ML19350E087
| ML19350E087 | |
| Person / Time | |
|---|---|
| Site: | 05000470 |
| Issue date: | 06/15/1981 |
| From: | ABB COMBUSTION ENGINEERING NUCLEAR FUEL (FORMERLY |
| To: | |
| References | |
| NUDOCS 8106160633 | |
| Download: ML19350E087 (111) | |
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CESSARf ,, Docket STN-50-470F The following sheets of CESSAR-F Chapter 4 are to be removed and inserted:
REtiOVE THE FOLLOWING: INSERT THE FOLLOWING: " EFFECTIVE PAGE LISTING TABLE OF CONTENTS TABLE OF CONTENTS Sheets v,vi, vii, viii, x1, xii, Sheets v,vi, vii, viii, xi, xiv, xy, xvi, xvii, xviii xii, xiv, xv, xvi, xvii, xviii
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TEXT TEXT , i l 4.3-1 to 4.3-34 4.3-1 to 4.3-35 i p TABLES TABLES 4.3-1 to 4.3-23 4.3-1 to 4.3-22 ; ,
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FIGURES FIGURES 4.3-1 to 4.3-73 4.3-1 to 4.3-70 ' i L i s I
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Chapter 4 ; 1 Table of Contents 1 , i P_ age, Amendment i ! i to iv y to vii 3 i , vii, ix, x, xi - xii to xy 3 .
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xvi xvii to xviii 3 l l Text ! ) Page, Amendment ' 4.3-1 to 4.3-35 3 4 i
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i i 4.3-1 to 4.3-22 3 i Figures Page Amendment 4.3-1 to 4.3-70 3 l NOTE : Please note that the entire Section 4.3 text has been re-typed. Therefore, the entire section, including unchanged pages, are to be replaced. l l l l J t
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__ __ - _ _ _ . - - . . __ . ._ , TABLE OF CONTENTS (Cont'd) CHAPTER 4 Section Subject Page No. 4.3.1.8 Power Distribution Control 4.3-2 4.3.1.9 Excess CEA Worth with Stuck Rod Criteria 4.3-2 4.3.1.10 Chemical Shim Control 4.3-2 4.3.1.11 Maximum CEA Speeds 4.3-2 4.3.2 *ESCRIPTION 4.3-3 4.3.2.1 Nuclear Design Descrfotion 4.3-3 4.3.2.2 Power Distribution 4.3-3 4.3.2.2.1 General 4.3-3
, 4.3.2.2.2 Nuclear Design Limits on the Power 4.3-4 Distribution
' 4.3.2.2.3 Expected Power Distributions 4.3-5 4.3.2.2.4 Allowances and Uncertainties on Power 4.3-6
/] 4.3.2.2.5 Distributions V Comparisons Between Limiting and Expected Power Distributions 4.3-7 4.3.2.3 Reactivity Coefficients 4.3-7 4.3.2.3.1 Fuel Temperature Coefficient 4.3-8 4.3.2.3.2 Moderator Temperature Coefficient 4.3-8 4.3.2.3.3 Moderator Density Coefficient 4.3-9 4.3.2.3.4 Moderator Nuclear Temperature Coefficient 4.3-9 4.3.2.3.5 Moderator Pressure Coefficient 4.3-9 4.3.2.3.6 Moderator Void Coefficient 4.3-9 4.3.2.3.7 Power Coefficient 4.3-10 4.3.2.4 Control Requirements 4.3-12 4.3.2.4.1 Reactivity Control at 80C and E0C 4.3-12 '
4.3.2.4.2 Power Level and Power Distribution Control 4.3-12 l3 4.3.2.4.3 Shutdown Reactivity Control 4.3-13 4.3.2.4.3.1 Fuel Temperature Variation 4.3-13 4.3.2.4.3.2 Moderator Temperature Variation 4.3-13 l3 4.3.2.4.3.3 Moderator Voids 4.3-14 4.3.2.4.3.4 Control Element Assembly Bite 4.3-14 4.3.2.4.3.5 Part-Length CEA Effects 4.3-14
. 4.3.2.4.3.G Accident Analysis Allowance 4.3-14 3 '
s 4.3.2.4.3.7 Available Reactivity Worth 4.3-14
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Amendment No. 3 v May 28, 1981
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TABLE OF CCNTENTS (Cont'd) CHAPTER 4 Section Subject Page No. 4.3.2.5 Control _ Element Assembly Patterns and Reactivity 4.3-15 Worths 4.3.2.6 Criticality of Reactor During Refueling 4.3-16 4.3.2.7 Stabili ty 4.3-16 4.3.2.7.1 General 4.3-16 4.3.2.7.2 Method of Anlaysis 4.3-16 4.3.2.7.3 Expected Stability Indices 4.3-17 3 4.3.2.7.3.1 Radial Stability 4.3-17 4.3.2.7.3.2 Azimuthal Stability 4.3-18 4.3.2.7.3.3 Axial Stability 4.3-18 4.3.2.7.4 Control of Axial Instabilities 4.3-18 4.3.2.7.5 Summary of Special Featut es Required by 4.3-18 Xenon Instability 4.3.2.7.5.1 Features Provided for Azimuthal 4.3-19 Xenon Effects 4.3.2.7.5.2 Features Provided For Axial Xenon 4.3-19 3 Effects and Power Distribution Effect and Control 4.3.2.8 Vessel Irradiation 4.3-19 4.3.3 ANALYTICAL METHODS 4.3-20 4.3.3.1 Reactivity and Power Distribution 4.3-20 4.3.3.1.1 Method of Analysis 4.3-20 4.3.3.1.1.1. Cross Section Generation 4.3-20 4.3.3.1.1.2 Fine Mesh Methods 4.3-22 4.3.3.1.1.3 Course Mesh Methods 4.3-23 4.3.3.1.2 Comparisons with Experiments 4.3-27 Critical Experiments 4.3-27 3 4.3.3.1.2.1 4.3.3.1.2.2 Power Reactors 4.3-28 4.3.3.1.2.2.1 Startup Data 4.3-28 4.3.3.1.2.2.2 Depletion Data 4.3-30 0 Amendment No. 3 vi May 28, 1981
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! ; TABLEOFCONTENTS(Cont'd) ,
CHAPTER 4 Subject Page No.
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i Section 4.3.3.2 Spatial Stability 4.3-31 1
, 4.3.3.2.1 Methods of Analysis 4.3-31
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, 4.3.3.2.2 Radial Xenon Oscillations 4.3-31 3 i ! 4.3.3.2.3 Azimuthal Xenon Oscillations 4.3-31 i
! 4.3.3.2.4 Axial Xenon Oscillations 4.3-32 '
! I 4.3.3.3 Reactor Vessel Fluence Calculation Model 4.3-32 l3 4 4.3.3.4 Local Axial Power Peaking Augmentation 4.3-33 l
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! 4.3-33 3 I 4.3.4 SECTION DELETED l 4.4- THERMAL AND HYDRAULIC DESIGN 4.4-1 ! 4.4.1 DESIGN BASES 4.4-1 , f 4.4.1.1 Minimum De)arture from Nucleate Boiling 4.4-1 1 Ratio (DNB1) , 4.4.1.2 Hydraulic Stability 4.4-1 , dJ.1.3 Fuel Design Bases 4.4-1 i
'4.4.1.4 Coolant Flow, Velocity, and Void Fraction 4.4-2
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4.2 DESCRIPTION
OF THERMAL HYDRAULIC DESIGN OF THE 4.4-2 i REACTOR CORE i ! 4.4.2.1 Sumary Comparison 4.4-2 ' 4.4.2.2 Critical Heat Flux Ratios 4.4-2 j 4.4.2.2.1 Departure from Nucleate Boiling Ratio 4.4-2 4.4.2.2.2 Application of Power Distribution and 4.4-3 :
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Engineering Factors 4.4.2.2.2.1 Power Distribution Factors 4.4-4 4.4.2.2.2.2 Engineering Factors 4.4-5 4.4.2.2.3 Fuel Densification Effect on DNBR 4.4-7 4.4.2.3 Linear Heat Generation Rate 4.4-7 4.4.2.4 Void Fraction Distribution 4.4-7 O l 1 hnendment No. 3 May 28, 1981 ; i i ! ! _ _ . _ _ _ _ . _ . _ _ _ _ _ . _ _ . _ _ _ . _ _ _ _ _ _ , _ . , _ , . . . _ , _ _ , . . .;
TABLE OF CONTENTS (Cont'd.) CHAPTER 4 Section Subject Page No. 4.4.2.5 Core Coolant Flow Distribution 4.4-7 4.4.2.6 Core Pressure Drops and Hydraulic Loads 4.4-8 4.4.2.6.1 Reactor Vessel Flow Distribution 4.4-8 4.4.2.6.2 Reactor Vessel and Core Pressure Drops 4.4-8 4.4.2.6.3 Hydraulic Loads on Internal Components 4.4-8 4.4.2.7 Correlations and Physical Data 4.4-9 4.4.2.7.1 Heat Transfer Coefficients 4.4-9 4.4.2.7.2 Core Irrecoverable Pressure Drop Loss 4.4-10 Coefficients 4.4.2.7.3 Void Fraction Correlations 4.4-11 4.4.2.8 Thermal Effects of Operational Transients 4.4-11 4.4.2.9 Uncertainties in Estimates 4.4-12 4.4.2.9.1 Pressure Drop Uncertainties 4.4-12 4.4.2.9.2 Hydraulic Loads Uncertainties 4.4-12 4.4.2.9.3 Fuel and Clad Temperature Uncertainty 4.4-12 4.4.2.9.4 DNBR Calculation Uncertainties 4.4-12 4.4.2.10 Flux Tilt Considerations 4.4-14 4.
4.3 DESCRIPTION
OF THE THERMAL AND HYDRAULIC DESIGN OF 4.4-14 THE REACTOR COOLANT SYSTEM (RCs) 4.4.3.1 Plant Configuration Data 4.4-14 l l 4.4.3.1.Y Configuration of the RCS 4.4-14 4.4.3.2 Operating Restrictions on Pumps 4.4-15 4.4.3.3 Power Flow Operating Map (BWR) 4.4-15 4.4.3.4 Temperature - Power Operating Map (PWR) 4.4-15 4.4.3.5 Load Following Characteristics 4.4-15 t 4.4.3.6 Thermal and Hydraulic Characteristics Table 4.4-16 O viii
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TABLE OF CONTENTS (Cont'd.) l
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CHAPTER 4 !
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1 , j - l Section Subject Page No. l 4.6.2.3.1 Pipe Breaks 4.6-2 i i ) 4.6.3 TESTING AND VERIFICATION OF THE CRDS 4.6-2 4 4.6.4 INFORMATION FOR COMBINED PERFORMANCE OF THE REACTIVITY 4.6-2 CONTROL SYSTEMS 4.6.5 EVALUATION OF COMBINED PERFORMANCE 4.6-2 I i l APPENDIX 4A QUIX COMPUTER CODE DESCRIPTION 4A-1 I i
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APPENDIX 48 SYSTEM 80 REACTOR FLOW MODEL TEST PROGRAM 48-1 4
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LIST OF TABLES CHAPTER 4 Table Subject 4.2-1 Mechanical Design Parameters 4.2-2 Tensile Test Results on Irradiated Saxton Core III Clading 4.2-3 C-E Poolside Fuel Inspection Program Sumary 4.3-1 Nuclear Design Characteristics 4.3-2 Effective Multiplication Factors and Reactivity Data 4.3-3 Comparison of Core Reactivity Coefficients with Those Used in Various Safety Analysis 4.3-4 Reactivity Coefficients 4.3 5 Worths of CEA Groups 4.3-6 CEA Reactivity Allowances 4.3-7 Comparison of Available CEA Worths and Allowances 4.3-8 Comparison of Rodded and Unrodded Peaking Factors for Various 3 Rodded Configurations 4.3-9 Calculated Variation of the Axial Stability Index 4.3-10 Maximum Fast Flux Greater Than 1 MeV 4.3-11 Control Element Assembly Shadowing Factors 4.3-12 C-E Criticals 4.3-13 Fuel Specifications (KRITE Experiments) 4.3-14 Comparison of Reactivity Levels for Non-Uniform Core 4.3-15 B0C,HZP, XE FREE, Unroeded Critical Boron Concentration 3 4.3-16 ITC Summary for ROCS /DIT 4.3-17 Compa 1 son of Calculp ed rad Masured CEA Bank Worths 4.3-18
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omparison of Pou a >, cients
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LIST OF TABLES (Cont'd.) V CHAPTER 4 Table Subject ' 4.3-19 Power Distributions Summary of Calculational And Measurement Uncertainties 4.3-20 Axial Xenon Oscillations 3 j 4.3-21 Densification Characteristics 4.3-22 Radial Pin Power Ireternal Census 4.4-1 Thermal and Hydraulic Parameters 4.4-2 Comparison of the Departure From Nucleate Boiling Ratios Computed with Different Correlations 4.4-3 Reactor Coolant Flows in Bypass Channels 4.4-4 Reactor Vessel Best Estimate Pressure Losses and Coolant Temperatures 4.4-5 Design Steady State Hydraulic Loads on Vessel Internals and Fuel Assemblies
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4.4-6 RCS Valves and Pipe Fittings 4.4-7 RCS Design Micimum Flows 4.4-8 Reactor Coolant System Geometry 4.4-9 Reactor Coolant System Component Thernal and Hydraulic Data j 4.6-1 Postulated Accidents ! l
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... Amendment No. 3 x111 May 28, 1981
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LIST OF FIGURES CHAPTER 4 Figure Subject 4.1-1 Reactor Verti ' Arrangement 4.2-1 Circumferential Strain vs Temperature 4.2-2 Design Curve for Cyclic Strain Usage of Zircaloy-4 at 7000F 4.2-3 Full Length Control Element Assembly (4-Element) 4.2-9 Full Length Control Element Assembly (12 - Element) 4.2-5 Part Length Control Element Assembly 4.2-6 Fuel Assembly 4.2-7 Fuel Spacer Grid 4.2-8 Fuel Rod 4.2-9 Burnable Poison Rod 4.2-10 Control Element Assembly Locations 4.3-1 First Cycle Fuel Loading Pattern 4.3-2 First Cycle Assembly Fuel Loading Waterhole and Shim Placement 3 4.3-3 Planar Average Power Distribution, BOC, Unrodded, Full Power, No Xenon, 0 MWD /T 4.3-4 Planar Average Power Distribution, B0C, Unrodded, Full Power, Equilibrium Xenon, 50 Mwd /T
- 4. 3 ', Planar Average Power Distribution,1000 Nd/T, Unrodded, Full Power, Equilibrium Xenon l 4.3-6 Planar Average Power Distribution, 6000 Nd/T, Unrodded, Full Power , Equilibrium Xenon 4.3-7 Planar Average Power Distribution, 9000 Nd/T, Unrodded, Full Power, Equilibrium Xenon 4.3-8 Planar Average Power Distribution, E0C, Unrodded, Full Power, Equilibrium Xenon,16500 Mwd /T 4.3-9 Planar Average Power Distribution, BOC, BankFull In, Full 3 Power, Equilibrium Xenon, 2000 Nd/T Amendment No. 3 xiv May 28, 1981
l r\ LIST OF FIGURES (Cont'd) (O CHAPTER 4 Figure Subject 4.3-10 Planar Average Power Distribution, 9000 N d/T, Bank 5 Full In, Full Power, Equilibrium Xenon 11 Planar Average Power Distribution, E0C, Bank 5 Full In, Full Power, Equilibrium Xenon,14000 Nd/T 4.3-12 Planar Average Power Distribution, BOC, Part-Length Rod as if Full Length, Full Power, Eouilibriun Xenon, 2000 Nd/T 4.3-13 Planar Average Power Distribution, 9000 Mad /T, Part-Length Rods as if Full Length, Full Power, Equilibrium Xenon < 3 4.3-14 Planar Average Power Distribution, E0C, Part-Length Rods as if Full Length, Full Power, Equilibrium Xenon,14000 Mwd /T
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4.3-15 Planar Average Power Disribution, BOC, Part-Length Rods as if Full Length, Bank 5 Full In, Full Power, Equilibrium Xenon, 2000 N d/T ^ O 4.3-16 Planar Average Power Distribution, 9000 hd/T, Part-Length V Rods as if Full Length, Bank 5 Full In, Full Power, Equili-brium Xenon 4.3-17 Planar Average Power Distribution, E0C, Part-Length Rods as if Full Length, Bank 5 Full In, Full Power, Equilibrium Xenon, 14000 Nd/T 4.3-18 Axial Power Distribution, B0C, Unrodded 4.3-19 - Axial Power Distribution at 50 Mwd /T, Unrodded 4.3-20 Axial Power Distribution at 4000 Mwd /T, Unrodded 4.3-21 Axial Power Distribution at 9000 Mwd /T, Unrodded 4.3-22 Axial Power Distribution at 13000 Mwd /T, Unrodded 3 , 4.3-23 Axial Power Distribution at E0C 16000 Mwd /T, Urreded 4.3-24 Planar Average Power Distribution at the Beginning of the Second Cy'le, Unrodded
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LIST OF FIGURES (Cont'd.) CHAPTER 4 Figure Subject 4.3-25 Planar Average Power Distribution at 50 N d/T of the Second Cycle, Unrodded 4.3-26 Planar Average Power Distribution, 6000 Mwd /T of the Second Cycle, Unrodded 4.3-27 Planar Average Power Distribution at the End of the Second Cycle, Unrodded 4.3-28 Planar Average Power Distribution at the Beginning of the Third Cycle, Unrodded 4.3-29 Planar Average Power Distribution at 50 N d/T of the Third Cycle, Unrodded 4.3-30 Planar Average Power Distribution at 6000 W d/T of the Third Cycle, Uniodded 4.3-31 Planar Average Power Distribution at the End of the Third Cycle, Unrodded 4.3-32 Planar Average Power Distribution at Beginning of the Fourth Cycle, Unrodded 4.3-33 Planar Average Powe.- Distribution at 50 Mwd /T of the Fourth Cycle, Unrodded 4.3-34 Planar Average Power Distribution at 6000 Nd/T of the Fourth Cycle, Unrodded 4.3-35 Planar Average Power Distribution at the End of the Fourth Cycle, Unrodded 4.3-36 Daily Reactor Power Maneuvering Near Beginning of Cycle (100% to 35% to 100% P]wer) 4.3-37 Daily Reactor Power Maneuvering Near End of Cycle (100% to 35% to 100% Power) 4.3-38 Daily Reactor Power Maneuvering Near Beginning of Cycle (100% to 50% to 100% Power) 4.3-39 Daily Reactor Power Maneuvering Near End of Cycle (100% to 50% to 100% Power) 4.3-40 Daily Reactor Power Maneuvering Near the Beginning of Cycle (2-Hour Ramps) 4.3-41 Daily Reactor Power Maneuvering Near End of Cycle (2-Hour Ramps) xvi
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LIST OF FIGURES (Cont'd) U/ CHAPTER 4 Figure Subject 4.3-42 F Q vs Time for a Load Following Transient 4.3-43 FhvsTimeforaLoadFollowingTransient 4.3-44 Normalized Power Distribution of Unshimed Assembly Used in 3 Sample DNB Analysis in Section 4.4.2.2 4.3-45 Fuel Temperature Coefficient vs Effective Fuel Temperature 4.3-46 Moderator Temperature Coefficient vs Moderator Temperature at B0C 1 4.3-47 Moderator Temperature Coefficient vs Moderator Temperature at E0C 1 4.3-48 Moderator Density Coefficient vs Moderator Density 4.3-49 Fuel Temperature Contribution to Power Coefficient at E0C 4.3-50 CEA Pattern Q < 4.3-51 CEA Bank Identification
4.3-52 Typical Power Dependent CEA Insertion Limit 4.3-53 Typical Integral Worth vs Withdrawal at Zera Power, E0C 1 Conditions 4.3-54 Typical Integral Worth vs Withdrawal at Hot Full Power, E0C 1 Equilibrium Xenca . Conditions 4.3-55 Reactivity Difference Between Fundamental and Excite States of Bare Cylindrical Reactor 4.3-56 Expected Variation of the Azimuthal Stability Index, Hot Full ! Power, No CEAs 4.3-57 PLCEA Controlled and Uncontrolled Oscillation 4.3-58 Rod Shadowing Effect vs Rod Position for Rod Insertion and
- Withdrawal Transient at Palisades V
Amendment No. 3 xvii May 28, 1981 _ . _ _ _ . _ . . _ .. _. _ . _ _ _ _ _ , . _ . _ _ _ . , _ _ ._ _ .. .__-.__,.._ _.. _ _ _ ___
LIST OF FIGURES (Cont'd) CHAPTER 4 Figure Subject 4.3-59 Typical Three Sub-Channel Annealing 4.3-60 Geometry Layout 4.3-61 Comparison of Measured and Calculated Shape Annealing Correlation for Palisades 4.3-62 Typical Temperature Defect vs Reactor Inlet Temperature 4.3-63 Calculation-Measurement ITC Different vs Soluble Boron 3D ROCS (DIT) 4.3-64 ROCS /DIT Reactivity from Core Follow Calculations,14X14 plants, Reload Cycles. 4.3-65 ROCS /DIT Reactivity from Core Follow Calculations,16X16 and 14X14 assembly Plants 4.3-66 A Divergent Axial Oscillation in an E0C Core with Reduced Power Feedback 3 4.3-67 Damping Coefficient vs Reactivity Difference Between Funda-mental and Excited State
4.3-68 Integral Radial Pin Power Distribution Used in Augmentation l Factor Determination 4.3-70 Augmentation Factor 4.4-1 Core Wide Planar Power Distribution for Sample DNB Analysis l 4.4-2 Rod Radial Power Factors in Hot Assembly for Sample DNB Analysis xviii Amendment No. 3 May 28,1981
i m 4.3 NUCLEAR DESIGN l
} \s / 4.3.1 DESIGN BASES The bases for the nuclear de. sign of the fuel and reactivity control systems are discussed in the following paragraphs.
4.3.1.1 Excess Reactfvity and Fuel Burnup The excess reactivity prrvided for each cycle is based on the depletion characteristics of the fuel and burrable poison and on the desired burnup for each cycle. The desired burnup is based on an economic analysis of the fuel cost and the projected operating load cycle for SYSTEM 80. The average burnup is chosen to ensure that the peak burnup is within the limits discussed in Paragraph 4.2.3.2.12. This design basis, along with the design basis in Paragraph 4.3.1.8, satisfies General Design Criterion 10. 4.3.1.2 Core Design Lifetime and Fuel Replacement Program The core design lifetime and fuel replacement program are based on approximately annual refueling with approximately one-third of the fuel assemblies replace at each refueling in later cycles. The first cycle design lifetime is longer than later cycles to permit a more orderly transition to equilibrium cycle conditions. 4.3.1.3 Negative Reactivity Feedback _ q"j In the power operating range, the net effect of the prompt inherent nuclear feedback characteristics (fuel temperature coefficient, moderator temperature coefficient, and modcrator pressure coefficient) tends to compensate for a , rapid increase in reectivity. The negative reactivity feedback provided by
the design satisfies General Design Criterion 11. 4.3.1.4 Reactivity Coefficients The values of each coefficient of reactivity are consistent with the design hasis for net reactivity feedback (Paragraph 4.3.1.3), and analyses that predict acceptable consequences of postulated accidents and anticipated operational occurrences, where such analyses include the response of the reactor protective system (RPS). 4.3.1.5 Burnable Poison Requirements The burnable poison reactivity worth provided in the design is sufficient to ensure that the moderator coefficients of reactivity are consistent with the design bases in Paragraph 4.3.1.4'. 4.3.1.6 Stability Criteria The reactor and the instrumentation and control systems are designed to detect and suppress xenon-induced power distribution oscillations that could, if not (~N suppressed, result in conditions that exceed the specified acceptable fuel ( ) design limits. The design limits. The design of the reactor and associated
'd systems precludes the possibility of power level oscillations. This basis
. satisfies General Design Criterion 12. 4.3-1 Amendment No. 3 May 28, 1981
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I 4.3.1.7 Maximum Controlled Reactivity Insertion Rate The core, control element assemblies (CEAs), reactor regulating system, and boron charging portion of the chemical and volume control system are designed so that the potential amount and rate of reactivity insertion due to normal operation and postulated reactivity accidents do not result in: A. Violation of the specif'., deceptable fuel des ci 'imits B. Damage to the reactor coolant pressure boundary C. Disruption of the core or other reactor internals sufficient to impair the effectiveness of emergency core cooling. This design basis, along with Paragraph 4.3.1.11 satisfies General Design Criteria 25 and 28. 4.3.1.8 Power Distribution Control The core power distribution is controlled such that, in conjunction with other core operating parameters, the power distribution does not result in violation of the limiting conditions for operation. Limiting conditions for operation and limiting safety system settings are based on the accident analyses described in Chapters 6 and 15 such that: specified acceptable fuel design limits and l3 other criteria are not exceeded for accidents. This basis, a hng with Paragraph 4.3.1.2, satisfies General Design Criterion 10. 4.3.1.9 Excess CEA Worth with Stuck Rod Criteria lhe amount of reactivity available from insertion of withdrawn CEAs under all power operating conditions, even when the highest worth CEA fails to insert, will provide for at least 2% excess CEA worth after cooldown to hot zero power, plus any additional shutdown reactivity requirements assumed in the 3 safety analyses. This basis, along with Paragraph 4.3.1.10, satisfies General Design Criteria 26 and 27. 4.3.1.10 Chemical Shim Control The chemical and volume control system (CVCS) (Subsection 9.3.4) is used to adjust the dissolved boron concentration in the moderator. After a reactor shutdown, this system is able to compensate for the reactivity changes asso-ciated with xenon decay and reactor coolant temperature decreases to ambient temperature, and it provides adequate shutdown margin during the refueling. This system also has the capability of controlling, independently of the CEAs, long-term reactivity changes due to fuel burnup and reactivity changes during xenon transients resulting from changes in reactor load. This design basis, along with Paragraph 4.3.1.9 satisfies General Design Criteria 26 and 27. 4.3.1.11 Maximum CEA Speeds Maximum CEA speeds are consistent with the maximum controlled reactivity insertion rate design basis discussed in Paragraph 4.3.1.7. Maximum CEA speeds are also discussed in Section 4.2. 4.3-2 Amendment No. 3 May 28, 1981
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O 4.
3.2 DESCRIPTION
4.3.2.1 Nuclear Design Description This section summarizes the nuclear characteristics of the core and discusses the important design parameters that affect the performance of the core in steaoy-state and normal transient operation. Summaries of nuclear design parameters are presented in Table 4.3-1, Table 4.3-2, and Figure 4.3-1. These data are intended to be descriptive of the first cycle design. Design limit values for these and other parameters are discussed in the appropriate sections. The first cycle design features a three-batch loading scheme in which the type B and C fuel assemblies contain rods of two different enrichments. In this approach, the three pins in each of the corners of every B assembly and every 3 C assembly were replaced by pins containing a lower fuel enrichment. This unique system of enrichment zoning offers improved long-term control over the local assembly power distribution. Fuel enrichment and burnable poison distributions are shown in Figures 4.3-1 and 4.3-2. The other three quadrants of the core are symmetric to the displayed quadrant. Physical features of the lattice, fuel assemblies, and CEAs are described in Section 4.2. Assembly enrichments, core burnup, critical soluble boron concentrations and worths, plutonium buildup, and delayed neutron fractions and neurton lifetime are shown in Table 4.3-1. The soluble baron insertion rates shown in this table, as discussed in Section 9.3.4, are sufficient to compensate for the maximum reactivity addition due to xenon burnout and normal plant cooldown. This maximum reactivity addition rate for which the CVCS will be required to ccmpensate is given in Table 4.3-1. The maximum value occurs for an end-of-cycle cooldown, where the moderator temperature coefficient is most negative. K ' p8Nr,, hot reactivity,andreactivitydefectdataassociatedwiththecoldzero stanby, hot full power without xenon or samarium, and hot full power, with equilbrium xenon and samarium conditions are shown in Table 4.3-2. l 4.3.2.2 Power Distribution ! 4.3.2.2.1 General At all times during operations, it is intended that the power distribution and coolant conditions be controlled so that the peak linear heat rate and the minimum departure from nucleate boiling ratio (DNBR) are maintained within operating limits supported by the safety analyses (Chapter 6 and 15) with due l3 regard for the correlations between measured quantities, the power distribution, and uncertainties in the determination of power distribution. Methods of controlling the power distribution include the use of full- or part-length CEAs to alter the axial power distribution; decreasing CEA in-nserti..# by boration, thereby improving the radial power distribution, and correcting off-optimum conditions which cause margin degradations (e.g. , CEA misoperation). C 4.3-3 Amendment No. 3 May 28, 1981 __ - - _ .
The Core Operating Limit Supervisory System (COLSS) will indicate continuously to the operator how far the core is from the operating limits and give an audible alarm should an operating limit be exceeded. Such a condition signifies a reduction in the capability of the plant to withstand an anticipated transient, but does not necessarily imply a violation of fuel design limits. If the margin fuel design limits continues to decrease, the RPS assures that the specified acceptable fuel design conditions are not exceeded by initiating a trip. The COLSS, described in Section 7.7 and Reference 32, continually generates an l3 assessment of the margin to linear heat rate and DNBR operating limits. The data requirec' for these assessments include measured in-core neutron flux data, CEA positions, and coolant inlet temperature, pressure, and flow. In the event of an alarm indicating that an operating limit has been exceeded, power must be reduced unless the alarm can be cleared by improving either the power distribution or another process parameter. The accuracy of the COLSS calculations are verified periodically as discussed in Chapter 16. In addition to the monitoring performed by COLSS, the RPS continually infers the core power distribution and thermal margin by processing reactor coolant data, signals from ex-core neutron flux detectors, each containing three axially stacked elements, and input from redundant reed switch assemblies to indicate CEA position. In the event the power distributions or other parameters are perturbed as the result of an anticipated operational occurrence that would violate fuel design limits, the high local power density or low DNBR trips in the RPS will initiata a reactor trip. The relationship between the des W power dir*ributions and the monitoring instrumentation is discussed in deta- in Reference 32. The dependence of the excore detector readings on the l3 power distributions is also detailed in Section 4.3.3.1.1. 4.3.2.2.2 Nuclear Design Limits on the Power Distribution The design limits on the power distribution stated nere were employed during the design process both as design input and as initial conditions for accident analyses described in Chapters 6 and 15. However, for the monitoring system, it is the final operating limit determination that is used to assure that the consequences of an anticipated operational occurrence or postulated accident will not be any more severe than the consequer.ces shown in Chapter 6 and 15. The initial conditions used in this operatinr, limit determ; nation are actually stated in terms of peak linear heat generation rate and required power margin for minimum DNBR. The design limits on power distribution are as follows: A. The limiting three-dimensional heat flux peaking fact, F"q, was estab-lished for full power conditions at 2.28 and 2.35 for first and equili-brium cycles, respectively. The lower va' for the first cycle reflects the presence of burnable poison shims in the fuel lattice and a corresponding reduction in the number of fuel rods. F"q is defined in Section 4.4.2.2.2.1, listing C and is termed the nuclear power factor or the total nuclear l3 peaking factor. An F"q of 2.28 in combination with uncertainties and allowances on heat flux which give the initial peak linear heat rate assumed in the safety 4.3-4 Amendment No. 3 May 28, 1981
/mi -( ,
analyses _ constituted one limiting combination of parameters for full
/
power operation in the first cycle. Other combinations of parameters which will result in acgeptable consequences of the safety analysis do exist, e.g. , a higher F q is acceptable at a reduced power level. Imple-mentation in the technical specification is via an operating limit on the monitored peak linear heat generation rate. B. The thermal margin to a minimum DNBR of 1.19 (using the C-E-1 CHF cor-relation as die sssed in Section 4.4.2.2. and 4.4.4.1), which is available to accommodate anticipated operational occurrences, is a function of several parameters, including i) the coolant conditions ii) the axial power distribution iii) the axially integrated radial peaking factor, F"r; where F"r is the rod radial nuclear factor or the rod radial peaking factor and is defined in Section 4.4.2.2.2.1, Paragraph A (referred to as Fr in that section). The coolant conditions assumed in the safety analyses, an F"r of 1.55, and the set of axial shapes displayed in Figure 4.4-3 constitute a set of limiting combinations of parameters for full power operation. Other combinations giving acceptable accident analysis consequences are equally acceptable. (mj f Implementation of these limits in the technical specificat is via an operating L/ limit on allowed mininum monitored DNBR underflow vs. measured incore axial shape index. This operating limit will be established prior to issuance of the operating license and will be cased on consideration of trany different allowed operating conditions (axial and radial power distributions as well as coolants) at any axial shape index. It will be shown in the following paragrapis that operation within these design limits is achievable. . i 4.3.2.2.3 Expected Power Distributions Figures 4.3-3 through A 3-17 and 4.3-18 through 4.3-23 show typical first cycle planar radial and unrodded core average axial power distributions, respectively. They illustrate conditions expected at full power for various times in the fuel cycle as specified on the figures. It is expected that the normal operation of the reactor will be with limited CEA insertion so that these power distributions represent the expected power distgibution during most of the cycle. The three-dimensional peaking factor, F r, expected during steady-state opergtior, is then just the product of the unrodded planar radial peakfng tactor (F q) and the axial peaking factor. The maximum expected value of F q is 1.88 during the first cycle and, as can be seen from the above figures, occurs near the beginning of-cycle for steady-state, base loaded operation with no full-length of part-length CEA insertion. Additionally, Figures 4.3-24 through 4.3-35 show typical planar radial power distributions for later cycles of System 80. Based on the similarity of these radial power distributions to those in the first cycle, it is clear that no significant differences in expected power distributior.s between cycles should exist. The (O,/ uncertainty associated with these calculated power distributions is discut in Section 4.3.3.1.2.2.6. 4 Amendment No. 3 4.3-5 May 23, 1981
The capability of the core to follow load transient without enceeding power distribution limitations depends on the margin to operating limits compared to the margin required for base loaded, unrodded operation. In order to illustrate the manuevering capability available in SYSTEM 80, the result 3 of calculations of the power distributions and power peaking factors during several load following transients are discussed below. The axial power distributions are calculated by QUIX (see Section 4.3.3.1.1), a one-dimensional spatial flux calculational model that considers the effects of the time and spatial variations of xenon and iodine concentration, full-length CEA position, part-length CEA position, thermal and moderator density feedback mechanisms, as well as the effect of the burnup distribution near end-of-cycle. Since QUIX does not have fuel depletion capability, axial-dependent depletion effects are included in end-of-cvcle calculations by using an end-of-cycle axial nuclide distribution t.n, ated frgm three-dimensional ROCS or PDQ depletion calculations. Estimates of F q and F r are obtained l3 by synthesis of the three-dimensional power distributions from two-dimensional (planar) PDQ and one-dimensional (axial) QUIX calculations. The QUIX model accepts values of radial peaking factors for each type of CEA bank insertion (unredded, bank 5 inserted, bank 5 plus PLCEA inserted, bank 4 inserted. e tc ) . These radial peaking factors are input for the appropriate core average bi ~ m condition and are applied over that axial region of the core having th . cified CEA bank configuration (e.g., unrodded, bank 5 inserted, etc.). Th a radial peaking factors are weighted by the axial power distribution to obtain an axially integrated radial peaking factor. The value cf this integrated radial peaking factor is conservatively large since the maximum of the radial peaking factors of each planar region is not, in general, eapected to occur at the same fuel pin location. The magnitude of the in,1ut radial peaking factors is determined primarily by the number and location of the inserted CEAs; it is evaluated at the full power ccndition and taken to be independent of power level. Figures 4.3-3f w ough 4.3-43 show the calculated axial power distributions and associated nuclear peaking factors during a typical day of a maneuvering transient to either 50 or 35% of the full power conditions. Also shown on these figures are the full-and part-length CEA locations during the transient. Throughout the calculation of the power distribution during these transients, it is assumed that the part-length CEAs are available for control of the axial power distribution. The part-length CEAs are moved to the position that minimized the dif ference between the current shape index and the reference value of shape index that existed prior to the initiation :f the maneuver. The detailed radial power distribution within any assembly is a function of tne location of that assembly within the core as well as the time in life, CEA insertion, etc. The normalized assembly power distribution used for the sample DNB calculation discussed in Section 4.4.2.2. is shown on Figure 4.3-44. In Section 4.3.3.1.2, the accuracy of calculations of the power distribution within a fuel assembly is discussed. 4.3.2.2.4 Allowances and Uncertainties on Power Distributions In comparing the expected power distributions and implied pe-k linear heat generation rate (PLHGR) produced by analysis with the design limits stated in Section 4.3.2.2.2, consideration must be given to the uncertainty and allowances associated with on-line monitoring by COLSS.
<.
Reference 1, a C-E Topical Reports on COLSS uncertainty assessment, contains the conclusion based on detailed numerical evaluations for cores similar to SYSTEM 80 that a penalty factor of on the order 7.5% should be applied to 4.3-6 by"b8" igg.3
.
A ( V
) COLSS determinations of F"q. The uncertainty analysis provided for SYSTEM 80 is described in Section 7.7.1.5. In addition, a power level uncertainty factor of 1.02, an engineering factor of 1.03, and an augmentation factor to account for power spiking associated with fuel densification are customarily included. The latter factor varies axially, but can be expected to have a value on the order of 1.03 at the elevation of the axial peak. It has been 3 demonstrated in Reference 32, for cores similar to SYSTEM 80, that an uncer-tainty of 4.6% is associated with the thermal margin calculation performed by COLSS.
4.3.2.2.5 Comparisons Between Limiting and Expected Power Distributions As was discussed in Section 4.3.2.2.3, Expected Power Distributions, the maximum expected unrodded F"q that cccurs during the first cycle at full l3 power is 1.88. Augmentii'g this value by the uncertainties and allowances discussed above provides an upper limit on F"q of 2.19 which is well below the design target of 2.28. Additionally, the calculations described in Section 4.3.2.2.3 show that, with proper use of the part-length CEAs, no appreciable increase in the peak linear heat rate occurs during these maneuvering transients. In the event that the part-length CEAs were not moved properly, the power distribution could have become unacceptable. In this case, the monitoring system would indicate if insufficient margin to indicate that action has tc be taken to improve the core power distribution, to improve the coolant conditions, or to reduce core power. O
's, /
Sim*larly, even allowing for +',e 1.6% uncertaingy on the monitoring of ther 1 margin, the maximum expected unrodded F r that occurs at full power
is we. below the design limit stated in Section 4.3.2.2.2 of 1.55. Again, as demonstrated by the calculations of the power distributions expected to occur during maneuvering transients, no appreciable loss in thermal margin is expected to occur during these transients. 4.3.2.3 Reactivity Coefficients ! Reactivity coefficients relate changes in core reactivity to variations in l fuel or moderator conditions. The data presented in this section and l associated tables and figures illustrate the range of reactivity coefficient l values calculated for a variety of operating and accident conditions. l3 Section 4.3.3 presents comparisons of calculated and measured moderator temperature coefficients and power doefficients for operating reactors. The good agreement shown in that subsection provides confidence that the i data presented in this section adequately characterize the SYSTEM 80 reactors. Table 4.3-3 presents a comparison of the reactivity coefficients calculated for SYSTEM 80 reactors with those used in the safety analyses described in Chapters 6 & 15. For each accident analysis, suitably l3 conservative reactivity coefficient values are used. Since uncertainties in the coefficient values, as discussed in Section 4.3.3.1.2, and other conservatisms are taken into account in the safety analyses, values used in the safety analyses may fall outside the ranges in a conservative direction of the data presented in this section. A more extensive list I p) t of reactivity coefficients is given in Table 4.3-4. V The calculational methods used to compute reactivity coefficients are discussed in Section 4.3.3.1.1. All data discussed in subsequent paragraphs Amendment No. 3 4.3-7 May 28, 1981
- _ - ._ - .- - _ ___ __ _ _ -____ _ . _ . - - - - -
cra calculated with two-dimensional, quarter-core nuclear models. Spatial distributions of materials and flux weighting are explicitly performed for the pirticular conditions at which the reactivity coefficients are calculated. Tha adequacy of this method is discussed in Section 4.3.3.1.2. 4.3.2.3.1 Fuel Temperature Coefficient Th fuel temperature coefficient is the change in reactivity per unit change in fuel temperature. A change in fuel temperature affects the reaction rates in both the thermal ano epithermal neutron energy regimes. Epithermally, the principal contributor to the change in reaction rate with fuel temperature is th: Doppler effect arising from the increare in absorption widths of the resonances with an increase in fuel temper re, nThe ensuing increase in absorption rate with fuel temperature causes a negative fuel temperature coefficient. In the thermal energy regime, a change in reaction rate with fu21 temperature arises from the effect of temperature dependent scattering properties of the fuel matrix on the thermal neutron spectrum. In typical PWR fu.'Is containing strong resonance absorbers such as U-238 and Pu-240, the mao # ude of the component of the fuel temperature coefficient arising from the Doppler effect is more than a factor of 10 larger than the magnitude of tha thermal energy component. Figure 4.3-45 shows the dependence of the calculated fuel temperature co-officient on the fuel temperature, both at the beginning and the end of the first cycle. 4.3.2.3.2 Moderator Temperature Coefficient Tha moderator temperature coefficient relates changes in reactivity to uniform changes in mode.rator temperature, including the effects of moderator density changes with changes in moderator temperature. Typically, an ircrease in the moderator temperature causes a decrease in the core moderator density and, th3refore, less thermalization, which reduces the core reactivity. However, wh:n soluble boron is present in the moderator, a reduction in moderator d:nsity causes a reduct on in the content of soluble boron in the core, thus producing a positive contribution to the moderator temperature coefficient. In order to limit the dissolved boron concentration, burnable poison rods (shims) are provided in the form of cylindrical pellets of alumnia with uniformly dispersed boron carbide particles. The number of shims is given in Table 4.3-1 and their distribution in one quadrant of the core is shown in Figure 4.3-1. The distribution is identical for the other three quadrants. The reactivity control provided by the shims is given in Table 4.3-1. This # rol makes possible a reduction in the dissolved boron concentration to the values given in Table 4.3-1. Th calculated moderator temperature coefficients for various core conditions at beginnings and end of first cycle are given in Table 4.3-4. The moderator tcmperature coefficients are more negative at end-of-cycle because the soluble boron in the coolant is reduced._ The buildup of equilibrium xenon produces a n:t negative change of -0.4 x 10 Wp/ F in the moderator temperature coefficient; this change is due mainly to the accompanying reduction in critical soluble boron. The changing fuel isotopic concentrations and the changing neutron spectrum during the fuel cycle depletion also contribute a small negative component to the moderator temperature coefficient. Amendment No. 3 4.3-8 May 28, 1981
The dependence of the moderator temperature coefficient on moderator temper-f]J t V ature at BOC and E0C (at constant soluble boron) is shown in Figures 4.3-46 and 4.3-47, respectively. These figures also show the expected moderator temperature coefficient at reduced power levels (corresponding to reduced moderator temperature) based on power reductions accomplished with soluble borcn only and with CEAs only. These two modes of power reduction result in the most productive and most negative moderator temperature coefficients expected to occur at reduced power levels. These figures show the expected moderator temperature coefficient for the full range of expected operating conditions and accident conditions addressed in Chapter 15. 4.3.2.3.3 Moderator Density Coefficient The monisator density coefficient is the change in reactivity per unit change in the a,erage core moderator density at constant moderator temperature. A positive moderator density coefficient translates into a negative contribution to the total moderator temperature coefficient, which is defined in Section 4.3.2.3.2. The density coefficient is always positive in the operating range, althoagh the magnitude decreases as the soluble baron level in the core is increased. The calculated density coefficient is shown in Table 4.3-4, and curves of density coefficient density coefficient as a function of density for several soluble boron concentrations are presented in Figure 4.3-48. These curves are based upon 2-D PDQ calculations and have been generated over a wide range of core conditions. The density coefficients explicity used in the accident analyses are based upon core conditions with t'le most limiting temperature [) coefficients allowed by the technical specification. Table 4.3-3 shows a C/ comparison of the expected values of the moderator temperature coefficients with those actually used in the accident analyses. 4.3.2.3.4 Moderator Nuclear Temperature Coefficient j The moderator nuclear temperature coefficient is the change in reactivity per unit change in core average moderator nuclear temperature, at constant moderator density. The source of this reactivity dependence is the spectral effects associated with the change in thermal scattering properties of water molecules as the internal energy, which is represented by the bulk water temperature, is { changed. The magnitude of the moderator nuclear temperature coefficient is i equal to the difference between the moderator temperature coefficient, defined
- in Section 4.3.2.3.2, and the moderator density coefficient, defined in
! Section 4.3.2.3.3. i 4.3.2.3.5 Moderator Pressure Coefficient The moderator pressure coefficient is the change in reactivity per unit change in reactor coolant system pressure: Since an increase in pressure, at constant moderator temperature, increases the water density, the pressure coefficient is merely the density coefficient expressed in a different form. The calculated pressure coefficient at full power is shown in Table 4.3-4.
I 4.3.2.3.6 Moderator Void Coefficient 1 , (m
'#
) The anticipated occurrence of small amounts of local subcooled boiling in the reactor during full power operation results in a predicted core average steam Amendment No. 3 4.3-9 May 28, 198
i (void) volume fraction of substantially less than 1%. Changes in the moderator void fraction produce reactivity changes that are quantified by the void coefficient of reactivity. An increase in voids decreases core reactivity, but the presence of soluble boron tends to add a positive contribution to the coefficient. The calculated values of moderator void coefficient are shown in Table 4.3-
- 4. Curves showing void coefficient vs. void content can be inferred directly from the density coefficient curves provided in Figure 4.3-48.
4.3.2.3.7 Power Coefficient The power coefficient is the change in reactivity per unit change in ccre power level. All previously described coefficients contribute to the power coefficient, but only the moderator temperature coefficient and the fuel temperature coefficient contributions are significant. The contributions of the pressure and void coefficients are negligible, because the magnitudes of these coefficients and the changes in pressure and void fraction per unit change in power leval are small. The contribution of moderator density change is included in the moderator temperature coefficient contribution. In order to determine the change in reactivity with power, it is necessary to know the changes in the average moderator and effective fuel temperature with power. The average moderator (coolant) temperature is controlled to be a linear function of power. O i l
%
4.3-10 a"y"N$"k9I
__- The core average linear heat rate is also linear with power. The average
'
y/ effective fuel temperature dependence on the core average linear heat rate is calculated from the following semi--empirical relation: 2 3 Tf=TMOD + (E B 4 I
*M ) *P + (I C. *MI )
- p 2 (j) i=o J J=0
'
T is the average moderator temperature ( F), M is the exposure in MWD /T, PM N the linear heat generation rate in the fuel in KW/ft, and T y is the average effective fuel temperature (*F). The coefficients B. and C. are determined from least squares fitting of the fuel temperaturd generdted by FATES. For a System 80 fuel pin, the following values apply: B o
= 146.526 Cg = -2.0355 B) = 0.8841
- 10-3 C) = -0.5121
- 10-3 B
2
= -0.2052
- 10 -6 C 2 = 0.5043
- 10-7 C3 = -0.1071
- 10 The basis for this relation is discussed in Section 4.3.3.1.2.2.4.
i The total power coefficient at a given core power can be determined by p ,, evaluation, for the conditions associated with the given power level, the ( following expression: U) dP
- aP -
aT f 3p , aT,
~dp aT ap aT 8P
(} f m The first term of the equation (2) provides the fuel temperature contribution to the power coefficient, which is shown as a function of power in Figure 4.3-49. The first factor of the first term is fuel temperature coefficient of reactivity discussed in Section 4.3.2.3.1 and shown in Figure 4.3-45. The second factor of the first term is obtained by calculating the deriative of equation (1). aT f I
= ($ B'.
- M ) + 2($ C 3 I
*M)*P (3) ap i=0 J=0 The second term in Equation (2) provides the moderator contribution to the power coefficient. The first factor, the moderator temperature coefficient, is discussed in Section 4.3.2.3.2 and shown in Figure 4.3-46 and 4.3-47.
The second factor is a constant since the moderator temperature is controlled to be a linear function of power. Since the factors aP/aT and aP/aT ar i variables; e.g., burnup,f temperatufe, esoluble functions of content, boron one or more xenonindependent worth , i and CEA insertion, the total power coefficient, dP/dp, also depends on gd these variables. 41endment No. 3 4.3-11 May 28,1981
The power coefficient tends to become more negative with burnup because the fuel and moderator temperature coefficients become more negative (see Figure 4.3-45 through 4.3-47). The insertion of the CEAs, while maintaining constant power, results in a more nenative power coefficient, because the soluble boron level is reduced and because of the spectral effects of the CEAs. The full power values of the overall ouwer coefficient for the unrodded core at BOC and E0C are shown in Table 4.3-4. f.3.2.4 Control Requirements There are three basic types of control requirements that influence the d sign of this reactor: A. Reactivity control so that the reactor can be operated in the unrodded critical, full power mode for the design cycle length. B. Power level and power distribution control so that the reactor power may be safely varied from full-rated power to cold shutdown, and so that the power distribution at any given power level is controlled within acceptable limits. C. Shutdown reactivity control sufficient to mitigate the effects of postulated accidents. R activity control is provided by several different means. The amount and enrichment of the fuel and burnable poison shims are design variables that d:termine the initial and end-of-cycle reactivity for an unrodded, unborated condition. Soluble boron and CEA poisons are flexible means of controlling long-term and short-term reactivity changes, respectively. The following paragraphs discuss the reactivity balances associated with each type of control requirement. 4.3.2.4.1 Reactivity Control at BOC anJ EOC The reactivities of the unrodded core with no soluble boron are shown in Table 4.3-2. This table includes the reactivity worth of equilibrium xenon and samarium, and shows the reactivity available to compensate for burnup and fission product poisoning. Soluble boron concentrations required for criticality at various core conditions are shown in Table 4.3-1. Soluble boron is used to compensate for slow reactivity changes such as those due to burnup, changes in xenon content, etc. The reactivity controlled by burnable poison shims is also giver in Table 4.3-1. At EOC, the reactivity worth of the residual poison is less than 1%, and the soluble boron concentra-tion is near zero. The reactor is to be operated in essentially an unrodded condition at power. The CEA insertion at power is limited by the power dependent insertion limit (PDIL) for short-term reactivity changes. 4.3.2.4.2 Power t.evel and Power Distribution Control The regulating CEA groups may be used to compensate for changes in reactivity associated with routine power level changes. In addition, regulating CEAs nay be used Amendment No. 3 4.3-12 Mr< 28, 1981
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'
' to compensate for minor variations in moderator temperature and boron concentrations during openation at power, and to dampen Axial Xenon oscillations. The reactivity worth of regdai.ing CEA control groups is shown in Table 4.3-5. Soluble boron is used to maintain shutdown reactivity at cold zero power conditions. The soluble boron can also be used to compensate for changes in reactivity due to power level changes and minor changes in reactivity which might occur during normal reactor operation. Thirteen part-length CEAs are provided in the design to help control the core power distribution. The function includes the suppression of xenon induced axial power oscillations. 4.3.2.4.3 Shutdown Reactivity Control The reactivity worth requirements of the full complement of CEAs is primarily determined by the power defect, the excess CEA worth with the stuck rod criteria discussed in Section 4.3.1.9. Table 4.3.6 shows the reactivity component allowances that define the total reactivity allowance. These data are based on the end-of-cycle conditions when the fuel and moderator temperature coefficients are the most negative and thus when the shutdown reactivity requirement is a maximum. Each allowance component is further discussed below. No CEA allowance is provided for xenon reactivity effects, e.g., undershoot, since these effects are controlled with soluble boron rather than with CEAs. The worth of all CEAs except the most reactive, which is assumed stuck in
'"
3 the fully withdrawn position, provides more shutdown capability than required by the total reactivity allowance shown in Table 4.3-6. This margin is shown in Table 4.3-7 for end of an equilibrium cycle. The excess CEA worth in cycle one is somewhat greater. The margin is more than sufficient to compensate for calculated uncertainties in the nominal design allowances and in the CEA reactivity worth. %us, the shutdown reactivity control provided in this design is sufficient at all times in the cycle. 4.3.2.4.3.1 'uel Temperature Variation. The increase in reactivity that occurs when the fuel temperature decreases from the full power value to the zero power value is due primarily to the Doppler W ect in U-238. The CEA reactivity allowance for fuel temperature variation shown in Table 4.3-6 is a conservative allowance for the end-of-cycle conditions. Measure-ments of first cycle power coefficients at Omaha, Calvert Cliffs, and Millstone-2 lead to a power defect of 1.2% Ap (paragraph 4.3.3.1.2.2.4). The slight increase in power defect with exposure due to the presence of plutonium isotopes is offset by the reduction in the fuel temperature resulting from fuel swelling and clad creep-down. 4.3.2.4.3.2 Moderator Temperature Variation. The moderator temperature variation allowance is large enough to compensate for any reactivity increase that may occur when the moderator temperature decreases from the full power l3 value to the zero power (hot standby) value. This reactivity increase, which is primarily due to the negative moderator temperature coefficient, is largest at the end-of-cycle when the soluble boron concentration is near
~
zero and the moderator coefficient is strongly negative. At beginning-of-cycle, when the moderator temperature coefficient is less negative, the (' < reactivity change is smaller. Amendment No. 3 4.3-13 May 28, 1981
The CEA reactivity allowance for moderator ten perature variation given in Table 4.3-6 is actually the sum of three allowances. The first, and most important, is the allowance for the moderator :emperature coefficient effect. The second is an allowance for the reihtction in CEA worth resulting from the shorter neutron diffusion length at the zero power moderator density relative to the full power moderator density. This allowance is necessary because the CEA worths shown in Table 4.3-5 were calculated at full power. The third allcwance is intended to cover the reactivity effects associated with the greatest expected axial flux redistribution resulting from the difference in moderator temperature profile between full and zero power, and the asymetric axial burnup distribution at E0C. 4.3.2.4.3.3 Moderator Voids. Reducing the power level from full power to zero power causes an increase in reactivitv resulting from the collapsing of steam bubbles caused by local boiling at full power. The amount of void in the core is small and is estimated to be substantially less than 1% at full power. As with the moderator temperature effect, the maximum increase in reactivity from full to zero power occurs at end-of-cycle, when the least amount of dissolved boron is present. The reactivity effect is small, and allowance for this effect is shown in Table 4.3-6. 4.3.2.4.3.4 Control Element Assembly Bite. The CEA bite is the amount of reactivity worth in CEAs that can be inserted in the core at full power to initiate ramp changes in reactivity associated with load changes, and to compensate for minor variations in moderator temperature, boron concentration, xenon concentration part-length CEA (PLCEA) movement, and power level. The reactivity allowance for this effect is shown in Table 4.3-6. 4.3.2.4.3.5 Part-Length CEA Effects. No reactivity allowance is provided or required for the PLCEAs because of their unique design. The l3 PLCEAs have a strong absorber section 4(B C) in the top 10% of their active length as described in paragraph 4.2.2.4. During normal operation with PLCEAs, the strong absorber section is not inserted into the active core since PLCEAs are not inserted more than 90% of the active core height. On reactor trip, the PLCEAs insert fully such that the strong absorber section is the top 10% of the active core. At E0C, where PLCEA reactivity effects are greatest, the negative reactivity insertion due to the strong absorber section will more than offset any positive reactivity insertion due co the shift in axial flux distribution between full and zero power. This flux redistribution occurs due to the asymmetric axial burnup distribution resulting from the integrated burnup effects of a negative moderator tempera-ture coefficient. Its effect is to increase the reactivity worth of the top of the core relative to the bottom of the core on reactor trip, thus increasing the importance of the 10% length strong absorber (B4 C) section of the PLCEAs as they insert on the reactor trip. 4.3.2.4.3.6 Accident Analysis Allowance. The allowance shown in Table 4.3-6 for accident analysis is consistent with that assumed under various postulated accident conditions addressed in Chapter 15, which result in predicted acceptable consequences. 4.3.2.4.3.7 Available Reactivity Worth. Table 4.3-7 shows the reactivity worths of the full complement of CEAs, and the highest reactivity worth of a single CEA in the fully withdrawn position, at end-of-equilbrium-Amendment No. 3 4.3-14 May 28, 1981
-. . __
p cycle. This table also compares the available net shutdown worth (including i J the effects of the stock CEA) to the reactivity worth requirements from D Table 4.3-6. As discussed in Section 4.3.3, the uncertainty in total CEA reactivity worth is 10% and the uncertainty in the stuck CEA worth is less than 10% Even allowing for the maximum calculated errors for total CEA worth in both the adverse directions, sufficient excess CEA reactivity in available. 4.3.2.5 Control Element Assembly Patterns and Reactivity Worths The locations of all CEAs are shown in Figure 4.3-50. The CEAs designated as regulating control rods are divided into five groups; the shutdown CEAs are divided into two groups; and the PLCEAs are divided into two groups. These groups are identified, for first cycle operation, in Figure 4.3-51. All CEAs in a group are withdrawn or inserted quasi-simultaneously. Shutdown groups are inserted after the regulating groups are inserted and are withdrawn before the regulating groups are withdrawn. The reactivity worths of sequen-tially inserted CEA groups are shown in Table 4.3-5 near the beginning and end of the first cycle where the maximum rod radial peaking factors (FQ) for these con- 3 figurations occur. The values of Fp for these times are shown in Table 4.3-8. It is expected that the core will be essentially unrodded during the full power steady-state operation, except for limited insertion of the first regulating group in order to compensate for minor variations in moderator temperature and boron concentration. For operation with substantial CEA insertion, the relationship between power level and the maximum permitted V CEA insertion is typified in Figure 4.3-52. This figure also illustrates ' the regulating group insertion order (5-4-3-2-1) and the 40% fixed overlap between successive regulating groups. Compliance with the power dependent insertion limits throughout the cycle insures that adequate shn.down margin is maintained and that the core conditions are no more severe than the initial conditions assumed in the accident analyses described in Chapter 15. Reactivity insertion rates for the safety aralysis of the core are presented in Chapter 15. The full power CEA ejection accident considers the ejection on one CEA from a fully inserted lead bank The ejected CEA worth is calculated by the difference between the tre-ejection and post-ejection reactivity of the core computed by static methods, and the effect of PLCEA insertion is explicity considered in these calculations. The maximum ejection CEA worth at hot full power used in the safety analysis is con-servative since the lead regulating bank is not expected to be fully inserted at full power. A similar analysis is performed for the CEA ejection analysis from zero power, except that the initial condition for this incident assumes that all regulating banks are fully inserted. The CEA withdrawal incident from low power is analyzed with the maximum calculated differential reactivity insertion rate resulting from a sequential CEA bank withdrawal with 40% overlap. The CEA withdrawal incident from full power is analyzed from the insertion of the lead bank which maximizes the reactivity irsertion and the power shape change during the CEA withdrawal. ReacHvity insertion rates are calculated by a static axial model of the v Amendment No. 3 4.3-15 May 28, 1981 _- - - - . .
System 80 core. The calculated reactivity insertion rate resulting from a sequential CEA withdrawal is presented in Figures 4.3-53 and 4.3-54. The full-length CEA drop incident is analyzed by selecting the dropped CEA that maximizes the increase in the radial peaking factor. A conservatively small negative reactivity insertion is used in the accident analysis. The part-length CEA subgroup drop incident is analyzed parametrically for various initial axial power shapes so that the combination of axial power shape and the inserted reactivity lead to the most adverse analyses results. A conservative reactivity insertion is used in a single PLCEA drop incident. This assumed reactivity is conservatively larger than the expected reactivity resulting from a single PLCEA dropping to the bottom of the core. The typical reactivity insertion during a reactor scram is presented in Section 15.0. This reactivity insertion is computed by static axial models at various scram CEA positions, and it is used for all accidents which are terminated by a scram, unless otherwise indicated. The reactivity insertion is conservative since (1) only the minimum shutdown worth of 10.0% Wp is assumed to be available at hot full power, and (2) the influence of delayed neutrons on the transient power shape is neglected. The scram reactivity insertion for the loss of flow is implicit in the kinetic axial analysis. 4.3.2.6 Criticality of Reactor During Refueling The soluble boron concentrations during refueling are shown in Table 4.3-1. These concentrations ensure that the keff f the core during refueling does not exceed 0.95. 4.3.2.7 Stability 4.3.2.7.1 General Pressurized water reactors (PWRs) with negative overall power coefficients are inherently stable with respect to power oscillations. Therefore, this discussion will be limited to xenon induced power d?stribution oscillations. Xenon induced oscillations occur as a result of rapid perturbations to the power distribution which cause the xenon and iodine distributions to be out of phase with the perturbed power distribution. This results in a shift in the iodine and xenon distribution that causes the power distribution to change in an opposite direction from the initial perturbation and thus an oscillatory condition is established. The magnitude of the power distribution oscillation can either increase or decrease with time. Thus, the core can be considered to be either unstable or stable with respect to these oscilla-tions. Discussed below are the methods of analyzing the stability of the core with respect to xenon oscillations. The tendency of certain types of oscillations to increase or to decrease is calculated, and the method of controlling unstable oscillations is presented. 4.3.2.7.2 Method of Analysis Xenon oscillations may be analysed by two methods. The first method consists of an explicit analysis of the spatial flux solution accounting for the space-time solution of the xenon concentrations. Such a method is useful Amendment No. 3 4.3-16 May 28, 1981
-- -- . . - - - - _ -
b for testing various control strategies and evaluating transitional effects (such as power maneuvers). The second methoo Sonsists of modal perturbation i pI g theory anaiysis, which is useful for the evaluation of the sensitivity of V the stability to changes in the reactor design characteristics, and for the determination of the degree of stability for a particular oscillatory mode. , The stability of a reactor can be characterized by a stability index or a damping factor which is defined as the natural exponent which describes the growing or decaying amplitude of the oscillation. A xenon oscillation may be described by the following equation. i
'
bt 4(F,t) = 4,(F) + A4,(F) e sin (wt + 6) where 4(F,t) is the space-time solution of the neutron flux 4,(F) is the initial fundamental flux I A4,(F) is the perturbed flux mode b is the stability index w is the frequency of the oscillation
!
j 6 is a phase shift
-
Modal analysis consists of an explicit solution of the stability index b using known fundamental and perturbed flux distributions. A positive stability index b indicates an unstable core, and a negative value indicates
. stability for the oscillatory mode being investigated. The stability index
! is generally expressed in units of inverse hours, so that a value of - L 0.01/h would mean that the amplitude of each subsequent oscillation cycle
'
decreases by about 25% (for a period of about 30 hours for each cycle). Xenon oscillations modes in PWRs can be classified into three general
- tyr.es
- radial, aximuthal, and axial. To analyze the stability for each i oscillation mode, only the first overtone needs to be considered since
! higher harmonic modes decay more rapidly than the first overtone. Further- , more, since the first overtone of a radial oscillation decays more rapidly l than the first overtone of an azimuthal oscillation, only the latter of ! these tow modes will be considered in detail. I ( 4.3.2.7.3 Expected Stability Indices j 4.3.2.7.3.1 Radial Stability. A radial xenon osdllations consists i of a power shift inward and outward from the center of the core to the i periphery. This oscillatorv m'de is generally more stable than an azimuthal mode. This effect is illustrated in Figure 4.3-55, which shows that for a bare cylinder the radial mode is more stable than the aximuthal mode. j () O Discussion of the stability for radial oscillatory mode is therefore deferred to the azimuthal mode. 1 Amendment No. 3 i 4.3-17 May 28, 1981
l l 4.3.2.7.3.2 Azimuthal Stability. An azimuthal oscillation consists of a X-Y power shift from one side of the reactor to the other. Modal analysis for this type of oscillation is performed for a range of expected r:: actor operating conditions. The expected variation of the stability index during the first cycle is shown in Figure 4.3-56. These results are obtained from analyses which consider the spatial flux shape changes during the cycle, the changes in the moderator and Doppler coefficient during the cycle, and the change in x non and iodine fission yield due to plutonium buildup during the cycle. As is shown_9n the figure, the expected stability index is no greater than -0.04h at any time during the cycle for the expected mode of reactor cperation. Comparison of predicted stability index with those actually measured on operating cores, as discussed in Section 4.3.3.2.3, provide a high confidence level in the prediction of aximuthal stability. Measure-ments of xenon spatial stability in large cores have been made (Reference 33) l3 which provide confidence in the methods that are used to predict the czimuthal stability of this core. 4.3.2.7.3.3 Axial Stability. An axial xenon oscillation consists of l3 e power shift toward the top and bottom of the reactor core. This type of oscillation may be unstable during the first cycle. Table 4.3-9 shows the calculated variation of the axial stability index during the first cycl.e. It is anticipated that control action with part-length rods and/or full-Icngth rods may be required to limit the magnitude of the oscillation. As discussed in Section 4.3.2.2, the axial power distribution is monitored by COLSS and the RPS. Based on the COLSS measurement of the axial power distribution, the operator may move either the full-length or the part length CEAs so as to control any axial oscillations. 4.3.2.7.4 Control of Axial Instabilities The control of axial oscillations during a power maneuver is illustrated in Figures 4.3-36 through 4.3-43. PLCEAs are used throughout these maneuvers to limit the change in the power distribution. The difference between an uncontrolled and a controlled xenon oscillation is illustrated in Figure 4.3-57. It was assumed in the calculation of the controlled oscillation that the PLCEAs wer moved in such a way as to preserve the initial shape in tha core prior to the initiating perturbation. The calculations are performed at the end of the first cycle which corresponds to the expected least stable condition for axial xenon oscillations. 4.3.2.7.5 Summary of Special Features Required by Xenon Instability Tha RPS described in Section 7.2.2 is designed to prevent exceeding accept-cble fuel design limits and to limit the consequences of postulated acc-d:nts. In addition, a means is provided to assure that under all allowed op; rating modes, the state of the reactor is confined to conditions not more severe than the initial conditions e~sumed in the design and analysis of the protective system. Since the reactor is predicted to be stable with respect to radial and azimuthal xenon l3 oscillations, no spec;al protective system features are needed to accommodate Amendment No. 3 4.3-18 May 28, 1981
\v) radial or azimuthal mode oscillations. Nevertheless, a maximum quadrant tilt is prescribed in the technical specifications along with prescribed operating restrictions in the event that the tilt is exceeded. The azimuthal power tilt is determined by COLSS and included in the COLSS determination of core margin. The azimuthal power tilt limit is accounted for in the RPS.
4.3.2.7.5.1 Features Provided for Azimuthal Xenon Effects A. Administrative limits on azimuthal power tilt B. Monitoring and indicating the azimuthal power tilt in COLSS as well as accounting for this tilt in the COLSS determination of core margin C. Accounting for azimuthal power tilt limit in the RPS. 4.3.2.7.5.2 . Features Provided for Axial Xenon Effects and Power Distribution Effect and Control A. PLCEAs or regulating CEAs for control of the axial power distribution, if required B. Monitoring and accounting for changes in the axial power distribution in COLSS C. Monitoring and accounting for the axial power distribution in the RPS
- 4.3.2.8 Vessel Irradiation The design of the reactor internals and of the water annuius between the active core and vessel wall is such that for reactor operation at the full power rating and an 80% capacity factor, the ves gl flugnce greater than 1 MeV at the vessel wall will not exceed 3.15 x 10 n/cm over the 40 year design life of the vessel. The calculated exposure includes a 10% uncertainty factor.
The maximum fast neutron fluxes greater than 1 MeV incident on the vessel ID and shroud ID cre as shown in Table 4.3-1s. The fluxes are based on a time averaged equilibrium cycle radial power distribution and an axial power distribution with a peak to average of 1.20. The models used in these calculations are discussed in Section 4.3.3.1 i l O '
- Amendment No. 3 4.3-19 May 28, 1981 l
4.3.3 ANALYTICAL METHODS 4.3.3.1 Reactivity and Power Distribution 4.3.3.1.1 Method of Analysis The nuclear design analysis of low enrichment PWR cores is based on a combination of multigroup neutron spectrum calculations, which provide cross-sections appropriately averaged over a few broad energy groups, and few group one, two, and three dimensional diffusion theory calculations of integral and differential reactivity effects and power distributions. Most of the calculations are performed with the aid of computer programs embodying analytical procedures and fundamcntal nuclear data consistent with the current state of the art. Comparisons between calculated and measured data that validate the design procedures are presented in paragraph 4.3.3.1.2. As improvements in analytical procedures are developed, and validation by comparison with related experimental data is performed, they are incorporated into the design procedures. 4.3.3.1.1.1 Cross Section Generation Few group cross sections for the various cell types or subregions of the core, such as fuel cells, burnable absorber cells, instrument and water hole cells, are g ated by the two dimensional integral transport assembly code DIT This code represents explicitely the heterogeneous geometry of the fuel assembly, and combines 85 group spectrum calculations in single cells or groups of cells (spectrum geometrics), with a four group assembly- ' wide flux distribution calculation. The multigroup spectrum calculation is performed for the heterogeneous cells in each spectrum geometry and uses integral transport theory based on the collision probability methods orginally developed by Carivik (2) with multigroup interface currents used to couple adjacent cells. Spectrum interactions between adjacent cells of a different type are thus explicitely taker into account. 3 The spatial mesh within cells may be both radial and azimuthal. Most frequently, one uses a radial mesh combined with an expansion in harmonics azimuthally. Following the spectrum calculations, each cell cross sections are collapsed to few groups (normally four), and an assembly-wide few group flux calculation is performed. Upon completion of the assembly calculation, excess reactivity (i.e., reactivity not accounted for by chemical shim) is r* roved by leakage in an 85 group B1 calculation with a buckling search. Thus, criticality is always maintained and the few group averaged cross-sections produced are always obtained in a critical spectrum. The DIT code utilizes a data library containing multigroup cross sections, fission spectra, fission product yields and other supplemental data. The principal source of data for the library is ENDF/B-IV. Three adjustments to the library data have been made to reflect changes to ENDF/B-IV recommended by the Cross Sections Evaluation Working Group (CSEWG) for incorporation into ENDF/8-V. Amendment No. 3 4.3-20 May 28, 1981
gy These adjustments include: I i V 1. A reduction of about 3% in the shielded resonance integral of U238
- 2. The adaption of the harder Watt fission spectra for U235 and PU239, later incorporated in ENDF/B-V.
- 3. A moderate upward adjustment of U-235 and Pu-239 thermal U-values of about 0.1% improving the G, n discrepancy but not going as far as ENDF/B-V.
In the epithermal region, the ENDF/B-IV files are processed with ETOG (3) to provide cross section resonance parameters and scattering matrices for the isotopes contained in the library. ET0G prepares this data in 99 energy groups spaning the range from 14.9 MeV to 0.414 eV. The GAM portion of GGC3 (4) is used to condense the 99 group data into 50 energy groups spanning the energy range 14.8 MeV to 1.855 eV weighted with a spectrum representative of that in a PWR assembly. In the resolved energy (9.1 kev to 1.855 eV), the capture and fission cross sections of resonant absorbers are replaced with resonance tables. In the thermal region, the ENDF/B-IV files are processed with FLANGE II (5) to provide cross sections and full scattering matrices in the thermal region (1.855 eV to 0.00025 eV). The cross sections of isotopes containing resonances in the thermal region are Doppler broadened. For hydrogen,
/ ; scattering matrices are prepared with FLANGE II using ENDF/B-IV thermal '\ ,/ scattering low parameters for H20.
The cross sections and scattering matrices are tabulated on the library for a sufficient number of temperatures to span the range expected during power reactor operation and te permit linear interpolation. 3 Cross sections for the resalved resonance region (9.1 kev to 1.355 eV) are prepared with C-E RABBLE, an extension of the RABBLE (6) code using resolved resonance parameters from ENDF/B-IV. The cosine current approximation in RABBLE was replaced with an integral transport routine. Group averaged resonance cross sections are generated with the modified RABBLE code which performs a space dependent calculation of the slowing down sources. The cross sections from the C-E RABBLE calculations are corrected to include the proper groun dependent smooth calculations which are derived from the FT0G/GGC3 calce.;tions. RABBLE is also used to validate interference effects among wonance absorbers as calculated by the DIT algorithm. Following the assembly spectrum calculation, a depletion time step takes place for each individual pin in the assembly and, when required for sub-divisions of a pin. At the end of the depletion step, new isotopic compositions are defined for use in the spectrum calculation of the next time step. This process is extended over the expected life of the fuel assembly. Editing routines provide cross sections for both fine mesh PDQ and coarse mesh tbdal codes. For fine mesh codes, all the non-fuel cell cross sections
/"'N are fitted in such a way that the diffusion calculations will preserve the .
reaction rates predicted by the transport code DIT. For those fuel cells () l ! Amendment No. 3 4.3-21 May 28, 1981
adjacent to the water holes, where spectrum interactions between fuel and water hole result in a substantial softening of the spectrum, cross sections are edited separately, and the thermal diffusion coefficient is fitted so that the water hole peaking factor calculated by the fine mesh PDQ model catches the DIT peaking value. For Control Element Assemblies (CEA), boundary conditions ano equivalent diffusion theory constants for individual elements of a CEA are calculated by the CERES program. For a one region CEA in cylindrical geometry, bouadary conditions are calculated in each multi group by the method of successive generations, with capture probabilities based on the tabulations of Stuart and Woodruff. (7) Two region CEAs are transformed to fictitious homogeneous CEAs by matching extrapolation lengths on the outer surface, as defined by Kear and Ruderman. (8) The homogeneous CEAs are then treated es above. Fictitious few group diefusion pa'ameters for use in multi-dimensional diffusien, theory calculations are obtained using methods defined by Wachspress (9) and Henry. (10) 4.3.3.1.1.2 Fine Mesh Methods. Fine mesh calculations are performed when detailed pin wise information is needed, such as fuel oin power distribution or in-core detector response. These calcul g ns are perfo g by the PDQ-X code, which is an extension of the PDQ-7 and Harmony programs, with the following added capabilities.
- 1) Pointwise Xenon feedback
- 2) Pointwise Fuel Temperature (Doppler) feedback.
A mesh rectangle is assigned to each pin cell. The inter-assembly wat' .- channel the core shroud and core barrel are each represented explicitly. 3 Because of the large number of mesh points needeu in such a model, PDQ-X is only used to perform two dimensional planar calculations, leaving all three dimensional treatments to coarse mesh models. The HARMONY scheme of cross section organization is used to input cross sections to PDQ-X. Cross section sets are edited from DIT for each cell type, i.e. for each fuel enrichment, water hole, depletable shim, in-core instruments etc., with a special set applicable to fuel cells adjacent to water holes. Through a complex sequence of DIT spectral and depletion crlculations, microscopic cross sections are generated and arrayed in HARMONY as a function of local exposure, fuel temperature, moderator temperature, soluble boron concentration and fuel enrichment. Interpolation is allowed against any one of these variables, taking full advantage of the multi-dimensional inter polation scheme of HARMONY. For non-fuel cells, fast neutron fluence is substituted for burnup as the interpolating variable. Use of the fitting procedures performed by the DIT cross-section processing routines, as described early, ensures improved accuracy in the prediction of local reaction rates and peaking factors in the planar PDQ-X calculation. The Doppler feedback option in PDQ-X makes use of the fuel temperature dependence of the cross sed ions built into the few group cross-sectico 4.3-22 b*y"NES"k9g.3
o [QW'\ array, together with a correlation relating local expoture and power density to fuel temperature. (13) This correlation is derived from the results of FATES (14) calculation, and takes into account the reduction of the fuel temperature with expowre, as a result of clad creep and pellet swelling and relocation.
!
In a large PWR core, the calculated power distributions are fairly sensitive i to the treatment of the reflector cross-sections. Terney (15) has compared transport and diffusion calculations of the albedo and shown that the latter substantially underpredicts the reflector albedos in the fast (top) group and !~tt the power distribution is shifted toward the core center when compared to multi group transport theory results. When the fast diffusion coefficients in the reflector are altered to make the transport and diffusion theory albelos m ?e, the power distributions are also brought into agreement. ap- er ^ The depletion of the core is accomplished by utilizing a set of linearized depletion chains specifying the coupling (neutron capture or fission or f decay) between the isotopes in the chain. Except for I-135, Xe-135 and Sm-149, the fission products are added together into a single lumped fission product. The exposure intervals of the depletion calculations are usually , chosen to be 1000 MWD /T during which it is assumed that the flux and micro-scopic cross-sections are unchanged. At the end of each exposu:e interval, the fluxes and s oss-sections are recalculated. Shorter exposure intervals (] j are used for recalculation of cross-sections which vary rapidly with depletion, sah as shim boron. ('" 3
;
4.3.3.1.1.3 Coarse Mesh Methods For those applications where three dimensional reactivity ard power distribution effects are important, the Reactor Operation and Control Simtlation (10,17) (ROCS) code is employed. ROCS is a digital computer code system for coarse-mesh, two-or three-dimensional neutronic analysis of nuclear reactor cores. The ROCS code was developed as a tool for core analysis which provides realistic trade-offs between accuracy and computational speed. To achieve this objective, the code incorporates a higher order expansion solution method for the flux based on two group diffusion theory. This method achieves a high degree of accuracy relative to traditional fine-mesh diffusion theory calculations, while requiring a low number of spatial mesh points. The kOCS neutreaic calculation is performed in two energy groups, using four nodes per fuel assembly. Albedo-like boundary conditions are used
,
instead of an explicit reflector representatian. The cross-section treatment is comparable to that of PDQ-X, in the sense that a generalized cross-section structure provides the flexibility needed for depletion calculations as well as thermal hydraulic feedback calculations.
,
A , I i The ROCS thermal-hydraulic calculation model represents the reactor core as V an array of closed, parallel flow channels. Each radial node corresponds Amendment No. 3 4.3-23 May 28, 1981
- .-_ _ - . - . -. -
1 l l to an individual flow channel, and the channels are axiu. y segmented for each ROCS plane. The thermal-hydraulic solution is thus discrete over the entire coarse-mesh nodal geometry for the active core. The thermal-hydraulic calculation basically consists of a steady state heat balance calculation for moderator enthalpy, and calculation of average fuel temperature in each segment using a burnup and power dependent correlation These calculations are used to provide the principal feedback mechanisms in the flux calculation. The steady-state heat balance calculation requires the specification of channel inlet conditions, and channel mass flow rate. The inlet values can be input as either uniform or disturbed values over all channels. The ROCS thermal-hydraulic module provides additional nodal calculations for user edit information independent of the feedback calculation. These include pressure drop, fluid quality and heat transfer information. The pressure drip edit includes components due to friction., momentum and gravity. The heat transfer information is based on calculations for a node-average pin using the Fourier heat conduction equacion and standard heat transfer correlations for subcooled forces convection, and for nucleate and bulk boiling. The ROCS system performs coarse-mesh depletion calculations for each node in a two-or three-dimensional core configuration. The depletion chains are internally modeled with fixed depletion equations so that beyond the input 3 cross-section data the user need supply only such data as initial concentra- + ions, decay constants and fission yields for each depletion nuclide. As the size of large power reactors increases, space-time effects during reactor transients becone more important. In order not to penalize reactor performance unduly with overly conservative design methods, it is desirable to have the capability of performing detailed space-time neutronics calcula-tions for both design and off-design transients. The HERMITE (18) computer code has been developed to meet this objective. It solves the few group, space and time dependent neutron diffusion equation including feedback effects of fuel temperature, coolant temperature, coolant density and control rod motion. The neutronics enuations are solved by a finite element method. The fuel temperature model explicitly represents the pellet, gep and clad regions of the fuel pin, ar.d the governing heat conduction equations are solved by a finite difference method. Continuity and energy conservation equations are solved in order to determine the coolant temperature and density. For one dimensional analysis of the Axial behavior of the core the QUIX code is usad. Three-dimensional RCCS depletion calculations are used to supply the necessary input to the QUIX code, which generates some of the data required by the monitoring and control systems. In addition to the eigenvalue problem, QUIX will perform three types of search calculations to attain a specific eigenvalue; viz., a poison search, CEA region boundary search, and a moderator density dependent poison search. The effects of moderator- and fuel-temperature feedback on the power distribution can be treated. The program can also perform power shaping searches to simulate the use of PLCEAs. Amendment No. 3 4.3-24 May 28, 1981
[w"* The QUIX code has the capability of simulating ex-core detector responses C} expected during operation. The calculated normalized core average power distribution is first corrected by the application of CEA shadowing factors to simulate the peripheral fuel assembly power distribution. Shape annealing factors (defined below) are then applied to the peripheral axial power distribution to simulate the integrated response of the subchannels of the three-element excore detectors.
- CEA shadowing is the change in excore detector response resultine from changing the core configuration from an unrodded condition to o Adition with CEAs inserted, while maintaining constant power operation. Although CEA shadowing is a function of azimuthal locations its effect is minimized by placing the ex-core detectors at azimuthal locations where minimum CEA shadowing occurs. CEA shadowing factors can be determined using detailed two-dimensional power distributions (XY-PDQ's) representing the cumulative presence of the various CEA banks and the shielding code SHADRAC. (19)
SHADRAC calculates fast neutron and gamma ray spectra, heating, and dose rates in a three-dimensional system untilizing a moments method solution of the transport equation. The core, vessel internals, vessel, and excore detector location are treated explicitly in the calculation. Normalized CEA shadr.<ing factors are relatively constant with burnup and pcwe. level changes made without moving CEAs. CEA shadowing factors at beginninn and end of first cycle life are as shown in Table 4.3-11. p) ( v Figure 4.3-58 shows the typical behavior of the CEA shadowing factor during a CEA insertion and withdrawal sequence. QUIX simulated factors and experi-mentally measured CEA shadowing factors during this transient situation are 3 shown to have quite good agreement over a significant range of CEA insertions. Shadowing factors account for the radial effects and annealing accounts for the axial effects on the excore detector responses. Due to neutron scattering in the various regions separating the core and the excore detectors, each detector subchannel responds to neutrons from the entire length of the core and not just from the section immediately opposite the subchannel. This effect is independent of the axial power shape and the azimuthal CEA shadowing factors. Typical axial annealing functions, given as fractional response per percent of core height for a three subchannel system, are shown in Figure 4.3-59. Axial shape annealing is determined utilizing a fixed source DOT calculation. DOT (20) is a two-dimensional discrete ordinates traasport code. The RZ geometry option is used for the anne ng calculation with representation of the core, vessel, vessel internal >, air gap, and biological shield. The actual detector walls are not represented in the 00T calculation. However, the self-shielding effect of the walls on the detector -esponse is small and uniform along the length of 'the detector and as a result will not change the normalized annealing. The fixed source is distributed over 20 uniform axial segments of the active core height. Figure 4.3-60 illustrates the radial regions represented in the DOT calculation as well as one of the m axial fixed source segments. By integrating the total pointwise response along the length of the subchannel, one can determine the subchannel response for each axial fixed source 4.3-25 by"b8"}9h. 3
- , . __ _-
._- .-. _ - - ... - . _
segment. Utilizing this same technique for all 20 axial segments of the core and normalizing to the total ex-core detector response, annealing curves similar to Figure 4.3-59 are detennined. As the annealing is decermined using a flat axial shape, the resulting annealing factors (S(z)) must be multiplied by the appropriate peripheral axial power shape, P(z), to obtain he total detector response. H D Lower = P(z)S(z) lower dz = lower detector response o H
=
D Middle P(z) S(z)middledz = middle detector response o H D Upper = P(z) S(z) upper dz = upper detector response o 3 The shape annealing factors are purely geometric correction factors applied to the peripheral axial power distribution. As such, .he effects of tim" in fuel cycle, transient xenon redistribution and CEA insertion, although affecting the peripheral bundle power shape, do not effect the geometric shape annealing correction factors. Figure 4.3-61 compares the peripheral axial shape index with the external shape index, inferred from detector signals at the core boundary, during a CEA and PLCEA motion test for the Palisades reactor. Shown are the results of QUIX simulations of the test as well as experimental data taken during the test. From this curve, we can conclude that even though the axial power distribution in the core and on the core periphery was changing during this transient the relationship between the ex-core response and the peripheral response was not. These results justify not only the separability of CEA shadowing and shape annealing as summed in QUIX but also demonstrates that shape annealing is purely a geometric effect, independent of the peripheral axial power distribution. The ex-core detector temperature decalibration effect is the relotive change in detector response as a function of reactor water inlet temperature. The temperature decalibration effect is calculated utilizing SHADRAC with explicit representation of core, vessel ir. cenals, vessel and detector location for various reactor inlet temperatures. Typical detector temperature decalibration effect as a function of inlet temperature normalized to an inlet temperature of 565F is as shown in Figure 4.3-62. Amendment No. 3 4.3-26 May 28, 1981
. - . . -- - -- - _ - - - _ - -. ~ _. . -_ - - . ,
Final normalization of the CEA shadowing, shape annealing, and temperature decalibration constants will be accomplished during startup testing. 3 4.3.3.1.2 Comparisons with Experiments The nuclear analytical design methods in use for SYSTEM 80 have been checked
! against a variety of critical experiments and operating power reactors. In i the first type of. analysis, reactivity and reaction rates and power distri-bution calculations are performed, which lead to information concerning the , validity of the basic fuel cell calculation. The second type of analysis '
consists of a core follow program in which power distributions, reactivity l' coefficients, reactivity depletion rate, and CEA worths are analyzed to provide a global verification of the nuclear design package, i 4.3.3.1.2.1 Critical Experiments !
,
Selected critical experiments have been analyzed with the DIT code. Selec-
! tions of criticals is based on the following criteria:
- 1. Applicability to C-E PWF fuel and assembly designs,
'
- 2. Self-consistency of measured parameters, and
! 3. Availability of adequate data to model the experiments. ! Two groups of critical experiments using rod arrays representative of the 14x14assemblyhavebeenemployedinthisevaluation.Thefg*isa series of clean experiments with UO fuel carried out in 1967 and the 7 i ' second is a set of experiments carrTed out in 1969 (22). Tables 4.3-12 and 4.3-13 give the principal parameters for each of the experimental configurations.
- The moderator-to-fuel volume ratios were varied by changing the cell pitch l of the fuel rod arrangement. The moderator and reflector material for all
) core was H2 0.
; Measurements included the criticality parameters and the fission rate
- distributions in selected fuel rods. This section addresses the comparisons l
between measured and calculated criticality, as well as between measured and calculated fissions rate distributions done to establish calculative ! biases and uncertainties iri predicting intra-assembly power peaking for . both 14 x 14 and 16 x 16 arrays. 3 ! Description of the Experiments
- a. Combustion Engineering Sponsored U0 2 Critical Experiments.
A series of critical experiments were performed for Combustion , Engineering by Westinghouse Corporation at the Westinghouse ! Reactor Evaluation Center (WREC) employing the CRX reactor. The ' experimental program consisted of approximately 70 critical
- configurations of fuel rods. The basic core configuration es a j 30X30 square, fuel rod array of Zr-4 clad 007 fuel having an i
] enrichment of 2.72 w/o U-235. Fuel rods were removed to create , internal water holes or channels to accommodate control rods or to simulate control rod channels and water gaps representative of j\ - the C-E 14X14 fuel assembly design. Amendment No. 3
- 4.3-27 May 28, 1981 i
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The majority of the experiments employed a lattice pitch of 0.600 inches with several experiments repeated with a lattice pitch of ' O.575 inches. These values of 0.600 and 0.575 inches, together with the fuel pellet dimensions and enrichment and the rod diameter, resulted in hydrogen to fuel ratios representative of the 14X14 design at room and at operating temperatures, respectively,
- b. KRITZ Experiments A program of critical experiments, sponsored jointly by Combustion Engineering and KWU, was performed at the KRITZ CRITICAL FACILITY of AB Atomenergi, Studsvik, Sweden. The program consisted of analyzing a number of core configurations of interest to C-E and KWU. The C-E configurations were representative of the 14X14 fuel assembly, including the 5 large control rod channels. A basic cell pitch of (0.5650) inches was used for all lattices. The cores were relatively large 3 both in cross-sectional area and height. Each core contained about 1450 rods 265 cm in length. The core was reflected with water on the four sides and the bottom. Soluble boron was employed for gross reactivity control.
Results of Analyses The results of the analyses of the six critical experiments are summarized in Table 4.3-14. The average Keff is 1.0016. As part of the C-E Criticals and the KRITZ CRITICALS experimental programs, pin by pin power distributions were also measured, to provide a data base with which to define biases and uncertainties in predicted water hole peaking factors. This analysis is described in detail in Reference 2,: When applied to two-dimensional PDQ calculations, the bias and 95/95 tolerance limit in assembly-wise peaking factor are 0.0 and + 2.4% respectively. 4.3.3.1.2.2 Power Reactors. The accuracy of the calculational system in its entirely can onl; be assessed through the analysis of experimental data collected on operating power reactors. The data under investigation consists of critical conditians, reactivity coefficients, and rod worths measured during the startup period, and of critical conditions, power distributions, and reactivity coefficients measured throughout the various cycles.
- 4. 3. 3.1. 2. 2.1 Startup Data. Because of the clean core configuration prevailing during the initial zero power operation, the startup data is extremely valuable in assessing the validity of the physics design package.
The soluble boron concentration that has to be added to the moderator to bring the unrodded reactor critical is a measure of the excess reactivity present in the core to accommodate thes negative reactivity insertion due to the power escalation and the fuel depletion. The boron concentrations measured at the hot condition for various reactors are compared with the predicted values in Table 4.3.15, showing an average error of 14+25 ppm at 3 the 95/95 conf idence level. In terms of reactivity, this corresponds to an underprediction by 0.18% ap 0.31% Ap. Amendment No. 3 4.3-28 May 28, 1981
The moderator temperature coefficient is the change in reactivity resulting _ from a unit change in moderator temperature. As the temperature increases N and the moderator density decreases, two phenomena take place, affecting [j ( the reactivity in two opposite directions. The reduced water density affects adversely the slowing down of neutrons to thennal energies, reducing the fission rates and thus the reactivity. Another effect of the reduced water density is a displacement of the soluble boron, which results in a reduction of the thermal absorption rate and an increase in reactivity. In a fresh core, when the soluble boron concentration is at its maximum, the second effect may overcome the first one and the moderator temperature coefficient may be positive at low power levels. As the core depletes and the boron concentration is reduced to maintain criticality, the first effect becomes predominant, and the moderator temperature coefficient becomes increasingly negative. The result of analyses performed on eleven different cycles is given in Table 4.3.16. The bias in calculated moderator temperature coefficients is linear w th soluble boron concentration, with a 95/95 confidence level of 0.18 x 10 j/ F. (Figure 4.3.63). CEA Reactivity Worth. Comparisons were made between the predicted and measured CEA worths for individual banks inserted sequentially for the 3 Calvert-Cliffs Units I and II, St. Lucie I and Arkansas reactors. Table 4.3-17 summarizes this comparison between calculation and experiment. In the evaluation of the experimental data, values of delayed neutron fractions published in Reference 24 were used. The comparisons demonstrate that CEA reactivity worths can be calculated to with 2% for first cycles and 6% for later cycles. Fuel Temperature and Power Coefficients. The power coefficient is expressed in terms of reactivity change per unit change in power. This coefficient consists mostly of two components: one results from the reactivity change associated with the core coolant temperature distribution, and the other, which will be discussed here, is due to the change in fuel temperature. At each power level, an equilibrium exists between the power produced in the fuel, the fuel temperature, the heat transfer between the fuel and the coolant, and the energy removed by the coolant. The fuel temperature can be calculated directly by a heat transfer calculation, or indirectly by the analysis of the reactivity effects associated with a fuel temperature change. Both approaches have been used aqd lead to very consistent results. The reactivity effects attributable to the fuel temperature are due to the Doppler broadening of the cross-section resonances, mostly those of uranium U238, as well as to the change in scattering properties of the oxygen present in the fuel. The power coefficient can be expressed as aP _ ao aT ap aT aP The determination of the first term, a p/aT, is performed by the lattice code DIT. The second term, aT/aP, is calculated by the FATES (14) code, n for a SYSTEM 80 fuel pin as well as for a 14X14 assembly fuel pin which has 3 different dimensions but the same fuel densification properties. FATES (w Amendment No. 3 4.3-29 May 28, 1981
shows that the fuel temperature is insenstive to the pellet diameter, and therefore that the measured power coefficient ap/3P for 14X14 cores can be used to establish the validity of the fuel temperature correlation. The results of the analyses are given in Table 4.3.18, 4.3.3.1.2.2.2 Depletion Data The two quantities which are monitored on a continuing basis during nominal full power operation are the reactivity depletion rate and the power distributions. The constant monitoring of these quantities establishes the validity of the nuclear design. The reactivity depletion rate is monitored by comparing measured critical steady states conditions with corresponding calculated conditions. These conditions are characterized by exposure, power level, boron concentration, inlet temperature and control rod insertion. Since the measured and the calculated critical conditions. cost likely differ in some respects, an interpolation scheme was devised to infer a calculated reactivity at each measured conditions. The results are displayed in Figure 4.3.64 for five later cycles of the 14X14 fuel assembly type core. It shows a small and consistent, burn-up independent reactivity bias of -0.25% ap, with a 95/95 probability level of + 0.22% ap. This bias is in good agreement with the hot zero power bias given earlier, demonstrating that Doppler and thermal hydraulics reactivity effects, as well as fission product worth are correctly treated throughout life by current design methodologies. 3 For corcs of the 16X16 assembly type such as System 80, the experimental data base is not as large. Figure 4.3.65 show a composite picture of the reactivity bias for both reactor types. lnis demonstrates that the reactivity predictive capabilities for System 80 are comparable to those for current reload cycles of the earlier design. Assembly Power Distributions The uncertainty to be attributed to calculated fuel assembly power distributions can be obtained by comparing detailed 3-0 calculations of the assembly powers with those inferred from incore measurements with the CECOR (25) system using fixed in-core rhodium detectors. The resulting differences are a reflection of both measurement and calculative errors. In order to determine the uncertainty to be attributed to the calculation, the measurement uncertainty has been subtracted out from these difference distributions as describmi below. The measurement uncertainty was taken from an evaluation of the uncertainty associated with the CECOR system (23). c , ,c c Estimates have been made of o and o for the standard deviations of the differences between calb01ateb$n,d true $s,embly power. The data base included Arkansas Nuclear One Unit 2 (AN02) cycle 1, Calvert Cliffs Unit I (BGE I) cycles 1-2, Calvert Cliffs Unit 2 (BGE II) cycles 1-2, and St. Lucie Unit 1 (FPL) cycles 1-3. AN02 is a 177 assembly core with a 16X16 fuel pin lattice, while the other cores have 217 assemblies with a Amendment No. 3 4.3-30 May 28, 1981
_ - _ 14X14 lattice. Overall, comparisons were made for these 8 cycles over 112 time points with about 40 instrument strings each, resulting in about 20,000 data points.
'
C) i' Table 4.3.19 summarizes the calculational uncertainties. 3 4.3.3.2 Spatial Stability 4.3.3.2.1 Methods of Analysis An analysis of xenon-induced spatial oscillations may be done by two classes of methods: time-dependent spatial calculations and linear modal analysis. The first method is based on computer simulation of the space, energy, and the time dependence of neutron flux and power density distributions. The second method calculates the damping factor based on steady-state calculations of flux, importance (adjoint flux), xenon and iodine concentrations, and other relevant variables. The time-dependent calculations are indispensable for studies of the effects of CEA, core margin, out-of-core and in-core detector responses, et., and are performed in one, two, and three dimensions with few-group diffusion theory, using tested computer codes and realistic modeling of the reactor Core. The linear modal analysis methods are used to calculate the effect on the damping factors of changes in fuel zoning, enrichment, CEA patterns, operating temperature, and power levels. These methods, using information at a m single point in time, are particularly suited to survey-type calculations. Methods are based on the work of Randall and St. John (26) as extended by [V } Stacy. (27) These methods are verified by comparison with time-dependent calculations. 3 i 4.3.3.2.2 Radial Xenon Oscillations .' To confirm that the radial oscill e. ion mode is extremely stable, a space-time calculation was run for a reflected, zoned core representative of System 80 l3 without including the dampir.g effects of the negative power coefficient. The initial perturbation was a poison worth of 0.4% in reactivity placed in the central 20% of the core for 1 hour. Following removal of the perturbation the resulting oscillation was followed in 4-hour time steps for a period of 80 hours. The resulting oscillation died out very rapidly with a damping factor of about -0.06 per hour. When this damping factor is corrected for a finite-time step size by the formula in Reference 28, a more negative r damping factor is obtained indicating an even more strongly convergent l3 oscillation. On this basis, it is concluded that a radial oscillation instability will not occur. 4.3.3.2.3 Azimuthal Xenon Oscillations Two-dimensional modal analysis techniques were used to calculate the damping factor for azimuthal oscillations, and included both the fuel temperature and moderator temperature components of the total power coefficient. These calculational techniques were used to predict the results of azimuthal v Amendment No. 3 4.3-31 May 28, 1981
. _
_ __ _
._ _ _ _ . . _ _. . _ _
oscillation tests at Maine Yankee at 75% power. The predicted damping factor of -0.045 per hour for azimuthal oscillations was found to agree well with the measured value of -0.047 +0.005 per hour. I3 4.3.3.2.4 Axial Xenon Oscillations To check and confirm the predictions of the linear modal analysis approach, numerical space-time calculations were performed for both beginning and end-of-cycle. The fuel and poison burnup distributions were obtained by depletion with soluble boron control, so that the power distribution was strongly flattened. Spatial Doppler feedback was included in these calcula-tions. In Figure 4.3-66, the time variation of the power distribution along the core axis is shown near end-of-cycle with reduced Doppler feedback. The initial perturbation used to excite the oscillations was a 50% insertion into the top of the core of a 1.5% reactivity CEA bank for 1 hour. The damping factor for this case was calculated to be about 0.02 per hour; however, when corrected for finite-time step intervals by the methods of Reference 28, the damping factor is increased to approximately +0.04. When l3 this damping factor is plotted on Figure 4.3-67 at the appropriate eigenvalue separation for this mode at end-of-cycle, it is apparent that good agreement is obtained with the modified Randall-St. John distribution of the moderator coefficient about the core midplane, and its consequent flux and adjoint weighted integrals of approximately zero. Axial xenon oscillation experiments performed at Omaha at a core exposure of 7000 mwd /T and at Stade at beginning of cycle and at 12000 MWD /T (29) were analyzed with a space-time one-dimensional axial model. The results are given in Table 4.3-20 and show no systematic error between the experi-mental and analytical results. 4.3.3.3 Reactor Vessel Fluence Calculation Model The calculated vessel fluence is obtained by combining the results of ANISN (30) and SHADRAC (19) in the following manner: 4 A (SHADRAC) t(E) = 4'(ANISN)
& B (SHADRAC) where 4 (E) is the neutron energy flux at the inner surface of the vessel, 4 (ANISN) is the neutron energy flux obtained from ANISN, 4 A (SHADRAC) is the neutron energy flux as calculated by SHADRAC in which the exact source geometry and a three-dimensional time aver aed power distribution are used.
4 8 (SHADRAC) is the neutron energy flux as calculated by SHADRAC O Amendment No. 3 4.3-32 May 28, 1981
using a cylindrical source geometry and the power distribution
/~N obtained from ANISN. ' (V) The neutron flux as calculated by the above method has uncertain.y limits of +10%, 40%. The total uncertainty is composed of 0%, -30% in the calculational method and +10% uncertainty in the combined radial and axial power distribution. The calculational uncertainty factors are obtained by
, comparing the ANISN-SHADRAC results with measurements from various oparational reactors.(31) l3 4.3.3.4 Local Axial Power Peaking Augmentation A reduction in U02 volume associated with fuel densification results in a shortening of the active fuel pellet stack height. If it is assumed that the reduction in the active stack length is not reflected by an equivalent increase in the length of the gas plenum but, instead, results in the formation of axial gaps within the fuel column, local power peaking is experienced in the vicinity of the fuel gaps. This arises because the decrease in neutron absorption, due to fuel removal, more than compensates for the fission loss. Since the magnitude of the local power peaking in a given rod is a function of both the size and number of gaps in surrounding fuel rods, and the distribution of gaps within a given volume of the core cannot be defined explicity, a statistical approach to the determination of the local peaking factor, resulting from the presence of gaps, is employed. This additional peaking due co gaps is called the augmentation factor. The augmentation [m factor, at any given plane, is defined as the ratio of th maximum power in V) that plane with the statistically expected distribution of gaps to the maximum power without gaps. The peaking augmentation factors are based on a 95% confidence level. That is to say, for each axial region, the augmented power is chosen so that there is a 95% probability that no more than one rod exceeds the augmented power. l3 Input information, which is specific to the reactor under consideration and is required for the calculation of augmentation factors, includes the fuel densification characteristics, the radial pin power distribution, end the single gap peaking factors. fh! fuel densification characteristics used in the calculation of augmentatirn factors for System 80 are presented in Table 4.3-21. The radial pin power census used in calculation of the augmentation factors is given in Table 4.3-22 and Figure 4.3-68. Figure 3 4.3-69 shows the specific assignment of fuel rod locations to radial groups and gives the power peaking associated with a single gap at each of these locations used in the calculation of the limiting augmentation factors for System 80. The axially dependent peaking augmentation factors for noncollapsed gaps 3 are presented in Figure 4.3-70 for System 80. Augmentation factors were calculated using limiting radial pin power distributions and single gap peaking factors, and therefore can be used throughout the first cycle. , 4.3.4 Section Deleted 3 V Amendment No. 3 4.3-33 May 28, 1981 1
--r,e> .v--.,a g -e -,e-- ,,,-,-,,-n, , , , - , , - , , - . , - . - .
, Referen _cas
- 1. A. Jonsson, R. R. Rcc U.N. Singh, " Verification of a Fuel Assembly Spectrum Code Based on Integral Transport Theory" Trans. Am. Nucl.
Soc. 28, 778 (1978)
- 2. I. Car 1vik, " Integral Transport Theory in One-Dimensional Geometrics,"
AE-227, AB Atomenergi, Studsvik (1966)
- 3. D. E. Kusner, et. al., "ETOG-1, A Fortran IV Program to Process Data from the ENDF/B File to the MUFT, GAM and ANISN Formats,"
WCAP-3845-1, (ENDF-110 , (1969). 4 J. Adin, K. D. Lathrop. " Theory of Methods Used in the GGC-3 Multigroup Cross Section Code." GA-715G. July 19, 1967.
- 5. H. C. Honeck, D. R. Finch, " FLANGE-II, A Code to Process Thermal Neutron Data from an ENDF/B Tape," DP-1278, ENDF-152, (1971).
- 6. P. Kier, A. Robba, " RABBLE, A Program for Computation of Resonance Absorption ir Multiregion Reactor Cells," ANO-7326, (1967).
- 7. Stuart, G. W. and Woodruff, R. W., " Method of Successive Generations,"
Nuclear Science and Engineering, Vol 3, P 339, 1958.
- 8. Kear, G. N. and Ruderman, M. J., "An Analysis of Methods in Control Rod Theory and Comparison with Experiment," GEAP-3937, May 1962.
" Thin Regions in Diffusion Theory Calculations,n 3
- 9. Wachpress, E. L.,
Nuclear Science and Engineering, Vol 3, p 186, 1958.
- 10. Henry, A.F., "A Theoretical Method for Determining the Worth of Control Rods," WAPD-218, August 1959.
- 11. Cadwell, W. R., "PDQ-7 Reference Manual," WAPD-TM-678, January 1968.
- 12. Breen, R. J., et al., "fiARMONY-System for Nuclear Reactor Depletion Computation," WAPD-TM-478, January 1965.
- 13. P. H. Gavin and P. C. Rohr, " Development and 'lerification of a Fuel Temperature Correlation for Power Feedback and Reactivity Coefficient Application," Trans. Am. Nucl. Soc., 30, 765 (1978).
- 14. "C-E Fuel Evaluation Model Topical Report" CENPD-139, Rev. 01 (1974)
- 15. Terney, W. C., " Albedo Adjusted Reflector Fast Diffusion Coefficient,"
Transactions American Nuclear Society, 18, 312, 1974.
- 16. T. G. Ober, J. C. f>tork, I. C. Rickard and J. K. Gasper, " Theory, capabilities and U..? of the Three-Dimensional Reactor Operation and Control Simulator (ROCSO, "Nucl. Sci. Eng., M , 605 (1977).
Amendment No. 3 4.3-34 May 28, 1981
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- 17. T. G. Ober, J. C. Stork, R. P. Bandera and W. B. Terney, " Extension n of the ROCS Co4rse Mesh Physics Simulator to Two Energy Groups,
( ) "Trans. Am. Nucl. Soc., 2_8,8 763 (1978).
%./ 3
- 18. P. E. Rohan, S. G. Wagner, S. E. Ritterbusch: "HERMITE, A Multi-Dimensional Space-Time Kiaetics Code for PWR Transients." Combustion Engineering Topical Reprrt CENPD-188-A, March 1976.
- 19. SHADRAC, " Shield Heating and Dose Rate Attenuation Lalculation,"
G30-1365, March 25, 1966.
- 20. Soltesz, R. G. and Disney, R. K. , " Users' Mar tal for the DOT-IIW Discrete Ordinates Transport Computer Code,- WANL-TME-1982, December 1969.
- 21. Taylor, E. G., et. al., " Combustion Engineering, Inc., Critical Experiments," WCAP-7102, October 16, 1967.
- 22. " Plutonium Experiments for KWU," A. B. Atomenergie, Studsvik, Sweden, October 1969.
- 23. A. Jonsson, W. B. Terney, M. W. Crump. " Evaluation of Uncertainty in the Nuclear Power Peaking Measured by the Self-Powered, Fixed In-Core Detector System." CENPD-153, Rev. 1. May 1980.
- 24. Tomlinson, L., " Delayed Neutrons for Fission, "UKAEA Report ALRJ-R-
-6993
, Harwell, Berkshire,1972.
g
- 25. "A Method of Analyzing In-Core Detector ta in Power Reactor,"
CENPD-145, C-E Topical Report, April 19,
- 26. Randall, D. And St. John, D.S., " Xenon Spatial Oscillations."
Nucleonics, 16, 3, pp 82-86, 1958.
- 27. Stacey, Jr. , W. M. , " Linear Analysis of Xenon Spatial Oscillations,"
Nuclear Science Engineering 30, pp 453-455, 1967. 3
- 28. Poncelot, C. G., "The Effect of a .tinite Time Step Length on Calculated Spatial Xenon Stability Characteristics in large PWR's, "Trans.
ANS.10, 2, p 571. l l 29. Gruen, A., "Messung Physikalischer Kenngroessen an Leistungsrcakoren," Atom Kernenergie, 25, 2, 1975.
- 30. Engle, Wald M. , Jr. , " A User's Manual for ANISN, A One-Dimensional Discrete Ordinates Transport Code with Anisotropic Scattering," K-i 1693, March 30, 1967.
l 31. Stephen, D. W., " Fast Neutron Attenuation by the ANISN-SHADRAC l Analytical Method," CENPD-105, June 1973.
- 32. "COLSS, Assessment of the Accuracy of PWR Operating Limits as i Determined by the Core Operating Limit Supervisory System," CENPD-169 C-E Proprietary Topical Report, July 1975.
'
- 33. L ebs, W. D. and Brinkman, H., Proceedings of Reaktortagung 1976,
'Jusseldorf, Germany, March 1976.
Amendment No. 3 4.3-35 May 28, 1981 l l
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TABLE 4.3-1 NUCLEAR DESIGN CHARACTERISTICS (Sheet 1 of 2) Item _ Value General Characteristics Fuel management 3-batch, mixed central znne Core Average Burnup (MWD /T),10 ppm soluble boron 16576 3 Core Average U-235 Enrichment (wt%) 2.64 Core Average 2H 0/U0 2 y lume ratio, first cycle, hot (core cell) 2.11 Number of control element assemblies Full length 76 Part length 13 p Burnable Poison Rods U Number 2112 l3 Material B4 C-Al230 Worth %Ap, at BOC Hot, 594 F 7.8 Cold, 68 F 6.3 Dissolved Boron Dissolved baron content for critic.ality, ppm (CEAs withdrawn, B0C) Cold, 68 F 1084 Hot, zero power, clean 565 F 1018 3 Hot, full power, clean, 594 F 911 liot, full power, equilibrium Xe 657 l ,
'
f"N
'\~)
Amendment No. 3 May 28, 1981 _ __ . - _ . _ _ - - _ - - __
- TABLE 4.3-1 NUCLEAR DESIGN CHARACTERISTICS (Sheet 2 of 2)
Item Value Dissolved boron content (pon: for: Refueling 2150 5% subcriticel, cold, first cycle 1419 (all CEAs ..:: : 0 MWD /T 5% subcritical, hot, first cycle 1453 (all CEAs out) 0 MWD /T Joron worth, ppm /%Ap (80C/E0C)
; lot, 594 F 91/83 Cold, 68 F 67/54 Neutron Parameters 3 Neutron lifetime (cycle average) microseconds 28.2 Delayed neutron fraction (cycle average) 0.0064 Plutonium Buildup (first cycle)
Gms Fissile Pu (final) kg U (original) 4.90 GMS Total Pu (final) kg U (original) 6.27 l l l . ' O knendment No. 3 May 28, 1981
TABLE-4.3d
,
EFFECTIVE MULTIPLICATION FACTORS AND REACTIVITY DATA (a)
f J J' d Condition K P eff - l Cold,68"F(0 PPM)BOCl 1.193 0.162 ! Cold (68'F) at minimum refueling boron 0. 91 3 -0.095 I concentration-(1,720 ppm)B0Cl-
.,
Hot, 564 F, zero. power, clean, (0 ppm) B0Cl 1.133 0.117
;
Hot, ful' power, no Xe or Sm, 594 F 1.112 0.101
! (0 ppm) 60C1
i Hot, full power, equilibrium Xe (0 ppm) 1.078 0.073 Hot, full power equiiibrium Xe and
;
Sm (0 ppm) 1.073 0.068 3 Reactivity decrease, hot
- Zero to full power, B0C (911 ppm) 0.012 i
Fuel temperature ' 0.011 Moderator temperature $ 0.001 Reactivity decrease, tho
!
- Zero to full power, E0C (0 ppm) 0.014 j Fuel temperature 0.0008
! Moderator temperature 0.0006 i
! , a. No control element assemblies or dissolved boron except as noted, j initial core. ! ( l i. l i , l[ l
1-l ' Amendment No. 3 May 28, 1981
TABLE 4.3-3 COMPARIS0N 0F CORE REACTIVITY COEFFICIENTS WITH THOSE USED IN VARIOUS SAFETY ANALYSES Moderator iemperature Density Coefficieng Doppler (a) Coefficiegt (Ap/ F X 10 ) Coefficient (Ap/gm/cm ) Coefficients from Table 4.3-4 Full power B0C -0.52 Figure 4.3-45 EOC 1 -2.51 Figure 4.3-45 0.066(b) N/A 3 Zero power, CEAs banks 5.4 dnd 3 inserted BOC -0.92 Figure 4.3-45 N/A E0C 1 -2.11 Figure 4.3-45 N/A Coefficients used in Accident Analyses CEA withdrawal Full /zero power 0/0(d) 0.85/0.85 N/A CEA misoperation (full length) Dropped CEA -3.5(d) 1.15 N/A CEA misoperation (part length) Dropped CEA 0(d) 0.85 N/A Loss of Flow 0(d) N/A N/A CEA ejection BOC, full /zero power 0/0(di' O.85 N/A Loss-of-coolant accident (c) Figure 4.3-45
"
(c'
" 3 Small Break 0
" "
Large Break +.5
- a. Nominal values of the Doppler coefficient (Ap/ F) es a function of the fuel temperature are shown in Figure 4.3-45. The numbers entered in the Doppler column of this table are the multipliers applied to the nominal value for analysis of designated accidents.
- b. Not applicable
- c. A curve of reactivity vs. moderator density is used for the LOCA evaluation.
The value of density coe{ficient used corresponds to a 0 MTC for the small break events and +.5x10~ Ap/ F for the large breaks resulting in rapid 3 depressurization.
- d. These values are the ones used at the nominal Tave = 594 F. For other temperatures the set of curves shown on Figures 4.3-46 and 4.3-47 corresponding to the extreme (i.e. , most positive at B0C, most negative 3 at E0C) will be used. Amendment No. 3 May 28, 1981
TABLE 4.3-4 (
/j\ REACTIVITY COEFFICIENTS Moderator Temperature Coefficient, Ap/ F Beginning-of-cycle (0-50 MWD /T)
Cold, 68*F, Clean,1084 PPM -0.05 x 10-4 Hot, zero power, 564 F, no CEAs Clean,1018 PPM l Hot full power, 594 F no CEAs, Clean, 911 PPM -0.52 x 10"
-0.17x10-l Hot full power, 594*F, no CEAs, Equilibrium Xe, 657 PPM -0.98 x 10-Hot zero power 564F regulating CEAs banks 5, 4 and 3 inserted, 50 MWD /T, 657 ppm, Hot full power equilibrium Xe -0.92 x 10-4 End-of-Cycle (10 ppm soluble boron,16576 MWD /T)
Cold, 68 F approximate +0.25 x 10-4 Hot zero power, 564 F, no CEAs, Hot full power i equilibrium Xe -1.88 x 10-4 Hot full power, equilibrium Xe, no CEAs,. 594*F -2.51 > 'O'4 Hot zero power 564 F rodded, regulating CEAs banks 5, 4 and 3 inserted, Hot full power equilibrium Xe -2.l' ,, 10-4 Moderator Density Coefficient, Ap/gm/cm y '
- Hot, operating, 594 F 3 Beginning-of-cycle, 911 ppm soluble boron, OMWD/T +0.066 Fuel temperature contribution to power co2fficient, ap/(kW/ft), 657 ppm, 50 MWD /T Hot zero power. -2.40 x 10-3 Full power, -1.68 x 10-3 Moderator void coefficient ap/% void Hot operation, E94 F
Beginning-of-cycle, 911 ppm soluble boron, OMWD/T -0.46 x 10-3
- Moderator pressure coefficient, ap/ psi l Hot operating, 594 F Beginning-of-cycle, 911 ppm soluble boron, OMWD/T +1.10 x 10 -6
? , Overall power coefficient, ap(KW/ft) Hot operating, 594 F Beginning-of-cycle, 657 ppm soluble boron, 50 MWD /T -2.21 x 10-3 j End-of-cycle, 10 ppm soluble boron, 16576 MWD /T -2.65 x 10-c t V
, Amendment No. 3 ' May 28, 1981 l ! _
TABLE 4.3-5 WORTHS OF CEA GPCUPS (%Ap) 2K MWD /T 14K MWD /T (635 ppm) (214 ppm) Shutdown CEAs -12.06 -13.18 Regulating CEAs Group 1 -1.21 -1.41 Group 2 -1.04 .94 3 Group 3 .79 .99 Group 4 .47 .47 Group 5 (LEAD Bank) .32 .32
-15.89 -17.26 O
O Amendment No. 3 May 28, 1981 _ _.
r_______..___.----- -_ __ -- --
. _ . - - - _ _ . _ _ _ _ - - - . _ _ _ _ _ _ _
l ! i i !- t ! ! TABLE 4.3-6 CEA REACTIVITY ALLOWANCES (%Ap) ;
!
- t Fuel temperature variation 1.04 !
l
!
Moderator temperature variation 3.08 Moderator voids 0.1 CEA bite 0.2 Part-length CEA effects 0.0 , Cooldown to Minimum Temperature '3.79 i 3 l , Total reactivity allowance - 8.21 ; i e , t 3 i
!
l l '
@ ;
!
!
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. e
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!
-
I ! l ! i l l 1 , I '
!
!
l l ! ! , l !
.
l Amendment No. 3 l May 28, 1981 l l
TABLE 4.3-7 COMPARISON OF AVAILABLE CEA WORTHS AND ALLOWANCES Reactivity Condition (%Ap) All full-length CEAs inserted, hot, 594 F 15.0 Total reactivity allowance full power 8.2 3 (from Table 4.3-6) Stuck rod worth 4.9 Excess, assuming inost adverse stack-up of uncertainties 1.9 3 O O Amendment No. 3 , May 28, 1981
- . . -
~ .___._ =___ _.._ _ _ _ _ _ _ _ _- _.._ __._... _ _. _ _
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_ _ _ _ _ . _ . _ _ . . . - . --
1 i i .
!
j TABLE 4.3-8 i
- !
- :
COMPARIS0N OF RODDED AND UNRODDED PEAKING FACTORS FOR l VARIOUS RODDED CONFIGURATIONS 3 [ _ i ! ;
'
!
- !
MaximumRodRadigl { Peaking Factor F 7
- .
j Configurations (2K MWD /T) (14K MWD /T) i j Unrodded 1.37 1.39 } Bank 5 1.54 1.40 il l PLCEA 1.44 1.49 i i !
'
j Bank 5 and PLCEA 1.55 1.41 Bank 5+4 (a) 1.53 1.59 3
>
i
- Bank 5+4+3-(a) 1.94 1.82 i
'
Bank 5+4+3+2 (a) 1.77 1.76 i
#
i Bank 5+4+3+2t1 (a) . 2.03 1.93 ! t - 1
!
l l a. No PLCEA I r ! i i i i : l \ !
!
l i ' l f i ! l
i l l
.
Amendment No. 3 . t May 28, 1981 1 ___ _-_
TABLE 4.3-9 CALCULATED VARIATION OF THE AXIAL STABILITY INDEX DURING THE FIRST CYCLE (a)(h"I) Powei Level (% of Full Power) BOC E0C 100 .006 +.115 3
- a. Equilibrium xenon conditions O
w~ O Amendment No. 3 May 28, 1981 . . . _ _ . . _
. - _ - _ . . _ . . - - . _ - - __
_ __ l l
!
!
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!@ l TABLE 4.3-10 i 2 MAXIMUM FAST FLUX GREATER THAN 1 MeV (n/cm -s) t i ! ! 4 i 1 < 4
i i i Lower Energy Flux, Shroud, ID i Neutron Group (MeV) and Core. Periphery Flux, Vessel, ID ' 1 7.41 4.29 (+11)(a) 1.27 (+9)
'
l 2 4.97 2.24 (+12) 3.76 (+9) 1 ' '
-
3 3.3a 5.04 (+12) 3.89(+9) 4 2.23 1.06 (+13) 7.30 (+9) 5 1.50 1.13 (+13) 7.49 (+9) 6 1.22 6.54 (+12) 4.01 (+9) i l l 7 1.00 5.60(+12) 3.44 (+9) l Total- -- 4.17 (+13) 3.12 (+10) i ! l ;
- a. ( ) denotes power of ten ;
!
l l i i !
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! t i I f i
Amendment No. 3 !
!
! May 28, 1981 . - - -. .. ... -- -- - -- . -
__. _ . _ _ _ _ . _ . . _ _ _ __ __ .___ __ _ . _ , . . _ _ . _ _ . . _ _ _ _ _ _ _ _ . . _ _ _ _ _ _ _
;
.; I TABLE 4.3-11 CONTROL ELEMENT ASSEi4BLY SHADOWING FACTORS l ? BOC E0C Bank 5 1.129 1.110 PLCEAs 0.992 0.993 Bank 5 + PLCEAs 1.133 1.117 ! I l
i O
.
l l
,
! l 1 Amendment No. 3 O May 28, 1981
.___ _ -__ . _ - ,
l O TABLE 4.3-12 l
,
i C-E CRITICAi.S Core Configuratiun soluble Fuel Cell Boron No. of Fuel Rod Pitch No. of Fuel Temp. of Conc. Control Rod Lattice Array (Inch) Rods Core PPM Channels
#12 30x30 0.600 880 68 F 0 5
#32 30x30 0.600 832 68 F 0 17 i
#43 30x30 0.600 880 68 F 323 5 68 F 3
#53 30x30 0.575 832 0 17 ,
#56 30x30 0.575 832 68 F 302 17 i
'
CN l '
.
l i Fuel Rod Design Clad OD 0.4683 in ' Clad Thickness 0.03145 in l Clad Material Zr-4 Fuel Pellet OD 0.400 in Fuel Density 10.40 gr/cc [' Fuel Enrichment 2.72 w/o l h V Amendment No. 3 ' May 28, 1981
- - _ _ _ _ - __-. . _ - _ - - -- -.-__ _ - . .-
. .
TABLE 4.3-13 FUEL SPECIFICATION (KRITZ EXPERIMENTS) Fuel material (pellets) UO 2 Fuel Density (dishing 3 including) g/m 10.15 U235 in U wt% 3.10 _____________________________________________________________ 3 Fuel length m 2650 Pellet length m 11 Ovide Diameter m 9.08 _____________________________________________________________ Clading material Zirocaloy 4 3 Density g/cm 6.55 Outer diameter m 10.74 Inner diameter m 9.30 l i i t l I O ! Amendment No. 3 May 28, 1981 _ __
__ () TABLE 4.3-14 Comparison of Reactivity Levels for Non-Uniform Core
, Measured Soluble Boron Vol Mod No. of Large Axial Buckling Conc.
Core Vol Fuel Water Holes M-2 PPM Keff
--
C-E Criticals 2.7% U-235 j 68'F , 3
- #12 1.49 5 3.53 0 1.0017
#32 1.49 17 3.70 0 1.0006
#43 1.49 5 1.64 323 1.0032
#53 1.26 17 2.82 0 1.0021 O #56 1.26 17 1.07 302 1.0006
, KRITZ UO2 (445 F) 1.79 21 2.20 959 1.0014 ,
,
Amendment No. 3 May 28, 1981 ! _ _ . _ _ _ _ _ _ _
TABLE 4.3-15 Beginning-of-Cycle, Hot Zero Power, Xenon Free, Unrodded Critical Boron Concentration Plant Cycle Critical Boron Concentration (PPM) Measured Calculated (ROCS /DIT) BG&E I 1 1096 1078 2 1013 984 3 1220 1216 4 1342 1340 3 BG&E II 1 1097 1087 2 1185 1175 3 1191 1181 FPL 1 962 944 2 1024 995 3 1127 1127 AN02 1 1012 999 Average Difference (Meas-calc) = 14 ppm 95/95 Confidence ' evel = 25 ppm O Amendment No. 3 May 28, 1981
. - _ . . .-. _ ... . . . _ - _
TABLE 4.3-16 ITC
SUMMARY
FOR ROCS /DIT 1 ' CORE CYCLE % Power PPM MEASURED CALCULATED DIFFERENCE (x10~4Ap/ F) BG&E II 1 HZP 1090 +.24 +.48 +.24 50 827 .10 +.03 +.13 96 745 - ?7 .17 +.10 2 HZP 1121 +.43 +.52 +.09 50 940 +.01 +.16 +.15 4 100 690 .50 .54 .04 3 HZP 1191 +.25 +.47 +.22 50 923 .11 .01 +.10 FP&L 1 HZP 962 +.10 +.36 +.26 50 696 .25 .10 +.15 1A 51 681 .21 .12 +.09 83 619 .35 .29 +.06 98 585 .42 .35 +.07 95 296 -1.01 .89 +.12 2 HZP 1024 +.27 +.48 F.21 100 670 .23 .28 .05 3
>
) 97 288 .89 -1.13 .24
'
/ 3 HZP 1137 +.32 +.54 +.22 100 757 .25 .24 +.01 BG&E I 1 HZP 1087 +.26 +.49 +.23 20 923 +.05 +.22 +.17 50 820 .11 +.06 +.17 80 764 .18 .00 +.10 100 740 .21 .14 +.07
, 100 365 .85 .81 +.04 t 95 83 -1.38 -1.48 .10 2 HZP 1013 +.07 +.35 +.28 50 765 .24 .13 +.11 100 593 .72 .67 +.05 3 HZP 1220 +.39 +.67 +.28 50 989 .04 +.24 +.28 100 660 .78 .60 +.18 . 4 HZP 1342 +.36 +.63 +.27 50 1066 +.19 +.24 +.05
- AN02 i HZP 1012 +.03 t.34 +.31
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20 825 .20 +.34 +.23 50 720 .33 .15 +.18 O Amendment No. 3 May 28, 1981 1
._ _ - . . - . . - . _ - _ . .-
.________ ________________ __ __________ _____
TABLE 4.3-17 COMPARIS0N OF CONTROL R0D BANK WORTHS 3D ROCS (DIT) vs MEASUREMENT UNITS OF % DIFFERENCE FROM MEASURED WORTH PLANT / CYCLE AN02 BG&EI BG&EII FPL BG&EI BG&EII FPL BG&EI BG&EII FPL BG&EI SECUENTIAL R0D BANK CY l CY l CY l CY l CY 2 CY 2 CY 2 CY 3 CY 3 CY 3 CY 4 7 - - -
-3.00 - -
+8.83 - -
+8.79 -
6 -2.82 - -
-0.96 - -
+3.02 - -
+6.67 -
5 -0.95 -5.98 -8.06 +0.87 +6.69 +3.55 +3.05 +18.36 -3.30 +5.97 +5.09 3 4 -2.15 +1.42 -4.61 +0.54 +12.46 +9.68 +12.14 +14.38 +3.59 +10.86 +12.50 3 -1.39 -4.29 -2.53 -7.53 +14.67 +2.40 +4.38 +6.92 +5.14 +4.34 +2.70 2 -2.07 +9.64 -6.08 +1.84 -1.70 +3.08 +3.38 -
-0.31 +10.50 +6.30 1 -1.88 -5.46 0.0 +1.88 +10.29 +5.67 -4.53 +11.82 +3.82 +0.39 +7.87 C -
-4.30 -0.08 - - - -
-0.60 - - -
B -
-1.60 -3.96 +3.08 - - - - - - -
CYCLE MEAN -1.88 -3.62 -1.51 -0.41 +8.48 +4.87 +4.32 +10.18 +1.79 +6.79 +6.89 + STD. DEV. +0.65 +2.99 +5.54 +3.44 +6.41 +2.95 +5.23 +7.31 +3.50 +3.70 +3.66 OVERALL 2 p o -1.80 + 3.66 ;' :: +6.12 + 5.14 : Amendment No. 3 May 28, 1981 O O O
=. - -_- - -.- _. ._.. -. . _ . - . . - . .
. . .. . . - . --
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t O O TABLE 4.3-18 i l COMPARISON OF POWER C0EFFICIENTS 1 ! 3D ROCS (0IT) vs MEASUREMENT f , UNITS OF 10-4Ap/%P i ,
I CALCULATED MEASURED DIFFERENCE PLANT / CYCLE POWER (%) BG&El 1 50 -1.28 -1.06 -0.22 100 -1.04 -1.03 -0.01 , 2 50 -1.06 -1.18 +0.12 ! 100 -0.94 -1.02 +0.08 ! 3 50 -1.03 -1.06 +0.03 ' 100 -0.86 -1.10 +0.24 3 4 50 -0.89 -1.08 +0.19 l i i ! BG&EII 1 50 -1.07 -1.07 0.00 ! 96 -0.86 -0.94 +0.08 2 50 -0.95 -1.12 +0.17 + 100 -0.7/ -0.94 +0.17
- 3 50 -0.92 -1.01 +0.09 i
! i l FPL 1 50 -1.15 -1.16 +0.01 ) 83 -0.96 -1.06 +0.10 98 -0.85 -0.90 +0.05
- l 2 100 -2.83 -0.72 -0.11 3 100 -0.77 -0.82 -0.05 i
Average Bias = 0.06 x 10 -4 Ap/%P -4 95/95 Confidence Level = +0.28
- 10
_ Ap/%P
/ Anendment No. 3 l May 28, 1981
- l
., _ _ . - _ .. _ _ , _
TABLE 4.3-19
SUMMARY
OF ROCS /DIT CALCULATIVE UNCERTAINTIES ROCS Calculational Uncertainty: F F F q R Absolute Standard Deviation, S .0357 .0432 .0224 C Degrees of Freedom, f 22 116 14 3 c Confidence Multiplier, k 2.20 1.85 2.41 95/95 O Percent Deviation, S 2.90% 3.52% 1.82% C 95/95 Confidence Internal, 6.39% 6.50% 4.39% KSC (%) l l l Amendment No. 3 May 28, 1993
- - . .. - -- .- - . - . - . ~ _ . - - . . . . _ - . _ - . - . . . . - . - _ _ . _ . - - - . . _ - . . . _
s- ! [e i TABLE 4.3 20 f AXIAL XENON OSCILLATIONS ! f i Period Dampjng
- (h) (h- )
3
-
-
i . Exposure
- Reactor (mwd /T) Measured Calculated Measured Calculated l
f Omaha 7075 29 32 -0.027 -0.030 i ! -Stade BOC 36 36 -0.096 -0.090~ f- i !- Stade 12200 27 30 -0.021 -0.019 l l < i ?
,
!@ i ', ! !
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! , I
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t Amendment No. 3 ,
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May 28, 1981 f f
,_ _ _ . _ _ _ . . - _ _ _ _ _ _
TABLE 4.3-21 DENSIFICATION CHARACTERISTICS Core Height 150 inches 3 Fractional Density Change 0.016 Clad Growth Allowance 0.005 inchet/ inch l
.<
. ! O l l O Amendment No. 3 May 28, 1981
. . . . - - - - ._
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i 1 1 l
! !
4
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i TABLE 4.3-22 l
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RADIAL PIN POWER INTERVAL CENSUS
- j. USED IN CALCULATION OF AUGMENTATION FACTORS t
Number of Pins with i Pin Power Interval Power in Interval ! ! 0-1.00 17,844
3
; 1.00-1.05 8,804 i
F 1.05-1.10 9,780 i
.
i 1.10-1.20 13,324 . l 1.20-1.26 5,204 [ ! i
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! i i i ! l , I i ! l ! t
!
i ! l l l l
,
@
Amendment No. 3 May 28, 1981 l
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l O N 1 2 3 4 F C-00 C-00 C-00 C-00 1 N = B0X NO. 5 6 7 8 9 10 ) F = ASSEMBLY TYPE C-00 C-Lo C-Lo B-Hi B -Lo B-Hi l 11 12 13 14 15 16 17 C-00 C-Lo A B-Lo A B-Lo B-Lo 18 19 20 21 22 23 24 25 C-00 C-Lo B-Lo B-Lo A B-Hi A B-Hi ,
!
26 27 28 29 30 31 32 33 C-Lo A B-Lo A B-Hi A B-Hi A
] 34 C-00 35 C-Lo 36 B-Lo 37 A
38 8-Hi 39 A 40 B-Hi 41 A 42 B-Hi 43 44 45 46 47 48 49 50 51 C-00 B-Hi A B-Hi A B-Hi B-Lo B-Hi A 52 53 54 55 56 57 58 59 60 C-00 B.-Lo B-Lo A B-Hi A B-Hi A B-Hi 61 62 63 64 65 66 67 68 69 C-00 B-Hi B-Lo B-Hi A B-Hi A B-Hi A l Amendment No. 3 _ May 28, 1981 FIRST CYCLE FUEL LOADING PATTERN 4 k
-
i ASSEMBLY NUMBER OF FUEL No. OF FUEL No. OF SHIM TYPE ASSEMBLIES ENRICHMENT RODS PER RODS / gm B l /IN. . W/T % '?d5 ASSEMBLY ASSEMBLY kj A 69 1.92 236 0 - BLO M 16 0.01842 2 B, 64 1.92 12 16 0.02532 H' 2.78 208 CO 40 2.78 12 o _ 3.30 224 1 24 2.78 12 16 0.01151 CLO 3.30 208 XX XX XX XX XS SS # 5'lX X X ___[] E SS 8 []___8 ___[] []___
~
M SS W SS
---O CJ--- ---EJ []---
XE & . E EX X t X XX l XX XX l XX BLO,BHI, AND Ct g CO WATERHOLE E LOWER ENRICHED FUEL PIN O HIGHER ENRICHED FUEL PIN O SHIM PIN Amendment No. 3 May 28, 1981 o ~
,
FIRST CYCLE AGSEMBLY FUEL LOADINGS 9# WATERHOLE AND SHIM PLACEMENT
&3-2
_ . . . _ . .. .. . . - _ - _ - _ . _ _ - _ _ -
FORMAT IS BO'( TYPE NO. MAX. VALUE IN BOX BATCH BOXES PWR.FR. AVG. PWR
] BOX, RPD 1.16 38 A 69 0.30 0.99
\ ,/ MAX 4-PIN 1.27 3 B-HI 64 0.29 1.11 MAX 1-PIN 1.35 4 B -LO 44 0.19 1.07 WITH CORE AVG. POWER 1.00 C 40 0.14 0.80 C 00 1 0-00 2 C -00 3 C -00 4 0.66 0.91 0.98 0.99 1.04 1.21 1.27 1.27 1.11 1.29 1.35 1.35 C-00 5 C-LO 6 C-LO 7 B-HI 8 B-LO 9 B-HI 10 0.61 0.82 1.03 1.02 1.04 1.02 0.98 1.12 1.19 1.13 1.13 1.13 1.05 1.19 1.29 1.21 1.22 1.21 C-00 11 C -L0 12 A 13 B-LO 14 A 15 aLO 16 B -L0 17 0.70 1.03 0.93 1.11 0.98 1.08 1.03 1.06 1.23 1.03 1.22 1.07 1.21 1.13 1.14 1.31 1.10 1.31 1.13 1.29 1.21 C-00 18 C -LO 19 B -LO 20 B -LO 21 A 22 B -HI 23 A 24 B -HI 25 0.62 1.02 1.04 1.13 1.02 1.15 1.00 1.10 0.98 1.23 1.13 1.24 1.12 1.22 1.10 1.20 O 1.05
^
1.31 1.23 1.33 1.18 1.32 1 16 1.29 C-00 26 A 27 B -LO 28 A 29 8 -HI 30 A 31 t 32 A 33 0.81 0.93 1.13 1.02 1.15 1.01 1.15 1.01 1.12 1.04 1.24 1.10 1.23 1.09 1.22 1.11 1.18 1.10 1.33 1.16 1.32 1.16 1.31 1.18 C -00 34 C 35 B -LO 36 A 37 B -HI 38 A 39 B -HI 40 A 41 B -HI 42 0.66 1.02 1.10 1.01 1.16 1.00 1.11 1.00 1.15 1.03 1.18 1.22 1.09 1.23 1.09 1.21 1.11 1.22 1.10 1.28 1.30 1.16 1.32 1.15 1.30 1.17 1.31
~
C -00 43 B -HI 44 A 45 B-HI 46 A 47 B -H I 48 B 49 B-HI 50 A 51 0.89 1.02 0.97 1.14 1.01 1.10 1.05 1.11 1.00 1.20 1.12 1.06 1.22 1.11 1.21 1.12 1.20 1.08 1.28 1.20 1.12 1.31 1.18 1.29 1.21 1,29 1.15 C -00 52 B-LO 53 B -LU 54 A 55 B-HI 56 A 57 B-HI 58 A 59 B-HI 60 0.96 1.03 1.07 0.99 1.14 0.99 1.10 1.00 1.15 1.25 1.13 1.20 1.07 1.21 1.08 1.20 1.08 1.22 1.33 1.21 1.28 1.14 1.30 1.14 1.29 1.15 1.31 C-oo 61 B -HI62 B-LO 63 B-HI 64 A 65 B -HI 66 A 67 B-HI 68 A 69 0.97 1.01 1.02 1.10 1.00 1.15 1.01 1.14 1.01 1.24 1.09 1.10 1.19 1.08 1.21 1.10 1.22 1.09 1.33 1.19 1.19 1.28 1.15 1.30 1.17 1.31 1.16 _ Amendment Nc.3 May 28, 1981 l C-E f PLANAR AVERAGE POWER DISTRIBUTION, BOC, UNRODDED Figure ggg j FULL POWER, NO XENON, O MWD /T 4.3-3
l FORMAT IS BOX TYPE NO. MAX, VALUE IN BOX BATCH BOXES PWR. FR. AVG. PWR O Q BOX, RPD MAX 4-PIN 1.17 1.26 38 12 A B -HI 64 69 0.30 0.29 1.02 1.11 MAX 1-PIN 1.35 21 B -L0 44 0.19 1.06 WITH CORE AVG. POWER 1.00 C 40 0.13 0.77 C-oo 1 0-00 2 C -00 3 C -00 4 0.64 0.87 0.92 0.93 1.02 1.17 1.20 1.20 1.09 1.25 1.28 1.28 C-00 5 C -b0 6 C-LU 7 B-HI 8 B-LU g 3 -Mi gg 0.62 0.83 1.03 0.99 0.99 0.97 1.00 1.15 1.20 1.11 1.06 1.06 1.07 1.22 1.30 1.19 1.15 1.14 C -00 11 C -LO 12 A 13 B-LU 14 A 15 B-Lu 16 B -LO 17 0.71 1.05 0.95 1.11 0.98 1.05 1.00 1.08 1.26 1.06 1.23 1.08 1.19 1.10 1.16 1.35 1.13 1.32 1.15 1.28 1.18 C -00 18 C -Lu 19 B-LO 20 B -LO 21 A 22 B-HI 23 A 24 B-H1 25 0.62 1.04 1.04 1.14 1.04 1.15 1.01 1.09 1.00 1.26 1.14 1.26 1.15 1.24 1.12 1.21 O 1.07 1.34 1.24 1.35 1.22 1.33 1.19 1.30
'
C-00 26 A 27 B-LO 28 A 29 B-HI 30 A 31 B -H1 32 A 33 0.82 0.95 1.14 1.04 1.17 1.04 1.16 1.04 1.14 1.07 1.25 1.13 1.25 1.12 1.23 1.15 1.21 1.13 1.35 1.20 1.34 1.19 1.33 1.21 C -00 34 C 35 B-L0 36 A 37 B -HI 38 A 39 B -HI 40 A 41 B -HI 42 0.64 1.02 1.10 1.04 1.17 1.03 1.12 1.03 1.16 , 1.01 1.19 1.22 1.12 1.25 1.12 1.23 1.14 1.24 l 1.09 1.29 1.31 1.19 1.34 1.19 1.32 1.21 1.33 C -oo 43 B-HI 44 A 45 B -HI 46 A 47 B -HI 48 B- 49 B-HI 50 A 51 0.85 0.98 0.98 1.14 1.04 1.11 1.06 1.12 1.03 ! 1.15 1.10 1.07 1.23 1.15 1.22 1.13 1.22 1.12 1.23 1.18 1.14 1.33 1.21 1.32 123 1.32 1.19 C -00 52 g-LO 53 B -LO 54 A 55 B -H1 56 A 57 B-"I 58 A 59 B-HI 60 0.90 0.98 1.05 1.00 1.15 1.02 1.11 1.03 1.17 i 1.18 1.08 1.19 1.09 1.23 1.11 1.22 1.12 1.24 1.26 1.16 1.27 1.16 1.32 1.18 1.31 1.19 1.34 C -00 61 B-H1 62 B-LU 63 B -H1 64 A 65 B-HI 66 A 67 B-HI 68 A 69 0.91 0.95 0.98 1.09 1.03 1.16 1.04 1.16 1.05 1.18 1.03 1.07 1.20 1.11 1.23 1.14 1.24 1.13 1.26 1.12 1.17 1.29 1.18 1.33 1.21 1.34 1.20 ! Amendment No. 3 May 28, 1981 l PLANAR AVERAGE POWER DISTRIBUTION, BOC, UNRODDED, Figure C-E i [ / FULL POWER, EQUILlBRIUM XENON,50 MWDff 4.3-4
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_
. _ . . .
[^ FORMAT IS BOX TYPE, NO. MAX VALUE IN BOX BATCH BOXES PWR.FR. AVG. PWR BOX, RPD 1.19 60 A 69 0.31 1.03 MAX 4-PIN 1.27 60 B-HI 64 0.29 1.12 MAX 1. PIN 1.37 60 B-t.u 44 0.19 1.06 WITH CORE AVG. POWER 1.00 C 40 0.13 0.74 C -00 1 C -00 2 C -00 3 C-00 4 0.61 0.82 0.87 0.87 0.98 1.11 1.14 1.13 1.05 1.18 1.21 1.21 C -00 5 C-LO 6 C. -LO 7 B -HI 8 B-LO 9 B-HI 10 0.61 0.82 1.01 0.97 0.96 0.94 0.99 1.14 1.19 1.10 1.05 1.03 1.05 1.20 1.28 1.17 1.14 1.11 C -00 11 C-LO 12 A 13 B-LO 14 A 15 B-LO 16 B-LU 17 0.70 1.04 0.95 1.11 0.98 1.05 1.00 1.06 1.26 1.06 1.23 1.08 1.20 1.10 1.14 1.35 1.13 1.32 1.15 1.28 1.18 C -00 18 C -LO 19 B -LU 20 B -LU 21 A 22 B-H1 23 A 24 B-HI 25 0.61 1.04 1.05 1.15 1.06 1.16 1.02 1.10 0.99 1.26 1.16 1.27 1.16 1.25 1.13 1.22 [) 1.05 1.34 1.25 1.36 1.23 1.34 1.20 1.31 , (\ C -00 26 A 27 B -L0 28 A 29 8-HI 30 A 31 B-H1 32 A 33 0.81 0.95 1.15 1.06 1.18 1.05 1.17 1.05 1.13 1.07 1.27 1.14 1.26 1.14 1.25 1.16 1.20 1.14 1.36 1.21 1.35 1.21 1.35 1.23 0-00 34 C 35 B -LO 36 A 37 B-HI 38 A 39 . B H1 40 A 41 B-HI 42 0.61 1.00 1.10 1.05 1.18 1.05 1.14 1.05 .1.18 0.97 1.18 1.23 1.13 1.26 1.13 1.24 1.16 1.25 1.04 1.27 1.31 1.21 1.35 1.21 1.34 1.23 1.35 C-00 43 B -HI 44 A 45 B-HI 46 A 47 B -HI 48 B 49 8-HI 50 A 51 0.81 0.96 0.98 1.15 1.06 1.13 1.09 1.15 1.06 1.09 1.09 1.07 1.24 1.16 1.24 1.16 1.25 1.14 1.17 1.17 1.14 1.33 1.23 1.33 1.26 1.34 1.21 i
~
C-00 52 B -LO 53 B -L0 54 A 55 B-HI 56 A 57 B-HI 58 A 59 B-HI 60 0.85 0.96 1.95 1.01 1.16 1.04 1.14 1.05 1.19 1.12 1.05 1.'9 1.10 1.24 1.13 1.24 1.14 1.27 1.20 1.13 1.13 1.17 1.34 1.20 1.34 1.22 1.37 l C-00 61 B -Mi 62 B -Lt ' 63 B -HI 64 A 65 B-HI 66 A 67 B-HI 68 A 69 0.85 0.93 0.91 1.10 1.04 1.18 1.06 1.19 1.07 1.11 0.99 1.01 1.21 1.13 1.25 1.17 1.27 1.15 1.19 1.08 1.17 1.30 1.20 1.35 1.24 1.37 1.23 h Amendment No. 3 May 28, 1981 C-E PLANAR AVERAGE POWER DISTRIBUTION,1000 MWD /T, Figure f4 UNRODDEO, FULL POWlER, EQUILIBRIUM XENON ggg 7
N I FORMAT IS BOX TYPE, NO. MAX VALUE IN BOX BATCH BOXES PWR. FR. AVG. PWR BOX, RPD 1.21 28 A 69 0.30 1.00 MAX 4-PIN 1.32 12 B-HI 64 0.28 1.10 MAX 1-PIN 1.40 12 B -L0 44 0.20 1.13 WITH CORE AVG. POWER 1.00 C 40 0.12 0.71 C -00 1 C -00 2 C-00 3 C-00 4 0.60 0.77 0.81 0.82 0.95 1.05 1.08 1.07 1.01 1.11 1.14 1.14 C -00 5 C -LO 6 C -LO 7 B -HI 8 B -LU 9 B-H1 10 0.62 0.86 1.06 1.00 1.03 1.01 0.99 1.18 1.25 1.13 1.13 1.10 1.05 1.24 1.33 1.20 1.20 1.17 C-00 11 C -LO 12 A 13 B-LO 14 A 15 B-LU 16 B-LU 37 0.71 1.10 0.98 1.15 0.99 1.13 1.11 1.08 1.32 1.07 1.25 1.06 1.23 1.19 1.16 1.40 1.13 1.33 1.12 1.30 1.27 C -00 18 C-LU 19 B-LU 20 B -LU 21 A 22 B-"I 23 A 24 B-HI 25 0.62 1.10 1.16 1.20 1.04 1.15 1.01 1.12 0.99 1.32 1.26 1.28 1.11 1.21 1.08 1.19 1.05 1.40 1.34 1.36 1.16 1.29 1.13 1.26 4 C.00 26 A 27 B-L0 28 A 29 B-HI 30 A 31 B-HI 32 A 33 0.86 0.98 1.21 1.04 1.15 1.00 1.12 0.99 1.17 1.07 1.28 1.11 1.23 1.06 1.19 1.06
# 1.24 1.13 1.36 1.17 1.31 1.12 1.27 1.12 C -00 34 C 35B-LO 36 A 37 B-HI 38 A 39 B .HI 40 A 41 B-HI 42 0.60 1.06 1.15 1.03 1.16 1.00 1.12 0.99 1.11 0.95 1.25 1.25 1.09 1.23 1.06 1.19 1.06 1.18 1.01 1.32 1.32 1.16 1.31 1.13 1.27 1.12 1.25 C -00 43 B-HI 44 A 45 B -HI 46 A 47 B-H1 48 B 49 B-H i 50 A 51 0.77 1.00 0.99 1.14 1.01 1.11
'
1.11 1.11 0.98 1.04 1.13 1.06 1.21 1.07 1.19 1.18 1.17 1.04 1.11 1.20 1.12 1.29 1.13 1.27 1.26 1.25 1.10 C -00 52 B -LO53 B -LU 54 A 55 B-HA 56 A 57 &HI 58 A 59 B-HI 60 0.81 1.03 1.13 1.00 1.12 0.99 1.10 0.98 1.11 1.07 1.13 1.23 1.06 1.19 1.04 1.17 1.03 1.17 1.14 1.20 1.30 1.12 1.27 1.10 1.25 1.09 1.24 C -00 61 B-H1 62 B-LU 63 B -HI 64 A 65 B -HI 66 A 67 &HI 68 A 69 0.81 1.00 1.10 1.12 0.99 ~ 12 0.99 1.10 0.98 1.07 1.09 1.18 1.19 1.05 i.18 1.05 1.17 1.03 1.13 1.17 1.26 1.26 1.11 1.25 1.11 1.24 1.09
^
Amendment NO. 3 May 28, 1981
'
C-E f PLANAR AVERAGE POWER DISTRIBUTION,6000 MWD /T, Figure l ggg jf UNRODDED, FULL POWER, EQUILIBRIUM XENON 4.3-6
.- . - _ __
i [ \ FORMAT IS BOX TYPE, NO. MAX VALUE IN BOX BATCH BOXES PWR.FR. AVG. PWR BOX, RPD 1.20 28 A 69 0.29 0.99 MAX 4-PIN 1.28 12 B HI 64 0.29 1.12 MAX 1. PIN 1.36 12 B-L0 44 0.20 1.15 WITH CORE AVG. POWER 1.00 C 40 0.12 0.69 e C-00 1 C -00 2 C -00 3 C-00 4 0.58 0.75 0.79 0.80 0.92 1.02 1.06 1.06 0.97 1.07 1.11 1.11 C-00 5 C -LO 6 C-LO ' 7 8 -HI 8 B-LU 9 8 -H1 10 0.59 0.84 1.06 1.02 1.07 1.06 0.94 1.15 1.25 1.14 1.19 1.17 0.99 1.20 1.31 1.21 1.26 1.24 C-00 11 C-LO 12 A 13 B -LO 14 A 15 B -LU 16 B-LU 17 0.68 1.07 0.95 1.14 0.99 1.17 1.18 1.04 1.28 1.04 1.23 1.05 1.26 1.25 1.10 1.35 1.09 1.29 1.10 1.32 1.33 C -00 18 C -LU 19 B-LU 20 B-A U 21 A 22 B -Mi 23 A 24 B-HI 25 0.59 1.07 1.16 1.19 1.01 1.15 1.01 1.15 0.94 1.28 1.26 1.27 1.06 1.21 1.06 1.22 0.99 1.34 1.33 1.33 1.11 1.27 1.12 1.29 d C-00 26 A 27 8 -LO 28 A 29 B-H I 30 A 31 B -HI 32 A 33 0.84 0.95 1.20 1.02 1.14 0.99 1.13 0.99 1.14 1.04 1.27 1.08 1.21 1.03 1.20 1.04 1.20 1.09 1.34 1.13 1.28 1.09 1.26 1.09 C-00 34 C 35 B -LO 36 A 37 B -LO 38 A 39 B-HI 40 A 41 B-HI 42 0.59 1.05 1.14 1.01 1.15 0.99 1.14 0.99 1.13 0.92 1.24 1.23 1.06 * ?2 1.04 1.21 1.04 1.19 0.97 1.31 1.29 1,11 1.28 1.09 1.27 1.09 1.25 C _oo 43 B-HI 44 A 45 B-HI 46 A 47 B-HI 48 8 49 B-H1 50 A 51 0.74 1.02 0.99 1.14 0.39 1.14 1.16 1.14 0.98 1.02 1.15 1.05 1.21 1.04 1.20 1.22 1.20 1.02 i 1.07 1.21 1.10 1.27 1.09 1.27 1.30 1.27 1.08 C -00 52 B-Lo 53 B -LO 54 A 05 8 'il F6 A 57 B-HI 58 A 59 B-HI 60 0.79 1.07 1.18 1.01 1.13 0.99 1.13 0.98 1.12 1.06 1.19 1.26 1.06 1.20 1.03 1.20 1.03 1.18 1.11 1.26 1.32 1.12 1.26 1.09 1.26 1.08 1.24 C -00 61 B-HI 62 B-LU 63 B -H1 64 A 65 B-H1 66 A 67 B HI 68 A 69 0.79 1.05 1.18 1.16 0.99 1.13 0.98 1.12 0.97 1.05 1.17 1.25 1.23 1.03 1.19 1.03 1.18 1.01 1.11 1.23 1.32 1.29 1.09 1.25 1.08 1.24 1.07 O
\ ,]
Amenciment No. 3 May 28, 1981
'^ PLANAR AVERAGE POWER DISTRIBUTION,9000 MWD /T, C-E f j Figure ggg j f UNRODDED, FULL POWER, EQUILIBRIUM XENON
- - - -
-. .. _ _ . ._ _.
. ___ - -_ _ _ _ _ _
FORMAT IS BOX TYPE, NO. MAX VALUE IN BOX BATCH BOXES PWR. FR. AVG. PWR (
\ BOX, RPD 1.26 17 A 69 0.29 0.98
/
V MAX 4 PIN 1.31 63 B-HI 64 0.30 1.17 MAX 1-PIN 1.38 17 B-L0 44 0.20 1.15 WITH CORE AVG. POWER 1.00 C 40 0.12 0.68 C -00 1 C -00 2 C -00 3 C -00 4 0.58 0.75 0.80 0.82 0.90 1.02 1.07 1.08 0.93 1.06 1.11 1.12 C-00 5 C-LO 6 C.-LO 7 B-H1 8 B-LU g g-H1 10 0.55 0.81 1.03 1.07 1.15 1.18 0.84 1.06 1.20 1.20 1.28 1.29 0.88 1.09 1.24 1.24 1.32 1.34 C-oo 11 C-LO 12 A 13 8 -LO 14 A 15 D-LO 16 B-L0 17 0.62 0.96 0.86 1.09 0.98 1.21 1.26 0.92 1.13 0.94 1.15 1.03 1.30 1.31 0.95 1.17 0.97 1.20 1.07 1.35 1.38 C-oo 18 C-LO 19 B -LO 20 B-LO 21 A 22 B-HI 23 A 24 B-H 1 25 0.55 0.96 1.06 1.08 0.95 1.15 1.01 1.21 0.84 1.13 1.14 1.14 0.99 1.21 1.07 1.29 (O 0.88 1.17 1.18 B-LO 28 A 1.19 1.03 B-HI 30 A 1.26 31 B-HI 1.11 32 A 1.34 33 C-oo 26 A 27 29 0.80 0 5t7 1.09 0.95 1.13 0.99 1.17 1.01 1.06 u.A 1.14 0.99 1.10 1.03 1.22 1.04 1.09 0.97 1.19 1.03 1.24 1.07 1.28 1.08 C 34 C 35 B-LO 3G A 37 B-HI 38 A 39 B-HI 40 A 41 B-HI 42 0.58 1.03 1.09 0.95 1.13 0.99 1.20 1.02 1.18 l 0.90 1.20 1.16 0.99 1.19 1.05 1.28 1.07 1.23 ' O.94 1.25 1.20 1.03 1.25 1.09 1.33 1.11 1.29 C-00 43 B -HI 44 A 45 B-HI 46 A 47 B-H I 48 8 49 D-HI 50 A 51 0.75 1.08 0.98 1.14 0.99 1.20 1.24 1.22 1.02 1.02 1.20 1.03 1.21 1.03 1.28 1.29 1.29 1.05 1.06 1.24 1.07 1.26 1.07 1.32 1.35 1.34 1.09 C-oo 52 B -L0 53 B-L0 54 A 55 B -HI 56 A 57 B-HI 58 A 59 B-HI 60 0.80 1.16 1.22 1.01 1.17 1.02 1.21 1.02 1.19 1.07 1.28 1.30 1.07 1.22 1.07 1.28 1.07 1.24 1.11 1.33 1.35 1.11 1.28 1.11 1.34 1.12 1.30 C.co 61 B -HI 62 B -LQi3 B -HI 64 A 65 B-H1 66 A 67 B-HI 68 A 69 0.P1 1.18 1.26 1.22 1.01 1.19 1.02 1.19 1.01 1.08 1.30 1.31 1.29 1.05 1.23 1.05 1.24 1.04 1.12 1.34 1.38 1.35 1.09 1.30 1.09 1.30 1.08 l
-
Amendment No.3 May 28, 1981
'
, C-E f PLANAR AVERAGE POWER DISTRIBUTION, EOC, UNRODDED Figure ' ggg / FULL POWER, EQUILIBRIUM XENON,16500 MWD /T 4.3-8 __ . _
__ [Yj FORMAT IS BOX TYPE, NO. MAX, VALUE IN BOX BATCH BOXES PWR.FR. AVG. PWR BOX, RPD 1.28 28 A 69 0.29 0.98 MAX 4-PIN 1.44 12 B-HI 64 0.27 1.04 MAX 1-PIN 1.54 12 B-L0 44 0.20 1.12 l WITH CORE AVC. POWER 1.00 C 40 0.14 0.80 C -00 1 C 00 2 C -00 3 C-00 4 0.67 0.88 0.92 0.92 1.07 1.18 1.20 1.19 1.14 1.26 1.28 1.27 B -LU g g-H1
'
C-00 5 C -LU 6 C -LU 7 B- H I 8 10 0.69 0.93 1.13 1.05 1.03 1.00 1.12 .1.28 1.31 1.19 1.11 1.08 1.19 1.36 1.41 1.27 1.20 1.17 C -00 11 C -LU 12 A 13 B-L0 14 A 15 B-LO 16 B-LO 17 0.80 1.20 1.07 1.21 1.03 1.06 1.00 1.22 1.44 1.19 1.34 1.13 1.20 1.08 1.31 1.54 1.26 1.44 1.20 1.28 1.16 C -00 18 C-LO 19 B -LO 20 B -LO 21 A 22 B-HI 23 A 24 B-HI 25 0.70 1.19 1.21 1.27 1.11 1.15 0.93 0.92 1.12 1.44 1.31 1.37 1.22 1.27 1.03 1.03 1.53 1.42 1.47 1,29 1.36 1.10 1.11 [m} 1.19 C -oo 26 A 27 &LO 28 A 29 B -H I 30 A 31 B-HI 32 A 33 0.92 1.08 1.28 1.13 1.20 0.99 0.95 0.59 1.28 1.20 1.37 1.22 1.32 1.10 1.10 0.68 1.35 1.27 1.47 1.30 1.41 1.16 1.18 0.71 C-oo 34 C 35 B -L0 36 A 37 B-HI 38 A 39 &HI 40 A 41 B-HI 42 0.67 1.12 1.21 1.11 1.21 1.02 1.06 0.89 0.92 1.07 1.31 1.34 1.20 1.32 1.12 1.18 0.98 1.03 l 1.14 1.40 1.44 1.28 1.42 1.20 1.27 1.03 1.11 C-oo 43 B-HI 44' A 45 B -HI 46 A 47 8 -HI 48 B 49 B -HI 50 A 51 l 0.87 1.05 1.03 1.14 1.00 1.05 1.00 1.01 0.91 1.17 1.19 1.13 1.26 1.10 1.18 1.08 1.10 0.99 1.25 1.27 1.20 1.35 1.17 1.26 1.17 1.18 1.06 C-oo 52 B -LO 53 BLO 54 A 55 B-HI SG A 57 8-HI 58 A 59 B -H I 60 ! 0.90 1.03 1.07 0.92 0.95 0.88 1.01 0.94 1.06 1.19 1.12 1.19 1.02 1.10 0.96 1.10 1.01 1.13 1.26 1.21 1.28 1.09 1.17 1.02 1.18 1.08 1.22 C -oo 61 B -HI 62 B -LO 63 B-HI 64 A 65 B -H I 66 A 67 B -H I 68 A 69 0.90 0.98 0.99 0.92 0.58 0.92 0.92 1.06 0.96 1.17 1.05 1.06 1.03 0.67 1.03 1.01 1.13 1.03 1.25 1.14 1.14 1.11 0.70 1.11 1.07 1.21 1.09 _ Amendment ha. 3 May 28, 1981
'
C-E f PLANAR AVERAGE POWER DLSTRISUTION, BANK 5 Fig ne ggg j f FULL IN, FULL POWEP, EQUILIBRIUM XENON,2000 MWD /T 4.3-9 <
>
_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _ _ _ - _ _ . FORMAT IS BOX TYPE. NO. MAX VALUE IN BOX BATCH BOXES PWR. FR. AVG. PWR BOX, RPD 1.19 28 A 69 0.29 0.99 MAX 4-PIN 1.27 12 B-HI 64 0.29 1.12 MAX 1-PIN 1.34 12 B -LO 44 0.20 1.15 WITH CORE AVG. POWER 1.00 C 40 0.12 0.69 C-oo 1 C-oo 2 C -oo 3 C -00 4 0.58 0.75 0.79 0.80 0.92 1.02 1.06 1.06 0.97 1.07 1.11 1.11 C -oo 5 C -L0 6 C -L0 7 B-HI 8 B -L0 g B-HI 10 0.59 0.84 1.06 1.02 1.07 1.06 0.93 1.14 1.24 1.15 1.19 1.17 0.98 1.20 1.31 1.21 1.26 1.24 C -oo 11 C-LO 12 A 13 B-L0 14 A 15 B-LO 16 B -LO 17 0.68 1.07 0.94 1.14 0.99 1.18 1.19 1.03 1.27 1.03 1.23 1.05 1.26 1.26 1.09 1.34 1 08 1.29 1.10 1.33 1.33 C-oo 18 C--L0 19 B -LO 20 B-L0 21 A 22 8-HI 23 A 24 B HI 25 0.59 1.06 1.15 1.19 1.01 1.15 1.01 1.16 0.93 1.27 1.25 1.26 1.06 1.21 1.07 1.23 0.99 1.34 1.32 1.33 1.11 1.27 1.12 1.30 C -oo 26 A 27 B -L0 28 A 20 BHI 30 A 31 B -HI 32 A 33 0.84 0.95 1.19 1.01 1.14 0.99 1.14 0.99 1.14 1.03 1.27 1.07 1.21 1.03 1.20 1.04 1.19 1.09 1.33 1.13 1.28 1.09 1.26 1.10 0-00 34 C 35 B -L0 36 A 37 B-HI 38 A 39 BHI 40 A 41 B-HI 42 0.59 1.05 1.14 1.00 1.15 0.99 1.15 0.99 1.13 0.92 1.24 1 23 1.05 1.21 1.04 1.21 1.04 1.19
- 0.97 1.31 . 29 1.11 1.28 1.09 1.27 1.10 1.25 C-oo 43 B-HI 44 A 45 B-H1 46 A 47 B-HI 48 B 49 BHI 50 A 51 0.74 1.02 0.99 1.14 0.99 1.14 1.16 1.14 0.98 1.02 1.15 1.05 1.21 1.04 1.21 1.22 1.20 1.03 1.07 1.21 1.10 1.27 1.09 1.27 1.30 1.27 1.08 C-oo 52 B -L0 53 B -L0 54 A 55 B-HI 56 A 57 B-HI 58 A 59 B-HI 60 0.79 1.08 1.18 1.01 1.14 0.99 1.14 0.98 1.13 1.06 1.20 1.26 1.07 1.20 1.03 1.20 1.03 1.18 1.11 1.26 1.33 1.12 1.26 1.09 1.27 1.09 1.25 C -oo61 B -HI 62 B -L0 63 B-HI 64 A 65 B -HI 66 A 67 B-H1 68 A 69 0.80 1.06 1.18 1.16 0.99 1.13 0.99 1.12 0.98 1.06 1.17 1.26 1.23 1.04 1.19 1.04 1.18 1.02 1.11 1.24 1.33 1.30 1.09 1.25 1.09 1.25 1.07 i
Amendment No. 3 May 28, 1981 C-E PLANAR AVERAGE POWER DISTRIBUTION,9000 MWD /T, Figure c B / BANK 5 FULL IN, FULL POWER, EQUILIBRIUM XENON 1 4.3-10
-
. _ _ _ _ _ _ _ _ _ _ _ _ _
,m
, FORMAT IS BOX TYPE, NO. MAX VALUE IN BOX BATCH BOXES PWR. FR. AVG. PWR l BOX, RPD 1.25 17 A 69 0.28 0.94 D) MAX 4-PIN MAX 1. PIN 1.33 1.40 63 17 8 -HI 64 B -LO 44 0.28 0.21 1.10 1.20 WITH CORE AVG. POWER 1.00 C 40 0.13 0.73 C -00 1 C -00 2 C-00 3 C-00 4 0.62 0.80 0.85 0.86 0.97 1.09 1.14 1.14 1.02 1.13 1.18 1.18 C -00 5 C -t0 6 C -LO 7 3 -HI 8 B-LO 9 B-HI 10 0.61 0.89 1.13 1.13 1.20 1.21 0.95 1.!7 1.31 1.25 1.32 1.32 0.99 1.22 1.36 1.30 1.37 1.38 C -00 11 C -L0 12 A 13 B-L0 14 A 15 B-L0 16 B-L0 17 0.69 1.09 0.96 1.18 1.02 1.23 1.25 1.04 1.27 1.04 1.25 1.07 1.32 1.33 1.09 1.34 1.08 1.31 1.12 1.38 1.40 0-00 18 C-LO 19 B -LO 20 B-L0 21 A 22iB-HI 23 A 24 B-HI 25 0.61 1.08 1.19 1.20 1.00 1.14 0.94 1.04
. 0.95 1.28 1.28 1.28 1.05 1.22 1.02 1.19 0.99 1.33 1.34 1.33 1.09 1.28 1.06 1.24 v 0-00 26 A 27 8 -LO 28 A 29 &HI 30 A 31 B-HI 32 A 33 0.88 0.96 1.21 1.01 1.15 0.94 0.98 0.59 1.17 1.04 1.28 1.07 1.22
- 00 1.09 0.73 1.22 1.08 1.34 1.11 1.28 1.04 1.13 0.76 34 C 35 B _t 36 A 37 B-HI 38 A 39 B-HI 40 A 41 B-HI 42 C-o03
- 0. 1.12 1. 8 1.00 1.16 0.98 1.13 0.89 0.95 0.97 1.31 1.25 1.05 1.22 1.02 1.20 0.97 1.06 1.02 1,36 1.31 1.09 1.28 1.06 1.26 1.01 1.00 C-oo 43 B -HI 44 A 45 B -HI 46 A 47 B-HI 48 B 49 B -HI 50 A 51 0.80 1.14 1.02 1.14 0.95 1.12 1.16 1.11 0.91 1.09 1.26 1.08 1.22 1.00 1.20 1.22 1.19 0.96 1.14 1.31 1.12 1.27 1.05 1.26 1.28 1.24 1.01 l C-oo 52 B-L0 53 B-LO 54 A 55 B-HI 56 A 57 aHI 58 A 59 B -H I 60 l 0.85 1.20 1.24 0.93 0.98 0.88 1.10 0.95 1.11 l 1.14 1.32 1.33 1.02 1. 0 0.97 1.19 0.99 1.15 1.18 1.38 1.39 1.06 1.14 1.01 1.24 1.04 1.22 C-oo 61 B -HI 62 B -LO 63 B-HI 64 A 65 B -H I 66 A 67 BHI 68 A 69 0.86 1.21 1.25 1.05 0.59 0.96 0.92 1.10 0.94 1.14 1.32 1.33 1.20 0.73 1.06 0.96 1.15 0.98 1.19 1.38 1.40 1.24 0.77 1.10 1.01 1.21 1.02
_ Amendment No. 3 May 28, 1981
-
PLANAR AVERAGE POWCR DISTRIBUTION, BANK 5 Figure C-E / FULL IN, FULL POWER, EQUILIBRIUM XENON,14000 MWD /T ggg 7 4.3-11 l
FORMAT IS BOX TYPE, NO. MAX. VALUE IN BOX BATCH BOXES PWR. FR. AVG. PWR p A 69 0.30 1.02 I BOX,RPD 1.26 66 lV MAX 4-PIN 1.34 42 B -HI 64 0.29 0.19 1.14 1.06 MAX 1-PIN 1.44 42 B -LO 44 WITH CORE AVG. POWER 1.00 C 40 0.13 0.73 C -oo 1 C -00 2 C -oo 3 000 4 0.60 0.80 0.85 0.86 0.96 1.08 1.12 1.12 1.02 1.15 1.20 1.20 C -00 5 C -LO 6 C -L0 7 B-HI 8 B-LO 9 8-HI 10 0.61 0.82 0.99 0.93 0.96 0.96 0.99 1.13 1.15 1.03 1.04 1.04 1.06 1.20 1.24 1.11 1.12 1.12 C -00 11 C-L0 12 A 13 aLO 14 A 15 BLO 16 b-LO 17 0.71 1.06 0.95 1.07 0.84 1.04 1.03 1.08 1.27 1.05 1.20 0.92 1.17 1.13 1.16 1.35 1.12 1.28 0.96 1.25 1.21 C -oo 18 C -LO 19 B -LO 20 B -LO 21 A 22 8 -HI 23 A 24 B -H1 25 0.62 1.05 1.06 1.11 1.02 1.13 1.04 1.14 0.99 1.27 1.14 1.22 1.12 1.24 1.14 1.26 O 1.05 1.35 1.23 1.30 1.18 1.33 1.21 1.35 C -oo 26 A 27 B-LO 28 A 29 B -HI 30 A 31 B-H1 32 A 33 0.82 0.95 1.12 0.90 1.15 1.07 1.22 1.10 1.13 1.06 1.21 0.97 1.25 1.16 1.31 1.21 1.19 1.12 1.30 1.02 1.35 1.23 1.41 1.28 C-00 34 C 35 B -LO 36 A 37 B HI 38 A 39 B-HI 40 A 41 RHI 42 0.60 0.99 1.06 1.01 1.15 1.06 1.20 1.12 1.25 0.95 1.15 1.19 1.09 1.26 1.14 1.30 1.22 1.34 1.02 1.23 1.28 1.16 1.35 1.22 1.39 1.29 1.44 C _oo 43 B -H I 44 A 45 B -HI 46 A 47 B-H1 48 B 49 B-HI 50 A 51 0.79 0.93 0.83 1.12 1.07 1.19 1.17 1.23 1.13 1.07 1.02 0.91 1.23 1.18 1.29 1.25 1.33 1.21 1.14 1.10 0.96 1.32 1.24 1.39 1.35 1.43 1.29 C -ooS2.B-LU 53 B -LU 54 A 55 B -H I 56 A 57 B-HI 58 A 59 B-HI 60 0.84 0.96 1.04 1.02 1.21 1.11 1.22 1.10 1.23 1.11 1.05 1.16 1.12 1.31 1.19 1.32 1.19 1.32 1.18 1.13 1.24 1.19 1.40 1.27 1.42 1.26 1.42 C-oo 61 B-HI 62 B -LO63 B -HI 64 'A 65 B-H1 66 A 67 B-H1 68 A 69 0.85 0.94 1.02 1.14 1.09 1.26 1.13 1.22 0.96 1.11 1.01 1.12 1.25 1.18 1.33 1.23 1.32 1.04 1.18 1.09 1.21 1.34 1.26 1.43 1.31 1.41 1.10 Amendment No.3 May 28, 1981 PLANAR AVERAGE POWER DISTRIBUTION, PART-LENGTH ' C-E Figure ROD AS IF FULL LENGTH, FULL POWER, EQUILIBRIUM XENON, [7bhdt /j/ 2000 MWD /T 4.3-12
;
1
n FORMAT IS BOX TYPE, NO. MAX. VALUE IN BOX BATCH BOXES PWR. FR. AVG. PWR BOX,RPD 1.25 49 A 69 0.29 0.98 [V')
'
MAX 4-PIN MAX 1-PIN 1.32 1.40 49 49 B -HI 64 B-LO 44 0.29 0.20 1.15 1.13 WITH CORE AVG. POWER 1.00 C 40 0.12 0.69 C-00 1 C -00 2 C -00 3 C -00 4 0.58 0.74 0.80 0.81 0.91 1.01 1.07 1.07 0.96 1.06 1.12 1.12 C-00 5 C -LO 6 C -LO 7 B -HI 8 B -LO 9 B -HI 10 0.59 0.84 1.03 0.98 1.06 1.07 0.94 1.13 1.19 1.08 1.19 1.18 0.99 1.18 1.26 1.14 1.25 1.24 C-00 11 C -LO 12 A 13 B-LO 14 A 15 B -LU 16 B -L0 17 0.68 1.06 0.93 1.09 0.84 1.15 1.19 1.03 1.26 1.01 1.19 0.92 1.26 1.27 1.09 1.33 1.06 1.24 0.95 1.33 1.34 C-00 18 C -LU 19 B -LU 20 B-L0 21 A 22 B -H I 23 A 24 B-HI 25 0.59 1.06 1.13 1.14 0.97 1.12 1.02 1.19 0.94 1.26 1, '" 1.20 1.02 1.21 1.08 1.25 0.99 1.33 1.28 1.27 1.07 1.27 1.13 1.32
'd C -00 26 A 27 8 -LO 28 A 29 B -HI 30 A 31 B-HI 32 A 33 0.83 0.93 1.14 0.86 1.12 1.01 1.19 1.05 1.13 1.01 1.21 0.94 1.21 1.07 1.26 1.10 1.18 1.06 1.28 0.96 1.27 1.13 1.33 1.16 C -00 34 C 35 B-LO 36 A 37 B HI 38 A 39 B-HI 40 A 41 B-HI 42 0.58 1.03 1.09 0.97 1.12 1.01 1.21 1.06 1.22 0.91 1.19 1.19 1.02 1.21 1.08 1.29 1.12 1.28 0.96 1.26 1.24 1.07 1.27 1.14 1.36 1.18 1.35 C -00 43 B -HI 44 A 45 B -HI 46 A 47 B-HI 48 B 49 B -HI50 A 51 0.74 0.98 0.84 1.11 1.02 1.20 1.25 1.23 1.06 1.01 1.08 0.92 1.20 1.07 1.28 1.32 1.30 1.11 1.06 1.15 0.95 1.27 1.13 1.36 1.40 1.37 1.17 C-00 52 B -LO53 B-LO 54 A 55 B-HI 56 A 57 B-HI 58 A 59 B -HI 60 l 0.79 1.07 1.16 1.02 1.19 1.06 1.23 1.05 1.18 1.07 1.19 1.26 1.07 1.26 1.11 1.30 1.11 1.27 1.12 1.25 1.33 1.13 1.33 1.17 1.37 1.17 1.33 C-00 61 B-HI 62 B -LO63 B-HI 64 A 65 0-HI 66 A 67 B-H I 68 A 69 0.81 1.06 1.19 1.19 1.04 1.22 1.07 1.18 0.90 1.07 1.18 1.26 1.25 1.09 1.28 1.12 1.26 0.96 1.12 1.24 1.34 1.32 1.15 1.35 1.18 1.33 0.99 -
Am.;ndment No.3 May 28, 1981
'
C-E / PLANAR AVERAGE POWER DISTRIBUTION,9000 MWD /T Figure gg t 7f PART LENGTH RODS AS IF FULL LENGTH, FULL POWER, EQUILIBRIUM XENON 4.3-13
_ FORMAT IS BOX TYPE, NO. MAX. VALUE IN BOX BATCH BOXES PWR. FR. AVG. PWR BOX, RPD 1.35 49 A 69 0.29 0.98 MAX 4-PIN 1,42 49 B -HI 64 0.31 1.19 MAX 1-PIN 1.49 49 B -LO 44 0.20 1.14 WITH CORE AVG. POWER 1.00 C 40 0.11 0.66 C-00 1 C -00 2 C -00 3 C -00 4 0.56 0.73 0.79 0.81 0.87 0.99 1.07 1.08 0.91 1.03 1.11 1.12 C-00 5 C-LO 6 C-L0 7 B -HI 8 B -L0 9 B-H I 10 0.54 0.78 0.99 1.01 1.12 1.16 0.84 1.04 1.14 1.13 1.26 1.28 0.88 1.08 1.19 1.17 1.31 1.34
~
C-oo 11 C -LO 12 A 13 8-LO 14 A 15 B -L0 16 B -LO 17 0.31 0.96 0.84 1.04 0.83 1.19 1.27 0.92 1.13 0.92 1.10 0.93 1.30 1.33 0.96 1.17 0.96 1.15 0.96 1.36 1.41 l C -oo18 C -L0 19 B -LO 20 B -Lo 21 A 22 B -HI 23 A 24 B - H .' 25 0.54 0.95 1.04 1.05 0.92 1.12 1.03 1.25 i
- 0.84 1.13 1.11 1.10 0.98 1.22 1.09 1.32 0.88 1.17 1.17 1.16 1.02 1.27 1.14 1.38 l
C-00 26 A 27 B-Lu 28 A 29 B-HI 30 A 31 B-HI 32 A 33 0.78 0.85 1.06 0.81 1.11 1.01 1.23 1.07 1.04 0.92 1.10 0.90 1.21 1.09 1.31 1.11 1.08 0.96 1.16 0.93 1.26 1.13 1.37 1.16 C -00 34 C 35 B-L0 36 A 37 yHI 38 A 39 B-HI 40 A 41 B-HI 42 0.56 0.99 1.04 0.92 1.12 1.02 1.28 1.10 1.28 0.87 1.15 1.10 0.98 1.21 1.11 1.38 1.16 1.35 0.91 1.19 1.15 1.02 1.26 1.16 1.43 1.21 1.41 C-oo 43 B -HI 44 A 45 B-HI 46 A 47 &HI 48 B 49 B-HI 50 A 51 0.72 1.02 0.83 1.12 1.02 1.27 1.35 1.33 1.11 0.99 1.13 0.93 1.22 1.09 1.38 1,42 1.40 1.15 l 1.04 1.18 0.96 1.26 1.13 1.43 1.49 5.47 1.20 l O oo 52 B -LO 53 8 -L0 54 A 55 B-HI 56 A 57 B-H1 58 A 59 B-H1 60 0.79 1.13 1.20 1.03 1.23 1.10 1.32 1.10 1.26 i 1.34 1 1.07 1.27 1.31 1.09 1.31 1.16 1.40 1.17 1.11 1.32 1.36 1.14 1.37 1.21 1.46 1.22 1.40 C -oo 61 B -HI 62 B-L0 63 B-HI 64 A 65 D-H1 66 A 67 B-H1 68 A 69 0.81 1.16 1.27 1.26 1.07 1.29 1.11 1.25 0.94 1.08 1.28 1.33 1.32 1.12 1.35 1.15 1.34 1.02 1.41 1.39 1.16 1.42 1.20 1.40 1.04 1.12 ( 1.34 i . kandment No.3 May 28, 1981 PLANAR AVERAGE POWER DISTRIBUTION, PART LENGTH Figure C-E f ROD AS IF FULL LENGTH, FULL POWER, EQUILIBRIUM I/ XENON,14000 MWD /T 4.3-14
- . .
. . - _ - _ .- - __ -. .. .
O' v FORMAT IS BOX TYPE, NO. MAX. VALUE IN BOX BATCH BOXES PWR. FR. AVG PWR BOX, RPD MAX 4-PIN 1.24 1.45 26 12 A 8-HI 69 64 0.29 0.27 0.96 1.06 MAXg) PIN 1.55 12 B -L0 44 0.20 1.11 WITH CORE AVG, POWER 1.00 C 40 0.14 0.82
! '
C-00 1 C -00 2 C -00 3 C -00 4 0.68 0.89 0.93 0.94 1.07 1.19 1.22 1.22 1.14 1.27 1.30 1.30 C -00 5 C-LU 6 C-LU 7 B-HI 8 B -LO 9 B-HI 10 , 0.71 0.94 1.12 1.02 1.03 1.01 1.14 1.28 1.28 1.12 1.10 1.10 1.22 1.36 1.37 1.21 1.20 1.19 C-00 11 C-L0 12 A 13 B-LO 14 A 15 aLO 16 aLO 17 0.82 1.22 1.07 1.17 0.88 1.04 1.00 1.25 1.45 1.18 1.31 0.96 1.14 1.09 1.34 1.55 1.25 1.41 1.01 1.22 1.17 C-00 18 C -LO 19 B -L0 20 E -LO21 A 22 BHI 23 A 24 B -HI 26 0.71 1.21 1.20 1.23 1.08 1.12 0.93 0.94 1.14 1.45 1.29 1.34 1.19 1.23 1.03 1.05 1.22 1.54 1.39 1.43 1.26 1.32 1.09 1.12
'
C -00 26 A 2'r d-L0 28 A 29 B-HI 30 A 31 B -HI 32 A 33 0.93 1.07 1.24 0.96 1.17 1.01 0.98 0.62 1.28 1.19 1.34 1.04 1.26 1.10 1.13 0.71 1.35 1.26 1,43 1.10 1.35 1.18 1.21 0.74 ! ! ' C -00 3d C 35 B-LO 36 A 37 B -HI 38 A 39 B -HI 40 A 41 B -HI 42 O.63 1.10 1.17 1.07 1.17 1.03 1.11 0.94 0.98 1.07 1.27 1.31 1.17 1.26 1.12 1.21 1.04 1.12 1.14 1.36 1.40 1.24 1.35 1.19 1.30 1.10 1.19 C -00 43 B -HI 44 A 45 B -HI 46 A 47 B -HI 48 8 49 B -HISO A 51 0.88 1.02 0.87 1.11 1.01 1.10 1.07 1.09 0.99 1.18 1.11 0.96 1.22 1.12 * ~1 1.14 1.18 1.07 1.26 1.20 1.00 1.31 1.18 s.30 1.24 1.27 1.14 C -00 52 B-LO 53 B-LO 54 A 55 B-HI 56 A 57 B-HI 58 A 59 B =HI 60 0.92 1.02 1.04 0.92 0.98 0.94 1.08 1.00 1.11 1.21 1.12 1.14 1.01 1.12 1.03 1.18 1.07 1.19 1.29 1.20 1.22 1.08 1.20 1.09 1.27 1.14 1.27 l C -00 61 B -Hl62 B -LU63 B -H164 A 65 B -HI 66 A 67 E-HI E A 69 0.92 1.00 0.99 0.94 0.61 0.98 0.99 1.10 0.88 1.20 1.07 1.06 1.04 0.71 1.11 1.09 1.18 0.95 1.28 1.16 1.15 1.12 0.74 1.19 1.15 1.27 1.00
Amendment No.3 May 28, 1981 C-E PLANAR AVERAGE POWER DISTRIBUTION BOC, PART LENGTH Figure ! f RODS AS IF FULL LENGTH, BANK 5 FULL IN, FULL POWER, EQUILIBRIUM XENON,2000 MWD /T
-
4.3-15
. .
. , - _ . . _ - _._m
__ FORMAT IS BOX TYPE, NO. MAX. VALUE IN BOX BATCH BOXES PWR.FR. AVG. PWR O BOX, RPD MAX 4 PIN MAX 1-PIN 1.27 1.43 1.50 20 12 12 A B-HI B-LO 69 64 44 0.28 0.28 0.21 0.93 1.07 1.18 WITH CORE AVG. POWER 1.00 C 40 0.13 0 76 i i C-00 1 C-00 2 C -00 3 C -00 4 0.64 0.81 0.86 0.87 1.00 1.10 1.15 1.14 i 1.06 1.16 1.20 1.20 C-00 5 C-LO 6 C =LO 7 8 -HI 8 B -LO 9 B -HI 10 0.68 0.94 1.14 1.06 1.12 1.12 1.07 1.26 1.31 1.15 1.23 1.21 1.13 1.32 1.38 1.22 1.30 1.28
'
C -00 11 C -LO 12 A 13 B-L0 14 A 15 B-L0 16 B -L0 17 0.78 1.21 1.04 1.19 0.88 1.15 1.17
? .18 1.43 1.13 1.29 0.95 1.24 1.24 1.25 1.50 1.19 1.36 0.97 1.31 1.32 C -00 18 O-LO 19 B-LO 20 B -L0 21 A 22 B HI 23 A 24 B HI 25 0.68 1.21 1.27 1.25 1.03 1.11 0.93 1.00 1.07 1.43 1.35 1.35 1.10 1.19 1.00 1.13 1.13 1.50 1.43 1.42 1.16 1.25 1.05 1.20 O' C-00 26 A 27 B -LO 28 A 29 B-HI 30 A 31 B -HI 32 A 33 0.94 1.04 1.26 0.92 1.14 0.96 0.98 0.61 1.26 1.13 1.35 1.03 1.21 1.02 1.10 0.72 1.32 1.19 1,42 1.05 1.27 1.07 1.15 0.75 C -00 34 C 35 B L0 36 A 37 B -HI 38 A 39 B-HI 40 A 41 B -HI 42 0.65 1.13 1.19 1.02 1.14 0.99 1.12 0.91 0.97 1.01 1.30 1.30 1.10 1.21 1.03 1.20 0.99 1.08 1.06 1.38 1.36 1.16 1.28 1.09 1.26 1.04 1.14
~
C -00 43 B -HI 44 A 45 B-HI 46 A 47 B -HI 48 B 49 B -HI50 A 51 0.81 1.06 0.88 1.11 0.96 1.11 1.15 1.11 0.94 1.10 1.15 0.95 1.19 1.02 1.20 1.21 1.18 0.99 1.16 1.22 0.97 1.25 1.07 1.26 1.29 1.24 1.04 C -00 52 B-L0 53 B -LO54 A 55 B -HI 56 A 57 E-HI 58 A 59 B-HI 60 0.86 1.13 1.16 0.92 0.98 0.91 1.10 0.96 1.08 1.14 1.23 1.24 1.00 1.10 0.99 1.18 1.01 1.15 1.20 1.30 1.31 1.05 1.15 1.04 1.24 1.07 1.21 ( C -00 61 B -HI62 B-L0 63 B-HI 64 A 65 B -HI 66 A 67 8 -HI 68 A 69 0.94 ' O.87 1.11 1.16 1.01 0.60 0.97 1.07 0.83 1.14 1.21 1.24 1.14 0.72 1.09 0.99 1.15 0.88 1.20 1.28 1.31 1.20 0.76 1.14 1.04 1.21 0.91 Amendment No.3 May 28, 1981 C-E PLANAR AVERAGE POWER DISTRIBUTION,9000 MWD /T Figure g PART-LENGTH RODS AS IF FULL LENGTH, BANK 5 FULL IN, s'U FULL POWER, EQUILIBRIUM XENON 4.3-16
. , - . . - . - - - - - . . . _ - - - - - . - . -
'
\ FORMAT IS BOX T.YPE, NO. MAX. VALUE IN BOX BATCH BOXES PWR. FR. AVG, PWR BOX, RPD 1.26 17 A 69 0.28 0.93 MAX 4-PIN 1.34 63 B -HI 64 0.29 1.12 MAX 1-PIN 1.41 17 B -LO 44 0.21 1.19 WITH CORE AVG. POWER 1.00 C 40 0.13 0.74 C-00 1 C -00 2 C -00 3 C-00 4 0.62 0.80 0.86 0.88 0.96 1.08 1.15 1.16 1.01 1.13 1.19 1.20 C -00 5 C-LO 6 C-L0 7 8 -HI 8 B -L0g B-HI 10 , 0.62 0.89 1.10 1.09 1.19 1.22 0.96 1.17 1.26 1.20 1.31 1.33 1.00 1.21 1.31 1.25 1.37 1.39 C-00 11 C -LO 12 A 13 B-LO 14 A 15 BLO 16 B -LU 17 0.70 1.09 0.95 1.13 0.87 1.20 1.26 1.05 1.28 1.02 1.20 0.96 1.32 1.34 1.10 1.33 1.06 1.26 0.99 1.37 1.41 C -0018 C-LO 19 B -LO 20 B -LU21 A 22 B -HA 23 A 24 B-ML 25 0.62 1.09 1.18 1.16 0.97 1.12 0.94 1.06 0.96 1.28 1.24 1.23 1.01 1.18 1.02 1.21 1.00 1.33 1.31 1.29 1.06 1.24 1.06 1.25 '
C -00 26 A 27 8 -LO 28 A 29 B-HI 30 A 31 B-HI 32 A 33 0.88 0.95 1.16 0.86 1.13 0.96 1.02 0.62 1.16 1.02 1.24 0.96 1.20 1.02 1.13 0.75 ' 1.21 1.06 1.30 0.99 1.25 1.06 1.18 0.79 C-00 34 C 35 B -LO 36 A 37 B-HI 38 A 39 B-HI 40 A 41 B-HI 42 , ' 0.63 1.10 1.13 0.97 1.13 1.00 1.19 0.95 1.03 O.97 1.26 1.20 1.02 1.20 1.06 1.28 1.04 1.15 1.C1 1.31 1.26 1.06 1.26 1.11 1.34 1.09 1.19 C -00 43 B -HI 44 A 45 B -HI 46 A 47 B -HI 48 B 49 B -HI 50 A 51 0.80 1.10 0.8'/ 1.11 0.97 1.19 1.25 1.20 0.99 1.08 1.21 0.97 1.18 1.02 1.28 1.31 1.29 1.04 1.13 1.26 0.99 1.23 1.06 1.33 1.38 1.35 1.09 C-00 52 B-LO 53 B -LO 54 A 55 B -H1 56 A 57 B -HI 58 A 59 B -HI 60 0.85
- 20
. 1.21 0.94 1.02 0.95 1.20 1.02 1.16 1.15 1.32 1,32 1.02 1.14 1.04 1.29 1.08 1.23 1.19 1.37 1.38 1.06 1.18 1.09 1.34 1.12 1.29 C-00 61 B-H1 62 B -LU 63 B -H1 64 A 65 B -HI 66 A 67 8-HI 68 A 69 0.87 1.21 1.25 1.07 0.62 1.04 0.99 1.16 0.87 1.16 1.33 1.34 1.1; 0.76 1.15 1.04 1.23 0.95 1.20 1.38 1.41 1.26 0.79 1.20 1.09 1.29 0.97 h
D Amendment NO.3 May 28, 1981 C-E PLANAR AVERAGE POWER DISTRIBUTION, PART LENGTH Figure RODS AS IF FULL LENGTH, BANK 5 FULL IN, FULL POWER, EQUILIBRIUM XENON,14000 MWD /T 4.3-17 l
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Amendment No.3 fiay 28, 1981 AXIAL POWER DISTRIBUTION AT 4000 MWD /T, UNRODDED 9
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_ Amendment No.3 May 28, 1981
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C-E AXIAL POWER DISTRIBUTION AT 9000 MWD /T , UNRODDED Figure [ 4.3-21
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' C-E AXIAL POWER DISTRIBUTION AT Figure n 13000 MWD /T, UNRODDED
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. . - - . _-. . _ - - . . ._ . _ . - _ - - . . ___ _.
X.X X BOX AVEHAGE (BOX POWERiLORE AVERAGE POWER) X.XX Md. ENTHALPY RISE FACTOR (MAX.OF AVERAGE POWER OF 4 NElGHBORING X.XX PINS / CORE AVERAGE POWER)
,-- i MAX.ONE PIN FACTOR (MAX. PIN POWER / CORE AVERAGE POWER)
( ,
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x.s No. O F No. O F SHIMS PER FUJL TYPE ASSEMBLil5 ENRICHMENT ASSEMBLY D 56 3.60 0 D* 24 3.05 0 D D D D 0.72 0.85 0.89 0.97 1.13 1.13 1.23 1.28 1.19 1.19 1.29 1.32 D D D BH BH D* 0.74 1.02 1.18 0.79 0.80 1.13 1.20 1.28 1.31 0.88 0.92 1.25 1.27 1.33 1.37 0.90 0.94 1.30 D D* BL BH BL BL BH 0.83 1.09 0.87 0.83 0.81 0.79 0.83 1.28 1.21 0.94 0.91 0.93 0.87 0.91 1.34 1.26 0.96 3.94 0.97 0.92 0.94 B C+ BL D* B D*
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~
.85 0.92 0.90 1.21 0.94 1.23
( ) 0.95 0.99 1.04 1.34 1.04 1.36 C/ 0.97 1.09 1.07 1.40 1.07 1.42 BL C BH C BH 0.92 1.17 1.02 1.21 1.05 1.08 1.33 1.10 1.35 1.13 1.14 1.38 1.13 1.41 1.16 C C*9 C C 1.24 1.17 1.24 1.24 1.36 1.25 1.34 1.34 1.41 1.29 1.39 1.38 C C+L C+L 1.24 1.24 1.26 1.34 1.34 1.33 1.39 1.39 1.38 C+L C 1.24 1.21 1.34 1.31 1.39 1.37 A 0.95 0.99 _ 1.02 C-E PLANAR AVERAGE POWER DISTRIBUTION AT THE BEGINNING
-l / Figure OF THE SECOND CYCLE, UNRODDED
,
Amendment No.3 May 28, 1931
.
X.XF BOX AVERAGE (BOX POWER,00RE AVERAGE POWER) X.X) MAX. ENTHALPY RISE FACTOR (MAX.OF AVERAGE POWER OF 4 NEIGHBOPING X.XX PINS / CORE AVERAGE POWER)
~N MAX.ONE-PIN FACTOR (MAX. PIN POWER / CORE AVERAGE POWER)
L] D D D D 0.71 0.83 0.8; 0.95 1.11 1.12 1.21 1.26 1.17 1.17 1.28 1.30 D D D By B H D* 0.73 1.00 1.16 0.79 0.80 1.13 1.18 1.26 1.29 0.87 0.92 1.24 1.25 1.31 1.35 0.90 0.94 1.30 D D* BL BH B L 9L BH 0.81 1.08 0.87 0.84 0.82 0.81 0.85 1.26 1.19 0.93 0.91 0.94 0.89 0.93 1.32 1.25 0.96 0.94 0.98 0.94 0.95 BH C+ BL D* B D* /^'s 0.86 .94 0.92 1.22 k.95 1,24 (d) 0.94 0.97 1.00 1.09 1.05 1.08 1.35 1.41 1.05 1.08 1.37 1.42 BL C BH C BH 0.94 1.17 1.03 1.22 1.06 1.09 1.33 1.10 1.35 1.13 1.14 1.38 1.14 1.41 1.16 C Cy* C C 1.24 1.17 1.23 1.24 1.36 1.25 1.33 1.33 1.42 1.29 1.38 1.38 C C+L C+ L 1.23 1.24 1.25 1.33 1.33 1.32 1.38 1.39 1.38 C+L C 1.24 1.20 1.33 1.31 1.38 1.36 A 0.95 0.99 1.02 O C-E / PLANAR AVERAGE POWER DISTRIBUTION AT 50 MWD /T Figure ggg 7f OF THE SECOND CYCL 8, UNRODDED Amendment No.3 May 28, 1981
X.XX BOX AVER AGE (BOX POWER, CORE AVERAGE POWER) X.XX MAY ENTHALPY RISE FACTOR (MAX.OF AVERAGE POWER OF 4 NEIGHBORING X.XX PINS / CORE AVEMAGE POWER) r~'y MAX. ONE PIN FACTOR (MAX. PIN POWER / CORE AVER AGE POWER) I )
'wJ D D D D 0.72 0.86 0.90 0.95 1.12 1.13 1.22 1.23 1.17 1.19 1.28 1.28 D D D BH BH D*
0.73 0.99 1.16 0.85 0.86 1.15 1.16 1.26 1.28 0.91 0.95 1.24 1.23 1.32 1.33 0.04 0.97 1.29 D D* BL BH BL BL BH 0.81 1.08 0.91 0.89 0.88 0.87 0.90 1.23 1.20 0.97 0.95 0.97 0.93 0.95 1.30 1.25 1.00 0.98 0.99 0.97 0.98 BH C+ L BL D* BH D* (~';1 0.91 1.00 0.96 1.21 0.95 1.21 1 0.96 1.08 1.04 1.31 1.03 1.31
\~2 1.00 1.15 1.07 1.36 1.05 1.36 BL C BH C BH 0.97 1.13 1.00 1.14 1.01 1.06 1.27 1.06 1.26 1.06 1.09 1.31 1.09 1.31 1.10 C C+H C C 1.16 1.12 1.14 1.14 1.26 1.17 1.25 1.23 1.31 1.21 1.30 1.27 C C+L C+L 1.15 1.17 1.21 1.25 1.26 1.30 1.30 1.30 1.34 C*L C 1.20 1.14 1.31 1.23 1.37 1.28 A
0.92 0.96 _ 0.99
-
"
I
/ PLANAR AVERAGE POV/ER DISTRIBUTION,6000 MWD /T 9" j@@ / / OF THE SECOND CYCLE, UNRODDED 4.3-26 i
Amendment No.3 May 28, 1981
X.XX BOX AVERAGE (BOX POWER / CORE AVERAGE POWER) X.XX MAX. ENTHALPY RISE FACTOR (MAX. OF AVERAGE POWER OF 4 NEIGHBORING X.XX PINS / CORE AVERAGE POWER) fN MAX.ONE PIN FACTOR (MAX. PIN POWER / CORE AVERAGE POWER) ( ) v D D D D 0.74 0.88 0.92 0.96 1.13 1.16 1.23 1.23 1.19 1.22 1.29 1.28 D D D BH BH D' O.75 1.00 1.17 0.88 0.89 1.16 1.17 1.27 1.29 0.94 0.97 1.24 1.23 1.33 1.34 0.97 0.99 1.29 D D* Bt BH BL BL BH 0.83 1.09 0.94 0.92 0.91 0.91 0.92 1.23 1.22 1.00 0.97 0.99 0.96 0.97 1.29 1.27 1.02 1.00 1.01 0.99 1.00 BH C*t BL D' BH D* es 0.93 1.03 0.98 1.21 0.96 ? 20 (d I 0.98 1.01 1.11 1.17 1.04 1.07 1.28 1.33 1.02 1.05 1.28 1.32 BL C BH C BH 0.98 1.12 0.99 1.11 0.99 1.9 1.24 1.04 1.22 1.03 1.Ld 1.28 1.07 1.27 1.07 C Cy* C C 1.13 1.08 1.09 1.09 1.23 1.13 1.19 1.18 1.27 1.17 1.23 1.22 C C+L C+L 1.10 1.11 1.14 1.20 1.19 1.23 1.24 1.23 1.27 l C+L C 1.14 1.08 ! 1.25 1.17 l 1.30 1.21 l
^
l 0.89 0.93 _ 0.95 ; C-E /, PLANAR AVERAGE POWER DISTRIBUTION AT THE END Figure
'l l OF THE SECOND CYCLE, UNRODDED 7/
Amendment No.3 May 28, 1981
X.X X BOX AVERAGE (BOX POWER l CORE AVERAGE POWER) X.XX MAX. ENTHALPY RISE FACTOR (MAX.OF AVERAGE POWER OF 4 NEIGHBORING X'XX PINS / CORE AVERAGE POWER) ) s' MAX. ONE. PIN FACTOR (MAX. PIN POWER / CORE AVERAGE POWER) (
\ m,)
No. O F No. O F SHIMS PER ,
'
FUEL TYPE ASSEMBLIES ENRICHMENT ASSEMBLY E 72 3.30 0 E' 8 3.00 0 E E E E 0.58 0.74 0.86 0.86 0.95 1.07 1.18 1.16 1.00 1.13 1.23 1.22 E E E BL E C+ L 0.61 0.85 1.05 0.76 1.23 0.95 1.01 1.13 1.20 0.85 1.36 1.06 1.06 1.17 1.26 0.88 1.42 1.11 E E C+L D C D* D* O.73 1.08 0.82 1.09 0.92 1.09 1.12 1.17 1.24 0.91 1.22 0.00 1.19 1.19 1.23 1.29 0.94 1.29 1.C2 1.23 1.23 D* D* C+L E* C D 1.04 1.02 0.92 1.22 0.96 1.24 [m') 1.10 1.14 1.13 1.17 1.02 1.06 1.33 1.39 1.07 1.35 1.40
'V 1.11 D D C+H D C 1.26 1.18 0.94 1.27 0.98 1.36 1.26 1.02 1.36 1.05 1.42 1.33 1.05 1.42 1.08 C+ D C D
.98 1.27 1.02 1.25 1.07 1.38 1.12 1.34 1.11 1.44 1.15 1.39 C D C 0.98 1.26 0.97 1 1.06 1.36 1.04 l 1.09 1.41 1.07 C D 0.98 1.19 1.08 1.31 1.11 1.37 BL 0.83 0.87 0.89
.-
C-E PLANAR AVERAGE POWER DISTRIBUTION
-
Figure AT BEGINNING OF THIRD CYCLE, UNRODDED S 4.3-28 I Amendment No.3 May 28, 1981
_ __ X.XX BOX AVERAGE (BOX POWER / CORE AVERAGE POWER) X.XX MAX. ENTHALPY RISE FACTOR (MAX.OF AVERAGE POWER OF 4 NEIGHBORING , X.XX PINS / CORE AVERAGE POWER) l rN MAX.ONE PIN FACTOR (MAX. PIN POWER / CORE AVERAGE POWER) l I
! )
V I E E E E 0.59 0.75 0.86 0.86 0.96 1.08 1.18 1.16 1.01 1.14 1.23 1.22 E E E BL E C+ L 0.62 0.86 1.06 0.77 1.23 0.96 1.01 1.13 1.21 0.86 1.3C 1.07 1.06 1.18 1.26 0.89 1.42 1.11 E E C+L D C D' D' O.73 1.08 0.83 1.10 0.92 1.09 1.12 1.17 1.24 0.91 1.23 1.00 1.19 1.18 1.23 1.29 0.94 1.30 1.03 1.23 1.23 L '- D* C+L E* C D ('3 1.04 1.03 0.92 1.22 0.96 1.23 i s ) 1.11 1.14 1.03 1.33 1.07 1.34 U t15 1.17 1.06 1.40 1.11 1.39 D D C+H D C 1.26 1.17 0.94 1.26 0.97 1.36 1.26 1.02 1.35 1.04 1.42 1.33 1.05 1.41 1.07 C+L D C D 0.98 1.25 1.01 1.23 1.08 1.37 1.10 1.32 1.10 1.43 1.14 1.37 C D C
,
0.97 1.24 0.96
'
1.05 1.33 1.03 1.08 1.39 1.06 C D 0.97 1.17 1.07 1.29 1.10 1.35 BL 0.83 0.87
- 0.89 l
'l C-E / , PLANAR AVERAGE POWER DISTRIBUTION AT 50 MWD /T Figure g j / OF THE THIRD CYCLE, UNRODDED
_ l Ameadment No.3 May 28, 1981 L
X.XX BOX AVERAGE (BOX POWER, CORE AVERAGE POWER) X.XX MAX. ENTHALPY RISE FACTOR (MAX. OF AVERAGE POWER OF 4 NElGHBORING PINS / CORE AVERAGE POWER) X.XX n MAX.ONE. PIN FACTOR (MAX. PIN POWER / CORE AVERAGE POWER) x m ,, E E E E 0.64 0.80 0.88 0.88 1.01 1.11 1.17 1.17 1.07 1.17 1.22 1.22 E I F. E BL E C+L 0.66 0.91 1.09 0.81 1.22 0.96 1.06 1.18 1.24 0.88 1.33 1.06 1.11 1.23 1.29 0.90 1.38 1.09 E E C+L D C D* D* 0.77 1.10 0.87 1.12 0.94 1.08 1.10 1.18 1.26 0.94 1.23 1.00 1.16 1.15 1.24 1.31 0.96 1.28 1.03 1.20 1.19 D* D* C+L E* C D 1.06 1.05 0.94 1.20 0.94 1.19 (% 1.12 1.13 1.02 1.29 1.03 1.29 l
\ 1.16 1.17 1.05 1.35 1.05 1.33 F
D D CH D C 1.23 1.15 0.93 1.20 0.94 1.33 1.24 0.99 1.31 1.01 1.38 1.29 1.01 1.37 1.04 C+ L D C D 0.97 1.20 0.98 1.17 1.06 1.31 1.06 1.26 1.09 1.36 1.09 1.30 C D C 0.94 1.18 0.93 1.01 1.26 0.99 1.04 1.31 1.02 C D 0.94 1.13 1.03 1.24 1.06 1.29 BL 0.33
- 0.fs7 0.89
-
C-E f , PLANAR AVERAGE POWER DISTRIBUTION AT 6000 MWD /T Figure
'l OF THE THIRD CYCLE, UNRODDED
/ _
Amendment No.3 May 28, 1981
l l X.X X BOX AVERAGE (BOX POWER, CORE AVERAGE POWER) X.XX MAX. ENTHALPY RISE FACTOR (MAX.OF AVERAGE POWER OF 4 NEIGHBORING X.XX PINS / CORE AVERAGE POWER) MAX.ONE PIN FACTOR (MAX. PIN POWER / CORE AVERAGE POWER)
'O' E E E E 0.68 0.83 0.88 0.89 1.04 1.12 1.15 1.15 1.09 1.17 1:20 1.20 E E E BL E C+ L 0.69 0.92 1.10 0.84 1.19 0.95 1.07 1.19 1.23 0 R9 1.29 1.04 1.12 1.23 1.28 0.91 1.34 1.07 E E C'L D C D* D*
0.78 1.09 0.89 1.12 0.95 1.07 1.08 1.16 1.24 0.94 1.21 1.01 1.14 1.13 1.21 1.28 0.96 1.26 1.03 1.18 1.16 D* D* C+L E' C D
,Ds 1.05 1.04 0.95 1.18 0.94 1.16 V)
I 1.10 1.14 1.11 1.14 1.02 1.04 1.26 1.30 1.01 1.04 1.25 1.30 D D C+H D C 1.20 1.13 0.93 1.18 0.94 1.30 1.21 0.98 1.29 1.01 1.34 1.26 1.01 1.34 1.03 C+L D C D 0.97 1.18 0.97 1.15 1.06 1.28 1.04 1.23 1.08 1.33 1.07 1.28 C D C 0.94 1.16 0.93 1.01 1.24 1.00 1.03 1.29 1.02 l C D 0.95 1.13 1.03 1.23 1.05 1.27 BL 0.86 0.90 . _ 0.92 l
'
'
l C-E f PLANAR AVERAGE POWER DISTRIBUTION AT THE END Figure i
#1EPJ // OF THE THIRD CYCLE, UNRODDED 4.3-31 Amendment No.3 May 28,1981
( 1
.
X.XX BOX AVERAGE (BOX POWER, CORE AVERAGE POWER) X.XX MAX. ENTHALPY RISE FACTOR (MAX.OF AVERAGE POWER OF 4 NElGHBORING PINS / CORE AVERAGE POWER) x'xx
/] MAX.ONE PIN FACTOR (MAX. PIN POWER / CORE AVERAGE POWER)
No.OF No. O F SHIMS PER FUEL TYPE ASSEMBLIES ENRICHMENT ASSEMBLY F 72 3.30 0 F* 8 3.00 0 F F F F 0.57 0.70 0.77 0.76 0.93 0.99 1.05 1.01 0.97 1.03 1.09 1.06 F F F D* F D* O.63 0.89 1.06 0.78 1.10 0.79 1.06 1.17 1.23 0.89 1.22 0.87 1.10 1.22 1.28 0.93 1.27 0.90 F F D E D E E' O.74 1.12 0.98 1.08 0.92 0.97 0.90 1.20 1.29 1.07 1.16 1.04 1.05 0.95 1.25 1.35 1.11 1.21 1.08 1.08 0.98 E* E D F* D* E i p- 1.05 1.10 1.10 1.17 1.04 1.17 1.22 1.36 0.88 1.01 1.04 1.16 1.13 1.22 1.21 1.41 1.04 1.20 E E O E D 1.21 1.13 1.10 1.2' 1.06 1.31 1.22 1.21 1.35 1.20 1.37 1.26 1.25 1.41 1.24 I D* E D E 0.94 1.24 1.15 1.23 1.02 1.37 1.25 1.34 1.05 1.43 1.29 1.39 D E D l 1.13 1.25 1.12 l 1.24 1.35 1.21 ! 1.28 1.40 1.25 D E 1.12 1.18 1.19 1.25 1.22 1.29 C 0.86 0.92 C-E PLANAR AVERAGE POWER DISTRIBUTION AT BEGINNING Figure OF THE FOURTH CYCLE, UNRODDED E 4.3-32 l Amendmcat No.3 May 28, 1981 l
-- - ._
1 X.X X BOX AVERAGE (BOY. POWER, CORE AVERAGE POWER) X.XX MAX. ENTHALPY RISE FACTOR (MAX. OF AVERAGE POWER OF 4 NElGHBORING PINS / CORE AVERAGE POWER) X.XX MAX.ONE PIN FAr: TOR (MAX, PIN POWER / CORE AVERAGE POWFR) frS ( ) NJ' F F F F 0.58 0.72 0.79 0.78 0.94 1.u1 1.08 1.03 0.99 1.06 1.11 1.08 F F F D* F D* 0.64 0.89 1.07 0.80 1.12 0.81 1.06 1.17 1.24 0.91 1.24 0.89 1.11 1.22 1.29 0.94 1.30 0.92 F F D E D E E' O.74 1.11 0.98 1.08 0.93 0.98 0.92 1.19 1.29 1.07 1.16 1.05 1.06 0.97 1.25 1.35 1.12 1.21 1.08 1.09 1.00 E' E D F' D* E 1.05 1.10 1.04 1.22 0.89 1.05 /7 1.10 1.17 1.16 1.36 1.01 1.16 ! ) V 1.14 1.21 1.21 i 1.42 1.04 1.21 E E D E D 1.20 1.12 1.09 1.20 1.06 1.31 1.22 1.20 1.33 1.18 1.36 1.26 1.24 1.39 1.23 D* E D E 0.93 1.22 1.13 1.21 1.00 1.35 1.22 1.31 1.03 1.41 1.26 1.36 D E D 1.11 1.22 1.10 1.21 1.32 1.18 1.25 1.37 1.22 D E 1.09 1.15 1.16 1.22 1.19 1.27 C i
' O.85 0.90 J
0.93
- '
PLANAR AVERAGE POWER DISTRIBUTION AT 50 MWD /T Figure C-E f ggg OF THE FOURTH CYCLE,UNRODDED _g 7f i _ Amendment No.3 May 28, 1981 _
X.XX BOX AVERAGE (BOX POWERfCPRE AVERAGE POWER) X.XX MAX. ENTHALPY RISE FACT ( :(MAX.OF AVERAGE POWER OF 4 NElGHBORING PINS / CORE AVERAGE POWEP) X'XX q MAX. ONE. PIN FACTOR (MAX. PIN POWER / CORE AVERAGE POWER) F F F F 0.64 0.79 0.86 0.85 1.01 1.09 1.14 1.11 1.06 1.14 1.18 1.16 F F F D* F D* O.67 0.92 1.10 0.85 1.18 0.87 1.08 1.20 1.25 0.95 1.28 0.94 1.13 1.24 1.30 0.98 1.34 0.96 F F D E D E E' O.76 1.11 0.99 1.09 0.96 1.03 0.97 1,18 1.28 1.07 1.17 1.03 1.10 1.02 1.24 1.33 1.10 1.21 1.06 1.13 1.05 E' E D F* D* E 1.04 1.09 1.03 1.21 0.90 1.06 D\ 1.09 1.16 1.12 1.31 0.98 1.16 (d 1.13 1.20 1.15 1.37 1.00 1.20 E E D E D 1.17 1,09 1.05 1.15 1.02 1.26 1.18 1.14 1.26 1.11 1.31 1.22 1.17 1.31 1.14
'
D* E D E 0.91 1.15 1.06 1.13 0.97 1.25 1.13 1.21 1.00 1.30 1.16 1.26 D E D l 1.03 1.13 1.02 ! 1.11 1.21 1.09 l 1.15 1.26 1.12 D E 1.02 1.08 1.07 1.14
1.10 1.18 ' C 0.82 ' 0.86 O.88
- - =
V C-E j PLANAR AVERAGE POWER DISTRIBUTION AT 6000 MWD /T Figure OF THE FOURTH CYCLE,UNRODDED 7 Amendment No.3 May 28, 1981
X.X X BOX AVERAGE (BOX POWER. CORE AVERAGE POWER) X.XX MAX. ENTHALPY RISE FACTOR (MAX. OF AVERAGE POWER OF 4 NElGHBORING X.XX PINS / CORE AVERAGE POWER)
^
MAX. ONE-PIN FACTOR (MAX, PIN POWER / CORE AVER AGE POWER)
, -
F I F F F 0.68 0.83 0.89 0.88 1.04 1.11 1.15 1.14 1.09 1.16 1.19 1.18 F F F D' F D* 0.69 0.93 1.10 0.88 1.18 0.90 1.07 1.19 1.24 0.97 1.27 0.96 1.12 1.23 1.28 0.99 1.32 0.98 F F D E D E E' O.78 1.09 0.98 1.03 0.97 1.04 1.00 1.15 1.24 1.06 1.16 1.03 1.10 1.04 1.20 1.29 1.09 1.20 1.05 1.13 1.07 E' E D F* b* E g' s 1.02 1.07 1.01 1.18 0.92 1.07 i, l 1.07 1.14 1.09 1.27 0.97 1.16
\~d 1.10 1.17 1.12 1.31 0.99 1.20 E E D E D 1.14 1.07 1.03 1.13 1.01 1.22 1.15 1.11 1.22 1.08 1.26 1.19 1.14 1.27 1.11 D* E D E 0.91 1.12 1.04 1.11 0.97 1.22 1.10 1.18 0.99 1.26 1.13 1.22 D E D 1.02 1.11 1.01 1.08 1.18 1.07 1.11 1.22 1.09 D E 1.01 1.07 1.06 1.14 1.09 1.17 C
0.84 0.88 0.90 O C-E f PLANAR AVEf AGE POWER DISTRIBUTION p;gg 7e
'l AT END OF THE FOURTH CYCLE, UNRODDED
_ Amendment No.3 May 28, 1981 __
._ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ x [~N ( h v
)
k-100% TO 35% TO 100% POWER -TWO HOUR RAMPS - 8 HOURS AT 35% POWER -WITH PART LENGTH RODS b POWER cCNTROL ON RODS, XENON CONTROL ON BORON - EQUILIBRIUM DAILY MANEUVERING CYCLE o 8 m 100% PO1VER, HOUR 0 100% POWER, HOUR 6 83.75% POWER, HOUR 6.5 67.5% POWER, HOUR 7 2.00
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S 1 S LEGEND: Amendment No.3 May 28, 1981 S= SPARE 4-FINGERED CEA
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5 - LEAD REGULATING BANK 4 - SECOND REGULATING BANK 3 -THIRD REGULATING BANK 2 - FOURTH RtEGULATING BANK 1 - LAST REGULATING BANK B - SHUTDOWN BANK B l A --SHUT 00WN GANK A ) P2 - PLR GROUP 2 P3 -PLR GROUP 1 S - SPACE CEA LOCATIONS 1 2 3 4 5 6 7 S 3 S 8 9 10 11 12 13 14 15 16 17 18 A 1 1 A 19 20 21 22 23 24 25 26 27 28 29 30 31 4 2 P 2 3 P 2 4 2 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 B B B B B B 47 48 49 50 51 52 53 lIi 55 56 57 58 59 60 61 2 P 3 5 P3 2 62 63 64 G5 66 67 68 69 70 71 72 73 74 75 76 77 78 A B 4 A A 4 B A 79 80 81 82 83 84 85 86 87 ab 89 90 91 92 93 94 95 S P P 8 2 2 96 97 98 99 100 101 102 103 104 105 106 107 108 109 110 111 112 1 B A 3 3 A B 1 113 114 115 116 117 118 119 120 121 122 123 124 125 126 127 128 129 3 3 5 P 5 3 3 3 130 131 132 133 134 135 136 137 138 139 140 141 142 143 144 145 146 1 B A 3 3 A B 1 147 148 149 150 151 152 153 154 155 156 157 158 159 160 161 162 163 S P P 8 2 2 164 165 166 167 168 169 170 171 172 173 174 175 176 177 178 179 180 A B 4 A A 4 B A 181 182 183 184 185 186 187 188 189 190 191 192 193 194 195 2 P3 5 P 2 3 196 197 198 199 200 ?O1 202 203 204 205 206 207 208 209 210 B B B B B B 211 212 213 214 215 215 217 218 219 220 221 222 223 4 2 P 3 P 2 4 2 2 224 225 226 227 228 229 230 231 232 233 234 A 1 1 A 235 236 237 238 239 240 241 S 3 S Amendment No.3 May 28, 1981 ' gggy CEA BANK IDENTIFICATION 4 3-51
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-0.08 -0 04 0 0.04 0.08 0.12 0.16 EXTERNAL SHAPE INDEX, I e
_ Amendment No.3 l May 28, 1981 C-E I i "8"* COMPARISON OF MEASURED AND CALCULATED
/ SHAPE ANNEALING CORRECTION FOR PALISADES 4.3-61
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l ! 555 565 575 l INLET TEMPERATURE, UF i { l Amendment No.3 ( - May 28. 1981 1-
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- TYPICAL TEMPERATURE DEFECT VS.
l S REACTOR INLET TEMPERATURE 4.3-62 1 1
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l C-E ROCS /DIT REACTIVITY FROM CORE FOLLOW Figure i g [ CALCULATIONS,14 x 14 PIANTS, RELOAD CYCLES 4.3-64
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0 't i I I I I O " 1.00 1.05 1.10 1.20 1.25 RADIAL PIN POWER Amendment No. 3 May 28, 1981 _
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C-E INTEGRAL RADIAL PIN POWER DISTRIBUTION Rgure USED IN AUGMENTATION FACTOR DETERMINATION 4.3-68 _ _ - _ _ _ . - _ _ _ . - ___. .- .
b) o I
,
i % INCREASE IN POWER (1) (3) (5) GAPPED
---
ROD 4.a?7 1.678 0.788 --- (1) (2) (4) (5) 4.477 3.007 1.447 0.788 (3) (A) (5) 1.678 1.447 0.788 (5) (5) - (D 0.788 0.728
\-)
'
I I LEGEND: I
.
l l (n) ROD GROUPING INDEX FOR THE LOCATION 0F UNGAPPED ROD RELATIVE TO THE GAPPED ROD x.xxx PEAKING IN UNGAPPED ROD DUE TO A SINGLE GAP AT INDICATED LOCATION Amendment No. 3 May 28, 1981 8 C-E LIMITING SINGLE GAP POWER PEAKING AND Figure ASSOCIATED ROD LOCATION RELATIVE T0 S THE GAP LOCATION a 3-69
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0 10 20 30 40 50 60 70 80 90 10'J DISTANCE FROM BOTTOM 0F CORE, % 2F wg N5
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