ML20056B034
| ML20056B034 | |
| Person / Time | |
|---|---|
| Site: | Susquehanna  |
| Issue date: | 08/01/1990 |
| From: | Kulick J PENNSYLVANIA POWER & LIGHT CO. |
| To: | |
| Shared Package | |
| ML17157A275 | List: |
| References | |
| PL-NF-90-001, PL-NF-90-1, NUDOCS 9008130272 | |
| Download: ML20056B034 (259) | |
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e i APPLICATION OF REACTOR ANALYSIS METHODS FOR BWR DESIGN AND ANALYSIS PL-NF-90-001 -:- August 1990 i Principal Enaineers Cleon E. Dodge Andrew Dyszel John J. Geosits I Kenneth C. Knoll Chester R. Lehmann i Contributina Enaineers John H. Emmett Kimberley I.R. Knippel Paul H. Lee Salvatore A. Somma Anthony J. Roscioli William J. Weadon I I I Approved- TI Ulkf/( h M. Ku'ick hb '
- kH L. ) , M 40 O Date Su visor-Nuclear Fuels Engineering I
l OwlN/a N ager-Nuclear Fuels S. Stefanko Systems Engineering A/, wo ate' I I
= _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . - . _ _ . . . . . . . I LEGAL NOTICE I I This topical report represents the efforts of Pennsylvania Power & Light Company (PP&L) and reflects the technical capabilities of its nuclear fuel analysis personnel. The information contained herein is completely true I and accurate to the best of the Company's knowledge. The sole intended purpose of this report and the information contained herein is to provide a description of PP&L's use of its steady state core physics and transient analysis methods for licensing applications. Any use of this report or the information contained herein by anyone other than PP&L or the U.S. Nuclear Regulatory Commission is unauthcrized. With regard to any unauthorized use, Pennsylvania Power & Light Company and its officers, directors, agents, and employees make no warranty, either express or implied, as to the accuracy, completeness, or usefulness of this report or I the information, and assume no liability with respect to its use. I I i 1 I I I I I I
y . I I ABSTRACT This topical report describes Pennsylvania Power & Light Company's
. -(PP&L's) use of its core physics and transient analysis methods for licensing applications.
.PP&L's steady state core physics and transient analysis methods are based mainly on the computer codes provided by the Electric Power Research Institute (EPRI). The codes used in the steady state core physics methods include: .the MICBVRN gadolinia fuel pin depletion code; the CPM-2 assembly lattice depletion code; the NORGE-B2 cross section data link I code; and the SIMULATE-E three-dimensional core simulation code. The EPRI
- codes used by_ PP&L for transient analyses include SIMTRAN-E, ESCORE, and g RETRAN02 MOD 004. The methodology for thermal margin analysis uses the E XN 3 Critical Power _ correlation developed by the Advanced Nuclear Fuels Corporation-(ANF).
Previously published PP&L reports describe PP&L's core physics and transient analysis methods =(References 1 and 2). These reports provide benchmarking analyses to demonstrate the validity of these methods and PP&L's ability to use them appropriately. The PP&L models and methods
& I described in those reports comprise a mainly best estimate methodology.
This report describes the adaptation of the steady state core physics and transient analysis methods to perform conservative analyses in support of licensing applications. The analysis methodologies described in this report are r ~anned to be used to establish MCPR operating limits, demonstrate compliance with the ASME overpressure criteria, and provide core physics input to the fuel vendor for use in their accident analyses
, (e.g.,LOCA). Sensitivity studies and evaluations based on first principles are presented and used to identify conservative modelling and input assumptions.
I events which PP&L plans to analyze for a typical reload core: rod Sample analyses are presented for those design basis I I _ _ _ _ _ _ _ - . _ _ _ - _ _ _ - . - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - ^ ' - - - - - ' - " - - " - - - - - - ^ ' - - - - " - - ^ - - -----^
I withdrawal error, fuel loading error, loss of feedwater heating, generator I g-load rejection, feedwater controller failure, recirculation flow controller failure, and the closure of all main steam isolation valves. Statistical Combination of Uncertainty (SCU) methodologies for the analysis of pressurization transients and the rod withdrawal error to estr.blish MCPR operating limits are described. The analysis methodologies described in this report will result in g conservative MCPR operating limits. The conservative ASME overpressure 5 analysis methodology described herein will be used to demonstrate that sufficient margin exists in the Susquehanna SES design and operation to comply with the ASME overpressure criteria. The methodologies presented herein will also provide appropriately conservative data to enable the fuel vendor to perform certain of their analyses. In addition, the use of CPM-2 to generate neutronics input to the POWERPLEX plant monitoring g system will result in accurate monitoring of thermal limits.
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r I ACKNOWLEDGEMENTS I The authors gratefully acknowledge the expert stenographic.' work provided I by Ms. Evelyn Lugo, Ms. Lisa Walsh, and Ms. Robin Abbott, the clerical support- provided by Mr. Francis E. Gria, and the assistance of Mr.- Paul Wasson's Systems and Computer Services group whose efforts have I contributed to the quality and timely cy.pletion of this topical report. The authors also acknowledge the' efforts of Mr.- Rocco R. Sgarro for his licensing reviews and coordination with the NRC. The assistance of the Advanced Nuclear Fuels Corporation is gratefully _ acknowledged. The consult Mg assistance provided by E. D. Kendrick of Nuclear Engineering Techno1cgy Corporation (NETCORP) is greatly f' appreciated. The authors also wish to gratefully acknowledge the continuing support of L the reload methods development effort by J. S. Stefanko, J. M. Kulick,. W..J. Rhoades, and G. D. Miller. I , l . i 4 I Y u I ; 'I l 'I I .
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APPLICATION 0F REACTOR ANALYSIS METHODS FOR WR DESIGN AND ANALYSIS l TABLE OF CONTENTS
-Section, gggg
'l' . 0 INTRODUCTION 1 2.0- STEADY STATE CORE PHYSICS METHODS APPLICATIONS 8 2.1- Rod Withdrawal Error 9
'2.1.1 Event Description i 2.1.2 Sensitivity Studies 2.1.3 Licensing Analysis Method 10 9
15 2.1.4 . Sample Licensing Analysis 18
=- 2 . 2 = Fuel' Loading Error 36 2.2.1 Event Description '36 1 2.2.2 Sensitivity Studies 2.2.3 ' Licensing Analysis Method' 38 43 2.2.4: Sample Licensing Analysis 46 1 '2.3 ' Loss of'Feedwater Heating 50 2.3.1 Event Description 50 2.3;2 Sensitivity Studies. . . . . 51
-2.3.3 Licensing Analysis Method 55 2.3.'4 Sample Licensing Analysis- 55 2.4 Shutdown Margin Determination' 59 2.4.1 Event Description 1> 2.4.2 Sensitivity Studies 59 60 2.4.3 Licensing Analysis Method 60 2.4.4 Sample Licensing Analysis 62
. _ _ _ _ _ . _ _ . _ _ . _ _ . _ _ _ . _ _ . _ _ _ . _ m-- - . _ ----- ---..--- -------- - - - "---------- - - - - - - - - - - -
Section PJlLgt I 8 2.5 Standby Liquid Control System Capability 64 2.5.1 Event Description 64 2.5.2 Sensitivity Studies 64 2.5.3 Licensing Analysis Method 66 2.5.4 Sample Licensing Analysis 68 2.6 RETRAN Transient Analysis Inputs 69 2.7 Loss of Coolant Accident inputs 71 1 2.8 Control Rod Drop Accident inputs 75 2.9 MCPR Safety Limit inputs 79 l 2.9.1 Core Physics Inputs 79 2.9.2 Uncertainties 81 2.10 Core Monitoring System Inputs 83 j 2.11 Fuel Storage Criticality Compliance 97 l 3.0 TRANSIENT ANALYSIS METHODS APPLICATIONS 98 3.1 Generator Load Rejection Without Bypass 104 j 3.1.1 Event Description 104 3.1.2 Sensitivity Studies 105 l 3.1.3 Licensing Analysis Method 107 3.1.4 Sample Licensing Analysis 110 3.2 Feedwater Controller Failure 125 3.2.1 Event Description 125 g 3.2.2 Sensitivity Studies 126 3, 3.2.3 Licensing Analysis Method 129 l 3.2.4 Sample Licensing Analysis 130 3.3 Recirculation Flow Controller Failure 143 l 3.3.1 Event Description 143 3.3.2 Sensitivity Studies 144 3.3.3 Licensing Analysis Method 147 l 3.3.4 Sample Licensing Analysis 148 l ll 1 I!
l Section Ra.gt 3.4 Non-Limiting' Events 159 3.4.1 ' Turbine Trip 160 3.4.2 .MSIV Closure (Position Scram) 160 i E 3.4.3 Loss of Condenser Vacuum- 160 ;
'R 3.4.4 Recirculation Pump Trip 161.. 1 3.4.5 Inadvertent HPCI Startup 161-7 3.5: ASME Overpressure Analysis -163 i
3.5.1 Criteria 163 : y 3.5.2 Event Description 164 3.5.3 :RETRAN Modelling 165 3.5=4 Sensitivity Studies
. -168 I 3.5.5 Licensing Analysis Method 3.5.6 Sample Licensing Analysis-1721 173
'4.0 TECHNICAL SPECIFICATIONS 193 4.1 MCPR Limits 193 a 4.1.1 Basic Approach :194-4.1.2 Scram Speed Dependent MCPR Operating Limits 195 4.2 Other Limits 197 5.0 .
SUMMARY
AND CONCLUSIONS 198-6.0, REFERENCES 199 it ' APPENDICES AL GAP CONDUCTANCE METHODS '203 i g.f* A.1 Introduction 204- 'I j' A.2 Definition of Conservative for Licensing Calculations 204 a A.3 Use of Rod Average Gap Conductance 205 !
~A.4 ESCORE Inputs 206
.. A.5 Power History. 208 L A.6 -Gap Conductance Methodology Summary 211 a! lA.7 Sample Results 213
.3l A.8 Summary / Conclusions 213 1 L >
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B.- -STATISTICAL COMB'aNATION OF UNCERTAINTIES- 221 .
.B.1 Introduction ~ 222 B.2' Pressurization Events 224 E
- B.3 Rod Withdrawal Error' 234 B.4 Interface with Non Statistical: Analyses 241 T g
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Ei LI . LIST OF TABLES Table - Number Iltle En.gt I
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1-1 General Design and Operating Features of the 4 Susquehanna Steam Electric Station Reactors 2.1 Rod Withdrawal Error Base Case Input Assumptions 21 2.1-2 Rod Block Monitor Response for a Sample Unit 2 22 Cycle 2 Rod Withdrawal Error 2.2-1 Maximum Changes in Local Peaking Factor and 48 S-Factors for the Rotated Bundle Analysis 2.2-2 Rotated Bundle Sample Reload Analysis: Maximum 49
' Calculated Changes -in Local Peaking Factor and S-Factors 2.3-1 Loss of Feedwater Heating Sample Analysis Compliance 56 B with Generic Regression Analysis 2.9-1 Uncertainties Used to Generate MCPR Safety Limit 82.
3.1-1 Gener'ator Load Rejection Without Bypass 113 Base Case :aput Assumptions-I-
, 3.1-2 Generator Load Rejection Without Bypass 114 Results of Sensitivity Analysis-3.1-3. Cases Analyzed for GLRWOB Response Surface 115 3.1-4 Coefficients for GLRWOB Response Surface. 116 3.1-5 Steam Line Parameters: Contribution to Code 117
- Uncertainty for GLRWOB i 3.2-1 Feedwater Controller Failure Base Case input Assumptions 132 I 3.2-2 Feedwater Controller Failure Results of Sensitivity Analyses 133~
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LIST OF TABLES (continued)- I Number Iltle f.nat 3.2 Steam Line Parameters: Contribution to Code 134-Uncertainty for FWCF 3.3 1 Single and Two Loop RFCF Events: Initial Power / 150 - Flow Conditions 3.3-2 Change in RCPR As a Function of Master Controller 151 Run-up-Rate for the 65/38 Power / Flow RFCF 3.4-1 Potentially Limiting ivents in Establishing 162 MCPR Operating Limits 3.5-1 ASME Overpressure Analysis Criteria- 174 I 3.5-2 Safety-Relief Valve Flowrates 175 3.5-3 MSIV Closure /ASME Overpressure Analysis 176- a Base Case Input Assumptions g 3.5-4 MSIV Closure /ASME Overpressure Analysis 177 Results of Sensitivity Study 3.5-5 MSIV Closure Sample Licensing An& lysis 178 Peak Calculated- Pressures A.4-1 Selection of Conservative Input Values for 214 ESCORE Licensing Basis Gap Conductance Calculation A.4-2 ESCORE Input Parameters with Negligible Effect on 217. E Gap Conductance (For Use in Both System and Hot B Bundle ~ Analyses)
'A.4-3 Noding and Time Step Selection.for ESCORE Gap 218 Conductance Calculations-(System and Hot Bundle Model)
A.7-1 System Model Gap Conductance Results for Susquehanna 219 SES Unit 2 Cycle 2 I I B
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- A.7-2 Hot: Bundle Model: Gap-Conductance Results for 220 Susquehanna SES Unit 2 Cycle 2 B.2 Uncertainties Used to Generate MCPR Safety. Limit 230
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,~ -l LIST OF FIGURES Figure Humber 111.13 Eagg l
1-1 ~ Susquehanna Steam Electric Station Typical Core 5 Power vs. Core flow I-2 PP&L Steady State Core-Physics Methods Computer 6 f Code. Flowchart 1-3 PP&L Reactor Transient Analysis Methods _ Computer 7 Code Flowchart.- B l 2.1-1 Control" Rod-Locations for Error Rod in Rod 30 Withdrawal Error Analysis 2.1-2 Minimu'm Channel 'A'-RBM Response as'a Function 31 of Rod Withdrawal for. 0,1,2,3, and 4 LPRM Failures- [ 2.1-3 Channel 'A' RBM Response as a Function of Rod 32 I Withdrawal for Four Different failure Combinations with Four LPRM Failures 2.1-4 RCPR Response as a Function of Rod Withdrawal for a 33-Rod Withdrawal Event 2.1-5 LHGR Response as a Function of Rod Withdrawal for a 34 l Rod Withdrawal Event l 2.'l-6 Core Thermal Power Response as a function of Rod- 35 i Withdrawal for a Rod Withdrawal Event 2.3-1 Loss of Feedwater-Heating Event.- Change-in Minimum 57 Critical -Power Ratio - 2.3-2 Loss of Feedwater Heating Event - 95/95 Tolerance 58 Limits on Regression Analysis 2,4-1 Exposure Effects on Core Shutdown Margin ~63 1 l l
I: I LIST OF FIGURES (continued) Il
. Figure I
Number 11.tle P_ASA a l g 2.10 1 Susquehanna SES Unit 1 Cycle 1 Relative Nodal RMS 85 q of.TIP Response Comparisons
, '2.10-2 Susquehanna SES Unit 1 Cycle 2 Relative Nodal RMS 86 i of TIP Response Comparisons 2.10 3 Susquehanna SES Unit 1 Cycle 3 Relative Nodal RMS 87 1 of TIP Response Comparisons i 2.10-4 Susquehanna SES Unit 2 Cycle 1 Relative Nodal RMS 88 of TIP Response Comparisons >
2.10-S Susquehanna SES Unit 2 Cych 2 Relative Nodal RMS 89 of TIP Response Comparisons 2.10 6 Susquehanna SES Unit 2 Cycle 3 Relative Nodal RMS 90 of TIP Response Comparisons 2.10-7 Susquehanna SES Unit 1 Cycle 1 Hot Calculated K-effective 91 l; 2.10 8 Susquehanna SES Unit 1 Cycle 2 Hot Calculated K-effective 92 2.10-9 Susquehanna SES Unit 1 Cycle 3 Hot Calculated K-effective 93 2.10-10 Susquehanna SES Unit 2 Cycle 1 Hot Calculated K-effective 94 2.10 11 Susquehanna SES Unit 2 Cycle 2 Hot Calculated K-effective 95 2.10-12 Susquehanna SES Unit 2 Cycle 3 Hot Calculated K-effective 96 3.0 1 RETRAN/CPRITER Code Relationships 101 3.0 2 Susquehanna SES RETRAN Model (Vestel) 102 3.0 3 Susquehanna SES RETRAN Model (Steamline and Bypass) 103 3.1 1 GLRWOB: Core Power 118 3.1-2 GLRWOB: Core Flow 119 3.1-3 GLRWOB: Dome Pressure 120 I L a
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I LIST OF FIGURES (continued) i { l Figure thahtt I1111 hst I i 3.1-4 GLRWOB: Vessel Steam Flow 121 3.1-5 GLRWOB: Narrow Range Level 122
3.1 GLRWOB
Feedwater Flow 123 1 I 3.1-7 GLRWOB: Average Heat Flux 124 3.2-1 FWCF: Core Power 135 3.2-2 FWCF:. Core Flow 136 3.2-3 FWCF: Dome Pressure 137 3.2-4 FWCF: Vessel Steam Flow 138 3.2-5 FWCF: Narrow Range Level 139 ,
.3.2a6 FWCF: 441 water Flow 140 3.2i7 FWCF: Average Heat Flux- 141 .
-3.2-8 FWCF:. Core Inlet Enthalphy 142 I 3.3-1 RFCF: Core Power 152 :
3.3-2 RFCF: Core Flow 153 , 3.3-3 RFCF: Dome Pressure 154 3.3-4 RFCF: Vessel Steam Flow 155 3.3-5 RFCF: Narrow Range Level 156 1.3-6 RFCF: Feedwater Flow 157 l 3.3-7 RFCF: Average Heat Flux 158 3.5-1 Assumed-SRV Flow Characteristics 179 3.5-2 Conservative SRV Flow Model 180 ~ I l
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I: LIST OF FIGURES (continued) Figure I Number Title fant gl g 3.53 MSIV Loss Coefficient 181 3.5-4 Overpressure Margin (1250 ?sig Design Pressure) 182 3.5 5 Overpressure Margin (1500 Psig Design Pressure) 183 3.5 6 MSIVC: Core Power 184 3.5 7 MSIVC: Core Flow 185 f 3.5 8 MSIVC:- Dome Pressure 186 3.5 9 MSIVC: Vessel Steam flow 187 l i 3.5 10 MSIVC: Narrow Range Level 188 3.5 11 MSIVC: Feedwater Flow 189 3.5 12 3.5 13 MSIVC: Average Heat Flux MSIVC: Pressure Margin /1250 Psig Category 190
.191 l l 3.5 14 MSIVC: Pressure Margin /1500 Psig Category 192 'l B.2 1 Statistical RCPP. Analysis 231 8.2-2 Safety Limit MCPR Calculation 232 B.2 3 Statistical Operating Limit Calculation 233 i B.3 1 Flow Path for Combination of Uncertainties in 238 Rod Withdrawal Error Analysis - Part 1 <
B.3 2 Flow Path for Combination of Uncertainties in 239 i Rod Withdrawal Error Analysis - Part 2 N B.3 3 Flow Path for Combination of Uncertainties in Rod Withdrawal Error Analysis - Part 3 240 L L l l I
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1.0 INTRODUCTION
Pennsylvania Power & Light Company (PP&L) operates the two unit Susquehanna Steam Electric Station (SES) near Berwick, Pennsylvania. Both ; Susquehanna SES units are General Electric Boiling Water Reactor 4 (BWR-4) product line reactor systems; each has a rated thermal power output of 3293 megawatts. The general core design and operating features are given I in Table 1-1. Figure 1-1 provides a typical power / flow relation for the Susquehanna SES units. l Reference 1 describes the steady state core physics methods used by PP&L - for BWR core analysis and provides qualification of the analytical
-methodologies which will be used to perform safety related analyses in ,
g support of licensing actions. PP&L's steady state core physics methods are based en the Electric Power Research Institute (EPRI) code package (Reference 3), as depicted in the flowchart contsined in Figure 12. The lgW main computer codes are the CPM-2/PP&L fuel assembly lattice physics, ! l depletion code (Reference 4), hereinafter referred to as CPM 2, and the SIMULATE E/PP&L three dimensional core sinslation code (Reference 5), hereinafter referred to as SIMULATE-E. The tilCBURN/PP&L code (Reference 6),hereinafterreferredtoasMICBURN,providesadetailed representation of the depletion of a single gadolinia (Gd,03 ) bearing fuel pin; the NORGE-B2/PP&L code (Reference 7), hereinafter referred to as t NORGE-B2, provides a nuclear cross section data link from CPM-2 into SIMULATE E as well as into the POWERPLEX core monitoring system I (Reference 8). TIPPLOT provides plotting and statistical analysis capabilities. The RODDK-E/PP&L code (Reference 9), hereinafter referred ' to as R00DK-E,1s used to select the strongest worth control rod locations for shutdown margin analyses and to provide estimatos of core shutdown ! margin. 'I I
I Reference 2 describes PP&L's transient analysis methods for the analysis I of a GE BWR-4 reactor. Qualification analyses are presented to demonstrate the accuracy of the codes, models, and methods and their adequacy for reactor transient analysis. PP&L's transient analysis , methods are also- based on the Electric Power Research Institute (EPRI) code package as depicted in the flow chart contained in Figure 1-3. The g, RETRAN 02 MOD 004/PP&L coupled ,;eutronic-thermal hydraulic code W ; (Reference 10), hereinafter referred to as RETRAN, is used to model the Nuclear Steam Supply System (NSSS) as a whole, and to model a single fuel ! assembly for thermal margin evaluations. The SIMTRAN E/PP&L code (Reference 11), hereinafter referred to as SIMTRAN E, collapses the three- f, dimensional neutronics data generated by the steady state core physics ' codes to the one-dimensional neutronics input required by RETRAN. Thermal ! margin' evaluations are performed using the CPRITER code, which is an ' automated version of the PP&L developed DELTACPR code described in Reference 2. The CPRITER and DELTACPR codes use the Advanced Nuclear - o Fuels Corporation (ANF) XN-3 critical power correlation (Refereace 12). The ESCORE/PP&L code (Reference 13), hereinafter -eferred to as CSCORE, is used to calculate gap conductances. This report describes the application of the steady state core physics and I : transient analysis methods to licensing analyses, including the g assumptions, methods, and application of uncertainties used to calculate conservative results. A portion of the analyses required to support licensing applications will be performed by the fuel vendor, including ! LOCA, MCPR safety limit, control rod drop, fuel handling accident, and g stability analyses. Several of these analyses will utilize input 5', generated using PP&L methods. Therefore, the data generation using PP&L methods to provide input to the fuel vendor methods is also described. The steady state core physics methods are used to perform explicit g analyses which' are used directly in determining the Minimum Critical Power Ratio (MCPR) operating limit (e.g., rod withdrawal error), to provide I I ,
,. . . . _ . ~ .-. .. - - .- - . - . - - . . . , .
I 1 I I assurance that the fuel rod Linear Heat Generation Rate (LHGR) does not ] exceed the Protection Against Fuel Failure (PAFF) limit (Reference 29), l and to provide inputs to the fuel vendor for use in their analyses. The RETRAN based transient analyses are used both to perform explicit analyses , which are used in determining the MCPR operating limit (e.g., generator load rejection)-and to perform reactor vessel ASME overpressure analyses. , Analyses described in this report represent work performed by PP&L. The computer codes, models, and calculations supporting this work are , documented, revi.wed, and controlled by formal procedures encompassed [ within PP&L's nuclear quality assurance pr) gram. I I . I LI i t I I I < I ul :
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fl -J! (! 1 i I TABLE 1-1 GENERAL DESIGN AND OPERATING FEATURES OF 1 THE SUSOUEHANNA STEAM ELECTRIC STATION REACTOR 5 . I; Reactor Type / Configuration: BWR-4/2 Loop Jet Pump Recirculation Ii System , l Rated Core /ower: J293 MW Thermal 31 Rated Core Flow:- 100x10' lbm/hr Reactor Pressure at Rated 1020 psia I, Conditions: 1
, Number of Fuel. Assemblies: 764 I' Number of Cont'rol Rods: 185 I;,
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l SUSQUEHANNA STEAM ELECTRIC STATION FIGURE 1-1 i I TYPICAL CORE POWER VS CORE FLOW I, 4 12 0 i . : : . : . . l l l l . : .
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i Ii FIGURE 1-2 I PP&L STEADY STATE CORE PHYSICS METHODS I COMPUTER CODE FLOWCHART g i MICBURN Gd Depletion
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v . il , l CPM-2 Laitice enysics g v g I V NORGE-B2 m'; Data Link -. 5 POWERPLEX-l' , Core Monitoring System - SIMULATE-E I y 3 D Simulation v
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I' 8 ! FIGURE 1-3 l PP&L REACTOR: TRANSIENT ANALYSIS METHODS COMPUTER CODE FLOWCHART I I I SIMULATE-E 3 D Simulation E SIMTRAN-E , 5 3 D to 1 D Link t I %/ I MODIFICATION OF CROSS SECTION DEPENDENCIES N/ 5 RETRAN ~ ESCORE I NSSS Model Gap Conductanc.e S/ RETRAN < Hot Bundle Model h/ ' LI CPRITER
. Thermal Margin Analysis 3
j' a 3 It 2.0 STEADY STATE CORE PHYSICS METHODS APPLICATIONS l The application of PP&L's steady state core physics methods described in Reference I to various licensing analyses is discussed in this section. g< W < Also discussed is the preparation of certain of the POWERPLEX core monitoring system inputs using PP&L methodology. The steady state core ' L physics methods are used for four major purposes:
- 1) To perform explicit event analyses to demonstrate compliance with Technical Specification requirements and to provide Technical Specification input such as MCPR operating limits (Sections 2.1 to g
2.5,2.11). For several of these events, a parameter of interest is E defined as the change in Critical Power Ratio (ACPR) divided by the W initial value of CPR, and is referred to as RCPR. :
- 2) To generate 1-D kinetics data for use by the RETRAN code for analyses of reactor system transients (Section 2.6).
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- 3) To generate physics data which is utilized as input to certain of the fuel vendor analyses such as LOCA, control rod drop accident, g
and MCPR safety limit analyses (Sections 2.7 to 2.9).
- 4) To generate lattice physics data for the POWERPLEX core monitoring ,
system (Section 2.10). o I I I 8-I L I I
I ! LI ! 2.1 Rod Withdrnal Error , 2.1.1 Event Descriotion The Rod Withdrawal Error (RWE) event is expected to be one of the limiting i . transient events in establishing the MCPR operating limit. The event is initiated by an operator erroneously selecting and continuously ' withdrawing a high worth control iad at its maximum withdrawal rate. This P control rod withdrawal results in the addition of positive reactivity, and therefore both the local power in the vicinity of the erroneously ' withdrawn control rod and the core average power increase. The local and g core average powers continue to increase as the control rod is withdrawn. The power increases are terminated when the Rod Block Monitor (RBM) I response reaches its flow biased trip setpoint. The RBM system uses inputs.from the Local Power Range Monitors (LPRMs) surrounding the L erroneously withdrawn rod. As a result of the event, the local bundle powers in the vicinity of the withdrawn control rod and the core average power increase and reach new steady-state conditions. Th'e power distribution in the immediate vicinity of the withdrawn control rod is highly peaked, thus reducing thermal margins in the adjacent fuel bundles. The two primary thermal margins of concern in this event are the Minimum - Critical Power Ratio (MCPR) and the maximum Linear Heat Generation Rate (LHGR). PP&L's Steady-State Core Physics Methods (Reference 1) are used in
'8 modelling the RWE event. In the RWE event, the rea<; tor power increases.
slowly, and the total increase is relatively small. The event may therefore be evaluated using steady-state analysis rethods. I g
. g.
LI l l-... w
-- +
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y I 2.1.2 Sensitivity Studies I The consequences of the RWE event are primarily driven by the reactor initial conditions, the worth of the control rod, and the RBM response, SIMULATE E analyses and statistical analyses were performed to determine g the srnsitivity s of calculated RCPR to changes in input assumptions. The parameters of interest examined were:
- 1) Control rod pattern
- 2) Error control rod location l
- 3) Xenon concentration g
- 4) RBM setpoint #
- 5) LPRM_ failure rate i
- 6) Failed LPRM location (s)
I, '
- 7) Failed RBM channel
- 8) Lore average _ power
- 9) Total core flow
- 10) Cycle exposure The base case values for the analyses are listed 'in Table 2.1-1.
l As previously mentioned, the main output of the RWE analysis is a RCPR for g a given RBM Trip Setpoint Setting (typically 108%). The consequences of the RWE event are primarily sensitive to the neutronic coupling of the 10 - L I I ~ . . . - .-
t
.5 I
limiting CPR bundle with the withdrawn rod and the RBM response. The neutronic coupling is increased as the limiting bundle is moved closer to the location of the rod being withdrawn, referred to as the error rod location. The sensitivity analyses showed that the control rod pattern, I error rod location, and the xenon concentration significantly affect the neutronic coupling. Control rod patterns that force the power distribution near the error rod location create large rod worths and large I RCPRs during a RWE event. These rod patterns, although technically achievable, will realistically not be used during normal reactor operation
;l_ due to the excessive power peaking. Therefore, use of this type of rod pattern is a significant conservatism built into the analysis method.
For the fuel management and operational strategies in use at Susquehanna ' I SES (i.e., quarter core symmetry for operational control rod patterns and reload design). 'the limiting rod locations are in the center of the core as shown in Figure 2.1-1. RWE analyses with the error rod locations I outside the area shown in Figure 2.1-1 will not exhibit higher peak powers j than the highest peak powers possible in the area shown. Tne impact of i changes to operational and reload strategies on the selection of the error ' rod location will.be evaluated on a cycle by cycle basis. I In addition to rod patterns, the xenon concentration significantly affects the capability of forcing the power distribution toward the error rod location. At full power xenon conditions, a significant amount of rod inventory is withdrawn in order to maintain criticality. Therefore, fewer
- l. control rods are available to force the limiting bundle toward the error rod location. RWE cases with equilibrium xenon show that the zero xenon
, assumption increases the RCPR by as much as 0.08. This assumption produces very conservative results.
. )) .
t g
-. _i___________________._____
Ii IL In conjunction with MCPR and LHGR operating limits, the RBM system acts to prevent the fuel from' exceeding its thermal limits by providing a rod l l block when the local power has increased to the flow-biased trip setpoint.
! As the RBM trip setpoint increases the coatrol rod may be withdrawn further. The RBM system is a dual channel system using input from the g{'
L four LPRM strings surrounding the rod selected for motion. The channel )
'A' subsystem uses input from the level 'A' and 'C' LPRMs; the channel 'B' subsystem uses input from the level 'B' and 'D' LPRMs. Either channel
'l must reach the trip setpoint to activate a control . rod block.
Operability of the LPRMs directly affects the RBM system response.
- l Because some LPRMs are allowed to be failed during core operation, the number and location of failed LPRMs will strongly influence the RBM response. Technical Specifications allow rod motion with one channel of g'
the RBM system inoperable. Also, the RBM system design allows up to one half of the LPRMs providing input to a given channel to be failed. Figure i , 2.1-2 shows the minimum RBM response for channel 'A' as a function of a typical rod. withdrawal for 0,1, 2, 3, and 4 failed LPRMs. As expected, the RBM response degrades rapidly as the number of failed LPRMs increase, l Typically, two or three LPRM failures may exist in the whole core but are spread out sufficiently so that one RBM channel very rarely uses less than 6 of its 8 LPRM inputs (i.e., 75%). j t D In addition to the number of failed LPR"s, location of the failed LPRMs can significantly affect the RBM response. A total of 163 failure ) combinations exist for each RBM channel assuming 0, 1, 2, 3, or 4 ' l failures. A failure combination is defined as a specific set of failed l LPRM locations. Table 2.1-2 shows typical channel 'A' and channel 'B' .! responses as functions of the error rod position for different failure combinations. Figure 2.1-3 shows a typical channel ' A' RBM response as a ') function of error rod position for four different failure combinations L each having four failed LPRMs. Both Table 2.1-2 and Figure 2.1-3 demonstrate the wide variability in the RBM response for different failure 1 I I
l l' I l ll 1
- combinations for a given error rod position. This variation in RBM response as a function of rod position is directly related to the location of the operable LPRMs with respect to the error rod tip location. As expected, when an operable LPRM is close to the error rod tip location, the LPRM is more responsive to the error rod movement. As the error rod I tip location moves away from an operable LPRM toward an inoperable LPRM, the RBM channel response may in fact be reduced due to the shifting of 'I. axial power shape. Failure combination 163 in Figure 2.1-3 shows this ef fect. Both Table 2.1-2 and Figure 2.1-3 also demonstrate that no one h failure combination for either RBM response is the most conservative throughout the whole rod withdrawal. The most conservative LPRM failure combination would be one which would provide the minimum RBM response throughout the entire withdrawal.
I The CPR response during the RWE event is very important in establishing the MCPR operating limit. As previously mentioned, the initial conditions I of the event place the limiting CPR bundle near the error rod location. This causes the CPR to decrease and RCPR to increase rapidly during the etror rod withdrawal. Figure 2.1-4 shows a sample Unit 2 Cycle 2 RCPR response as the' error rod is withdrawn. The limiting CPR may be in different bundles during the event, but the limiting bundle is always close to the error rod location. The RCPR response is steadily increasing and flattens out when the error rod is nearly full out. Figures 2.1-5 and 2.16 show the responses of peak LflGR and core thermal 8 power which are also of interest. For the case shown, the conservative initial control rod pattern results in the steady state LiiGR limit being exceeded. This would not be allowed by the plant Technical Specifications. For all cases analyzed, including the example shown in the figures, the peak LHGR does not exceed the transient LHGR Protection Against Fuel failure (PAFF) limit, and the core average power increase is B l I
i E1
,S I'
small. Therefore, the RWE event does not result in LHGRs exceeding the fuel mechanical design limits. The variation in the steady state LHGR and PAFF limits with control rod tip position occurs because these limits are l{ functions of assembly planar exposure. Initial core power and core flow conditions can also significantly affect ; the magnitude of the RCPR response during a RWE. RWE cases along the 100% rod line showed that the RCPR is essentially constant with a variation of about 0.01 in RCPR. This variation in RCPR is well within the conservative assumption of zero xenon. At powers below the 100% rod line, however, the Technical Specification flow-biased RBM trip setpoint is farther away from the initial RBM response. To determine the effect of the RWE at less than the 100% rod line, RWE analyses were performed on the 90%, 80%, 70%, and 60% rod lines to - determine initial CPR and RCPR sensitivity. As power decreases, calculated RCPR does increase but the initial CPR increases more rapidly, g, To place a bundle on MCPR or LHGR limits, the rod patterns had to be 5 adjusted. This adjustment required major rod insertion in the periphery of the core. These adjusted rod patterns had many secondary *ods at l intermediate and deep positions which are not typical of power operation and caused violation of the MAPLHGR limits used to protect against LOCA. g Because no credit is taken for the intermediate and low Rod Block Trip Setpoints, the error rod could be fully withdrawn. All of these assumptions result in excessive conservatism and a RCPR increase of approximately 0.04. To provide a more realistic assessment of RCPR, the zero xenon assumption was removed and equilibrium xenon was assumed. g W. These additional cases showed that the 100% rod line RWE analyses with ' zero xenon bound the equilibrium xenon cases at lower powers, h The last parameters analyzed for the RWE were the exposure effects due to , previous cycle shutdown exposure and current cycle exposure. Beginning of Cycle 2 of Unit 2 RWE analyses were performed for different previous cycle I L I_
i
\
j i exposures (11.220 GWD/MTV,12.050 GWD/MTV, and 13.000 GWD/MTV). The results show a minor effect of previous cycle exr'sure on the RWE l g analysis. All the RCPRs were within 0.01 of each other. Typical reload end of cycle exposure windows (i.e., the range of previous cycle exposures l on which the current reload cycle analyses are based) are less than the
,B 1.78 GWD/MTU exposure difference analyzed. Therefore, future RWE analysis !
{
-results should not be dependent on the previous cycle window. ' .I i
The other exposure effect, cycle exposure of the design cycle, does affect yf the RCPR for the RWE event. As expected, this effect derives from the , critical control rod density required at different cycle exposures. As peak hot excess reactivity is approached, the control rod density reaches its peak. Peak control rod density conditions allow for a greater number of inserted control rods than at BOC to force the limiting bundle location l to be closer to the error rod location. This effect is similar to the !- zero xenon assumption. In addition to the rod pattern effect, the fresh bundles are very close to their peak reactivity. Therefore, neutronic coupling between the limiting CPR bundle and the error rod location may be increased at the cycle exposure corresponding to peak hot excess i_ reactivity. In order to determine the sensitivity of these effects on L RCPR, RWE analyses were performed at 5.5 GWD/MTV cyc*re exposure which I corresponds to Unit 2 Cycle 2 peak hot excess rentivity. These analyses showed that the RCPR increased by approximately 0.03 at 5.5 GWD/MTU u I compared to BOC which demonstrates that a cycle exposure effect on the : RCPR for the RWE event exists.
.g 2.1.3 Licensino Analysis Method The assumptions made in the analysis of the RWE are based on the results and' discussions of the sensitivity studies described in Section 2.1.2.
Four coder are used to perform the analysis: SIMULATE-E, kBM, RBMSTAT, 15 - I I
I and STATOL. The SIMULATE-E ccde provides the detailed neutronic and I thermal hydraulic feedback information during the RWE event. The RBM code uses the SIMULATE-E predicted detector readings and predicted CPR values l to generate a RWE response surface of RCPR as a function of RBM trip setpoint, operable RBM channel, and LPRM failure combination. The RWE g response surface is used as input to RBMSTAT along with inputs of nominal g RBM setpoint, LPRM failure rate, the uncertainty _in measured RBM trip a setpoint (drift, accuracy, and calibration), the uncertainty in calculated RBM response, and the uncertainty in calculated RCPR. RBMSTAT then statistically combines the input through a Monte Carlo approach es described in Section B.3. This method results in a comprehensive and
- detailed description of the analyzed RWE event covering all p
- :=ible l
failure combinations and RBM responses. The output of RBMSTAT is a RCPK cumulative probability distribution with at least 95% confidence. The ! RCPR cumulative probability distributioe is used to determine the MCPR operating limit using the methods described in Appendix B. i l The following assumptions and methods are used for the RWE licensing ' analysis:
- 1) Cases are run at 100% power /100% flow.
- 2) A control rod pattern that forces the limiting core MCPR I
I location to be within one control cell of the error rod and is predicted to be near critical is developed. l
- 5) The error rod location is located in the area shown on Figure 2.1-1.
- 4) The error rod initial position is full in.
- 5) Zero xenon is assumed as the xenon concentration for 100%
power rod line cases. I
i g. i l I i
- 6) At least one-half of the LPRM inputs will be assumed operable as enforced by the RBM system design. Every possible failure l combination will be considered. A conservatively high LPRM failure probability is assumed. j
- 7) One RBM channel is assumed to be inoperable. Each channel is '
assumed to have an equal probability of being the inoperable I channel.
- 8) The RBM measured trip setpoint uncertainties, calculated RCPR uncertainty, and the calculated RBM s ponse uncertainty are
'l independent. The 95% probability /95' onfidence level uncertainties are used (see Table 2.1 1).
I 9) Analyses will be performed for both BOC and peak hot excess I reactivity. The Statistical Combination of Uncertainty (SCU) method described in Appendix B can be divided into a number of steps:
- 1) Create a RWE-response surface from SIMULATE-E and RBM calculations that relates calculated RCPR to the variables to be analyzed statistically. The variables are RBM trip setpoint, LPRM failure combination, and RBM channel.
- e
- 2) Define measured and calculated uncertainty distributions for the parameters used in the response surface.
- 3) Perform a Monte Carlo analysis with RBMSTAT (as described in Section B.3) to produce a cumulative probability distribution of calculated RCPR.
I I.
B B
- 4) Perform safety limit type analyses to produce cumulative probability distribution functions of fraction of pins in boiling transition for a range of MCPRs. For the sample calculation presented herein, cumulative probability g-distribution functions were generated by ANF for a core containing all ANF 9x9 fuel, since these distributions wiil be typical of future licensing applications.
- 5) Assume a MCPR operating limit and perform STATOL calculations '
that use Monte Carlo analysis to' combine the cumulative probability distribution functions for fraction of pins in boiling transition and the cumulative probability distribution function of transient RCPR, thus producing a combined safety g limit and transient analysis.
- 6) Select the value of number of rods expected to be in boiling transition at the 95% confidence level. If the value is greater than 0.1% of the pins, increase the assumed MCPR' operating limit and repeat steps 5 and 6.
I 2.1.4 Samole Licensina Analysis This section presents a sample licensing analysis using SIMULATE-E, RBM, RBMSTAT, and STATOL using the previously described methodology. 2.2,4.1 Response Surface A response surface of calculated RCPR as a function of RBM setpoint, LPRM failure combination, and RBM channel was produced with the SIMULATE-E and h RBM codes. The SIMULATE-E calculations produce predicted detector responses and limiting CPR as the error rod is withdrawn beginning at notch 0 (full in) and ending at notch 48 (full out). For this case, the error rod is incrementally withdrawn by 4 notches so a total of 13 sets of I E
I I
- 6) At least one-half of the LPRM inputs will be assumed operable as enforced by the RBM system design. Every possible failure combination will be considered. A conservatively high LPRM failure probability is assumed.
I 7) One RBM channel is assumed to be inoperable. Each channel is assumd to have an equal probability of being the inoperable I Chai?'r i .
- 8) Ihe RBM measured trip setpoint uncertainties, calculated r.CPR uncertainty, a: , the calculated RBM response uncertaint' are j independent. The 95% probability /95% confidence level uncertainties are used (see Table 2.1-1).
l 9) Analyses will be performed for both BOC and peak hot excess reactivity. < The Statistical Combination of Uncertainty (SCU) method described in Appendix B can be divided into a number of steps:
- 1) Create a RWE response surface from SIMULATE-E and RBM calculations that relates calcu'ated RCPR to the variables to i
be analyzed statistically. The variables are RBM trip setpoint, LPRM failure combination, and RBM channel.
- 2) Define measured and calculated uncertainty distributions for the parameters used in the response surface.
3)' Perform a Monte Carle analysis with RBMSTAT (as described in Section B.3) to produce a cumulative probability distribution of calculated RCPR. Eg LI ug
B B 4)' Perform safety limit type analyses to produce cumulative probability distribution functions of fraction of pins in boiling transition for a range of MCPRs. For the sample calculation presented herein, cumulative probability g distribution functions were generated by ANF for a core containing all ANF 9x9 fuel, since these distributions will be g typical of future licensing applications, a
- 5) Assume a MCPR operating limit and perform STATOL calculations that use Monte Carlo analysis to combine the cumulative probability distribution functions for fraction of pins in boiling transition and the cumulative probability distribution function of transient RCPR, thus producing a combined safety g limit and transient analysis.
- 6) . Select the value of number of rods expected to be in boiling transition at the 95% confidence level. If the value is greater than 0.1% of the pins, increase the assumed MCPR operating limit and repeat steps 5 and 6.
2.1.4 Samole Licensina Analysig
.I This section presents a sample licensing analysis using SIMULATE-E, RBM, RBMSTAT, and STATOL using the previously described methodology.
2.1.4.1 Resoonse Surface I A response surface of calculated RCPR as a function of RBM setnint, LPRM failure combination, and RBM channel was produced with the SIMULATE-E and RBM codes. The SIMULATE-E' calculations produce predicted detector responses and limiting CPR as the error rod is withdrawn beginning at notch 0 (full in) and ending at notch 48 (full out). For this case, the error rod is incrementally withdrawn by 4 notches so a total of 13 sets of 1 I
.~ _ _ .__ _ _ . _ _ . . _ . _ _ . -
e ! i i predicted detector response and CPR result. The RBM code uses this information.to calculate the RBM response based on every failure combination for each RBM channel at each axial position. The RBM code i then creates the RCPR response surface, an array of RCPR as a function of RBM setpoint, LPRM failure combination, and RBM channel. 2.1.4.2 Treatment of Uncertainties I e The uncertainties in measured RBM setpoint, calculated RBM response, and , calculated RCPR are included in the RBMSTAT calculations discussed in Section B.3. Table 2.1-1 shows the list of uncertainties and their 95% probability /95% confidence values. The measured RBM setpoint uncertainties are derived from the RBH system performance with respect to instrument drift, accuracy, and calibration. The calculated RBM response is based on the ability of.StMULATE-E to predict the detector response at the location of the LPRMs. The TIP response comparisons presented in I Reference 1 and additional comparisons obtained since that time have ' provided a large data base to derive a cumulative probability distribution of predicted LPRM response error and therefore predicted RBM response error. The calculated RBM response uncertainty is calculated as follows: , onsa " otenM i I where: N = the number of operable _PRMs. N is set to the minimum allowable value of 4 for conservatism. l I otp,, - the uncertainty in calculated LPRM response. The calculated RCPR response is based on the ability of SIMULATE-E to predict the radial bundle power response. A linear correlation of RCPR as a function of the change in radial bundle power is first established to
'l I
g-
E I derive the RCPR uncertainty as a function of radial bundle power I uncertainty. The radial bundle power 'incertainty is discussed in Section 2.2.3. The slope of the correlation, B, is conservatively adjusted to its l 95% probability level and used in the following relationship to calculate the RCPR uncertainty for the RWE event, g o,c,, = B (2 o,, wherg: o , - nodal code bundle power uncertainty l For this sample analysis, the 95% probability based value of B is 0.155 MW". The correlation coefficient of RCPR as a linear function of radial l bundle power based on analyses is 0.999. This correlation coefficient demonstrates that the RCPR is directly related to the radial bundle power ' during the RWE. l 2.1.4.3 Results
.The resulting MCPR operating limit for the sample Unit 2 Cycle 2 RWE using I
j the Statistical Combination 'of Uncertainty (SCU) methodology presented in Appendix B is 1.26 for a 108% rod block trip setpoint at 100% flow. It l' should be noted that if other events require higher operating limits, then. the actual-Technical Specification MCPR operating limit would be the higher value. I 3 I I' I' a;
t
- I TABLE 2.1-1 Rod Withdrawal Error Base Case Inout Assumptions I
Parameter Assumotion
- 1. Core Power 100% !
I 2.- Core Flow 100% g 3. Xenon' Concentration 0%
- 4. RBM Setpoint 108%
- 5. LPRM Failure Probability 15%
- 6. RBM Measurement- ,
Uncertainties at 95%/95% l Probability / Confidence I a. b. c. Drift Accuracy Calibration 3.0% 1.6% 1.6%
- 7. RBM Calculational 6.4%
Uncertainty.at'95%/95% s Probability / Confidence t- 8. RCPR Calculational 0.042 i L Uncertainty at 95%/95% , Probability / Confidence I E e r L3 33 n
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.WW M. e.N. W.e.T.e.N. @.h.o.e.. . S. =.N.W.T.h.M.e N. M. .g .e.. p p o. M. .. M. I 9 e . . . .M W @ . e $ $ 9 h. h ..S e. h h o p h h. m a c e g M N. -l g .....=. ................ . . . . . . N. M. == N. . . . . M. M. ....= g e. .m. ......=. .e M. M e m e I 9' ; e
.i l
I
~**
o s. c.o.p. .e e.c.h . .m e. M e.o.N.h e.N.e.S.N.e.e.e.s. o.e.M.e.o.N.o.m.e. e. s,e.e.m.e.M, l, e
====.ecesN oMNoop M S e W e . N = = N N N N .N N NTM.TeTM s
' g a
s . . . . . . . . . . . . . . . . . . N. N . . . . .. . . . e. e. .M e. ...........= . e o c e. N N N N = = = .m. .p h. e. T N. e. e. H
;l i
- i e
3@ 't
-s M. ccho*NeoGTM9@MMe
@-. T. S. c. o. e. =. T. o. N. M. . M. e. h S. . s. o. N. d.. W.
. .N. . . O.... S. o. N. S. . N. N . O N ,j.l MM=Sh@hhe@
OM
= '
h e g a
=NN
= = . N. N. N..N. N N. N. N. N. N. . ..N. .N N. N .N N N.=.=.NN .N.
.N ohoeh.
..N .N .=N=
. h h T.h.
p a== N
-1 l
= s- l i ,! .' e s. N.h. @ . e.v.e. e.=.T. e.e.h.h.*.h .e.e. S. p
- l
- q. o N si T=ohoepeh o. S. M. . e . c. N =..o... .N = M. N e.. ..
. .. N. o e. . n o 1 p-
, -&. M. t ==.MooMNo.Sc .sheMeh amesh
' ====. N N = N.
4 e .t .h e .W..= N. . N. .e w .c h t et L: ii g. . . . N. N... N N N. ..N .. N. =.= .N .=h N N .N .N ......N. ...=N
-g o e.
. .T. o. S. =. T. N. T. e. M. o. o. M. M. o. m. . N. . $..0 .= p. M..M . M M o SP P 4
9 .-I ; oNThe . . 4@@@T. . .. s N - 'N . . o d. N = = . s. e @ e h d h e T h e T M N S M c M . 't@eh ,f' g et t e .NN.==...=.
.=======..N.N .=..
........... N=em =
. = . .N K t i
1 a .NhMhhhdh.co.oTSTNS
.. . .. ..... .. .... S.f se.
i 9 . V. . . - . . j F,- g N o
-I-s ' .... cN .=. cS.
4; W S h e h W T = oopeSM*SM === T M. o m . . e. N. 9. N. c. N. h. h s. e
.o .c o .= .= .* .= .
. . a=oM
====.. .. N o. N. .== = ..o .=N..*.=o.=.* .N. .S .=o.=e .* @M. e c e @ . w I
I m' W-o s a . W e. @4h e. S. W. @. e. h. T. M. e. h S. S. W.. T.
. . e. .e .........
T 8 4.e S e h e h o m o m T S N e T 1 S*S@S$h@TNSeN= I- 3 e
,%. w N- I 9N.coe@ece /l l .
.g w e t a .o .N... = . c .o .c .o c .o .= O.====
. = =e c o .= N.=S.oe.=N.= ..*e S====.N.=.=.=N..de.-
.J. e w ' O. .
..M. 'l
' W >
3 *. s o- M. o. c. h. o. o. M. o. c. o. c. S. e. .*..S.. p o o. p e. S. o. m. e. .=.. M . N . . N. S. . e. e p ')
-&* W . ... . . .. .. M hl .C
- e'. w -
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.- .c =
e-o
.e a
.i i
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..-.... .. ..eoN.
. .m
- o. .e h = S h e .@ M W J
=
o & o a y e 3 ;s h e. c. o. c..o. . S. N. . M. M. e e. N. o.. T. w..o.o.T.e.e.M. e. . .c .o .. T. o. e. W '3
. .h . n .= .h = N 14 'W e N OW l '#
E S = 4 t o.T
= = = ==oc.
WMWNNe@@eh@SeeSe@TMcMSe@Se@@McM.W.TowT i
~ N 's t a = = . o .c .c .o r. .C==.o..o. C ..o C. o c. o = = =. o. c. o. c. c o = .= *. *. . . .
- N.
=. .. i e.
- I -
t. eq g .
. 1 e e ' MSee9*M*
M. T. W. T.@MMOTRWMMWW .
@. h o.
- e. N. N. W...o . W O. M. h. h. M. h a e. o. c. c. T.e. N.M. h a e a .q;.
t
..t Nt e . t 4
.: .o.==. o 0 0 o .o .0 0 .. .o 0 0 0 .. 0 0 ... 0 0 0 000o000000000000000
. 0...=0 .M= e..e c... e@MM@TM e e @ e S e @ e .T W .' !
== y
# N- ,1
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. c. e. =. h W o. W. T. h
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=MNNNoace=====o====..NN=N.m.*.=NN=NNNNNT W t ..o .o .o .o .o ... .
o .o .= .o ... .o =. .o o o o o. . ..o .... o .o..o o o o. o o .o o. o o. o o o. o . > ,I o
'I e
,e e
Il '
- o. c.o. o.o. o.o.o. o.o.o. o. o. o.o.o.o.o. o c o c o o .c .o .o. .O c .o .O .o .C .o o o o o o o o o o ooooooooooooo
.o .o..=-===.
o o o o
.o .o .o .o o o. ooooooooo o o.o.o.o.o.o.o.o.o.o.o.o.o.
o....... o o. o o o o o o o o o. .o o o o o I i 1 1 4 1 I ll1' o e T I M MMX M M MM M MM MMM
; e M MX MMM MMMk 8
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- 4 M M M M M XX MM M M MM MMX MMX i
e t.
"1 o' l
~
3 N 6 MM M M M MM M MM MK MM e 6 MM MM MMM MM a s :i e o - t
- 8 'MMM M MMM MMM MMM MMM MMM- MMM W- P M
4 i 3 e e l- I 4 9 3 W 6 0 i M MMMM MMMMMM MMMMMX MMMM
'I I
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o M e MMMMMMMMMX t MMMMMMMMMX DI 'I .
+ ' e e N
i 4 MMMMMMMMMMMMMMMMMMMM 6 1_ e t *
. I MMMMM
. - ., e
', s
- g 8
3 I w = o I i E b 4 3 4 i TetheSc.NMTeemeSo=NMTechemo.NMWeeheme=NM l 3
.J 4
E
.I e t
i 6
.N.N. N.N.N.N .M ....M. M M M M.M.M.M.M.T . .. .. . ec.=.....
T T. T T==.T T T .T T m. @= m m m. e . W m m e @ W @ W l 29 a
- I';, Y
I, FIGURE 2.1-1 7 g, L CONTROL ROD l LOCATIONS FOR ERROR ROD 1 IN ROD WITHDRAWAL' ERROR ANALYSIS g L 3:
,, iiiiiii.iiiiii. g; r=: ,+ + :,+ +:.+ + '+
u . H L $7 '
- ss 53
+ + + + + + + + + - IR
+ +, + +, + + + + + + +
i si 49 i,
, - r- -
I h h h y 47
'% h h h y y L h $ 5 5 g..
43 - 4,_
+ ++ +++ ++ + + ++ + . . .+ +:
+ ~+ EV QS Sb y + W+ f +
I
;;- + + 4 _
;;- + + + .
+ + R N sN 98 Sh + +. + + + : 31 si- + + +
+ + SF R hy gs '+ + "+ + "+ :
f
';- + + + + + RS Sb M N sN + + + +. + : IL 21
+ v '+ +" + EQS Sb & N + + "+ +i + : I, 1e - + + + + + + + + + + + + + + +:
17 :
- 3
+ +
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+
ii 9
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7 5 I Y, 3
+ + + ++ + + -
g l l l l l l l 000204060810 12 14 16 18 202224262830323436384042444648505254565860 x I
@ Error Rod Locations l
9 LPRM 1.ocations - 43 l
?(
g3 s l'
' FIGURE 2.1-2' MINIMUM CHANNEL 'A' RBM RESPONSE AS Ai "l ' FUNCTION-OF ROD 1 WITHDRAWAL FOR o 0,1,-2, 3,? AND 41LPRM: FAILURES- " ,5 -
125
.+ .
12 0 - *
~
0* LPRM - i
.] FAILURES ' l N
lI L l b 115 -
-y 1 - <
*1* LPPM ~ /
[ FAILURES 3
$ /
- ' LE- *2* LPRM .
\
3: FAILURES - 11 0 -
*3" LPRM FAILURES 105- - :
4
. g:
y/ ~ ,4. LP M A, LURES 8 1ee O 4 8
/ . .
20
-12 16 24 28 32 36 40 44 48 IN AXIAL ROD POSITION OUT yg LI
,.,,,m.,,,. . , . .
d v ., ll 2"' FIGURE 2.1-3: IL
' CHANNEL 'A' RBM RESPONSE AS A FUNCTION 0F 4
' ROD WITHDRAWAL FOR'FOUR DIFFERENT:
-FAILURE COMBINATIONS WITH FOUR-l:
LPRM FAILURES: g 130 12 5 - FAILURE ,
- COMB 163 N .I
-- 120 - - -
8 FAILURE
- w. COMB 136 k 116- - .
g . a: (( .I
=
11 0 - - jjp I 105-
/
/
FAILUP.E COMB 124 N FAILURE COMB 94
/ .
7 i=
/
'O 4
.8
. . . . . . . . . . a 12 16 20 24 28 32 36 40 44 48 W AXIAL ROD POSITION I
n-
) J '. ' FIGURE 2.1-4L L- # RCPR RESPONSE ASDA FUNCTION , OF. ROD 1 WITHDRAWAL FOR AL RODiWITHDRAWAL EVENT e l, a 0.25- " i- i i z i' . O.20 - . . . . . . . . . . I I h O.15 - - - e
.j-
= 0.10 - t i .i. i .;. a. .;. , .;
I .
, 0.05 - . . . . . i-I O.00
~
0- 4 8 12 16 20 24 28 32 36 40 44 48 IN AXIAL ROD POSITION OUT '
--4, e' I t .
f i*', i .. 3, W* FIGURE 2.1-5 m LHGR RESPONSE- AS A FUNCTION g OF ROD WITHDRAWAL FOR A-ROD WITHDRAWAL EVENT g'; 19 11 % .
. g h ( O O O O O O 0% >
.g:
3 IF 18 - - - h I 17 - Legend- - 9 LHGR l' O PAFF LIMIT jp
' C X' STEADY STATE 16 -
LIMIT LHGR s
- igr ,P g.
j l . O /.
.I 15 - --
c; ~ -
,/ ;
1 14 - - - , I I s > ' _, y
; ; ; e O O- O~~ '
X xA'
? A A .- s 7, ;;, X ;z x : -
o L12 , , , , , , , , , , , n 0 4 8
, -12 16 20 24 28 32 36 40 44 48 IN AXIAL ROD POSITION OUT g ~ .
I 34 - l 3 .
~
,y ,
"h r.l FIGURE 2.1-6
" CORE THERMAL POWER- RESPONSE AS: A l' FUNCTION OF ROD WITHDRAWAL FOR A ROD:; WITHDRAWAL EVENT
!l (105:
h e ../ -
'g-104 .
I B / I I - /
, . 10 3 - - -
It >
-L 3
g 2: j 102 - - l ir / I ,.. e O l' '
" 101- -
a.
;3
/
- 100: ;
/. , , , i. _,. , , , , , ,
O 4- 8 12 16 20 - 28 32 36 40- 44 48 A A D POWlON IN OUT I
I 1 2.2 Fuel Loadina Error r 2'2.1: Event ^Descriotion p ' ? ,1
- A fuel loading error can ' occur during core loading prior to startup of a
. reload cycle. Two_ types of loading errors'can occur. -In the first-type,.
.g_
or mislocated bundle error, a fuel assembly is incorrectly loaded into'a- 5' location in the reactor core causing a discrepancy =between the intended'
- reload pattern and the actual core .anfiguration. In the second_ type, or l
- rotated bundle error, a fuel assembly is loaded into the correct fuel cell but is rotated by 90* or 180 from its correct orientation. For the analysis,.it is assumed that the misloading error goes undetected during
.(
the final core configuration verification procedure. Core monitoring, however, is based on the intended reload pattern and may thus overestimate the thermal margin. Reactor operation is assumed such that either the misloaded or adjacent fuel assembly is placed on the Technical g W
. Specification thermal limits.
The consequences of a fuel mislocation error are highly dependent on the- .
- exposure and enrichment of the misloaded and immediately adjacent- fuel -
assemblies. For Susquehanna SES, the most limiting case occurs when a ~ high' reactivity fuel assembly is .misloaded into a locatian intended for a i low reactivity assembly. At the beginning of cycle, the highest
- reactivity fuel is usually a once-burned gadolinia bearing assembly which q
, . is' as 'close as possible to peak reactivity (i.e., the exposure atz which '
the gadolinium has been reduced to residual l levels.) The location for the . misloaded assembly will be a location which was intended _for a highly exposed'twice-burned-or thrice-burned assembly or a heavily gadolinia j loaded fresh assembly. At the middle of cycle, the high reactivity assembly is usually a fresh assembly which has depleted to peak
-fl '
reactivity. The location of the assumed misloading is the location g planned for a depleted assembly which is adjacent to as many high R
- reactivity assemblies as possible. The placement of the additional L
l
,' >j.
+ I (misloaded) high reactivity assembly in a location adjacent to other high - eeactivity assemblies exacerbates the power peaking:in that core location. resulting in lower thermal margins. End of cycle is usually not limiting-since, by this- time, the fresh assemblies are _ past peak reactivity and the exposed fuel is relatively low in power. For the rotated bundle event, the fuel assembly is loaded into the correct I location but is rotated by 90' or 180' with respect to the intended - orientation. As a result, the channel fasteners and spacer pads force the top of the assembly to tilt toward the center of the control cell, changing the size of the water gaps outside of the channel. The water gaps at the lower part of the assembly will not change. Because the. Susquehanna units are C-lattice .nlants (i.e., equal water gaps on all sider.af the fuel channel), the consequences of this event are relatively small. Yhe pin enrichment distribution within a fuel assembly is symmetric about the diagonal.- Further, enrichments for pins which were - 5 ': intended 19 be adjacent to the control rod are typically low to help reduce the pin power peaking during control rod motion. Following the h assembly rotation, these pins will be adjacent- to the larger water gap which causes a corresponding. change in the pin power distribution. This l change will affect the actual CPR. Since the rotation was assumed to.be undetected during core verification, the core monitoring system will monitor the core as if the core had been loaded correctly and, consequently, the calculated thermal margin will not reflect the effect of the rotation. The intent of the fuel loading error analyses is to demonstrate that the CPR and LHGR safety limits are not violated during normal operation.- For both the mislocated and rotated bundle analyses, the critical parameter is the change in the CPR. For Susquehanna, the CPR is determined via the XN-h 3 correlation developed by ANF (Reference 12). A measure of the event I I l 3 .
g . 4p; < ' - l l-Q * - r' -severity is determined by ' calculating tn ' maximum RCPR which could be-M 's/ expected from all credible operating scenarios. - The RCPR is-defined as: - A> y
, MCPRu '- MCPRu MCPR u-I -
l where: MCPRete = the minimum CPR for the correctly ldaded
*g core, and MCPR,te = the' minimum'CPR for the misloaded core - -
(i.e., either mislocated or rotated g bundle) 5 u For the mislocated bundle analysis on a given reload, the worst credible misloci; tion is identified and analyzed. All applicable code uncertainties-are factored into the analysis to assure a conservative ~estima,te-of the RCPR. Fo,* the rotated bundle analysis on a given; reload, maximum changes _ ; in local; peaking and S-factors are determined. These changes are compared
-to the changes assumed for a bounding rotated bundle calculation to assure g that the bounding' analysis is still valid. g 2.2.2 Sensitivity Studies-2.2.2'.1 Mislocated Bundle Analyses The mislocated bundle analysis is performed using the SIMULATE-E computer program (Reference 5). The same models which have been benchmarked
~
against data and approved by thU NRC (Reference 1) are used for the g evaluation. 'To assure that the worst credible mislocations could be B identified, the effects of several key parameters were investigated to I 1 1 I' 1
4m y e; jip X: datermine their-impact on the analysis. From these studies, an analysis
. approach was' developed to assure a conservative estimate of the RCPR from the event. The items investigated included:
1)' Assembly type / Core location I- ~) 2 Core / cycle exposure Control rod pattern I 3)
- 4) Core monitoring system The most-important input assun.gtion affecting the RCPR for the mislocated bundle analysis is -the 'solection of the assembly and the location of the misloaded assembly. As mentioned earlier, the event RCPR will be ,
-maximized by_ replacing a low reactivity assembly with a high reactivity
-assembly. At beginning of cycle, this is done by misloading a once-burned assembly which is close to peak reactivity into a location intended for a heavily gadolinia loaded fresh or a highly exposeo' fuel assembly. By I '
middle of cycle, the gadolinium in the fresh fuel assemblies in the
- highest power location will have depleted to residual levels and the assembly will therefore be at peak reactivity. Thus at middle of cycle, the worst' case mislocation error results from replacing a highly exposed fuel assembly with a fresh assembly. In all-cases,'the mislocation must be at least one control cell away from the core periphery.- Otherwise, the
- neutron leakage prevents the assembly power from becoming excessive.
For.most reload analyses, the most severe mislocation occurs at the middle I of cycle. For beginning of cycle cases, the once burned assemblies have exceeded their peak reactivity exposures. The fresh assemblies are heavily loaded with gadolinia to reduce the overall core reactivity. This reduces the consequences of the mislocation. If the reactivities of the once-burned assemblies are at or before peak reactivity, then beginning of cycle cases can be limiting and will be evaluated. At middle of cycle,
.I 1
I !
I the worst mislocation occurs when a fresh assembly was misloaded in place I of an exposed assembly ~which was to be surrounded by other fresh l; assemblies..'The mislocation then places several fresh assemblies face m, adjacent-to one another. The higher power in this region _will accelerate - the gadoliniim depletion rate resulting in higher assembly reactivities ' earlier _in cycle. This effect, combined with the higher core average g
- reactivity, increases the power peaking which reduces the CPR. The lower u l: - CPR for the misloaded co.e will increase the RCPR. l d The end of the previous cycle exposure also affects the results from the I
mislocated assembly analysis. Higher cycle exposures from the previous >
- cycle will reduce the assembly reactivities of the exposed fuel. The h1 reduced reactivity in the exposed fuel shifts more power generation to the ?
fresh assemblies. When a fresh assembly is misloaded adjacent to other i
~
fresh assemblies, the power peaking is even higher. This results in reduced CPR and, hence, increased RCPR values. The-presence of control rods can also has-a significant effect on the i severity of the mislocation analysis. If a control rod is inserted in the L' control cell or adjacent control cell to the misloaded assembly, the RCPR ~ will be dramatically minimized or completely eliminated. For reload licensing applications, no credit is taken for the-presence of control rods. Many control rod pattern sensitivity analyses were performed to g; identify the control rod pattern which would maximize the RCPR for the : event. Most of these rod patterns intentionally shifted the power into ' the region of the mislocation. Based on these calculations, it was a determined that an all-rods-out calculation provides a consistently conservative RCPR. As a result, an all-rods-out rod l pattern will be used to evaluate the mislocated assemblies analysis for reload licensing applications. f1
-The core monitoring system will also influence the degree to which the effects of the fuel assembly mislocation will be factored into the thermal I
I
I' 1 I l limits l evaluation. For the Susquehanna SES units, the core monitoring system in-use is the ANF POWERPLEX system (Reference 8). In POWERPLEX, a nodal reactor code is used to calculate a three-dimensional power distribution consistent with the current state of the reactor core (i.e., I core: exposure, power, flow, control rod pattern,' etc.). This power
~ distribution'is then adjusted based on the core LPRM readings. During ,
this adjustment, which is performed in the UPDATE algorithm of POWERPLEX,
~
no credit is taken for core symmetry and each fuel assembly power distribution is adjusted based on measurements from its closest LPRM
. strings. Even though the calculated nodal' power distribution is based on I a correctly loaded core, the LPiM readings will be influenced by the ;
effects of the mislocated assemblie:. The UPDATE algorithm will increase the calculated assembly powers in the aica of the mislocation. This would I result in a lower RCPR relative to the case when the monitoring system is not considered. Typically, the monitoring system is not credited in the event analysis.- i ( 2.2.2.2 - Rotated Bundle' Analyses l 1The rotation of a. fuel: assembly in the reactor core causes-a significant shift'in-the pin power distribution. This is caused by the change:in the ' size of the water gap surrounding the fuel channel. The change in power , distribution is determined by using the CPM-2 computer code (Reference 4). l 1 This computer program has already been reviewed and approved by the NRC for use in licensing applications for Susquehanna SES (Reference 1). As a I . result of the power distribution change, the assembly secondary local , peaking factors (S-factors) will also change. These changes in local peaking factor and S-factors will affect the XN-3 calculation of CPR. Sensitivity studies were performed with both CPM-2 and SIMULATE-E to determine the effect of water gap size, assembly exposure, assembly axial ! 41 ! I ;
power distribution,- and fuel assembly design. CPM-2 was used to determine l the effect of each parameter on the calculated peaking factors while l' 4 [E - SIMULATE-E-was used to determine the impact on the calculated thermal margins.1
, Each parameter was varied to determine its impact on;the calculated. l
- peaking factors. These calculations were then used to determine a set of local peaking factor and S-factor changes which would conservatively bound all anticipated future designs. The magnitude of the conservatism was q
~ determined by. comparing CPR calculations using these bounding peaking factor changes with: calculations using'a best estimate change in peaking
- h. i factors. The bounding peaking factor changes used in the analysis are
- l. listed in Table 2.2-1. The use of these peaking factor changes results in a maximum event RCPR of 0.17, as discussed in Section 2.2.4. {' .
The. calculation of the bounding peaking factor changes has several significant conservatisms. .The most significant conservatism is the g', assumption that the maximum possible fuel assembly displacement as 3.
-determined 'at the _ top of the fuel is the same throughout the: entire axial .
- length. Second, the bounding peaking factor changes were selected to
. provide significant margin to the expected peaking _ factor changes for f]'
actual. assembly designs.. Third, the rotation of the fuel assembly results - g
- in a decrease in fuel. assembly reactivity of up to'0.01- Ak. Although-the decrease in reactivity will provide.a decrease in the assembly. power and a ;
. corresponding increase in the assembly CPR, the effect is not credited.
Finally, the rod-patterns used to estimate the RCPR were designed to shift the axial power distribution to the top of the core and into the rotated E 5 [ assembly. Since the peaking factor changes are largest at the 70% void
. levels, this further exacerbates the calculated RCPR.
l
- To determine the magnitude of the conservatism in the bounding analysis, several calculations were performed using best estimate peaking factors.
-The axial tilt of the assembly was also considered such that no assembly L ,
I I
t I , displacement was assumed at.the bottom of the core while maximum assembly displacement was assumed at the top of the core. ~ The rod patterns used
, for the analysis _ were typical of operationally acceptable patterns. The- ;
assembly chosen for the rotation was the fuel assembly closest'to-the CPR I limit. The maximum RCPR from these cases was 0.07 which illustrates the large amount of conservatism provided by the use of bounding peaking. I factor changes. 'As a' result, no additional conservatism or uncertainties I ' need to be-applied to the analysis.
.If'the rotated bundle analysis should ever become the limiting event, some relaxation could be made to the bounding peaking factor changes presented in Table 2.2-1. If.any relaxation is made, thorough. analyses will be performed'to ensure that the new set of peaking factor changes continues to. provide conservative results for this event.
2.2.3 Licensino Analysis Methods 2.2.3.1 Mislocated Bundle Analysis The mislocated bundle' analysis is typically performed at the middle of l' cycle for the reload being analyzed. If fuel assemblies from the previous cycle are at or before peak reactivity exposure, then beginning of cycle cases =will also be examined. Typically, cycle step-out analyses for the correctly loaded core using anticipated operational control rod patterns L , will be performed-from the upper end of the expected exposure range of the j 1 previous cycle. Restart calculations are made from the cycle step-out in-which all control rods are withdrawn. These calculations are typically done at cycle exposure intervals of 1.0 GWD/MTV. The exposure points around peak core reactivity are of most interest as they usually provide the limiting values for RCPR. The radial power distribution from the , all-rods-out calculation can then be used to select the worst location for LI .
I the misloaded assembly which will occur at the location which has the highest assembly power for a non-fresh assembly. If multiple batches are present for the exposed fuel, several locations may require in:1y n . If several batches are present for fresh fuel, only the most reactive type need be considered (i.e., lowest gadolinia content and/or highest enrichment.) A cycle step-out depletion calculation is performed for the misloaded core with the same rod patterns which were used for the correctly loaded core analysis. Thermal margin is evaluated for all rods out restart calculations. Comparisons to the thermal margin calculated for the correctly loaded core are needed to generate RCPR values for the specific eisload. The largest RCPR from all mislocation evaluations at all exposure points becomes the event RCPR. Altho: gh the analysis uses inputs which generate conservative RCPR's (i.e., assembly location, control rod pattern), the SIMULATE-E calculation providts a best estimate calculation of the MCPR. As a result, it is necessary to add appropriate calculational uncertainty into the result to assure conservatism. For the mislocated bundle event, the RCPR can be conservatively estimated as: RCPR = C (RP2- RP3 ) where: RP 2 - Relative assembly power in MCPR location for the misloaded core, RP, = Relative assembly power in MCPR location for I the correctly loaded tore, RCPR = Event RCPR, and I I a_
, - . _ . , _ , , , - -. .pr------
y 7
- I ,
'C -
Constant from regression analysis, including appropriate uncertainty factor Note that this correlation is u' sed only for the purpose of developing the I .RCPR uncertainty. Actual SIMULATE-E calculations using the XN-3 correlation will be used to evaluate the event RCPR. The RCPR standard.
) deviation can then be derived as:
onCPR = o,p (qarg: o,p = nodal code asstmbly power uncertainty i fg - l The assembly power uncertainty can be conservatively estimated by using L the integral TIP' comparisons which are available from Susquehanna SES
, operation. -The database includes all steady-state TIP sets taken for Units-1 and:2 through April 15,.1990. Each TIP set for Susquehanna SES-lI l consists of measurements for each of 43 individual strings. Only TIP strings with-known measurema.it problems or excessively large errors (i.e.,
errors greater than five standard deviations) were eliminated. This- f resulted in a total of 6976 integral TIP string comparisons. The assembly i power uncertainty (based:on assembly relative power) is: L- o,p = 0.0277 This produces'an RCPR standard deviation for the mislocated bundle analysis of: ( o,cp, = 0.0320 l A tolerance factor on the event RCPR was calculated by statistically combining the RCPR .1 certainty calculated above with the actual RCPR g I)
I distribution from the mislocated bundle analysis. The RCPR code error is I assumed to be normally distributed which, based on data, is conservative since data from the assembly power uncertainty indicates that the distribution is actually more peaked than a normal distribution. The resulting tolerance factor bounding 05% of all RCPR values is: Koun m = 0.037 2.2.3.2 Rotated Bundle Analysis The CPM-2 code is used to rvaluate the pin power distributions for all ( fresh assemblies being lor.ded into the new reload cycles using the maximum g possible assembly displar.ement resulting from a 180 rotation. The changes { in the peaking factors Ii.e., local peaking factor, interior S-factor, and { peripheral S-factor) are calculated by comparing the rotated assembly I calculaticns to the non-rotated assembly calculations at various exposures and void levels. The maximum change in each peaking factor is compared to l the bounding value provided in Table 2.2-1. If the changes in peaking factors are less than the values presented in that table, the bounding { analysis can be used and the resultant RCPR for the event will be less ! than 0.17. If the changes in peaking factors are greater than the values given in Table 2.2-1, explicit SIMULATE-E calculations will be performed using the calculated peaking factor changes. Due to the large ' conservatism in the bounding analysis described in Section 2.2.2, no additional uncertainties are applied. { t 2.2.4 Samole Licensina Analysis l A sample reload analysis was performed for the Unit 2 Cycle 2 reload core. For the mislocated bundle analysis, the effects of depletion were modelled using a cycle step-out approach. The maximum RCPR including code ( uncertainty for this event was: Ii l B! i El
I RCPR ,= 0.144 g. u Assuming a safety limit of 1.06, the event ACPR becomes:
.ACPR ,= 0.18 w
)I For the rotated bundle analysis, the maximum peaking factor changes were.
. computed for the fresh fuel. assemblies. The maximum changes'in peaking factors are.given in Table 2.2-2. Comparing these values to the bounding
~
s i values'in Table 2.2-1 demonstrates that the bounding analysis is valid and ,. the RCPR for this event will be: 1 RCPR,u.< 0.17 sg .
. Assuming a safety limit of 1.06, the event ACPR becomes: i ACPR,u < 0.22 ;
The-large magnitude of this value is caused by the significant L. :.onservatism in- the bounding analysis as discussed in Section 2.2.2.2. u ! 1: g I __ _ _ _ _ < _ _ _ . _ _ _ _ _ _ . _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ . .~
, . _ _ _ _ _ _ _ . . . . . . . , . . _ _ _ . . . ~ . . _ _ _ _ . . . . _ _ . _ . . _ . . . _
TABLE 2.2-1 I ' Maximum Changes'In Local Peaking Factor and S-Factors For The Rotated Bundle Analysis Void. Maximum Change Level. Local Peaking- Interior Peripheral. ,
._(%1, Factor S-Factor S-Factot_ >
0 0.20 0.040 0.050 '
. 40 ' O. 2' 5 0.050 0.080
-70 0.30 0.080 0.100 l
I I, I gi I: Io i I' L s ,
E EBf l: 1I' .,
.x -
h TABLE 2.2-2!
~, _ _ Rotated Bundle Sample Reload Analysis.
Maximum Calculated Chances in Local Peakina Factors and S-Factors
. Void .
B Level Local Interior- Peripheral - 4
.c ,
& .Peakinq Factor S-Factor S-Factor t -0.164
_0 0.0103 0.0455 40 0.225 0.0295 0.0692 70 0.285- 0.0514 0.0959 I I I I I I I I I . I
__ 7-ifj/. i 3 : 7 s s ,
= 2.3- -Lass"of'Feedwater Heatina- 4
~, .
2.3.1 Event Desqriotion
.f
~
Theiloss of feedwater heating transient at Susquehanna SES can be caused =l
-- by either of. the following system malfunctions:- g{
Bu
. 1) the closure of a steam extraction line to' one of the feedwater a
- v -heaters, or i
- 2) the' bypass of one or more of the feedwater heaters. .
- The result-from either malfunction is a decrease in the feedwater enthalpy ,
-as it enters the: reactor vessel. The reduction in feedwater enthalpy will cause?a corres'ponding-decrease in the core inlet enthalpy. As'a result, g;;
, . the neut'ron' flux and corresponding core thermal power increase due to the' B' additional moderation'at the-bottom of the core. This change in thermal j x power isl accompanied by an increase in steam flow. The i'ncrease'in steam
? flow results .in-an increase in' the. steam line pressure drop' and a -
corresponding; increase in the core pressure. If the-increase in power. h; level does not cause a scram, the core thermal power and core pressure. will' reach a new steady st' ate level with a- small shif t in the- axial power distribution toward the bottom of the' core, t
. At Susquehanna SES, there are three feedwater heater trains in each unit. I1j
'Each train is composed of five separate feedwater heaters. The as-built
. system was analyzed to identify the worst single failure of equipment- or single operator error which would cause the largest possible change in' j:
feedwater temperature. The two most severe failures were: it: 1) f.ailure of the-relay coil closing the extraction steam lines to feedwater heaters 3, 4 and 5 on a single train. This would I I s
I I j result in an approximately 50'F decrease 16- feedwater
' temperature. !
2)- drain failures resulting from the loss of control air to four
.feedwater heaters in all three strings and moisture i separators. . This would result in an approximately 46'F-decrease in the-feedwater temperature.
I The assumption used for the licensing analysis is a 100*F decrease in the 1 g feedwater temperature. This provides additional conservatism in the lL analysis. 'I L The consequenceof s a loss of.feedwater heating are an increase in the core thermal power and a corresponding decrease in the limiting bundle L CPR. _The operating limit CPR must be sufficiently high such that the core-
.g MCPR after the transient is greater than.or equal:to the safety limit d W =MCPR.
2.3.2 Sensitivity 'tudies The' loss of feedwater heating analysir is performed using the SIMULATE-E- i computer program (Reference 5). The same model development process benchmarked by PP&L and approved by the NRC in Reference 1 is used for the ! e' valuation. To develop a conservati.ve an'alysis approach, several key
~
parameters were investigated to. determine their impact on the results. These included core pressure, control rod patterns, assembly reactivity, 1 and core exposure, i The first effect investigated was the change in CPR due to a change in core pressure. As noted in the previous section, the core pressure will increase as a result of the increase in core power. The amount of the I i l- I 1 l
I change will be dependent on the change in steam flow which is in turn Ii
' dependent on the change in core thermal power. This increase in core pressure has the effect of decreasing the calculated post-transient CPR , resulting in a larger ACPR. Typical changes in the core pressure are relatively small. A conservative increase of 5 psi will be assumed in the licensing application of this analysis.
The goal of the licensiag analysis is to produce a conservatively large ,
'ACPR starting at any possible initial core state such that the final core I state MCPR will be at the safety limit. The resulting operating limit MCPR must be greater than or equal to this ACPR plus the tafety limit. l, Because of the small change in CPR associated with this transient, it is i difUcult if not impossible to define a set of operationally realistic !
initial conditions which will provide a post-transient MCPR equal to the safety limit. ' Two mechanisms were used to study the effect on ACPR as the final state MCPR approached the safety limit. The first method used control rods to shift the power distribution into a localized area in the core. Although - many of these control rod patterns were operationally unacceptable (i.e., violating control rod sequences, excessive power peaking, etc.), the l '!I j trends observed in these calculations were consistent with the results l' obtained from more realistic control rod patterns. The second method which was used to reduce the post transient MCPR was to increase the g, L reactivity of the limiting assembly. Gradual increases in the assembly W reactivity were made until the post-transient MCPR was at the safety L.- limit. The results from these cases showed trends consistent with the rod I L - pattern study. Finally, the calculations were reprodlced at a variety of , cycle exposures. Both proposed control rod patterns as well as Haling 4 calculations were used to provide the cycle depletions. h'i The results from all of the calculations showed that the post-transient core MCPR was strongly dependent on the pre-transient core MCPR and l
I it- i I ! insensitive to cycle exposure or control rod pattern. The results from these calculations are shown graphically in Figure 2.3 1. Due to the high . degree of correlation between the post-transient core MCPR and the pre-transient core MCPR, a regression analysis was performed using the two
~ variab1:s. The resulting regression line is: '
MCPR, = 1.108
- MCPR, - 0.051 Eq. 2.3.1 Wher.t: HCPR i - pre-transient core MCPR ;
MCPR, = post transient core MCPR ' I The largest deviation from the regression line for all cases analyzed was less than 2%. The mean square error from the analysis was 0.000096, indicating a high' degree of accuracy in the fit.
- Tolerance limits on the regression analysis were generated such that, with 95% confidece, 95% of the residuals will be bounded b 'U. limits. The tolerance iirrits were' evaluated for the regression 1.nv and are plotted with the data in Figure 2.3-2. Note that all of the data is bounded by the toleran:e limits. Throughout the range of interest, this tolerance interval it bounded by a constant adder of 0.024. For simplicity of l
application, a constant of 0.024 will be added to the resulting pre transient core MCPR to adequately cover analysis uncertainty. Thus, Equation 2.3.1 becomes: MCPRi = 1.108 MCPR, 0.027 Eq. 2.3-2 , The RCPR for the transient is defined as:
- I '
I ,
, . . . - . , . . , . e~- ,- -e .- -r ,
I! ; I; RCPR = 1
~
MCPR 1 ! Substituting Equation 2.3-2 into this definition gives: ' 0.108 MCPR, - 0. 027 Eq. 2.3 3 RCPR = 1.108 MCPRg - 0.027 Althogh this equation has been adjusted to account for analysis approach uncertainty via tha cvsluated tolerance limits, the regression analysis is ; based on calcuiations from SIMULATE-E. As a result, a SIMULATE-E RCPR I code uncertainty for this transient must be applied to the results from ; Equation 2.3-3. A conserystive adder has been developed based on : SIMULATE E's ability to predict a change in core thermal power. This , bounding value applicable to the loss of feedwater heating analysis is: Ko,c,,, % = 0.01S Eq. 2.3-4 The RCPR of interest is the value given a post-transient core MCPR equal to the safety limit. Substituting this into Equi. tion 2.3-3 and adding the bounding RCPR uncertainty, the loss of feedwater he.iting ' licensing RCPR can be calculated as: ; 0.108 SLMCPR - 0. 027 RCPR" " = 1.108 SLMCPR - 0. 027 + 0. 015 ; where: SLMCPR - safety limit for the reload being licensed I 1
. S. .
E' I
r s 'I 2.3.3 LLeensino Analysis Method i The analysis described in the previous section is generic to Susquehanna ' SES, provided significant changes to the operating strategy are not made I which would affect the regression line in Equation 2.3 1. To assure continued applicability of the generic analysis, specific SIMULATE-E analyses will be performed for each reload licensing application at 800, MOC and EOC using the anticipated operational control rod patterns. If new data-indicates that the regression coefficients are no longer l applicable, a net: :ot will be generated to maintain a conservative evaluation of tho loss of feedwater heating event. 2.3.4 Samole Licec ina Analysis I To demonstra'.e the results of a loss of feedwater heating evaluation, a i l sample r'il',ad analysis was performed for Susquehanna SES Unit 1 Cycle 3. ' Table 2.3 1 shows the BOC, MOC, and E00, SIMULATE-E calculated initial and l final ECPR values and the corresponding initial MCPR values using the ! regression line. Since the calculated values are within the 95/95 tolerance limits, .the generic regression analysis is applicable to Unit 1 Cycle 3. For thi:: unit and cycle, the safety limit MCPR was 1.06. Using { this in Equation 2.3 5 provides the maximum RCPR for this transient ! (including all applicable uncertainties): RCPR, = 0.091 ' I Converting to a ACPR yields: i ACPR = 0.106 I ! 1I ; 4 L
T TABLE 2.3-1 I Loss of Feedwater Heating Sample Analysis como11ance with Generic Reoression Analysis Cycle $1MULATE-E I Exposure RegressionAnalgsis Cal Bounds on MCPR, 8 (GWD/MT) M,LglationMi$ (pwit VoDer BOC(0,0) 1.570 1.689 1.665 1.713 MOC(5.0) 1.469 1.585 1.553 1.601 l E0C(10.35) 1.439- 1.529 1.519 1.567 l' a h l
- 1) MCPR initial
'- MCPR,,- = final'(po(pre st-transient)transient) MCPR MCPR l
- 2) Values of MCPR, based on the calculated SIMULATE-E MCPR, and the 95/95 tolerance limits to Equation 2.3.1 u I ,
i I I i
I l
. I FIGURE 2.3-1 l LOSS OF FEEDWATER HEATING EVENT CHANGE IN MINIMUM CRITICAL POWER RATIO :
I 1.7 1 f X 1.6 - x I X I 1.5- . . X
- x.. ,
X 3 X
$i .4 k :
fl 2 x x 3 e '
- 1,3-
; . .x . ,
XX! , 1.2- - y , 7 x ,
- I l g g 1.2 1.3 1.4
~
1 1.1 1.6 1.6 POST-EVENT MCPR
~
I f
\
I! , l
- FIGURE 2.3-2 i
LOSS OF FEEDWATER HEATING EVENT 95/95 TOLERANCE LIMITS ON l L 2.2 REGRESSION ANALYSIS l s I:
,// I.
1,3 ... . . . 't 4 . i D. O
' / .* .
2 . 1.6 - - E, 2 . I
.i. ',
j,4_ .
.h .i . i.
l 1.2- - - -
- - . l -
tiQ.TES: 1) CALCULATIONS USING SIMULATE E - FOR SSES 1 CYCLES 2 & 3
- 2) TOLERANCE LIMIT BASED ON REGRESSION ANALYSIS 1 i i i i i i i i i 1 1.1 1.2 1.3 1.4 1.5 1.6 1.7 1.8 1.9 2 POST-EVENT MCPR I
I q tI 2.4 Shutdown Marain Determination l 2.4.1 Event Descriotion I Core Shutdown Margin (SDM) is defined as the amount of negative core reactivity at 68 degrees F and 0% xenon with all the control rods fully inserted except for the strongest reactivity worth control rod, which is I fully withdrawn. The SDM analysis is performed using SIMULATE-E and RODDK-E as a part of the reload analyses and to demonstrate compliance wicn plant Technical Specifications. The SDM evaluation is not considered l to be related to an accident or abnormal operating occurrence. The SDM analysis demonstrates that the core will be subtritical in the most reactive (i.e., cold) condition with sufficient margin using only control 1- rods. The Technical Specification required margin assures that under the specified conditions, the core will not be critical considering the uncertainties associated with manufacturing tolerances and calculational ; method. As a result of these uncertainties, the minimum required demonstrated SDM-is 0.38%-Ak/k. For predictive calculations typically performed during the reload design process, an additional uncertainty is added to the 0.38% Ak/k value to cover uncertainties in predicting the l critical k-effective bias for SIMULATE-E. This prediction uncertainty is evaluated prior to finalizing the reload design and updated if appropriate. This evaluation considers all past critical calculational f { comparisons and the current fuel and core design. The limiting SDM (i.e., I the smallest SDM) for the operating cycle occurs near the peak cold reactivity core condition with the most reactive control rod fully withdrawn. The core and local fuel exposure and void history significantly affect the reactivity worth of the strongest rod and the ; core SDM. Therefore, the core SDM analysis is performed as a function of cycle exposure and uses conservative values for the previous cycle fuel exposures, t I LI .
u i 2.4.2 Sensitivity Studies SIMULATE E and RODDK-E calculations were performed for Unit 2 Cycle 4 to Ii
]
determine the sensitivity of calculated SDM to changes in input assumptions. The parameters considered were: -l
- 1) Previous cycle exposure, and It
- 2) Cycle exposure for the design eveln.
I.[ , y The results of these various sensitivities are graphically represented in ' Figure 2.4-1 'As the figure depicts, a lower previous cycle exposure , assumption results in less SDM. This is expected since the core is in a more reactive state due to the irradiated fuel being less exposed and, hence, more reactive. I
- Cycle exposure for the design cycle also affects the SDM. Figure 2.4-1 shows that SDM varies significantly as a function of cycle exposure. For typical reloads, the limiting SDM occurs either at BOC or near EOC. The j
limiting SDM occurs near the time when the fresh fuel reaches its peak reactivity (near 'lij) or when the once-burned fuel is at its peak reactivity (BOC or near B00). 2.4.3 Licensina Analysis Method-For each reload design, the 0.38% Ak/k SDM Technical Specification Limit l and the related predictive uncertainties are verified to be applicable g, considering the most recent data, and if necessary, the uncertainties are W. reevaluated. Core SDM is then evaluated using the RODDK-E code and the SIMULATE-E core simulation code. The strongest worth rod location is ' l determined by selecting a number of rod locations ranked by RODDK-E as , potential strongest worth rod locations at the specific exposure point of l interest. The exposure point of interest will be based on a contred rod step-out depletion using the lowest expected energy production of the L
g' ; I . l previous cycle. The strongest worth rod candidates will be used as input in the SIMULATE-E evaluations to determine the analytically determined strongest worth rod at the exposure points (within at least 1.0 GWD/MTV of each other) throughout the entire design cycle. Core SDM is evaluated using the equation below: 1 - (K-efffE) - Bias (E)) I SDM(E)= (K-eff(E) - Bias (E)) x 100% . Etuttg: SDM(E) = the core SDM at cycle exposure, E; Keff(E) = the SIMULATE-E core k-effective with the strongest worth control rod fully withdrawn at 68'F and xenon-free conditions at cycle exposure, E; and Bias (E) = the SIMULATE-E cold critical k-effective bias at cycle exposure, E. h The bias equals the expected ( SIMULATE-E critical k-effective minus l.0. The minimum core SDM of all the SDM(E) values shall be greater than the < SDM design limit which includes allowances for manufacturing tolerances - and calculational uncertainty. This value is then used to establish the R value which is defined as the difference between the B0C SDM value and the minimum SDM value. The R value and other calculated parameters (e.g., control rod worths) are utilized in the Shutdown Margin Technical Specification test performed at BOC. 5 :
I 2.4.4 inmole Licensina Analysis I This section presents a sample core SDM analysis for Unit 2 Cycle 4. The sample SDM design criterion of 0.8% Ak/k was used which includes the 0.38% Ak/k Technical Specification value and the 0.42% Ak/k prediction 'E uncertainty. From Figure 2.4-1 the minimum core SDM value is 1.28% Ak/k. g The BOC SDM value is 1.29% Ak/k and the R value is 0.01% Ak/k. The a minimum core SDM value exceeds the 0.8% Ak/k design criterion, and therefore sufficient SDH has been assured for the Unit 2 Cycle 4 core. I I I I I I I. I I I
- g. ,
I I
I m a
- s
=
=
D~3 a mOsb E2 es 0 0 1 1 2 2 3 3 4 0 6 0 . 0 .5 0 5 0 O - E 4/ X =
=
8 P 1,
- O s u
e - S s U 2, - . ex R = E = U E E m 23, - . . F m F C 0 4 E C 8-0 C s u Y 4, 0 A T m C L 3 8 K-w - - SF E % / K. I a OG m E K S. A. X 5, P
/
K M P
/ NU O
CR T E L. OO S U C H E D. E . L O ECP XYR PCE OE e R 6, S OLV S R.2 E P I W setO s ( E C G. N.
~
*A U R U S E4 - u s
G L I C.. R. l E W 7, M I . T. S1 e I T E. D R. H
/ l O. } -
u m M N. T 8 i U
)
s_ u u m e O 9 ~ m 2
. W n a
- N s e M 1 e 0
A
~
~
m s
/ R s 1
1 G I o m N e m n a s es '
I 2.5 Standby Liouid Control System Canabil Ry I 2.5.1 Event Description The Standby Liquid. Control System (SLCS) is designed to provide an
.alterr. ate' method of bringing the reactor from fM1 power operation to a cold (68'F), xenon-tree shutdown condition withod reliance on control E u
rods. At Susquehanna SES, this emergency shutdown ;apability is provided by injection of a sodium pentaborate solution through the standby liquid . control sparger located in the lower plenum of the reactor core. The l minimum boron concentration in the reactor after injection is 825 ppm. However, the analysis of the SLCS is intended to demonstrate that the l I- reactor will be at least .01 AK shutdown with only 660 ppm boron present. The SLCS will only be used in the event that the control rod drive system fails to insert sufficient rod inventory to bring the reactor to a cold xenon-free shutdown. Due to the undesirable consequences of 1'aovertent system operation, there are no automatic initiations. The SLCS is - manually initiated by the operators in accordance with the Susquehanna'SES
- emerger.cy operating procedures. The analysis assumes that the control rods remain at their hot operating positions with no additional insertion.
l 2.5.2 Sensitivity Studies The SLCS analysis is most dependent on the concentration of boron assumed in the reactor coolant. Although the system is designed to insert at least 825 ppm boron, only 660 ppm boron will be credited in the analysis.
.The reduced value is to account for imperfect mixing of the sodium pentaborate solution.
h' The effect of a given concentration of boron will be dependent on the ; water to fuel ratio of the fuel assembly. A higher ratio means more boron atoms will be present in the core, thus increasing the boron reactivity , I
I worth per ppm. Fuel design parameters which affect the water to fuel ratio will consequently affect the boron worth. The analysis developed l here is for application to the current fuel designs which are in use at Susquehanna SES (ANF 8x8 and ANF 9x9-2). If significant changes e made I to the. fuel assembly design parameters that would alter the boron worth characteristics, new boron worth parameters will be developed. I Two methods can be used to evaluate the effectiveness of SLCS. The first method applies a conservativFy low boron reactivity worth to a non-borated nodal core k effective calculation. The second method modifies the thermal absorption cross section in the nodal calculation to directly l model the boron in the reactor coolant. For both methods, the nodal evaluation is performed with the SIMULATE-E computer code (Reference 5). I To determine a conservative boron worth, numerous lattice physics j calculations using the CPM 2 computer program (Reference 4) were made to B. determine the reactivity worth of 660 ppm boron. The reactivity worth of . the boron was evaluated for numerous fuel designs at various lattice exposures as: Apkon - N'"5" ~
- PP" K6 soppe
- Koppe / 660 ppm Whgra: Kuo , = lattice K. with 660 ppm boron at cold xenon-free conditions, K,
o
= lattice K, with 0 ppm boron at cold xenon-free conditions, and I Ap' % = boron reactivity worth per ppm boron es .
I a
I l A conservatively low boron reactivity worth was developed to bound all I anticipated fuel designs and lattice exposures. The value selected base d on the CPM 2 evaluations was -20x10~5 Ap/ ppm. This provides at least 5xlC~; l Ap/ ppm margin to the minimum calculated boron worth. If the fuel design changes such that the bounding reactivity worth is no longer applicable, the value will be updated accordingly. The CPM-2 computer program was also used to determine the change in the g thermal absorption cross section at various lattice exposures for a different fuel designs. For a constant boron concentration, the change in the thermal absorption cross section is relatively insensitive to exposure for a given lattice design. Therefore, a beginning of life calculation can be used to determine the necessary adjustment to the cross section. Recalculating the lattice reactivity with the modified thermal absorption l cross section shows tttt this approach conservatively estimates the lattice reactivity until well past the exposure of peak assembly reactivity. At high exposures, these modified cross sections may g underestimate the lattice reactivity by as much as 0.006 AK. Although a fuel assemblies at these exposures are generally on the core periphery and contribute little to the overall core reactivity, the calculated core eigenvalue with.660 ppm boron in the reactor coolant will be conservatively increased by 0.01 AK when using this approach. 2.5.3 Licensina Analysis Method The purpose of the SLCS licensing analysis is to show that for the reload - being analyzed, the SLCS will provide adequate shutdown margin assuming cold xenon-free reactor core conf tions. The core eigenvalue calculation is performed with the SIMULATE E computer code. At least 0.01 AK shutdown margin will be demonstrated for each reload cycle to assure shutdown capability and to account for code uncertainty. The SLCS analysis can be performed using either of the two methods described in Section 2.5.2. The l first uses a conservative boron reactivity worth. If this method fails to g I I_i
I = demonstrate the required shutdown margin, the cross section modification process will be used to mt,1e accurately account for the effects of the l boron worth.
'g . The boron reactivity worth used in the first method was developed to bound all anticipated ANF 8x8 or 9x9 2 fuel assembly designs. If the fuel I mechanical design changes in such a fashion as to alter this value, new' conservative boron worths will be developed and applied. The analysis assumes that the control rods remain at their hot operating positions.
Cold, xenon free SIMULATE-E calculations are performed at various points in the cycle. The cold k-effective from these calculations are adjusted g for bias and boron wortb. The amcunt of shutdown margin is calculated as: SD = K[M' - K,Y/ " I ' wht.tt: Klif = critical eigenvalue after adjustment for bias KU; " - eigenvalue after adjustment for bias and boron worth l SD - amount of shutdown in units of AK If.the conservative boron worth approach does not demonstrate the required
.g 0.01 AK shutdown margin, a more accurate approach is used. in the more W. accurate method, the effect of the 660 ppm boron concentration is accounted for by a change in the thermal absorption cross section. This change is based on beginning of life lattice calculations for each fuel type in the core. The change in thermal absorption cross section based on g
1 l I-
' beginning of life calculations is made in the SIMULATE-E calculation. The resulting eigenvalue is adjusted for model bias plus a conservative adder
- before, at least 0.01 AK shutdown must be demonstrated.
l.< of 0.01 Af 2.5.4 Samole Analysis A sample SLCS licensing evaluation is provided using both approaches for Unit 1 Cycle.2. Using the conservative boron reactivity worth method, 660 I' , g ppm boron,prvvides 0.047 AK shutdown margin. Using the cross section E modification approach, 660 ppm boron provides 0.070 4K shutdown margin. The difference in the results can be attributed to the conservative value of boron reactivity worth used in the first method in which the bot i reactivity worth was intentionally reduced to assure a conservative evaluation of the SLCS capability. Therefore, both analysis methods l' demonstrate greater than 0.01 AK shutdown margin. I I' I. I I 3u l I .; I l
B lI 2.6 RETRAN Transient Analysis Inouts This section describes the generation of core physics data required for RETRAN transient analyses. The initial conditions for the reactor must be defined to begin a transient analysis. RETRAN transient analyses to B develop MCPR operating limits and perform the ASME overpressure analyses are performed for a range of initial conditions. The parameters required I~ to define the initial reactor state are system pressure, core power, core flow, core inlet subcooling, and fuel exposures (i.e., time in cycle). Core physics data is generated once the necessary parameters are defined. The process to develop one dimensional kinetics input to RETRAN is described in detail in Appendix A of Reference 2. The process consists of three major steps. First, using SIMULATE-E (Reference 5), an initial core state is calculated in three dimensions. This case is referred to as the base case. If I modelling a scram is required during the transient, additional SIMULATE-E cases are run with intermediate and all-rods in control states. The information from these cases must be collapsed to one dimension to be used by RETRAN. This collapsing is performed by SIMTRAN E (Reference 11). Second, SIMTRAN E performs perturbations on the three-dimensional
'g moderator density and fuel temperature arrays from SIMULATE-E to determine the dependence of the cross sections on these parameters. The perturbed three-dimensional data is then collapsed to one dimension in the axial I- direction. All of the collapsed cross sections and their associated one dimensional density and temperature distributions are used to produce
. polynomials that represent the cross sections as functions of the change in density and fuel temperature from the initial values. These polynomials are used as input to RETRAN.
f I I g:
I; I' The dependence of cross section change on moderator density change is extremely important for transients with large void changes (i.e., severe l; pressurization transients). Since RETRAN models the core as a single channel and SIMULATE-E models the core as individual bundles, the SIMTRAN E generated cross sections must be modified to account for this difference. The axial moderator density change in a tnree dimensional array of fuel bundles and in an average one dimensional model of a fuel g' W bundle may be quite different for a given pressure change. Thus, the cross section polynomials are recorrelated to minimize this effect of collapsing from three to one dimensions. This recorrelation assures that : i-the change in cross sections predicted by RETRAN for a given instantaneous pressure change is the same as that predicted by SIMULATE-E. This l . modification procedure is described further in Reference 2. Uss'of this cross section modification procedure is not required for transients in
'g' which a small change in the void fraction occurs.
- I I:
tu I. I I. I I
I I 2.7 Loss of Coolant Accident Inouts l The Loss of Coolant Accident (LOCA) analyses that support the Susquehanna SES units were performed by Advanced Nuclear Fuels Corporation and are contained in References 30 and 31. PP&L's design and analysis methods (References 1 and 2) will be used to provide fuel and core design
]
I information for input to the cycle specific heatup calculations performed by ANF and to confirm that the core neutronics parameters are bounded by the assumptions used in the Susquehanna SES blowdown and reflood calculations. The LOCA inputs calculated by PP&L are (1) local power distributions for each lattice type, (2) fuel rod power history, (3) core f scram reactivity versus time, (4) moderator density reactivity (i.e., reactivity as a function of moderator density), and (5) Doppler reactivity (i.e., reactivity as a function of fuel temperature). g The local power distributions are calculated as a function of void history 5 and exposure using the CPM 2 computer code for each reload fuel assembly. These are provided to ANF for use in the heatup calculation. ANF selects the iimiting local power distribution to analyze using their NRC approved LOCl< methods. The peak cladding temperature is nkvated and verified to ,l be liss than 2200'F.
- The fuel rod gap conductance and, hence, initial stored energy are dependent on the fuel rod power history. Rod power histories are generated by combining the CPM-2 local power distribution with the
~
g SIMULATE-E cycle depletion calculations. ANF reviews this data to assure that the initial fuel rod stored energy for the reload fuel is within the bounds of that assumed in the LOCA analysis. ANF used a conservative scram reactivity insertion rate, moderator density , reactivity, and Doppler reactivity in the LOCA blowdown calculations of L I
I References 30 and 31 to help ensure that cycle specific variations of I these parameters would not require a new LOCA blowdown calculation. methods will be used to determine cycle specific scram, Doppler, and PP&L l moderator density reactivities to demonstrate that the ANF blowdown calculations are conservative. All rods out (end of cycle) conditions will be evaluated because the scram reactivity insertion rate is worse
- when control rods are initially at their full out position. At other points in the cycle, the more adverse moderator density reactivity coefficient and Doppler reactivity coefficient are less significant than the higher scram reactivity insertion rate. Therefore, PP&L will calculate scram, moderator density, and Doppler reactivities for end of cycle conditions and compare them to the ieactivities assumed in the ANF LOCA calculations,
{ gl RETRAN will be used to calculate the scram reactivity as a function of rod position. First, SIMULATE-E and SIMTRAN-E cases are run for end of cycle conditions. Then, the RETRAN base model inputs are modified so that moderator density and fuel temperature feedback is eliminated from the neutronics calculation. With this revised model, the RETRAN 1 one-dimensional kinetics calculation is performed to simulate a reactor j L scram. Total calculated reactivity from RETRAN is converted to units of dollars using the core average delayed neutron fraction (froa the SIMTRAN-E calculation) and compared to the values used in the ANF LOCA g{ i analysis. l The reactivity as a function of moderator density is obtained from .l SIMULATE-E calculations. The SIMULATE-E base case used for the scran curve calculation is also used as the base case for the moderator density reactivity calculations. Additional SIMULATE-E cases with no temperature feedback are run using various pressures, holding inlet enthalpy constant, l) i The resulting SIMULATE-E calculated core average moderator densities and g core k-effectives are used to calculate a reactivity table. Normalized J 72 - p I;
densitics are calculated relative to the base case and the corresponding reactivity is calculated using the following equation:
~
1.0 '1 1 1.0-(A b-A 1) Y whgtg: pi = reactivity in dollars, I 1,= core k-effective from the base case, li = core k-effective from perturbed pressure case i, B = core average delayed neutron fraction from SIMTRAN-E. I During the initial few seconds of the LOCA, the cycle specific moderator density reactivities must be demonstrated to be more negative than the values used by ANF in order to validate tie conservatism of the LOCA analyses. After the first few seconds, the negative ' cram reactivity shuts down the reactor. The fuel temperature (Doppler) reactivities are calculated by SIMULATE-E. The SIMULATE-E base case used for the scram calculation is also used as the base case for the Doppler reactivity calculations. Additional SIMULATE-E cases with no moderator density feedback are run using various core average fuel temperatures. The reactivities for each perturbed case are calculated using the same equation shown above for the moderator density reactivity table, e cept with 1, defined as the core k-effective from the tempeiature pertoebed case i. During the initial few seconds of the LOCA, the cycle specific fuel temperature reactivities must be less than the reactivities used by ANF in order to validate the conservatism of the LOCA blowdown calculations.
1 g; In summary, the data described above is generated using PP&L's analysis Il i: methods and used by PP&L and ANF to assure the validity of the reported LOCA results and the Technical Specificatinn MAPLHGR limits, l r I: l
)
3: I I. I 4 I I I r 17 IL I. l
.If l
74 I. I
I I , 2.8 Control Rod Droo Accident Inouts The generic parametric based methodology used by Advanced Nuclear Fuels l Corporation (Reference 28) to calculate the peak deposited fuel rod ' enthalpy will be used by PP&L in order to calculate the consequences of I- the Contiol Rod Drop Accident (CRDA). The CRDA has been shown to be the bounding reactivity insertion event (Reference 14). The CRDA is the postulated dropping of a fully inserted and decoupled control rod of maximum worth. The control rod is assumed to fall at its maximum velocity and the worth of the control rod is assumed to have the I maximum incremental worth consistent with the withdrawal constraints of 1' the Banked Position Withdrawal Sequence (Reference 33). The initial l plant response of a CRDA is a prompt power burst which is initially terminated by Doppler feedback from the fuel temperature increase. Void reactivity also assists in terminating the power burst, but this effect is I not considered in the analysis. The prompt power burst can lead to significant fuel cladding failures and an increase in reactor pressure.
-The final shutdown via control rod insertion is achieved by the high neutron flux scram. '
I The specific criteria used to assess the cycle specific CRDA are to verify the following:
- 1. The radially averaged fuel rod enthalpy is less than 280
, I cal /gm at any axial location in any fuel rod.
- 2. The maximum reactor pressure shall be less than the pressure that will cause stresses to exceed the ASME " Service Limit-C" i :- l . stresses. Past evaluations (References 14 and 28) have shown that the CRDA is not a concern for reaching the ASME " Service LI
I Limit C" values, and that the transient evaluations for the I ASME overpressurization event bound those of a CRDA.
- 3. The radiological consequences based on the assumed failures of all rods that exceed 170 cal /gm shall be less than the General Electric evaluation for the initial core (i.e., the radiological consequences.of 770 GE 8x8 fuel rods failing)
(Reference 14). Otherwise, new radiological releases will be determined and results submitted with the reload application, i The important parameters used in the analysis are the worth of the dropped l' L control rod, local power peaking factors, delayed neutron fraction, j Doppler reactivity, and the scram reactivity. These values are used in g ANF's generic parametric analysis where the radially averaged fuel rod ! enthalpy deposition is determined as a function of control rod worth, four i bundle peaking factor, delayed neutron fraction, and Doppler coefficient of reactivity. The ANF generic parametric analysis used both a conservative scram curve and a conservative positive reactivity insertion rate for the dropped rod. I,' To assure consistency with the ANF CRDA methods as well as consistency with Technical Specification requirements, the following assumptions are g used for the initial core conditions: 1
- 1) Hot zero power ( .l
- 2) Zero void fraction -
- 3) Zero xenon concentration j l
- 4) Equal moderator and fuel temperature g i
- 5) Inoperable Rod Worth Minimizer l I;
I
l l
- 6) Operable Rod Sequence Control System i
- 7) Eight inoperable rods may exist and are fully inserted, g 8) Inoperable rods are separated by at least two control rods in all directions.
- 9) One or.two inoperable control rods may only exist ><u us Rod Sequence Centrol System rod group.
The maximum control rod worth and the four bundle peaking factor at the limiting time in the cycle will be determined by CPM-2 and SIMULATE-E l calculations. The Doppler coefficient of reactivity and the delayed neutron fraction are core wide best estimate values which are evaluated at the limiting times in the cycle. , Results for a sample Unit 2 Cycle 2 CRDA analysis are shown below which will be applied to the ANF-generic parametric analysis:
- 1) Maximum rod worth (mk) = 12.2 3
- 2) Doppler coefficient (Ak/k/deg F) = -11.0 x 10
i L 3) Delayed neutron fraction (B) = 0.00571
- 4) Four bundle local peaking factor (P4BL) = 1,47 Based on items 1) through 4) above and the generic parametric analysis (Reference 28), the peak radial average fuel rod enthalpy is 209 cal /gm.
'h This enthalpy is less than the 280 cal /gm limit. In addition, less 395 fuel rods exceed the 170 cal /gm limit, and the radiological than 1
7 I 1
}
i i' consequences are less than the initial core evaluation. Therefore, the I L. consequences of the Unit 2 Cycle 2 CRDA fall within the acceptance criteria. l
.I Ii ,
I. . I, I; Is i I? g, l . I. . I' I. I' I . I
1 I ! l I. 1 l 2.9 MCPR Safety limit inouts l The methodology used by ANF to perform MCPR safety limit calculations is described in References 17 and 18. As part of PP&L's reload licensing , I effort, data generated with PP&L's codes and methods will be Lsed by the fuel vendor to perform MCPR safety limit analyses. Tbc data will be used in both the conventional safety limit analysis an<' the " safety limit type"
.I analyses perforraed as part of the Statistical Combination of Uncertainty (SCU) methods described in Appendix B.
The codes and analysis techniques utilized by the fuel vendor are not { changed from those described in References 17 and 18. The only difference l is that some of the inputs are supplied based on PP&L's methodology. The l' data calculated to support the analyses using the fuel vendor's methodology consists of two major parts. The first part contains data r-g generated by PP&L's NRC approved core physics methods (Reference 1). The ) 5 second part consists of the system and fuel related uncertainties (See !
. Table 2.9 1) .
2.9.1 Core Physics Inouts ' E PP&L's NRC approved core physics methods (Reference 1) will be used to "
.g- generate three sets of data as input to the safety limit analyses: rod l relative power distributions, secondary local peaking factors, and core
, wide. radial peaking factor histograms. For the purposes of a safety limit calculation, a " flat" local power distribution in a fuel assembly is conservative. This fact is evident since, when an assembly is operated at the Safety Limit MCPR, a flat ul distribution contains more pins at higher powers. 1hus, a flat LI . , , . ll -
- g
m . gp = - y j n ;6 i distribution has more pins close to trar.: tion boiling than a strongly I
-peakeddEtribution.- l The local power distributions which will be used in the analyses .using~ the L fuel vendor's mathodology to perform the safety limit anelyses are (
generated using PP&L's-NRC. approved steady state core physics methods-
'(Reference-1). This data is generated by the CPM-2 code for a range of-q Y
void levels and lattice exposures which span the range'of expected values for the specific reload cycle. From this data, the most conservative g' m-local power distribution (s) will be selecteC (i.e., lowest maximutu local .;
;pekking factor for the exposure range expected during the cycle). -For l peripheral-assemblies,l a maximum local peaking 1'ctor equal to 1.0 is j
assumed (1.e., all fuel pins at the same power). The use of these local
, peaking. factor distributions 'is consistent with current ANE methodology; t only the lattice. physics code used to generate the local power
' distr!butichs 1s different, g s
L In addition to the local power distributions, the fuel vendor's y methodology also requires the limiting peripheral and interior secondary. . L local peaking factors.(S-factor:) for each of the lattice types in the .< [ reload core.. As-with the local power distributions, the S-factors are I j generated for a range of void levels as a function of lattice exposure. L This-data is generated using the CPM-2 lattice physics code. Tha third important core physics input to the safety limit calculation is ' h
- a; histogram of the number of fuel bucdles having a given radial peaking.
]
fcdor_ versustradial- peaking factor.. PP&L will perform a-cycle step-out .- analysis with the nodal physics code SIMULATE-E and generate bundle power- - Sisto0 rams at=each exposure point. The criteria used by ANF to select the :
" bounding" histogram (s) are: "
1). Histogram;from the exposure point which has the largest bundle IL a m radial peaking factor. J l J
. I, m .
1B L3
.- 2)- Histo vram having the largest number of bundles with powers close to the maximum bundle power.
!8 To comply with the fuel vendor's methodology in selecting the bounding
. histogram, PP&L will generate the information ne assary to produce these-
, histograms at each exposure point. 2.9.2 Uncertainties Both: the conventional safety limit and the " safety 1 tit type" analyses utilize a Monte Carlo procedure to combine various uncertiinties in order to demonstrate that 99.9% of the fuel rods in the core are not expected to exparience boil _ing transitio.n, in conformance with Standard Review Plan 4.4 (Reference 19). The uncertainties considered are listed in Table 2.9-
- 1. The " system uncertainties" listed in Table 2.9-1 concern measurement uncertainties of system parameters. The values used are the same as those I listed in References 17 and 18.
E-As described in Section 2.10, the CPM-2 lattice physics code will be used to generate inputs to the POWERPLEX core monitoring system. Of the " fuel
=ll '
related uncertainties" listed in Table .2.9-1, the only ones which are potentially affected by PP&L's methodology are: 1) radial' bundle power,
; 2) local power, and 3) axial power. Evaluations by PP&L were performed that demonstrate that the uncertair, ties on those three parameters used by
_ the fuel vendor (Reference 17) bound the uncertainties obtained using CPM-L : 2 input-in the core monitoring system. Thus, for conservatism, the _r : Reference 17 uncertainties for all the fuel related uncertainties are
- ; used. If other values of these uncertainties are used, further
. justification will be provided.
]j .' m
,..,,,p ,
m :q
. . o ,
. IbN
,I + . ;l I..[)\7 s,
- i W
9:y q
# TABLE 2.9-1
.).
g; l; g , g, W Uncertainties Used To Generate MCPR Safety Limit ' h 1 N* System Uncertainties- . l l -. Feedwater Flow Rata Feedwater Temperature .n Core Pressure yn Core Flow Rate l Core Inlet Terr.perature l l l u Fue1~Related Uncertainties 1 I iq XN-3 Correlation I Assembly Flow Rate . l
< Radial-Bundle Power
, Local' Power o l Axial Power
' h' ; .
l ii. 1 } J:: 82 19 ..
\ ? -
l' ' y ; .
/j . ,
li, ____-______2____ __-----___--_---_L'_-_------____-____---___-__--____--_----_-_.___----_-
i 1 q 2.10_ Core Monitorina System Inouts I PP&L' utilizes the POWERPLEX* Core Monitoring Software System (Reference 8) l to perform the on-line thermal margin calculations for Susquehanna SES; I ANF.is the developer of POWERPLEX, and the original application of' t POWERPLEX at Susquehanna SES was based on input from ANF's lattice physics- 1 code XFYRE (Reference 28). POWERPLEX input decks.are fuel cycle specific and contain nuclear physics data, peaking factors, and thermal limits ' inforriation. PP&L develcp::d, validated, and will "se an (nput deck generation methodology based on results from the CPM-2 lattice physics code (References 1 tad 4). PP&L has developed modifications to the NORGE-B2 code (Reference 7)_ to create the lattice physics input data needed by
- the POWERPLEX core monitoring system (e.g., cross sections, peaking L;Wg factors,etc.). The calculation flow path is shown in' Figure 1-2. CPM-2 L
. performs lattice physics calculations for specific fuel designs; NORGE-82 formats .the CPM-2 results into the POWERPLEX input deck format;- and then-L the NORGE-B2_ output and other POWERPl.EX inputs- (i.e., thermal hydraulic
- data', core geometry, calculation options, neutron detector response data,
~'
L - etc.) are combined to complete the POWERPLEX input deck for a specific cycle.- i g Benchmarking was performed using this process for Susquehanna SES Unit:1, E Cycles 1, 2, and 3 and Unit 2 Cycles 1, 2, and 3. The results were , evaluated against measured data, SIMULATE-E results, and POWERPLEX results using ANF generated input. The use of'PP&L',s input deck generation
-methodology for POWERPLEX~ produced better comparisons to measured
,f Traversing-Incore Probe'(TIP) data than the ANF generated input decks, and H results which are approximately equi /alent to the SIMULATE-E TIP response comparisons (Figures 2.10-1 through 2.10-6). The k-effective calculated
- by POWERPLEX using PP&L's input decks is very predictable and is compared
-'to the SIMULATE-E calculated k-effective in Figures 2.~10-7 through I
I
/
- n ,
4
.'I'
~ .9 +
o;jg 2.1012. ;T,%:e.results and statistical ' evaluations of:the.results lead to- J p-the conclusions that for Susquehanna SES,'POWERPLEX using PP&L gen ?ated g;, CPM-2; input'. data generates results better than those obtained.from- ! a POWERPLEX using.ANF generated inputs, and similar to the SIMULATE-E. L results., Because of this, PP&L's use of the ANF power distribution f uncertainties in the Safety Limit MCPR calculations is conservative and. f
~
? justified (See Section 2.9. of this report). O e '
u I Io l, ll i i - o 4 I y
- i
n. g7 84 - L- ! R.
) . t , ?
1 s< t
f 3 I
- l: FIGURE 2.10-1
.SUSQUEHANNA SES UNIT 1 CYCLE 1 l RELATIVE NODAL-RMS OF TIP: RESPONSE COMPARISONS 19.0 - ,
..i.
10.0 - . . , n o- Legend
- to.o' e O SIMULATE-E 15.0 -
I . Y. 14.0-A POWERPLEX (CPM) _ g - 13.0 - I'
>, t 2 12.o-e :
-a
- 4 _ 11.0 - . s <- ,
h 1o.0- -
..g - .
~
..o 1 g . .
g ..e. . a1a . , . . a g i p IA e .o . .
.at .
..O o O CD a ciOOD"
~
.o .g O o00 .6 6.... O'
;l: ... .. a. 00; .<
O to g . pa ZA
' 4,0 - - o Qi . . . . . .
3.0 - 5 2.0 - + + l.0 - 0.0 , , , ,. , , ., , , , ,
; O 1 2 3 4 6 6 7 8 9 10 11 12
;. CYCLE EXPOSURE (GWD/MTU) ja
.I.
y . l;-
. Im FIGURE 2.10-2 I-SUSQUEHANNA SES UNIT 1 CYCLE 2 RELATIVE NODALL RMS OF I::l TIP RESPONSE COMPARISONS ;:
20.0 -
,s.o. . .
. 3
' 1a.o- - ;- .
n o-is.o - Legend O SIMULATE-E'
' ~ ~
+
g' ' 16.o - A POWERPLEX (CPM) h 1 ,
.w - 14.0 - -
t
* {
] is.o- 4 g
2 12.0- - a . g 11.0-i
@ - io o. .
!O
$ e.o- -
e .o - - - g
- 7. o - .
- -4 s.0 -
0~ . , 6.0 - hh 90 4 O i- ( 4 o-j. 3.0 -
.i..
2.0 -
,. 1. 0 -
- l -
^
o.O 0 2 3 4 E-1 6 6 7 8 9 10 11 12 5: CYCL" EX/OSURE (GWD/MTU) I. Il,
-[
/k
_ . . - . . . . - . . - , - - , - - - + - - ,
a I ' I. FIGURE 2.10-3 SUSQUEHANNA SES UNIT 1 CYCLE 3. l RELATIVE NODAL RMS OF TIP RESPONSE COMPARISONS to.o- s < < *- is.o - ,, . 17.o- Legend
- I
= t o .o -
O SIMULA1E-E 16.o- ' t-6 POWERPLEX (CPM) 8 _ 14.0- -
; 13.o- , , , ,
! 12.0- - - - -
ti.o- .O O' OO g _ go.o. . . . . . o. l j e.o-fOO O-- , ? i b s.o- O , O! r -O- . -
. '. . R .O J 000
;'wg E -
.A; B-'
.7,o. 42 6. . ., 6:
g
.o . . . :-
s.o - O g s + 6 O' 6
-[A AI 6 4.0 - + - - - - - -
3.0 - i i 3.0 - ' 8 -i 1.0 - 0.o -
; O 1 2 3 4 6 6 7 8 9 10 11 12 CYCLE EXPOSURE (GWD/MTU)
.lL I l
- j. .
5 1
I I FIGURE 2.10-4 I SUSQUEHANNA SES UNIT.2 CYCLE 1
; RELATIVE NODAL RMS OF I 3.. TIP RESPONSE COMN'GISONS .l -j , . ,
20.0- u 10.o - -o- *
- 1
- is.o- .- ,
3 .ne - Legend - > > 18:D-O SIMULATE-E- ' " g'!i 15 4 ~ - - 6 POWERPLEX (CPM) 4 >
- 14.0-
; 13.o- '
6- - - E. 2 g 12.0 - + - 5.-- e J g 11.0- . . t . .
.-6 'i a ' '
o Z-- 10 4 - ' 'A ! 'A- ' s L & 'A A > > j@ ' e.o- a a j s.o - j. da go g;!o - o-0
= 3 A at.
~
g .o. A lj to o - c
~
$: *7.o- a S- - -
og og . g go o; s.o - . . . .
.p .i. # .
.i . . .O[ "O .
O. A- . 4.o . . . . ; . 3 , . , .i , .o - 3 '. g j- a Q. 44 . . . . . . . . . . . . ,
'O 1 2 3 4 6 6 7 8 9 10 11 12 13 3' k CYCLE EXPOSURE (GWD/MTU) l'.
g
,/,e
.. . . - . - ____________-_________:_---__2___--_____-___--_:
?I LI
!I FIGURE 2.10-5
;SUSQUEHANNA SES UNIT-2 CYCLE'2-I RELATIVE NODAL RMS OF TIP RESPONSE-COMPARISONS te.o- ',
. -< -i-. . -9
+-
1s.o - - - - . . . u.o - Legend: i e 1s.o' O SIMULATE-E 15 0" a POWERPLEX (CPM)
" 'i 8 14.0 - - . .
3R -- .
; 13.0- .. .
. , + , .i.
^
h 12.0-a p Y
.i.
O. .
.i..,
- g. 11.0 - . , ,
h 10.0- - go! ' O. .
.E.o.o a . .; . ,
g.o - g g ' O ;6! 7.0 O. f , .ai. .. $ 3 l a a .a
~
i Sai s .o - b. , s.o - U O i Y O-. 4.o . . . . . . . g .- 3,0
' 2.0 -
3.o . l g '. o.o , , , , , , , , , , ,
-gg
. 0 '1 2 3 4 6 6 7 8 9 10 11 12 '
CYCLE EXPOSURE (GWD/MTU) lI
- g lg,
, 1 m -- -_ - _ _ _ _ _ - - - _ _ _ _ _ _ _
B FIGURE 2.10-6L I SUSQUEHANNA SES UNITT2LOYCLE 3 RELATIVE NODAL RMS~OF- l TlP RESPONSE COMPARISONS . 20.o . 1s.o - .i-is.o-17*0 - Legend > 18 0 - O SIMULATE-E-15 0 - - A POWERPLEX (CPM)
.- 14.o- - .
+ ,
] t s.o --
2 gg; 12.0 - .
.J 11,o- , .. , ,
a ,0.o. . . e .o - : . a S.o - . . . 7.o - g 3 ,. ; - , . s.o - . 3g . .
, + + .
s.o - CO 99 . 4.o -
- 00. .o., , , ,
s.o - .i. 2.0 - ,
' ).o- ,
e.e . . . . . . . . . . . . O 1 2 3 4 6 6 7 8 9 10 11- 12 g. CYCLE EXPOSURE (GWD/MTU) I I
, I E
=
em' as as aum :em :amae.-as:- em , :mme r sus m :em ::ess: !am :ami en_
. amn': mam= am -
~ = '~
FIGURE 2.10-7 - f~ SUSOUEHANNA SES. UNIT 1 CYCLE 1-
- - HOT CALCULATED K-EFFECTIVES 1.01
. + . :.. . . . . . . . . ..
. . . . .; . .. ... . :L . . . . . + . . .i .... . -
Legend . .. .,. . . . m . . .. . . _ O SIMULATE-E , ,, [ , _, , ,_, _ , , ,; . , , , , , A POWERPLEX (CPM) 1.00- - = - - - - - - ~ * -
.g
. s.
. w _ . =
2 N u. 0.99- - (D :iC g,
. - ..m . . . . . . . . . - . . . . L V . . . .
-e e . w . . = = . . e - . . . Ye4m 0.98- - - - - - - - - -
. . . . . . 2. . .. . . .: .
M . . . 9 . p as . W . O.97 . . . . . I 0 1 2. ~3 4 5. 6 7 8. ~9 10 .11 ' 12 - 13 14- 16 CYCLE EXPOSURE (GWD/MTU)?
_ _ e : FIGURE 2.10-8 J . SUSQUEHANNA?SES UNIT 1 CYCLE E HOT CALCULATED K-EFFECTIVES-1.016 . . . .
,._.x , . . _ . . . .
1.010 - - -
.g
. . Legend . . . . . .
~~ "~ ~
- ~ " ~ ' - - ~~ ~ ~
~.
O SIMULATE-E . . . . . A POWERPLEX (CPM) . . . . . . .
- =- - - - -
1.005- - - - LU g . . .
.2, . . . ..
- F-- . . .
8 1.000- - -
, g . _ ,
.2.
tu . . U . . 0.996- - L- - - -" =- - - - 0.990-
. . . ~. . .
0.985 , 0 1 2 3 4 6 .0 ~7 8 9 10 11 - 12-
. CYCLE EXPOSURE (GWD/MTU) .
c -n= se em aus an as aus: e r aus uma sus me : ms: san as ama aus semi uma
~ :c i,;O1 ..
1 FIGURE 2.10-9
- SUSQUEHANNA18ES UNIT 1 CYCLE 3:
HOT CALCULATED KaEFFECTIVES . 1.016 .
. . . . .:. . . . a. .;... . . .. . . .-
1.010 - - -
+-- - =- - - - - -- - - ' = - -
~ '
. .a . Legend: . .: . . ..
- ~~' - -
e O SIMULATE-E .'".
. . . . . . . . . . . . . . . = . . . . .
1.006- -- - --
. P . . . . . .2. .
O -
" uJ 1.000- - -- - -
u g . . . . . . .. ._. . . . .. . . . .
~
[ [ .
..,~ . . .
- O.
~
~""~
~: .. ._ :
. _ . t . . . . .
0.990- - - 1 0.985 , , , , . , , , , , , 0 1 '2 3 -4 '6 . 6 :. 7 8: 9 10. 11: 12-CYCLE EXPOSURE (GWD/MTU)
+ _ _
^
. . [A '
- L; .; ..
FIGURE 2.10-10 y ~
- SUSQUEHANNA"SES. UNIT 2 CYCLE 1 -
HOT CALCULATED K-EFFECTIVES 1.01
. Legend- . .
..;... . . ~ . .
. O SIMULATE-E . . . . . . _ . ...
A POWERPLEX (CPMi . . . .. . . . .- .. . . . 1.00- - - - p- o - - - - . - hg% id It g e bd o...-
. Wa. .
ofo 7
. . . . .. . . .;. 2.
. ~
0.98- -
* * . <* .4, , . e . p.
0.97' i . . 2 . i i. . . . i i 0 1 2 3 4 5 6 7 8 9 10 11 12 13 14 16 CYCLE EXPOSURE (GWD/MTU) aus aus aus . ass ss ass ses mus ms. sua sus aus ' ass sus .mus sus use: sus ass- -
= . _
sus aus aus
~ sum sum aus am~ as ame sum = an as as amm see tamm amm aus amm
~ .+.
~ ~
FIGURE 2.10-11 SUSQUEHANNAiSES: UNIT 2. CYCLE 2: -m HOT CALCULATED K-EFFECTIVES!
'1.0 16 1.010 - -
- -- - - - - - - -- ~ -
O SIMULATE-E - A POWERPLEX (CPM) ,
,~
w 1.006- - - - - - - - - --
~
q) ..
$ W g . . .
e m w .. :. . . . . .j . . . . . .. M 1.000- - --
.i. .
6
- 1 '- -- J -"- - -- -
0 995- ' - - - 0.990 , , . , , , ,. , , , . 0 1 '2 3 :4 6 . 6. 7- 8 9: '10-11 .12 CYCLE EXPOSURE (GWD/MTU) A -
+ ,~ : a' .,t ' ! r tli t}, :! \1{;ff{. j' ! ,>
..y _ . . .
, p'-
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= - '. _ '. d a
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- = L P e ,6 X GN U E =
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, 7' Ea e
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I 2,11 Storace Criticality Comoliance
' Fuel storage. criticality analyses nave been performed by ANF for both the new fuel vault'(Reference 26) and the spent fuel pool (Reference 27).
I Soth of4 these analyses a'e used to assure that the vault or pool k-effective is less than 0.95 under normal canditions and less than 0.98 lg - under all credible accident scenarios. All-appropriate calculational and 3 geometric encertainties have been accounted for in these analyses.. Continued compliance with these analyses is performed for each new fuel
' design as part of the reload licensing analysis. For both the new fuel l vault and'the spent fuel pool, the ANF-analyses remain valid provided the reload fuel is an ANF 9x9-2 design with a lattice nominal enrichment plus enrichment tolerance less than 4.00 wt% U-235. For axially zoned assemblies, this applies to the maximum enrichment zone. Fur the new fuel-vault, an additional criterion is placed on the maximum enrichment zone I
reactivity. The beginning of life k-infinity evaluated using in-core geometry shall not' exceed 1.388 for the maximum reactivity zone'of the reload fuel assembly. This criterion is-verified using the CPM-2 computer program'(Reference _4) using the PP&L analysis methods previously approved by the'NRC (Reference 1). Lll il
<g l:
Io 97
;I?
is ! 3.0 TRANSIENT ANALYSIS METHODS APPLICATIONS The application of PP&L's transient analysis models and methodologies to I
- licensing' analyses is discussed in this section. Reference 2 describes -
PP&L's transient analysis models and methods. The Susquehanna SES system
- and hot bundle RETRAN base models are described in Sections 3.0 and 4 0 of
' Reference 2, respectively. Since the publication of Reference 2, a minor modification ~ was made to the DELTACPR code used to compute Critical Power Ratios (CPRs) and Critical Heat Flux Ratios (CHFRs). The modification implemented a means of automatically iterating on hot bundle power until a minimum CHFR equal to 1.0 is achieved. The thermal limits calculations in the modified code, referred to as the CPRITER code, remain unchanged. An l
outline'of PP&L's RETRAN/CPRITER analysis methodology is provided below g and illustrated in Figure 3.0-1. The RETRAN system model (see Figure 3.0-2 and 3.0-3) simulates the core
~
and system response to an event. The system model calculates normalized g; core power, _ core inlet temperature, upper plenum pressure, and lower 5
- p1_enum pressure as functions of time. These time dependent parameters are input as boundary conditions to the RETRAN hot bundle model. The hot
- bundle mode 1~ output is used by the CPRITER code for thermal margin h
, calculations.
In a manner similar to Advanced Nuclear Fuels Corporation's methodology
- (References 17 and 18), the system model calculated initial core average Axial Power Distribution -(APD) and the calculated normalized power versus g time are used as input to the RETRAN hot bundle model. The all-rods-out B-axial power distributions for MCPR limiting assemblies at end of cycle tend to be more. bottom peaked than the core average. These assemblies are actually affected by scram control rod insertion sooner than the use of h
the core average APD would predict. Therefore using the core average normalized power versus time curve in the hot bundle model is conservative for licensing basis pressurization events. l
~ - - - _ -_ _ -
a i
' The RETRAN hot bundle model calculates the following parameters as functions of time and axial location: heat flux, enthalpy, flow rate, and pressure. The CPRITER code uses these hot bundle calculated parameters to calculate CHFR-versus time and the initial value of CPR. The CPRITER~ code iterates on initial hot bundle power and performs successive RETRAN hot bundle model calculations until the transient minimum CHFR equals 1.0.
The transient ACPR is thea equal to the initial CPR minus 1.0.
'For the events whose results factor into the MCPR operating limits, the parameter of interest is defined as the change in Critical Power Ratio (ACPR) divided by the initial- value of CPR, and is referred to as RCPR.
Results of the sensitivity studies presented in Sections 3.1 to 3.3 are l expressed in terms of RCPR.
-I With the exception of the ASME overpressure analysis described in
'Section 3.5, all analyses in this section are considered as Anticipated Operational Occurrences (A00s) for the purpose of establishing MCPR operating:llmits. The A00s presented in Chapter 15 of the.FSAR (Reference
- 14) were.-evaluated and the potentially limiting events were analyzed. .A subset of the FSAR events was evaluated and determined to be non-limiting as described in Section 3.4.
The three potentially limiting MCPR events which use the RETRAN/CPRITER methodology are: Generator Load Rejection Without Bypass (GLRWOB), I : Feedwater Controller Failure (FWCF), and Recirculation Flow Controller Failure (RFCF). These events will be analyzed for each reload core. The conservative method of analysis, along with sensitivity studies and sample
!g 5_ licensing analyses are presented for these events (Sections 3.1 to 3.3).
The sample _ licensing analyses were performed for Susquehanna SES Unit 2 Cycle 2. The Unit 2 Cycle 2 core consisted of 324 ANF 9x9 fuel assemblies I . gg . I a 1
+ ',
j and'440 GE 8x8R fuel assemblies. As discussed in Reference 2, the RETRAN' I) system model used GE 8x8 fuel pin geometry to represent the core, sinceLit was; the " dominant" (i;e., most prevalent) fuel type.- The hot bundle model represented-an ANF 9x9 fuel bundle for the sample analyses, since this' fuel type isl expected to produce higher calculated ACPRs. An actual ) reload.' licensing analysis would include hot bundle /ACPR analyses for each I potentially limiting fuel type. ' f The RETRAN system model and hot' bundle model gap conductances were
]
generated with ESCORE using the methodology described in Appendix A'. , g L I . I: r I
-l
(.I
- I' s.
s l 0, <
.100 -
{ ; I
. s_
w , ,--, -
"^
e FIGURE 3.0-1 I RETRAN/CPRITER CODE RELATIONSHIPS .l ' INPUT DATA' 5> ; V-
.a RETRAN SYSTEM
]g MODEL
; w
~
i Upper Plenum Pressure vs Time 5 g" a Lower ~ Plenum Pressure vs Time Normalized Core Power vs Time inlet Temperature vs Time ig V RETRAN HOT < Hot Bundle Power
' BUNDLE MODEL
< Gap Conductance 2t
-fg' .- Heat Flux (z,t)-
.g Enthalpy (z,t)
. Flow (z,1)
Pressure (z t)' y
& Hot Bundle Geometry Local Peaking Factors
[ Secondary Local Peaking Factors i g', V
.CHFR vs Time
.g V
.EF CPR (initial)
_. - 101 - a+
<'n-.
I g.:
. FIGURE-3.0-2 L .
LSUSOUEHANNA SES.RETRAN MODEL p : (VESSEL) L v0LtNES b
'JUNC7 IONS HEAT-l; CONOUCTORS SIfD4N AN .
~
. CORE L, . -
d
, HEAT ipi
. CONOUCTORS 6 no, -
, j g--
1 eas g
- QOwho,7 SEPARAfg
'O s g
FEEDWATER 810 = 1,QQdW WI 820 m m - RCIC _830 STAND PIPES
"-0 a060 l' '/ UPPEDPL[NUMN
'N #
h- _- SUB g g f00LE .
, 135 .
$nNl268 AC-215W76@ v
'7 M2 O'
140 a '#2 r l25 9 AC 24Nf75@ W h '24 r,7 i 4 p. "' O DOM' " b.-hk - h@ - S ,,pp pump h3b _a^ in ml Q 2 [ 4 @ AC 2 QlSa
/ <
hhg :
. , A @ AC-t boy @l3, MARGE NG i
i
+ a .' i @ LRR 16Td '
.g' o
~
161 l65 E 215 h h 2'o > g < Q ,
+.
[ LOWER PLENUM Ahha%gg r RECIRCULATION P(pp LOOP S s', ANDSgDNPIPING 4 102 - l
- sus am ass aus em en ams aus ass t .- ; ass -
um uma sus as ese - y II Il ll I
-- P j P 1 P j >
320 330 Jto
- O
- O +o- O O ~'
U .ascitos- . g , , RE C10R - g h star aseuttots
&Plh _ _
b 8 342 5 0 l e 343 250 360 3M U h --a
'h EEwE
- h Y4a'.h .
10 tunerseE 381-B WASS VALVE h# 390 400' 4IO O ->. 4d04- O - O - O - O n-; wmvz l FIGURE 3.0-3 O ~ '" SUSOUEHANNA SES RETRAN MODEl (STEAM LINE--AND BYPASS)
L 3.1: -Generator load Rehetion without Byoass C o 3.1.1; Event Descriot_1.QD The Generator Load Rejection Without Bypass (GLRWOB) event is expected to
' be one of the limiting pressurization events in establishing the MCPR g operating limits. W 1 For the Susquehanna SES units, there are two events.that are considered. 'I One is initiated by a power load imbalance signal which initiates fast.
closure of the Turbine Control Valves (TCVs). ~ The second event can be -l caused by.a number of other generator conditions and actuates the: I generator primary lockout relay, thus causing closure of both the TCVs and ! h Turbine:Stop Valves (TSVs). Both of these events were analyzed (Section 3.1.2) and calculated ACPR results were virtually identical, ' since' at'most power levels, the TCVs close more rapidly than the TSVs in the full arc mode of pressure regulation. As changes are made to the mode
=of TCV: operation, this conclusion will be reviewed for accuracy. >
The GLRWOB produces a fast closure signal to the TCVs. Closure of the TCVs causes a rapid increase in reactor pressure. A reactor scram and an End of Cycle-Recirculation Pump Trip (E0C-RPT) occur.as a result of low TCV oil pressure. Control rods begin moving into the core and the = recirculation pumps begin coast.ing down, thus reducing core flow. The , H system pressurization causes a rapid decrease in core void fraction and a - subsequent < void reactivity induced power increase. The increase in core power causes a degradation in thermal margin (i.e., a reduction in CPR).
;The event is typically analyzed at end of cycle, since the core is unrodded, and the void reactivity coefficient is more negative. h The Susquehanna SES Technical Specifications allow operation with the lE E0C-RPT inoperable. Analyses of the GLRWOB to produce MCPR operating limits for this mode of operation will be performed using the methodology l- - 104 -
l I i \. i -
r y , q l
)
described _in this section with:the exception that the E0C-RPT will not be credited. 1
-l 3.1.2 Sensitivity Studies I RETRAN/CPRITER analyses were performed to determine the sensitivity of calculated RCPR to changes in input assumptions. The parameters I considered were:
- 1) Control rod insertion rate
- 2) Recirculation Pump Trip (RPT) l
'3) Turbine Control Valve (TCV) cloture time
- 4) Safety Relief Valve (SRV) opening setpoints and flow capacity
- 'Feedwater temperature 5)
Pressure regulator setpoint 6)~
- 7) Core power-
- 8) Core flow
- 9) Steam line inertia 10)_ Steam line volume 11)- Steam..line pressure drop In addition, two extra analyses were performed: 1) a Turbine Trip 'l Without Bypass (TTWOB) which, for the Susquehanna_SES Units, closes both I the TSVs and the TCVs, and 2) a Gl.RWOB which occurs due to a power load ,
I imbalance signal i'd closes the TCVs only. , The base case values of the above mentioned parameters are listed in Table 3 .1 - l~ . Fast closure of the TCV in the RETRAN model is accomplished by a manual trip on elapsed time. Results of the study are shown in L. - Table 3.1-2. Conclusions an'd observations from these analyses are L. discussed below.
- 105 -
I; l l.~
4 o
.1 A faster control rod insertion (scram curve) clearly produces a y significant reduction in peak core power and, hence, RCPR. A reduction in
- /* _RPT delay results in a faster reduction in core flow, which produces a reduction in peak core power and RCPR. Use of the Technical Specification maximum RPT delay (currently 0.175 seconds) is, thus, conservative. A faster TCV closure produces a more severe pressurization and results in a higher calculated RCPR.
For the Susquehanna SES units, there are two modes of operation for the Safety Relief Valves (SRVs), referred to as the safety mode and the relief mode. In the safety mode, the valves are opened mechanically when system pressure overcomes the force of a spring. The number of operable valves la j required and the pressure setpoints are specified in the Susquehanna SES of units' Technical Specifications .(Refmcer 15 and 16). In the relief mode, the valve opening is initiated by electrical signals when measured dome pressure exceeds the specified setpoints. The relief mode pressure , setpoints are lower than the safety mode pressure setpoints. [ The difference in assumed SRV operation between the Technical Specification safety mode (6 valves out of service) and the nominal relief mode-(all valves operable) produces only a small change in RCPR. However, h' ' the Technical Specification safety mode assumption produced a more severe - pressurization,. and gave a slightly higher calculated RCPR. The effect of decreased feedwater temperature on the GLRWOB is an increase It L in-peak core power due to a more top peaked Axial Power Distribution (APD). However, the effect on calculated RCPR is negligible. An increase _
-)
in pressure regt.lator setpoint, whick ncreases dome pressure [ approximately the same amount, re*.ss in a decrease in calculated RCPR. ~ The effects of decreased feedwater temperat'ure and increased pressure
.{
regulator setpoint on peak power and calculated RCPR are complex. Both n cases show slightly more top peaked APDs than the base case, which
. accounts for the slight increase in peak core power. However, the changes
- 106 -
e amas
I in hot bundle enthalpies produced by the. pressure, flow, and heat flux transients are difficult to predict for the different cases. In any case,- < for changes of 10*F in feedwater temperature and 32 ' psi in pressure l regulator setpoint, the impact on RCPR is very small, as shown in Table ! 3.1-2.
.g Increased initial core power and increased initial core flow produce a 5 higher calculated RCPR. At increased power, the initial steam flow and, hence, vessel pressurization rates are higher. At increased flow, the -
core APD is more top peaked causing a significantly higher peak core power and, hence, a larger calculated RCPR. The steam line parameters. (inertia, volume, and pressure drop) all affect
; the pressure wave in the steam line. Higher inrrtia creates a higher pressure peak at the TCV.and a higher pressurization rate in the steam dome. This adverse pressurization produces a higher and broader power
!'I; versus time curve, which produces a larger calculated RCPR. A larger steam line volume reduces the vessel pressurization and RCPR, since more mass and energy are required to pressurize the larger volume. A larger steam line pressure drop means a higher resistance to flow (i.e., larger losscoefficients). The larger loss coefficients will attenuate the - pressure wave more as.it travels from TCVs to the vessel, producing a lower peak power and a lower RCPR. r 3.1.3 Licensina Analysis Method L' - - The GLRWOB analysis method-is based on the results of the sensitivity , studies discussed in the previous section. The analysis approach may use yg the Statistical. Combination of Uncertainties (SCU) methodology described
- LE in Appendix B. Unless treated statistically with the SCU methodology, the following input assumptions are used:
107 -
-- F'"'-
j
- 1) Technical / Specification scram curve.
- 2) -Technical Specification (maximum) scram and EOC-RPT delay times. _
3)- Technical Specification safety mode forf the SRVs -(6 valves inoperable).
- 4) Analysis is performed at end of cycle /all-rods-out. If cycle exposure depende'nt thermal limits are to be specified, additional _ cycle exoosure conditions will be analyzed.
- 5) Analysis is performed at the most limiting condition from 0%
to 104.4% of rated core power and 100% core flow. l 6). Code uncertainty is applied as specified in Section 8 of { Reference 2.
- 7) Use nominal feedwater temperature.
I The uncertainties on steam line AP, steam line inertia, steam
.8) line volume, TCV closure time, and pressure regulator setpoint are _ statistically combined with the Raference 2 code {
uncertainty.- Results of an analysis for Unit 2 Cycle 2 using the above assumptions,
- with the-exception that initialLeore-power was 100%, are presented in Figures 3.1-1 to 3.1-7. ,
3.1.3.1 Treatment of Steam Line Uncertainties-l , The RETRAN code uncertainty derived in Reference 2 excludes the steam line i
' parameter uncertainties. 'The steam line' parameters of importance are: .
L - 108 I , _;f s s - - ,-
s .
- 1) Steam line AP
- 2) Steam line inertia
- 3) Steam line volume 4). TCV closure time
-9 5 ) -- Pressure regulator setpoint J These uncertainties were combined with the Reference 2 code uncertainty to i produce a total model uncertainty by the root sum square method. Table 3.1-5 illustrates the calculation. The total model uncertainty was used in the SCU analyses described in this Section, it is clear from Table { 3.1-5 that the steam line parameter uncertainties, when combined with the
.. code uncertainty generated in Reference 2, will- have only a small effect on calculated RCPR. The larger uncertainty, which includes the steam line uncertainties, will be used in the analysis.
3.1.3.2 SCU Method
= - As'previously stated, several of the assumptions listed above may be modified if a Statistical Combination of Uncertainties (SCU) approach is used. The-SCU methodology described in Section B.2 consists of the following steps:
- 1) Create a response surface from RETRAN/CPRITER calculations that relates calculated RCPR to the variables to be analyzed statistically. The variables chosen for the sample Gl.RWOB are core power and scram insertion speed.
- 2) Define uncertainty distributions for the parameters used in the response surface.
- 109 -
1 1
+ ;
11 3)' Perform Monte Carlo analysis to produce a cumulative
~
probability distribution of. calculated RCPR. h
'4) _ Perform safety limit type analyses to produce cumulative probability distribution functions of fraction of pins in boiling tra'isition for a range of MCPRs. For the sample -
calculation presented in Section 3.1.4, cumulative probability distribution functions were generated by the ANF methodology , for a core co .aining all ANF 9x9 fuel. These distributions are used since they are typical for future licensing applications.
- 5) Assume a MCPR operating limit. ,
6)- Perform Monte Carlo analysis to combine the cumulative , probability distribution functions for fraction of pins in boiling transition and the cumulative probability distribution 'g function of transient RCPR, thus producing a combined safety E limit and transient analysis. .
- 7) Select the value of number of rods expected to be in boiling transition at-the 95% confidence'. level. -If the value is ;
greater than 0.1% of the pins, increase the assumed MCPR operating limit and repeat steps 5, 6, and 7. : 3 .' l . 4 Samole licensina Analysis To demonstrate the application of the SCU methodology for the GLRWOB, a -
~
; sample analysis was performed for Susquehanna SES Unit 2 Cycle 2. This is -
the method PP&L plans to use for reload licensing applications. The- , - variables- treated statistically were core power and control rod scram insertion speed. L ,
- 110 -
I l '
y A response surface of calculated-RCPR as a function of core power and average scram speed.was produced with CPRITER. The choice of scram speed was based on statistical analyses of over 3800 individual Susquehanna SES control rod scram time measurements. These analyses demonstrated that average scram speed can be treated as a normally distributed variable. The nine RETRAN cases run to produce the response surface'are shown in Table 3.1-3. The response surface-is of the form:
] RCPR = A, + A,X + A2 X, + A12Xt X, + A33X* + A22 X,*,-
3 yhtrg: Xi- change in core power from 100% power (% rated) X, - change in scram speed from 4.167.ft/sec (ft/sec) The coefficients of the response surface for the U2C2 sample GLRWOB
; analysis are given in Table 3.1-4.
The mean and standard deviation of the core power error were assumed to be 0.0 and 2A respectively. The mean and standard deviation of average scram speed were chosen, based on a conservative evaluation of plant data, to be 4.167-ft/second and 0.2 ft/second, respectively. The core power uncertainty is common to both the-safety limit type and transient RCPR analyses for the GLRWOB. A higher core power adversely affects both' analyses (i.e., higher number of rods in boiling transition and larger calculated RCPR). As stated in Section B.2.3, Item 7, the
]
_ , safety limit type and RCPR analyses are not strictly statistically independent, and an additional conservatism is required. The maximum variation. in calculated RCPR over the core power range of 96% to.104.4% was only .007. Thus, as an additional conservatism to cover the 111 - 1
^ ^ ' ' ' - ^ ~ ^ ^
p<>O, y a. . n y.
~
f4[ , i
- assumption of statistical independence of the safety limit and RCPR-i '
3_ ; analyses, a value of .007 was added to the transient calculated RCPRs, and _ the operating limit was recalculated. Another valid-approach would be to 7
- q. select the core power level which produces the maximum calculated RCPR l
'within two standard deviations (96-104% power,) and use that power level for all response surface analyses (i.e., conservatively treat power uncertainty and exclude it from the response surface). .~
The calculated MCPR operating limit for the sample U2C2 GLRWOB, using the .i SCU methodology described in Appendix B was 1,30. It should be noted that i if other events require higher operating limits, then the actual Technical-ff y Specification MCPR operating limits would be-the highest value obtain0d.
, for any of the events.
' I ~'
a g L g1 I l I. i
.m. I I'
I
J 3 p TABLE 3.1-1 ]
'~
Generator Load Rejection Without Bypass : Base Case Innut Assumntions Parameter Assumptien l 1. Control Rod Insertion Technical Specification insertion rate (slowest allowed)
- 2. Recirculation Pump Trip Technical Specification value (maximum d 1 Delay Time (sec.) allowed)
- 3. Turbine Control Valve 85% of the best estimate value based on .l Closure Time (sec.) plant data '
- 4. . Safety / Relief Valve Technical Specification required safety Operation valve mode of SRVs with six SRVs out of service H 5.. Feedwater Temperature Best estimate val'ue based on plant data i
- 6. Pressure Regulator Setpoint Best_ estimate value based on plant data-
- 7. Core Power 3293 Mwt.(100%)
- 8. Core F_ low 100 M1b/hr
- 9. Steam Line Inertia- Best estimate values (Reference'2)
- 10. Steam Line Volume Best estimate values (Reference 2)
- 11. Steam Line Pressure Drop. Best estimate values (Reference 2)
I I
- 113 -
Y
1 I l TABLE 3,1-2 : Generator Load Rejection Without Bypass Results of Sensitivity Analyses 1 Change in Peak Core Power Change in l' Parameter Chance from base value (% rated) RCPR
. l. Control Rod Best estimate curve -71.3
.080
-Insertion based on plant. data >
- 2. RPT Delay (sec.). .05 seconds -25.0 .008
- 3. TCV Closure Time +.01 seconds -2.9 .002 I (sec.) -l y
- 4. SRV Operation Relief mode /all valves 0.0- .002 E operable 3,
- 5. Feedwater -10 *F +10.1- +.000 Temperature
- 6. ' Pressure Regulator +32 psi +2.3 .007 i Setpoint -
'7 Core Porer +4.4% (104.4/100) -
+ 002
-35% (65/100)- -
' 024 ,
- 8. Core Flow -1.,% (100/87) -45.2 .027
- 9. Steam Line Inertia +26% +64.8 +.015 g ,
- '10.~ Steam Line' Volume +5% -2.6 .004
,' 11. Steam Line Pressure- +29.5 psi -58.6 3 28 m
. Drop
- 12. TSV Closure-(yes/no) no +14.4 +.001 l> 13. Event' TTWOB -1.0 .002 *I y -
- 114 -
1 g 4
~
- 5 H
TABLE 3.1-3 I Cases Analyzed for GLRWOB Response Surface l l Scram Speed (ft/sec) Core Power Error 4.902 id52 L_Q11
+4.4% X X X I +0.0% X X X 1
-4.0% X X X Mll: X - means case was used in response surface generation.
I i
- 115 -
1 a s L \.
. TABLE 3.1-4 .!
Coefficients for GLRWOB. L Response Surface l-j '. Ao = +.174109
, Ag =
.0005875 A2 = .066197 A12 - .0002802 -
Ai , = +.0002237 L A22 - .0079919 l 1 l l l Response Surface Standard Deviation = .0002 E/, p 3 l II , I' l . I !
.i i - 116 -
1 l 1,
F 1 gg I TABLE 3.1-5 I Steam Line Parameters: Contribution to Code Uncertainty for GLRWOB B Parameter Uncertainties") dRCPR/dx(3) ' )," Steam Line AP(psi) 5 .0009/ psi .0045 Steam Line Inertia (%) 5 +.0006/% .0030 Steam Line Volurte (%) 5 .0008/%. .0040 Turbine Control Valve 10 .0002/% .0020 Closure Time
.(% of best estimate)
Pressure Regulator 5 .0002/ psi .0010
. Setpoint (psia)
I
'I - (1) S,; based on an engineering- assessment of the parameters.
I R
- (2)c. S,CPR = S, dx (3)' Steam Line Parameters' Cotribution t; Code Uncertainty (S".'"." mum)
R RCPR 2 ! ug 3.CPR
.. n. -
- s ' -
117 - l l
I Ill FIGURE 3.1-1 : GLRWOB COREPOWER
.Ill 400
.n . .
50 ........ . 300 . . . . . . . . . .. .... .. . ..... . ........... . . .. pas .
~o,. .
w . . s . x 0 ....
.u.
O . s., . . . . . 00 ... ... x . w . O. x
-w.
x . O .
~ . . . . . e 10 0 ..
5n........... o . i ! i i RBRAN 0 o i i i i i i 0 0.5 1 1.5 2 2.5 3 3.5 4
. IlME (SEC)us - -
l
n ?L 5
'I l^
-- i I.}, f
-, s i
- r. r, O
J: - : ,:; RGURE 3.1-2 : GLRWOB CORE R0W; i
.L I s
I
. ,/
/ 12 0
- - e. r . . . .
zu . ~ .
.c .
...........s..... . .. ......... .. .. ... .......... . .... ..
~ . . . . .
&v < .
4......... .>......... 300
- n. .
- w. . .
r . 30 ..............................
........3.....................3..........
y .. 4o. . . . . . . . . . . .......... .. ............ .. ... .................. : ........
. 4........ .
O .
. e . . .
. 3 .
0- 0.5 1 1.5 2 2.5 3 3.5 4
' 11ME(SEC)
- 119 -
I t . f
.. I t;
- RGUE 3.1 : GLRWOB DOME PRESSURE
?
' e I
1500 . t
+ .
1250 ........... . .. . .. ....
. lj .......... l
; i .
-A
- v. . . .
i
= . m .
c t - . o~p ,. ,. +.
+
+. .
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, v,, .
j j j j j j egend a
! i : i i i RETRM -
L 1000 i i , i i i i . Um L 0 0.5 1 1.5 2 2.5 3 3.5 4 i' TIME (SEC)
- 120 - g
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i k [l t FIGURE 3.1-4 GLRWOB VESSEi. STEAM R0W J
.l l
o, 4 I 30 ...... E
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-a - 121 -
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m ;.. . i I N RGURE 3.h5 GLRWOB NARROW RANGE EVEL 4 q I1
- 40 I
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. RGURE3.1-6 GLRWOB FEEDWATERR.0W c I' l
- ; 100 .
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- 123 -
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. , ' . . . ..~ , ,_ , - - , - . . - - . , - . , , - y, ,
I I 1 a 3 ~. 2 Feedwater Controller Failure 3.2.1 Event Description 1 i The Feedwater Controller Failure (FWCF) event is expected to be one of ths limiting events in establishing the MCPR operating limits. As , I demonstrated by the analyses presented in Section 3.2.2, the event produces larger calculated ACPRs at lower power levels. ' The FWCF event is assumed to result in an increase in the feedwater demand , signal 'to its maximum value. The failure causes the feedwater control system to increase feedwater flow to its maximum flow rate.. This produces a cooldown of the middle and lower downcomer fluid and a reduction in core inlet temperature. The decrease in core inlet temperature causes a slow increase in core power due to a reduction in the core void fraction. The I water level in the downcomer increases'until a turbine trip signal is generated on high narrow range vessel level. As~ a result of the turbine trip signal, both the TSVs and TCVs begin to. close, causing-an increase in reactor pressure. The rapidly increasing pressure causes a rapid core power increase due to void reactivity a feedback. Closure of the TSVs and TCVs generates reactor trip and E0C-RPT signals. The turbine trip produces a fast opening signal for the main-I turbine bypass valves. Thus, the portion of the FWCF following turbine trip,is'similar to the Generator Load Rejection event presented in Section-
; 3.1,'except that the pressurization is mitigated by steam relief through the main turbine bypass valves. After a small power decrease due to initial' scram rod insertion, the rapid pressurization produces a core ! power increase due to positive void reactivity feedback. The increase in core power causes a degradation in thermal margin (i.e., a reduction in
.I - 125 -
i j . . O; s CPR). The event is typically analyzed at end of cycle, since the core is ' unrodded-(i.e., worst case scram response) and the void coefficient is at L its most negative value at this point in the cycle.
'The Susquehanna SES Technical Specifications also allow operation with E1 either the turbine bypass system or the EOC-RPT inoperable. Analyses of the FWCF to generate MCPR operating limits for these modes of operation will be performed using the methodology described in this section, with ,
the exception that either the bypass system or E0C-RPT will be assumed to-be inoperable, respectively. 3.2.2 lensitivity Studies
.I ,
RETRAN/CPRITER analyses were performed to determine th: sensitivity of calculated RCPR to changes in input assumptions for the FWCF. The - parameters' analyzed were:
-1) Control rod insertion rate ,
- 2) Recirculation Pump Trip (RPT) delay time
- 3) --Turbine Control Valve-(TCV) closure time f;,
- 4) Safety Relief Valve (SRV) pressure setpoints and flow capacity 1 5)- Feedwater temperature
- 6) Pressure regulator setpoint l 7) Core power
' 8) Core flow -
- 9) Steam line inertia g
- 10) Steam line volume 5
- 11) Steam line~ pressure drop
- 12) Bypass valve capacity ,
- 13) Turbine trip high level setpoint
- 14) Maximum feedwater flow ,
- 126 -
I- 1 I, t
I The base case values. for the sensitivity analyses are listed in L
. Table 3.2-1. The FWCF is initiated in the RETRAN model by a manual trip on elapsed time. Results of the sensitivity analyses' are shown in Table 3.2-2. Conclusions and observations from these analyses are discussed below.
A faster control rod insertion (scram curve) clearly produces a
,I ,
significant reduction in peak core power and, hence, RCPR. A reduction in
=
RPT delay results in a faster reduction in core flow, which produces a reduction.in peak core power and RCPR. Thus, use of the Technical l Specification maximum RPT delay (currently 0.175 seconds) is conservative. A. faster TCV closure produces a more severe vessel pressurization and results in a higher calculated RCPR. I The effects of using the relief mode of the SRVs (as opposed to the %fety I mode) are complex, because the pressure and core flow transients are different. A separate set of cases was run without bypass valve operation. The effect of assuming the relief valve mode of SRV operation
-for the case without bypass operation was to~make the event less severe'as it did for the GLRWOB event (see Section 3.1.2). Analysis results show that use of the relief mode for the FWCF (bypass available) and safety mode ~for the FWCF (bypass unavailable) is conservative.
A lower _ initial feedwater temperature causes a higher power rise before . turbine trip. Although the effect is small, the lower initial feedwater temperature case produces a slightly-higher peak power and calculated I RCPR. An increase in pressure regulator setpoint, which raises dome pressure approximately the same amount, produces a small increase in RCPR.
- l. The effects of decreased feedwater temperature and increased pressure regulator setpoint on peak power.and calculated RCPR are complex. Both cases show slightly more top peaked APDs than the base case, which h - 127 -
)U tlj
I! L accounts-for the' slight increase in peak core power. However, the I: relative effects of changes in hot bundle enthalples produced by the different pressure, flow, and heat flux transients are difficult to evaluate. The impact on RCPR is very small, as shown in Table 3.2-2 for i changes of 10'F and 32 psi in feedwater temperature and pressure regulator l setpoint, respectively. E W.. The FWCF event is more adverse (i.e., higher calculated RCPR) at lower
. initial core power. This is due to the fact that the initial feedwater g
5 flow -is lower. -Thus, an increase of feedwater flow to its maximum value has a larger effect on tht core inlet temperature reduction and core power - increase prior'to turbine trip. Therefore, licensing analyses must consider the FWCF at different power levels. At 100% core flow' compared to 87% core flow, with- all other factors the same, the core axial power
- distribution is more top peaked. The more top peaked axial power '
distribution produces a higher peak core power due to scraia effects and, hence, a higher calculated RCPR, g u! , The steam line parameters (inertia, volume, and pressure drop) all affect
'the pressure wave in the steam line. Higher inertia creates a higher L
pressure peak at the TCV and a higher pressurization rate in the steam f I dome. This adverse pressurization produces a higher and broader power versus time curve, which produces a larger calculated RCPR. A larger ' steam-line volume reduces the vessel pressurization and RCPR,_since more l mass and energy are required to pressurize ;he larger volume. A larger steam _line pressure drop means a higher resistance to flow (i.e., larger- g L loss coefficients). The larger loss coefficients will attenuate the W L pressure wave more as it travels from the TCVs to the vessel, producing a lower peak power and a lower RCPR. The capacity of the bypass valves has a significant effect on the l pressurization portion of the FWCF and, hence, the peak core power and RCPR. A higher turbine trip. water level setpoint allows a longer cooldown
- 128 -
I I
4 > LI: g and power increase prior to turbine trip and produces a higher calculated RCPR.- The FWCF to maximum demand is more adverse than the FWCF to a. smaller /than maximum demand, since the FWCF to maximum demand producas a more severe cooldown:and power rise prior to turbine trip. 1 3.2.3' Licensina Analysis Method i The FWCF event will be analyzed at'100% core flow and various power
- levels, since calculated RCPR Increases as initial power level decreases.
Additional assumptions made in the analysis are based on the sensitivity 1 analyses described in:Section 3.2.2. These assumptions include: 1)- Technical Specification scram curve, s I 2) Technical Specification maximum scram and EOC-RPT delays. , 3)' Conservative bypa'ss valve opening times based on startup test acceptance criteria.
- 4) High water level turbine trip setpoint is equal to the Technical Specification allowable value, plus accuracy and calibration uncertainty, plus 5.0 inches for initial level variations
(= 63.7 inches). 5). Conservative turbine stop valse closure time (= 0.1 seconds). I 6) Failure is assumed to the maximum flow rate.
- 7) Analysis performed at the maximum allowed cycle exposure with all rods out. If cycle exposure dependent MCPR operating limits L~ li l - 129 -
ol-
le , f; l f Ib are to be specified, ' additional cycle exposure points will be gl l ff wr analyzed, 8)- Code uncertainty is applied to calculated RCPR as_specified in { Section 8 of Reference 2. The Reference 2 ccn uncertainty is
-]
[ increased to include steam line uncertainties using the root sum m-g ,; ji ' square technique in the same manner that was used for the M GLRWOB. The uncertainty in feedwater temperature was included ! "'~ k along with th: tcam line parameter. uncertainties, and Table 3.2-3 illum rates the calculation. The effect of the steam line uncertainties on the FWCF event is less than their effect on the - GLRWOB, as can be seen by comparing Tables 3.1-5 and 3.2-3. [ > This is expected since the pressurization effect for the FWCF is - less severe than it is for the GLRWOB, due to operation of the ( turbine bypass valves. Also, the effect of these uncertainties' . on the total code uncertainty is small. ! The sample licensing _ event presented in-Section 3.2.4 uses the above m f assumptions. However, the Statistical Combination of Uncertainties (SCU) it J : methodology described in Appendix' B may also be used for the FWCF event. ; I' . 3.2.4 Samole Licensina Analysis
- This section presents a sample licensing analysis for a Unit 2 Cycle 2 ,
FWCF at 100% power /100% flow using the conservative assumptions described h, in Section 3.2.3. The transient behavior of key system parameters is g1 1
, -presented in Figures 3.2-1 through 3.2-8. As shown'in Figure 3.2-6, the 5, K , FWCF causes a rapid increase in feedwater flow. The immediata effect of-4 the flow increase is to increase the downcomer level and elevation head
- which produces an increase in core flow at about 2 seconds. As shown in -,
Figure 3.2-8, the core inlet enthalpy begins to decrease rapidly around 6 seconds into the event, which causes a rapid power increase. A second l
" wave" of cold water comes by way of the recirculation loop piping. After
- 130 - '
I'
~
~ .
a g transitting the recirculation loops, additional colder water enters the
- jet pumps as drive flow at around 21 seconds. At approximately 25
[ seconds,.the turbine trip causes a scram and rapid pressurization as shown in Figure 3.2-1. Due to the operation of'the bypass valves, the resulting-power peak is significantly less severe than.the GLRWOB described in - i Section 3.1. The calculated ACPR for the event, including RETRAN code uncertainty is 0.20.
'f -
I
'E:
8 I! g I. I I I; I - 131 - I 3
,.,; s l,,+, -
~
l '
, TABLE 3.2-1 11 '
Feedwater Controller Failure
- l Base Case Inout Assumptioni Parameter Assumption I" .
- 1. Control Rod Insertion Technical Specification minimum .
insertion rate
- 2. Recirculation Pump Teip' Delay Technical Specification value ;
Time (sec)
- 3. Turbine Control Valve Closure
. Time (sec)
Best estimate value based on plant data f
- 4. Safety / Relief Valve Operation Technical Specification required safety valve mode of SRVs with six SRVs out of service. i
- 5. Feedwater Temperature Best estimate value' based on plant ;
data ' h
- 6. Pressure Regulator Setpoint Best estimate value based on plant ,
data
~7. Core Power- 100%
-8. Core Flow 100% -'
-9. ' Steam Line Inertia Bestestimatevalues(Reference 2) '
- 10. Steam Line Volume Bestestimate' values-(Reference 2) 1 11; Steam Line Pressure Drop
- 12. . Bypass Valves Capacity Best estimate values (Reference ~2) Ii e
. Best estimate values based on plant measurements 5
r.'
- 13. Turbine Trip Level Setpoint Technical Specification allowable, e o value, plus accuracy and g; calibration uncertainty plus 5 ic.ches for initial level variations (= 63.7 inches) lu
- 14. Maximum-Feedwater Flow 135% rated P - 132 - l I'
E u.
7 'n ~ TABLE 3.2-2 E Feedwater controller Failure Results of Sensitivity Analyses Change in Peak {' Parameter Ghance from Base Value Core Power Change in (% ratedL RCPR
- 1. Control Rod Best estimate curve -31.9 .053 Insertion based on plant data
- 2. RPT Delay (sec) .025 seconds -5.2 .003
- 3. TCV Closure. . 008 seconds ( 10%) +2.8 +.002 T'.me (sec)
- 4. SRV Operation Relief mode /all valves 0.0 +.006 operable
- 5. Feedwtter -10'F +13.4 +.003 Temperature
- 6. Pressure Regulator +32 pst +7.9 +.003
, Setpoint
- 7. Core Power -35% (65% power /100% flow) -
+.072
- 8. Core Flow -13% (100% power /87% flow) -25.1 .031
- 9. Steam Line Inertia +25% +12.0 +.00d
-{
- 10. Steam Line Volume +5% -3.2 .003
- 11. Steam Line Pressure +29.5 psi -22.8 .015
. Drop
- 12. Bypass Capacity -10% +24.7 +.016 100% +232.1 + 120
- 13. Turbine Trip Level -5 inches -5.6 .006 Setpoint
! 14. Maximum Feedwater -10% rated -5.7 .004
- 133 -
I I: TABLE 3.2-3 Steam Line Parameters: Contribution to code Uncertainty ) for FWCF Il ! Parameter Uncertainties") dRCPR/dx(8) [,"_. Steam Line AP(psi) 5 .0005/ psi .0025 h' , Steam Line Inertia (%) 5 +.00032/% .0016 l Steam Line Volume _(%) 5 .0006/% .0030 Turbine Control Valve 10 .0002/% .0020 : Closure Time (% of best estimate) Pressure Regulator 5 +.0001/ psi .0005 Setpoint (psia) , Feedwater Temperature ('F) .0003 .0015 l 5 g I; (1) S,; based on an engineering assessment of the parameters, t S"" dE2.8 dx (2) ,
- S, 3 (3) Steam Line Parameters' Contribution to Code Uncertainty ( S",[",g i, )
S[",t i, - [(S"C)2
, = .005
- 134 -
I I L g
?l4 o{, flGURE 3.2-1 : FWCf COE POWER ! l ug L 200 i E . E .
. . . i 0
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. ! RDRM 0 i i , i i i 0 5 10 15 20 25 30 35 l'I IlME(SEC) 235 -
-we g-eywv-
_ . . _ - _ _ . _ _ ~ _ . _ _ . _ _ _ _ _ _ . _ _ _ _ _ E. i , flGURE 3.2-2 : FWCF CORE FLOW lli Y 12 0 I ' no. . . ion . ...
. . . . r i
w 90 ..... ... r . x o so. ... . .,. ...... .. .,.. ...... ., .. .... ...... . , ... . ) . . . [ . y . . n . . . . g 70 ........
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. . 8
. . 8 3o . . , , , ,
0 5 10 15 20 25 30 35 M(SEC) I'
- 136 -
l
il g RGURE3.2-3 : FWCF DOMEPRESSURE i 1250 q . . . . .l i 4 l o, . l 1200
~ .
1 8a . 7 Vi. . . a v 333o. ........... ......... . ........... ..... .... ..... .. . .... ...... . . .... w - e . v a . m .
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- 137 -
n. e - .. -- ,-.,,e , . . . , . . . -
l I!' flGURE 3.2-4 FWCF VESSD. STEAM Fl.0W g! i I
-f 12 0 '
300 ........ .:.. ... ... .. ....... ......... . r5 80 . 0 . 3 w F< . . x a . O . v - - 4o. ... .. .,. ........ .,......... .,.. :
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1
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. _8900
- : . RDRAN .
h 1 0 5 10 15 20 25 30 35 TIME (SEC) I.
- 138 - i 1
i i
flGURE 3.2-5 FWCF NARROW RANGEl.EVEl. ss s0 . .......... ...... ..... ... ........ . n . .... .. .. w .
.t U . .
e.
.Z v . ........... . ......... ...... ...
w .
=. W : . : : :
a . w o 2 x 4O. .. .. . ..... .. O . e . t x z 35 ...... 30 ....... ... e9ent
- : i RURM s .
k k k k k k
. 0 5 10 15 20 25 30 35
%SEC)
- 139 - ,
4 I' , RGURE3.2-6 FVCf FEDWATERR0W li 14 0 Ii ,
. . .:........ . + -......;~... .....;. ... . .
.... t 12 0
. .:. .. ... ,.. . ...... t
. . 1
<s O .
i w . . 1 F( j00- t t- - t-x -
- t. .
L 0 . i
. . i. ,
M : - v : - - 80- ii- i- 4- -- W : : - F . . x .
- a 60-k.
!- "" i- ""i
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x .
.w .
F . . 40- ; ; ; ;. ' O . w . . W. . . L. 0 .... .... .
.i : i ! i
_09end 0 i i RETRAN g
, i i ,
0 5 10 15 20 25 30 35 IlME(SEC) I' 140 -
ncuREM-7 FWCf 3 AVERAGEHEATFLUX j 20 ? 11 0 - gi x - 7 ' 4 x g 90-I a g [80-3 I {70- . . I e y60-I e l 50-E \ g e. . 30-i Legend a RURAN 0 5 10 N $0 $5 $0 55 TIME (SEC)
- 141 -
6 I IlCURE 3.2-8 CORE RET ENTHAl.PY l 520 .. I 519-n . m 52- . . t- 4 t-4<<- v .
>- i i <-
- n. SU- : -
i-i s . .
.Z .
W 54- i i i F -l s - i- - w . . s u Z-
. 1 l w .
e 515 3 ...........:....... ...:...........::. .........:.. ..... ..:...........:. n 5 .... ......
.i !, .i .i !
.i
. 09end i
513 i e RE_T.RAN i i i . . 0 5 10 15 20 25 30 35 g TIME (SEC)
- 142 -
{
.; I I 3.3 Recirculation Flow Controller Failure , I 3.3.1 Event Descriotion The Recirculatiun Flow Controller failure (RFCF) event with increasing : flow is expected to be one of the limiting events in establishing the ' l Minimum Critical Power Ratio (MCPR) operating limits at low reactor core
- flow conditions. A RFCF event can result in either an increase or a decrease in core flow. The RFCF event with decreasing flow is not analyzed for each reload, because the power decreases during the event.
The decrease i'n power for the RFCF event with decreasing flow makes it I non-limiting from a MCPR standpoint. The RFCF event is initiated when either the master controller or a single j individual loop controller fails. The failure of the master controller h causes the individual loop controller input demands to increase. This causes both reactor recirculation pumps to increase core flow thereby increasing reactor power. The maximum rate at which the individual l controller inputs can change is limited electronically to prevent rapid
- I core flow increases while still providing adequate operating maneuverability. Failure of an individual loop controller causes the ;
demand input to the MG-set scoop tube positioner to increase. The rate of I -increase of the positioner is limited by the physical size of the system. Since there is no mechanical or electrical limiting of the positioner
- demand input, an individual controller failure can cause a faster recirculation loop flow increase than a master controller failure. In l either case, the events are either terminated by a reactor scram on high neutron flux or when a new steady state is reached.
I LI l
- 143 -
l l !I L LI
l ll 8 3.3.2 Sensitivity Studies 3.3.2.1 Initial Studiti .Ii Sensitivity studies were performed to determine the most limiting master controller failure event and individual loop controller failure event. These events werS analyzed from the three different power / flow conditions shown in Table 3.3-1. This range of conditions was chosen to determine the sensitivity of RCPR to core power and flow, i Analyses demonstrated that the RFCF event in which an individual loop controller fails is less limiting than the event in which the master l controller fails. To determine this, the m'sto limiting controller demand increase rate (hereiaafter referred to as run up rate) for both types of g' failures was determined. The RFCF event with an individual loop controller failure is more adverse when the run up rate is high. This is - also true for the master controller failure provided the high neutron flux trip setpoint is not reached. The slow rates are more adverse for the master controller failure if the high flux trip setpoint is reached. Each event is described below to explain this behavior. l The RFCF event in which the individual loop controller fails is terminated g either by the high neutron flux scram ! rip or when the reactor reaches a new steady state, depending on the run-up rate that is used and the rod l line from which the event is initiated. The slow run-up rate case has a ' peak core power and therefore a peak heat flux that is lower than the fast g run-up rate case. Thus, a slower run up rate produces a smaller change in W CPR. This behavior was observed for all analyzed power / flow conditions with or without a reactor scram. l The RFCF event in which the master controller fails either is terminated by the high neutron flux reactor trip or reaches a new steady state, depending on the run-up rate that is used and the rod line from which the
- 144 -
l l E I
.. - _ - _ _ . - - - . . . . ~ . - .---. -- . -.
I I event is initiated. If the event is initiated with a fast run-up rate at a high enough rod line to cause a scram on high neutron flux, then there exists a slower run-up rate which is limiting. A slower run up rate will l be limiting because the heat flux follows power more closely than it does
'g- for a faster run-up rate, thus producing a higher heat flux at the time of scram. The flow is also highs r at the time of scram but this is not enough to compensate for the effect on CPR of the increase in heat flux.
I Therefore, the ACPR is greater for the slower run up rate case. However, the ACPR will decrease if the run up rate is too slow to cause a scram on high neutron flux. Results from a run up rate study are presented in Table 3.3-2. The decrease in RCPR occurs for the slowest run-up case because the run-up rate is not fast enough to offset the effect that the increasing teedwater temperature has on the core power. This causes the g core power to peak below the high neutron flux reactor trip setpoint. A j smaller RCPR is calculated for this case because the power and heat flux . peaks are lower than the peaks calculated for the faster run-up cases. If the rod line that is chosen is too low to cause a scram on high neutron flux with the fastest allowable run-up rate, then the fastest rate will be limiting. This is because the power peak from the fast rate case is not l limited by scram and will therefore be higher than=the power peak from any case with a slower rate. I 3.3.2.2 Limitina Case Studies I The most limiting RFCF event was determined by the RETRAN analyses discussed in Section 3.3.2.1 to occur when the master controller fails. The change in calculated CPR for this event is a strong function of the initial core power and flow condition. As the initial power level is h lowered at a given core flow rate for this event, the ACPR increases, Even though the ACPR becomes greater for events initiated from lower i
- 145 -
I I
I I, I. initial rod lines, the event will be analyzed along the 100% rod line. This is-justified by the fact that as the core is operated on a lower rod line, the actual operat1ng MCPR increases more than the calculated ACPR. l Thus, the master controller failure from points along the 100% rod line is selected as the limiting event. Since this event will result in a high g flux scram, the slow run-up rate will be limiting. Analyses of.the master controller slow run up failure were performed for the 65% core power, 38% core flow initial condition. Sensitivities for g 5 this event are expected to be similar to those obtained from the low flow point' of the 100% rod line. Two important parameters examined were the system model fuel pin gap conductance and the Maximum Combined Flow limit l (MCFL) setting. The effect of the cross section modification on the RFCF results was also investigated. As gap conductance increases, the change in CPR increases. However, the E. increase in the ACPR was small. A larger value of gap conductance produces conservative results. g W The MCFL setting determines the maximum amount of steam flow that is l allowed to pass through the main steam line control valves and bypass valves. Once the steam flow reaches the MCFL setpoint, the valves are not allowed to open further. This prevents the valves from regulating vessel l , pressure. Once pressure regulation is lost, the vessel pressurizes l_ rapidly causing an equally rapid power increase. The power increases ; until the high neutron flux scram setpoint is reached. The additional pressurization that occurs makes the event more severe. The amount of g W pressurization, if any, depends on the amount of steam flow generated during the event. Slow run-ups that are initiated from high power rod lines could produce steam flows high enough to cause a loss of pressure control depending on the MCFL setting. The effect of the MCFL setting on I l this event will be evaluated on a cycle by cycle basis, l
- 146 -
Ii 1 I I-
m-s Modification of the SIMTRAN-E cross sections is required for events with rapid and/or large amounts of core void change. Since the limiting RFCF event is not rapid and the amount of void change is small, the cross section modification process is not required for this event. The sensitivity study performed determined that using unmodified SIMTRAN-E cross sections produces conservative results for the RFCF event. Additional sensitivity studies were performed to account for uncertainties in calculating the RCPR for the most limiting RFCF event. The code uncertainty in Section 8 of Reference 2 can not be applied to this transient, since it was generated for an event that is not similar to the RFCF event. A separate uncertainty was generated specifically for this event. The uncertainty was derived based on the uncertainties associated with calculating void and Doppler coefficients of reactivity for the event. The uncertainties for the calculated void and Doppler coefficients are estimated to be 25 and 10 percent, respectively. The I sensitivities of RCPR to changes in Doppler and void coefficient were determined and multiplied by their respective uncertainties to calculate the uncertainties in terms of RCPR. The uncertainties in RCPR due to Doppler and void coefficients are combined by the Root Sum Square (RSS) method into a net uncertainty. This net uncertainty is doubled for conservatism (i.e., 2-sigma value), and expressed as a multiplier on the calculated RCPR for the event. The calculated multiplier for this ev at is 1.04. This multiplier can then be applied to the calculated RCPR to produce a conservative value of RCPR. This multiplier is applied to the I RFCF event described in Section 3.3.4. 3.3.3 Licensino Analysis Method The RFCF event will be analyzed on the 100% rod line for various core flows, since calculated RCPR increases as core flow is decreased along the
- 147 -
I I; rod line. Additional assumptions made in the analysis are based on the , sensitivity analyses described in Section 3.3.2. These assumptions ' include: E 1). The Technical Specification scram curve is used. !
- 2) The Technical Specification maximum scram delay is assumed.
- 3) The Simulated Thermal Power Monitor scram trip is not credited.
1
- 4) The High Neutron Flux Trip setpoint is at least equal to the allowable Technical Specification limit, plus calibration and l'
accuracy uncertainties, g
- 5) The limiting run-up rate will be determined for the event.
- 6) The RCPR uncertainty generated in Section 3.3.2 is applied to- ,
the calculated RCPR for this event. This multiplier is applied to the RFCF event described in Section 3.3.4. The sample licensing analysis presented in Section 3.3.4 used the above I' assumptions. 3.3.4 Samole licensino Analysis
- This section presents a sample licensing analysis for a Unit 2 Cycle 2 RFCF from 65% power /38% flow using the conservative assumptions described in Section 3.3.3. The transient behavior of key system parameters is .
presented in Figures 3.3 1 through 3.3-7. The calculated change in CPR
'for the event, including the RFCF event specific uncertainties is 0.44.
l For an actual reload application, several RFCF analyses will be performed g
- 148 -
1 I _. .. a
8 I from points on the 100% rod line to generate t5e MCPR operating limit as a function of ~ core flow. I I I I I I I I g . I
\
l B B B
- 149 -
3;
Il i Ii i TABLE 3.3-1 ; Sin!11e and Two Loop RFCF Events I' _ In' tial Power / Flow Conditions l Approximate g :- core Power Core Flow Rod Line f% rated) (% rated) (%) g 65 38 118 40 38 73 40 65 52 I i I I L r I' I I I I
- 150 -
I I I
l l I. l I TABLE 3.3-2 I Change in RCPR as a Function of Master Controller Run-un Rate for the 65/38 Power / Flow RFCF : I Run-up Rate ! f%/Secl Chance in RCPR j I ; T 0.3 .006* l 0.4 0.0 > I i 0.5 .002 ! I 2.0 .007 I 10.0 I .032 l ,
- Did not result in reactor scram
- I ll I - 151 -
I I
I . W - RGURE 3.3-1 : RfCf CORE POWER 14 0 I 12 0 .................. ..... ..... ..
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. . . 4 RURM 0 10 $0 $0 /0 $0 IO70 $0 NO Id0 11O 120 150 ll0 1$0 150 170 TIME (SEC)
- 152 -
I
) ) FIGURE 3.3-2 : RFCF CORE FLOW ) l 100 l 1 90- - - 1 I 80-O 1 W
=
I g70-l 8 i S 360-l e 1 8 so- - - - - 1 I 40-l.egend 30 , , , , , , , o io 20 30 40 so so 20 80 20 ioo iio no ao uo iso iso i,o 1 M(SEC)
- 153 -
l
1 I1 i RGURE3.3-3 : RfCf DOMEPRESSURE I , 1030 l! t I, 1020- >- -
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- 154 -
, ., . . - - ,_ _-.,.,_____.-,,,,,,.,m.e-- ,. . , . , , . _ . _ _. m _ _ . , , _ - . ..%_,m_
I l I RGURE3.3-4 RFCf VESSELSTEAMFLOW I i 12 0 I I I
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- 155 - l I '.. ,
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l . ~ ' ' FIGURE 3.3-5 RfCf NARROW RANGE LEVEL I' t 60 I: . g- . . l
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5 . . . . . . g i 50 . ................ i
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RURAN 0 l0 2O $0 /0 $0 $0 IO 8O i0150110 d0 d0 ll0 t$0 l$0170 WE(SEC) - 156 -
-l
I l FlGURE3.3-6 RfCf FEEDWATERFLOW I 12 0
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I 1j0. . . . . . . . . . . . . . . .. . . . . . . . . . . . . . . . . . ........,
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- . i :
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i - g g g g g g o . 8 L99end ( 60 , . . . . , , , I 0 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170 IlME(SEC) I - 157 -
+ .,.a. .. ..a-. ..o a -a L.s - e: - n.-r- 4 as--r,- eu. '.s a asa .a* <-.-+-a es..-e, -a--n,.+.,m.+n-..<na_ w-,-mau, I;
nGURE3.3-7 RfCF AVERAGEHEATR.VX ! I' 130
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l . . . . s j . . . . . . . . RURAN 60 ' ' ' 0- l0 $0 $0 /0 $0 b f0 8O 90 Id0110 N0 N0 N0 $0 $0170 WE(SEC)
- 158 -
P 4 3.4 Non-Limitina Eventi I
'his section describes an evaluation of the Susquehanna SES FSAR (Reference 14) Chapter 15 transient events (i.e., anticipated operational occurrences) to determine which events need to be evaluated in I establishing the MCPR operating limits for each reload licensing submittal. .The detailed analysis methods for the events selected for I analysis on a typical reload are described in Sections 2.1 to 2.3 and 3.1 to 3.3 of this report. This section describes the evaluations performed to establish that the other Chapter 15 events are non-limiting and, hence, that they will not be analyzed for each reload.
The FSAR Chapter 15 transient events for which calculated ACPR is of importance are divided into 5 main categories: 1'
- 1) Decrease in Coolant Temperature (15.1) i 2) Increase in Reactor Pressure (15.2)
- 3) Decrease in Reactor Coolant System Flow Rate (15.3)
- 4) Reactivity and Power Distribution Anomalies (15.4)
- 5) Increase in Reactor Coolant System Inventory (15.5)
I Calculated values of ACPR for Cycle 1 are given in Table 15.0-1 of the i FSAR (Reference 14). Events which show a calculated ACPR greater than zero in FSAR Table 15.0-1 and potentially limiting events in each category were evaluated. The events selected on this basis for evaluatirn are listed in Table 3.4-1. Evaluations of specific events identiftad as non-limiting in Table 3.4-1 are provided in the following sections, i 1
- 159 -
l . .
I, 3.4.1 Turbine Trio I! A Turbine Trip Without Bypass (TTWOB) was analyzed and compared to the I GLRWOB as described in Section 3.1 of this report. The events are nearly identical with only small differences in the times for the start of TCV g' motion. The results presented in Table 3.1-2 show the TTWOB to be slightly less limiting than the GLRWOB, due to the fact that the TCVs . j start to close slightly later with i apect to time of reactor trip for the 1 TTWOB. Thus, the TTWOB need not be analyzed as it is less limiting than I the GLRWOB. 3.4.2 MSIV Closure (Position Scram) Il i The MSIV closure from 100% power /100% flow with reactor trip on valve i position (85% open) was analyzed. The calculated ACPR was equal to 0.0. Because of the rapid' post-trip power reduction, which occurs prior to the pressurization induced power increase, the core average heat flux never g exceeds its i .tial value. As indicated by these results, the MSIV 3 closure with valve position reactor trip is less limiting than the GLRWOB for calculated ACPR. 3.4.3 Loss of Condenser Vacaum ' l The Loss of Condenser Vacuum event is described in Section 15.2.5 of the Susquehanna SES FSAR (Reference 14). The loss of Condenser Vacuum consists of a turbine trip with partial bypass capacity. Reactor scram g7 occurs as the result of the turbine trip. Because of steam release 5 ! through the bypass valves, the pressurization and, hence, the calculated ACPR for this event would be less than those calculated for the TTWOB and GLRWOB events. hi I
- 160 -
I' I,
. . . . . . - . . _ _ . ~ -
I t
!3.4.4 Recirculation Pumo Trio i
- An analysis was performed of the simultaneous trip of both recirculation
~
pumps. The- rapid decrease in flow produces increased voiding in the core - ' and.causes'a rapid decrease in core power. The calculated ACPR was
'l approximately 0.02, which is an order of magnitude lest than typical limiting events. Thus, this event is judged to_be non-limiting, s
3.4.5 _ Inadvertent HPCI Startuo
' A RETRAN analysis of the inadvertent startup of the High Pressure Coolant '
-Injection (HPCI) system was performed at 100% power /100% flow. Plant _ test
, . data was used to produce a table of HPCI flow versus time. The condensate storage tank supplies the initial-HPCI flow to the reactor vessel. Plant data indicates that the actual HPCI injection flow temperature is usually -
I greater than 100'F. Thus, .the analysis conservatively assumed that the
~
HPCI system delivered 40'F water to the reactor vessel. Also, the high-
. neutron flux and high thermal power reactor trips were disabled to allow the event to reach a new steady state. The calculated ACPR for the HPCI startup was less-than 0.10. Since the event is_ expected to be relatively f ,
independent of unit and- cycle, -and more limiting = events wili produce significantly higher calculated ACPRs, the event is considared to be non-
'g limiting. 4 I
E B B
- 161 -
I
i- . -;
-l
-1 TABLE 3.4-1 .
Potentially Limiting Events in Establishino MCPR Operatino Limits: 1 Event Classi fication* -. l.. Loss'of Feedwater Heating R
- 2. Feedwater Controller Failure R a
- 3. : Generator Load Rejection -5, R
4. 5. Turbine Trip MSIV Closure (position scram) N/L
.N/L l
- 6. Loss of Condenser Vacuum N/L
- 7. Recirevhtion Pump Trip .N/L
- 8. Rod Withdrawal Error R
- 9. Recirculation Flow Controller failure (increasing) R :,
- 10. - Fuel Lt " r R
- 11. Inadver^en:. PCI Startup N/L ',
R = Anab ' ow 'ar each reload N/L LNot g' r anstrated by analysis or evaluation; .do not ce n. Ah reload Ii lu: I'
- 162 -
.l-e x. - . . . - . . , ,
I LI LI .3.5 ASME Overoressure Analysis i 1 l-l L 3.5.1 frjfterji.a " I The Susqueht.nna MS nuelsar pressure relief system has been designed to prevent overpetsuritation which could lead to failure of the reactor I coolant pressute boundary. The #dE Roiler and Pressure Vessel Code, ' Section III (Reference 24) requires that the peak pressure be maintained less than or equal'to 110%'of the vessel design pressure under upset conditions. I The combination of the Reactor Protection System and the Safety / Relief
- Valves (SRVs) must ensure that this criterion.is met. Analyses of the limiting' overpressure event for each reload will be performed to c I demonstrat'e that the peak calculated pressures are within 110% of the design pressures of-the reactor coolant pressure boundary components. As specified .in Standard Review Plan 5.2.2 (Reference 19), credit can be taken for the. spring loaded safety valves and the second safety grade reactor protection signal.
i :. The-components which comprise the reactor coolant pressure boundary have been designed with several different design pressures, as detailed in FSAR Table 5.2-3 (Reference 14). The criteria used for overpressure analyses < on the Susquehanna SES units are given in Table 3.5-1. g The criteria specified in Table 3.5-1, which represent 110% of the '.5 pressure boundary component design pressures, must be met for the most adverse pressurization transient. The two potentially limiting pressurization events are the closure of all Main Steam Isolation Valves h' '
'(MSIVs) and the Generator Load Rejection Without Bypass (GLRWOB). Both L
- 163 -
I IR .1 these events were evaluated by PP&L using the RETRAN models and methods I described in this section and assuming that the first safety. grade reactor trip does not function. Thus, both events produce a reactor scram on high lR neutron flux. Results of the RETRAN analyses demonstrate that the MSIV-closure event with high neutron flux scram produces higher vessel g, pressures than the GLRWOB with high neutron flux scram. g W L 3.5.2 Event Description In the Susquehanna SES, MSIV closure is initiated by any one of a number I: of trips (e.g., by low reactor water level, high radiation, low condenser vacuum,-operatoraction,etc). The MSIVs are large valves (two in each of f' ' the four steam lines) which have a range of allowable closing times. specified in the Technical Specifications. The maximum allowable closure time is-established to contain fission products and to ensure the core is not uncovered following steam line breaks. The minimum allowable closure time'is limited by vessel component overpressure criteria. ! The analyses presented in this section assumed an inadvertent simultaneous t closure of all MSIVs. As the MSIVs close, valve position limit switches activate the Reactor Protection System (RPS), which produces a reactor hi scram. For overpressure. analyses, however, this trip is not credited, and g reactor. scram occurs on high neutron flux, g l The decrease in steam flow caused by the MSIV closure produces an increase ! in vessel pressure. The pressure rise continues and some or all of the Safety Relief Valves (SRVs) open. Normally, the SRVs will open as the E g
- i. result of a vessel pressure signal 'in the relief mode. However, for the overpressure analyses, the relief mode is assumed to be unavailable, and L
p spring pressure safety mode operation is assumed.
- 164 -
I' I-
I ' I l
. As. the pressure increase collapses voids in the core, neutron power rises
'and'downt.arer water level decreases. The decrte ing level will generate a l reactor recirculation pump runback. signal.
System pressure typically increases further until the Anticipated Transient Without Scram Recirculation Pump Trip (ATWS-RPT) high pressure ' setpoint is reached. This occurrence does not have a pronounced effect on I overall vessel pressure, but it does affect the overpressure margin 1 locally in the external recirculation loop discharge piping. i feedwater response to the MSIV closure event is assumed to provide a constant flow. Although the steam driven feedwater turbines are isolated , from the reactor steam supply, the feedwater pumps continue to inject-
~
feedwater into the vessel for approximately 30 seconds utilizing steam from the steam line~ piping downstream of the closed MSIVs. The additional g mass and energy of'the feedwater flow added to the subcooled downcomer !W fluid effectively reduces the steam space in the vessel and results in an increased system pressure rise. The decreased power level following scram and steam release through the SRVs limit the vessel peak pressures. Ll 3.5.3 RETRAN Modellino The base RETRAN system model used for the overpressure analysis of the MSIV closure is described in Section 3 of Reference 2. In order to I conservatively model the MSIV closure event, three changes were made to the'models described in Reference 2:
- 1) A conservative model of the safety mode of the SRVs was developed, since the Reference 2 "best est%te" model represents the relief mode of the SRVs.
IL - 165 - Lt .
} ;f ,
y
'2) t An improved MSIV model was-developed to better represent th'e -
' transient flow rate during valve closure.
This model, which is based on data'from the valve manufacturer ~, was then-
~
conservatively biased to reduce steam flow more rapidly. than-would be expected. R
.j
- 3) A ' trip and control system was added that automatically .
'l calculates _the pressure boundary pressures and compares them to a their respective peak pressure criteria. This control system is '
c for information only, and it has no effect on the calculated ! system response, f; L 3.5.3.1 Safetv/ Relief Valve Model SusquehannaL SES Technical Specification 3.4.2. (Reference 16) specifies l, that 10 of the 16 SRVs must be operable and specifies the valve setpoints b
~ and tolerances. The model developed for the overpressure analysis assumes -
the 6 SRVs with the lowest pressure setpoints are inoperable.and that the M L . valve setpoints are at the high pressure end of the tolerance band i specified in the Technical, Specification,(i.e., +1%). The SRV vendor L ~ specified maximum opening times were used to minimize the' pressure relief h 11 through the SRVs.' The assumed.SRV opening characteristics are:shown in 1 yi Figure 3.5-1. Table 3.5-2 supplies the valve capacities used in the b g '
' analyses.
The-RETRAN system model.was modified to incorporate the assumptions' described above. Valve openings are initiated by means of RETRAN trips, E L 5
> and the mass flux through the SRVs is calculated via the control system illustrated in. Figure 3.5-2.
I E
- 166 -
IL I. I
_ 3.5.3.2 MSIV Model Since the publication of Reference 2, PP&L obtained detailed loss coefficient data on'the Atwood & Morrill MSIVs. This data, which consists of loss coefficient versus valve position, was used to produced more-
~
7 accurate modelling of the reduction in steam flow caused by closure of the - a MSIVs. The Reference 2 MSIV model was replaced by the one described herein. Since none of the Reference 2 benchmarking analyses involved closure of the MSIVs, those analyses remain valid. The reduction in steam , flow during MSIV closure is modelled as an increase in loss coefficient for a constant-junction area (i.e., valve area assuined constant), The control system to vary the valve loss coefficient as a function of valve position (i.e., time) is illust' rated in Figure 3.5-3. The ratio of loss coefficient to full open lass coefficient as a function of stem position
=was-increased of 25% to produce a conservatively rapid reduction in steam flow as the MSIV'. close. The effect of closing both inboard and outboard MSIVs is modelled.
3.5.3.3 Trio and Control System Two RETRAN control systems were set up to track peak vessel pressures for comparison to the relevant ASME criteria. -These two control systems provide edits for analyst convenience, and they do not ' affect the transient being analyzed. The first control system calculates pressures for components'which have a 1250 psig design pressure (Figure'3.5-4). The
-RETRAN volumes whose pressures are tracked with this control system are:
J 1) Steam Dome (Volume 120) Upper Downcomer (Volume 130)
, 2) j 3)- LowerDowncomer(Volume 148)
- 4) Steam Line Inside Containment (Volumes 320 & 330)
- 167 -
2
u.
!ih The pressure.at_the bottom of the reactor vessel is taken to be the RETRAN lower plenum (Volume 160) pressure plus the elevation head associated with U
lt one-half the volume height, since RETRAN pressures are calculated at the middle of a volume. Based on GE design data, the pressure in the feedwater= lines:.is less than or equal to 28 psi higher than the dome pressure. Thus, the RETRAN dome pressure (Volume 120) is increased by 28 ; psi for comparison with the limit. The control' system illustrated in - Figure 3.5 4. calculates the maximum of the above defined pressures for 4
, comparison with the pressure criterion. ;
The control system.that calculates pressures for the vessel components whose design pressure is 1500 psig is illustrated in Figure 3.5-5. The l pressure in the recirculation pump discharge piping (Volume 220) and the pump discharge pressure (based on the RETRAN junction pressure plus the l contribution of the recirculation pump minus- the local- frictional loss) : ., are tracked. This control system assumes that the event is symmetric with - - respect to.the vessel and should not be used for asymmetric events. It g should be noted that results of MSIV closure analyses clearly indicate 5 that the components having a 1500 psig design pressure will not be the 1Imiting' components. 3.5.4 Sensitivity Studies c . L RETRAN system model analyses were performed to determine the sensitivity i I 4 of_ peak power and vessel pressures to changes in input assumptions for the MSIV closure event. - The. parameters analyzed were: [ 1) Core power
- 2) Feedwater temperature
- 3) Pressure regulator setpoint p _4) Steam line pressure drop .
- 5) Control rod insertion rate
- 6) Gap conductance >
- 168 -
I t
B. I
- 7) MSIV closure time
- 8) Neutronics input / Cross section modification l 9) Feedwater flow transient behavior 10): ATWS recirculation pump trip setpoint
- 11) Steam dome volume I The base case values of the above mentioned parameters are listed in Table' 3.5-3. Closure of the MSIVs was initiated at time zero by activating the MSIV manual trip logic in the RETRAN model (Reference 2).
The study results demonstrated that the components whose design pressure is 1250 psig had the closest approach to their ASME overpressure limit.
-l.. The components whose design pressures are 1230 and 1240 psig were slightly.
less limiting, and the components whose design pressure is 1500 psig lg showed at least 200 psi margin to the _ applicable ASME overpressure limit.- The study results are presented in Table 3.5-4. The peak pressure changes I are presented for the 1250 psig design pressure category only, since this category exhibited the closest approach to the applicable ASME overpressure limit. Conclusions and observations drawn from the analysis results are discussed below. The results gsven in Table 3.5-4 indicate that the peak pressure-is most strongly affected by: core power, core flow, feedwater temperature, oressure regulator setpoint, scram curve, and MSIV closure ' time. .The other parameters investigated produced a less than 2 psi change in calculated peak pressure. I A higher initial core power implies a higher steam flow (all other things being unchanged), which produces a larger pressurization of the vessel on MSIV closure. The higher pressurization rate also increases the peak core power due to positive void reactivity feedback, which increases the energy
- 169 -
I I -
.; 1 I,
'1 f j
~
input to the vessel and,. hence, the pressure. Thus, higher initial core power is conservative for MSIV closure overpressure analysis. .l, A higher initial core flow has two main effects on the MSIV closure ~ ' events. First, higher core flow means less voids in the core, which makes ~ the vessel water'less compressible and increases the pressurization. Second, the higher core flow produces a more top peaked core axial power distribution. The more top peaked power distribution results in a higher peak power due to scram effects, which increases the energy deposited to 4~ g the vessel and increases the pressurization. Therefore, higher core flow L! is conservative for MSIV closure overpressure analysis. Based on the j sensitivity study.results (cases 1 and 2), however, the effect of a change l in flow on the order of the flow uncertainty (2.5%,- Reference 17) is only t about 1 psi. ' A lower feedwater temperature et 'a given power / flow condition has the effect of reducing the initial vessel steam flow. A lower initial steam g:; flow makes the vessel pressurization less severe for the MSIV closure, E! resulting in lower peak power and pressure. The effect is relatively minor for changes on the order of the temperature measurement uncertainty
.(0.76%,' Reference 17).
h_ The main effect of a higher _ pressure regulator setpoint is to raise the A L 4 initial dome pressure. While this places the initial pressures closer to- ] the ASME criteria, it' also_ places the steam line and vessel pressures a closer to the.SRV setpoints and ATWS RPT setpoint, respectively. Thus, g the net effect is only a 6 psi increase in peak pressure for a 32 psi n, increase in pressure regulator setpoint (Table 3.5.4, Case _4).
; Since the larger part of the steam line is downstream of the MSIVs, the Ij steam line pressure drop has negligible impact on the MSIV closure event.
L [ - 170 - I L
i I t g
-l A faster reactor scram control rod insertion reduces the peak core power and the integrated heat added to the vessel, which reduces the peak. vessel l pressures. The effect of using the Technical Specification scram curve versus.a best estimate scram curve is substantial (increase in peak pressure of 33 psi, Cases 1 and 6).
I The gap conductance has little effect on peak vessel pressure (Case 7). However, use of a lower value results in a slightly higher peak pressure, 1 l
, , Thus, the same conservatively low value generated for ACPR calculations R
., (See Appendix A) will be used for the MSIV closure overpressure analyses. '
A shorter MSIV closure time creates a more significant vessel pressurization since less energy is removed from the system.
' t The modification cf cross sections as described in Section 2.6 has a i I negligible effect on peak vessel pressure. This is due to the fact that a MSIV elosure is a less rapid event than a GLRWOB,'resulting in smaller f changes in core nodal water densities prior to the insertion of significant scram reactivity. Since the cross section modification process is performed to account for differences between RETRAN and SIMULATE-E calculated moderator density changes for rapid pressurization ,
events, the effect is much less than it is for the GLRWOB. Also, small changes in peak core power have only a small effect on peak pressure, I although they may have a big effect on ACPR. Therefore, use of the cross ; section modification process is not necessary. .! The impact of feedwater flow differences during the MSIV closure event is I extremely small (Cases I and 10), mainly due to the fact that peak pressure is attained in about 5 seconds, and the feedwater flow I I differences have not had sufficient time to significantly influence vessel l conditions.
- 171 -
l'
I; The Anticipated Transient Without Scram. (ATWS) Recirculation Pump Trip (RPT) setpoint has a negligible effect on peak vessel pressure and power l, since.it occurs after peak power and not long before peak pressure. Small changes in steam volume in the vessel, such as could be caused by a j different vessel . water level, will have a negligible impact on peak vessel ' l pres'sure for the MSIV closure event, due largely to the very large steam g-l space upstream of the MSIVs. 5 1 3.5.5 Licensino Analysis Method The ASME overpressure analysis will be b -sed on~ the event defined by a ' simultaneous closure of all MSIVs. Th following conservative assumptions ' .
-l will be used to perform the analysis ror a licensing application:
- 1) The event will be analyzed at 104.4% of rated core power (105%
steam flow) and 100% rated core flow. - q
- 2) The pressure regulator setpoint is assumed equal to 966 psia, which corresponds to a dome pressure of approximately 1040 psia. .
- 3) =The' Technical Specification (maximum) average scram times will be assumed.
- 4) ~ A value of MSIV closure time which is less than or equal to the Technical Specification minimum closure time will be assumed.
}
[ 5) .Feedwater Flow will be held constant at its initial value. l
- 6) A conservatively low value of gap conductance will be used (see
- Appendix A).
- 7) The maximum value of the ATWS-RPT pressure setpoint including uncertainties will.be used. '
- 172 - !
l
~
n, q.. < !
-in I
- 8) The conservative MSIV model' described in Section 3.5.3.2 will be - I used. I I -The- results 'of the RETRAN system model benchmarking analyses presented in
. Reference 2-show that the PP&L best estimate methodology can accurately predict vessel pressurization (Reference 2, Sections 5'and 6). Therefore, the conservative assumptions outlined above introduce significant ;;
conservatism into the calculated peak vessel pressures...The cross section modification process will=not be used for MSIV closure overpressure analyses. l- 3.5.6 Samole Licensino Analysis This section presents a sample licensing analysis.for-~a Unit _2' Cycle' 2 MSIV: cloture event. The analysis was performed using the set of i I assumptions described in Section 3.5.5, with the exception that the nominal pressure regulator setpoint'was used. However, on an actual reload licensing application, the higher pressure regulator setpoint will be used, q E The transient behavior of selected system parameters is presented in l. Figures 3.5-6 to 3.5-13. The calculated peak pressures for-the event and l- their locations are given in Table 3.5-5. Adequate margin to the ASME overpressure criteria was demonstrated. .
- 173 -
7 it -' s 3 TABLE 3.5-1 ASME Overpressure Analysis
. 0 ' Criteria Design Pressure Criterion Component- (osia) (psia) j Reactor' Vessel 1250' 1375 -
Recirculation Suction Piping 1250 1375 Main Steam Pipin9- 1250 1375 Recirculation Discharge Piping 1500 1650 Recirculation Pump 1500 1650
~HPCI and RCIC Steam Piping. 1230 1353 Core Spray System 1240 1364 I;
E,: I, Il I?
- 174 - +
l __________u1________________________
y wp ,
- l b g_
mg TABLE-3.5-2 y
.' Safety / Relief Valve p Flow Rates ig,,
f . Flow Rate at 103% Nominal Spring Spring Set Pressure Number of Valves Set Pressure (osia) (1b/hr valve)** 2 .* - 4 * - y 4 1185 891,380'
'3 1195 898,800 h 3 1205 906,250 I *These valves were assumed to. be inoperable.
I -** Flow rate =from the Susquehanna SES FSAR-(Reference 14). I I I I
- 175 -
g I ..
d4 y 7
.g -
TABLE 3.5-3 4 MSIV Closure /ASME Overpressure Analysis Base Case Input Assumptions
+t Parameter Assumption Ir:
- 1. Core Power 3293 Mwt (100% rated)
- 2. Core Flow 100 Mlbm/hr (100% rated). f
- 3. Feedwater Temperature Best estimatevalue based on A plant data =
- 4. Pressure Regulator Setpoint Best estimate value based on
! J plant data ~
- 5. Steam Line Pressure Drop Best estimate values
",< (Reference 2)
- 6. Control Rod Insertion Technical Specification ,
insertion rate '
- 7. Gap Conductance Value calculated using lj
. Appendixpmethodology(855 BTU /hr ft F) .
- 8. MSIV Closure ~ Time - 2.0 seconds 1
- 9. Cross'Section Modification Cross section dependencies:
modified (see Section 2.6)
- 10. Feedwater Flow Transient Determined by the feedwater Behavior controller (Reference 2)'
'l l . ATWS Recirculation Pump Analytical setpoint (1184.7 5 Trip Setpoint psia) 1 i 12. Steam Dome Volume Best estimate value a L (Reference.2) g p
L !, 176 - L
"l'll . ..Iii i i iii si i i iii ., m i is iiiii,iigi si e i ,ii is ii i m i l
l TABLE 3.5-4 MSIV Closure /ASME Overpressure Analysis Results of Sensitivity Study L3 Change in 3 Change from Peak Core Change in Peak Parameter Base Value Power (% rated) Pressure Insi)
- 1. Core Power +4.4% rated +76% +7.
- 2. Core Flow -13% rated 31% -6.
- 3. Fee ~ '. iter -10*F 9% -7.
T' s erature
- 4. Pressure +32 psi -9% 40.
Regulator I 5. Setpoint Steam Line +8.65 psi +5% +0. Pressure Drop
- 6. Control Rod Best estimate based -23% -33.
Insertion Rate on plant data
- 7. Gap Conductance o 1000 -14% -2, Increase}F)
(BTV/hrft l 8. MSIV Closure Time Increased to 3 sec.- -31%
-16.
- 9. Cross Section Not Modified +10% -1.
Modification
- 10. Feedwater Flow Held constant at 100% +1% +2.
Behavior
- 11. ATWS RPT Decreased to 1149.7 +0% +0.
l Setpoint psia (nominal)
'E 12. Steam Dome -138 ft 3 +4% +0.
1W Volume l - 177 - B I.
. - - . - - - - - . . - - . . - . . . . . - . . . . ~ - . . . - - - . . - . . - - _ . - . .
m *' l I. TABLE 3.5-5 MSIV Closure Sample Licensing Analysis Peak Calculated Pressures-- c
~ Design Pressure- Maximum Time of Peak cateaory criterion (osia) Pressure (osia) Pressure (sec) -
1230 psig 1367.7 1301 5.3 1240 psig 1378.7 1308 5.1 1250 psig 1389.7 1331 5.2 1500 psig 1664.7 1412 3.5 I; I a
\
?I I 1 L - 178 - l u . s - , . - - = . _ , - e, .-,
,7 11
$l 4 h l , FIGURE 3.5-1 1 $1 ASSUMED SRV FLOW CHARACTERISTICS 12 0
;I l- >
100- - ____ m . l. t.e ,
-r 80-8
- I L 3R '
- ). /
o 60- , E l L ,e Lo / A
.e 40- /
I.,
/
/
20- /
/
l : /
,: /
2 0 , , , , , 98 99 100 101 102 103 104 Pressure / Pressure Setpoint (%)
- 179 -
FIGURE 3.5-2 : CONSERVATIVE SRV FLOW MODEL - TRIP 201 s SRV MUL s FNG s MASS
-901 g = 1/1211.6 ' -904 g = 2234.
' (B 3)
PRES 310 TABLE 97 TRIP 202 s SRV Mut s FNG s MASS
~
-902' h s g = 1/1221.7 g = 2253. (B N 4)
PRES 310 TABLE 97 TRIP 203 s SRV MUL s FNG s MASS
-903 g = 1/1231.8 ' -906 g = 2271.
(BAN 5) PRES 310 TABLE 97 NOTE: Banks 1 and 2-(6 values) are assumed out of service e HEEll' ~ M M ME E
, g e- e e e aus Bigg
I~~,,, ,, ,,, ?,,, :,,,s ,,, ,,, : ,,, ,, . .e ,,, . : ,,, ,, 9 FIGURE 3.5-3~: MSIV LOSS COEFFICIENT . CONSTANT) g g = .086 SUM 4 -320
-> -708 LOSS COEFFICIENT OR ZED MSiv TRl('
O
. MUL ' s INT POSITION s FNG COEFFICIENT a; -705 -706 g = CIC* -707 g = .727 CONSTANI 7 (=1.0) '
. TABLE 3 JUNCTION 9 SUM 330
-709 4 LOSS CONSTANT > COEFFICIENT g = .322
- gain = -(11 MSIV closure time); SEC"
4 RGURE 3.5-4 : OVERPRESSUliE MARGIN (1250 PSIG DESIGN PRESSURE) . PRES:120 O6- MAX PRES 120 -
-908 SUM
= 28
\
s MAX 0404 ' MAX s
-913 y PRES 148 -909 '
, o MAXIMUM
' PRES 160 PRESSURE y MAX FOR 1260 i Oo SUM -915 3 DESIGN
-910 PRESSURE i
1 AVED 160> > COMPONENTS ' 0=.06 - hlAX PRES 210 911 MAX PRES 320 -914 MAX l
-912 t 40 PRES 330 >
l ; M 'M M M ^M M M M- M M M M 8 MI M M M
.t I-
~RGURE 3.5-5-: OVERPRESSURE MARGIN (1500 PSIG DESIGN PRESSURE) y AVDJ 215
., O414 13 WP" 215 '
' l' MUL' DIV s i
WP" 215 -916 - 917 g = CIC* SUM s O413 g -e.7 sex 10e )
-918 '
SUM MAXfMUM PRESSURE i t
. PJUN 215 > MAX _y FOR 1600 -!
4 11 -920 DESIGN E >
"' PRESSURE !
COMPONENTS PMPP. 215 i O415 PRES 220 I O410 - i l
- gain = lnitial values'of loss coefficient in Junction 215 from the HETRAN initialization run f
._ . _ _ - e.-.,. , m .-- ~ .,, L ,- , -...-l. ~
r , v s. ,
]'
Ili l RGURE 3.5-6 : MSIVCCORE POWER lj s 400 lL ~v . . 50 ..... .......... .. 300 ....... t
- n. .
- w. .
l . j, . ..;; I . . . - l . F
. .)
%/ . ..... .
.w ..
w- . e . . O . . . 10 0 ....... 1 l . . 4 +
.0 .... .
p . I* . . . 0 , , , , , , , , , 0- 1 2 3 4 5 6 7 8 9- 10 TIME (SEC) I
- 184 - ,
~-l
t a B l g RGURE 3.5-7 : MSIVC- CORE R.0W I 12 0 ' I 1 mo. ......., ........ . t . 10 0 ....... ...
...s..
. . . . . . .... . ... . . . . . . . , .. . . . . . . . , . ... .. t 1
.n .
-w- . ..... .
- s. .
o;80- >.- i _y . a . . w . .
. O 60 ... . . .. .. .... ..
.1 l . .
j r .
. i .
j 3o. ......
.........>... .y...... .. . . , ....... .... . ..,.. . .
l . . .
.I l
i ! i !. -
.! . RETRAN 30 i i , , , ,
i i i ; o 0 1 2 3 4 5 6 7 8 9 10 i IlME(SEC) r
- 185 - , ,-. I-D- V g-v -
m I RGURE 3.5-8: MSIVC DOME PRESSURE- l
'1350 I ,
. i3c.o. ....,......... . . . . . . , . .
u3o_ . . . . . . .
.....m.
u . w .
. . .\
l y.000- -) l x . 3 . on: y( , . . A '1150- -. .> < i
.'- i'- -)
n . . \ O .
. 4 .
- l. .
1100 - " " " :" " "4 ,'
- 4. "- a. - " -
. . . . . 4 i
c . i 1M0 . o . . . . . . . \- . . . O en t . k
- . ! : REIRAW 1000 e i i i i i i e i g,
- 0- 1- 2 3 4 5 6 7 8 9 10 L
TIME (SEC) I
- 186 -
I
*I
. Ii flGORE 3.5-9 : MSIVC VESSEL SEAM R.0W y
4000
'a ,
;y I
. . . . \ .
. 1 l 3000 .......
l
- a. .
. . . s .
x . 2000- :.
.a
. ,1 1
l. v .
+
l 1000 .
. ~.
- l. w ,. ~l
+ .
.r m .
l
.J 0- .
.w- 1. .
I m w ... . l .
, _3000 .... .. . ...... . . . . . . . . . : . .. .
3- . 09.end g . . - -
-2000 ; , , , ; , i i ;
O 1 2 3 4 5 6 7 8 9 10 M(SEC)
- 187 -
l
l.. 1 3' r
. FIGURE 3.5-10 : MSIVC NARROW RANGE LEVEL lJ l
l 35
)
- g. . . .
i , r 30 .. ..... ..... . ... . . . . . .. . 2+~. n .
.w, . .
.I .
0 .
. .. ....... .... . . . .. .... .. 1 z 25
--v
. 1
.a . .
.w w .
.J . .
w 20 .
.o . .
- z. .
1
.g. .
-1
. . . . . . -1
. O. . .......
g :. . g.
- l. . .
I - . . . . . . s - z ', . 0 l i : .
'. RURAN .
5- h. h h 0 1 2 3' 4 5 6 7 8 9 10 TIME (SEC) ll
- 188 -
1
i
. Lk 4
4 RGURE 3.5-11:WSIVC FIEDWATER fl0W I no 1
]
8 0 ....... ............. . .
. . 1 8
x I t . , 0 g 90 .. I ... .. . ...... ........ .. . . . . . . .......... ....... v . . s< I x I
. . t a 50 ....... ...... ..
,r . .
I>. w . r< . 1 o . w- . w . . .I.. t . . .
- t. . ....... ......... .... .
I t . l _e n E , l 1 i i i i i i 1 0 1 2 3 4 5 6 7 8 9 10 l M(SEC) 189 -
.5.
wt 'yv--- -% -wM s' -%ywr# hy .esT-rw--rg-e--+hwww-e+-mvp -me w-ve+ 4rtowwwwa*wetwe-- w%W+t-vmw-W--1*e e er,metyv -e.--_ rs-,my.-mww-+-__---m -
! '. .9 ,('
) f g4 l
flGURE3.5-12:WSIVC AVLRAGEHEATPR l' wo O
. 1 l- 0 ... .. .. ..... ..
I Q L. . . . . . .
- i W. .
- j. . .
.s k .
0 l< ....... .. ... .. . . 0 . . y .
.}}