompassed: derivation of measured power distributions for cycles 1 tnrough 5, simulations of these 5 cycles using EPRI-N0DE-P, development of fitting procedures for the assembly axial power, and development of a statistical basis fer estimating the calculational accuracy of EPRI-NODE-P.

1.2 Structure Of Supplement 2 This supplement is structured to provide smooth transitions between various major topics. Secti,on 2 will describe the assembly measured power data base. Section 3 will compare calculated and measured local radial factors.

Section 4 describes the EPRI-N0DE-P Oconee simulations and presents comparisons of assembly radial and peak powers. Section 5 quantitatively compares calculated and measured powers from Section 4. EPRI-NODE-P reliability factors are calculated based on observed differences. Appendix A outlines the fitting procedure for the assembly axial power.

i l' . 3 Summary Local radial factors predicted by PDQ07 were examined and found consistently censervative relative to either measurements or more elegant neutronics codes.

A large data base consisting of five cycles of Oconee-1 measured and EPRI-NODE-P calculated fuel assembly powers was assembled.

Calculated and measured powers were statistically combined to derive 95/95 Observed Nuclear Reliability Factors (ONRF) for EPRI-NODE-P. Using a variety of fuel cycle combinations, ONRF's were calculated for both assembly peak and radial powers.

S2 1-1

ONRF's of 1.05 for the radial and 1.075 for the peak were found to be conservative for unrodded (feed and bleed) core operation.

\

i 1

1

(

D S2 1-2

-2. MEASUREMENT DATA 2.1 Measured Assembly Power Data The measured power data base used in this supplement comprises assembly power data from Oconee Unit I for cycles 1 to 5. All assembly power data are direct-

'ly traceable to raw signals received from the incore detector system.

i

-The measured assembly power data for Cycle 1 and at'30.6 EFPD in Cycle 2 were 1

'those used by Babcock & Wilcox in their calculational nuclear uncertainty anal sis (5) The remaining data was generated by B&W using similar methods to reprocess the raw detector signals (6,7) ,

2.2 Measurement System Description The.incore detectors at Oconee consist of pure Rhodium emitters which respond

'to the incident neutron flux. With each neutron absorption, a beta particle ($)

is released according to the reaction:

0 3Rh + I n+ Rh + . $ + energy (2-1)

The current measured from the emitter to ground is proportional to the net emitter loss. After the emitter current has stabilized (M i minutes), the current is then proportional to the local neutron flux (in the neighborhood of the eight pins surrounding the emitter) This emitter is called a S_ elf

.P_owered Neutron Detector (SPND). The SPND's are physically located inside the F_uel Assembly (FA) Instrument T_ube (IT). The IT is situated in the center of the FA.

i

[ SPND signal magnitude is of the order of nanoamps. The reactor's on-line com-F puter (OLC) performs a signal to power conversion approximately at ten minute intervals, logging signal data, core power, power distribution, and assorted l, other. data pertinent to core operation.

I S2 2-1

SPND's are distributed in fixed positions to provide an adequate three di-mensional assembly power measurement. In'each instrumented FA, seven SPND's are located equidistantly along a " string". Each string also has a thermocouple '

and an insulated leadwire which is used to correct for gamma induced signals

.in the seven detector leads.

In Oconee, 52 of 177 FA contain detector strings. The locations of instru- I mented FA form a spiral as shown in Figure 2-1. Eight strings are located symmetrically in the interior; another 8 symmetrically farther out toward the periphery. The two sets of eight strings are used to supply corewise quadrant tilt information. An eighth core map of 29 FA three dimensional powers can be i

obtained; and using. tilt data, full core maps of 177 radial FA powers and 1239 sagaent FA powers can be obtained.

l The measured powers used in this supplement will be collapsed from 52x7 (full {

core) to 29x7 (eighth core) at each reactor state point. As shown on Figure 2-2, 11 of the 29 eighth core locations have symmetrically located detector strings. l Relative powers (radial and seven-level) for each symmetric pair or symmetric octet were averaged to obtain the best estimate of the "true" measured power.

Power measurements were taken at approximately equilibrium Xenon conditions.

Reactor power was also as close to 100% full power as practicable.

2.3 Measured Powers: Cycle 1 and Cycle 2 - 30.6 EFPD This data base consisted of the sama PDO data as Babcock and Wilcox used in its calculational nuclear reliability analysis. The raw emitter signals were reprocessed off-line at B&W using software which represented then current state-of-the-art experience obtained from the operation of Cycle 1 of Oconee 1.

The core state-points where these measurements were taken are shown in Tables 4-7 and 4-8. Seventeen state points were used in Cycle 1.

S2 2-2

- 2.4 Measured Powers Beyond 30.6 EFPD - Cycle 2 i

l In order to complete the measured assembly power data base using a consistent procedure, raw detector signal data were processed by B&W for the remainder of the cycles. As in Cycle 1, all data were taken at equilibrium Xenon conditions so as to mitigate any. transient effects. Tables 4-8, 4-9, 4-20, and 4-11 show the selected reactor state-points.

No explicit estimate is made here of the measurement system accuracy during Cycles 1 to 5, since this component is conservatively treated in Section 5.

For failed detectors substitute signals are derived through a spline fitting procedure, provided that operating SPND's are adjacent on either side to the failed SPND. If two or more adjacent SPND's fail on a string, substitute signals a're derived from either symmetric or adjacent locations. The same procedure is employed for entire string substitution. Since most of the failed strings in Cycle 3 had operable symmetric counterparts, it was judged that the radial power measurements were reasonable. However, the number of individual detector failures would disqualify the Cycle 3 power distributions l from being used in a reliability factor program.

Cycle 3 radial powers and peak powers will be shown in this report for com-parison only. No Cycle 3 data will be used in Section 5 where EPRI-NODE-P reliability factors will be derived.

1 l.

I i-l l

S2 2-3 L

l I

Figure 2-1 Oconee Instrument String Locations I

A -

i 3 31 30 C

32+ 29 28+ 52

• l D 33 27 51 E 34 7* 5* 26 . ___

F 35+ 6 4 24 .23+

36 9* 8 3 25* 22 G

g 37 10 1 2 21 _

I 11* I 19* 20 K ..

38 391 12 18 50+

L 40 13* 16* 17 49 g _..

41 14 15 ~

3 42 43+ 47+ 48 0 -

p 44 45 46 )

g .-._

. g 2

1 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

• - Inner Eight Symmetric Detector Strings

+ - Outer Eight Symmetric Setector Strings S2 2-4

Figure 2-2 Oconee Fuel Assembly Map Detector String Number Assignment For Eighth-Co're Averaging 4

8 9 10 11 12 13 14 15 H 1 2 4 10 14 21 30 37 0 15 22 31 K 3 1 13 1 45 8 20 29 36 19 25 24 23 28 32 38 17 46 L 12 35 39 43 18 27 40 47 50 t

33 40 M 26 49 34 42 N 41 48 51 O 52 l l

l' S2 2-5 1

L

i l

3. LOCAL RADIAL ANALYSIS 3.1 Local Radial Reliability Factor Analysis

~

'In this' document, the local radial is defined as the ratio of the maximum pin power to the assembly radial power (assembly average X-Y power). Commercial

reactors, such as Oconee, are not instrumented to measure local radial factors.

The only pinwise power data available to Duke Power are the room temperature B&W critical experimentsa where pinwise powers were measured.

The calculational tools available for predicting the local radial factor are PDQ07, EPRI-CPM8 , and CASM0 10 Both EPRI-CPM and CASMO are two dimensional transport theory codes. EPRI-CPM uses collision probabilities to solve the 4 Boltzman' transport equation, while CASMO employs transport probabilities.

Both transport theory codes have been benchmarked-against the KRITZ hot criticals.11 3.2 Comparisons of ARMP PDQ07 To Cold Criticals In order to establish the local radial predictive accuracy of the ARMP PDQ07

. code, a series of calculations 12 were perforced and compared to measured pin powers in the B&W uranium criticals.

All criticals consisted of a central region of'nine 15 x 15 lattices sur-

! rounded by a buffer region of fuel rods. This buffer region, in turn was l-l reflected radially by borated water. Criticality was achieved by raising the moderator height. -

The overall measurement error has been estimated to be 1.11% in the deter-mination of relative pin powers. Additional descriptivt information of the l B&W uranium fuel criticals can be found in Reference 8.

I

(. -In the two dimensional simulation of the criticals, the ARMP procedure was used in generating PDQ07 cross sections and usage of the PDQ07 code. Although the simulations of these criticals were not performed at Duke Power, the methods.u~ sed would have been similar, resulting in identical PDQ07 results.

I S2 3-1 l

t l

. - _ _ _ . _ _ .- ,- ~ _ . . _ - _ _ . _ . _

A typical B&W Mark B 15x15 fuel assembly (FA) was mocked by two such criticals described in Table 3-1. Load number 2 corresponds to an unrodded FA. While load number 8 corresponds to a FA with lumped burnable poison (LBP). Figure 3-1 depicts the two loading patterns. The results are shown in Figures 3-2 and 3-3.

In load 2, shown in Figure 3-2, the PDQ07 predicted local radial was 1.3% higher than the measured local radial. Likewise, the PDQ07 local radial in load 8 fFigure 3-3) was 3.3% higher.

To assess the overall local radial predictive behavior, the ten maximum pins were compared in each critical. In load number 2, the (C-M) mean was .0091 with a standard deviation of .01775. In load number 8, the mean was .0239 with a standard deviation of .01356. When the (C-M) differences from loads 2 rad 8 are lumped together, the 20 data points give a mean of .0165 and a standard deviation of .01715.

The positive means of the C-M differences and small standard deviations shown above indicate that the ARMP PDQ07 discrete pin model yields radial local factors which are conservative.

3.3 ARMP PDQ07 Local Radials Compared to Simulated Hot Full Power Conditions To assess the ARMP PDQ07 radial predictive capability at Hot Full Power (HFP) conditions, quarter assembly simulations were performed at Duke Power Company.

The codes CASMO, EPRI-CPM, and PDQ07 were used to model a 2.75 w/o 23sU i

unrodded FA at beginning of life (BOL).

I i

In an unrodded FA, usually the worst peaking occurs at BOL, since Xenon has I

not yet built up to equilibrium concentrations. And so this is a good time r

I to test t two group diffusion theory model against multigroup transport theory

codes

a) EPRI-CIM which provides a solution using collision probabilities, and

! b) CASMO which uses transmission probabilities.

( The results of these calculations are shown in Figure 3-4. The ARMP PDQ07 conservatively overpredicted the maximum CASMO pin power by 2.5% and likewise overpredicted the CPM maximum pin by 1.8%. The mean difference of the 10 l

S2 3-2 I

_ - - - - _.._~ , . _ .-. -- _ _ . _

' highest PDQ-CASMO pairs was .0143 RPD with a standard deviation of .01233.

Comparisons of the 10 maximum pin powers for PDQ-CPM likewise gave a mean dif-i ference of .0084 RPD and a standard deviation of .00902.

3.4 Conclusions In Section 3.1'it was demonstrated that the ARMP PDQ07 discrete pin model con-servatively overpredicts measured peak powers from critical experiments. Also the ARMP.PDQ07 discrete pin model conservatively overpredicts the hot full power pin powers when compared to two independent transport theory solutions.

Since this conservatism was demonstrated at het and cold conditions', it is concluded that an uncertainty component applied to the assembly local radial f

is not necessary. This component would only be necessary if the ARMP-PDQ07

. model were to underpredict the maximum pin power.

t W

f e

i i

S2 3-3 w

• w -

g.__.,.,

Table 3-1 B&W COLD CRITICALS USED FOR LOCAL RADIAL UNCERTAINTY FACTOR ANALYSIS Core -Compositions of Central Core -

Load Type of Fuel Assembly Control and Number -Number Fuel Instrument Unit Cells j XI '2 UO2 17 H2O XI 8 UO2 1 H2 0, 16 Pyrex 4

?'

l I

S2 3-4

)

) l FIGURE 3-1

-B&W URANIUM CRITICAL EXPERIMENTS GEOMETRY OF INTERIOR " ASSEMBLIES" LOAD 2 LOAD 8 LOCATION 1 WATER WATER LOCATION 2 WATER POISON LOCATION 3 WATER POISON 1

9

- 2 .

3 2

! S2 3-5 l

FIGURE 3-2 l

RELATIVE PIN POWERS IN CENTRAL ASSEMBLY OF B&W CRITICALS UNRODDED LATTICE LOAD 2 - 1335 PPM E

EFF EXPERIMENT 1.0007 ARMP PDQ07, STANDARD 1.0052

• MAXIMUM PIN EXPERIMENT 1.108 ARMP PDQ07, STANDARD 1.122 0.

O.

1.072 1.033 1.085 1.075

.993 1.040 0.

1.010 1.074 0.

.968 1.002 1.080 1.050

.985 1.033 1.087 1.096 1.013 1.082 1.108* 0.

.992

.978 1.028 1.087 1.122 0.

1.062 0. 1.096 1.073 .982

.993 .995

.978 1.050 0. 1.102 1.056

.999 .974 .941 .939

.948 .993 1.035 1.003 .959 .936 .917

.957 .998 1.036

.942 .940 .919 .890 EXPERIMENT

.951 .970 .930 .894 ARMP PDQ07,

.925 .914 .905 .896

.934 .940 .947 .939 Standard S2 3-6

4#

+

)

FIGURE 3-3

-RELATIVE PIN POWERS IN CENTRAL ASSEMBLY OF B&W CRITICALS BURNABLE POISON LATTICE LOAD 8 794 PPM h2 _

EXPERIMENT 1.0007 Y

ARMP.PDQ07, STANDARD .9993

~c.

• MAXIMUM PIN I7 EXPERIMENT 1.158 ARMP PDQ07, STANDARD 1.196

-- 0.

O.

1.102 .995 1.114 .999

.986 .907 0.

.998 .899 0.

1.003 .961 .264 .873

.990 .924 .851 .839 l

1.015 .943 .864 .820 0.

.996 .931 .855 .816 0.

.'s 1.005 .924 0. .896 .931 1.028

.1.005 .914 0. .855 .926 1.021 l

1.C'.5 1.027 .961 1.007 1.053 1.095 1.118 1.046 .993 .945 .994 1.067 1.115 1.155 1.079 1.061 1.044 1.045 1.093 1.147 1.151 1.158* EXPERIMENT

-1.087 1.077 1.070 1.090 1.123 1.153 1.178 1.196 ARMP PDQ07,

, Standard l

l S2 3-7 Yr -- * - , e---.. -n, .-~-. . _ _ _ . , - - - -

l FIGURE 3-4 OCONEE UNRODDED ASSEMBLY PIN X PIN POWER DISTRIBUTION HOT FULL POWER SIMULATION 1

CASMO EPRI-CPM AhMP PD007 CODE USED 1.2973 1.2968 1.2962 K-INF 1.2972 1.2967 1.2962 K-EFF 0 0 0 PPMB O O O EXPOSURE

• MAXIMUM PIN CASMO 1.072  ;

EPRI-CPM 1.080

-ARMP PDQ07 1.099 0.

O.

O.

g 1.009 1.012 1.015 1.022 1.007 1.025 1.001 1.039 0.

1.005 1.045 0.

1.002 1.045 0.

0.997 1.013 1.057 1.048 0.999 1.020 1.064 1.061 0.995 1.030 1.073 1.087 0.995 1.011 1.057 1.072* 0.

! 0.995 1.017 1.063 1.080 0.

l 0.991 1.027 1.074 1.099 0.

l 0.992 1.030 0. 1.056 1.036 0.989 c 0.991 1.032 0. 1.063 1.036 0.988 0.988 1.033 0. 1.075 1.042 0.992 L 0.977 0.990 1.023 0.996 0.976 0.960 0.947 0.975 0.991 1.023 0.997 0.972 0.953 0.937 0.967 0.994 1.020 1.003 0.970 0.945 0.925 0.963 0.967 0.972 0.967 0.958 0.948 0.940 0.936 CASMO 0.959 0.963 0.967 0.962 0.950 0.939 0.930 0.924 CPM l 0.947 0.954 0.962 0.955 0.940 0.926 0.915 0.910 PDQ l

l S2 3-8

.- ,. w . m.m- - , ,__,--,r .-- .._.---.+-=,-e,-- -- < - .

4. EPRI-N0DE-P POWER DISTRIBUTION COMPARISONS 4.1 EPRI-N0DE-P Model The primary three-dimensional nuclear code employed at Duke Power is EPRI-N0DE-P.

This code is used for all maneuvering analyses, cota follow, and physics test data In this section, where three-dimensional core power distributions are required.

comparinons of measured and EPRI-N0DE-P calculated values will be shown for both .

radial and total peak powers. Comparisons were performed on a total of 55 reactor state points covering Cycles 1 through 5 of Oconee Unit I.

Each fuel assembly was The Oconee core was modeled using qaprter core symmetry.

modeled with one radial and 12 equidistant axial nodes.

The active stack height was set at 144 inches. Control rods couiJ be positioned ~

continuously in this model, with a maximum inserted length of 139 inches.

Methods described in the ARMP System Documentation 13 were used to gener;te fuel neutronics characteristics. Figure 4-1 shows a flow chart of the generat methodology employed. In generating the in. 2 for EPRI-N0DE, "as built" as-sembly uranium enrichments and loadings were used to model more closely actual core ccaditions.

EPRI-N0DE-P employs fuel km fits versus moderator temperature for rodded, half-rod'ded, and nonrodded conditions. Assemblies having burnable poison n

rods are treated similarly. All km fits are referenced to either moderator temperatare or fuel exposure. Also fuel temperature reactivity changes

• are included. Therefore, EPRI-N0DE-P explicitly models the effects of thermal-hydraulic and Doppler (THD) fr 4 c.ks.

Since EPRI-N0DE-P does not accou.,1 ce: 4 tly for the core reflector and baffle, the following normalization procedure was employed. The assembly radial powers itcc EPRI-N0DE-P were normalized once at B0C to discrete pin model PDQ07 radial powers when the core reached equilibrium Xenon and Samarium conditions (s25 EFPD). Since the two dimensional PDQ07 does not have thermal and hydraulic feedbacks, all EPRI-NODE-P nornalization runs were performel likewise.

S2 4-1

Plots comparing calculated and measured assembly axial powers are given in Appendix A.

4.2 Oconee Fuel Cycle Simulations Using the EPRI-NODE-P model described in Section 4.1, Cycles 1 through 5 of Oconee 1 were depleted using THD feedbacks. The depletions were performed in a core follow mode, utilizing critical boron searches at each exposure step.

Since all depletions were performed at greater than 90% full power, control banks 1 through 5 were not inserte- into the core.

Table 4-1 shows by cycle: cora operation mode, control bank interchange /

withdrawal exposures, and core reload data. Assembly peaks were derived by the method described in Appendix A. The results of these depletions will be described below.

Cycle 1 operated in a rodded mode until end of cycle which occurred at 309.6 EFPD.14 Control bank locations are shown in Figure 4-2 and the core loading pattern is shown on Figure 4-3.

There were Assembly radial powers in this cycle were normalized at 88.6 EFPD.

two control rod bank iN.erchanges during this cycle causing considerable global power shifts.

There were 25 depletion steps in this cycle; 17 of these time points are shown in Table 4-2. Assembly radial power comparisons are shown in Figures 4-4 to 4-20. Assembly peak powers are compared in Figures 4-21 to 4-37.

b

h a

E ;;;p

'a S2 4-2

i Cycle 2 operated in a rodded mode, having a transient bank interchange at 53 EFPD, and transient bank withdrawal at 237 EFPD. The reactor continued operation until 292 EFPD.15 Figure 4-38 shows the control rod bank locations.

Five 2.60 and 56 3.20 w/o 2ssU enriched assemblies were loaded. Figure 4-39 shows the assembly enrichments, and Figure 4-40 shows the quarter core assembly shuffle used to rearrange fuel from EOC-1 to BOC-2.

The EPRI-NODE-P radial powers were normalized to a discrete pin quarter core POQ07 depletion at 22 EFPD. EPRI-NODE-P simulation of Cycle 2 employed 20 timesteps, 10 of these are shown in Table 4-3. These 10 state points are ,

where comparisors to measured power data were made. Figures 4-41 to 4-50 l show comparisons for assembly radial powers. Assembly peak powers are com-pared in Figures 4-51 to 4-60.

During Cycle 3, Unit 1 operated in the rodded mode. A transient bank inter-change occurred at 100 EFFD, and transient. bank withdrawal occurred at 245.8 EFPD. The reactor continued operation until 303.8 EFPD.is Figure 4-61 shows the locations of control banks 6, 7, and 8. Sixty 2.75 w/o 23sU enriched fuel assemblies were loaded as shown in Figure 4-62. The Cycle 2 to 3 assembly shuffle pattern is shown in Figure 4-63.

EPRI-NODE-P was normalized at 38.5 EFPD to a parallel quarter core discrete pin model PDQ07 depletion. Cycle 3 was simulated using 16 depletion time steps, I 10 of which are shown on Table 4-4.

Figure 4-64 through 4-73 compare assembly radial powers, and Figures 4-74 through 4-83 compare total peak powers.

l

\

S2 4-3 l

l During startup testing at B0C-4, a quadrant tilt occurred that limited reactor The reactor operational mode operation to 75% full power for the first 10 EFPD.

The reactor was changed to feed and bleed to provide additional safety margin.

continued operation at full power until a control rod dropped about 2 EFFD before shutdown. The reactor power level'was lowered to 50% full power, and operation continued until shutdown at 245.9 EFPD.17 During the refueling preceding Cycle 4, 34 of 52 detector strings were replaced.

The Therefore, the power measurement uncertainty was considerably improved.

measured power data from this cycle was included in the analysis of EPRI-N0DE-P reliability factor.

The EPRI-N0DE-P Cycle 4 model was normalized to the Duke Power discrete pin model PDQ07 depletion at 25 EFPD. The exposure steps used for power comparisons The control bank patterns are shown on Figure 4-84.

are shown on Table 4-5.

Core loading and quarter core assembly shuffle patterns are shown on Figures 4-85 and 4-86, respectively.

Calculated assembly radial and peak powers agreed very well with measured.

Figures 4-87 to 4-95 comptre the radials ana Figures 4-96 to 4-104 compare the peaks.

Locations During Cycle 5, the core also operated in a feed and bleed mode.

of control rod banks are given in Figure 4-105. Fifty six 3.02 w/o 23sU enriched assemblies "ere loaded as shown in Figure 4-106. Figure 4-107 shows the quarter core assembly shuffle pattern.

Ten equilibrium reactor state points were used for power comparisons as shown in Figure 4-6. As in Cycle 4, the EPRI-NODE-P radial assembly power distri-bution was normalized to the PDQ07 parallel depletion at 25 EFPD.

In Figures 4-108 to 4-117, the assembly EPRI-N0DE-P and measured radial powers are compared. Figures 4-118 to 4-127 show the calculated and measured assembly peak powers.

S2 4-4

4.3 Conclusions EPRI-NODE-P yielded consistently good power distributions when compared to measured power distributions. This conclusion applies for both radial arm peak power comparisons. Although the conclusions in this section are qualitative, quantitative statistical results of these comparisons will be shown in Section 5.

S2 4-5

Table 4-1 OCONEE-1 CORE CONDITIONS Nominal Bank 7 Bank 7 Assembly Loading Mode of Interchange Withdrawal (w/o U-235/

Cycle Operation (EFPD) (EFPD) # Assembly 1 Rodded 91.5 None 2.00/41, 2.10/56, 196.0 2.15/80 2 Rodded 53.0 237.0 2.60/5, 3.20/56 3 Rodded 100.0 245.8 2.75/60 4 Unrodded - -

2.79/56 5 Unrodded -

3.02/56 i

l i

i-I-

i-i.

I i

I l

I S2 4-6 l

Table 4-2 I

OCONEE UNIT 1 STATE POINTS CYCLE 1 J

Control Bank Position (% WD) Axial Offset, %

Point Power EFPD (Percent) Group 6 Group 7 Group 8 (Meas / Calc) 1 88.6 96.4 94.9 22.6 2.0 -4.74/-3.10 2 122.2 98.1 96.7 24.3 0.5 -0.22/-1.98 3 129.0 97.4 98.6 27.2 1.6 -4.07/-2.34 4 145.6 100.7 98.4 24.5 0.7 -2.96/-2.62 5 169.7 100.1 92.9 20.4 4.1 -2.87/-2.47 ,

6 186.4 100.5 98.6 26.0 0.6 -4.23/-4.33 7 195.8 99.5 99.6 27.8 1.3 -3.70/-4.28 8 208.9 99.8 96.4 20.2 3.9 -4.17/-4.51 9 225.5 98.8 93.3 18.1 0.7 -3.04/-4.91 10 232.4 99.0 90.1 16.3 3.4 -1.81/ '.85 ,

11 239.6 100.8 96.7' 18.8 4.3 -0.02/-4.28 12 246.4 100.8 96.0 20.8 4'. 3 -1.68/-5.97 13 253.4 100.4 94.3 20.3 2.7 -4.67/-6.24 14 260.4 100.2 91.8 17.0 4.9 -3.11/-4.50 15 267.4 100.1 92.1 16.1 4.9 -2.05/-4.05 16 270.7 99.2 96.3 21.3 6.9 -1.34/-6.27 17 281.6 90.5 92.1 17.2 12.0 -2.35/-6.37 l

l l

S2 4-7

l I

Table 4-3 OCONEE UNIT 1 STATE POINTS CYCLE 2 Power Control Bank Position (% WD) Axial Offset, %

Point Group 6 Group 7 Group 8 (Meas / Calc)

1. ' EFPD (Percent) 1 .30.6 98.9 84.7 10.2 6.2 -2.28/-3.30 2 52.5 97.9 85.5 9.5 4.7 -3.30/-3.24 3 83.0 98.7 84.9 6.4 3.4 -2.65/-5.35 4 103.5 98.8 85.8 7.6 3.5 -7.61/-5.58 5 129.0 97.2 83.2 3.4 4.0 -5.57/-3.20 6 156.0 98.7 85.2' 6.3 3.3 -2.07/-4.51 7 184.0 99.9 85.3 6.6 2.9 -4.19/-4.63 8 203.8 99.6 84.6 5.6 5.7 -1.53/-3.78 9 222.9 99.7 81.5 2.9 6.1 -0.39/-4.05 10 259.9 99.9 100.0 83.7 15.6 -2.77/-3.90 1

I S2 4-8 l

ee, - .,- . . - , , , - . .. . . --

Table 4-4 OCONEE UNIT 1 STATE POINTS

- CYCII 3 Axial Offset, %

Point Power Control Bank Position (% WD)

Group 6 Group 7 Group 8 (!;eas/ Calc)

1. EFPD (Percent) 99.8 94.1 20.3 18.2 -6.78/0.07 1 25.3 99.7 94.6 18.3 18.6 -4.88/-1.02 2 58.5 99.6 94.9 18.3 17.5 -4.72/-1.93 3 91.2 99.9 92.3 18.4 25.6 -5.97/-5.15 4 121.9 99.4 97.6 22.4 19.5 -2.38/-1.14 5 143.9 99.7 94.6 21.1 24.6 -4.54/-6.02 6 179.1 98.9 98.5 22.5 23.8 -4.88/-3.48 7 203.3 232.6 99.9 97.8 22.5 25.4 -5.92/-5.49 8

99.9 100.0 85.3 25.8 2.00/ 1.38 9 266.0 90.2 100.0 84.5 31.6 -2.55/-4.62 10 303.1 S2 4-9

Table 4-5

-OCONEE UNIT 1 STATE POINTS CYCLE 4 l l

Power Control Bank Position (% WD) Axial Offset, % j Point Group 6 Group 7 Group 8 (Meas /Cale) )

1. EFPD -(Percent) 1 28.3 99.4 100.0 83.8 35.3 -2.16/2.44 2 56.6 99.7 100.0 80.7 32.5 -1.13/0.68

-3 83.2 99.5 100.0 86.0 34.7 -2.00/-0.57 4 103.4 99.9 100.0 85.6 32.6 -2.15/-0.01 5 125.1 99.6 100.0 82.7 29.4 -1.61/-0.33 6 150.6 99.0 100.0 84.6 29.4 -2.90/-0.43 7 174.8 98.8 100.0 88.1 30.5 -3.15/-0.32 8 194.5 98.8 100.0 84.3 28.3 -2.04/-2.02 9 234.7 99.6 100.0 88.7 29.6 -2.42/-1.55 s

i S2 4-10

- . _ ~ --

Table 4-6 OCONEE UNIT 1 STATE POINTS CYCLE 5 Control Bank Position (% WD)

Axial Offset, %

Point Power Group 6 Group 7 Group 8 (Meas / Calc)

1. EFPD (Percent) 97.0* 100.0 89.6 26.1 -0.41/-0.78 1 24.5 44.4 97.0 100.0 90.0 28.6 -2.54/-3.62 2

99.5 100.0 88.7 25.6 -4.61/-3.17 3 90.4 99.0 100.0 90.1 26.1 -1.84/-2.86 4 118.1 99.5 100.0 87.8 22.3 -2.89/-2.70 5 145.4 98.8 100.0 87.8 22.3 -3.33/-3.22 6 175.3 99.1 100.0 90.1 25.9 -4.87/-4.29 7 207.5 98.2 100.0 88.4 19.8 -0.94/-1.81 8 239.7 98.5 100.0 89.8 20.3 -1.40/-1.59 9 263.6 93.8 100.0 91.2 20.4 -1.53/ 0.44 10 288.1

• Calculation performed at 97.5% Full Power d

S2 4-11

Figure 4-1 Outline of P9R Data Flow BOL CPM

' BOL EPRI-CELL for EPRI-CELL Depletion Control for Element *~ """ ' "* '

Burnable Poison Water Hole Representa-Represen-tation. Parameterization tion

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PDQ-7/ HARMONY Core X-Y Core X-Y Assembly X-Y Depletion Beginning of Cycle (BOC) Depletion d

1 Core

_ Rearrangement for Next Cycle l

q o

K=,M 2 ,KIf,vEf, etc. fits as a fBOL Normalization \

of Global Code to function of depletion end other kassemblypower ) )

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i! 1I Global (3-D) Nodal .- -

Calculation with Depletion 1

K, critical boron concentration, power distribution, reactivity .-

coefficients as functions of operating conditions L

S2 4-12

Figure 4-2 i Oconee 1, Cycle 1 Control Rod Configuration 12 13 14 15 B 9 10 11 Rg H Tr2 Tr3 Rg-Tr1 Ap t Tr3 ,

M T:2 Tri N Tr3 Ap 0

Rg Rg = Regulating Bank (6)

Tr1 = Transient Bank 1 (7)

R (0-92 EFPD)

Tr2 = Transient Bank 2 (7)

(92-196 EFPD)

Tr3 = Transient Bank 3 (7)

(196-310 EFPD)

Ap = APSR (8)

S2 4-13

Figure 4-3 Oconee 15 Cycle 1 Quarter Core Loading Diagram 11 12 13 14 15 8 9 10 2.00 2.00 2.10 2.15 2.15 H 2.00 2.00 2.00 3 1 1 2 3 1 1 1 2.00 2.10 2.10 2.15 2.15 2.00 2.00 2.00 3 K 1 2 2 3 1 1 1 2.10 2.10 2.10 2.15 2.15 2.00 2.00 2.00 3 L 2 2 2 3 1 1 1 2.10 2.15 2,15 2.00 *2.00 2.10 2.10 2 2 3 3 M 1 1 2 2.10 2.10 2.15 2.15 2.15 2.00 2.10 3 N 2 2 3 3 1 2 2.10 2.10 2.15 2.15 2.15 0 2.10 3 2 2 3 3 2

2.15 2.15 2.15 2.15 2.15 P 3 3 3 3 3

2.15 2.15 2.15 R 3 3 3

-W/0 U-235 Batch No.

t 1

S2 4-14

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S2 4-15

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I S2 4-16

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S2 4-17 i.

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S2 4-18 l

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S2 4-19

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S2 4-21 I

f Figure 4-11 OCONEE*i CY=i EPRI=N00E CALC VS Meas RA0!al ASSEMBLv t'cw e3 99.45 FP CONTROL BANKS ..re. 9. 4, 20.2e 4.01 N!TdDR AWN 204.9 EFPD in is i. is

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S2 4-22

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t S2 4-23

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S2 4-24

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S2 4-25

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S2 4-26 mm -

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S2 4-28

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Figure 4-19 =

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S2 4-35

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l Figure 4-26 1 occhEE.1 CY.1 epa!.400E CALC vs MEAS ASSE#8LY PEAK POWERS l 18..s EPPO 100.4s PP ConfROL Bawus 6e7 8 98.6, 26.e. 0..t h!THORAWN-13 14 15 e 9 to 11 12

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Figure 4-28 OcnNEE=1 CY=1 EPRI 400E CALC vs MEAS ASSEW8LY PEAM POWERS 204.9 EFPO 99.45 FP CONTROL SANKS .e7,0 9. 4e 20.2e 4.C% WITHORAWN 9 to 11 12 13 to 15 e

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OCONEE=1 CV=1 EPRI= NODE CALC VS #EAS ASSEWOLT PEAR POWERS 99.pt FP CONTROL BANWS 6ere$.9.... 20.2e ..et w!THOWAWN .I 225 6 EFPD 12 13 1 tC e 9 1. 11 N

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h DCONEE=1 CY=1 EPR!=N00E CALC VS MEAS ASSEMBLY PEAK PO.ERS 1...RI FP CONTROL BANES 6,7e8 9. 7e 14.ge 4 33 WITHORAWN I 239.. EFPD 4 9 10 11 12 13 la 15 l

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Figure 4-32

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OCONEE=1 CY=1 E##!=N00E CALC V8 #EAS ASSEMBLY PEAN POWERS

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i figure 4-33 OCONEE=1 CY=1 (PRI= NODE CALC VS MEAS ASSEMSLT PEAM POWERS CONTROL BANMS 6,7,0 94.3e 20.3e't.7% WITHORAWN i 253.s EFP0 100.at FP 8 9 10 11 12 13 to 15 i w  : 1:11 : 1:11 7 1:ll

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L OCONCEst CYet EPR! N005 CALC VS 88t AS ASSEW9LY PE AK PC.tRS CONTROL BANWS 6,7et.91.Se 17.@e 4.91 WIT 180RAWN 260 4 EPPD 100.it PP e 9 le 11 12 13 la 15

13 :

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f Figure 4-35 0Co#EE=1 CY=1 EPRI NODE CALC Vs *EAS A8stpsLY , Eau pc. Ens 267.a EFP0 10..nz FP contact sawns ser.4 92.1. t.1, s.ex w!TMonawu 12 13 la 15 8 9 to 11

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S2 4-46

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1 Figure 4-36 OcchEEe1 CV=1 EP#!*N00E CALC VS WEAS ASSEPSLY PEAK POWERS 270.7 (FPO 100.at FP CONTROL BANRS 6e7et.92.1e 14.le 4.95 N!THORAWN 12 13 18 15 8 9 10 11

. 7 . . . .e .

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OCONEE=1 CY=1 EPe!=N00E CALC vs MEAS ASSEMBLY PEAx Powcas Rei.6 'EPPO 9..StPP CONTROL SAMus 6efe8 92.5ei7.3 12.st m!?M0aAWN

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Figure 4-38

- Oconee 1, Cycle 2 Control Rod Configuration 4

9 10 11 12 13 14 15 i 8 i

H Tr1 Tr2 Tr1 K

Tr2 Tr2 Ap Rg

-L M

N 3

Ap Tr1 0 Tr2 4

P Rg Rg = Regulating Bank (6)

R Tr1 = Transient Bank 1 (7)

(0-53 EFPD)

Tr2 = Transient Bank 2 (7)

(53-292 EFPD)

Ap = APSR (8)

. S2 4-49

.,-,,am.., . --. ---.--..----.m.-,+. - - . - - - - -

Figure 4-39 Oconee 1, Cycle 2 Quarter Core loading Diagram 8 9 10 11 12 13 14 15 i

2.60 2.15 2.15 2.60 2.15 3.20 2.15 3.20 H

4A 3 3 4A 3 4B 3 4B 2.15 2.10 2.15 2.15 2.10 2.10 2.15 3.20 K 3 2 3 3 2 2 3 4B 2.15 2.15 2.15 2.15 2.15 2.10 3.20 3.20 L 3 3 3 3 3 2 4B 4B 2.60 '2.15 2.15 3.20 2.10 2.15 3.20 M 4A 3 3 4B 2 3 4B 2.15 2.10 2.15 2.10 2.15- 2.15 3.20 N 3 2 3 2 3 3 4B 0 3.20 2.10 2.10 2.15 2.15 3.20 4B 2 2 3 3 4B 2.15 2.15 3.20 3.20 3.20 P 3 3 4B 4B 4B R 3.20 3.20 3.20 4B 4B 4B W/0 U-235 Batch tio.

S2 4-50

~ 1 Figure 4-40 Oconee 1, Cycle 2 Quarter Core Shuffle Pattern 8 9 10 11 12 13 14 15 H F 0-13* K-15 F H-14 F H-15 F

~

~ * ~ ~ ~ ~ ~

K 1

1 R-9** P-11 R-9** L-14 L-15 K-13 F F g

M F P-9 P-10 F M-12 N-13 F i l

P-8 0-10 R-10 N-11 N-12 N-14 F N

0 F M-10 0-9 0-12 P-12 F R-8 0-11 F F F P

R I I I Previous cycle loc.

• Assembly location formerly on diagonal placed on the core flat.

• • Moved from eighth core symmetric location to quarter core symmetric location.

S2 4-51

- - _ _ - _ - _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ -- - ~ _ _ _ _ , _ _ _ _ _ _

l Figure 4-41 OCOMEE*1 CY=2 EPRIe400E CALC VS MEAS RACIAL ASStutLY POWERS

3. 6 EFPO 98.9tFP CONTROL BANAS 6,7, essa.7,10.to. 22 WITN3RAWN 12 13 la 15 4 9 it 11

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Figure 4-45 OCONEE=1 CYet EPRI. NODE CALC VS MEAR RA0!AL ASSEMBLY P0.ERS 129.8 EFP0 97.2t PP CON 7ROL BAhKS 6e7 4 83.2e3.s.a.0 I WITHORA.W

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l Figure 4-48 OCONEEe, CYet E*R!aN00E CALC VS MEA 8 RA0!AL ASSEMRLY #0.ERS 203 8 EFPO 99..tFP CONTROL Samus 6 7.a.ta.4 5..e5.7 5 n!THDaanu 15

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l Figure 4-50 OCONEE=1 C?ot EPRI=N00E CALC VS MEAS RA0!AL ASSEM9LY PD.nR8 259.9 EFPO 99.9tFP CONTROL tihKS 6e7et. 200e83.ve15.6I h!7MORA.N 13 1 15 10 11 12 8 9

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l S2 4-61

Figure 4-51 OC04EE*1 CY=2 EPRI N00E CALC Vs MEAS ASSEW9LY PEAK P0.ERS

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i Figure 4-52 EPRI*N00C CALC VS MEAS ASSEM8LY PEAM POWEWS .

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S2 4-63

Figure 4-53 0CowtE-1 CYet EPe! NODE CALC vs PEAS ASSEW8LY Pfax POWERS 83 0 EPPD 98.7tFP CONTROL #aNM8 6e7. eses.9,6.a.3 4 1 1THORAWN 12 13 14 15 .

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r Figure 4-54 l -

OCONEE 1 CY=2 EPRI.400E CALC Vs *EAS ASSEwSLY PEAK POWEe8 103 5 EFPO 98.8tFP C0wfa0L BANKS . 7,3 85.8,7.. 3.31 WITHORAWN 13 la 15 to 12 .

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Figure 4-55 oCnktE=1 cv=2 E.t!=N00E CALC VS WEAS AS$fMSLY PEAK POWERS 129.0 EFPD 97.2tFP CONTROL 8ANMS 6,7,8 83.2e3.4e4. 1 WITMORAwN 12 13 to 15 8 9 to 11

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t Figure 4-56 i

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Figure 4-57 y-OCONEE=1 CY=2 EPR!* NODE CALC Vs *EAS ASSENSLY PEAK P0.'de$ ,

l 184.0 EFPO 99.9tF# CON,ROL SANKS 6 1.8 85.3*6.6 2.9 I n!, HORA.N 12 13 14 15 8 9 to 11

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Figure 4-61 Oconee 1, Cycle 3 Control Rod Configuration 10 ' 11 12 13 14 15 8 9 Rg1 Tr2 ,

Tr1 K

Rg1 Ap Rg2 L

M Rg1 N Tr2 Ap Tr2 Tr1 0

P Rg2 ,

Rg1 = Regulating Bank 1 (6)

(0-100 EFPD)

R Rg2 = Regulating Bank 2 (6)

(100-308 EFPD)

Tr1 = Transient Bank 1 (7)

(0-100 EFPD)

Tr2 = Transient B.ank 2 (7)

(100-308 EFPD)

Ap = APSR (8)

S2 4-72

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Figure 4-62 Oconee 1, Cycle 3 Quarter Core Loading Diagram 10 11 12 13 14 15 8 9 2.15 2.60 3.20 2.15 2.75 2.15 2.75 e H 2.60 4A 4B 3 5 3 5 4A 3 l

2.15 3.20 2.15 3.20 3.20 3.20 2.15 2.75 K 4B 3 4R 4B 4B 3 5 3

e 2.60 2.15 2.15 3.20 2.15 3.20 2.75 2.75

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• - -Batch No.

I S2 4-73

f Figure 4-63 Oconee 1, Cycle 3 Quarter Core Shuffle Pattern 9 10 11 12 13 14 15 8

H-10 H-11 H-15 H-14 F H-9 F H H-8 i

L-8 0-13 K-10 L-14 K-15 M-14 K-14 F M-8 L-9 L-10 L-15 L-12 N-14 F F M R-8 P-10 R-10 M-11 N-13 M-13 F P-8 R-9 N-10 0 - 1.* H-13 F F y

i.

O F P-11 P-12 0-11 F F f

K-8 P-9 F F F p

l R F F F

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f

. S2 4-74

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Figure 4-70 DCONEE*i CYe3 EPR! ANODE CALC V3 NEAS RA0!AL ASSEMBLY PC.ERS 2 3 3 EPPO 9. 9tFP CONTROL .ANM3 . 7e. 9. 5e 22.5, 23. 1 u!THORAWN 9 1. it it i3 is il

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