ML20058H351
| ML20058H351 | |
| Person / Time | |
|---|---|
| Site: | Millstone  |
| Issue date: | 08/31/1990 |
| From: | Brown U, Novendstern E, Skaritka J WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20058H326 | List: |
| References | |
| NUDOCS 9011150180 | |
| Download: ML20058H351 (390) | |
Text
{{#Wiki_filter:. . . , _ - t s WESTINGHOUSE PROPRIETARY' CLASS 3 - PLANT SAFETY EVALUATION FOR MILLSTONE GENERATING STATION UNIT 3-- VANTAGE SH FUEL UPGRADE" .,
.. AUGUST 1990i Editors: J. Skaritka :
-U. L., Brown . ;
Approved: E. H. Novendstern ' Thermal Hydraulic, Design and Fuel Licensing c - WESTINGHOUSE ELECTRIC CORPORATION Commercial Nuclear Fuel Divicion P. O.i Box.3912 ' Pittsburgh, Pennsylvania 15230' e '
; 00309-13.10 082790f ~
Attachment to' THFL 90 520 t I 9011150180 901101 P '
.I PDR ADOCK 05000423 ,
.e, P PDC ' -
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ACJ(NOWLEDGMENTS j G The editors would like to acknowledge the efforts of the many contributors to the design l - Sections to this Plant Safety Evaluation. Major contributors include: l i Mechanical Design - R. W. Brashier, Y. C. Lee ! Nuclear Design R. J. Fetterman .! Thermal and Hydiaulic Design A. J. Friedland - 'l Fuel Rod Performance C. A. Bly Ncn LOCA Analysis - P. W. Rosenthal, D. S. Huegel, 1 LOCA Analysis C. Mi Thompsoni M. P. Kachmar, l
. J. Hansen ,
Radiological Assessment - K. Rubin Control System Design- - R. A. Carlson Containment Design 'L. C. Smith Instrument Uncertainty for RTDP W. H. Moomau
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l PLANT SAFETY EVALUATION for THE MILLSTONE UNIT 3 FUEL UPGRADE l i TABLE OF CONTENTS . ;
'i SECTION TITLE PAGE
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( 1 Introduction and Summcry 1 1- l 1.1 Upgraded Fuel Features (VANTAGE 5 Hybrid) 1 1- I 1.2 Increased Pedking Factors 12 1.3 Conclusions 12 2 Design Features- 21 2.1 Introduction 21 I 2.2 VANTAGE SH Fuel Assembly 21 2.3 Intermediate Flow Mixer (IFM) Grid 22. I 2.4 Fuel Rod Performance 22 4 3 Nuclear Design i
.3-1 3.1 Introduction and Summary -31 '
3.2 Increase in FaaN 3-1 > j 3.3 - Increase in Fo Limit '31 !
- 3.4 Methodology :32 .i 4 Thermal and Hydraulic Design 41 4.1 Introduction and Summary . 41 ~ '
4.2 Methodology 41 4.3 Hydraulic Compatibility 42 4.4 Effects of Fuel Rod Bow on DNBR - ' 42-4.5 Fuel Temperature Analysis 43 4.6 Transition Core Effect "j 44 4.7 Thimble Plug Removal 43 ,i 4.8 Conclusion = , 43-5 Accident Analysis ' '51 - 5.0 Introduction 51 i 5.0.1 General -S1 5.0.2 VANTAGE 6H Design Features 5-1 1 5.0.3 Fuel! Core Related Features and Assumptions 53. I 5.0.4 Plant' Tech Spec Related Changes to be implemented 56 i 5.0.5 Additional Assumptions Incorporated into Safety Analysis 57 O D030913:10.082790 -i , d e
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1 TABLE OF CONTENTS (Continued) SECTION TITLE PAGE 5.1 Non LOCA Accident Analysis 5 11 -! 5.1.1 Introduction 5 11 ; 5.1.2 Accidents Analyzed and Evaluated 5-22' i 5.1.3 Increase in Heat Removal by the Secondary Systen: 5 25 j 5.l.4 Decrease in Heat Removal by the Secondary System 5 40 '! 5.1.5 Decrease in Reactor Coolant System Flowrate - 5 57 , 5.1.6 Reactivity and Power Distribution Anomalies 5 CS .j 5.1.7 increase in Reactor Coolant Inventory -5 83 " 5.1.8 Decrease in Reactor Coolant inventory . -5 86 5.2 LOCA Accidents 5 89 l 5.2.1 Large Break LOCA (4 Loop Operation) 5 89- :l 6.2.2 Large Break LOCA (3 Loop Operatien) 5 100' l 5.2.3 Small Break LOCA (4 Loop Operation) 5110) ! 5.2.4 Small Break LOCA (3. Loop Operation) 5 113: g 5.2.5 Blowdown Reactor Vessel and Loop Forces.- 5 114 5.3 Mass and Energy Release and Containment integrity 5119. 5.3.1 LOCA Mass and Energy Release. 5 119- 1 5.3.2 Steamline Break Mass and Energy Release 5 121 l 9 5.3.3 Containment Integrity 5-122' 5.4 Radiologica_I Consequences of Accidents 5 123 5 4.1 Radiological Consequences of Extended Discharge 3 Burnup 5123 1 5.4.2 Radiological Consequences to Fuel and Plant Changes ' 5 i23 ; rs Normal Operation Control System Operation . i and Margin to Trip ~ 5-125 1 5.5.1 Introduction 125 5.5.2 Method of Analysis '5125= 5.5.3, Results 126 9 5.5.4 Conclusion L5130'. 5.6 RWST/ Accumulator Boron Concentration increase Evaluation 5 131. , 5.6.1 Introduction .5 131 .i 5.6.2 Non-LOCA' Safety Analysis 5 131 5.6.3 FSAR LOCA ~ Analysis (10CFR50.46)-.
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5-134 5.6.4 Post Accident Chemical Environme 5-136- '
-5.6.5 Conclusions - 5139 5.7 RCCA' Parked Position - .5-140 5.7.1' Introduction ,
5-140 $ 5.7.2 Analyses 5-140: j 5.7.3 Conclusions , 5 142: o 9 1 D0309' 13:1D/082790 ~ 11 :
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L - TABLE OF CONTENTS (Continued) j SECTION TITLE -PAGE : I 5.8 Instrument Uncertainty Revised Thermal Design - i Procedure (RTDP) -5143 j
' 5.9 Safety Evaluation for VANTAGE SH and '
Asseclated Changes - 5145 c 5.9.1 Background .-5 145 ! 5.9.2 Licensing Basis 5147 5.9.3 Evaluation 5 150' : 5.9.4 Determination of Unreviewed Safety Question .5 154 ! 5.9.5 Conclusion - -5 158-5.10 Accident Analysis Figures. ' 6.0 - References '61 ! O 1 i
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= 1 LIST OF TABLES TABLE TITLE: - PAGE- .;
'21 Comparison of 17x17 Standard (Region 5) and VANTAGE 5 24-Hybrid Fuel Assembly Mechanical Design Parameters !
41 Millstone Unit 3 Thermal and Hydraulic Gasign Parameters . 4 51 l - a 42 DNBR Margin Summary '47 I
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i 5.1.1 1 Nuclear Steam Supply System Power Ratings ' 5 15 l 1 I 5.1.12 Summary of Initial Conditions and Computer Codw, used 5 16 ] E l 5.1.13 Nominal Values of Selected Plant , ;ter 5 20 Utilized in the Non LOCA Safety Analyses = j! A g 5.1.1 4 Reactor Trip Setpoints and TripTime Delays Assumed: 5 21' ' for the Millstone Unit 3 Non LOCA Analyses , R 5.1.2 1 Millstone Unit 3 Non LOCA Events - 5 23 Extent of Consideration for VANTAGE SH Transition- ; 5.1.3 1 Time Sequence of Events for incidents that Result . 5 38 in an increase in Heat Removal by the Secondary System
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_ 5.1.4 1 Time Sequence of Events for Incidents which Result ' 5 51 in a Decrease in Heat Removed by the Secondary, System - ; 5.1.5 1 Time Sequence of Events for Incidents which Result 5 64 in a Decrease in Reactor Coolant System Flowrate 5.1.5 2 Surn w )f Results for Locked Rotor Transients - 5 65, 5.1.6 1 Time Sequence of Events for Incidents which Result - 5 79 in Reactivity and Power Distribution Anomatics = i 5.1.62 Parameters Used in the Analysis of the- 5-82, , Rod Cluster Control Assembly. Ejection Accident i 5.1.71 Time Sequence of Events for incidents which Result. 5 85
. In an increased in Reactor Coolant Inventory th 1
00309-13:10 082790 .iv - ii i -
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LIST OF TABLES (Continu3d) ' i; TABLE TITLE PAGE < 1 5.1.8 1 Time Sequence of Events for incidents which Cause : 5 88 a Decrease in Reactor Coolant Inventory 5.2.11 N Loop Plant Configuration for Large Break LOCA 5-95 ~ !' Analysis ! 5.2.12- N Loop Large Break LOCA Analysis Parameters- 5 96 5.2.13 N Loop Large Break LOCA for VANTAGE SH FuGl- 5 97- ; Time Sequence of Events l 5.2.14 -l' Large Break N Loop LOCA Results Fuel' 5 98 Cladding Data Spectrum of Breaks for VANTAGE SH Fuel-5.2.15 Large Break N Loop LOCA Results Fuel 5 99 Cladding Data - VANTAGE 5H Vs Standard Fuel l 5.2.2 1 N 1 Loop Plant Configuration for Large Break LOCA Analysis- 5105 l 5.2.2 2 N 1 Loop Large Break LOCA Analysis Parameters 5106 i 5.2.2 3 N 1 Loop Large Break LOCA for VANTAGE SH Fuel .5 107 i Time Sequence of Events [! 5.2.2 4 Large Break N 1 Loop LOCA'Results Fuel 5 108-Cladding Data VANTAGE SH Fuel
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5.2.2 5 Larr,e Break N 1 Loop LOCA Results Fuel- 5 109 Cladding Data. VANTAGE SH Vs STD Fuel- - i 5.2.3 1 N Loop Plant Configuration for Small Break LOCA A'n alysis -5 1.15 5.2.3 2 N Loop Small Break LOCA Analysis Parameters- 5 116 i 5.2.3 3 Time Sequence of Events for Incidents which 5 117 q Cause a Decrease in Reactor Coolant loventory - ' 5.2.3 4 Small Break Results 5 118 l 1 /' 0030913:10 V82700 V' ( , f{ 1
l LIST OF FIGURES FIGURE TITLE l 21 Comparison of the 17x17 V5H Fuel Assembly'and the l 17x17 STD ' Fuel Assembly (Page 2 6) 5.1.11 Doppler Power Coefficient Used in the Non LOCA Accident Analysis j 5.1.12 Moderator Density Coefficient Used in the Non LOCA Accident Analysis l
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5.1.13 RCCA Position Versus Time Normalized to Dashpot ! 5.1.14 Normalized RCCA Worth-Versus Fraction Inserted 5.1.15 Normalized RCCA Bank Reactivity Worth Versus Drop Time 5.1.16 Illustration of Overtemperature and Overpower w AT Protection (Four Loop Operation) 5.1.1 7 Illustration of Overtemperature and Overpower AT Protection (Three Loop Operation) G 5.1.18 Illustration of f(al) Penalty . 5.1.31& Feedwater Control Valve Malfunction (Four Loop a 5.1.3.2 Operation) 5.1.31 A & Feedwater Control Valve Malfunction (Three Loop 5.1.3.2A -Operation) j 5.1.33& Ten Percent Step Load increase, Minimum 5.1.3 4 Reactivity Feedback, Manual Reactor Control 4 5.1.3-5 & Ten Percent Step Load increase, Maximum -l 5.1.3 6 Reactivity Feedback, Manual Reactor Control-5.1.3 7 & Ten Percent Step Load increase, Minimum I 5.1.3 8 Reactivity Feedback, Automatic Reactor Control 5.1~.3-9 & Ten Percent Step Load increase, Maximum ., 5.1.3 10 Reactivity Feedback, i / )matic Reactor Control 00309 13:10:082790 vi 1 J
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LIST OF FIGURES (Continued) FIGURE TITLE - j 5.1.3 12 & Failure of Steam Generator Safety or Dump Valve (Four Loop: . 5.1.3 13 Operation) i 5.1.312A & Failure of Steam Generator Safety or Dump Valve (Three Loop 5.1.3 13A Operation) . H
- I 5.1.3 14 Doppler Power Feedback 5.1.3 15, 1.4 Ft2 Steam line Rupture, Offsite Power Available 5.1.3 16 & (Four Loop Operation) 5.1.3 17 i
'i 5.1.3 15 A, 1.4 Ft2 Steam'line Rupture, Offsite Power Available '!
5.1.316A & (Three Loop Operation) . a 5.1.317A u 5.1.41& Turbine Trip with Pressurizer Pressure Contro'l, 5.1.4 2 Minimum Reactivity Feedback (Four Loop Operation) . ; 5.1.41 A & Turbine Trip with Pressurizer Pressure Control, 5.1.4 2A Minimum Reactivity Feedback (Three Loop Operation) 5.1.43& Turbine Trip with Pressurizer Pressure Control, 5.1.4 4 Maximum Reactivity Feedback (Four Loop Operation): 5.1.4 3A & Turbine Trip with Pressurizer Pressure Control,. 5.1.4 4A Maximum Reactivity Feedback (Three Loop Operation) 5.1.45& . Turbine Trip without Pressurizer Pressure Control, . 5.1.4 6 Minimum Reactivity Feedback (Four Loop Operation) 5.1.4-5A & Turbine Trip without Pressurizer Pressure Control, - 5.1.4 6A Minimum Reactivity Feedback (Three Loop Operation) - 1
- . 5.1.47& Turbine Trip without Pressurizer Pressure Control, i
- 5.1.4 8 . Maximum Reactivity Feedback (Four Loop Operation) .
5.1.4 7A & Turbine Trip without Pressurizer Pressure Control, 5.1.4-8A Maximum Reactivity Feedback (Thrco Loop Operation) 9 5.1.4 9 Pressurer Pressure and Water Volarne. Transients for Loss of Normal FeedWater with Offsite Power (Four Loop Operation)
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D030913:1 D.'082790 vil o
'l G FIGURE LIST OF FIGURES (Continued)
TITLE l i 5.1.4 10 Core Average _ Temperature Transient and Steam Generator Water Volume ' Transient for Loss of Normal Feedwater.with Offsite Power (Four Loop : > Operation) , q 5.1.4 11 Pressurize. Pressure ar'd Water Volume Transients for Loss of Normal Feedwater with Loss of Offsite Power (Four Loop Operation) i 5.1.4 u Core AverageLTemperature Transient and Steam Generator Water Volume- ' Transient for Loss of Normal Feedwater with (Loss of Offolte Power (Four. Loop Operation) t 5.1.413 to Main Feedline Rupture with Offsite Power Available (Four Lo'op a 5.1.4 19 Operation) 5.1.413A to Main Feedline Rupture with Offsite Power Available ' 5.1.4 19A (Three Loop Operation) 5.1.4-20 to Main Feedline Rupture Without_Offsite Power (Four.Loopt 5.1.4 26 Operation) ! O 5.1.4-20A to Main Feedline Rupture Without Offsite Power,(Three Loop 5.1.4 26A Operation) - 5.1.51 to Transients for Four Loops Operating, One Pump Coasting Down 1 5.1.5 4 5.1.5 5 to Transients for Four Loops Operating, Four Pumps. Coa' sting Down 5.1.5 8 .j
; 5.1.5 9 to Transients for Fou' cops Operating and One Locked Rotor .
5.1.5-12
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5.1.5 9A to Transients for Three Loops Operating and One Locked Rotor : 5.1.5 12A ' 5.1.'6-1 & Uncontrolled RCCA Bank Withdrawal from Subcritical . 5.1.6 2 5.1.6 3 to Uncontrolled RCCA Bank Withdrawa: '. rom Full Power - 1 5.1.65 at Maximum P. ate of 70 pcm/sec (Four Loop Operation) . 5.1.6 3A to Uncontrolled RCCA Bank Withdrawal from. Full power , 5.1.6 5A at# .num Rate of-70 pcm/sec'(Three_ Loop Operation) . ' O ' 00309-13:10/082700 vill:
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.- LIST OF FIGURES (Continued)- i FIGURE -. TITLE 5.1.6 6 to Uncontrolled RCCA Bank Withdrawal from Full Power at Minimum Rate 5.1.68 of 3 pcm/sec (Four Loop Operation); !
.. 8A of 3 pc ' ec Th ee Loop O erat on) 5.1.6 9 Minimum DNBR vs Reactivity insertion Rate,v .;
RCCA Withdrawal from 100% Rated Power (Four Loop Operation) ; 5.1.6 9A - Minimum DNBR vs Reactivity Insertion Rate, . ^ RCCA Withdrawal from 75% Rated Power (Three Loop Operation) 5.1.6 10 Minimum DNBR vs. Reactivity insertion Rate, RF 'A Withdrawal from 60% Rated P_ower (Four Loop Operation) 5.1.6 10A - Minimum DNBR vs Redctivity insertion Rat _e, RCCA Withdrawal frcm 10% Rated Power (Three Loop Operation)- a 5.1.6 11 Minimum DNBR vs Reactivity insertion Rate,- ACCA Withdrawal from 10% Rated Power (Four Loop Operation) 5.1.6 12 & BOL HFP RCCA Ejection Transients (Four Loop Operation) ! 5.1.6 13 5.1.612A & BOL HFP'RCCA Ejection Transients (Three Loop Operation) -! 5.1.613A L 5.1.6 14 & EOL HZP RCCA Ejection Results (Two RCPs. Running).- 5.1,6 15 5.1.71 to Inadvertent Operr. tion of ECCS During Power Operation (Four 5.1.' 3 Loop Operation) 5.1.81& Inadvertent Opening of a Pressurizer Safety' Valve (Four Loop-5.1.8-2 Operation) 5.1.81 A & Inadvertent Opening of a Pressurizer Sarety Vane (Three Loop 5;1.8 2A - Operation) i 15.5 1 Steady State Margin to OTAT Trip Versus al O J ix d
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I l ! i i LIST OF FIGURES (Continued) FIGURE TITLE
-l 15.6 5 Code Interface Description for Large Break Model 15.6-6 Code Interface Description for Small Break Mooel- 1 15.6 Small Break Power Distribution 15.6-8 Peak Clad Temperature DECLG (Co = 0.6) 15.6-8A Core Pressure - DECLG (Co = 0.6) d I
15.6 9 Downcomer and Core Water Levels' During Reflood DECLG (Co = 0.6')
'15.6 10 Core Inlet Velocity During Reflood DECLG (Co:= 0.6) 15.6 11 Core Power Transient DECLG'(Co _= 0.6)_
15.6 12 Core Flow Top and Bottom DECLG (Co = 0.6)- 15.6 . Core Heat Transfer Coefficient DECLG (Co = 0,6)1 i J 15.614 Fluid Temperature DECLG (Co = 0.6), ! j 15.6 15 Break Mass Flow Rate DECLG (Co: = 0.6)' i 15.6 16 Break Energy Release Rate DECLG (Co = 0.6)- 15.6 17 Fluid Quality DECLG (Co = 0.6) ' 15.6 18 Accumulator Flow During Blowdown DECLG (Co = 0.6): 15.6 19 Mass Velocity DECLG (Co = 0.6) 15.6-20 Pumped ECCS Flow During. Blowdown DECLG (Co = 0.6); I 15.6 21 Peak Olad Temperatu'r e DECLG (Co .= 0.8)
.15.6 22 -Core Pressure-DECLG (Co = 0.8) 15.6 23 Downcomer and Core Water Levels During Reflood DECLG (Co =.-0.8)' l 15.6 24 , Core Core inlet Velocity During Reflood DECLG (Co_ = 0.8)-
1
.I ,'00309-13:10082790 x i ,
' i LIST OF FIGURES (Continued)'
FIGURE TITLE l 15.6 25 Core Power: Transient DECLG'(Co = 0.8) 15.6 26 Peak Clad Temperature DECLG (Co = 0.4) . i
.15.6 27 Core Pressure DECLG (Co.= 0.4)
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15.6 28 Downcomer and ' Core Water. Levels During Reflood DECLG (Co =: 0.4) - d 15.6 29 Core Inlet Velocity During Reflood DECLG (Co =- 0.4) . ,
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15.6 30 Core Power Transient DECLG (Co' =' 0.4) 15.6 31 Reactor Coolant System Depressurization-Transient (3 inch Break) 1 15.6 32 Core Mixture Level (3 inch Break) 15.6 33 Clad Temperature Transient (3 Inch Break) 4 15.6 3 t Core Exit Steam Flow (3 inch Break)
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15.6 35 Rod Film Heat Transfer Coefficient (3 Inch Break)- 1 15.6 36
'i Hot Spot Fluid Temperature (3 Inch Break) l q
15.6-37 Reactor Coolant System Depressurization Transient (2 Inch ~ Break) l' 15.6 38 Reactor Coolant System Depressurliation Transient'(4 inch' Break) ' 15.6 39 Core Mixture Level (2 inch Break) l 15.6 40 Core Mixture Level (4 Inch Break) ' 15.6 41 Clad Temperature Transient (2 Inch Break) - 15.6 42 Clad Temperature Transient (4 inch Break) 15.6 43 Safety injection Flowrate i 15.6-44 Hot Spot Clad Temperature (Active Loop Break Co = 0.4, N 1 Loop)- i5.6-45 Coolant Pressure in Reactor Core (Active Loop Break.Co = 0.4,- O- N 1: Loop) .
-00309-13:1o?082700 xi i., r ,
q I, LIST OF FIGURES (Continued) 1 FIGURE TITLE
-I 15.6 46 Water Levelin the Core and Downcomer During Reflood i pctNe Loop Break Co-0.4, N 1 Loop) !
t 15.6 47 Core Reflooding Rate (Active Loop E reak Co = 0.4, N 1 Loop) 'I q 15.6 48 Thermal Powr During Blowdown (Actue Loop Greak Co = 0.4, N 1 Loop) 'i y 15.6 49 Core Flow During Blowdowniinlet and Outlet (Active Loop Break-Co = 0.4, N 1 Loop) i q 15.6 50 Core Heat Transfer Coefficients (Active Loop Break Co = 0.4, N 1. !, Loop) { 15.6 51- Hot Spot Fluid Temperature (Active Loop Break Co = 0.4, N 1 Loop) 'i 15.6 52 Mass Released to Containment During Blowdown (Active Loop Break - 4 Co = 0.4, N 1 Loop)1 i 15.6 53 Energy Released to Containment During Blowdown (Active Loop Break - , Co = 0.4, N 1 Loop) < 1 15.6 54 Fluid Quality in th Mot Assembly During Blowdown (Active-Loop Break Co - 0.4, N 1 Loop) ~ l 15.6 55 Mass Velocity During Blowdown (Active Loop Break Co = 0.4 N 1 Loop)
... . . y 15.6 56 Accumulator Water Flow Rate During Blowdown-(Active Loop Break.
l Co = 0.4, N 1 Loop) ; i 15.6 57 Pumped Safety injection Water Flow During Reflood (Active Loop- .; Break Co = 0.4, N 1 Loop)- -! 15.6 58 Hot Spot Clad Temperature (Active Loop Break Co = 0.6, N 1 Loop) i L 15659 Coolant Pressure in the Reactor Core (Active Loop Break Co = 0.6,. L N 1 Loop). ' i f[i,
.15.6 Water Level in the Core and Downcomer During Reflood (Active Loop .
Break Co = 0.6, N 1 Loop) : e
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1..~ f' .(Q:gjf;. ;. 7g g LIST OF FIGURES (Continued) - , 7-FIGURE TITLE 15.6-61 Core Reflooding Rate (Active Loon Break Co = 0.6, N 1 Loop) 15.6-62 Thermal Power During Blowdown _(Active Loop B'reak 00 0.6, N 1 Loop) s' O 0 1 00309 13: 10.082790 xiii >> i '
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- 1. INTRODUCTION AND
SUMMARY
The Millstone Generating Station Unit 3 plans to refuel and operate with upgraded Westinghouse fuel features, and increased peaking factors and the capability of operating with up to all of the assembly thimble tube plugs being removed. ~This report summarizes the safety evaluations that were performed to confirm the acceptable use of these options for four loop and three loop operation. Sections 2.0 through 5.0 of the Plant Safety Evaluation (PSE) provide the results of the Mechanical, Nuclear, Thermal and Hydraulic, and Accident Analysis, respectively. The Millstone Unit 3 Plant Safety Evaluation is to serve as a reference safety evaluation / analysis report for the region by region reload transition from the Millstone Unit 3 Cycle 3 core with 17x17 standard (STD) fuel assemblies to a core containing the VANTAGE SH upgraded features described in Section 1.1. Thus, the PSE will be used as a basic reference document in support of future Millstone Unit 3 Reload Safety Evaluations (RSEs) for upgraded fuel reloads. The PSE utilizes the Westinghouse standard reload methodologyW.(Consistent with this methodology, parameters are chosen to maximize the applicability of the PSE evaluations for future cycles. The objective of subsequent cycle specific RSEs will be to verify that applicable safety limits are satisfied based on the reference evaluation / analyses established in this safety evaluation. The transition to and all VANTAGE SH core evaluations / analyses in the PSE were performed at a reactor thermal full power levul of 3411 MWt (100% rated power) for four loop operation and 2560 MWt (75% rated power) for three loop operation. The following conservative assumptions were made in the safety evaluations / analyses: 10% uniform steam generator tube plugging
- with RCS thermal design flow rates of 378,400 gpm and 294.900 gpm for four and three loop operation respectively, core bypass flow increases from 6 to 8.6% to account for complete thimble plug removal and the use of IFMs, a positive moderator temperature coefficient (PMTC) of + 5 pcm/'F from 0 to 70% power-and then decreasing linearly to O pcm!'F between 70 to 100% power, the use of RAOC power distribution control for both 3 and 4 loop operation, an' d increases in Fo and Fan as noted in Section 1.2. For 4 loop operation, a vessel / core inlet coolant temperature of 557.0'F and a vessel coolant average temperature of 587.1*F were used. For 3 loop operation, a vessel / core inlet coolant ternperature of 550.2*F and a vessel coolant average temperature of 579.6'F were used.
1.1 UPGRADED FUEL FEATURES (VANTAGE 5 HYBRID) Millstone Unit 3 Cycle 4 and subsequent core loadings will have fuel assemblies that incorporate the low pressure drop Zircaloy grid and the Intermediate Flow Mixer (IFM) Assumes 10% of steam generator tubes in each generator are plugged and 1 corresponds to the worst plugging level of any steam generator. D030913:10 082790 11
, 4 ,
l grid. This upgraded fuel feature is known.as VANTAGE 5 Hybrid (VANTAGE 5H) with IFMs and has been submitted as an Addendum
- to the " VANTAGE 5 Reference Core Report," WCAP 10444 P A*. VANTAGE SH has received generic NRC approvalW. A' '
brief summary of those upgraded fuel features is presented in Section 2 of this report.' 4 In addition to the above mentioned VANTAGE SH design features, Millstone Unit 3 reloads will also contain several VANTAGE 5 design features? and other upgraded fuel design-features used in the Cycle 3 Core. - Millstone Unit 3 Cycle 3 Region 5 fuel has already-incorporated the VANTAGE 5 Integral Fuel Burnable Absorbers (IFBAs), the VANTAGE 5 1 Axial Blanket design the VANTAGE 5 Extended Burnup design, the VANTAGE 5 i Reconstitutable Top Nozzle (RTN) design feature, as well as Debris Filter Bottom Nozzles :! (DFBNs), snag resistant grids and standardized fuel pellets.- 1 1 1.2 INCREASED PEAKING FACTORS j Beginning with Cycle 4 the future cycles of operations for Millstone Unit 3 will use ' increased FagN and Fa peaking factors. The full power FaHN' peaking factor design limitI l will increase from the current value of 1,55 to -1.70. The maximum 4 I . 3 Fo peaking . j factor limit willincrease from the current value of 2.32 to 2.60. The ma-mum 3 loop Fo ' peaking factor limit willincrease from the current value of 2.25 to 3.0 at the maximum 75% rated power. The K(Z) envelopes will be modified. These changes will permit more { flexibility in developing fuel management schemes (i.e., longer fuel cycles," improvement of ; fuel economy and neutron utilization, vessel fluence reduction).
1.3 CONCLUSION
S l The results of evaluation / analysis described herein lead to the following conclusions:,
- 1. The Westinghouse fuel assemblies containing VANTAGE SH and the additional upgraded fuel features for Millstone Unit 3 are mechanically compatible with the !
current fuel assemblies, control rods, and reactor internals interfaces. The. current design bases for Millstone Unit 3 have been changed as described in this report to accommodate the VANTAGE SH design. t
. 2. Changes in the nuclear characteristics due to the transition to upgraded fuel will- !
be within the range normally.seen from cycle to cycle due to fuel management i
- 3. The reload upgraded fuel assemblies are hydraulically compatible with^the fuel l assemblies from previour .eload cores.
- 4. The changes in the design full power FaHN limit from 1.55 to 1.70 (with ,
appropriate treatment of uncerti .ples) is supported by design basis safety ' analyses summarized in this evaluation.
- 5. The changes in the full power maximum Fo limit from 2.32 to_2.60 for 4 loop operation and from 2.25 '.o 3.0 for 3 loop' operation at 75%~ rated power, and 1
00309-13:1D.082700 -12 i m__ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _i__
O modification to the K(Z) envelopes are supported by the design basis safety analyses suminarized in this report? G. Evaluation / Analyses have shown that all or'any combination and pattern of thimble tube plugs may be removed from the Cycle 4 Core and subsequent cores.
- 7. The core design and safety. analyses re.sults documented in this report show-the core's capability for operating safely at the 100% rated Millstone Unit 3 J design thermal power, for 4 loop operation and 75% rated power for 3 loop -
operation with up to 10% steam ~ generator tube plugging in each steam. generator.
- 8. This report establishes a reference upon which to base Westinghouse reload ~
safety evaluations for future reloads with the upgraded fuel features and increased peaking factor limits; O
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- 2. DESIGN FEATURES
2.1 INTRODUCTION
k The mechanical design of the VANTAGE 5H upgraded fuel asse' .blies for Millstone Unit 3 is the same as the Region 5 reload fuel assemblies in the Cycle core, except that the upgraded fuel assemblies will incorporate two major fuel design 1provements. These improvements are the VANTAGE SH Zircaloy grid, and Intermediate Flow Mixer (IFM) grids. The design changes are described in more detailin the following sections. Design features previously introduced, which will continue to be used in the upgraded fuel assemblies, include the Reconstitutable Top Nozzle, Debris Filter Bottom Nozzle, (DFBN), snag resistant grids, standardized fuel pellets, Integral Fuel Burnable Absorbers, Axial Blankets and extended burnup design. These design features have beer. used in the Cycle 3 core, and, therefore, they are not discussed in the remainder of Section 2.0. For Cycle 4 and subsequent cycles, as a DFBN minor modification, skirts have been added to the bottom periphery of the plate between the nozzle legs, as shown in Figure 2-1, to strengthen the plate and legs. m.2 VANTAGE SH FUEL ASSEMBLY The VANTAGE SH fuel assembly design evolved from the current VANTAGE 5. Optimized Fuel Assembly (OFA) and Standard (STD) fuel assembly designs. It is based on a 9 substantial design and operating experience. Design features from each of these previous designs are incorporated into the VANTAGE SH fuel assembly design. The VANTAGE SH design is characterized by the use of Zircaloy grids with 0.374 inch OD standard fuel rods. To accommodate the Zircaloy grids, the VANTAGE SH thimble tube
.___ diameter was modified to be the same as the 17x17 OFA or VANTAGE 5 fuel. A l
_ comparison of the STD and VANTAGE SH fuel assembly design parameters is given in Table 2-1. Figure 21 demonstrates the similarity of the two designs and shows a i
, comparison of overall dimensions.
Hydraulic tests on the VANTAGE SH (with IFM's) fuel assembly indicate that the extent of the overall fuel assembly and individual component pressure drop differences, relative to the 17x17 STD, OFA, or VANTAGE 5 (without IFM's) fuel assemblies, is comparable to or withir- a bounds of previous variances identified in the VANTAGE 5 program. Thus, prevt , analyses and tests performed as part of the VANTAGE 5 program demonstrate ,
= '
that ruel rod contact wear with spacer grids will be within the allowable design limits. The major components that determine toe structuralintegrity of the fuel assembly are the grids. Mechanical testing and analysis of the VANTAGE SH Zircaloy grid and fuel assembly have demonstrated that the VANTAGE SH structural integrity under seismic /LOCA loads will provido margins comparable to the STD 17x17 fuel assembly design and will meet all design bases
- 0 l 1
00309-131 D 082790 2-1 n _
l ! ! 1 { p The VANTAGE SH Zircaloy grid is based on the OFA Zircaloy grid design and operating ( experience. The grid strap thickness, type of strap welding, basic mixing vano design and pattern, method of thimble tube attachment, type of fuel rod support (6 point), material and j I envelope are identical to the OFA Zircaloy grid. This evaluation of the VANTAGE SH grid j performance is based on the extensive design and irradiation experience with previous . l grid designs and full grid testing completed with the VANTAGE SH grid designi e I l In order to demonstrate early performance of the Zircaloy grid design, fuel assembly ' l- demonstration programs were conducted inserting OFA fuel assemblies containing Zircaloy grids into 14x14,15x15 and 17x / cores. Subsequent to the satisfactory performances observed in these programs, the OFA with Zircaloy grids were loaded and have operated successfully since the early 1980's in many Westinghouse cores (5'; ; i 2.3 INTERMEDIATE FLOW MIXER (IFM) GRID , i Mi'! stone Unit 3 Cycle 4 and subsequent reloads will contain fuel assemblies that incorporate IFM grids as well as the VANTAGE SH Zircaloy grids described in the previous section. The IFM feature is currently part of the licensing basis in other plants'5) . and meets all fuel assembly and fuel rod design criteria. l The IFM grid in the VANTAGE SH assemoly is an adaptation of the existing VANTAGE 5 ' l IFM grid design to a 0.374 inch OD standard fuel rod. As shown in Figure 21, IFM's are p located in the three uppermost spans between the VANTAGE SH Zircaloy structural grids I g to promote flow mixing in the hottest fuel assembly spans. The IFM's are fabricated of ' Zircaloy in the same manner as the VANTAGE SH Zircaloy grid but are not intended to be . structual members. The IFM grid envelope is slightly smaller than the VANTAGE SH + Zircaloy grid. Each IFM grid cell provides four (4) point fuel rod support. The simplified cell arrangement allows tho IFM to accomplish its flow mixing objective with minimal pressure drop. 2.4 FUEL ROD PERFOPMANCE The VANTAGE SH fuel rods are designed to satisfy the Millstone Unit 3 FSAR fuel rod. , design bases for the fuel's irradiation life. These same bases are applicable to all fuel l rod designs, including Westinghouse STD fuel rod designs, with the only difference being - that VANTAGE 5H fuelis designed to achieve a higher burnup and to operate with a higher FaH limit. The design bases for Westinghouse VANTAGE SH fuel are discussed in Reference 2. Considering the potential impact on fuel rod designs, there is no effect having fuel with more than one type of geometry simultaneously residing in the core during the transition cycles. The mechanical fuel rod design evaluation for each region incorporates all appropriate design features of the region, including any changes to the fuel rod or pellet geometry from that of previous fuel regions (such as the presence of IFBA fuel rods, axial-p blankets or changes in the fuel rod and plenum length)._ Analysis of IFBA rods includes. _Q any geometry changes necessary to model the presence of the burnable boride absorber,
- 00309 13:1D'082790 22 ,
~~ ~
j p and conservatively models the helium gas release from the thin boride absorber coating, 4 Fuel performance evaluations are completed for er.ch fuel region to demonstrate that the ; design criteria will be satisfied for all fuel rod types in the STD and VANTAGE SH assemblies under the planned operating condPaons for each reload core. Any changes from the plant operating conditions originally evaluated for the mechanical design of a fuel -l region (for example, any increase in burncp cr peaking factors) are addressed for all
- affected fuel regions when the plant change is to be implemented. ;
Fuel rod design evaluations for the Millstone Unit 3 transbon core fuel are performed l using the NRC approved models in References 6 and 7 and the NRC approved extended
- burnup design methods in Reference 3 to demonstrate that all of the FSAR fuel rod design l bases are satisfied. i i
The VANTAGE SH Upgrade fuel rod analysis includes the following: an increase in the full power FaHN limit from 1.55 to 1.70 RAOC power distribution control for 3 loop and 4 loop operation, twenty four month fuel cycles, low leakage loading ": patterns, and a range of backfill pressures for VANTAGE SH IFBA and non IFSA rods which include top and bottom 6 inch axial blankets (Natural UO2 ). l OV 5 n
/
l u . D030913:10 082790 23 .
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H Table 21 Comparison of 17X17 Standard (Region 5). and Vantage 5 Hybrid Fuel Assembly-- Mechanical Design Parameters f Standard (Realon 5) VANTAGE 5H
- t Fuel assembly Weight,1b. 1467 1470 l Fuel Assembly Overall Length, inch 159.9 160.0 Fuel Rod Overall Length, inch 152,16- 152.2'
- l. Active Fuel Stack Height, inch 144.0- 144.0 l
Assembly Envelope, inches 8,426 8.426 Fuel Rod Pitch, inch 0.496 0.496 , I Number of Fuel Rods' Assembly 264 264= '! Number of Guide' Thimbles Assembly 24 24 Number of insirumentation Tube Assembly 1 1 Fuel Tube Material Zirc 4 Zirc.41 $ Fuel Tube Clad OD, inch 0.374 0.374 Fuel Rod Clad Thickness, inch 0.0225 0.0225 Fuel Clad Gap, mil (uncoated pellets) 6.5' 6.5 j Fuel Pellet diameter, inch (uncoated pellets) .3225 0.3225 j Fuel Pel!et Length Enriched fuel, inch 0.387 0.387 s Unenriched fuel, inch 0.545 -0.462 l Guide Thimble Material Zirc 4 Zirc 4 ; Guide Thimble 00, inch 0.482' .0.474 , Guide Thimble Wall Thickness, inch 0.016 0.016 i Grid Material, inner Mid Grid (6) Inconel ' Zirc 4' Edges Modified .No Yes Grid Material. End Grids (2) Inconel' incon'el . j IFM Material (3) None 'Zirc 4 i l l 1 1 6 v D0309-13:10 091490 24 Rev.1 - 9/90' f
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1 Table 21 (Contiriued) : Comparison of 17X17 Standard (Region 5) j 1 and Vantage 5 Hybrid Fuel Assembly; J ' Mechanical Design Parameters Standard (Reolon 5)- VANTAGE SH .f Grid Types Utilized Inconel Mid grids Yes No .'i Zircaloy Mid grids No Yes' Inconel Top & Bottom Grids. Yes Yes No- Yes i Flow Mixers Inner Spring (Mid Grids)- Vertical Non vertical ;l
.i Grid Fabrication Inconel Grids Brazed joining of '
Interlocking
~
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stamped straps. stamped straps L.aser weld joining of = i Zircaloy Mid Grid None- interlocking stamped ctraps ' ! t Grid / Guide Thimble Attach, i Inconel Grids ' Thimbles bulged- Thimbles bulged .- { together with sleeve together with sleeve prebrazed prebrazed I L l I Thimbles bulged .
; together with- . Zircaloy Mid -
None , sleeves laser ; Grids'and Flow Mixers prewelded.to grid ; straps Top Nozzle Reconstitutable - Reconstitutable - stainless steel ~ stainless steel reduced height reduced heightJ removable design , removable design J Bottom Nozzle 304 SS 304 SS . t Reconstitutable Reconstitutable Debris Filter" ; Debris Filter" Compatible with Fuel Handling Equipment - Yes- .Yes. i
- VANTAGE SH design in Reference 2 did not have a d 3bris filter ' '
s 00309 13:10 082790 25 i
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= 152.2 O 1
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= 152.I60 O i
MlHHHHHB1 :l Figure 2-1 an he 17x17 S Fue Ass m y -
. 0p30912:10/081490 26
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l l l Q 3. NUCLEAR DESIGN l V
3.1 INTRODUCTION
AND
SUMMARY
The effects of using upgraded Westinghouse fuel features and increasing the core , peaking factor limits on the nuclear design bases and methodologies for Millstone i Generating Station Unit 3 are evaluated in this section. 1 The grid material and grid volume of the VANTAGE SH fuel are different than that of 17x17 standard fuel assemblies. The effects of these changes on core physics characteristics are small and are explicitly modeled in the neutronics models. The specific values of core safety parameters, e.g., power distributions, peaking factors, rod worths, are primarily loading pattern dependent. The variations in the loading pattern dependent safety parameters are expected to be typical of the normal cycle to cycle variations for the standard fuel reloads. In addition, the present Millstone spent fuel pool criticality analysis is applicabla to the upgraded Westinghouse fuel features, including the use of VANTAGE SH fuel. The increase in peaking factor limits allows more flexibility when developing the fuel management scheme. Specific items concerning the evaluation are given in Sections 3.2 and 3.3. p In summary, the change from the current standard fuel core to a core containing the V upgraded fuel product will not cause changes to the current Millstone Unit 3 FSAR nuclear design bases. However, the design bases will be modified due to the increases to the peaking factor limits Nuclear design methodology is not affected by the use of upgraded fuel features or increased peaking factors. 3.2 INCREASE IN F 3gN The limit on the nuclear enthalpy rise _ hot channel factor, FagN . will be; FaHN= 1.70 (1 + 0.3(1 P))- where P = THERMAL POWER / RATED THERMAL POWER. The increase in the nuclear enthalpy rise hot channel factor lin'it will allow additional l flexibility for fuel management and for determining core loading patterns. The new limit is i applicable to standard and VANTAGE SH fuel for both 4 loop and 3 loop operation. 3.3.lNCREASE IN Fo LIMIT The 'imit on the heat flux hot channel factor, Fo, will take the following form in the i Technical Specifications: O V 00309-13:10 082790 31 i
( - Four Loop Operation O P Fo(z) s (2.60!P) x (K(z)) for P > 0.5, and i Fo(z) s (5.20) x (K(z)) for P - s 0.5 Three Looo Operation q Fo(z) s (2.25/P) x (K(z))- for P > 0.375, and ! Fo(z) s (6.0) x (K(z)) for.P s 0.375 t where P = THERMAL POWER / RATED THERMAL POWER, and > K(z)= The function obtained from the Millstone Unit 3 Core Operating Limit Report: for a given core height location. ! J The increased heat flux hot channel factor limit, Fo, will allow additional flexibility in ' fuel management and core operation as well as accommodate the increased nuclear enthalpy rise hot channel factor limit and use of fut that incorporates axial blankets. With the longer cycles and the necessary higher enrichments that will be used in the future, the radial peaking factor (FagN discussed in Section 3.2) willincrease._ This increase will' result in higher total peaking factors in the Millstone Unit 3 core; Fuel using natural Os uranium axial blankets may also result in~ slight increases in the axial component of the total peaking factor. However, this results in minimal peaking factor increases,'since I integral Fuel Burnable Absorbers (IFBAs) are also incorporated into the fuel design to ' I reduce axial peaking factors. l 3.4 METHODOLOGY i No changes to the nuclear design philosophy or methods are necessary because'of the. I upgraded fuel product or the use of increased peaking factors. The reload design o philosophy includes the evaluation of the reload core key safety parameters which comprise the nuclear design dependent input to the FSAR safety evaluation.for each ( reload cyclem. These key safety parameters will be evaluated for each Millstone Unit 3 reload cycle. If one or more of the parameters fall outside the' bounds assumed in the safety analysis, the affected transients will be re evaluated and the results documented in the RSE for that cycle. f 4 l The 0.374 inch diameter fuel rod has had extensive nuclear design and operating ' experience with the current Millstone Unit 317x17 STD fuel assembly design. The Zircaloy grid material has also had extensive nuclear design and operating experience with the current 17x17 VANTAGE SH,17x17 VANTAGE 5 and 17x17 OFA fuel assembly designs. These changes'have a negligible effect on the use of standard nuclear design- ~ analytical models and methods to accurately describe the neutronic behavior of the .
.A VANTAGC SH fuel- 1 00309-13:1D.082700' 32 a
, . 5
. . ~ - -7 Analyses of Millstone Unit 3 reload cores are performed in accordance with standard reload methodologyW to ensure compliance with the new peaking factor limits.
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G 4. THERMAL AND HYDRAULIC DESIGN
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4.1 INTRODUCTION
AND
SUMMARY
This section describes the calculational methods used for the thermal hydraulic analysi_s, the DNB performance, and the hydraulic compatibility during the transition from mixed fuel cores to an all VANTAGE SH core. Based on minimal hardware design differences and prototype hydraulic testing of the fuel assemblies, it is concluded (2) that the STD and VANTAGE SH fuel assembly designs are hydraulically compatible. Table 41 summarizes the thermal hydraulic design parameters for Millstone Unit 3 that were used in this analysis. The thermal hydraulic design for the upgraded fuel product was analyzed for an increase in the design limit value for the nuclear enthalpy rise hot channel factor (FAHN) from 1.55 to 1.70. This increase is achieved by removing unnecessary conservatism from the analysis through use of improved methods and DNB correlation as described in the following section. The thermal hydraulic design criteria and methods remain the same as those presented in the Millstone Unit 3 FSAR with the exceptions noted in the following sections. All of the current FSAR thermal hydraulic design criteria are satisfied. , 4.2 METHODOLOGY The existing thermal hydraulic analysis of the 17x17 STD fuel used in the Millstone Unit 3 plant is based on the standard thermal and hydraulic methods and the W 3 (R-Grid) DNB A correlation as described in the Millstone Unit 3 FSAR. The DNB analysis of the core l V containing both 17x17 STD and VANTAC'I SH fuel assemblies has beea mcdified to incorporate the WRB 1 and WRB 2 DNB correlations (3.9), the Revised Thermal Design Procedure (RTDP)(10), and an Improved THINC IV Modeling(11). .The W 3 correlation and standard methods are still used when conditions are outside the range of the WRB 1 or l WRB 2 DNB correlation, and of the RTDP. The WRB 1 DNB correlation is based entirely on rod bundle data and takes credit for the
, significant improvement in the accuracy of the critical heat flux predictions over previous DNB correlations. The approval of the NRC that a 95/95 limit DNBR of 1.17 is appropriate-for the 17x17 STD fuel assemblies has been documented (12).
As documented in VANTAGE SH Fuel Assembly Report (2), the use of the WRB 2 DNB correlation with a 95/95 limit DNBR of 1.17 is applicable to the VANTAGE SH fuel assembly. l The W 3 correlation with a 95/95 limit DNBR of 1.30 is used below the STD and VANTAGE SH fuel assembly first mixing vane grid. The W 3 correlation with a 95/95 limit DNBR of 1.45 is used for steamline break analyses in the pressure range of 500 to 1000 pslaW.' With RTDP methodology, variations in plant operating parameters, nuclear and thermal parameters, fuel fabrication parameters, and DNB correlation predictions are considered
,. statistically to obtain the overall DNBR uncertainty factor which is used to define the design limit DNBR that satisfies the DNB design criterion. This criterion is that the 00309 13:10 082790 41 i
O probability that DNB will not occur on the most limiting fuel rod is at least 95% (at 95% !' confidence level) for any Condition I or ll event. Conservative uncertainty values used to calculate the design limit DNBR are given in WCAP 12621, as described in Section 5.8, Since the uncertainties are all included in the uncertainty factor, the accident analysis is ' done with input parameters at their nominal or best estimate values. RTDP analyses use - a new flow parameter, minimum measured flow (MMF), equal to thermal design flow (TDF) ; plus a flow uncertainity. Analyses by standard methods continue to use TDF, l The improved THINC IV modeling scheme improves the accuracy of the solution by minimizing the inaccuracies which result from the use of the perturbation technique in the solution of the governing equations, i For this application, the design limit DNBR values for 17x17 STD fuel are 1.25 for typical cells and 1.24 for thimble cells for both 4 and 3 loop operation. The design limit DNBR values for VANTAGE SH are 1.26 for typical cells and 1.24 for thimble cells for both 4 and 3 loop operation. For use in the DNB safety analyses, the limit DNBR is conservatively increased to provide DNB margin to offset the effect of rod bow, transition core and any other DNB penalties that may occur, and to provide flexibility in design and operation of the plant. For 4 Loop operation, the STD fuel safety analysis limit DNBR values _of 1.27 for typical cells and 1.26 for thimble cells are employed in the analysis, and for VANTAGE.5H fuel, values of 1,69 for typical cells and 1.65 for thimble cells are used. For 3 loop l operation, the STD fuel safety analysis limit values used are 1,39 for typical cells and 1.38 O for thimble cells, and for VANTAGE SH fuel, values of 1.75 for typical cells and 1.72 for d thimble cells are used. Table 4 2 summarizes the available DNBR margin for Millstone Unit 3. 4.3 HYDRAULIC COMPATIBILITY The STD fuel assembly and VANTAGE SH designs have been shown to be hydraulically compatible in the VANTAGE SH Fuel Assembly Report (2). 4.4 EFFECTS OF FUEL ROD BOW ON TR The phenomenon of fuel rod bowing must be accounted for in the DNBR safety analysis of ; Condition I and Condition 11 events. j in the IFM region of a VANTAGE SH assembly, the grid to grid spacing is approximately 10 inches compared to approximately 20 inches in the current fuel assemblies in the Cycle 3 core. Using the rod bow topical report methods, Reference 13, and scaling with the NRC approved factor (L2/EI) results in predicted channel closure in the limiting spans of less than 50% closure; therefore, no rod bow DNBR penalty is required in the 10_ inch spans in the VANTAGE 5H safety analyses. In the VANTAGE SH (ower assembly spans, and in the other fuel assembly spans, rod bow is accounted for in available DNBR margin Q O as summarized in Table 4 2. D030913:10 082790 42
l l p The maximum rod oow penalties accounted for ni the design safety analyses are based-() on an assembly average burnup of 24,000 MWD /MTU, which has NRC approval via L Reference 14. At burnups greater than 24,000 MWD /MTU, credit is taken for the effect of ) FaH burndown, due to the decrease in fissionable isotopes and buildup of fission product >
~
inventory. No additional rod bow penalty is required. 4.5 FUEL TEMPERATURE ANALYSIS There is no dibence in the fuel temperatures used in the safety analysis calculations betw < n the VANTAGE SH fuel and the other fuel assemblies. l 4.6 TRANSITION CORE EFFECT ~! l The VANTAGE SH fuel assembly has IFM grids located in spans between mixing vane grids, whereas no IFM grid exists in the STO assembly. The additional grids introduce ; localized fiow redistribution from the VANTAGE SH fuel assembly into the STD assembly at the axial zones near the IFM grid position in a transition core. Between the IFM grids, flow returns to the VANTAGE SH fuel assembly. The localized flow redistribution described above actually benefits the STO assembly. Thus, the analysis for a full core of STD fuel assemblies remains appropriate for that fuel type in a transition core. Transition cores are analyzed as if they were full cores of one assembly type (full STD or l A full VANTAGE SH), applying the applicable transition core penalty. For VANTAGE 5 fuel, V penalties are a function of the number of VANTAGE 5 fuel assemblies in the core, as per , Reference 15, which has been soproved by the NRC. The same methodology is used to calculate VANTAGE SH transiti'.., core penalties. The DNBR penalty is less than 12.5%. 4.7 THIMBL5! PLUG REMOVAL Detailed evaluations have shown that the main effect of thimble plug removalis the increase in core bypass flow which is reflected in Table 41. This increase has been incorporated into the non LOCA and LOCA safety analyses that have been performed in support of the VANTAGE SH/STD fuel transition cores. I Based on the assessment of the impact of the thimble plug removal on system and-component structural adequacy and core plant safety, it is concluded that it is acceptable to remove all or any combination of these devices in any pattern from the Millstone Unit 3. - core.
4.8 CONCLUSION
The thermal hydraulic evaluation of the fuel upgrade and peaking factor increase for Millstone Unit 3 has shown that 17x17 STD and VANTAGE SH fuel assemblies are d hydraulically compatible and that the DNB margin gained through use of the RTDP methodology and the WRB 1 and WRB 2 DNB correlations is sufficient to allow an O
.s increase in the design FaHN from 1.55 to 1.70. More than sufficient DNBR margin in the o0309 13.10 082790 43
- i safety limit DNBR exists to cover any rod bow and transition core penalties. All current-thermal hydraulle design criteria are satisfied.
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' 00309 13:10 082790 44
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.O Table 41- !
I s Millstone Unit 3 Thermal and Hydraulic Design Parameters , Thermal and Hydraulic Desion Parameters Design Parameters. l l
> 'i 4 Loop 3 Loops -
Reactor Core Heat Output, MWt '3 111 2,560 , Reactor Core Heat. Output,106 BTU /Hr 11,639 8,735 ; Heat Generated in Fuel, % 97.4 97.4 Core Pressure, Nominal, psia 2250 2250 1 Radial Power Distribution
- 1,70 [1 + 0.3(1 P)) 1.70[1 + 0.3(1 P)]
DNB Correlation (STD) WRB1 ~ WRB 1 l l (V5H) WRB 2 WRB 2 W 3- W3" : HFP Nominal Coolant Conditions i Vessel Thermal Design Flow pg Rate (including Bypass),106 lbm/hr - GPM 140.8 ~ 378,400 110.8 294,900 Core Flow Rate" , (excluding Bypass, based on TOF) 106 lbm!hr 128.7 101.3 GPM 345,858 269,539
-i l
Core Flow Area, ft2 (STD) 51.1 51.1 (V5H) 51.3 51.3 , Core inlet Mass Velocity, ! l 106 lbm/hr ft2 (Based on TDF) (STD)- 2.52 1.98 -l ' l (V5H) 2.51- 1.98 l 1.70 represents 1,64 plus 4% measurement uncertainty. Based on design bypass flow of 8.6% without thimble plugs. Based on 75% of rated thermal power. _ W 3'Is used for conditions outside the range of applicability of the WRB 1 or WRB 2 , correlation. ; 00309-13:10 082790 - 45- i _ _ _ . ._ . i -. ~ - ,
k Table 4-1 (Continued)- O Millstone Unit 3 Thermal and Hydraulic Design Parameters
-l Thermal and Hydraulic Desion Parameters ' Design Parameters i 4 Loop 3 t. oops
(' Nominal Vessel / Core inlet Temperature, 'F 557.0 .550.2 Vessel Average Temperature, 'F 587.1 = 579.6 ; Core Average Temperature, *F 591.4 583.6 I Vessel Outlet Temperature', 'F 617.2 -609.0. -' Average Temperature Rise in Vessel, "F 60.2 58.8-
~l Average Temperature Rise in Core 'F-- 65.2 '63.8 I i
. I Heat Transfer Active Heat Transfer Surface Area, ft2 (STD) 59,742 59,742 (V5H) 59,742 59,742 -'
Average Heat Flux, BTU /hr ft2 (STD). 189,800 142.450 (V5H) 189,800 142,450 Average Linear Power, kw/ft 5.45 4.09: Peak Linear Power for Normal ~ Operation + kwift 14.16 12.26
. 1 Temperature at Peak Linear Power for Prevention of Centerline Melt, 'F 4700 4700-f i
+ Based on maximum Fo of 2.6 for 4 loops and 3.0 for 3 loops, t O
. D0309-13:10 082790 46 1
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l Table 4 2 ] v DNBR Margin Summary . , 1 I l 4 Loon Operation 17x17 STD ' VANTAGE 5H DNB Correlation" WRB1 WRB 2 Correlation Limit" 1.17 1.17-Design Limit (Typ)= 1.25 1.26 (Thm) 1.24 1.24 Safety Limit 1.27 1,69 ; ! (Typ)) (Thm 1.26 1.65 l DNBR Margin * (Typ) 1.6% 25.4% ' (Thm) 1.6% 24.8% Rod Bow DNBR Penalty 1.3% 1.3% (20 inch span) 0% (10 inch span) Transition Core DNBR Penalty 0% < 12.5%" Available DNBR Margin 0.3% > 11% 3 Loon Operation 17x17 STD VANTAGE SH j l DNB Correlation" WRB 1 WRB 2 Correlation Limit" 1.17 1.17 Design Limit (Typ) 1.25 1.26 (Thm) '1.24 -1.24 Safoty Limit (Typ). 1.39. 1 ~.75 (Thm) 1.38 1,72 DNBR Margin * (Typ) .10.1 % 28.0% l (Thm) 10.1 %- 27.9 ] Rod Bow DNBR Penalty 1.3% 1.3% (20 inch span). 0% (10 inch span) .. Transition Core DNBR Penalty. 0%. < .12.5% " Available DNBR Margin. 8.8%
> 14%
r
.DNBR margin between the safety limit and the design limit DNBP.a.
Function of number of VANTAGE SH fuel assemblies in core. For conditions outside the range of applicability of WRB 1~or WRB 2. the W 3 correlation is used with a correlation limit of 1.45 in the pressure range of 500 to 1000 psia and 1.30 for pressures above 1000 psia.
. D0309-13:10 082700 47
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i l 5. ACCIDENT ANALYSIS-1
5.0 INTRODUCTION
5.0.1 General j This section contains a discussion of the accident analyses which were performed to : 2 address the following: ., o Transition from 17X17 Standard fuel to VANTAGE 5H fuel l e _lmplementation of fuel / core related changes-- l
- Implementation of plant related changes e incorporation of contemplated plant changes into the analyses ,
j The analyses presented in this section include:
- 5.1 Non LOCA j 5.2 LOCA 5.3 Mass and energy release and containment integrity .
- 5.4 Radiological eonsequences !
- 5.5 Margin to trip ,
These analyses address transition core effects as well as N 1 loop operation, The impact of the following items have been incorporated into the above analyses; but,. [ separate sections have been devoted to them to discuss their impact more fully: 5.6 RWST boron concentration increase [ 5.7 RCCA parked position change Section 5.8 provides a summary of the instrumentation uncertainty calculations.which 4 were performed in support of the change to the Revised Thermal Design Procedure. . Section 5.9 provides an integrated safety evaluation which specifically confirms the - l acceptability of these changes'using the criteria of 10CFR50.59, . 5.0.2 VANTAGE SH Desian Features . The design features of the VANTAGE 5H fuel' assembly that were considered in the safety i !' analysis effort include the following: , t
- Fuel Rod Dimensions .
- Intermediate Flow Mixer (IFM) Grids e Axial Blankets' O e ,
Integral Fuel Burnable Absorbers (IFBAs)- 1 00309 1:10 000890- 51 . _ _ _ - - _ _ _ _ _ _ - _ - - - _ . - . ._- - - . . .-. ~ . . . .- .
e q s e Debris Filter Bottom Nozzle (DFBN)1
.* Reconstitutable Top Nozzle (RTN),
~ ~
6 Zircaloy Grids l i A brief deceription of each of these items and its treatment in the safety analyses follows. 7 Fyel Rod Dimensions . The fuel rod dirnensions which determine the temperature versus linear power densityh . ; relationship inr,lude rod diameter, pellet diameter, initial pellet tr. clad gap size, and stack; ! height. At Fown in Section 2, there are no changes in these parameters for the_ , VANTAG'c oH design versus the standard design ,The fuel temperature and rod geometry. ! as'sumptions used in the safety analysis bound both the 17X17. STANDARD and VANTAGE l Sh fuel. , ) i IFh,'ji@lg ;
-- . . i The IFM Grid tsed in the' VANTAGE SH fuel assembly design provides an increase'in i DNB ma@.tover standard fuel. As a result, the VANTAGE SH fuel safety analysis limit -
DNB values contain significant DNB margin (see Section 4). For the transition cycles the ' ; core limits dictated by the 17X17 STANDARD fuel without IFM grids are more restrictive ) than those for the VANTAGE SH fuel. Any transition core penalty resulting from the presence of the two fuel types is accounted for with the available DNBR margin.( The IFM grid feature of the VANTAGE SH fuel design' increases the core pressure drop. ., One result is that the control rod scram time to the dashpot has been increased from 2.2 to 2.7 seconds. Tnis increased drop time primarily affects the fast reactivity transients, all . , of which were explicitly reanalyzed for this report. The increased scram time was modeled in all reanalyzed events and the remaining transients have been evaluated, j Axial Blankets and IFBAs j Axial blankets reduce power at the top and the bottom of the rod,' thereby increasing axial' , power peaking at the center of the rod. : This effect is offset by the presence _of the part '
~
length IFBAs which flatten the power distribution. The net effect on the axial shape is a function of the number and configuration of IFBAs in the core'and the time in core life. The effects of axial blankets and IFBAs on the reload safety analysis parameters are j taken into account in the reload design process. The axial power distribution assumptions -
}
in the safety analysis kinetics calculations for this report are sufficiently bounding to i accommodate the presence of IFBAs and axial blankets. Although both these features - ) are part of the VANTAGE SH fuel product for. Millstone Unit 3, they are not new features for - 4 this plant since they were introduced for Cycle 3 operationt l 00309 1;10 080990 , 5 . I
I d t RTNs and DFBNs j. RTNs and DFBNs are features that have been used extensively in Westinghouse designs. 1 Analyses and tests have been performed that confirm the hydraulic compatibility of these particular designs to existing designs so that these components do not impact any parameters important to the safety analysis. Similar nozzle designs have been previously used at Millstone Unit 3 and were introduced for Cycle 3 operation (see Section 2). ' Zircalov Grids Zircaloy structural grids have replaced Inconel grids in the VANTAGE SH fuel assembly with the exception of the top and bottom grids which remain inconel. The effects of i Zircaloy grids have been incorporated into the safety analysis. t 5.0.3 Fuel / Core Related Features and Assumotions The following core related features and assumptions have been incorporated into the-safety analysis. ! e increase in FdH and Fq :
- Reactivity coefficient changes l e Revised thermal design procedure e Revised non RTDP uncertainties O e Relaxed axial offset control Reduced shutdown margin T
e Thimble plug removal ; e Modified OTAT and OPAT trip setpoints e RCC A parked position change
- e. RWST boron concentration increase i e Deletion of clad temperature acceptance criterion for RCCA ejection event -l A discussion of the effect which these items have on the safety analysis is given below.
Increased Power Distribution Peakino Factors (Fm Si Fnj The non LOCA safety analyses have conservatively modeled a full' power F AHf or both 'the g 17X17 STANDARD and VANTAGE SH fuel of 1.64 for RTOP applications and 1.70 for non-
'RTOP. A value of FAH = 1.7 has been used in the LOCA analyses, '
l i l The Technical Specification maximum Fo for four loop ~ operation has been increased from - L 2.32 to 2.60 for both the 17X17 STANDARD and the' VANTAGE SH fuel.' The three' loop : maximum value of Fo has increased from.2.25 to 3.0 (75% power). l i
- D03091:ID=080990 ' 53 a
_________i______ s .'we - <-c_, + + E +- - 4 +- t *I
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Reactivity Coefficient Chances To accommodate longer fuel cycles and extended fuel burnup; a moderator density. coefficient of .50 Ap/gm/cc corresponding to end of cycle, full power conditions, was conservatively incorporated into the safety analyses performed for this report.- , The minimum feedback Doppler power coefficient as a function of power (see the lower bound case in Figure 5.1.1 1) has been reduced (i.e. -' smaller absolute value) in the non-LOCA analysis. The Doppler power coefficient remained unchanged for the LOCA analysis. The limiting care for the complete loss of flow event is analyzed using a 0.0 pcm/'F full l power moderator temperature coefficient.'- The reference licensing basis analysis for this-event used an overly conservative value of + 5.0 pcm/*F at full power. Revised Thermal Desian Procedure (RTOP) . The calculational method utilized to meet the DNB design basis is the RTDP (see Section i 5.8), which is described in Reference 10. Conservative uncertainties in the plant operating ' parameters are statistically incorporated in the design limit DNBR value. - The ' , conservative values are + 6.0/ 6.6'F for reactor coolant system average temperature,
+ 50/ 53 psi for pressurizer pressure,12.0% of rated thermal power (RTP) for core power .
(N loop analysis), 2.3% RTP for core power (N 1 loop analysis),fi2.4% flow for reactor coolant system flow (N loop analysis), and i2.8% flow for reactor coolant eftem flow (N 1 loop analys!s). The calculated uncertainties are reported in Reference 43. Since the - r parameter uncertainties are considered in determining the design DNBR value, the new methodology permits the associated plant safety analyses to be performed using nominal input parameters without uncertainties. Revised Non_ RTOP Uncertainties ; The use of RTOP represents a change in the methodology for-applying u_ncertainties for - i those events used to show the DNB design basis is met. However, the change in power.- pressurizer pressure, reactor coolant system average temperature, and RCS flow - , uncertainty allowances from previously used values.must be evaluated for those events using Standard Thermal Design Procedure (i.e., non RTDP).: The core power uncertainty I used for three loop operation is 2 2.3% RTP as compared to the 2 2.0% 'RTP in the ; current licensing basis (the core power uncertairity'used for N loop operation remains I
!2.0% of rated thermal power). . The LOCA analysis continued using 2.0% RTP as the ;
core power uncertainty for three loop operation. The reactor coolant system average = temperature uncertainty allowance is + 6.0/ 6.6* F as compared to the previous a 6.1
- F; ;
The conservative pressurizer pressure uncertainty is + 50/ 53_ psi as compared to the !
!45 psi noted currently in the i:SAR. The higher pressurizer pressure uncertainty resulted ir. using some of the existing analysis PCT margin. - Specificallyc0.2'F of large break margin and 13.4*F of small break margin was utilized. Finally, the conservative value of the uncertainty for reactor coolant system flow for N loop operation is 2.4% flow i 00309 1:10 082890 54 i
. . . _ i
i i versus the previous i1.8% flow and for N 1 loop operation is i2.8% flow versus i2.0%- f flow. f The values of these uncertainties used in the analysis increased in order to conservatively - d bound plant data which has recently been obtained. Relaxed Axial Offset Control (RAOC) : For four loop conditions, RAOC operation with'a + 12%,17% al band at 100% Rated Thermal Power (RTP) formed the basis of the non LOCA safety evaluations, For three' ! loop conditions, a RAOC a! and of + 10%,15% at 75% RTP formed the basis for the- { l non LOCA safety evaluations. The LOCA analyses used a + 13%, 20% al band for both ; 4 and 3 loop operation. } Reduced Shutdown Maroin The shutdown rnargin considered in the non LOCA safety analysis was reduced from 1.6% AK to 1.3%. AK. 1 Thimble Pluo Removal Thimble plug removal affects the core pressure drops and increases core bypass flow. -L These effects have been conservatively incorporated into the safety analyses performed , for this report. The analyses are actually bounding for either the thimble plugs in place or ; removed. There are no restrictions in placement of the remaining thimble plugs in the. ' core in this situation. _j Modified Over Temoerature oT and Over Power AT Trio Setooints [ The implementation of Vantage SH fuel and'the inclusion of more conservative uncertainty values in temperature and pressure cause the DNB core limits to' change .With the core- l limit change and implementation of RAOC, the over temperature and over power AT: reactor trip setpoints are changed in the analysis.~ The values of the setpoints are ' + changed in the analysis. The values of the setpoints are given in Table 5.11-4. RCCA Parked Position Chanoe The safety analysis performed for this report included consideration of the RCCA full out [ position being in the range from 222 to 231 steps. RWST Boron Concentration increase The non LOCA safety analysin provided in this report considered a range in the RWST boron concentration of from.2500 to 2900 ppm. The LOCA analysis is considered a range in RWST boron concentration of 2700 to 2900 ppm. These values will support the expected Technical Specification limits for this parameter. 00309-1:10 082890 55
'1 l
Deletion of Clad Temperature Accootance C'iterion r for the RCCA Election Event l The analysis of the RCCA ejection event for the VANTAGE SH transition does not include . ; consideration of the 2700'F clad average temperature criterion previously' applied to this accident, as described in FSAR Section 15.4.e
- The deletion of this criterion is not i related to the fuel change. Rather the analysis oquired to support the fuel change provided an opportunity to implement the deletion which was actually explained and ,
Justified in Reference 28. 5.0.4 Plant / Tech Spec Related Chances to be implemented The following plant changes have been addressed completely in this report and can be implemented:
- Expanded accumulator ievel band e Revised pressurizer level program ~
o Response time assumed for the high pressurizer water _ level reactor trip A discussion of the effect of these changes on the safety analysis is given below. , , Exoanded Accumulator Level Band A nominal accumulator water volume of 912.5 ft3 was considered in the non LOCA safety 9 analysis, The LOCA analysis used 900 ft3 per accumulator and have evaluated the impact of 912.5 ft3 Revised Pressurizer Level Prooram The non LOCA safety analysis used a revised pressurizer level program that specified nominallevels of 61.5% and 28% indicated span for full and zero power conditions, respectively. Previously, the nominal zero power level had been 25% The LOCA analysis used a value of 62% for full power conditions'. Response Time Assumed for the Hioh Pressurizer Water Level Reactor Trio The safety analysis performed for the VANTAGE 5H transition assumed a response time of 2.0 seconds for the high pressurizer water level reactor trip.- The pr.evious licensing j basis analysis did not require a response time for this function and the current Technical Specification indicate "N.A.' as the associated response time for this trip function. The assumption of the response time in the analysis presented in this report dictates that a , corresponding Technical Specification change be made to require the 2.0 second . - response. 1 l 00309- t . I O 082890 56 o
i L :l u 5.0.5 Additional Assumptions incoroorated into' Safety Analysis'. l The impact of the following items on the results of the safety analysis han been I specifically determined. These additional assumptions are conservative and require no i i Tech Spec changes for the analysis to be valid. However, additional work.is required j before these Tech Spec changes can be implemented at the plant. l l J l e increased safety valve drift (Pressurizer and Steam Generator) -j ! e Revised auxiliary feedwater flow rates i e Nogative flux rate trip deletion l e- 10% steam generator tube plugging . ! e Degraded safety injection and charging parameters e Modified N 1 loop conditions _ e Modified safety injection system flow source for the steamline break event e Modified setpoints A discussion of these items is given below. 1 Increased Safety Valve Drift (Pressurizer & Steam Generator)- l I The non LOCA safety analyses performed for the VANTAGE 5H transition effort include - revisions in the treatment of both the pressurizer and steam generator safety valves.' With regard to the pressurizer safety _ valves, they were modeled as not opening until pressurizer pressure reached 2575 psia.- The flow through the pressurizer safety valves - was modeled with 3% accumulation,'i.e., it ramps up from zero to full rated flow over the-range of 2575 to 2652 psia (2500 x 1.03 x 1.03). Previously the pressurizer. safety valves had been modeled as opening with the pressurizer pressure at 2500 psia and full rated-flow being reached at 2575 psia. For events that may challenge the secondary system pressure protection, the non LOCA i' analysis in this report employed a model that allowed steam generator pressure to rise to-1320 psia, at which point it was assumed that sufficient safety, valve capacity exists.to stabilize the pressure and prevent further. increases. The 1320 psia figure was selected ' as a conservative bounding value since it is equal to 110% of the steam generator design-pressure. Previous analyses had used a model that allowed steam generator pressure to rise to 1236 psia, which represented the design pressure with a 3% allowance for - accumulation. The current FSAR analysis auxiliary flow rates correspond to a steam generator pressure of 1236 psia which reflects the assumed capability of the steam generator. safety valves. However, as noted above, the analyses for the VANTAGE SH transition use a more conservative model that allows the steam generator pressure to reach 1320 psia.. I The LOCA analyses are not affected by the increased. pressurized safety valve drift. The 3% -increase in the steam generator safety valve drift was considered in both the large-and smail break LOCA analysis. 00309 4 )0 082890 57
1 l [ Revised Auxiliary Feedwater Flow Rates For the loss of non emergency AC power to the station auxiliaries (Section 5.1.4.6), loss of ' normal feedwater (Section 5.1.4.7), and feedwater system pipe break (Section 5.1.4.8) - events, the cases analyzed for the VANTAGE 5H transition assume auxiliary feedwater- j i flow ratas similar to the current licensing basis FSAR analysis. Specifically, the four loop' ; loss oi normal feedwater analysis uses an auxiliary feedwater flow rate of 510 gpm - J l delivered to four steam generators, while the four loop feedwater system pipe break assumed 480 gpm delivered to three steam generators. The 510 gpm four loop value. represents a slight reduction from the 520 gpm used in the current loss of normal feedwater FSAR analysis. The three loop feedwater system pipe break assumes 300 gpm delivered to two steam generators. . In addition, the analyses performed for steam system piping failure (FSAR Section 15.1.5) i and the inadvertent opening of a steam generator relief or_ safety valve (FSAR Section , 15,1.4) assumed a reduction in the maximum auxiliary feedwater flow rate that can be ' I delivered to the steam generators. The current licensing basis steam system piping ~ failure analysis assumes a total auxiliary feedwater flow rate of 2850 gpm delivered to the l faultod steam generator for four loop operation and 2529 gpm to two steam generators' (2229 gpm to the faulted) for three-loop operation. The VANTAGE 5H analyses assume 1200 gpm and 900 gpm delivered to one steam generator for four loop.and three loop ' i operation, respectively. The current licensing basis analysis for inadvertent opening of a steam generator relief or safety valve assumer a total of 2850 gpm delivered to three , steam generators (2229 gpm to the faulted) fu four loop operation and 2529 gpm to two - a steam generators (2229 gpm to the faulted) for three loop operation ! The VANTAGE 5H - analyses assume 1200 gpm delivered to four steem generators and 900 gpm delivered to three steam generators for four loop and three locp. operation, respectively. Necative Flux Rate Trio Deletion 7 The analysis for the dropped RCCA event (FSAR Section 15.4.3) did.not credit the intervention of the power range negative flux rate trip function. Previous Millstone Unit 3 licensing basis analysis had taken credit for this trip. 10% Steam Generator Tube Pluaaina All of the events reanalyzed for this report have incorporated modeling assumptions that bound up to a maximum of 10% ' steam generator tube plugging (or the hydraulic equivalent of plugs and sleeves) in each steam generator, it is assumed that no one. .' steam generator exceeds 10% tube plugging. Dearaded Safety Inlection (SI) and Charaino Parameters - -I The non LOCA analyses performed for steam system piping failure (FSAR Section 15.1.'5).- [ the inadvertent opening of a steam generator relief or safety valve (FSAR Section 15.1.4) ' J D0309 U D 082890 58
and feedwater system pipe break (FSAR Section 15.2.8) assume a 42.0 second delay in j , Si initiation. This delay time was cus'ervatively assumed for the cases with offsite power '! available as well as for those considering the lo_ss of offsite power. Diesel generator.. - l starting and sequence loading delays are included in the 42.0 seconds. The large break LOCA analysis incorporated a safety injection delay time of 40 seconds. The small break l LOCA analysis incorporated a safety injection delay time of 45 seconds. , A 10% reduction in Sl and charging flow was incorporated in the LOCA and nor' LOCA analysis as well as a 10 gpm Si and charging pump flow imbalance. 't l l Modified N 1 Looo Condition ; i For all safety analyses performed for this repod,75% of the four loop Rated Thermal-i Power was the assumed nominal ? full. power" for three loop operation.1 This is a change from the current licensing basis for Millstone Unit 3,'where all three loop non LOCA analyses used this assumption', with the exception of the rods in DNB evaluation for the locked rotor event that assumed a nominal three loop power of 65E The previous LOCA j analysis had also used a three loop power of 65% ; A decrease in the reactor coolant system thermal design flow from 99,600 to 98,300 gpm per loop was considered in the safety analyses.: With the use of RTDP for the VANTAGE . 5H transition analysis, the non LOCA analyt,es performed to confirm the DNB design basis ; p use minimum measured flow, which for three loop operation at Millstone Unit 3 is 101,066 j gpm per loop. These events are unaffected by the decrease in thermal design flow. For all three loop non LOCA analyses that consider thermal design flow, the reduction in flow 1 was either addressed explicitly via reanalysis or evaluated. Modified Safety Iniection System Flow Source for ths Steamline Break Event-Previous steamline break analysis had taken credit for minimum safety injection flow as provided by a single high head centrifugal charging pump. The analyses performed for ; this report instead credit the minimum safety injection flow as provided by.a single high ; head safety injection pump. Using either of these assumed water sources represents a 1 conservative analysis model. Modified Getooints The high steam generator water level setpoint for the safety analysis was conservatively ~I assumed to be at 100% naitow range span _(NRS), in contrast to the_ previous safety i ana'ysis setpoint of 85.7%.NRS. Using the 100% NRS figure in the ana!ysis was intended to provide maximum operational flexibility for the plant. A reduced setpoint for safety injection initiation and steamline isolation on low steam . ] pressure of 435 psia was used in the safety analyses for this. report as compared to the I previous 444 psia setpoint. 1 3 003091:1D'080990 59
i 4 The setpoint for safety injection on low pressurizer pressure was reduced from 1680 psia - d to 1600 psia for the non LOCA safety analysis. The large break LOCA analysis used 1600 - psia for the reactor trip and safety inj6cdon setpoint. The 'small break LOCA analysis.' i used 1600 psia as the, low pressurizer pressure safety injection setpoint/ The small break.- ; LOCA analysis of the 3_and 4 inch break used 1600 psia as the reactor trip setpoint. i However, in the small break LOCA analysis of the 2 inch break, a reactor. trip setpoint of 1700 psia was required to achieve acceptable results. A revised trip setpoint for low reactor coolant flow has been incorporated in the analysis i (see Table 5.1.1-4), This vclue was us'e d in the analysis to conseNatively account for the t l l anticipated flow measurement uncertainty. Based on the final uncertainty allowance i reported in Reference 43, no change to the Technical. Specification values associated with. J' this reactor trip are required as a result of the VANTAGE 5H transition analysis, , t l , ,t ( O' i o 1 9- , 00309 1:10 080990 5 10 .
.t
- . . - , ..-_v ~ . - -
I 5.1 NON LOCA ACCIDENT ANALYSIS 5.1.1 Introduction j This section addresses the impact on the FSAR Chapter 15 Non LOCA accident analyses of the transition of Millstone Unit 3 from Westinghouse 17x17 STANDARD fuel to Westinghouse VANTAGE SH fuwl. The general approach taken will be to present event specific evaluations or analyses that address the impact of this transition on the non LOCA accidents within the current Ilconsing basis for Millstone Unit 3, as defined by the FSAR. Both N and N.1 (i.e., four and three) loop operation are considered. The non LOCA g analyses and evaluations presented here include assumptions that reflect the direct impact of certain VANTAGE SH features as well as the other changes discussed in ! SYon 5.0 and in Sections 5.1.1.1 through S;1,1.5. ; In most cases, the basic methodologies applied to the non LOCA analyses presented for VANTAGE 5H fuel in this report are the same as those currently documented in the Millstone Unit 3 FSAR for standard fuel. Whenever this is the case for the non LOCA analyses presented in this report, repetition of the explanaiory discussion found in the ; FSAR will be kept to a minimum More detailed discussion will be provided here only as required to deal with deviations from the FSAR methodology. , The limiting single failures have not changed from those assumed in the current Millstone ! Unit 3 FSAR for standard fuel. j 5.1.1.1 Plant Characteristics and Initial Conditions Assumed in the Accident Analyses ! Table 5.1.1 1 lists the principal power rating values assumed in analyses performed in this ) report. As noted earlier, the three loop power rating used in the non LOCA analysis for the ! VANTAGE 5H transition represents 75% of the four loop RTP, which is conservative with . respect to the current licensed maximum three loop Mwer for Millstond Unit 3 of 65% RTP. For most accidents analyzed to demonstrate that the DNB design basis is met, the Revised Thermal Design Procedure (RTDP) is employed in defining the initial conditions ' (see Section 4.2), The other accidents obtain initial conditions by adding the maimuri steady. state errors to rated values (this procedure is commonly known as the Standate-Thermal Design Procedure or STOFi. The following steauy. state errors are considered: , Core power i 2.0% anG t 2.3 f. allowance for four loop and three loop, ! ! respectively, for calorimetric error. This error is l conservatively applied in the positive direction in non LOCA > accident analyses. Average reactor + 6/ 6.6'F allowance for controller dead band- ; coolant system measurement error; the asymmetric allowance is due to l temperature consideration of a transmitter bias (see Reference 43)._ j i
. 1 D0309 2:10/080890 5 11 I
v
- . - -___ - - - - . - - - . - - - - - . -- . - ~___- - -_ . . - -___
I 1
-f i i l
i ! Pressurizer pressure + 50/ 53 psi allowance for steady state fluctuations and . measurement errors; the asymmetric allowance is due to consideration of a transmitter bias, j For clarification it is reiterated that accidents employing RTDP assume minimum measured flow (MMF) while accidents analyzed with STDP assume thermal design flow '
]
(TDF). Table 5.1.12 summarizes the initial conditions and computer codes used in the , i accident analyses. The values of other pertinent plant parameters utilized in the accident analyses are given in Table 5.1.13. l L 5.1.1.2 Reactivity Coefficients Assumed in the Accident Ana!vses i The transient response of the reactor coolant system is depenoont on reactivity feedback i effects, in particular the moderator temperature coefficient and the Doppler power coefficient. Depending upon event specific characteristics, conservatism may dictate the ' use of either large or small reactivity coefficient. values. Figure 5.1.1 1 shows the upper and lower Doppler pcwor coefficients, as a function of power, used in the transient ] analyses. Figure 5.1.12 shows the modarator density coefficient, as a function of temperature, used in the transient analyses. The justification for use of conservatively large versus small reactivity coefficient valuet is treated on an event by event basisc The ; values used are given in Table M.12. The limiting case for the complete loss of flow event is analyzed using'a 0.0 pcmi'F full power moderator temperature coefficient. The reference licensing basis analysis for this l event used an overly conservative vdue of + 5.0 pcm/'F at fu!I power. In addition the t locked rotor rods in DNB event is analyzed using a + 5.0 pcm/'F full power moderator temperature coefficient. The reference licensing basis analysis for this event used a value of 0.0 pcm/'F at full power. 5.1.1.3 Rod Cluster Control Assembly insertion Characteristics The negative reactivity insertion following a reactor trip is a function of the acceleration of the RCCAs and the variation in rod worth as a function of rod position. With respect to the accident analyses, the critical parameter is the time from the start of insertion up to the j dashpot entry or approximately 85 percent of the rod cluster traveli For the accident _ i analyses, the insertion time to dashpot entry is conser.atively taken as 2.7 seconds. The RCCA position versus time assumed in accident analyses is shown on Figure 5.1.13. Figure 5.1.14 shows the fraction of total negative reactivity insertion versus normalized , l rod insertion. This curve is used in all transient analysis point kinetics core models.- l There is inherent conservatism in the use of this curvo in that it is based on a bottom skewed axial power distribution. For cases other than those associated with axial xenon D oscillations, significant negative reactivity would have been inserted due to th'e more - favorable axial power distribution existing prior to trip. 00309 2 10 080990 5 12 i
,y , , , . - r--. .--, i- ,e , . - - + -- 1
l ~! j t i I i l The normalized RCCA negative reactivity insertion versus time is shown on Figure 5.1.15. l The curve shown in this figure was obtained by combining Figure 5.1,13 and Figure 5.1.1 l 4. A total negative reactivity insertion following a trip of 4.0 % Ak/k is assumed in the - l transient analyses except where specifically noted otherwise. This assumption is verified ; to be conservative with respect to the core design. , I i l For analyses requiring the use of a dimensional diffusion theory code (see discussion of the TWINKLE code in FSAR Section 15.0.11.5 and Reference 26), the negative reactivity insertion resulting from reactor trip is calculated directly by the code and is not separable from other reactivity feedback effects, in this case, the RCCA position versus time of ; Figure 5.1.13 is modeled in the code. i 5.1.1.4 Trio Setooints and Delavs Assumed in Accident Analyses ( l Limiting trip setpoints assumed in accident analyses and the time delay assumed for each ! trip function are given in Table 5.1.14. Where appropriate, differences between the trip i setpoints for four and three loop operation are indicated. A reactor trip signal acts to open two trip breakers connected in series, feeding power to the control rod drive mechanisms. The loss of power to the mechanism coils causes the mechanisms to release the rod cluster control assemblies, which then fall by gravity into the core. There are various instrumentation delays associated with each trip function, O i including delays in signal actuation, in opening the trip breakers, and in the release of the rods by the mechanisms. The total delay to trip is defined as the time delay from the time l that trip conditions are reached to the time the rods are free and begin to !all. The OTAT and OPAT trip functions referred to in Table 5.1.14 are graphically illustrated in 1 Figures 5.1.16 and 5.1.17 for four and three loop conditions, respectively. These trip setpoints were determined to bound the transition cores as well as a full core of VANTAGE 5H fuel. New axial offset. limits that bound the various cores were employed to determine a new f(AI) penalty shown in Figure 5.1.18. which is applicable to both four and three loop operatia. i The difference between the limiting trip setpoint assumed in the analysis and the nominal t trip setpoint represents an allowance for instrumentation channel error and setpoint error. Nominal trip setpoints are specified in the plant technical specifications. During plant startup tests it will be demonstrated that actual instrument time delays are equal to or less > than the assumed values. Additionally, protection system channels are calibrated and - instrument response times determined period!cally in accordance with the plant's technical specifications. 5.1.1.5. Non LOCA Safety Evaluation Methodoloov r The reload safety evaluation methodology is described in Reference 1. The methodology confirms that, if a core configuration is bounded by existing safety analyses, then the applicable safety criteria are satisfied. The methodology systematically identifies :
\
00309 2:1o<080990 5 13 - '
i i parameter changes on a cycle by cycle basis which may violate existing safety analysis I assumptions and identifies the transients which require evaluation. This methodology is l
- applicable to the evaluation of VANTAGE 5H t
- ansition and full cores.
Any required evaluation identified by the reload methodology is one of two types. If the 4 identified parameter is only slightly out of bounds, or the transient is relatively insensitive to i that parameter, a simple evaluation may be made which conservatively evaluates the magnitude of the effect and explains why_the actual analysis of the event does not need to 1 be repeated. Alternatively, should the deviation be large and/or expected to have a significant or not easily quantifiable effect on the transients, analyses are required. ; The analysis approach utilizes Westinghouse codes and methods which have been i accepted by the NRC and have been used in previous submittals to the NRC. These. . . methods are those which have been present6d to the NRC for a specific plant, reference i SARs or reports for NRC approval. Section 15.0 of the Millstone Unit 3 FSAR provides information on the analysi6 codes and methods that is applicable for the non LOCA analyses performed for the VANTAGE SH transition. , The key safety analysis parameters are documented in Reference 1. Values of these parameters which bound both fuel types (17x17 STANDARD and VANTAGE 5H) were , assumed in the non LOCA safety analyses. For subsequent fuel reloads, these parameters will be evaluated to determine if violations of these bounding values exist. An . evaluation of the effected accidents will be performed as described in Reference 1, I i a 9 00309 21 0 080990- 5 14 .
- l. i I ,. _ . . .- _ _ , - - _ , - . - . _ -
t i
- t TABLE 5.1.1 1 1 NUCLEAR STEAM SUPPLY SYSTEM POWER RATINGS ,
Four Loop Three Loop r (MWt) (MWt) l NSSS thermal power output (O MWt Pump Heat) 3411 2558 ; l i NSSS therma! power output (14/11
- Mwt nom Pump Heat) 3425 2569 ;
NSSS thermal power output (20/15
- Mwt max Pump Heat) 3431 2573 NSSS thermal power output (14/11
- Mwt nom Pump Heat) 3494" 2628" ;
~ NSSS thermal power output (20/15
- Mwt max Pump Heat) 3500" 2652"-
Engineered safety features design rating 3656" 2771" , O i i i i i i L Four loop pump heat'three loop pump heat ; Includes 2.0% power uncertainty l includes 2.3% power uncertainty l l O- ,. I 00309 2.10 080930 5 15
.i r
TABLE 5.1.1-2 (SHEET I OF 4) sun 0 MARY OF INITIAL CONDITIONS AND COMPUTER CODES USED ; Reaci tv CoeNicient.as ed ; neoderaeor On.stal NSSS 1hermal FSAR Computer Denssty Power Output Assumed Section Faults Codes Utahzed (Ak/gm/cm3) Doppler (RSWI) 15.1 increase in heat removal by the secondary system. . t Feedwater system LOFTRAN 0.50 Lower (see Figure o and 3425(a) malfunctions that result in 5.1.1-1) ' ! an increase in feedwater b , Excessive increase in LOFTRAN Fig. 5.1.1-2 Lower and upper 3425(a) - secondary steam flow Most negahve (see Figure 5.1.1-
& 0.50 1)
Inadvertent openiaxj of a LOFTRAN Functon of See Secton 0 (Subcnhcal) sicam generator sal (sty or moderator densely 5.1.3.4. Figure , telief valve Fig. 5.1.3-11 5.1.3-14 ; Steam system pepeng LOFTRAN. THINC Funchon of See Seckon 0 (Subcritical) Iailure moderator 3.1.3.5 Figure t l density - 5.1.3-14 i Fig. 5.1.3-11 > 15.2 Decrease in heat removal [ by the secondary system Loss of external electrical LOFTRAN Fig. 5.1.1-2 Lower and upper 3425 (4 loop) i -- load and/or turbine trip Most negative & . (see Figure 5.1.1- 2569 (3 loop) i ! 0.50 1) seruwe 4 is**waum 5-16 9 _ ._ . _. 9 . - _ . _ _ . . 9
l O , t TABLE 5.1.1-2 (SHEET 2 OF 4) , i SURAMARY OF INITIAL CONDITIONS AND COtePUTER CODES USED Reactevity CoeNicsonts Assumed
- Moderator gnielal NSSS Thermal FSAR CC. se.r Denssty Power Output Annassned Section Faults Codes Unlued (ak/gm/cm3) Doppler (agwt)
Loss of non-emergency Results bounded AC power lo the station by loss of normal 1 auxiliaries feedwater with loss ; of AC power . Loss of normal feedwater LOFTRAN Fig. 5.1.1-2 Upper, see 3656(a) flow Most negative Fig. 5.1.1-1 Feedwater system pipe LOFTRAN 0.50 Upper, see 3656 (4 loop) break Fig. 5.1.1-1 2771 (3 loop) i 15.3 Decrease in reactor coolant system flow rate Partial and complete loss LOFTRAN. Fig. 5.1.1-2 Upper, see 3431(a) of forced reactor coolant FACTRAN THINC Most negative for Fig. 5.1.1-1 Now. partial and 0.0 for limiti.eg complete Reactor coolant pump LOFTRAN. Fig. 5.1.1-2 Upper, see 3500 (4 loop) , shalt seizure (locked FACTRAN Most negative Fig. 5.1.1-1 2652 (3 loop)
; rotor)-pressure, i temperature transient Reactor coolard pump LOFTRAN. Fig. 5.1.1-2 Upper, see 3431 (4 loop) ,
shaft seizure (locked FACTRAN Most negative Fig. 5.1.1-1 2573 (3 loop) rotor)-rods in DNB THINC . i
, 6WillWI 9 f t)(19tilSNatt 5-17 . . .- . . . . . . . . . . . , . . ,...---~--..e..-._-- - . , . - - . . . ~ . . . - . - .c.,-_s . . . . . _ . - . . . .. . . _ _ . . _ . _ . , . - . . - - - - - . . , . . . _ . . , . , - , _ . . - . - - - . _ .
l TABLE 5.1.1-2 (SHEET 3 OF 4) l l
SUMMARY
OF INITIAL CONDITK)NS AND COMPUTER CODES USED i Reactswsty Coeffaceents Assumed , Moderator kWhal NSSS Thermal
'FSAR Computer Denssey Power Output Assumed Section Faults Codes Unlized (ak/gm/cm3) D3ppler (wwt)
- 15.4 Reactivity and power
~
i distribuhon anomalies Uncontrolled RCCA bank TWINKLE. Refer to Section Coelhcient is O (subc.%:al) withdrawal from a FACTRAN. THING 5.1.6.1.2 consistent with . subcritical or low power Doppler Defect of
- startop condstion -0.90% AK Uncontrolled RCCA bank LOFTRAN Fig. 5.1.1-2 Lower and upper 3425 (4 loop) withdrawal at power most negative & see Fig. 5.1.1-1 2569 (3 loop)
O.50 RCCA misalignment THINC. TURTLE 3425(a) LOFTRAN, LEOPARD Spectrum of RCCA TWINKLE. Refer to Section Icefficient for all 3500 (4 loop) ektion accidents FACTRAN. THINC 5.1.6.6.2 caos is consestent 2652 (3 loop) with a boundeng 0 (4 & 3 loop) Doppler Detect of 2
-0.90% AK r
~
- is.. n e in n .. r s.1 a O - - - . - -. _ . - -
O - _ - _ - _ - . - _ _ O
{- T I i 1 I l TABLE 5.1.1-2 (SHEET 4 OF 4) j
SUMMARY
OF NTIAL CONDITIONS AND COGAPUTER CODES USED ! Reactivity Coh AssunuMS . Rsodersoor iniaW NSSS Thewmal FSAR CC.;- w Denssty Power Output Assumed i Section Faults Codes Utdized (ak/gm/cm3) Doppler (RSWR)
', 15.5 increase in reactor coolant inventory i i
inadvertent operation of LOFTRAN Fig. 5.1.1-2 Lower & upper. 34114) the emergency core Most nc-;:Sn & see Fig. 5.1.1-1 ! cooling system during 0.50 power operation 15.6 Decrease in reactor coolant inventory inadvertent opening of a LOFTRAN Fig. 5.1.1-2 Upper. see 3431 (4 loop) pressurizer safety or relief Most negatwe Fig. 5.1.1-1 2573 (3 loop) 4 valve !
- (a) For this case, explicit analysis performed for four-loop operaton only. Three loop operation was evaluated. ;
t I l i . ] 00309-2J D 000890 5-19
. ._ - ._. . ~_ ._ -.. _. _ _- _
i l I l .! TABLE 5.1.13 l NOMINAL VALUES OF SELECTED PLANT PARAMETERS l UTILIZED IN THE NON-LOCA SAFETY ANALYSES ! ! Pour Loop Three Loop , ! I NSSS thermal power (MWt) See Table 5.1.11 ! Core inlet temperature ('F) 557.0 550.2 j Vessel average temperature ('F) 587.1 579.6 ! l Reactor coolant system pressure (psia) 2250 2250 l l Reactor coolant flow per loop, TDF (gpm) 94600 98300 l t Reactor coolant flow per loop, MMF (gpm) 96700' 101066 i NSSS steam flow, total (106 lbm/hr) 15.03 10.76 f Steam pressure at SG outlet (psla) 957 900 l Maximum steam moisture content (%) 0.25 0.25 ; i Feedwater temperature at SG inlet (*F) 436.2 404.2 . Average core heat flux (8tu/hr ft2) 189800 143450 l 9 1 h 1 t i 9i l 00309 2,10/080890 5 20 ,
. ~ . , , . - . - , __, . - . . . _ . , -.,-,_,,-n_ . . . - - , _ . . . . . ~ ~ . ,
i l l TABLE 5.1.14 ; l REACTOR TRIP SETPOINTS AND TRIP TIME DELAYS ASSUMED 1 l FOR THE MILLSTONE UNIT 3 NON-LOCA ANALYSES { i Assumed Non-LOCA Trip Setpoint(d) Y, Trip Function (For Four & Three Loop, Unless ! (sec nds) l Otherwise indicated) - Power range high neutron 118 % . 0.5 [ flux. high setting four loop ; Power range high neutron 89% 0.5
- flux high setting three loop ,
l Power range high neutron 35% 0.5 l i flux. Iow setting Overtemperature AT Variable (See Figures 5.1.16&5.1.17) 7.0(a) Overpower AT Variable (See Figures 5.1.16 & 5.1.17) 7.0(a) ; High pressurizer pressure 2425 psia 2.0 Low pressurizer pressure 1860 psia 2.0 i Low reactor coolant flow 85% loop flow O' (from loop flow detectors) 1.0 ! RCP underspeed trip 92% nominal 0.6 Turbine trip N/A 2.0(e) Low low steam gen, water 0.0% of narrow range span 2.0 level (NRS) for feedline break ; 10% of NRS for loss of 2.0 ! normal feedwater/ loss of ' offsite power ! l High steam gen. level trip of 100% of NRS 2.5(b) f l the feedwater pumps, closure ; of feedwater system valves, '
& turbine trip 7.0(c) ,
Pressurizer water level high 100% of span 2.0 , t
- a. Total time delay from the time the temperature difference in the coolant loop exceeds the tnp setooint until the - f RCCAs are free to fall. Delay includes the response charactenstics of the RTD/thermowell' scoop configuration, electronics delays, top breaker opening delays and gnpper openmg delays. i
- b. From time setpoint is reached to turbine trip.
- c. From time setpoint is reached to feedwater isolation. }
- d. Tabulated values conservatively bound Toch Spec values with uncertainties O 4
- e. Some non LOCA analyses have assumed 0.0 for conservatism.
D0309 2:10 080990 5 21 t
.-, --,,---v - , ,
- 4 l l l !
5.1.2 Accidents Analyzed and Evaluated
+
i l For each of the potentially impacted non LOCA transients, consideration was given to - ; effects of the VANTAGE 5H fuel design features and modified safety analysis assumptions l discussed in Section 5.1 of this report. As dictated by event specific sensitivities, a ! ! decision was made for each transient with regard to the need for formal analysis, as i opposed to simply evaluating the impact of the subject features and assumptions. An i additional issue that requires consideration for the Millstone Unit 3 analyses is the l possible difference in relative behavior of the three loop cases as compared to those with j four loops operating. For certain events, previous licensing basis analyses clearly ; demonstrate that four loop results bound those for three loops operating. In those cases, l an evaluation of tnres loop operation was performed for the VANTAGE 5H transition, even - ! if analysis was determined to be necessary for the associated four loop case. l i Table 5.1.21 documents for each of the potentially impacted non LOCA transients whether an analysis was performed or an evaluation was sufficient to assess the impact of the VANTAGE SH transition. ! i o; i k i b h 9 L l 9' 5 22 00309 2:1D/080890
_ .__ __ _ _ _ _._. _ ._~.__ _..__ _ _ _ . _ . .
?
TABLE 5.1.21 (SHEET 1 OF 2) MILLSTONE UNIT 3 NON LOCA EVENTS ( EXTENT OF CONSIDERATION FOR VANTAGE SH TRANSITION i Four Loop Three-Loop ; Event Category or Event FSAR Section Operation Operation i t increase in Heat Removal 15.1 , By the Secondary System l P Feedwater (FW) System 15.1.1 Evaluated Evaluated j i Malfunctions that Result in a ' , Decrease in FW Temperature I FW System Malfunctions that . l' Result in c;, incrossc in FW Flow 15.1.2 Analyzed Analyzed i Excessive increase in Secondary 15.1.3 Analyzed Evaluated l Steam Flow l 5 Inadvertent Opening of a SG 15.1,4 Analyzed Analyzed
- O Rollet or Safety Valve _
Steam System Piping Failure 15.1.5 Analyzed Analyzed Decrease in Heat Removal By 15.2 the Secondary System t Loss of External Electrical Load 15.2.2 Bounded by Turbine Trip ; l Turbine Trip 15.2.3 Analyzed Analyzed ' Inadvertent Closure of MSIVs 15.2.4- Bounded by Turbine Trip ; Loss of Condenser Vacuum /Other 15.2.5 Bounded by Turbine Trip
- Turbine Trip Events -
Loss of Non Emergency AC Power 15.2.6 Bounded by LONF/AC Loss , to the Station Auxillaries i Loss of Normal Feedwater Flow 15.2.7 Analyzed . Evaluated g Feeewaie, System ripe Br.,ax .
,5.2.5 Anaiyzed Anaiyzed 00309 2:10/080890 5 23
I TABLE 5.1.21 (SHEET 2 OF 2) l MILLSTONE UNIT 3 NON-LOCA EVENTS I l EXTENT OF CONSIDERATION FOR VANTAGE SH TRANSITION . Four-Loop Three Loop [ ( Event Category or Event FSAR Section Operation Operation ! Decrease in RCS Flow Rate 15.3 Partial Loss of RCS Flow 15.3.1 Analyzed Evaluated l t Complete Loss of RCS Flow 15.3.2 Analyzed Evaluated j i RCP Shaft Seizure (Locked Rotor) 15.3.3 .- Analyzed Analyzed i ' I RCP Shaft Break 15.3.4 Bounded by Locked Rotor Reactivity and Power 15.4 } Distribution Anomalies l Uncontrolled RCCA Bank 15.4.1 Analyzed Withdrawal from Subci;tical r Uncontrolled RCCA Bank 15.4.2 Analyzed . Analyzed I Withdrawal at Power j Dropped RCCA (RCCA Misalignment) 15.4.3 Analyzed - Evaluated l i inadvertent Loading and Operation 15.4.7 Evaluated Evaluated of a Fuel Assembly - l RCCA Ejection Accidents 15.4.8 Analyzed Analyzed , increase in RCS Inventory 15.5 Inadvertent Operatica of ECCS 15.5.1 Analyzed Evaluated Decrease in RC3 Inventory 15.6 Inadvertent Opening of Pressurizer' 15.6.1 Analyzed Analyzed .; Safety or Reliel Valve
=
The reference "four loop" analysis performed for this event assumes only two RCPs , to be operating, That analysis bounds any conditions' associated with three loop - operation. ' i D0309 2:10/080890 5 24 i
. ~ _ . . . - . . . - _ - . - . . - ._. ---- _- - - . _._ _ _
i i 5.1.3 Increase in Heat Removal by the Secondary System [ A number of postulated events that could result in increased heat removal from the RCS l by the secondary system have been identified and the appropriate limiting cases are ' i presented in Section 15.1 of the Millstone Unit 3 FSAR. The analyses and evaluations for these events, necessary to support the introduction of VANTAGE 5H fuelinto Millstone ; Unit 3, are discussed in tnis section. ~ 5.1.3.1 Feedwater System Malfunctions that Result in a Decrease in Feedwater i Temocrature [ 5.1.3.1.1 Introduction As described in FSAR Section 15.1.1, the cooldown of the RCS resulting from reductions- l in feedwater temperature produce an in, crease in core power in the presence of a ; negative moderator temperature coefficient of reactivity. Such transients are attenuated j by the thermal capacity of the secondary plant and of the RCS. The high neutron flux, overtemperature AT (OTAT), and overpower AT (OPAT) reactor trips prevent any power increase which could lead to a departure from nucleate boiling ratio (DNBR) less than the - limit value. l The not effect on the RCS from the reduced feedwater temperature is similar to that from Meressed secondary steam flow, i.e. the reactor reaches a new equilibrium condition at a f O' power level corresponding to the new steam generator AT. A decrease in normal ! feedwater temperature is classified as an ANS Condition 11 event, a fault of moderate frequency. The general acceptance criteria for this event category are discussed in t FSAR Section 15.0.1.2. 5.1.3.1.2 Method of Analysis j As described in FSAR Section 15.1.1.2. this transient is analyzed by computing conditions i at the feedwater pump inlet following opening of the low pressure heater bypass valve at {~ full power. These feedwater conditions are then used to cerform a heat balance through the high pressure heaters. This heat balance gives tN, new feedwater conditions at the steam generator inlet. The calculations made fm inis event reflect the modified plant conditions associated with the VANTAGE Sri fuel analysis and discussed in Section ; 5.1.1.3. t 5.1.3.1.3 Results i Opening of a low pressure heater bypass valve and the trip of the heater drain pumps, as l discussed in the FSAR, causes a calculated reduction in the feedwater temperature of - less than 35'F. This means that the excessive increase in secondary steam flow event of Section 5.1.3.3 which considers a 10% step load increase, groduct s a greater feedwater
; temperature reduction than the cases considered here.- Therefore, opening of the low pressure heater bypass valve would result in a transient very similar to, but of reduced :
00309 3 10 080990 5 25
4 : 1 O! magnitude relative to the Section 5.1.3.3 event. Therefore, no explicit results are presented for the opening of the heater bypass valve. ! i i 5.1.3.1.4 Conclusions j The decrease in feedwater temperature transient is less severe than the increase in i secondary steam flow event. Based on results presented in Section 5.1.3.3, the ; applicable acceptance criteria for the decrease in feedwater temperature event have been ; met and the conclusions of the FSAR remain valid for both four and three loop operation. l 5.1.3.2 Feedwater System Malfunctions that Result in an Ingrease f in Feedwater Flow ( i 5.1.3.2.1 Introduction . As described in FSAR Section 15.1.2 the addition of excessive feedwater will produce a decrease in reactor coolant temperature which, in the presence of a negative moderator ! temperature coefficient of reactivity, causes an increase in core power. Such transients ! are attenuated by the thermal capacity of the secondary plant and of the RCS. The high > neutron flux OPAT and OTAT reactor trips prevent any power increase which could lead to ; a DNBR less than the safety analysis value. Continuous addition of excessive feedwater l would ultimately be limited by the steam generator high high level trip which closes all ; feedwater control and isolation valves, trips the main feedwater pumps and trips the . l turbir.e. The increase in normal feedwater flow incident is classified as an ANS Condition ; 11 event, a fault of moderate frequency. The general acceptance criteria for this event t category are discussed in FSAR Section 15.0.1.2. ! 5.1.3.2.2 Method of Analys,i,1 ; The excessive heat removal due to a feedwater system malfunction transient is analyzed + using the general methodology, as described in FSAR Section 15.1.2.2. As appropriate, j the safety analysis assumptions discussed in Section 5.1.1.2, including a conservatively l large negative moderator temperature coefficient, are reflected in the analysis of this : report. The analysis is performed to demonstrate acceptable consequences in the event > of an excessive feedwater addition, due to a control system malfunction or operator error i which allows a feedwater control valve to open fully Specifically, the feedwater system malfunction is analyzed to show that the DNB design basis is satisfied for the transition to , VANTAGE 5H fuel Consistent with the FSAR, a step increase of 140% and 200% of ' nominal feedwater flow to one steam generator is assumed for full and zero power. ' respectively. The following cases have been considered: -
- 1. Accidental full opening of a feedwater control valve with the reactor just critical at i zero load conditions.
- 2. Accidental full opening of one feedwater control valve with the reactor at power _ f assuming automatic and manua.' rod control. '
00309 3:10 080890 5 26 ' i
- i. [
l i i l l O i
- For the analysis of this event in support of VANTAGE SH, the initial reactor power, [
i pressure, and RCS temperatures are determined using RTDP methodology, which is a f change from the previously used STDP methodology currently presented in the FSAR. , Consistent with the RTDP approach, nominal values are assumed for the initial reactor j power, pressure, and RCS temperature while minimum measured flow is assumed rather than thermal design flow, Uncertainties in initial conditions are included in the limit DNBR,
; - as discussed in Section 4.0 of this report. The specific initial conditions and plant -
l characteristics are discussed in Section 5.1.1.3. l Normal reactor control system and engineered safety systems are not required to function. The reactor protection system function is expected to trip the reactor due to overpower or turbine trip on high high steam generator water level conditions. - 5.1.3.2.3 Results ! i As stated in the FSAR, for the case at zero load conditions, the results of analysis with ! four loops in operation bovad those with three loops in operation. Additionally, it has been j determined that the full p'Jwer four loop results bound those for three loop operation. ! i Also, no results are presented for an accidental full opening of one feedwater control valve t with the reactor at zero power and the above mentioned assumptions. That particular i case is bounded by the results for RCCA bank withdrawal from a subcritical or low power O startup condition as analyzed in Section 5.1.6.1. {
- i The full power case (maximum reactivity feedback coefficients, automatic rod control) ;
results in the greatest power increase. Assuming the reactor to be in manual rod control ! results in a slightly less severe transient, so the result for that case are not presented. l When the steam generator water level in the faulted loop reaches the high high level j setpoint, all feedwater control and isolation valves are closed and the main feedwater i pumps are tripped thereby terminating the cooldown event. This prevents continuous I addition of feedwater, in addition, a turbine trip and a resulting reactor trip are initiated. A high neutron flux reactor trip initiates a reactor trip prior to the high high level setpoint for j the N 1 full power case. Transient results for four loops in operation are presented in Figures 5.1.31 and 5.1.3 2
- that show nuclear power, pressurizer pressure, DNBR, faulted loen AT, and TAVO as functions of time. Transient results for three loops in operation are presented in Figures !
5.1.31 A and 5.1.3 2A that show the same parameters as functions of time. As noted l earlier, the four loop results bound those for three loop operation for zero power and thus j three loop results are not presented. The DNBR does not drop below the limit value at ! any time during the transient, thereby maintaining the ability of the primary coolant to : remove heat. Following reactor trip the plant approaches a stabilized condition and i O standard plant shutdown procedures may then be used to further cool down the plant. ; Q 'i i
. 1 o0309 3:10 080890 5 27. 1 i
_ _. . _ . _ , _ . _, . _ . . 1______4 . - . , _
- - - -... - ~-- . - - - - - - - . - - - . - .- _
i Since the power level rises during the excessive feedwater flow incident, the fusi temperatures also rise until reactor trip occurs. The core heat flux lags behind the neutron ! l flux response due to the fuel rod thermal time constant; hence, the peak heat flux does not [ exceed 118 percent of its nominal value (i.e., the assumed high neutron flux trip setpoint). i . Thus, the peak fuel temperature will remain below the fuel melting temperature. l i The calculated sequence of events for the full power increase in feedwater flow event is { shown in Table 5.1.31. 1 5.1.3.2.4 Conclusions The results of the analyses show that the predicted DNBR for an excessive feedwater addition at power remains above the limit value (see Table 4 2) so that the DNBR design i i basis is met. Similarly, the results demonstrate that neither the reactor coolant system nor !
- the secondary system are overpressurized during this event. It was also determined that j the reactivity insertion for the RCCA bank withdrawal from subcritical event (Section
- 5.1.6.1)is larger than that for an excessive feedwater addition at zero power. Therefore.
the results of the RCCA bank withdrawal from subcritical event bound those for the zero power excessive feedwater addition event. For all the cases of excessive feedwater
- addition considersd, the conclusions of the FSAR with respect to both four and three loop !
operation remain valid. ; j 5.1.3.3 Excessive Increase in Secondary Steam Flow 5.1.3.3.1 Introduction i As described in FSAR Section 15.1.3. an excessive load incsease incident is a rapid ' increase in steam flow that causes a power rnismatch between the reactor core power ! and the steam generator load demand. This accident could result from either an t administrative violation such es excessive loading by the operator or an equipment malfunction in the steam dump control or turbine speed control, The excessive load increase transient is considered to be an ANS Condition 11 event, i.e.'s fault of moderate frequency. The general acceptance criteria for this event category are discussed in i FSAR Section 15.0.1.2. l The reactor control system is designed to accommodate a 10 percent step load increase and a 5 percent per minute ramp load increase in the range of 15 to 100 percent of full ; power. Any loading rate in excess of these values may cause a reactor trip actuated by i the reactor protection system. Steam flow increases greater than 10 percent are discussed in Sections 5.1.3.4 and 5.1.3.5. l Protection against an excessive load increase accident is provided by the following ! reactor protection system signals: ,
- OTAT
! - OPAT ' l
. i I
00309 31 0 080990 5 28 .
t
- Power range high neutron flux
- Low pressurizer pressure f
j 5.1.3.3.2 Method of Analysis . I i The excessive load increase event is analyzed using the same basic methodology f described in FSAR Section 15.1.3.2, and ac 1'dingly four cases are analyzed to l demonstrate the plant behavior following a t- Wrcent step load increase from rated load. } These cases are as follows: l
. i"
- 1. Manual rod control with minimum reactivity feedback i
- 2. Manual rod control with maximum reactivity feedback - l
- 3. Automatic rod control with minimum reactivity feedback 1 l
- 4. Automatic rod control with maximum reactivity feedback [
l The primary consideration in the analysis of the excessive load increase event is to demonstrate that the DNB design basis is satisfied for the transition to VANTAGE SH fuel. , As appropriate, the modified safety analysis assumptions discussed in Section 5.1.1.2 are j reflected in the analysis of this event. For the analysis of this event in support of i VANTAGE 5H, the initial reactor power, RCS pressure, and RCS temperatures are O determined using RTDP methodology, which is a change from the previously used STDP methodology currently presented in the FSAR. l j i Consistent with the RTDP approach, nominal values are assumed for the initial reactor ! power. RCS pressure, and RCS temperature while minimum measured flow is assumed l rather than the thermal design flow applicable to STDP. Uncertainties in initial conditions ; are included in the limit DNBR, as discussed in Section 4.0. The specific initial conditions l and plant characteristics are discussed in Section 5.1.1.3. ; i For the minimum moderator feedback cases a zero moderator temperature coefficient of ] reactivity was assumed which results in the least Inherent transient response capability. i ! For this cooldown event, the zero moderator temperature coefficient produces bounding ! results relative to those with a positive moderator temperature coefficient. The maximum l moderator feedback cases assume conservatively the most negative moderator j temperature coefficient of reactivity which produces the largest amount of reactivity j feedback from changes in coolant temperature. The most negative moderator j temperature coefficient used in the VANTAGE 5H analysis is more limiting than that used j previously. ! As discussed in FSAR Section 15.1.3.2, normal reactor control systems and engineered 4 safety systems are not required to function for this event. The reactor protection system is -] assumed to be operable; however, reactor trip is not encountered for any case primarily 00309 31 0 080890. 5 29 . i d
l 1 l due to the conservative safety analysis setpoints. No single active failure will prevent the
- reactor protection system from performing its intended function, if required. ;
i ! , A 10 percent step increase in steam demand is assumed, and all cases are analyzed ! ! without credit being taken for pressurizer heaters. The cases which assume automatic ! rod control are analyzed to ensure that the worst case is presented, although the l automatic function is not required. l l t F 5.1.3.3.3 Results ; i All of the results presented here are for four loop operating conditions, From the results of l previous licensing basis analysis performed specifically for Millstone Unit 3, it has been ! determined that the four loop results bound those with three loops operating. This { conclusion is based on the higher initial DNBR for three loop operation and the fact that , the excessive load increase event does not produce a significant erosion in that initial , difference. Therefore, no explicit three loop results are reported in this section. ! I - i Figures 5.1.3 3 through 5.1.3 6 illustrate the transient results with the reactor in the manual l control mode. As expected, for the minimum moderator feedback case there is a slight t power increase and a large average core temperature decrease Thil results in a ONBR I which increases above its initial value. For the maximum moderator feedback, manually I controlled case there is a large increase in reactor power due to the moderator feedback. [ A reduction in DNBR is experiencad but DNBR remains above the limit value. Figures 5.1.3 7 through 5.1.310 illustrate the transient results assuming the reactor is in l the automatic control mode. Both the minimum and maximum moderator feedback cases ! show that core power increases, thereby reducing the rate of decrease in coolant average l temperature and pressurizer pressure. For both of these cases, the minimum DNBR j remains above the limit value. - i For all cases, the plant rapidly reaches a stabilized condition at a higher power level ! corresponding to the increase in steam flow, Neither the primary nor secondary system ' pressures approach their respective overpressure limits. Reactor trip is not predicted to j occur for any of the cases analyzed. However, due to conservative safety analysis ; setpoints assumed, it is possible that reactor trip could actually occur for the automatic , control cases, Normal plant operating procedures would then be followed to reduce : power; the plant would then reach a stabilized condition following the trip, As noted - earlier, the four loop results are bounding with respect to three loop operation. - The calculated sequence of events for the excessive load increase incident are shown in i Table 5.1.31. ; 5.1.3.3.4 Conclusions The analysis results presented show that for four loop operation with a ten percent step- i load increase, the DNBR remains above the limit value (see Table 4 2), thereby ensuring - J l 00309 3:1D'080890 5 30 -! l J - . _ _ - - .-. -. . - - - ~ ... . - -. _
i 5 that the DNB design basis is met. Also, neither the primary nor secondary system pressures exceed 110% of the design values. The results of the analysis also ; demonstrate that for the modeled transient, the plant reaches a stabilized condition l I l following the load increase without a reactor trip occurring. Because the four loop results are bounding with regard to three loop operation, tne conclusions of the FSAR continue to j be vaild for both three and four loop operation. ! 5.1.3.4 Inadvertent Ooenina of a Steam Generator Relief or Safety Valve j F 5.1.3.4.1 Introduction .; As described in FSAR Section 15.1.4, the most severe core conditions resulting from an i accidental depressurization of the main steam system are associated with an inadvertent { i opening of the largest cf any single steam dump, relief, or safety valve. The main steam system depressurization results in an initial steam flow increase which decreases during l the accident as the steam pressure falls. The energy removal produces a cooldown and . depressurization of the RCS in the presence of a negative moderator temprature '! l coefficient the cooldown then results a positive reactivity insertion. [ Accidental depressurization of the r ;eam system is classified as an ANS Condition 11 event, i.e. a fault of moderate frequ, .f. The general acceptance criteria for this event { category are discussed in FSAR Section 15.0.1.2. The event is analyzed to demonstrate l that the DNB design basis is satisfied assuming hot shutdown conditions with a stuck l RCCA, offsite power available, and a single failure in the the engineered safety features # l system. The primary plant safety features that provide protection against an accidental j depressurization of the main steam system due to the opening of a steam generator safety , or relief valve are as follows:
- 1. Safety injection System actuation from low pressurizer pressure or low steamline l pressure signals i
- 2. Reactor trip from high neutron flux, overpower AT, or in conjunction with the l safety injection signal, t
- 3. Redundant isolation of the main feedwater lines to minimize the additional cooldown that would otherwise occur from continued main feedwater flow, i
- 4. Shutting r/ the fast acting steam line stop valves I
t ! This list summaritas material presented in greater detail in FSAR Section 15.1.4.1. t O : 00309 3 1o:000890 5 31 :
,a , - - .,..,L.
l i l l l i 5.1.3.4.2 Method of Analysis The accidental depressurization of the main steam system is explicitly ant lyzed for both l four and three loop operation using the same basic methodology as is currently described l in FSAR Section 15.1.4.2. As noteci earlier, the primary intent of the analysis of this event j is to verify that the DNB design basis is met throughout the course of the transient. Of ! course, it is also necessary to confirm that no other RCS or secondary design system i design limits are exceeded. j l t i Consistent with FSAR Section 15.1.4.2, the analysis assumes that the steam system . : depressurization begins with the reactor at hot zero power conditions, the control rods fully ; inserted in the core and the most reactive rod stuck out. The analysis of this accident is : performed using STDP (not RTDP). Though the primary methodology of FSAR Section ; 15.1.4.2 has been followed in the analysis of this event, a number of revised safety { analysis assumptions discussed in Section 5.1.1.2 directly impact the accidental l depressurization of the main steam system. In some cases those revised assumptions l l contradict the current FSAR information. The key modified assumptions for this event are l as follows: ;
- 1. A negative troderator coefficient, corresponding to the EOL rodded core with the most reactive RCCA assembly in the fully withdrawn position is assumed. The :
analysis includes coefficient variation with temperature and pressure. The Ken j versus temperature at 1000 psia corresponding to the negative moderator temperature coefficient used is shown on Figure 5.1.311. !
- 2. Reflecting the assumption of minimum capability for safety injection flow, the previous analysis for this event had taken credit for minimum safety injection flow as provided by a single centrifugal charging pump. The a 1alysis performed for ,
this report instead credits the minimum safety injection flow as provided by a i single high head safety injection pump, which still represents a conservative analysis model. I
- 3. The analysis of this event conservatively assumed a 42.0 second delay in SI l Initiation for the cases with offsite power available as well as for those l l considering the loss of offsite power. Diesel generator starting and sequence l loading delays are included in the 42.0 seconds. ;
- 4. A reduced low steam pressure setpoint of 435 psia is used for safety injection :
and steamline isolation in contrast to the previously used 444 psia setpoint.
- 5. A reduced shutdown margin of 1.3% Ak is used in contrast to the previously. ;
used shutdown margin of 1.6% ak. '
- 6. A redu;ed auxiliary feedwater flow rate of 1200 gpm for four loop operation and !
900 gpm.for three loop operation is assumed. D0309 3.10 080890 5 32
l i q l 1 0 i l The specific steam system depressurization considered is a steam flow of 277 lbm/sec at i 1200 psia pressure with offsite power available and corresponds to a safety valve being ' i stuck open. This flowrate is the maximum capacity of any single steam dump, safety, or. ;
+
i relief valve. 5.1.3.4.3 Results , The calculated sequence of events for this accident is listed in Table 5.1.31, The results presented for this event are a conservative indication of the transient which would occur for the subject main steam system depressurization since it is postulated that all the assumed bounding conditions described above and in FSAR Section 15.1.4.2 occur . simultaneously. Figures 5.1.312 and 5.1.3-13 show the results for a steam flow of 277 lbm/sec at 1200 . psia from one steam generator with the plant operating under four loop conditions. Figures 5.1.312A and 5.1.313A present the' corresponding results for three loop operating conditions. For the three loop case,' the steam release is assumed to occur on one of the operating loops. l Safety injection is initiated automatically by the low pressurizer pressure function, with the flow assumed to be provided by one of the high head safety injection punips. The 2500 ppm boron provided from the RWST enters the RCS rapidly enough and provides O sufficient negative reactivity to prevent core damage. This is confirmed by verifying that the DNB design basis is satisfied throughout the course of the event. The calculated transient is conservative with respect to the predicted cooldown, since no credit is taken for the considerable amount of energy stored in the system metal, other than that of the fuel elements or the energy stored in the steam generator tubes. 'Since the limiting portion of the transient occurs over a period of five to ten minutes, the neglected stored energy would likely have a significant effect in slowing the cooldown. < 5.1.3.4.4 Conclusions The analyses show that for both three and four loop operation, the DNB design basis (see Table 4 2) is met for the accidental depressurization~of the main steam system event. i Additionally, it is noted that this event is less limiting than the rupture of a main steam line presented in Section 5.1.3,5. , 5.1.3.5 Steam System Pioina Failure t l 5.1.3.5.1 Introduction As described in FSAR Section 15.1.5, the steam release arising from a main steam line rupture produces an initial increase in steam flow followed by a gradual decrease during the accident as the steam pressure falls. The energy removal produces a cooldown and depressurization of the RCS. In the presence of a negative moderator temperature O' coefficient the cooldown then results in a' positive reactivity insertion. If the most reactive v h 00309 31 0'080990 5 33 i
,n , -
RCCA is assumed stuck in its fully withdrawn position after reactor trip, there is an increased possibility that the core could become critical and return to power. A post trip return to power following a steam line rupture is a potential problem.mainly because of the high power peaking factors induced by assuming the most reactive RCCA to be stuck in its fully withdrawn position. The core is ultimately shutdown by the boric acid injection l delivered by the safety injection system. A major steam line rupture is classified as an l ANS Condition IV event, i.e. a limiting fault. The general acceptance criteria for this event ; . category are discussed in FSAR Section 15.0.1.4. As noted in FSAR Section 15.1.5.1, I although DNB and possible clad perforation may be acceptable following a main steam line rupture, the analysis does, in fact, show that the DNB design basis (see Section 4) la met for any such rupture even assuming the most reactive RCCA stuck in its fully , withdrawn position. l The primary plant safety features that provide protection against the consequences of a main steam line rupture are as follows.
- 1. Safety injection System actuation from low pressurizer pressure, low steamline pressure, or Hi 1 containment pressure signals.
- 2. Reactor trip from high neutron flux, overpower AT, or in conjunction with the safety injection signal, j i
- 3. Redundant isolation of the main feedwater lines to minimize the additional cooldown that would otherwise occur from continued main feedwater flow. ;
- 4. Shutting of the fast acting steam line stop valves This list summarizes material presented in greater detail in FSAR Section 15.1.5.1. For steam line ruptures downstream of the isolation valves, closure of all valves would '
completely terminate the blowdown. The limiting single failure is loss of a single train of ESF. All MSIVs are assumed to close. For any break, in any location, no more than one steam generator would blow down even if one of the isolation valves fails to close. A description of steam line isolation is included in Chapter 10 of the FSAR.: i Steam flow is measured by monitoring dynamic head inside the steam pipes. Nozzles that are of considerably smaller diameter than the main steam pipe are located in the I steam generators and serve to limit the maximum steam flow for any break at any location. i 5.1.3.5.2 Method of Analysis The steam pipe rupture event was explicitly analyzed for both four and three loop 3 operation using the same basic methodology as is currently described in FSAR Section 15.1.5.2. As noted earlier, the conservative criterion applied to this event is to verify that the DNB design basis is met throughout the course of the transient. 00309-31 0 080990 5 34 l t l
1 l i O The specific cases considered here and in the FSAR to address verification of the DNB design basis are as follows: l l
- 1. Complete severence of a main steam pipe with the plant initially at no load [
' conditions. full reactor flow and offsite power available. j i 2. Case 1 with the loss of offsite power occurring simultaneous with the steam fine i break and initiation of the safety injection signal. The loss of offsite power results ! in the coastdown of the reactor coolant pumps. j i
- 3. Cases 1 and 2 with one reactor coolant loop out of service, Consistent with FSAR Section 15.1.5.2. the analysis assumes that the steam pipe rupture l event begins with the reactor at hot zero power conditions. the control rods fully inserted in i the core with the most reactive rod stuck out. The analysis of this accident is performed '!
using STDP (not RTDP). Though the primary methodology of FSAR Section 15.1.5.2 has i been followed in the analysis of this event, a number of revised safety analysis ' , assumpt!ons discussed in Section 5.1.1.2 directly impact the main steamline rupture. In l some cases those revised assumptions contradict the current FSAR information. The key modified assumptions for this event are as follows: l l
- 1. A negative moderator coefficient, corresponding to the EOL rodded core with the !
most reactive RCCA assembly in the fully withdrawn position is assumed. The O analysis included coefficient variation with temperature and pressure. The Keff versus temperature at 1000 psia corresponding to the negative moderator l i temperature coefficient used is shown on Figure 5.1.311. The effect of power , generation in the core on the overall reactivity is shown in Figure 5.1.314. ,
- 2. Reflecting the assumption of minimum capability for safety injection flow, the i previous analysis for this event had taken credit for minimum safety. injection flow !
as provided by a single centrifugal charging pump. The analysis performed for . this report instead credits the minimum safety injection flow as provided by a ! single high head safety injection pump, which still represents a conservative l analysis model.
- 3. The analysis of this event conservatively assumed a 42.0 second delay in Si i initiation for the cases with offsite power available as well as for those i considering the loss of offsite power. Diesel generator starting and sequence- l loading delays are included in the 42.0 seconds. ;
- 4. A reduced low steam pressure setpoint of 435 psia was used for safety injection ;
and steamline isolation in contrast to the previously used 444 psia setpoint. I 1
- 5. A reduced shutdown margin of 1.3% ak is used in contrast to the previously . ]
used shutdown margin of 1.6% ak.
l l .1 l
- 6. A reduced auxiliary feedwater flow rate of 1200 gpm for four loop operation and 900 gpm for three loop operation is assumed.
O! l l r 5.1.3.5.3 Results l The calculated sequence of eventu for this accident is listed in Table 5.1.31. The results -l presented for this event are a conservative indication of the transient which would occur i for the subject main steam system depressurization since it is postulated that all the assumed bounding conditions described above and in FSAR Section 15.1.5.2 occur .! simultaneously. l Figures 5.1.315 through 5.1.317 show the response of pertinent system parameters j foi;cwing a main steam line rupture with offsite power available such that full reactor i coolant flow exists. The transient shown assumes an uncontrolled steam release from v only one steam generator, as discussed in FSAR Section 15.1.5.2. Figures 5.1.315A i through 5.1.317A present the corresponding results for three loep operating conditions. )
- As shown in Figures 5.1.315 and 5.1.315A, for both four and three loop cases, the core j attains criticality with the RCCAs inserted, assuming the design shutdown margin and one ;
stuck RCCA. Criticality occurs before boric acid solution at 2500 ppm enters the RCS ; from the safety injection system that is injecting fluid from the RWST. A peak core power ; well below nominal full power is predicted for both the four and three loop cases. The calculation of core power assumes the boric acid is mixed with and diluted by the ! water flowing in the RCS prior to entering the reactor core. The concentration after mixing l depends upon the relative flow rates in the RCS and in the safety injection systems The ! variation of mass flow rate in the RCS due to water dimsity changes is included in the ! calculation as is the variation of flow rate from the safety inject. ion system reflecting ! changes in the RCS pressure. The safety injection system flow calculation includes the l line losses in the system as well as the pump head curve. The accumulators provide an i additional source of borated water if the RCS pressure decreases below 600 psia. Explicit results are not presented for the cases with loss of offsite power since, as ! discussed in FSAR Section 15.1.5.2, it has been demonstrated that the loss of offsite i power case is bounded by the analysis with offsite power, it should be noted that following i a steam line break only one steam generator blows down completely. Thus, the remaining intact steam generators are available for dissipation of decay heat after the , initial transient is over, in the case with loss of offsite power, the decay heat is dissipated i into the atmosphere via the steam line safety valves and the power operated relief valves. j The results of the limiting case with offs!!e power were then used to perform a DNB i analysis. The DNB design basis is discussed in Section 4.2. : 00309 31 0 080890 5 36 l
. - - _1 ._ _
. . . . - .-. .. . . . _ - . - . - . . . . . . -. ~ -- - .-
i .i ! The steam generator mass and energy release data for inside'and outside containment , i have been evaluated. The results of the evaluation are reported in Section 5.3.2. The ; ] radiological consequences of the mass and energy release associated with this event are reported in Section 5.4.2. ! 5.1.3.5.4 Conclusions The analyses show that for both three and four loop operation, the DNB design basis (see ! Table 4 2) continues to be met for the rupture of a main steam line. Therefore, the ' conclusions of the FSAR for this event remain valid.
'\
i O : i
+
. I 3
i i I r D0309 3:10:08'J890 5 37 :
. . _ _ , . , ..v
, , - - y , , . ~ . . . - - -
-.,J,.. ., ,- ...,r., - , . - . , _ _. . .
i e i TABLE 5.1.31 TIME SEQUENCE OF EVENTS FOR INCIDENTS THAT HESULT IN AN INCREASE IN HEAT REMOVAL BY THE SECONDARY EYSTEM N Loop N1 Accident Event Time (sec) (sec) Feedwater system One main feedwater control valve falls . 0.0 0.0 _ ! malfunctions producing fully open increased feedwater flow: ' High neutron flux setpoint reached -- 15.1-High high steam generator water level 103.2 21.1 ; signal generated Turbine trip occurs due to high high 105.7 15.6 steam generator level Minimum DNBR occurs g2.5 16.5 i Reactor trip occurs 107.2 15.6 Feedwater isolation valveo close 110.2 28.1 automatically Excessive increase in : secondary steam flow:
- 1. Manual reactor 10 percent step !oad increase 0.0 (1) i (minimum mode'rator feedback) Equilibrium conditioris reached 180 t (approximate time only) '
- 2. Manual reactor 10 percent step load increase 0.0 (maximum moderator i feedback) Equilibrium conditions reached 50 ,
(approximate time only) 3 Automatic react 7 10 percent step load increase 0.0 control (minimum moderator feedback) Equilibrium conditions reached 100-(approximate time only) :
- 4. Automatic reactor 10 percent step load increase 0.0 control (maximum moderator feedback) Equilibrium conditions reached 50 (approximate time only) !
(1) Explicit three loop analysis not performed for this event; the four loop results bound those for three loops operating. 'l 00309 3 10 080990 5 38 , l
. - . = , . .
L TABLE 5.1.31 (conQ TIME SEQUENCE OF EVENTS FOR INCIDENTS.T!iAT RESULT IN AN - INCREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM
"'I
- N Loop '-
Accident Event. ; Time n (sec)- (sec)- Inadvertent opening of a Inadvertent opening of cao main steam - b.0 0.0 steam generator relief or safety or relief valve safety valve: Pressurizer empties 328.2 333.2 Criticality attained 334 7- 289,7 2500 ppm boron reaches core - 454,7 = NA" Steam system piping failure:
- 1. Offsite power Steam line ruptures 0.0 t. 0.0 '
Criticality attained 20.0 20.0-O- Pressurizer empty _ . 25.0 35.2 ~ 2500 ppm boron reaches core : 78.0; ' 86.0 ' l
- 2. With loss of offsite Steam line ruptures. 0.0 i : 0.0 '
power: Criticality attained 22.0- 26.0 Pressurizer empties 28.6 50.4 2500 ppm boron reaches core 1122.0- NA t 00309 3.10 080990 -5 39_ .i
)
7__.__. t I l 5.1.4 Decrease in Heat Removal by the Secondary System. A number of transients and accidents have been postulated which result in a reduction of i U the capacity of the secondary system to remove heat generated in the reactor coolant system. These events are discussed in this sections ,
~
- 5.1.4.1 Steam Pressure Reaulator Malfunction or Failure that Results'in Decreamino - ;
Steam Flow
.l As stated in the FSAR, there are no steam pressure regulators in Millstone Unit 3 whose .
failure or malfunction could cause a steam flow transient. .
]
5.1.4.2 Loss of External Electrical Load l ! As discussed in FSAR section 15.2.2, this transient is bounded by the analysis for the turbine trip event. The turbine trip event is described in the next section (5.1,4.3). , 5.1.4.3 Turoine Trio - 1 ' l 3 ! 5.1.4.3,1 Introduction As described in FSAR Section 15.2.3, for a turbine trip event, the reactor would be tripped ' directly (unless below approximately 50 percent power) from a signal derived from ther turbine stop emergency trip fluid pressure and turbine stop valves. The turbine stop 'j ' valves close rapidly (typically 0.1 second) on loss of trip fluid pressure actuated by one of a number of possible turbine trip signals. Among the' turbine trip initiation signals noted in- , the FSAR are generator trip, low condenser vacuum, turbine overspeed and manual trip. , 1 Upon initiation of stop valve closure, st6am flow to the turbine stops abruptly. The loss of 1 steam flow results in an almost immediate rise'in secondary system temperature and ' pressure, with a resultant increase in primary system temperature and pressure. Due to a 4 more rapid loss of steam flow caused by faster valve closure, the turbine trip event produces a slightly more severe transient than the loss of' electrical load event. The turbine trip event is classified as an. ANS Condition 11 event.' a fault of moderate 'I frequency. The general acceptance criteria for this event category are discussed in. , FSAR Section 15.0.1.2. The specific acceptance criteria applied to this event are that the DNS design basis be satisfied. while also assuring that the neither the primary nor I secondary system pressure limits are exceeded. The primary plant safety features' that provide protection against damage to the RCS or .I the steam system-in the event of a turbine trip event are as follows: 3
]
- 1. High pressurizer pressure reactor trip ;
- 2. Overtemperature AT reactor trip
- j. D0309 41D.080890 5 40 i
- 3. Overpower AT reactor trip
- 4. Opening of the pressurizer safety valves y 4
- 5. Opening of the steam generator safety valves
- 6. Low low steam generator water level reactor trip 5.1.4.3.2 Method of_ Analysis
, H The turbine trip event was explicitly analyzed for both four and three loop operation using the same basic methodology as is currently described in FSAR Section 15,2.3, However,.
some of the revised safety analysis assumptions discussed in Section 5.1.1.2 of this report directly impact the response of the plant to this accident. For this event, the behavior of the unit is analyzed for a complete loss of steam load from nominal full power, without direct reactor trip, primarily to show the adequacy of the pressure relieving devices, and also to demonstrate core protection margins.L That is, the turbine is not tripped until conditions in the RCS result in a trip.- The turbine is assumed to-trip without actuating all the turbine stop valve limit switches, This assumption delays , reactor trip until conditions in the RCS result in a trip due to other signals. Thus, the analysis models a ivorst case transient ' In ac: m, no credit is taken for the steam dump system or the steam generator power operated relief valves. Main feedwater flow is terminated at the time of turbine trip l with no credit taken for auxiliary feedwater (except for long term recovery) to mitigate the consequences of the transient. ; The analysis performed for this event to support the VANTAGE 5H transition considers the
)
same set of cases currently defined in the'FSAR. That is, both minimum and maximum
~
reactivity feedback conditions are considered. Because of the potential for delaying reactor trip and thereby producing worse results, cach of the reactivity cases are in turn analyzed for two modes of pressurizer spray and_ power operated relief valve operation.
- 1. Full credit is taken for the effect of pressurizer ' spray and power operated relief valves in reducing or limiting the coolant pressure, Safety valves are also -
available.
- 2. No credit is taken for the effect of pressurizer spray and power operated relief.
valves in reducing or limiting the coolant pressure, J Safety valves are operable. As noted above, the revised safety analysis' assumptions of Section 5.1.1.2 dictate modifying certain assumptions that are part of the current licensing basis for the turbine O trip event presented in Ssetion 15.2.3 of the FSAR. The key modified assumptions for this event are as follows: 00309-41 0,080990 5 41 1 r ,
1
'l
,' 1. For the VANTAGE SH transition, RTDP methodology was employed in analyzing ! this event, which is a change from the previous STDP, methodology currently _ ' presented in the FSAR. Consistent with the RTDP approach nominal values are-assumed for the initial reactor power, pressure, and RCS temperature while - minimum measured flow is assumed rather than the. thermal design flow applicable to STOP. Uncertainties in initial conditions are included in,the limit I DNBR as discussed in Section 4.0 of this report. The specific initial conditic.ns i and plant characteristics are discussed in Section 5.1.1.3. With regard to the use of nominal values for l'nitial conditions, it is noted'thatu previous studies have shown that the peak pressurizer pressure reached for the turbine trip event is relatively insensitive to the initial conditions'of temperature , l and pressure, and peak pressurizer pressure is only slightly' sensitive to the - initial power condition. Therefore, the use of these initial conditions is appropriate for trils event.' -l
- 2. The turbine trip is analyzed with both maximum and minimum reactivity 5 feedback. The minimum moderator feedback used in.the' VANTAGE 5H analysis is consistent with the FSAR licensing basis analysis for standard fuel. .j However, the most negative moderator temperature coefficient used for -
VANTAGE 5H effort is more limiting than that used previously.? 3.
~
The VANTAGE 5H turbine trip analysis includes revisions in'the treatment of safety valve drift for both the pressurizer and steam generator safety valves, as O discussed in Section 5.1.1.2. ' For the pressurizer safety valves this means opening at a pressurizer pressure of 2575 psia with an additional 3%1 accumulation until full flow is reached. For the; steam generator safety valves. ! the revised model assumes that the safety valve actuation will allow steam-generator pressure to rise to 1320 psia (110% of design pressure), but go no' higher. As noted in Section 5.1.1.2, this is a much more conservative model , l than used for the current FSAR licensing basis analysis. , 5.1.4.3.3 Results The transient responses for a turbine trip from nominal four loop', full power operation are I shown for four cases: two cases for minimum reactivity feedback and'two cases for maximum reactivity feedback'(Figures 5.1.41 through 5.1.4-8). : Results for the four. )
- corresponding cases are also shown for a turbine trip from nominal three loop, full power l operation.(Figures 5.1.4.1 A through 5.1.4 8A). l Figures 5.1.41,5.1.4 2,5.1.4-1 A. and 5.1.4 2A show the four and three loop transient'
, responses for the turbine trip event with minimum reactivity. feedback, assuming full' credit I I for the pressurizer spray and pressurizer power operated relief valves. No credit is taken l L for the steam bypass. The_ reactor is tripped by the high pressurizer pressure trip D channel. The minimum DNBR remains well above the limit value_(soe Table 4 2). The 1 00309-41 0,080890 5 42 ' q I4 '_ , ___U______[ '
)
t' pressurizer and steam generator. safety valves prevent overpressurization of the primary. O and secondary systems.- j Figures 5.1.4 3,5.1.4 4,5.1.4 3A. and 5.1.4 4A show the four and three loop responses for -- ( the turbine trip event with maximum reactivity feedback. All other plant parameters are the same as the above. The DNBR increases throughout the transient and never drops below ~ 1 its initial value. The pressurizer power operated relief valves and steam generator safety i ' valves prevent overpressurization in tha primary and secondary systems, respectivel". The reactor is tripped by the high pressurizer pressure trip channel. . The pressurizer- ,( safety valves are not actuated for this case. .[ t The turbine trip accident was also analyzed assuming the plant to be inidally operating at ; nominal full power with no credit taken for the pressurizer spray, press'.srizer power. : operated relief valves, or the turbine bypass system. The reactor is tr#pped on the high pressurizer pressure signal. Figures 5.1.4 5, 5.1.4-6, 5.1.4 5A, and 5.1.4 6A show the four . ; and three loop transient responses with minimum reactivity feedbacl<. The DNBR never ; goes below its initial value throughout the transient. In this case the pressurizer safety. valves and steam generator safety valves are actuated to maintain the RCS and main i steam system pressure below 110 percent of their respective design value. , Figures 5.1.4 7,5.1.4 8,5.1.4 7A and 5.1.4 8A'show the four and three loop transient h responses with maximum reactivity feedback while the other assumptions remain the - ! same as in the preceding case. The reactor is tripped by the high pressurizer pressure
~
{ ( signal and the DNBR increases throughout the transient / The pressurizer safety valves and steam generator safety valves are actuated to limit primary and secondary system' ; pressures. : The calculated sequence of events for the turbine trip incident are shown in Table'5.1.4-1. f 5.1.4.3.4 Conclusions 1
)
Results of the analyses show that the plant design is such that a turbine trip without a . j direct reactor trip presents no hazard to the integrity of the RCS or the. main steam j system. Pressure relieving devicer incorporated in the two systems'are adequate to limit the ma.ximum pressures to within the design limits. The integrity of the core is also - , maintained since the DNBR remains above the limit value (see Table 4 2). Thus the conclusions presented in the FSAR remain valid for.the turbine trip event. n v 5.1.4.4 Inadvertent Closure of Main Steam isolation Valves , j See discussion in Section 5.1.4.5 below.- '. i~ : l ' 00309 4:10 080890 .5 43 ,
1 5.1.4.5 Loss of Condenser Vacuum and Other Events Resultino in a Turbine Trio ! Inadvertent closure of the main steam isolation valves, loss of condenser vacuum, and other events resulting in a turbine trip are all bounded by the analysis in 5.1.4.3. Additional information for these event categories is provided in FSAR Sections 15.2.4 and 15.2.5.' 5.1.4.6 Loss of Non Emeroency AC Power to the Plant Auxiliaries-l A complete loss of non emergency AC power to the' plant auxillaries accident is described l! in Section 15.2.6 of the FSAR. This event may result in the loss of all power to the station auxiliaries: 1.e., the reactor coolant pumps, condensate pumps, etc. The loss of power ; may be caused by a complete loss of the offsite grid accompanied by'a turbine generator l trip at the station, or by a loss of the onsite AC distribution system. Upon the loss of power to the reactor coolant pumps, the RCS flow necessary for core cooling and the removal of residual heat is maintained by natural circulation in the reactor
~
coolant loops. A loss of non emergency AC power to the station auxiliaries is_ classified as an ANS Condition 11 event, a fault of moderate frequency. The general acceptance ' criteria for this event category are discussed In' FSAR Section 15.0.1.2. This is a long-term heat removal event which is specifically analyzed to determine if the auxiliary - feedwater (AFW) system is capable of removing the stored and residual heat, tnereby preventing either RCS overpressurization or water relief from the pressurizer (assuring that the core remains covered). As discussed in FSAR Section 15.2.6.1, a loss of AC power to the station auxillaries as t described above could also result in a loss of normal feedwater if the condensate pumps l lose power to operate. A loss of normal fe * ; water is the mo'st limiting Condition ll event in the decrease in secondary heat removal ct.tegory and, in Section 5.1.4.7 of this report, is ; analyzed for a case which includes the loss of AC power. Since these results of Section i 5.1.4.7 are bounding, just as in the FSAR, no detailed analytical results are presented here for the loss of non emergency AC power to the station auxiliaries. 5.1.4.7 Loss of Normal Feedwater Flow 5.1.4.7.1 Introduction The loss of normal feedwater flow accident is described in Section 15.2.7 of the FSAR. This event, which may arise from pump failures, valve malfunctions, or the loss of AC power, results in a reduction in the capability of the secondary system to remove the heat i generated in the core. If the reactor were not tripped during this accident, core damage could possibly occur from the sudden loss of heat sink. If an alternative supply of feedwater were not supplied to the plant, core residual heat following reactor trip would : heat the primary system water to the point where water relief from the pressurizer would occur. A sicmificant loss of water from the RCS could conceivably lead to core damage.- It should be noted that for this event, the plant is tripped well before the steam generator. ' l
- D0309 410 080990 5 44
)
4
I 1 heat transfer capability is reduced and the primary system conditions never approach a j violation of the DNB design basis. q The loss of normal feedwater event is classified as an ANS Condition 11 event, a fault of ~ _l moderate frequency. The general acceptance criteria for this event category are i discussed in FSAR Section 15.0.1.2. This is a long term heat removal' event which is . ] specifically analyzed to determine if the auxiliary feedwater (AFW) system is capable of . { l removing the store'd and residual heat, thereby preventing either RCS overpressurization-or water relief from the pressurizer, thus assuring that the core remains covered. The analysis assumes that the reactor is protected for the loss of normal feedwater event by reactor trip on a low low narrow range water level signal in any steam generator. The - ; AFW system is started automatically by the signals identified in FSAR Section 15.2.6.1. ) 5.1.4.7.2 Method of Analysis The loss of normal feedwater event was analyzed using the same_ basic methodology as is j currently described in FSAR Section 15.2.7.2. The primary criterion used to establish the- ', acceptability of the accident analysis for this event is that the pressurizer shall not be , allowed to fill (i.e., become water solid).following the loss of normal feedwater. . This event is not limiting with respect to DNB concerns,' and is therefore not analyzed using RTDP. Instead the analysis continues to employ STDP methodology, consistent with the current
"*^"
O Both four and three loop results are currently'_ documented in the FSAR. The 1 ' VANTAGE SH transition analysis in this report presents an analysis for four loop operation - only. From the licensing basis FSAR analysis for Millstone Unit 3.-it has been oetermined' that the four loop results also bound those with three loops operating for the loss of normal 'i feedwater event. The FSAR results specifically show that for three loop operation with a i nominal power level of 75% rated thermal power,- the initial. level in the pressurizer is lower 'i than for four loop. Additionally, for the three loop loss of normal feedwater event, the' absolute increase in the pressurizer water volume during the transient is less than or- : equal to that with four loop operation. The net result is that the three loop transient is clearly less limiting. The basic methods of FSAR Section 15.2.7.2 were used, although the analysis reflects the $ changes dictated by the modified safety analysis assumptions of Section 5.1.1.2. Most _of 1 the FSAR assumptions for this event remain unchanged for the loss of normal feedwater analysis of this report, even though a number of the safety analysis assumptionc ofj i Section 5.1.1.2 could impact the results and are identified below: . 1
- 1. Secondary system steam release, as modeled in the-analysis, continues to be ' !
through the steam generator safety valves. However,lthe performance of these valves is modeled as being consistent with allowing the steam generator pressure to reach 1320 psia, as noted in the increased safety valve drift discussion of Section S (; 1 5.0.5. L l , l 00309 4 10 080990 5 45
!l
- 2. The initial RCS average temperatures used in the analysis reflect ine revised non- '
RTDP uncertainties indicated in Sections 5.1.1.2 and 5.1.1.3. Consistent with the j methodology of the FSAR, the initial RCS average temperature is 6.0'F higher than . i nominal for the case where off site power is maintained. This results in greater : l expansion of the RCS fluid from pump and decay heat during the transient, and -l subsequently, a higher water level in the pressurizer.' For the. case with loss of ! offsite power, the initial average temperature assumed is 6.6'F lower than nominal-since this results in greater RCS fluid density with lower natural circulation. ;
- 3. The initial pressurizer pressure used in the analysis reflects the revised non RTDP ,
uncertainties indicated in~5.1.1.2 and 5.1.1.3. Consistent with the methodology of the - FSAR, the initial pressurizer pressure is 50 psia higher than nominal; 1
- 4. An auxiliary feedwater flow rate of 510 gpm ' delivered to'four steam generators is assumed for four loop operation. !
Additionally,'It is noted that a reactor trip on low low steam generator water level was the protection system function assumed to actuate for this event. A trip setpoint of 10% narrow range span was assumed, as indicated in Table 5.1.14. I 5.1.4.7.3 B g.3,y.1,tg Figures 5.1.4 9 and 5.1.410 show the behavior of significant plant parameters following a loss of normal feedwater with four loops in. operation and offsite power available. Figures ; 5.1.411 and 5.1.412 present the same plant parameters following a loss of normal. ! feedwater with four loops initially operating, with a subsequent loss of offsite power. ! Following the reactor and turbine trip from full load, the water level in the steam generators will fall due to the reduction in steam generator void fraction and because steam flow through the safety valves continues to dissipate the stored and generated heat. .One . ' l minute following the initiation of the low low steam generator water level trip, both motor-I driven auxiliary feedwater pumps are automatically started, thereb'y reducing the rate of water level decrease. L The capacity of the auxiliary feedwater pumps i.s such that the. water level in-the steam generator being fed does not drop below the'the lowest level at which sufficient heat transfer area is available to dissipate core residual heat without water relief from the RCS . relief or safety valves. That is, from the results shown in Figures 5.1.4 9 through 5.1.412, it can be seen that at no time during the transient does the pressurizer fill. 1 The calculated sequence of events for this transient with four loop operation is shown in Table 5.1.41. Using this table and the associated figures noted above, it is seen that the plant approaches a stabilized condition.following reactor trip and auxiliary feedwater' , initiation. Standard plant procedures may then be.followed'to further cool down the plant. E f D0309-410'080890 5 46 - 4 a_ .. __ _
5.1.4.7.4 Conclusions The results of explicit analysis for four loop operation demonstrate that a loss of normal feedwater event does not adversely affect the core, the RCS, or the steam system. Auxiliary feedwater capacity is such that reactor coolant water is not relieved from the pressurizer relief or safety valves, and the water level in all steam generators receiving auxiliary feedwater is maintained above the lowest level at which sufficient heat transfer area is available to dissipate core residual heat without water relief from the RCS relief or - safety valves. The four loop results bound those for three loops operating; hence the - results of this report confirm that the conclusions of the FSAR remain valid. 5.1.4.8 Feedwater System Pios Break 5.1.4.8.1 Introduction The major feedwater line rupture event is described in Section 15.2.8 of the FSAR. This event is defined as a break in the feedwater line large enough to prevent the addition of: sufficient feedwater to the steam generators to maintain shell side fluid inventory in the steam generators. If the break is postulated in. a feedline between the check valve and the steam generator, fluid from the steam generator may also be discharged through the break. A break upstream of the feedline check valve would affect the nuclear steam supply system only as a loss of feedwater and that case is covered by the evaluation of Section 5.1.4.7. As noted in the FSAR, depending upon the size of the break and the plant operating-conditions, a major feedline rupture could cause either a RCS cooldown (by excessive energy discharge through the break) or a RCS heatup? The p'otential cooldown resulting from a secondary pipe rupture is evaluated in Section 5.1.3.5, so that only the RCS heatup effects are evaluated for a feedline rupture. A feedwater _line rupture event reduces.the ability to remov'e heat gen 6 rated by the core. from the RCS As discussed in FSAR Section 15.2.8.1, the analysis must show that the-auxiliary feedwater system can assure that even in the event of a major feedline rupture, no substantial overpressurization of the RCS shall occur and decay heat is removed in
~
order to maintain sufficient liquid in the RCS to keep the reactor core covered. A major feedwater line rupture is classified as an ANS Condition _IV event. The general acceptance criteria for this event category are discussed in FSAR Section 15.0.1.4. To conservatively assure that the general Condition IV criteria are met, the specific criterion applied to this event is that no bulk boiling occurs in the primary coolant system following a feedline rupture, prior to the time that the heat removal capability of the steam , generators receiving auxiliary feedwater flow exceeds nuclear steam supply system heat' generation, in addition to the actuation of the auxiliary feedwater system, the following plant protection system functions can act to mitigate the effects of a major feedwater line rupture: 00309 4 10 080890 . 5 47 1
')
b
'd
-]
- 1. Reactor trip on high pressurizer pressure,'overtemperature AT, and low low steam - O 1 generator water level in any. steam generator.
- 2. Safety injection signal from low steam line pressure or high containment pressure- j (Hi 1).
l 5.1.4.8.2 Method of Analyg,ia l .
] !
l The major feedwater line rupture event is explicitly analyzed for both four'and three loop _l t operation using the same basic methodology as is currently described in FSAR Section l l 15.2.8.2. As noted above, the primary criterion for which this event is analyzed ~ relates to l bulk boiling in the RCS. . Meeting the DNB design basis'is not an acceptance criterion for .
; major feedwater line rupture,~ therefore, the analysis employs the STOP methodology, I i
i~ Consistent with the FSAR, the four cases analyzed for this event are double ended ruptures of the largest feedwater line with the following conditions.
-I
- 1. Four loop operation initially at 102% of engineered safety features design rating, with.
l offsite power available, i . 2. Four loop operation initially at 102% of engineered safety features design rating, with loss of offsite power.
- 3. Three loop operation initially at 77.3% of rated thermal power, with' offsite power available.
- 4. Three loop operation initially at 77.3% of rated thermal power, with loss of offsite .
power. I \' The use of 77.3% power for the three loop initial power l'evel is a slight increase over the 5 j 77.0% currently in the FSAR analysis, but this change' simply reflects the revised non- i
- RTOP uncertainties and associated initial conditions described in Sections 5.1.1.2 and i 5.1.1.3. respectively. For all the cases, reactor trip is assumed to be initiated _when the !
low fow steam generator water level trip setpoint ls reached in the ruptured steam generator. A conservative setpoint of 0.0% narrow range span is assumed, as indicated in Table 5.1.14. , I Modified safety analysis assumptions from Section 5.1.1.2 that represent deviations from ! those currently shown in FSAR Section 15.2.8.2 are as follows:
- 1. Initial reactor coolant average temperature is 6.0'F above' the nominal value and the initial pressurizer pressure is 50 psi above its nominal value.
- 2. The assumed initial pressurizer and steam generator levels are chosen to reflect I conservative level errors. The initial pressurizer level is at the nominal programmed, 5
,00309 41 0 080890 . 5 48 i
q
.__ M i __ n _ ___n _ __ _ .. , , ~ , --
~
value plus 6.0%; initial steam generator water level is at the nominal va'lue plus 6.0% for the steam generator in the ruptured loop and minus 6.0% for the steam generators in the intact loops.'
- 3. The steam generator safety valves are mcdeled as being consistent with allowing . I the steam generator pressure to reach 1320 psia as noted in the increased safety-valve drift discussion of Section 5.0.5.
Also, as noted in the FSAR, the only reactor trip function that is actually credited in the analysis is the low low steam generator water level signal. With regard to safety injection, the actuating signal uced is low steam line pressure; high containment pressure is not
~
modeled. Other important FSAR assumptions that continue to be used in the feedwater; line rupture analysis of this report include the two motor driven auxiliary feedwater pumps delivering 480 gpm to the three intact steam generators for the four loop operation case and 300 gpm to the two intact steam generators for the three loop case. 5.1.4.8.3 Results Predicted plant parameters following a major feedwater line rupture are shown in Figures 5.1.413 through 5.1.4 26 for initial four loop operation, while Figures 5.1.413A through 5.1.4 26A are the corresponding figures with three loops initially in operation. Results for the cases with offsite power available are presented in Figures 5.1.413 through 5.1,419 and 5.1.413A through 5.1.419A for four and three loop operation, respectively. Results _ for the cases where offsite power is lost are presented in Figures 5.1.4 20 through 5.1.4 26 and 5.1.4 20A through 5.1.4-26A for four and three loop operation, respectively. - The calculated sequence of events for all the ca ses is presented in Table 5.1.41, The system response following the feedwate'line rupture is similar for all cases analyzed. Results presented in the figures show that for all of the cases involved, the.RCS and main steam system pressure remain below 110% c' the respective design pressures.- Pressurizer pressure increases until reactor trip on low low steam generator water level occurs. Pressure then decreases, due to the loss of heat input, until the safety injection system is actuated on low steam line pressure in the ruptured loop. Coolant expansion occurs due to reduced heat transfer capability in the steam generators and the pressurizer safety valves open to maintain primary pressure at an acceptable.value. The - , eventual addition of the safety injection flow aids in cooling down the' primary system and helps ensure that sufficient fluid exists'to keep the' core covered with water. Figures 5.1.413,5.1.413A,5.1.4 20, and 5.1.4 20A show that following reactor trip, the core remains subcritical except for a brief return to criticality for the four loop case with offsite power. This reactivity excursion is due to the cooldown caused by the. steam generator blowdown, but it is terminated when boron from the safety injection system-reaches the core at approximately 206 seconds. DNBR remains above the safety analysis limit (see Table 4 2) at all times during the-T transients, as shown in Figures 5.1.419,5.1.419A. 5.1.4 26, and 5.1.t26A. 2 As stated in 1 00309-4:1D 080890 5 49
~
the FSAR, the release of radioactivity due to the steam generator blowdown for this event is less than that calculated for the steam line rupture event (Section 5.1.3.5). ' l : RCS pressure is maintained at the PORV setpoint until safety injection flow is terminated ' by the operator, as dictated by emergency operating procedures and discussed in FSAR Section 15.2.8.2. The reactor core remains covered with water throughout the transient, ! as conservatively confirmed by the analysis which predicts that no bulk boiling occurs in - the RCS. 5.1.4.8.4 Oonclusions Results of the analyses for both four and three loop operation show that for the postulated . feedwater line rupture, the assumed auxiliary feedwater system is adequate to remove i decay _ heat, prevent overpressurizing the RCS, and prevent uncovering the reactor core, l Therefore, the conclusbns of the FSAR remain valid.
-1
'i
.i I
l l ( O> 00309 4:10 080890 5 50 r
i i l TABLE 5.1.41 (Sheet 1 of 6): ; TIME SEQUENCE OF EVENTS FOR INCIDENTS ; WHICH RESULT IN A DECREASE IN HEAT ' REMOVAL BY THE SECONDARY SYSTEM . i N1- 1 N Loop Accident Event - . Time . j0P: m (sec) (,,c) - -j Turbine Trip :
- 1. With pressurizer. Turbine trip; loss of main feedwater' flow . 10.0 0.0 ; j control (minimum reactivity feedback) ..
.i High pressurizer pressure reactor trip 7.6 = 10.4- ti '
point reached ~ , Rods begin to drop 3 9.6 12.4 initiation of steam release from steam 11.5 c 14.5 - generator safety valves O Peak pressurizer pressure' occurs 11.5- ;15.0 V Minimum DNBR occurs'. 13.5 (1)- l
- 2. With pressurizer Turbine trip: loss of main feedwater flow - 0.0 : 0.0 ' .
control (maximum reactivity feedback) . High pressurizer pressure _ reactor trip - ,
- 9.2 ~ -
point reach , Rods begin to drop- 11.2 - 114.5 5 Initiation of steam release from steam-- .11.5 15.5 ' ' s generator safety valves: Peak pressurizer. pressure occurs 12.5. ;7.5 ' , Low low steam' generator water level 1 - 112.5-reactor trip point reached ' Minimum DNBR occurs (1).. (1)- i i u l
.- 00309 4:10'080890 5 , .. - + . . _ .
. . . _ . . . ~ _ _ _ ..
=
TABLE 5.1.41. (Sheet 2 of 6) .
' TIME SEQUENCE OF EVENTS FOR INCIDENTS- "
WHICH RESULT IN A DECREASE IN HEAT. . REMOVAL BY THE SECONDARY SYSTEM N Loop-N1-Accident . Event Time: - P' (sec) - (sec) - l Turbine Trip ' l
- 3. Without pressurizer- Turbine trip; loss of main feedwater flow ' 0.0 - 'O.0; control (minimum -i reactivity feedback)- High pressurizer pressure reactor trip 4.8- 6.0. - !
point reached-Rods begin to drop 6.8 8.0' Peak pressurizer pressure occurs - 9.0 10.5 l Initiation of steam release from steam' 11.5- .14.5 generator safety valves
- 4. Without pressurizer Minimum DNBR occurs :(1)- (1) 9 .
Turbine trip; loss of main feedwater flow 0.0 ~ 0.0 control (maximum - reactivity feedback) High pressurizer pressure reactor trip' 4.7 6.0 point reached Rods begin to drop . 6.7 8.0 Peak pressurizer pressure occurs . 8.5 10.0 initiation of steam release from steam :11.5 15.5 generator safety valves:- i l Minimum DNBRl occurs .(1) (1) i i e1 i 00309 4 10 080990 ~ 5 52
, q
l 9 TABLE 5.1.41 (Sheet 3 of 6)
=t TIME SEQUENCE OF EVENTS FOR INCIDENTSL 'i WHICH RESULT IN A DECREASE IN HEAT j REMOVAL BY THE SECONDARY SYSTEM l o
N Loop .
.N1 Time P Accident . Event l
- (sec) . f
- (sec)
Loss of normal feedwater Main feedwater flow stops .140 < (2) i flow and loss of offsite power Low steam generator water level tripi 59 l' Rods begin to drop- 61- a Reactor coolant pumps begin to' '63- , coastdown - J Peak water level in pressurizer occurs 74- ! 4 steam generators begin to receive . 119 i 510 gpm of auxiliary feedwater from two ; motor driven auxiliary feedwater pumps Core decay heat decreases to auxiliary- #400 ' feedwater heat removal capacity - . Cold auxiliary feedwater delivered to the 472- ', steam generators; Loss of normal feedwater Main feedwater flow stops - 10.0 (2) flow with offsite power .; Low steam generator water level trip : 58 Rods begin to drop 60 . r y f O. ; 00309 4:10,080890 ' 5 53. l
( TABLE 5.1.4-1 (Sheet 4 of 6)l TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH RESULT IN A DECREASE IN HEAT - REMOVAL BY THE SECONDARY SYSTEM N Loop N1 . Accident Event ; Time -
-(sec)
(sec): Loss of normal feedwater Peak water level in pressurizer occurs- 67 (2). flow with offsite power (continued) 4 steam generators begin to receive 118 510 gpm of auxiliary feedwater from two motor driven auxiliary feedwater pumps. Cold auxiliary feedwater delivered to the 471 steam generators Core decay heat and pump heat #2700 decreases to auxiliary feedwater heat removal capacity -1 Feedwater system pipe break 1, With offsite power Main feedline rupture occurs 10.0' 10.0 available - Low low-steam generator water level - '.16.5 17.1 reactor trip setpoint reached in ruptured steam generator Rods begin to drop .18.5 19.1 3 intact steam generators begin to 78.5 - receive 480 gpm auxiliary feedwater from two motor driven auxiliary. feedwater pumps , 2 intact steam generators begin to - 79.1 receive 300 gpm auxiliary feedwater from two motor driven auxiliary feedwater pumps O
.- 00309 4:1D.080990 '5 54 q
TABLE 5.1.41 (Sheet 5 of 6) TIME SEQUENCE OF EVENTS FOR INCIDENTS I WHICH RESULT IN A DECREASE IN HEAT' ( REMOVAL SY THE SECONDARY SYSTEM : si l N 1- } N Loop. Accident Event Time , f,0P l
- (sec) -
!(sec)
Feedwater system pipe Low pressurizer pressure safety 116.4 ,- break injection setpoint reached 118.4 77.8. l
- 1. With offsite power Safety injection pumps start '
available (continued) Low steam line pressure setpoint _ 184.1 175.8_ l reached in ruptured steam generator.
}
All main steam line isolation valves 1191.1 82.8) closed : Pressurizer power operated relief va'Ive '375.0- 241.0 ! setpoint reached - :
' Steam generator safety vaive setpoint--
658.0 L420.0 ' reached in intact steam generators . i
- 2. With loss of offsite Main feedline rupture occurs , .10 11 0 - !
power _ i Low low steam generator water. level . -16.5 L '17.1 ; reactor trip setpoint reached.in ruptured steam generator < y Rods begin to drop, power lost to the 18.5 > 19.1. l reactor coolant pumps - -) 3 intact steam generators begin to ' . ,78.5 :- receive 480 gpm auxiliary feedwaterc , from two motor driven auxiliaryL ' feedwater pumps 2 intact steam generators begin to- .- 179.1 receive 300 gpm auxiliary feedwater-from two motor driven auxiliary- - feedwater pumps 9 i
\ .
oo309ki0-080990 5 55- - L ' '
, .u.) ,
-.__l-
TABLE 5.1.41 (Sheet 6 of 6) TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH RESULT IN A DECREASE IN HEAT REMOVAL BY THE SECONDARY SYSTEM N Loop N1 Accident Event Time- '00P Time (sec) (sec) Feedwater system pipe Low steam line pressure setpoint. 60.1- 38.6' break reached in ruptured steam generator
- 2. With loss of offsite All main steam line isolation valves; 67.1 45.6 I power (continued) closed Safety injection pumps start 77.1 55.6 i
Pressurizer power operated relief valve 210.0- 133.0 setpoint reached Steam generator safety valve setpoint 339.0- 275.0 reached in intact steam generators (1) DNBR does not decrease below its initial'value O (2) Explicit three loop analysis not performed for this e,ent; the four loop results bound those for three loops operating.
+1 4
1 O> 00309 4:10 080990 -5 56- 3
y 5.1.5 Decrease in Reactor Coolant System Flowrate A number of faults which could result in a decrease in the reactor coolant' system flowrate 'i are postulated. These events are discussed in this section 5.1.5.1 Partial Loss of Forced Reactor Coolant Flow . e 5.1.5.1.1 Introduction ;
?
The partial loss of coolant flow accident is described in Section-15.3.1 of the FSAR'.' This . event may arise following a mechanical or electrical failure in a reactor coolant pump, or ; from a lault in the power supply to the pump or pumps supplied by the a reactor coolant j
?
bus. If the reactor is at power at the time of the accident, the immediate'effect of loss of- ' coolant flow is a rapid increase in the coolant temperature. This increase could result in DNB with subsequent fuel damage if the reactor is not tripped promptly. , The partial loss of forced reactor coolant flow event is classified as an ANS Condition 11 l I event, i.e. a fault of moderate frequency The ge_neral criteria for this event category are - discussed in FSAR Section 15.0.1.2. The actual limiting criterion which the analysis for . this event primarily addresses is to demonstrate that the DNB design basis is met, j Protection against the partialloss of coolant event is provided by the low primary coolant 5 flow trip signal which is actuated in any reactor coolant loop by two out of three low flow ( signals. Above Permissive 3, low flow in any. loop actuates a reactor trip. Between 4 approximately 10% rated thermal power (Permissive 7) and the' power level corresponding to Permissive 8clow flow in any two loops actuates a reactor trip. l 5.1.5.1.2 Method of Analysis The only case analyzed for this report is the .l.oss of one pump from an initial condition of four pumps operating, in FSAR Section 15.3.1, th'e analysis for this event considers both ~ l three and four loop initial operating conditions.c However, the results of that licensing basis ; analysis demonstrate that the four loop case bounds those for three loops initially operating. That conclusion is based on the higher initial DNBR for three loop operation - and the fact that the licensing basis analysis dem.onstrates that the partial loss of coolant 'j flow event does not produce a significant erosion in that initial difference. Therefore, no 3 explicit three loop results are reported in this section. The general methodology used in the analysis of this event is consistent'with that of FSAR ' , Section 15.3.1. However, as appropriate, the modified safety analysis assumptions discussed in Section 5.1.1.2 are reflected in the. reanalysis of.this event. For the analysis , of this event in support of VANTAGE SH, the initial; reactor power, pressure, and.RCS g temperatures are determined using RTDP methodology; which is a change from the o previously used STDP methodology currently presented in the' FSAR. b 00309 5 10,080890 5 57 i si *
, ! I
..J-. . _ . .
4 Consistent with the RTDP approach, nominal values are assumed for the initial reactor power, pressure, and RCS temperature while mhimum measured flow is assumed rather ; than thr' thermal design flow applicable to STDP. Uncertainties in initial conditions are 1 included in the limit DNBR, as discussed in Section 4.0 of this report. The specific initial conditions and plant characteristics are discussed in Section 5.1.1.3.j The initial. j conditions assume nominal power, nominal reactor coolant average temperature, and :l nominal reactor coolant system average pressure = Uncertainties are included in the~ determinatlan of the limit DNBR.- l Additionally, it is noted that the analysis performed for this event continues to use the -! LOFTRAN, FACTRAN,'and THINC computer codes as indicated in FSAR Section- ! 15.3.1.2. However, the associated DNBR calculations performed for the VANTAGE 5H-transition use the WRB 1 and WRB 2 DNB correlations, rather than the W 3 (R Grid) that' applies for the current FSAR analysis. Section 4.0 of this report provides a complete' discussion of this topic. 1 5.1.5.1.3 Results Figures 5.1.51 through 5.1.5-4 show the transient response for the loss ^of one reactor i coolant pump with four loops initially in operation. As indicated tiy Figure 5.1.5-4, the ! DNBR design basis is met throughout the course of the event. Since DNB does not occur, the ability of the primary coolant to remove heat from the fuel rod is not greatly reduced. Thus, the average fuel and clad temperatures do not increase significantly i above their respective initial values. The calculated sequence of events table for the analyzed four loop operation case is shown in Table 5.1.51. The affected reactor coolant pump coasts down while the _ core , flow reaches a new equilibrium value corresponding to the number of pumps:stillin operation. With the reactor tripped, a stable plant condition is eventually' reached and l normal plant shutdown may then proceed. 5.1.5.1.4 Conclusions t , The analysis shows that the minimum DNBR remains above the limit value (see Table 4 2) ' ~ at all times during the transient. Thus, no adverse fuel effects or clad rupture;is predicted, , and all applicable acceptance criteria are met.' Therefore, the conclusions of the FSAR' ' remain valid for this event. I 5.1.5.2 Comolete Loss of Forced Reactor Coolant Flow
- 5.1.5.2.1 Introduction t I
FSAR Section-15.3.2 describes a complete loss of forced reactor coolant flow, which may result from a simultaneous loss of electrical power to all reactor coolant pumps.If the reactor is at power at the time of the accident, the immediate effect of a loss of coolant 00309 5:1o.080800 5 58 l
- -.-_ __ _ f
l l l 1 l . flow is a rapid increase in the coolant temperature.xThis increase could result in DNB with ! L subsequent adverse effects to the fuelif the reactor were not tripped prompt!/. - ! I The complete loss of forced reactor coolant flow event is classified as an ANS _ Condition l 111 event. i.e. an infrequent fault. The general criteria for this event category are discussed , in FSAR Section 15.0.1.3. The actual limiting criterion that is conservatively applied to this -; event by Westinghouse is to demonstrate that the DNB design basis is met. l The signals which provide the necessary plant protection for this event at Millstone Unit 3! j are:
- 1. Reactor coolant pump underspeed j
- 2. Low reactor coolant loop flow. ;
~l These trip functions are fully described in the FSAR. For the analysis of the complete-loss of coolant flow event, the reactor trip actually assumed in the analysis is the pump -
underspeed trip. . 5.1.5.2.2 Method of Analysis l 1 The general method of analysis and the assumptions made regarding initial operating. O conditions are identical to those discussed in Section 5.1.5.1, except that following the-V loss of power supply to all pumps at power, a reactor trip is actuated on reactor coolant - pump underspeed rather than the low reactor coolant loop flow trip.- . The only case analyzed for this report is theLloss of power to all reactor coolant pumps ' from an initial condition of four pumps operating.. In FSAR Section 15.3.2, the analysis for , this event considers both three and.four loop initial operating conditions. : However, the l results of that licensing basis analysis demonstrate that the four loop case bounds that for. . l three loops initially operaung. That conclusion is based on the higher initial DNBR for _ ; , three loop operation and the fact that the licensing basis analysis demonstrates that the ' complete loss of coolant flow event does not produce a significant erosion in that initial , difference.' Therefore,- no explicit three loop results are reported in this section.- a j The !imiting case.for the complete loss of flow event is analyzed using a 0.0 pcm/*F full. j power moderator temperature coefficient. The reference-llcensing basis analysis for this - event used an overly c,nservative value of.+ 5.0 pcm/'F at full power. 5.1.5.2.3 Results i Figures 5.1.5 5 through 5.1.5 8'show the transient resp'onse for the loss of power to all' l RCPs with four loops in operation and the assumed trip on the reactor coolant pump ; underspeed function. Figure 5.1.5 8 shows the DNBR to be always greater than the safety analysis limit value for the most limiting fuel assembly cell. l,
. 00309 5.t 0 080890 5 59 1
+ , ,
j Since DNB does not occur, the ability of _the primary coolant to remove heat from the fuel ll rod is not significantly reduced. Thus, the average fuel and clad temperature do not -! increase significantly above their respective initial. values, i
-l The calculated time sequence of events for the analyzed case is shown in Table 5.1.51.
5.1.5.2.4 Conclusions i l The analysis shows that the minimum DNBR remains above the limit value at all times ; during the transient. Thus, no adverse fuel effects or clad rupture is predicted, and all _ 'j applicable acceptance criteria are met. Therefore, the conclusions of the FSAR for this , event continue to be valid. : 5.1.5.3 Reactor Coolant Pumo Shaft Seizurellocked Rotor)- . ! i 5.1.5.3.1 Introduction ' l l The reactor coolant pump shaft seizure event is discussed in FSAR Section.15.3.3. For ; .[
- the instantaneous seizure of a reactor coolant pump rotor, flow through the affected ;
reactor coolant loop is rapidly reduced, leading to a reactor trip on a low flow signal.
~
j Following the trip, heat stored in the fuel rods continues to pass into the core coolant, causing the coolant to expand. At the same time, heat transfer to the shell side of the ' steam generator is reduced, first because the reduced flow results in a decreased tube ' side film coefficient and then because the reactor coolant in the tubes cools down while : the shell side temperature increases (turbine steam flow is reduced lto zero upon' plant trip). The rapid expansion of the coolant in the reactor core,~ combined with the reduced heat transfer in the steam generator causes an insurge into the pressurizer and a pressure increase throughout the Reactor Coolant System. l i l The insurge into the pressJrizer Causes a pressure increase which in turn actuates the ! automatic spray system, opens the power operated relief valves,- and opens the
- pressurizer safety valves in a sequence dependent on the rate of insurge and pressure -i increase. The power operated relief valves are designed for reliable operation and would ,
be expected to function properly during the accident; however, for conservatism, their pressure reducing effect as well as the pressure reducing effect of the spray are not included in this analysis. The locked rotor event is classified as ari ANS Condition IV incident, i.e. a limiting fault. . The general acceptance criteria applicable to this event category are discussed in FSAR Section 15.0.1.4. The specific criteria-considered in the analysis are as'noted in the FSAR. That is, at no time during the transient can the the RCG pressure exceed that-which corresponas to the faulted condition stress limit.-'Also, the peak clad surface - temperature calculated for the hot spct must remain below 2700*F (Reference 28). Finally, the analysis for this event at Millstone Unit 3 ihcludes'the calculation-of the. percentage of fuel rods postulated to experience DNB. This value is:Used as input to the consideration of the radiological consequences associated with this event.. ! Il 00309 5:10 080890 5 60 , i
. - - , e -- u- . _. .- 'A
l l 5.1.5.3.2 Method of Analysis The general methodology used in the analysis of this event is consistent with FSAR Section 15.3.3.2. Explicit analysis.is performed for four and three loop operating : ! conditions, without offsite power available. The results for the cases without offsite power bound those with offsite power available. As in the FSAR, the analysis for this' report l includes consideration of the pressure transient, the consequences of the assumed DNB, - a conservative film boiling coefficient, a fuel clad gap coefficient that maximizes clad ! temperature. and the zirconium steam reaction. However, as appropriate, the analysis includes the design changes associated with the transition to VANTAGE 5H fuel and certain other modified safety analysis assumptions. .These items are as follows:
- 1. For the evaluation,of the pressure' transient and consideration of the peak clad ;
temperature the initial conditions continue to reflect the STDP (i.e. non RTDP) .j methodology as in.the current FSAR. The actual Initial uncertainties for core j power, reactor coolant average temperature, and pressurizer pressure reflect' j the STOP values indicated in Section 5.1.1.3..rather than those in the current I FSAR. RTDP methodology is used to calculate fraction of rods in DNB. i
- 2. As indicated in Table 5.1.14, the low reactor coolant loop' flow trip setpoint is [
85%. This is a decrease from the current FSAR setpoint of 87%.
- 3. The pressurizer (and steam generator) safety valves are modeled as discussed O
i in Section 5.1.1.2. The revised characteristics of the pressurizer safety valves are that the valves are not full open and passing their rated flow until the j pressurizer pressure reaches 2652-psla.1 As noted in FSAR Section 15.3.3.2,- , the current licensing basis analysis assumes full open/ full flow conditions at- 2 2575 psia, , The rod power at the hot spot is conservatively assumed to be 3.0 and 4.0 4. times the average initial rod power. (i.e.'FQ =' 3._0 & 4.0) for four and three loop ! cases, respectively, :The current FSAR analysis reports 2.5 and 2.75 for four , and three loop cases, respectively. ,
- 5. The full power FaH is conservatively assumed to be 1.64 and 1.76'for four and three loop operation, respectively,'fo'r the rods in DNB Analysis.
5.1.5.3.3 Results ;
- 1. Locked Rotor Analysis for Initial Four Loop Operation - .
The transient results for the four loop locked rotor accident are shown in . Figures 5.1.5. 9 through 5.1.5-12l:The calculated sequence of events for this . l l- case is shown in Table 5.1.5-1', while the results are summarized in Table 5.1.54 I l' ,
- 2. The peak RCS pressure reache'd during the transient is less than that which
\
would cause stresses to exceed the faulted condition stress limits of the ASME 00309 51 0 080890 1 15 61 . 3 a 1 ___i____i_.__________.________ _ _ . . . . _ _ . . , , , . . . , ,- a,.- .-. , y _
h Code, Section Ill. ' Also the peak clad temperature is considerably less than .! 2700'F. It should be noted that the clad temperature was conservatively t calculated assuming that DNB occurs at the initiation of the transient; These - i results represent the most limiting conditions with respect to the locked rotor.. l event or the pump shaft break. j As a result of this accident, the conservative locked rotor ' analysis predicts that , a fraction of the fuel rods will undergo DNB and are assumed to release gap 1 inventory to the reactor. 6% of the fuel rods in the core are postulated to have ; clad damage for four loop operaticin. .
?
- 2. Locked Rotor Analysis for Initial Three Loop Operation c The transient results for the three loop locked rotor accident are shown in: _
j Figures 5.1.5. 9A through 5.1.5-12A. The calculated. sequence of events for.this' case is shown in Table 5.1.51, while the results are summarized in Table 5.1.5- '
- 2. The peak RCS pressure reached is slightly higher than in the four loop case,-
but is still less than that which would cause the faulted condition stress limits to - _, ' be exceedsd. The peak transient cladding temperature remains well below the . 2700'F. i For three loop operation, the conservative locked rotor analysis predicts that a ~ ; fraction of the fuel rods will undergo DNB and are assumed .to release gap inventory to the reactor. 8% of the fuel rods in the. core m(postulated to have clad damage for three loop operation. i 5.1.5.3.4. Conclusions
~
For both the four and three loop locked rotor analyses, the peak RCS pressure reached for any of the transients is less than that which would cause stresses to exceed the faulted conditions stress limits, thereby assuring that the integrity of the primary coolant system is - maintained. Similarly, since the peak clad surface temperature calculated for the hot spot during the worst four and three loop transients. remains considerably less than 2700'F (the temperature at which clad embrittlement may be expected), the core will remain in place s and intact with no loss of core cooling capability? The percentage of. fuel rods in the core postulated to undergo.DNB during the locked rotor event is 6% and 8% for initial four and , three loop operating conditions, respectively. Therefore. the con'clusions of the FSAR. with. respect to the locked rotor event continue to be valid.
'5.1.5.4 Reactor Coolant Pumo Shaft Break h
The evaluation for this event remains consistent with that pmsented in FSAR Section , 15.3.4. That is, the model used for analyzing the locked rotor event of Section 5.1.5.3 '
- bounds the most limiting conditions with respect to either a locked rotor or pump shaft break event. With a failed shaft, the impeller could conceivably; be free to spin in the reverse direction, as opposed to being fixed in position as would occur for;a locked rotor.
00309 5:10 080890 5-62 .
. . , ,_ , 4
The effect of such reverse spinning is a slight decrease in the end point core flow when: compared to the locked rotor. The model for the analysis of Section 5.1.5.3 explicitly . includes this effect, so that the re Its bound the most limiting conditions for the locked . rotor and shaft break; therefore, a separate analysis for the shaft break is not presented. O > O u 00309 5:10,080890 5 63-1 I
L! a TABLE 5.1.5-1 TIME SEQUENCE OF EVENTS FOR !NCIDENTS WHICH RESULT IN A - DECREASE IN REACTOR COOLANT SYSTEM FLOWRATE : 4 Loop 3 Loop Accident . Event - Time ' Time -
, (sec) (sec) -
Partialloss of forced reactor coolant flow 1 Four pumps ~ initially Coastdown begins 0.0 '(1) operating, one pump . coasting down Low flow reactor trip 1.5 , Rods begin to drop 2.5 ! Minimum DNBR occurs 3.8 Complete loss of forced 4 (1) reactor coolant flow i-All operating pumps lose power and 0.0 l begin coasting down Reactor coolant pump underspeed_ trip 0.9 setpoint reached Rods begin to drop ~ 1.5 l i i Minimum DNBR occurs 3.6 Reactor coolant pump shaft seizure (locked rotor) (withoul offsite power) Rotor on one pump locks - 0.0 L 0.0 ' s Low flow trip point reached 0.03 0.04 Rods begin to drop .1.03 1.04 '! Remaining pumps begin to coast dowr. 3.03 L 3.041-Maximum RCS. pressure occurs 3.6' 3.4 i Maximum clad temperature- 3.6- 3.8 i (1) Explicit three loop analysis not performed for this event; the four loop results bound those for - ! three loops operating.
,i D0309 5.10 080890 5 64
.I i i I
i l
~!
l
. TABLE 5.1.5 2
SUMMARY
OF RESULTS FOR LOCKED ROTOR TRANSIENTS Four Loops- Three Loops ; l Initially Operating initially Operating , Without Offsite Power , Maximum RCS Pressure (psla) 2652 - 2682 ; 1 I Maximum Clad Ter,perature (* F) '1969 2327 at Core Hot Spot , l Zr H 2O Reaction at Core Hot Spot -0.5 - ' 1'.7 (Percent by weight) f O ! 4 if e e a I - 00309 5:10 080800 5 65
,g:
5.1.6 Reactivity and Power Distribution Anomalies As discussed in Section 15.4 of the FSAR, a number of. faults have been postulated which ' could result in reactivity and power distribution anomalies, included among the possible causes of reactivity changes are RCCA motion or ejection. Possible sources of power distribution changes include RCCA motion, misalignment, or ejection, or by static means such as fuel assembly mislocation. These events are discusseo in this section. 5.1.6.
- Uncontrolled Rod Cluster Control Assembly Bank Withdrawal from a Suberitical or Low Power StartuD Condition 5.1.6.1.1 Introduction As described in FSAR Section 15.4.1, an RCCA withdrawal incident is defined as an ,
uncontrolled addition of reactivity to the reactor core caused by the withdrawal of RCCA , banks resulting in a power excursion. While the occurrence of a transient of this type is highly unlikely, such a transient could be caused by a malfunction of the reactor control or - control rod drive systems. This could occur with the reactor either subcritical, at hot zero: power, or at power. The "a' nower" case is discussed in Section 5.1~.6.2 This event is classified as an ANS Condition 11 event, i.e. a fault of moderate frequency. The general acceptance enteria for this event category are preseneted in FSAR Section 15.0.1.2. The specific acceptance criterion applied is to demonstrate that the DNB design basis is satisfied for this event. As discussed more fully in FSAR Section 1'5.4,1,1, should a continuous RCCA withdrawal accident occur, the automatic features of the reactor protection system available to terrtJnate the transient are as follows:
~
- 1. Source range high neutron flux reactor trip
- 2. Intermediate range idgh neutron reactor flux trip
- 3. Power range high neutret l:ux reactor trip (low setting)
- 4. Power range high neutron flux reactor trip (high setting)
- 5. High nuclear flux rate reactor trip in addition, control rod stops on high intermediate range flux and high power range flux serve to discontinue rod withdrawal and _may prevent the need to actuate the intermediate range flux trip and the power range flux trip, respectively.
1 O, u l 00309 10:10 080990 ~ 5 66 :l
j l J 5.1.6.1.2 Method of Analysis The uncontrolled RCCA bank withdrawal from suberitical accident for the VANTAGE 5H ; transition is analyzed following the same basic methodology discussed la FSAR Section l 15.4.1.2. The analysis is performed in three stages: first, an average core nuclear power i transient calculation, then, an average core heat transfer calculation and, finally, the . !
- DNBR calculation. The computer codes used are just as described in the FSAR and j indicated.in Table 5.1.1.2. .
j For the limiting transient associated with this event, the core axial power distribution is l severely peaked to the bottom of the core.- The W.3 DNBR correlation is used to evaluate DNBR in the span between the lower non mixing vano grid and the first mixing vano grid. ! The WRB 1 and WRB 2 correlations (see SeWon 4.0) remain applicable for the rest of the . fuel assembly for the 17x17 STANDARD and VANTAGE 5H fuel types, respectively, i i The general assumptions indicated in the FSAR for this event continue to apply, though ) the specific analysis parameters used reflect the VANTAGE 5H transition effort. Key l assumptions are as followe: ! I
- 1. A conservatively 'ow value for Doppler power defect ( 0.90% aK) is used, j i
- 2. The analysis employs a moderator coefficient which was at least + 5 pcm/*F at ;
( the zero power nominal temperature, and which becomes less positive for j
\ higher temperatures. This was necessary sinc e the TWINKLE computer code j (see FSAR Section 15.4.1.P.) used in the analysis is a diffusion theory code i rather than a point kinetics approximation and the moderator temperature t feedback cannot be artificially held constant with temperature. ;
- 3. The reactor is assumed to be at hot zero power (557'F). As discussed in the i FSAR, this assumption is more conservative than that of a lower initial system temperature.
- 4. Consistent with the FSAR, reactor trip for this event is initiated by the power j range high neutron flux (Iow setting) function. A safety analysis trip setpoint of 35% is assumed, which constitutes an increase of 10% from the nominal 25% l value. Also, the reactor trip insertion characteristic is based on the assumption
! that the hignest worth RCCA is stuck in lis fully withdrawn position, i i 5. The maximum positive reacitivity insertion rate assumed is greater than that for the simultaneous withdrawal of the combination of the two sequentkl control
- _ banks having the greatest combined worth at maximum speed (45 in.. min.).
- 6. The most limiting axial and radial power shapes, associated with having the two ,
highest combined worth banks in their high worth position, are assumed in the ! DNB analysis. 1
,. 00309 10.10 080890 5 67 i
_ _ _ _ _ _ _ . _ __. ._ __ _ _. - - . _ ~ i 1 I
- 7. The initial power level is assumed to be below the power level expected for any ;
shutdown condition (10 9 of nominal power). The combination of highest ; reactivity insertion rate and low initial power produces the highest peak heat - : flux, ; l l 8. Three reactor coolant pumps are assumed to be in operation, which is ; l conservative with respect to the DNS transient. Also, with the three pump assumption, the single case analyzed is bounding for both four and three trop l j cperation. ! 9. The accident is analyzed using the STOP methodology, just as in the current [ FSAR analysis. . For this event, the use of STOP dictates that the RCS flow rate for three pumps operating ; is based on a fraction of Thermal Design Flow and that the RCS pressure is 53 psia below l nominalin accordance with the initial condition assumptions of Section 5.1.1.3. Since the 4 event is analyzed as being initiated from hot zero power, the steady. state errors on core power and RCS average temperature indicated in Section 5.'.1.3 need not be considered in defining the initial conditions. < 5.1.6.1.3 Results The nuclear power, heat flux, fuel average temperature, and clad temperature as functions of time are shown in Figures 5.1.61 and 5.1.6 2. The peak transient nuclear power < exceeds full power nominal, but only for a very short time. This short duration limits the ; resulting energy release and and associated fuel temperature increase. The thermal flux, I which is a primary factor in determining the potential for DNB. remains well below the full l power nominal value as a result of the inherent thermal lag in the fuel. The analysis results show that the minimum DNBR remains above the limit value (see Table 4 2) at all times during the transient. The calculated time sequence of events is presented in Table 5.1.61. Once the reactor , trip occurs, the plant returns to a stable condition at which time normal plant shutdown procedures can be followed. 5.1.6.1.4 Conclusions i In the event of an RCCA withdrawal from suberitical accident taking place, the core and the RCS are not adversely affected since the DNB design basis continues to be met throughout the transient. Therefore, the conclusions presented in the FSAR for this event i remain valid. G' 00309 10.10 080390 5 68 _ , , . . . _ . _ ,- . . ~ -_ _ _._ _ _ _ - . . - .I
.i 1
5.1.6.2 Uncontrolled Rod Cluster Control Assembly Bank Withdrawal At Power I 5.1.6.2.1 Introduction l 4 i The uncontrolled RCCA bank withdrawal at power event, as described in FSAR Section ! i 15.4.2. results in an increase in core heat flux. Since the heat extraction from the steam , generator lags behind the power generation until the steam generator pressure reaches f the relief or safety valve setpoint, there is a not increase in reactor coolant temperature. Unless terminated by manual or automatic action, the power mismatch and resultant coolant temperature rise could eventually result in DNB. ; i This event is classified as an ANS Condition ll,i.e. a fault of moderate frequency. The l general acceptance criteria for this event category are discussed in FSAR Section { 15.0.1.2. The primary acceptance criterion explicitly considered in the analysis of this l event is to venty that the DNB design basis is satisfied for the transition to VANTAGE 5H fuel, ; t As discussed more fully in the FSAR, the following reactor trips are intended to provide the required protection against an uncontrolled RCCA bank withdrawal at power event: i
- 1. Power range high neutron flux !
- 2. Overtemperature AT
- 3. Overpower AT . ,
- 4. High pressurizer pressure I
- 5. High pressurizer water level ;
i In addition to these reactor trips, there are RCCA withdrawal blocks available on high neutron flux, overpower AT, and overtemperature AT, , l The manner in which the combination of overpower AT and overtemperature AT trips ; provide protection over the full range of RCS conditions is illustrated in Figures 5.1.16 and i 5.1.17 for four and three loop operation, respectively. These figures present allowable reactor coolant loop average temperature and AT for the design power distribution and j flow as a function of primary coola7t pressure. The boundaries of operation defined by ; the overpower AT and overtemperature AT trip functions are represented as " protection ; lin ;.," on this diagram. The protection lines include all adverse instrumentation and a setpoint errors so that under nominal conditions a trip would occur well within the area i bounded by these !ines, i The utility of these diagrams is the fact that the limit imposed by any given DNBR can be - ! i represented as a line. The DNB lines represent the locus of conditions for which the l
\ .. DNBR equals the safety analysis limit value. All points below and to the left of a DNB line 00309-101 0 080990 5 69
c i
'l for a given pressure have a DNBR greater than the limit. These diagrams show that DNB i is precluded for all cases if the area enclosed with the maximum protection lines is not >
l traversed by the applicable DNBR line at any point. , The area of permissable operation (power, pressure, and temperature)is bounded by the combination of reactor trips: high neutron flux, high pressurizer pressure, low pressurizer ; pressure, overpower AT, and overtemperature AT, ; 5.1.6.2.2 Method of Anaissis i
}
The RCCA bank withdrawal at power event is explicitly analyzed for both
- four and three loop operation using the same basic methodology as described in FSAR Section 15.4.2. The purpose of the analysis for this event is to demonstrate that the protection functions described above act in concert to ensure that the DNB design basis is met over a wide range of reactivity insertion rates and at various initial conditions. For .
any individual case, the specific reactivity insertion rate and initial conditions govern which protective function actuates first. The analysis performed for the VANTAGE 5H transition considers the same set of cases as currently defined in the FSAR. However, within this general framework certain specific analysis assumptions in the FSAR are modified to reflect the various revised safety , analysis assumptions of Section 5.1.1.2. The key modifications are as follows:
- 1. For the VANTAGE SH transition, RTDP methodology is employed in analyzing this event, which is a change from the previous STDP methodology currently 9i presented in the FSAR. Consistent with the RTDP aporoach nominal values are assumed for the initial reactor power, pressure, and RCS temperature while i
minimum measured flow is assumed rather than the thermal design flow applicable to STDP. Uncertainties in initial conditions are included in the limit l l DNBR as discussed in Section 4,0 of this report. The specific initial conditions and plant characteristics are discussed in Section 5.1.1.3.
- 2. As in FSAR Section 15.4.2.2, minimum and maximum reactivity feedback conditions are considered. The most positive moderator temperature
! coefficient used in the current analysis is consistent with the FSAR licensing basis analysis. However, the minimum moderator temperature coefficient used ' for VANTAGE SH effort is more limiting than that used previously.
- 3. The primary protection functions actually modeled in the analysis are as in the FSAR, .That is, the overtemperature AT and high neutron flux reactor trips, both ,
with four and three loop setpoints as defined in Section 5,1,1.6. However, the VANTAGE SH transition analysis also considers the high pressurizer water level trip with a setpoint and response time as defined in Section 5.1.1.6 and Table 5.1.14. The current licensing basis analysis did not model the actuation of the high pressurizer water level reactor trip. 0030910.1 D 080990 5 70. ;
i 4 5.1.6.2.3 Results Figures 5.1.6 3 through 5.1.6 5 show the transient response for a rapid RCCA bank j I
! withdrawalincident starting from full power. Reactor trip on high neutron flux occurs !
shortly after the start of the accident. Since this is rapid with respect to the thermal time '! constants of the plant, small changes in Tavg and pressure result and margin to DNB is j maintained. ; i i ! The transient response for a slow RCCA bank withdrawal from full power is shown in l Figures 5.1.6 6 through 5.1.6 8. Reactor trip on overtemperature AT occurs after a longer i period and the rise in temperature and pressure is consequently larger than for rapid ~ i RCCA bank withdrawal Again, the minimum DNBR is greater than the limit value. ; Figure 5.1.6 9 shows the minimum DNBR as a function of reactivity insertion rate from initial full power operation for minimum and maximum reactivity feedback. It can be seen ! that two reactor trip channels provide DNB protection over the whole range of reactivity _ ) insertion rates. These are the high neutron flux and overtemperature AT channels. The , minimum DNBR is always greater than the limit value, I Figures 5.1.610 and 5.1.611 show the minimum DNBR as a function of reactivity insertion ! rate for RCCA bank withdrawal incidents starting at 60 and 10 percent power respectively. [ The results are similar to the 100 percent power case, except as the initial power is l decreased, the rarsge over which the overtemperature AT trip is effective is increased, in O 4 neither case does the DNBR fall below the limit value. Figures 5.1.6 3A through 5.1.610A show similar results for RCCA bank withdrawal at j power events for three loop operation, j
. The shape of the curves of minimum DNBR versus reactivity insertion rate in the $
referenced figures is due both to reactor core and coolant system transient response and to protection system action in initiating a reactor trip. The calculated sequence of events for the uncontrolled RCCA bank withdrawal at power ! l Incident are shown in Table 5.1.61. i 5.1.6.2.4 Conclusions ! The high neutron flux and overtemperature AT trip channels provide adequate protection ; l over the entire range of possible reactivity insertion rates to ensure that the minimum i l value of DNBR is always larger than the limit value (see Table 4 2). In addition. credit for the high pressurizer water level reactor trip ensures that the pressurizer does not go water { Solid. , l O - l l i 0030910 t 0 080890 5 71 :!
- l-
I l 5.1.6.3 Rod Cluster Control Assembly Misalionment 5.1.6.3.1 Introduction FSAR Section 15.4.3.1 contains a general discussion of the RCCA misalignment accidents that identifies causes and provides descriptions of the specific events involved. As presented there, the RCCA misalignment accidents include: l
- 1. One or more dropped RCCAs within the same group.
- 2. A dropped RCCA bank.
. 3. Statically misaligned RCCA.
- 4. Withdrawal of a single RCCA, 1 l
The dropped RCCA, dropped RCCA bank, and statically misaligned RCCA events are classified as ANS Condition Il accidents,i.e , faults of moderate frequency. The single l l RCCA withdrawal is classified as an ANS ' >ndition lli event, i.e., an infrequent fault. The general acceptance criteria for these evo categories are discussed in FSAR Section 15.0.1. For the three Condition ll events, the analysis is performed to verify that the DNBR design basis is met. As explained in FSAR Section 15.4.2.4, since the single RCCA withdrawal i event is a Condition ill accident, the applicable general criteria allow a small fraction of A the fuel to experience damage. For Millstone Unit 3, the current licensing basis analysis ! reports that an upper bound on the number of fuel rods predicted to experience DNB is l 5% of the total fuel rods in the core. The single RCCA withdrawal analysis for the VANTAGE 5H transition uses the same 5% limit as a measure of acceptability for this i event. The transition to VANTAGE 5H fuel, itself, has a:most no impact on the discussion in FSAR Section 15.4.3 with regard to the various RCCA misalignment accidentsc For example, the means of detecting a dropped RCCA bank, misaligned RCCA, or single RCCA withdrawal remain just as described in FSAR Section 15.4.3.1 However, certain of the modified safety analysis assumptions, while not impacting the general methodology, do alter the explicit DNB related calculations for a given case. Among the assumptions ,. falling into this category are the use of the Revised Thermal Design Procedure, the ' I increased RCCA drop time for the VANTAGE 5H fuel, and.the increased F3s power - distribution peaking factor, , There is also one modified assumption that has a more direct impact on RCCA ! misalignment analysis. With regard to the dropped RCCA transient (one or more dropped RCCAs), the VANTAGE SH transition analysis does not take credit for any direct reactor trip or for an automatic power reduction due to the dropped RCCA(s) Specifically, no credit is taken for the negative flux rate trip which, in the current FSAR licensing basis analysis,is assumed to produce a reactor trip for dropped RCCA(s) that produce a l reactivity insertion greater than 400 pcm.. The methodology represented by this approa.ch ; l to the dropped RCCA event is fully documented in Reference 29. i 00309 10.10 080890 5 72
l 6 h l
' 5.1.6.3.2 Method of Analysis l t !
f As in the current licensing basis analyses for the RCCA misalignment events, the cases j considered are bounding for both four and three loop operating conditions. For evaluation l of DNB during the dropped RCCA event (one or more dropped RCCAs), the general analytical methods described in FSAR Section 15.4.3.2 are modified to be consistent with : no credit being taken for the negative flux rate trip. Additionally, the statopoints calculated i for evaluating the DNB design basis are based on plant conditions as defined by RTDP l and a hot channel peaking factor consistent with the increased FAH assumed for the j VANTAGE 5H transition. The transient response, nuclear peaking factor analysis, and . DNB design basis confirmation are performed in accordance with the metnodology of j Reference 29. For Millstone Unit 3, the use of Reference 29 as the basis for the analysis l represents the replacement of the current licensing basis methodology of Reference 30 i which is discussed in the FSAR, , A dropped RCCA bank results in a symmetric power change in the core. As discussed in Reference 29, assumptions made for the dropped RCCA(s) analysis provide a bounding i analysis for the dropped RCCA bank. { For the statically misaligned RCCA, the DNB evaluation includes consideration of steady. state power distributions analyzed using the computer codes and general methods as i described in FSAR Section 15.4.3.2. However, the initial condition assumptions used are ! O nominal power nominal RCS pressure, nominal RCS average temperature, and minimum measured flow as dictated by RTDP. Uncertainties in these initial conditions are lncluded in the limit DNBR, as discussed in Section 4.0. In contrast, the current FSAR analysis for this event uses thermal design flow and is based on power, pressure, and temperatures 1 which directly reflect the uncertainties as required by STDP. ! For the single RCCA withdrawal event, the calculation of power distributions and the j l associated peaking factors within the core, is performed using the methods and computer l codes defined in FSAR Section 15.4.3.2. For the VANTAGE SH transition, the , determination of the minimum DNBR includes consideration of initial conditions that are ! defined consistent with the RTDP methodology. 5.1.6.3.3 Results Single or multiple dropped RCCAs within the same group result in a negative reactivity insertion. The core is not adversely affected during this period, since power is decreasing , rapidly ~ Power may be re established either by reactivity feedback or control bank ! withdrawal. Following a dropped RCCA event initiated from manual rod control, the plant will establish a new equilibrium condition. The equilibrium process without control system interaction is ! monotonic, thus removing power overshoot as a concern and establishing the automatic - O rod control mode of operation as the limiting case.' 1 l 00309-10.10 080890 5 73 !
1 j l i l l For a dropped RCCA event in the automatic tod control mode, the rod control system j detects the drop in power and initiates control bank withdrawal. Power overshoot may occur due to this actio'1 by the automatic rod controller after which the control system will insert the control bank to restore nominal power. In all cases, the minimum DNBR remains above the limit value. , Following plant stabilization normal rod retrieval or shutdown procedures are followed. The operator may manually retrieve the RCCA by following approved operating procedures. A dropped RCCA bank typically results in a negt.har reactivity insertion greater than 400 pcm. The core is not adversely affected during the insertion period, since power is decreasing rapidly. The transient will proceed as described above, however, the return to < power will be less due to the greater worth of an entire bank. The analysis demonstrates , that a reactor trip is not required to mitigate the consequences of the transient. Following plant stabilization, normal rod retrieval or shutdown procedures are followed to further cool down the plant. Consistent with the description for statically misaligned RCCA found in FSAR Section 15.4.3.2, the most limiting misalignment case is considered. As noted above, the analysis l for the VANTAGE 5H transition considers initial conditions based on RTDP. The DNB l design is confirmed to be met for the RCCA misalignment incident, and thus the ability of the primary coolant to remove heat from the fuel rod is not reduced, Following the identification of an RCCA gwup misalignment condition by the operator, the ( operator is required to take action as required by the plant Technical Specifications and i operating instructions. For the single rod withdrawal event, the two cases found in the FSAR have been I considered. Those are single RCCA withdrawal with the reactor in automatic and manual control. For both of these cases, the analysis demonstrates that although a reactor trip on the overtemperature AT trip would actually be expected, that trip does not occur sufficiently fast in all instances to prevent the minimium DNBR in the core from falling
~.ow the safety analysis limit. The evaluation for this event at the power and coolant conditions at which the overtemperature AT trip would be expected to trip the plant shows that an upper limit for the number of rods with a DNBR less than the limit value is 5% 't 5.1.6.3.4 Conclusions For cases of dropped RCCAs or dropped RCCA banks, the DNBR remains greater than the limit value (see Table 4 2). Therefore, the DNB design basis is met and the '
conclusions of the FSAR for this event remain valid. 9 0030910.1 D/080890 5 74
l l, i ) i j For the limiting cases associated with the statically misaligned RCCA event the DNB l remains above the safety analysis limit value (see Table 4 2). Therefore, the DNB design ! basis is met and the conclusions of the FSAR remain valid for this event. ! t For the accidental withdrawal of a single RCCA event, with the reactor in olther automatic [ or manual control mode and initially operating at full power, an upper bound on the i number cf fuel rods experiencing DNBR is 5 percent of the total fuel rods in the core. 5.1.6.4 A Malfunction or Failure of the Flow Controller in a Boillna Water Reactor Looo that Results in an increased Reactor Coolant Flowrate ! This subsection is not applicable to Millstone Unit 3, 5.1.6.5 Inadvertent Loadina and Operation of a Fuel Assembly in an Imorocer Position ! Tne in core instrumentation's ability to detect gross differences between measured and 7 predicted thimble reaction rates is unaffected by fuel type: therefore, the conclusions of l' the FSAR remain valid. 5.1.6.6 Soectrum of Rod Cluster Control Assembly Election Accidents l 5.1.6.6.1 Introduction ! This accident, as discussed in FSAR Section 15.4.8,is defined as the mechanical failure of a control rod mechanism pressure housing resulting in the ejection of a RCCA and . drive shah. The consequence of this mechanical failure is a rapid positive reactivity , insertion together with an adverse core power distribution, pbssibly leading to localized 7 fuel rod damage. j This event is classified as an ANS Condition IV event,-l.e. a limiting fault. The general acceptance criteria for this event category are discussed in FSAR Section 15.0.1.4. The , specific limiting criteria evaluated in the analysis for this event are summarized as follows: 1
- 1. Average fuel pellet enthalpy at hot spot below 225 cal /g for unirradiated fuel and 200 callg for irradiated fuel, j
- 2. Peak roactor coolant pressure less than that which could cause stresses to !
exceed the faulted condition stress limits. ;
- 3. Fuel melting will be limited to less than ten percent 10% of the fuel volume at the hot spot even if the average fuel pellet enthalpy is below the limits of j criterion 1 above. l it should be noted that the current FSAR includes an additional criterion that the average clad temperature at the hot spot must remain below 2700 'F. The elimination of the clad l.
l 00309 10 10 080890 5 75 !
~ '
i
.s i
tempe. Pure criterion as a basis for evaluating RCCA ejection is consistent with the l revised Wesn,5 house acceptance criteria for this event, as defined in Reference 28. i 1 ! The only reactor trip function assumed in the analysis for this event is power range high ! neutron flux, both high snd low setting, e i 5.1.6.6.2 Method of Analysis l The basic methodology used for the analysis of this event is fully consistent with that I described in FEAR Section 15.4.8.2. The calculation of the RCCA ejection transient is f f l performed in two stages, first an average core channel calculation and then a hot region } l calculation. The average core calculation is performed using spatial neutron kinetics , methods to determin3 the average power generation with time including the various total i core feedback effects, i.e., Doppler reactivity and moderator reactivity. Enthalpy and ; temperature transients in the hot spot are then determined by multiplying the averago core ! energy generation by the hot channel factor and performing a fuel rod translent hedt l transfer calculation. The power distribution calculated without feedback is possimistically s assumed to persist throughout the transient. A detailed discussion of the method of f i analysis can be found in Reference 27, [ The average core, hot spot, and system overpressure analyses are performed using the l same methodology and computer codes as described in FSAR Section 15.4.8.2. i Similarly, the general discussion provided in FSAR Section 15.4.8.2.2 continues to be ! appropriate with regard to the calculation of such key analytical parameters as: the > l ejected rod worths, hot channel factors, reactivity weighting factors, moderator coefficient, !
. Doppler coefficient, delayed neutron fraction, and trip reactivity insertion. However, certain specific values indicated in the FSAR text have been modified to reflect the i VANTAGE SH transition.
Table 5.1.6 2 shows the VANTAGE SH transition analysis values for the ejected rod- { worths. hot channel factors, reactivity weighting factors, delayeo neutron fraction, and trip l reactivity insertion while Table 5.1.12 indicates the Doppler coefficient. Just as discussed ! in the FSAR, the critical boron concentration at the beginning of life and end of life aro I adjusted in the nuclear code (i.e. TWINKLE) to obtain moderator density coefficient curves ! which are conservative relative to the actual design conditions for the plant. i The trip reactivity insertion indicated in Table 5.1.12 includes the effect of one stuck RCCA adjacent to the ejected rod. These values are reduced by the ejected rod ' reactivity. The shutdown reactivity was simulated by dropping a rod of the required worth [ into the core. The start of rod motion occurred 0.5 seconds after the high neutron flux trip : point is reached. The actual negative reactivity insertion characteristic as a function of : time following trip is consistent with the 2.7 seconds to dashpot entry, as discussed in ! Section 5.1.1.5. i The set of casos analyzed for the VANTAGE 5H transition, as in the current FSAR, consists of the following: 1 00309 10:10.080890 5 76 . !
i i i !
, i l I
\ 1. Beginning of life, full power for both four and three loop operation
- 2. Beginning of life zero power; as analyzed, a single case bounds both four and ;
three loop operation ! A
- 3. End of life, full power for both four and three loop opvation
- 4. End of life, zero power; as analyzed, a single case bout is both four and three !
loop operation j 5.1.6.6.3 Results f a i Table 5.1.6 2 summarizes the results for all the cases, which covers beginning of life, end- i of life, zero power, and full power as appropriate for both four and three loop operation, j
- 1. Beginning of Cycle, Full Power ;
-)
Control Bank D is assumed to be inserted to its insertion limit. The worst ! ejected rod wcth and hot channel factor are conservatively calculated to be : 0.25 % Ak/k and 6.0 respectively. For the four loop case, the peak hot spot fuel i center line temperature reaches molting, conservatively assumed at 4900 'F. ! However, melting is restricted to less than 10% of the pellet at the hot spot. !
\ The peak hot spot fuel center temperature for three loop operation is 4419 'F, {
thus no fuel melting is predicted. i
- 2. Beginning of Cycle, Zero Power
,t For this condition, Control Bank D'Is assumed to be fully inserted and banks B I and C are at their insertion limits. The worst ejected rod is located in_ Control !
Bank D and has a worth of 0.78 % Ak/k cnd a hot channel factor of 11.5. The fuel center line temperature is 4078 'F. i
- 3. End of Cycle, Full Power ,
Control Bank D is assumed to be inserted at its insertion limit. The ejected rod I worth and hot channel factors are conservatively calculated to be 0.25 % Ak/k ; and 7.0, respectively. For four loop operation, the peak hot spot fuel center line i temperature reaches the conservative melting temperature of 4800 'F, based . upon a high burnup. However, melting is restricted to less than 10% of the pellet at the hot spot. The peak hot spot fuel center line temperature for three i loop operation is limited to 4166 'F, so no fuel melting is predicted. O . 00309 10:10 080890 - 5 77 :
5
?
l . 4. End of Cycle. Zero Power The ejected rod worth and hot channel factor for this case is obtained assuming f Control Bank D to be fully inserted and banks B and C at their insertion limits. l The results are 0.85 %aWk and 26.0 respectively. The peak fuel center line : temperature is 3g31 'F. The Doppler' weighting factor for this case is ! l significantly higher than for the other cases due to the very large transient hot j i channel factor. 'j
. l For all the cases analyzed, average fuel pellet enthalpy at the hot spot remains below 200 cal /g.
The nuclear power and hot spot fuel and clad temperature transients for the worst cases (beginning of life, full power and end of life hot zero power, respectively) are presented in. i Figures 5.1.612 through 5.1.615 for four loop operation. Figures 5.1.612A and 5.1.613A show these same parameters for the worst case (beginning of life, full power) for three . , loop operation. The calculated sequence of events for the worst case rod ejection accidents are presented in Table 5.1.6.1. For all cases, rod insertion occurs after the nuclear power , excursioriis terminated by Doppler feedback. As discussed previously, the reactor . ; remains subcritical following reactor trip. : The ejection of an RCCA constitutes a break in the RCS, located in the reactor pressure vessel head. Following the RCCA ejection, the operator would follow the same 9 l i emergency instructions as for any other LOCA to recover from thenvent. , The installation of the VANTAGE SH fuelin the Millstone Unit 3 core and the results of the ! transition analysis for the RCCA ejection event, as described above, do not invalidate the : discussion of FSAR Section 15.4.8.3 with regard to fission product release, pressure i surge, and lattice deformation. That is,less than 10% of the rods enter DNB, the accident ! does not result in an excessive pressure increase or further RCS damage, and no i mechanism exists for a net positive feedback resulting from lattice deformation. ! 5.1.6.6.4 Conclusions i l The analyses indicate that the described fuel limits are not exceeded it is concluded that i' there is no likelihood of sudden fuel dispersal into the coolant. Since the peak pressure does not exceed that which would cause stresses to exceed the faulted condition stress limits it is concluded that there is no likelihood of further consequence to the RCS. The I analyses have also demonstrated that, for fission product release calculations. the number of fuel rods entering DNB is limited to less than 10% of the fuel rods in the core. O,l
- 00309 10 10 080990 5 78 i
_ . . . _ , ,. __,,_a,_
i Table 5.1.61 l l Time Saouence of Events for incidents Which Result in ' Ret.ctivity and Power Distribution Anomalies N Loop N 1 Loop i Time Time Accident Event (sec) (sec) ,
?
- Uncontrolled RCCA bank initiation of uncontrolled RCCA 0.0 NA -
withdrawal from a withdrawal from 10 9 of nominal subcritical or low power power ; start up condition ! l Power range high neutron flux 6.90 . , low setpcint reached
] ,
Peak nuclear power occurs 7.03
- l Rods begin to fall into core ~7.40 Minimum DNER occurs 8.6 ;
Peak heat flux occurs 8.5 ;
. Peak average clad temperature .9.06 occurs .
Peak average fuel temperature 9.56 occurs 1 Uncontrolled RCCA bank withdrawal at power
- 1. Case A Initiation of uncontrolled RCCA 'O.0 0.0 i- (Maximum feedback) . withdrawal at a high reactivity _ !
Insertion rate (75 pcm/sec) ; i i Power range high neutron flux 5.51- 4.99 . high trip point reached ; Rods begin to fall into core 6.01 5.49' Minimum DNBR occurs 6.20 ~5.80 i i
\
4 00309 10:10'080990 5 79 i
- - . _ , _ _ _ _ . _ - , . - .s. . . . . _ . , , _ . . . . . . - -
_~ _. _ _ _ . . _ . _ _ _ . _ _ _ . . __ _. __ __ _. . _ . L 4 l l
- i Table 5.1.6-1 (cont.)
Time Sequence of Events for incidents Which Resuit in- i i Reactivity and Power Distribution Anomalies i i ! N Loop N 1 Loop - Time . Tim 4 , ; Accident Event (sec) (sec) ' l i
- 2. Case B Initiation of uncontrolled RCCA 0.0 0.0 i (Maximum feedback) withdrawal at a small reactivity l insertion rate (3 pcm/s) +
Overtemperature AT setpoint , 410.6 620.g ! reached Rods begin to fall into core 412.1 622.4 .l Minimum DNBR occurs 412.1 621.1 f RCCA ejection accident l i
- 1. Beginning of life, initiation of rod ejection 0.0 0.0 full power i Power range high neutron flux 0.04 0.04 setpoint reached i Peak nuclear power occurs 0.13 0.13 j Rods begin to fall into core 0.54 0.54 I Peak fuel average temperature 2.06 '. 2.04 I occurs -l Peak clad average temperature . 2,15 2.14 occurs i Peak heat flux occurs 2.16 2.17 !
l > ! t l l e; f 00309 10.10-080990 5 80- ! li
- . , . . . . . ~ . - , .,.-,..-_4- - . - - . . -
i
- i Table 5.1.61 (cont.)
Time Sequence of Events for incidents Which Result in i Reactivity and Power Distribution Anomalies ! N Loop N-1 Loop i ! Time Time l Accident . Event (sec) (sec) i 2. End of life. In'(lation of rod ejection 0.0 NA l zero power l Power range high neutron flux 0.18 ! , low setpoint reached ' !
'{
Peak nuclear power occurs 0.21: Rods begin to fallinto core 0.68 ; Peak clad average temperature 1.38 i occurs l Peak heat flux occurs 1.38 ! Peak fuel average temperature 1.63 occurs ! I
-l l
6 l l i i
?
I i O ; I
. 00309 10 1O/080990 5 81 i
i l l Table 5.1.6 2 Parameters Used in the Anolysis of the Rod Cluster . 1 Control Assembly Election Accident l 1 ! 4 I Egp*:.vus Three Loop Time in Life Beginnine Leginning End End Beginning End l
% 102 0 77 77-Power level (%) 102 t 1
Elected rod 0.25 0.78 .0.25 0 85 0.25 0.25 ] worth (% DK) ; Delayed neutron 0.50 0.50 0.40' O.40 0.50 'O.40 - fraction (%) Feedback reactivity 1,415 2.15 1.55 3.78 1.415 1.55 , welghting Trip reactivity 4.0 2.0 4.0 2.0 4.0 4.0 (% DK) Fq before rod 2.85 - 2.85 - 4.0 4.0 < ejection Fq after rod 6.0 11.5 7.0 26.0 6.0 7.0 ejection Number of RCPs 4 2 4 2- 3 3 running Maximum fuel pellet 4135 3546 3928 3507 3457 3227 + avg. temperature ('F) ; Maximum fuel center 4978 4078 4873 3931 4419 4166 . temperature ('F) I Maximum fuel 182 151 171 149 146 135 stored energy (callgm) 1 Percent of fuel melted <10 0 <10'- 0 0 0' -j O 00309 10.10 080890 5 82
-. . . . - - . ~ . . - - _ . . . - - _ - . - - . - . - - . - _ - . . _ - - __
1 5.1.7 Increase in Reactor Coolant Inventory 1 The events which could result in an increase in the reactor coolant system inventory are ; discussed in this section. 5.1.7.1 Inadvertent Ooeration of the Emeroency Core Coolina System Durina Power Oooration . 5.1.7.1.1 Introduction 1 1 As described in FSAR Section 15.5.1, the spurious ECCS operation at power could be ! caused by operator error or a falso electrical actuation signal. A spurious signal may originate from any of the safety injection (SI) actuation channels. : i An SI signal normally results in a reactor trip followed by a turbine trip. However, it cannot l be assumed that any single fault that actuates the SIS will also produce a reactor trip. If , the reactor protection system does not produce an immediate trip as a result of the spurious SIS, the reactor experiences a negative reactivity excursion due to the injected , boron, causing a decrease in reactor power. The transient is eventually terminated by the reactor protection system low pressurizer pressure trip or by manual trip, i This event is classified as an ANS Condition 11 accident, i.e. a fault of moderate frequency, f The general acceptance criteria for this event category are discussed in FSAR Section 6O 15.0.1.2. With regard to the actual criterion considered in the analysis, it should be noted , that the DNBR is increasing throughout the transient so that the DNBR design basis is not i challenged. The analysis is therefore performed to confirm that there is no hazard to the ' integrity of the RCS. This is accomplished by verifying that the RCS does not go water solid, i.e. that the pressurizer does not fill. 5.1.7.1.2 Method of Analysis The spurious operation of the ECCS is analyzed using the basic methods and computer ! coo 0s described in FSAR Section 15.5.1.2. The current licensing basis FSAR analysis for this event explicitly considers both four and three loop operating conditions. However, using the results of the previous licensing basis analysis performed specifically for Millstone Unit 3, it has been determined that the four loop results bound those with three i loops operating. The basis for this assessment is that the DNBR for the three loop case , remains above that for four loop operation, while the pressurizer volume for the three loop - case is less than that with four loops operating. Therefore, no explicit three loop results ! are reported in this section. . The spurious ECCS event is relatively benign and no plant limits are actually challenged. . , As noted in the FSAR, because of the power and temperature reduction during the transient, operating conditions do not approach the core limits. The results have also t been shown to be relatively independent of time to. trip. ; i 00309 12 10.080890 5 83
?
, - . - -, . - ,.~~ , ,
l I . Certain of the analysis assumptions reporied in the FSAR for this event are affected by the changes made in support of the VANTAGE SH transition. The two modified assumptions ; i are as follows: . i 1. RTDP methodology is employed in analyzing this event, which is a change from j the previous STDP methodology currently presented in the FSAR, Consistent l with the RTDP approach nominal values are assumed for the initial reactor - j power, pressure, and RCS temperature while minimum measured flow is assumed rather than the thermal design flow applicable to STDP. Uncertainties in initial conditions are included in the limit DNBR as discussed in Section 4 of this report. The specific initial conditions and plant characteristics are
! discussed in Section 5.1.1.3. j
- 2. At time zero, two high head charging / safety injection pumps deliver 2900 ppm borated water into the cold leg of each loop. The boron concentration used i reflects the increase included in the analysis for VANTAGE 5H.
]
5.1.7.1.3 Results 1 Figures 5.1.71 through 5.1.7 3 show the transient response to inadvertent operat!ons of I the ECCS during power operation. Neutron flux starts decreasing immediately due to l boron injection, but steam flow does not decrease until later in the transient when the l turbine throttle valve goes wide open. The mismatch between load and nuclear power 1 causes Tavg, pressurizer water level, and pressurizer pressure to drop. When the pressurizer low pressure trip setpoint is reached, the reactor trips and control rods start
]
moving into the core. The DNBR increases throughod the transient. l The calculated sequence of events is shown in Table 5.1.71. After reactor trip, pressure I and temperature slowly rise since the turbine is tripped and the reactor is producing some l power due to delayed neutron fissions and decay heat. I 5.1.7.1.4 Conclusions l i Results of the analysis show that spurious ECCS operation without immediate reactor trip l presents no hazard to the integrity of the RCS. If the reactor does not trip immediately, the low pressurizer pressure reactor trip is I actuated. This trips the turbine and prevents excess cooldown thereby expediting I recovery from the incident. 5.1.7.2 A Number of Bollino Water Reactor Transients This event is not applicable to Millstone Unit 3. t G' D030912:10.080800 5 84 ) l
. - - . . - . - . . - . . . . - - - . - ~ .
4 . J 1 Table 5.1.71 Time Sequence of Events for incidents Which Result- l in an increase in Reactor Coolant inventory l 3 Time ! l Accident Event (sec). ; 3 ' i l Inadvertent ECCS actuation Spurious Si signal generated: 0.0 l l during power operation two centrifugal charging pumps i begin injecting borated water ! i n ' Low pressurizer pressure reactor : 56.1 i trip setpoint reached . l Control rod motion begins 58.1 : O ,
. I t
1 I
' 00309 12:10 080890 5 85 .
l 5.1.8 Decrease in Reactor Coolant Inventory f The inadvertent opening of a pressurizer safety or relief valve, which results in a decrease in RCS inventory, is discussed in this section. l 5.1.8.1 Inadvertent Openino of a Pressurizer Safety or Relief Valve 5.1.8.1.1 Introduction An accidental depressurization of the RCS, as described in FSAR Section 15.6.1, could occur as a result of an inadvertent opening of a pressurizer relief or safety valve. Since a
- pressurizer safety valve is sized to relieve approximately twice the steam flow rate of a relief valve,its opening produces a much more rapid RCS depressurization Therefore, the most severe core conditions resulting from an accidental depressurization of the RCS are associated with an inadvertent opening of a pressurizer safety valve. . initially the event results in rapidly decreasing RCS pressure, and the effect of the pressure decrease is to increase power via the moderator density feedback. The average coolant temperature decreases slowly, but the pressurizer level increases until reactor trip.
This. event is classified as an ANS Condition ll accident, i.e. a fault of moderate frequency. The general acceptance criteria for this event category are discussed in FSAR Section 15.0.1.2. The specific acceptance criterion applied in the analysis of this event is that the DNB design basis be satisfied. The reactor protection system signals that would be expected to trip the reactor for this event are overtemperature DT and pressurizer low pressure. ! i 5.1.8.1.2 Method of Analysis The RCS depressurization event is explicitly analyzed for both four and.three loop operation using the same basic methodology and computer program as described in - FSAR Section 15.6.1.2 However, some of the revised safety analysis assumptions discussed in Section 5.1.1.2 may directly impact the response of the plant to this accident. The key changes in the assumptions of the FSAR analysis are as follows:
- 1. RTDP methodology is employed in analyzing this event, which is a change from the previous STDP methodology currently presented in the FSAR. Consistent with the RTDP approach nominal values are assumed for the initial reactor power, pressure, and RCS temperature while minimum measured flow is assumed rather than the thermal design flow applicable to STDP Uncertainties .
In initial conditions are included in the limit DNBR as discussed in Section 4.0 of this report. The specific initial conditions and plant characteristics are ' discussed in Section 5.1.1.3. 00309 12.10 080890 5 86 i c--,
I L j O A least negative Doppler only power coefficient is assumed, as defined by -
- 2. '
Figure 5.1.1 1. such that the resultant amount of negative feedback is conservatively low to maximize any power increase due to moderator feedback. 5.1.8.1.3 ,R,,g,g d The system responses to the inadvertent opening of a pressurizer safety valve with four loop operation is shown in Figures 5.1.81 and 5.1.8 2. Figure 5.1.81 illustrates the nuclear power transient following the depressurization. Nuclear power increases slowly i until reactor trip occurs on overtemperature AT. The pressure decay transient and average temperature transient following are given in Figure 5.1.8.12. Pressure drops ; more rapidly while core heat generation is reduced via the trip and then slows once ' l Saturation temperature is reached in the hot leg. The DNBR decreases initially, but increases rapidly following the trip as shown in Figure 5.1.81. The DNBR remains above i the limit value (see Table 4 2) throughout the transient. The calculated sequence of events is shown in Table 5.1.81. ; 5.1.8.1.4 Conclusions ; The results of the analysis shows that the low pressurizer pressure and overtemperature AT reactor protection system signals provide adequate protection against the RCS .
, Q depressurization event. l h
l l ?
?
i I l l i l 00309 12.1O'080890 - 5 87 i i
J l i i l
- I l
l Table 5.1.81 l Time Sequence of Events for incidents Which Cause a J Decrease in Reactor Coolant inventory ; i
. N Loop N 1 Loop Time Time Accident Event (sec) (sec) I
)
Inadvertent opening of a Pressurizer safety valve opens 0.0 0.0 pressurizer safety valve fully ; OTAT reactor . 23.5 --- trip setpoint reached Low pressurizer pressure --- 33.0 reactor trip setpoint reached . Rods begin to drop 25.0 35.0 Minimum DNBR occurs 25.6 35.8 < O' f i l 00309 t210 080890 5 88
1 e 1 3 i 5.2 LOCA ACCIDENTS' , 5.2.1 Larae Break LOCA (4 LOOP OPERATION) 5.2.1.1 Description of Analysis / Assumptions for 17X17 VANTAGE 5H Fuel The large break loss-of coolant accident (LOCA) analysis for Millstone Unit No. 3,
- applicable to a full core of VANTAGE 5H fuel assemblies, was performed to develop 1 Millstone Unit No. 3 specific peaking factor limits. This is consistent with the methodology ,
J employed in the Reference Core Report for 17X17 VANTAGE SH, Reference 2. The Westinghouse 1981 Evaluation Model + BASH, References 31 and 32, was utilized and a i spectrum of cold leg breaks were analyzed for Millstone Unit No. 3. Additionally, a single break analysis at the most limiting discharge coefficient (Co) was performed assuming a full core of the previous fuel type (17X17 Standard) anel assuming fuel parameters
; consistent with one cycle of burnup. Fuel data used in the analysis was based on the new fuel thermal model, reference 7. Other pertinent analysis assumptions are provided in Tables 5.2.1 1 and 5.2.12. J l
Conservative assumptions were used in modeling the performance of the ECCS. Both the charging and high head SI pumps were modelled as having 10% less flow than is actually available. Further, the RHR/LHSI pump performance was based on having the RWST l return line valve open. RHR/LHSI performance is severely reduced if the RWST return l
, line is open. Reductions in flow and shut off head occur which would result in degraded I \ ECCS performance. Thus, incorporating this assumption into the LOCA analysis provides l a basis for future plant operational flexibility. I VANTAGE SH fuel features, as applied to Millstone Unit No. 3, result in a fuel assembly . l with a greater hydraulic resistance than that of 17X17 Standard fuel. This difference can '
result in a fuel assembly that is more limiting than the 17X17 Standard fuel assembly with j i l respect to large break LOCA ECCS performance, (Reference 2). As such, VANTAGE SH fuel has been analyzed and a single confirmatory analysis for 17X17 Standard fuel was I l performed. 5.2.1.2 Method of Analysis The methods used in analyzing Millstone Unit No. 3 for VANTAGE 5H fuel, including: computer codes used and assumptions are described in detail below.. The analysis of a large break LOCA transient is divided into three phases: (1) blowdown, (2) refill, and (3) reflood. There are three distinct transients analyzed in each phase, # including the thermal hydraulic transient in the RCS, the pressure _and temperature transient within the containment, and the fuel and cladding temperature transient of the i hottest fuel rod in the core. Based on these considerations, a system of interrelated . codes has been developed for the analysis of LOCA. l LO - o0309-6.10 082790 5 89 . N-.. n - -, , - -~, , - - - - , ,a < ,. . - - - . - , - - - . . ~ - - - - ~
t l l The description of the various aspects of the LOCA analysis methodology is given in References 33,34,35,36, and 37. These documents describe the major phenomena ! modeled, the interfaces among the computer codes, and the features of the codes which ensure compliance with the Acceptance Criteria. The SATAN VI (Reference 38), WREFLOOD (Reference 39), COCO (Reference 40), BART (Reference 36), BASH (Reference 37) and LOCBART (Reference 41,37) codes are used to assess the core heat transfer geometry and to determine if the core remains amenable to cooling throughout and subsequent to the blowdown, refill and reliood phases of the LOCA. These codes j have been previously approved by by the NRC staff for use in LOCA analysis, 1 I SATAN.VI is used to calculate the RCS pressure, enthalpy, density, and the mass and - ( energy flow rates in the RCS, as well as steam generator energy transfer between the primary and secondary systems as a function of time during the blowdown phase of the , LOCA, SATAN VI also calculates the accumulator water mass and internal pressure and i the pipe break mass and energy fl(nv rates that are assumed to be vented to the containment during blowdown. At the end of the blowdown phase, these data are ' (' transferred to the WREFLOOD code. Also at the end of blowdown, the mass and energy. release rates during blowdown are transferred to the COCO code for use in the i determination of the containment pressure response during the first phase of the LOCA. l Additional SATAN.VI output data from the end of blowdown, including the core inlet flow rate and enthalpy, the core pressure, and the core power decay transient, are input to the LOCBART code. l WREFLOOD, using input from the SATAN VI code, calculates the time to bottom of core . l recovery (BOC), RCS conditions at BOC and mass and energy release from the break ; during the reflood phase of the LOCA. Since the mass flow rate to the containment depends upon the core flooding rate and the local core' pressure, which is a function of the Containment backpressure, the WREFLOOD and COCO codes are interactively linked. The BOC conditions calculated by WREFLOOD and the containment pressure transient ' > calculated by COCO are used as input to the BASH code, Data from both the SATAN VI code and the WREFLOOD code out to BOC are input to the LOCBART code which ' calculates core average conditions at BOC for use by the BASH code. BASH provides a more realistic thermal hydraulic simulation of the reactor core and RCS during the reflood phase of a large break LOCA. Instantaneous values of the accumulator- i conditions and safety injection flow at the time of completion of lower plenum refill are provided to BASH by WREFLOOD. Figure 15.6 5 illustrates how BASH has been substituted for WREFLOOD in calculating transient values of core inlet flow, enthalpy, and ' pressure for the detailed fuel rod model, LOCBART. A more detailed description of the BASH code is available in Reference 37. The BASH code provides a much more sophisticated treatment of steam / water flow phenomena in the reactor coolant system , during core reflood. A more dynamic interaction between core thermal hydraulics and ~ system behavior is expected, and experiments have shown this behavior; The BART - r code has been coupled with a loop model to form the BASH code and BART provides the entrainment rato for a given flooding rate. The loop model determines the loop flows and ( pressure drops in response to the calculated core exit flow determined by BART. The .; o0309 61 0'082790 5 90-. e i
, -, -,n. - - - ,- ~ . , . . - - . . . - _- , .--- --. -- . . - , -
)
l ! updated inlet flow is used by BART to calculate a new entrainment rate to be fed back to l l the loop code. This process of transferring data between BART, the loop code and back i to BART forms the calculational process for analyzing the reflood transient. This . l coupling of the BART code with a loop code produces a more dynamic flooding transient, j i which reflects the close coupling between core thermal hydraulics and loop behavior. ; i The cladding hett up transient is calculated by LOCBART which is a combination of the - l LOCTA code with BART. A more detailed description of the LOCBART code can be found ; in Reference 41.37. During reflood, the LOCBART code provides a significant ! improvement in the prediction of fuel rod behavior. In LOCBART the empirical FLECHT ! correlation has been replaced by the BART code. BART employs rigorous mechanistic models to generate heat transfer coefficients appropriate to the actual flow and heat { transfer regimes experienced by the fuel rods. ! l 5.2.1.3 Results i The analyses performed for Millstone Nuclear Power Station Unit 3 are based in part on , sensitivity studies reported in References 44,47 and 48. These studies demonstrated that ; a break in the Reactor Coolant System cold leg piping results in the highest calculated . { Peak Cladding Temperature for Double Ended Guillotine breaks. Therefore, a spectrum i covering a range of discharge coefficients for Double Ended Cold Leg Guillotines (DECL- i G) were analyzed, i The results of this analysis are given in Tables 5.2.13 through 5.2,15 and the transient i results are given in Figures 15.6 8 through 15.6 30. The spectrum for breaks in the cold leg resulted in the Cd = 0.6 being more limiting that the Cd = 0.4 or Cd = 0.8 break cases. ; l Figures 15.6 8 through 15.6 30 present the parameters of general interest for the N loop large break ECCS analyses. The parameters selected are: ! i e Hot Spot Clad Temperature; e Coolant pressure in the reactor core; j e Collapsed Liquid level and Quench front level in the core and the Downcomer level : during reflood; f e Core flooding rate; The following parameters are presented for the limiting break only, e Thermal Power during blowdown; f i e Core flow during blowdown (inlet and outlet);- ; C
* , Core Heat Transfer coefficient; I
D0309 6.10'082790 5 91 l l ~ w e .. -,,w...+.,. ,-.. , , . . . . , . . . - . - - - - , , + . , - - w
e
- Hot Spot fluid temperature; e Mass released to the containment during blowdown; o Energy released to the containment during blowdown: t e Fluid quality in the hot assembly during blowdown; i
e Accumulator water flow rate during blowdown and q e Mass velocity during blowdown; , I e Pumped safety injection water flow rate during reflood. The NRC SER for reference 31 stated three restrictions related to the use of the 1981 EM
+ BASH calculational model. The application of these restrictions to the plant specific large break LOCA analysis was addressed with the following conclusions:
Millstone Unit No. 3 is neither an Upper Head injection (UHI) nor Upper Plenum injection (UPI) plant so restriction 1 does not apply, l The Millstone Unit No. 3 plant specific LOCA analysis analyzed both minimum and maximum ECCS cases to address restriction 2. The Cd= 0.6 Double Ended Cold Leg Guillotine (DECLG) with minimum ECCS flows was found to result in the most limiting - consequenct,s. Restriction 3 required demonstration of the most limiting power shape selection. The [ Westinghouse design process of using a chopped Cosine was used in performing the Millstone Unit 3 Analysis. 5.2.1.4 Conclusions . ; 1 The large break LOCA analysis performed for the Millstone Unit No. 3 has demonstrated that for breaks up to a double ended severance of the reactor Coolant piping, the Emergency Core Cooling System (ECCS) will meet the acceptance criteria of Title 10 CFR Part 50 Section 46. That is:
- 1. The calculated peak cladding temperature will remain beiow the required 2200'F.
- 2. The amount of fuel cladding that reacts chemic;ily we ine water or steam does not exceed 1% of the hypothetical amount that would be generated if all the zirconium metalin the cladding cylinders surrounding the fuel, excluding the cladding; <
surrounding the plenum volume, were to react. , 00309 6;I0 082790 5 92 i
I i O 3. The localized cladding oxidation limit of 17 percent is not exceeded during or after quenching. t
- 4. The core remains amenable to cooling during and after the LOCA.
- 5. The core temperature is reduced and decay heat is removed for an extended period ,
of time. This is required to remove the heat produced by the long lived radioactivity ' remaining in the core. The time sequence of events for all breaks analyzed is shown in Table 5.2.13. The large [ break LOCA analysis for Millstone Unit No. 3 assuming a ful; core of VANTAGE 5H fuel 7 utilizing the 1981 EM + BASH calculational model, resulted in a peak cladding temperature of 1973.5'F for the limiting DECLG break at a total peaking factor of 2.60. The maximum local metal wa.ter reaction was 4.55%, and the total core wide metal water reaction was less than 1.0% for all ca es analyzed. Further, the clad temperature , transients turned around at a time when the core geometry was still amenable to cooling, The effect of the transition core cycles are conservatively evaluated to be at most 50'F higher in calculated peak cladding temperature. The transition penalty is always applied - to the fuel type having the higher hydraulic resistance, in this case VANTAGE 5H. ' Addition of the 50'F transition penalty yields a transition core PCT of 2023.5'F. The transition core penalty can be accommodated by the margin to the 10 CFR 50.46,2200'F limit. The most limiting discharge coefficient found for VANTAGE 5H fuel (Cd = 0.6) was re-analyzed assuming 17x17 Standard fuel parameters. Since VANTAGE 5H and 17x17 l Standard fuel have nearly identical overall hydraulic resistance and the hydraulic transient ' does not model grid effects, the two fuel types will have identical hydraulic transients. Since it is the hydraulic transient that determines the most limiting PCT, the most limiting ; break size (discharge coefficient) will not change between the fuel types. Thus, only the [ l clad heatup transient need be reanalyzed with 17x17 Standard fuel parameters, which only [ l differs from VANTAGE SH in the description of the grids. Thus a complete spectrum , assuming 17x17 Standard fuel was not performed. The 17x17 Standard fuel case with i Cd = 0.6 and using once burnt fuel parameters resulted in a peak cladding temperature of i 2133.2 'F which exceeds the VANTAGE 5H result and requires'that 17x17 Standard fuel [ be considered the limiting fuel type for LOCA consideration until the remaining 17x17 Standard assemblies have obtained sufficient burnup to be bounded by the VANTAGE 5H ( L fuel assemblies or until all 17x17 Standard fuel assemblies have been removed from the , l Core. l That 17x17 Standard fuel was found to result in a higher calculated PCT than [ VANTAGE SH fuel with IFMs differs from the licensing position established in Reference 2. An examination of the results indicates that VANTAGE SH fuel experienced favorable grid : rewetting due to the presence of the IFM grids which rewet first, due to their lower mass,- :i which created conditions for the structural grids to rewet. This effect resulted in'a i D0309 610 082790 5 93
t i ! i l reduction of the enthalpy rise in the hot channel compared to the 17x17 Standard fuel case and produced a lower calculated Peak Cladding Temperature, i t i I i
.I l
1 I 4! i 'l
?
I l
. Oi 4
h L 5 0 00109 6.IO 082790 5 94.
.l I
1 Tabb 5.2.1-1
- j. N-Loop Plant Configuration for Large Break LOCA Analysis; [
t
- Charging Flow Reduced 10%
)
l-
- Safety injection Flow Reduced.'10% . l e RWST Recirculation Valve Open e 10 gpm Safety injection and Charging Pump Flow Imbalance'.
.)
*- 10% Sleam Generator Tube Plugging = .
- Thimble Plugs Removed ,
'i e 17x17 Vantage 5H Fuel with IFMs; .i i
17x17 Standard Fuel, once burnt O 1
?
Y
- . t l
l
- l. i M
i! 1
,'t.'
o . 1 .
-j
- D0309 6:10 082790 - 5e95 2
,4 n
-e w $ *t-s J w, - _ vr-_-a u - e _ m.m._.- Aw- + s m .
-Table 5.2.12 -
N Loop Large Break LOCA Analysis Parameters.
* -3479 MWt' Rated Thermal Power (1.02 x 3411)
- Total Peaking Factor (Fo) - 2.60.
e Hot Channel Enthalpy Rise Factor (FaH)- - 1.70 e - 900 'cul ft. Accumulator Water Volume _(nominal)' e Total Thermal Design F!ow_ 378400 gpm . e Hot Leg Temperature (nominal). 617'F .
- Initial RCS Pressure (nominal)- 2280 psia e Reactor Trip Setpoint' 1600 psia e Safety injection Signal Setpoint 1600 psia e Safety injection Delay Time 40 s:
.i e1 0 0 00309 6:10 082790 5 96 4
_ - . _ _ _ - , - - - .--_._n. _ _ . . _
i Table 5.2.13
' N-Loop Large Break LOCA for VANTAGE 5H Fuel j Time Sequence of Events j I
Co = 0.4 ' Co = 0.6 - Co = 0.8 Co = 0.6 t
~ Min SI ' Min SI LMin SI . Max SI .
1 Start (sec) 0.00 ~ 0.00 - 0.00- ' O.00'
]
Reactor. Trip _ j Signal (sec) 0.934 0.901 0.881 :0.901L SI Signal . . i (Sec) ~ 3.38 2.62 2.23- L 2.62 -. ; Accumulator Injection (sec) '20.40 15.30 '12.70.
- 15.30- l t
End of Bypass - (sec) ' 39.95 32.24- *27.80 -32.24.
- l. t l End of Blowdown (sec) 39.95 :32.24 .27.80 .32.24 'r l
t Pump Injection Begins (sec) 43.38-- 42.62 '38.23- 42.62 7 Bottom of Core l Recovery (sec) 51.92 46.66-- 38.57 46.33 l Accumulator .. Empty (sec) 57.5 55.63 46.39 56.88 , 1 4 1 a 1 5 00309 6 10 082700 5 97 ,
-.-.y.. . . .
q'- i ll
,l Table 5.2.14 Large Break N Loop LOCA Results Fuel Cladding Data l Spectrum of Breaks for VANTAGE 5H Fuel-i DECLG .!
Co = 0.4 Co = 0.6. Co = 0.6 Co = 0.8 .I (Min SI). . (Min SI) (Max SI)1 -(Min SI): . t Results for N Loop 4 I Peak Clad Temp. ('F) 1600.24 -1973.5' 1867.8 ::1775.5. l Peak Clad Location (Ft) 7.0 8.75 7.0 7.0 l. Local Zr/H2O Rxn (Max)(%) 1.15 4.55 3.47 2.52- ; I Local Zr/H 2O Location (Ft) 7.0 8.0 L 7.0 7.0: Total Zr/H2b Rxn (%) < 1.0 < 1.0 < 1.0 - - < 1.0 - Hot Rod Burst Time (Sec) NB1 43.33- ~43.33 46.41 Hot Rod Burst Location (Ft) N/A 6.0/ 6.0 .6,0'- Calculation Rated Thermal Power, Mwt Peak Linear Power, Kw/Ft 3,479 ~
,14.44-j Peaking Factor (at License rating) 2.60 f Accumulator Water Volume, ft3 900/ Accumulator ' 4
-j gy,gjg ' Recion . !
Fuel Region - cycle analyzed - All - All; i 1 1 Burst was not calculated to occur for this case. NB "No Burst". , 1 r l 100309 6:10 082790- 5 98-i
i
-)
a L Table 5.2.15 Large Break N Loop LOCA Results Fuel Cladding Data' i VANTAGE SH Vs Standard Fuel-j DECLG 5 Co = 0.6 Co = 0.6. j (Min SI) . (Min SI) ! VANTAGE SH.- Standard 1 Results for N Looo j Peak Clad Temp. ('F) 1973.5 2133.2 1-Peak Clad Location (Ft) ~ 8.75 8.75 Local Zr/H2O Rxn (Max)(%) 4.55 8.00 j Local Zr/H 2O Location (Ft) 8.0 8.50- l Total Zr/H 2O Rxn (%) . < 1.0 - < 1,0 ! Hot Rod Burst Time (Sec) '43.33 50.34 Hot Rod Burst Location (Ft) 6.0 6.0 ' Calculation ~ O Rated Thermal Power, Mwt 3,479-Peak Linear Power, Kw/Ft 14.44 ' Peaking Factor (at License rating) 2.60 Accumulator Water Volume, ft3 900/ Accumulator gy.gjlg . Realon Fuel Region - cycle analyzed All . All 1 Standard fuel was analyzed based on once b'urnt fuel parameters. Standard fuel will lead the core until sufficient burnup has been obtained to roduce the calculated Peak- ' Cladding Temperature below the VANTAGE SH value 00309 6:10.082790 5 99
- _ n L
l l 5.2.2 Laroe Break LOCA (3 LOOP OPERATION) 5.2.2.1 Description of Analysis / Assumptions for 17 x 17 VANTAGE SH Fuel The large break loss of coolant accident (LOCA) analysis for' Millstone Unit No. 3 for' '! operation with a loop out of service (N 1 or 3 loop operation), applicable to a-full core of VANTAGE SH fuel assemblies, was performed to develop Millstone Unit No. 3 specific i 31oop peaking factor limits. The Westinghouse 1981 Evaluation Model +l BASH,- y References 31 and 32 was utilized with modifications necessary to account for the loop; out of service configuration. Two breaks in an active loop cold leg break were analyzed l .l ' for Millstone Unit No. 3. Fuel data'used in the analysis was based on the new fuel thermal . model, reference 7. Other pertinent analysis assumptions are provided in Table 5.2.21 1 and 5.2.2 2. J Additionally, a single break analysis at the limiting discharge coefficient (Co) was - performed assuming a full core of the previous fuel type (17x17 Standard) and assuming j4 fuel parameters consistent with one cycle of burnup. ; 5.2.2.2 Method of Analysis I The methods used in analyzing Millstone Unit No. 3 for VANTAGE 5H fuel and with a loop l out of service are identical to the models described in detail in section 5.2.-1.2 with regard L ; to computational features. Modifications were required for the modeling features of noding [ and input necessary to account for the loop out of service. These changes are consistent a with past practice for loop out of service analyses where loop isolation valves have been employed (Reference 42). Since application of the newer 1981EM + BASH technology , has not previously been used for loop out of service analyses and specific NRC review; and approval for such an application has not been received additional justification for use i of the 1981EM + -BASH is provided. In determining if a model satisfies the requirements ' j of Appendix K to 10CFR50 the following must be demonstrated.
- 1. Selection of the most limiting active single failure of the ECCS.
, 2. A spectrum of break sizes and location have been considered with the most limiting analyzed. L 3. Selection of the most limiting axial power distribution. , 4 Selection for the most limiting assumption e Offsite power availability. j
- 5. Computer program solution convergence demonstration by studies of system q modeling or noding and calculation time step.
- 6. Sensitivity studies for variations in noding, phenomena assumed in the calculation to predominate and values of parameters over their applicable ranges ,
i 00309 6:1D 082790 5 100 i
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v Since modeling of the PWR Reactor Coolant System (RCS) primary with a loop out of service (N 1) where loop isolation valves are employed does not significantly alter the j i geometry.of the system, the existing noding used.for the case where all the loops (N) are l l active will still apply. This conclusion is based on consideration for modeling of the. vessel ; L upper plenum using a one dimensional assumption and arrangement of the ECCS l injection points with respect to the vessel.in a configuration identical to N loop case. This. 1 L demonstrates that the operation with a loop out of service'does not introduce a new ~ d geometry requiring extensive noding or convergence studies ; Thus, for application of the I 1981EM + BASH Evaluation Model to N 1 situations, noding, convergence criteria, and l' l the treatment of phenomena presumed to predominate do not require modification. l l l Review of the N 1 large break LOCA transient for breaks in an active loop show a great . similarity to the N loop cases. Thus, similar phenomena must be dominating both the ti N and N 1 large break LOCA such that. assumptions for the most limiting single failure? :l most limiting axial power distribution and consideration of offsite power availability . , established for the N loop case apply equally to the N 1 case for breaks in the active loop.- i a Analyses for a break in the inactive loop during N 1 operation were not performed.- White: required by the NRCs SER for N 1 LOCA analyses (Reference 42) this additional break. j location was not analyzed based on the following evaluation which accounts for the ! different effects seen in earlier analysis of inactive loop breaks. The Millstone Nuclear' : Power Station Unit 3 FSAR,- Amendment 18, contains an analysis of large break LOCA for 1 l N 1 (3 loop operation) analyzing a spectrum of active loop breaks and an inactive loop i V break based on the most limiting discharge coefficient determined from the active loo'p . ; break spectrum. These analyses were performed using the 1981.EM (WCAP 9220 P A, 1 Rev.1) and demonstrated that breaks in the inactive loop are less limiting than active loop - breaks. Application of the newer 1981EM + BASH (Reference 31) model to anal'yses of N 1 LOCA' ' introduces newer technology for calculating the reflood flooding rate and heat transfer- ,
- coefficient. while the calculation for blowdown and refill has remained virtually unchanged, ,
l Thus, there would be no change in calculated ECCS performance during blowdown-and ; refill from the previous calculation except for those changes introduced by changes in fuel design. peaking factor assumptions and ECCS flows which would not affect the difference seen between active and inactive loop breaks. ' A review of the analysis of record for the ' i * ,. 1981EM analysis revealed that the inactive loop blowdown calculated a lower cladding - temperature at Bottom of Core Recovery (BOC) of nearly 245'FJ This result demonstrates that analysis of the inactive loop break using the'1981EM + BASH would ' calculate a similar
- benefit out to Bottom of Core Recovery, l
The analysis of the inactive loop break appearing in the Millstone Nuclear Power: Station - Unit 3 FSAR, Amendment, was carried out through reflood using the WREFLOOD computer code (Reference 39). A comparison of the flooding rate history for the active j l and inactive loop breaks shows a higher flooding rate for the inactive loop break and a ; lower cladding temperature rise during reflood of about 45'F compared to the active loop. i
-break. Thus, the break in the inactive loop calculated at 290*F lower Peak Cladding ;
l
.j 00309 6.t0<082790 5101 w 2., . - . .--.- . 4 . . ...__ _ _ , _______________________m_______.'____ _ _ __ _ _ _ _ i_ . _ . _ _ _ _ _ _ _ . _ _ _ _ _
i I l: Temperature (PCT). The higher flooding rate seen for the inactive loop break is due to the - smaller break area associated with the inactive loop break (single sided guillotine).- resulting In a higher RCS pressure to vent the steam. A similar effect has been seen for , intermediate breaks (Reference 44),1.0 ft2, such that this effect is considered realistic. ( While BASH is a more mechanistic code than WREFLOOD for calculating thermal. hydraulic behavior during reflood the mechanism resulting in a higher flooding rate for the; t inactive loop break would also be calculated by BASH such that the inactive loop break. , analyzed with the 1981EM + BASH would calculate a higher flooding rate compared to an . ; active loop break. The higher flooding rates for the inactive loop break would result in a l
'I smaller increase calculated cladding temperatures during reflood when compared to a break in the active loop.
l l On this basis, an analysis of a break in the inactive loop was not considered necessary to. demonstrate compliance with the criteria of 10CFR50.46 and Appendix K to .10CFR50! - 5.2.2.3 Results of Analyses The results of this analysis are given in Table 5.2.2 3 through 5.2.2 5 and the translEnt - ' I results are given in Figures 15.6 44 through 15.6 62. A limited spectrum for breaks'in thel active loop were analyzed which resulted in the Cd = 0.6 being more limiting than the Cd = 0.4 break, The Cd = 0.4 case was analyzed to assure that the longer blowdown- a transients seen for N 1 cases would not result in a short fall for accumulator injection- 'l leading to reduced flooding rates and higher calculated Peak Cladding Temperatures i (PCT). A Cd = 0.8 case was not analyzed based on the N loop results which have shown J a lower PCT for this case. A chopped cosine power shape was used in'the large break -l' N 1 LOCA analysis for the Millstone Unit No. 3. Figures 15.6 44 through 15.6 62 present the parameters of general interest for the N 1 loop break ECCS analyses. The parameters selected are: j
- Hot spot clad temperature;.
- Coolant pressure in the reactor core; 1 .
- Collapsed liquid level and quench front level in the core and the downcomer level ,
during reflood; ! i e Core flooding rate; I q
- Thermal power during blowdown; The following parameters are presented for the Cd =.0.4 DECL G break only. i
- Core flow during blowdown (inlet and outlet);
- Core heat transfer coefficient; D0309 6.10:082790 5-102 .
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- Hot spot fluid temperature;
- Mass released to the containment during blowdown;-
I l e Energy released to the containment during blowdown; j
- Fluid quality in the hot assembly during blowdown; 1
l
- Mass velocity during blowdown; L a l
- Accumulator water flow rate during blowdown and.
i
- Pumped safety injection water flow rate during reflood. :
.1 5.2.2.4 Conclusions l
The N 1 large break LOCA analysis performed for the Millstone Unit No. 3 has _ i demonstrated that for breaks up to a double ended severance of the reactor coolant- , piping in an active loop, the Emergency Core Cooling System (ECCS) will. meet the acceptance criteria of Title 10 CFR Part 50 Section 46. -That is;
- 1. The calculated peak cladding temperature will remain below the require'd 2200'F. :
( 2. The amount of fuel cladding that reacts chemically with the water or steam does not j exceed 1% of the hypothetical amount that would be generated if all the zirconium:
- l metal in the cladding cylinders surrounding the fuel, excluding the cladding . ;
, surrounding the plenum volume, were to react. i . a
- 3. The localized cladding oxidation limit of 17 percent is ndt exceeded dufing or after quenching. :
- 4. The core remains amenable to cooling during and after th's LOCA.
l 5. The core temperature is reduced and decay.h' eat is removed for an. extended period -
- of time. This is required to remove the heat produced by the long lived radioactivity L remaining in the core. .,
I The time sequence of events for all breaks analyzed'Is shown in Table 5.2 2. The N 1. loop large break LOCA analysis for Millstone Unit No. 3 assuming a full core of ,
~ VANTAGE 5H fuel utilizing the 1981 EM + BASH calculational model resulted in a peak-l cladding temperature of 1824.7'F for the limiting DECLG break at a total peaking factor of i 3.00. The maximum local metal water. reaction was 3.01N and ths total core wide metal p -water reaction was less than 1.0% for all cases ana!yzed. The additional;17X17 Standard j V fu'el case with fuel parameters consistent with one . cycle of burnup resulted in a' peak j
.I 00309-6:10:082790 5 103'
.-n-, . ., . . . . , , . . _ , , s. . ,.- . .._:, . , , . . . . . . .
R cladding temperature of 1822,16'F which is bounded by the. VANTAGE 5H result. Further, , the clad temperature transients turned around at a time when the core geometry'was still ,
)
amenable to cooling, ,
. . 1 The effect of the transition core cycles are conservatively evaluated to be at most 50*F .
~
higher in calculated peak cladding temperature.'. The transition penalty is always applied' a to the fuel type having the higher hydraulic resistance, in this case VANTAGE 5H. 'h Addition of the 50'F transition penalty yields a transition core PCT of 1874.7'F, The . , transition core penalty can be accommodated by the margin to the 10 CFR 50.46, 2200* F; ! l i limit. r
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i 1 Table 5.2.21 N-1 Loop Plant Configuration for' Large Break LOCA Analysis !
- Charging Flow Reduced 10% -i
,. ( l e Safety injection Flow Reduced 10% i I r e RWST Recirculation Valve Open . l. e -10 gpm Safety injection and Charging ~ Pump Flow Imbalance - , e 10%' Steam Generator Tube Plugging
- Thimble Plugs Removed ,
e 17x17 Vantage 5H Fuel with IFMs .; 4 O u i V I L. I-l $ 1 t:
'00309-6.1D/082790 5 105 [
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U Table 6.2.2 2-- l N-1 Loop Large Break LOCA Analysis Parameters -
- Power Level 2609 MWt; i}
(.75 x 3411 x 1.02) l d
- Total Peaking Factor _ (Fo) ..3.0 '
- Hot Channel Enthalpy Rise Factor (FAH)- 1.83
- Accumulator Water Volume (nominal) 900 cu, ft< , ..
t e Total Thermal Design Flow - 294900 gpm f i e Hot Leg Temperature (nominal): 609?F j q e- Initial RCS Pressure (nominal); ' P280 pala~' - I
- Reactor Trip Setpoint 1600 psiai e Safety injection Signal Setpoint - 1600 psia : -
e Safety injection Delay Time '40 seconds
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- Table 512 3 - '
N 1 Loop Large Break LOCA for VANTAGE SH Fuel . j Time Sequence of Events l t Co = 0.4 - : Co = 0.6 (Min SI) - (Min SI) Start (Sec) 0.0 0.0 Reactor Trip 0.84 0.815 .3 Signal (Sec) ,
?
SI Signal 4.04- 2.98 (Sec) . Accumulator 18.2 13.9 Injection (Sec) End of Bypass 37.6 29.63 : (Sec) : End of Blowdown 37.7 29.63: ,, (Sec) l Pump Injection 44.04 '42.98' Begins (Sec) "e Bottom of Core 49.50' 40.8 Recovery (Sec) . Accuraulator 72.66 64.'89' Empty (Sec)
~
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1 I Table 5.2.2 4 Large Break 41 Loop LOCA Results Fuel Cladding Data l VANTAGE SH Fuel '! DECLG Active Looo Co = 0.4 ' Co = 0.6 : ,
'(Min SI) 1 (Min. SI) -
Results for N 1 Looo. . Peak Clad Temp. ('F) 1822.2 . 1824.7-Peak Clad Location (Ft) 6.25 ' 6.25: Local Zr/H 2O Rxn (Max) (%) 3.00 3.01 Local Zr/H 2O Location (Ft) 6.25 : 6.25' Total Zr/H 2O Rxn (%) < 1.0 < 1.0 . , Hot Rod Burst Time (Sec) 54.54 50.7 Hot Rod Burst Location (Ft) 6.0 6.0 - Calculation i Rated Thermal Power, Mwt 2,609 O (.75 x 3411 x 1.02) L Peak Linear Power, Kw/Ft '12.47 . L Peaking Factor (at 75% power) 3.00 . Accumulator Water Volume, ft3 900/ Accumulator. Q1qlg Reaion Fuel Region - cycle analyzed All All L l 1. t l I c P O' - D0309-6.1 D<082700 . 5 108:
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.----_.----___________-,b.-- . . - , s , - -, .- . , . - - -
( . Table 5.2.2 5 l Large Break N 1 Loop LOCA Results Fuel Cladding Data . ! VANTAGE 5H Vs Standard Fuel: i i DECLG -,
. Active Loon - l Co = 0.6 - Co = 0.6 ; !
(Min SI)' (Min SI)- ,
. VANTAGE 5H ~ Standard- i Results for N Looo Peak Clad Temp. (* F) . 1824.7' 1822.16 1' ' !
Peak Clad Location (Ft) 6.25 -8.50 ': Local Zr/H2O Rxn (Max)(%) 3.01. 6.37 , Local Zr/H 2O Location (Ft) 6.25 :7.0 i Total Zr/H 2O Rxn (%) . < 1.0 , < 1.0 i Hot Rod Burst Time (Sec) . 50.7 110.8 : Hot Rod Burst Location (Ft) 6.0 . 7.0 ; O Calculation - 5 V l Rated Thermal Power, Mwt 2,609 (.75 x 3411 x 1.02) '! Peak Linear Power. Kw/Ft 12.47 j Peaking Factor (at 75% power) 3.00
- Accumulator Water Volume, ft3 900/ Accumulator -'
I Cycla - Reaion - : Fuel Region - cycle analyzed All All ! a 4 l l 1 Standard fuel was analyzed based on once burnt fuel parameters.- o . l
. 00309 6.10/082790 5-109- .
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3 5.2.3 Small Break LOCA f4 Loco Operation) L l 5.2.3.1 Description of Analysis and Assumptions j The small break analysis was performed with the Westinghouse ECCS Small Breaki q Evaluation Model using the NOTRUMP and LOCTA IV codes (Section 5.2.3.2 of this: - report). The small break (less than one square foot) loss of coolant accident (LOCA) was ! l analyzed assuming a full core of VANTAGE 5H fuel to determine the peak claddingi l l temperature. This is consistent with the methodology employed'in Reference 2 for ,17x17i I VANTAGE 5H transition. The currently approved.NOTRUMP Model Small Break
- Emergency Core Cooling System (ECCS) Evaluation Model was utilized for a spectrum of.
cold leg breaks. Pertinent assumptions include an FNAH of 1.70 for the Millstone Unit No. 3 and a total-peaking factor corresponding to 2.6 at the core mid plane. The analysis is performed at . , 102 percent of the licensed core thermal power level of 3411 MW!.' Other pertinent analysis assumptions are provided in Tables 5.2.31 and 5.2.3 2.~ An auxiliary feedwater flow delivery of 130 gpm per steam generator,was assumed in the analysis. Figure 15.6 7 presents the hot rod power shape utilized to perform the small break analysis. This power d [ shape was chosen because it provides an appropriate distribution of power versus core - height and because local power is maximized in the upper regions of the reactor core q (10 feet to 12 feet), als power shape'is skewed to the top of the core with the peak' local i power occurring at the 10.0 foot core elevation. This is limiting for the small break j analysis because of the core uncovery process for small breaks.: As the core uncovers,' ; the cladding in the upper elevation of the core heats up and is sensitive to the local power at that elevation. The cladding temperatures in the lower elevation of the core, below the i two phase mixture height, remain low. The peak clad temperature occurs above 10 feet. A For these analyses, the safety injection delivery considers accumulator flow, where : , applicable, as well as pumped injection flow which assumes minimum emergency core' cooling cooling system capability and operability. Figure 15.6 43 depicts the pumped' -l safety injection flow as a function of the RCS system pressure. - This figure represents injection flow from a high head safety injection pump, a charging pump, and a low head residual heat removal pump. The head / flow characteristic curve for each pe ip is based on the minimum allowable design flow minus 10 percent of the designt.esd subtracted-
-)
uniformly over over the entire curve. The relationship shown in Figure l15.6 43 includes the pump delivery minus the spillage to containment, with the contWment at atmospheric pressure. Further, it was assumed that one diesel generator failed to start, and, therefore, only one pump in each subsystem is available. A 45 second delay was assumed prior to the initiation of pumped Si flow following reactor trip. This includes the time required for diesel startup and loading of the safety injection pumps'onto the emergency buses. !ln the transients examined, the residual heat removal pump flow is not utilized since its shutoff ' head is lower than the minimum RCS pressure achieved during the relevant portions of .l the transients. 1 01 00309 6.10<082790 5 110 l
1 i I i
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l- The Steam Generator Secondary was assumed to be operating with the Main l Steam; _ .l t-Safety Valves set 3% above the Technical Specification setpoint value and an additional j l 3% accumulation required to fully open the MSSVs. ! l At the initiation of the break, the system was assumed to be operating in steady state.at . l l 102 percent of licensed power, and at 2280 psia,; The analysis was performed.with the ! initial upper head fluid temperature equal to the RCS hot leg fluid temperature, in addition,-
.l' the effect of 10 percent uniform steam generator tube plugging is incorporated into the analysis. A spectrum of 2,3,and 4 inch diameter breaks was examined, with the 3 inch l break proving to be the limiting case, For all cases except the 2 inch break, the reactor . ->
trip was set to 1600 psia. For the 2 inch break, a 1700 psia set point was used.uThe. , results obtained for the larger breaks are conservative, as the use of the higher setpoint would have resulted in earlier initiation of Si flow.
?
Evaluations performed to determine which fuel type is more limiting have determined that l VANTAGE SH with IFMs is more limiting than 17x17. standard fuel in calculated ECCS performance. Since VANTAGE 5H fuel.and 17x17 standard fuel use fuel rods of equal: diameter there will be no difference in stored energy or fuel rod heatup rates due to rod . .j size differences. Therefore, the assembly with the highest hydraulic resistance will result : in the highest calculated PCT. This conclusion is' based on the resistance to steam flow - i during the core uncovery period. For the small break LOCA, the.effect of the fuel-difference is most pronour'ced during core uncovery periods and .therefore, shows up i predominantly in the LOCTA IV calculation in the evaluation model analysis.- On this r i O- basis, only VANTAGE SH fuel was analyzed, since it is the most limiting of the two types of j fuel (17x17 standard and VANTAGE 5H) that would reside in the core. : 5.2.3.2 Method of Analysis ! 1 The NOTRUMP and LOCTA IV computer codes are used to perform the analysis of loss- ' of coolant accidents due to small breaks in the Reactor Coolant System'(RCS). The 1 NOTRUMP computer code, approved for this_use by the Nuclear Regulatory Commission in May 1985 (Reference 23), is used to calculate the transient depressurization of the RCS ! as well as to describe the mass and enthalpy of the flow through the break. This code is j i a state of the art one dimensional general network code incorporating a number of - advanced features. Among these features are the utillzetion of non equilibrium thermal calculations in all fluid volumes, flow regime dependent drift flux calculations with counter- I current flooding limitations, mixture level tracking logic in' multiple stacked fluid nodes and regime dependent heat transfer correlations. The NOTRUMP small break LOCA ! l emergency core cooling system evaluation .model was developed to determine the RCS
- response to design basis small break LOCAs and to address the ,NRC concerns.
expressed in NUREG 0611. " Generic Evaluation of Feedwater Transients and Small Break Ji Loss of Coolant Accidents in Westinghouse-Designed Operating Plants."- In NOTRUMP, the RCS is subdivided into fluid filled contrbi volumes'(fluid nodes) and metal nodes interconnected by.flowpaths and heat transfer links. The transient behavior of the system is determined from the governing conservation equations of mass, energy, and 00309 61 0,082790 5-111 l l
f momentum applied to these nodes. The broken loop is modeled explicitly, and the intact ' loops are lumped into a second loop. A detailed description of the NOTRUMP code is , l provided in References 23 and 25. l l l In the NOTRUMP model, the reactor core is represented as a vertical stack of heated , control volumes with an associated bubble rise model to permit a transient mixture height- ' ! calculation ' The multinode capability of the program enables an' explicit and detailed-spatial representation of various system components.; In particular, it enables a proper . calculation of the behavior of the loop seal during a loss of coolant accident.
'l Clad thermal analyses are performed with the LOCTA IV code (Reference'24) which uses : lI as input the RCS pressure, fuel rod power history,' steam flow past the uncovered part of the core, and mixture height history from the NOTRUMP hydraulic calculations as input.
. For all computations, the NOTRUMP and LOCTA runs were terminated slightly,after the time the core mixture level returned to the top of the core following uncovery.
1 A schematic representation of the computer code interfaces is given in Figure 15.6 6. l 5.2.3.3 Results Based on the results of these studies, the limiting small break was found to be 3 inch diameter rupture of the RCS cold leg, with a peak clad temperature of 1890.5'F. ' In none t of the cases simulated did the clad rupture. In addition, sensitivity studies combined with ;
- the results of the present study indicate little or no uncovering will occur for break sizes 1 that are less than 2 inches, A range of small break analyses is presented which l
establishes the limiting small break'at three inches.- The results'of these analyses are 1 summarized in Tables 5.2.3 3 and 5.2.3-4. Figures 15.0 31 through 15.6 42 present the principal parameters of interest for the small break ECCS analyses. For all cases analyzed, the following transient parameters are - j i presented: , o RCS pressure
- Core mixture height e Hot spot clad temperature For the limiting (3 inch) break analyzed the following additional transient parameters are .
presented: q e Core steam flow rate ' e Core heat transfer coefficient j e Hot spot fluid temperature
- Core power after reactor trip The maximum calculated peak clad temperature for all small breaks analyzed is 1890.5',F. These results are well below all acceptance criteria limits of 10 CFR 50.46.
g T 00309 6:10 082790 5 112 h i
~ _ _ . _ _ . _ . . _ ._ __ _ .
5.2.3.4 Conclusions The srnall break VANTAGE 5H LOCA analysis,~ utilizing the currently approved NOTRUMP . 3 Evaluation Model, resulted in a peak cladding temperature (PCT) of 1890.5'F for the: } 3 inch diameter cold leg break. The analysis assumed a limiting small break power _- i shape consistent with a LOCA Fo(z) envelope of 2.60 at the core midplane elevation linearly decreasing to 2.40 at the top of the core. The maximum local metal water.
. 4 reaction is 4.56 percent, and the total core metal water reaction is less than 1.0 percent i for all cases analyzed. The clad temperature transients turn around at a time when the core geometry is still amenable to cooling. An analyses for 17x17 Standard fuel was not performed to determine which fuel type could be more limiting since the current small . I break methodology does not model the fuel rod grids. VANTAGE 5H fuel and 17x17 : I Standard fuel are identical in both design and response to LOCA transients, in the 1 e absence of grid effects.
The analyses described in the preceding sections demonstrate that one centrifugal pump and one high head safety injection pump, together with the accumulators, provide . . j sufficient core flooding to keep the calculated peak clad temperature well below the required limits of 10 CFR 50.46 (for all units). Adequate protection is therefore afforded by-the ECCS in the event of a small break LOCA. 5.2.4 Small Break LOCA f3 Looo Ooeration) The current Westinghouse methodology 'or calculating' ECCS performance for a small , break LOCA with a loop out of service is described in Reference 42.: This methodology was based on the older W FLASH computer code which has since been replaced by : NOTRUMP Reference 42 did not perform analyses for a loop out service but rather -l presented arguments for not analyzing the small break LOCA.: The arguments c' entered ' on the margin in calculated Peak Cladding Temperatures between the more limiting N l loop large break LOCA analyses and the N loop small break analyses and the increase in . the ratio of pumped Si to core power for operation in N 1 loop modes These arguments were used to state why N 1 small break LOCA would never be more limiting than the N and N 1 large break LOCA results.- The current analyses for Millstone Nuclear Power Station Unit 3 did demonstrate j significant margin between the N loop large and small break LOCA results, in fact, the-small break PCT (1890'F) is'less limiting than the large break PCT (1973*F). The. . . . l argument that the Si to power. ratio increases when a loop is isolated for operation in the N 1 loop mode applies and NOTRUMP would calculate improved ECCS performance'for- i a small break LOCA for N 1 loop operation. Given that operation with a loop out of service q would improve the Si to power ratio and that small break is not as limiting as large break, i small break LOCA analyses for N 1 loop operation were deemed unnecessary in ! demonstrating compliance with the criteria of 10CFR50.46. . O
'i ;
00309-6.10/082790 5 113 . l
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5.2.5 Blowdown Reactor Vessel and l'ooo Forces The forces created by a hypothesized break in the reactor coolant system (RCS) piping - ' i are caused principally by the motion of the decompression wave throughout the RCS,- The strength of the decompression wave is a function of the assumed break opening time, . the pipe break area and relative location of the break, and the RCS steady state operating l i conditions of power, temperature and' pressure. 'These parameters are not significantly l affected by a change tc VANTAGE 5H fuel at Millstone Unit 3. The magnitude of the j LOCA hydraulic forces in the fuel region is a function of the core flow area / volume and the' f associated loss coefficients. ' Previous sensitivity studies performed for these changes ! .] have demonstrated that there will be no adverse change in the forces calculated for a- i hypothesized LOCA as a result of this fuel change, tThose LOCA hydraulic forces acting- . upon the RCS loop piping are also not significantly influenced by changes in fuel-~ } assembly design. A review of the LOCA hydraulic forcing functions representative of the-Millstone Unit 3 N. loop plant design, concludes that the use of limited displacement inlet , and outlet reactor vessel nozzle breaks is acceptable for the evaluation of the - l VANTAGE 5H fuel. The use of these breaks, yields bounding forcing functions for the N- 1 loop configuration in consideration of the margin made available as a result of Leak. . Before Break methodologies. ~With respect to the N 1 loop configuration at Millstone _ Unit!
] t 3, branch line breaks of the accumulator line and pressurizer surge line are appropriate .
for the evaluation of the VANTAGE 5H fuel. The use of these breaks generated specifically for the Millstone Unit 3 plant, is appropriate in consideration of the effect of-
}
l l those small plant changes associated with the VANTAGE 5H fuel reload and the; i application of Leak Before Break methodology. o 1 i i i
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, D0309 6:10.08279o 0*lI4
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1 l-I O - TABLE 5.2.31 ' N LOOP PLANT CONFIGURATION FOR SMALL BREAK LOCA ANALYSIS [ ( e Charging Flow Reduced 10% Jr e Safety injection Flow Reduced 10% - 1 e i e RWST Recirculation Valve Open for RHR/LHSI performance
- 10 gpm Safety lnjection and Charging Pump Flow Imbalance -
- 10% Steam Generator Main Steam Safety Valves l '
Setpoint pressure increased by 3% over the Technical. Specification c
= value plus an additional 3% for accumulation - i
- Thimble Plugs Removed A 4
e 17X17 VANTAGE.5H Fuel with IFMs o ; t I i l l *
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{ -, I. TABLE 5.2.3 l N LOOP SMALL BREAK LOCA ANALYSIS PARAMETERS < 1 L - t .. . .
- _ e Reactor Power Level 3479 MWt; L
j (3411 Mwt X 1.02) ,i 1 e Total Peaking Factor (Fo) 2.6t - I l e Hot Channel Enthalpy Rise Factor (FaH)l :1.70 l l e Accumulator Water Volume 900 Cu. Ft.
- Total Thermal Design Flow - 1378400 gpm
- Hot Leg Temperature N
~ 617' F .'
1 e initial RCS Pressure- - 2280 psia i e Reactor Trip Setpoint 1700 psia .! e Safety injection Signal Setpoint'- L1600 psla : , L e Safety injection Delay Time :45 Seconds
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l TABLE 5.2.3 3 1 TIME SEQUENCE OF EVENTS FOR INCIDENTS WHICH CAUSE 1 A DECREASE IN REACTOR COOLANT INVENTORY - _l
)
N Loop Time Accident Event ,
. (sec)_ ;
a Small Break LOCA f ( 1. 2 in Start 0.0 - l Reactor tr'ip signal 151.3 l - l L Top of core uncovered 1987 Accumulator injection begins , l Peak clad temperature occurs 2742 Top of core covered 3680 , l
- 2. 3 in Start 0.0 -
Reactor trip signal - - 76.7 - :. l Top of core uncovered ~ '894 ! i Accumulator injection begins 1738 i Peak clad temperature occurs 11541' q Top of core covered- 2607
- 3. 4 in - Start 0.0 - q Reactor trip signal 39.4 [
Top of core uncovered 554 -i Accumulator injection begins - 833 d Peak clad temperature occurs 896' ' Top of core covered 1154-i
- Note: No accumulator injection for 2 in break. ;i O 1
- i 00309 6:1D/082700 5 117
. -- , . a. .. w .. . -
1
.i TABLE 5.2.3-4 SMALL BREAK RESULTS '
?
'{
2-inch - 3-inch 4-inch i
- (0.0218 fta) ,
(0.0491 ft2)~ l(0.0873 ft2) Peak Clad Temperature, deg F ' 1117.0.l 1890.5 1622.2 Peak Clad Location, ft ' '11.50 ~ 12.00- .11.50 '
~[ ,
Local Zr/H 2O Reaction (MAX), %
~
0.077 4.558 -0.720. ] Local Zr/H 2O Reaction, ft 11;50 12.90 L 11.50 j Total Zr/H 2O Reaction, % - < 1.0 . < 1.0 < 1.0 i Hot Rod Burst Time, sec -N/A N/A' .N/A-~ ; Hot Rod Burst Location, ft N/A N/A N/AL ! 1:
- l. ]
(. 'i l
?
=
s i r Gl f I 00309 6:10 082790 :5 118. j
- -
- w- w. , . - (
~
l l l l l O 5.3 MASS AND ENERGY RELEASE AND CONTAINMENT INTEGRITY 1. Q 5.3.1 LOCA Mass and Enerov Release j l l' 5.3.1.1 Introduction =l
\
The purpose of the evaluation described herein is to determine the impact of Vantage SH fuel and the increase in peaking factors on the current LOCA mass and energy releases. L 1 Also, the impact of other items including RTOP,. revised non RTDP uncertainties 10% l l steam generator tube plugging, RAOC,' reactivity coefficient changes, reduced . shutdown ! margin, thimble plug removal, increased safety valve drift, deletion of the negative flux rate trip, increasing the delay time for diesel start. degraded safety injection flow, expanding . the accumulator level band, and revising the pressurizer level prog'am were considered; in the investigation. -' y 5,3.1.2 Method of Analysis 5.3.1.2.1 Current FSAR' Analysis The current licensing basis evaluation model for the long term LOCA mass and energy release calculations is the March 1979 model described in Reference 17. This evaluation ~ ! model has been reviewed and approved by the NRC, and has been used in the analysis-of many other ple.nts. j For the current licensing basis long term mass and energy release calculations, j l described in Reference 16, operating temperatures for the highest average coolant- i l temperature case were selected as the bounding analysis conditions.~ The use of higher: temperatures is conservative because the initial fluid energy is based on coolant temperatures which are at the maximum levels attained in steady state operation. Additionally, an allowance is reflected in the temperatures to account for instrument error-and deadband. The plant power level was adjusted for. calorimetric error (+ 2 percent of g power) in the analysis. The initial RCS pressure in the analysis is based on a nominal ' value and includes an presaure allowance to account for.the uncertainty on pressurizer pressure. The following items ensure that the mass and' energy releases are conserutNely calculated in order to determine the' maximum containment pressure: '
- Maxin:um expected operating temperature of the reactor coolant system- j
- Allowance !n temperature for instrument error and dead band
- Margin include:i in RCS volume (for thermal expansion and uncertainty).
* ' Maximum power evel 00309 7.IO 080990 5119 l
.5
t .l 1 i,
- Allowance for calorimetric error e1 :
}
- Conservative coefficients of heat transfer (i.e., steam generator ~(
primary / secondary heat transfer and reactor coolant system metal heat J transfer) . .;
.i
- Allowance in core stored energy effect of fuel densification ,
i e Margin included in core stored energy - J
- Allowance for RCS pressure uncertainty _ 'l El The sources of energy considered in the LOCA mass and energy release analysis - '
- include:
- Reactor Coolant System Water
- Accumulator Water-
- Pumped injection Water -i I-
- Decay Heat i
- Core Stored Energy 1 l
- Reactor coolant System Metal I l
- I
- Steam Generator Mr :
- Steam Generator Secondary Energy
- Secondary Transfer of Energy (feedwater into and steam out of the steam -
generator secondary) ,; 1 Consen>a""a analysis variables were also used in determining the LOCA shorI term mass u l ande "ses. The methodology described in Reference 18 was utilized to. f gener ; term releases utilized in the subcompartment analyses for the-posti 4 accidents. l
- 5. 3.1.. oe 5H Fuel Evaluation Current analyses and evaluations and other relevant information were utilized to evaluate -
the impact of Vantage SH fuel and the increase in peaking factors. Also,-the impact of
~ RTDP, revised non RTDP uncertainties,10% steam generator tube plugging. RAOC, .
j
. reactivity coefficient changes,' reduced shutdown margin, thimble plug removal, increased safety valve drift, deletion of the negative flux rate trip, expanding the accumlator level i
Em e?!mM$E@ . 54120;_ , _
- . _ - - - - - _ _ ~ . ______
, t
'l
./ band and revising the pressurizer level progrannvere evaluat,ed. Increasing the delay! l
\ time for diesel start and degraded safety injection flow would require reanalysis to quantify._
Therefore, these items were considered _oulof s_ cope for this evaluatiort 5.3.1.3 Results The Vantage 5H fuel to be used in Millstone Unit 3 has the same fuel rod diameter and pellet type as used in the current 17 x 17 standard fuel. The core stored energy is therefore unaffected due to the change in fuel type. The effect of the IFM grids is to- . increase the core bypass flow, which increases. core Tavg, but the reactor vessel Tavg i and primary loop temperatures remain the same with the change in fuel type. For LOCAL mass and energy releases it is conservative to assume 0% steam generator tube plugging, which maximizes the delivery of secondary side heat to the primary system.' ! Therefore, the current analysis bounds the 10% plugging level increase in peaking-factors, effect of RTDP, revised non.RTOP uncertainties, RAOC, reactivity coefficient changes, reduced shutdown margin, thimble plug removal. Increased safety valve drift . 4 deletion of the negative flux rate trip deletion, expanding the accumulator level band and revising the pressurizer level program have no impact on the long term LOCA mass and _- ! energy releases. For the LOCA short term mass and energy re' leases used in the subcompartment' pressure analysis, there is no impact of any of the previously mentioned items, including _ l P the change in fuel type, because the transient is typically only concerned with the first 3 L' seconds of the transient. i 5.3.1.4 Conclusions - l There is no impact on LOCA short and long term mass and energy analysis due to the replacement of the 17 x 17 standard fuel with the Vantage 5H fuel for Millstone Unit 3. Furthermore, the results of the four loop LOCA mass and energy release evaluation bound the three loop results. f l Increased delay time for diesel start and degraded safety injection.would require -; l reanalysis to quantify the impact on the long term releases _for a postulated LOCA. j Therefore, these items are considered out of scope of this evaluation. , 5.3.2 Steamline Break Mass and Enerav Release 1 1 The steam generator mass and energy release dat'a inside containment are generated to l determine the containment pressure and temperature response following a steamline -; break event. As discussed in Section 6.2.1.4 of the FSAR, the mass and' energy' releases, j are generated for a variety of power levels, break sizes, and single failure essumptions; { The mass and energy releases presented in the'_FSAR are based on analyses performed , generically for plants with model F steam generators. As_such, the analyses are-conservative with respect to the actual Millstone Unit 3 design, including the changes. [ [\ discussed in Section 5.0 for the Vantage 5H fuel upgrade.. Therefore,'it is concluded that 00309 7:10 080990 5121- ) i w , ., < , . +
the current steamline break mass and energy releases inside containment for Millstone Unit 3 remain valid for the Vantage 5H fuel upgrade. The steam generator mass and energy release data outside containment are generated to. j ensure that equipment environmental qualification limits are met following a steamline break. As was the case for the mass and energy releases inside containment, the outside containment analyses were performed generically such that the analyses bound a number ' fe of four loop Westinghouse plants. These analyses are conservative with respect to the j actual Millstone Unit 3 design, including the changes discussed in Section 5.0 for the. ; Vantage SH fuel upgrade. Therefore, it is concluded that the current steamline break mass and energy releases outside containment for Millstone Unit 3 remain valid for the > Vantage 5H fuel upgrade. 5.3.3 Containment Inteority Containment Integrity Analyses are performed during nuclear plant design to ensure that the pressure inside containment will remain below the containment building design. '! pressure if Loss of Coolant (LOCA) or a rupture of a main steamline pipe (MSLB)inside containment should occur during plant operation. The analysis ensures that the containment heat removal capability is sufficient to remove the maximum possible discharge of mass and energy to containment from the Nuclear Steam Supply System without exceeding design pressure. A spectrum of accidents and single failures combined with simultaneous occurrences, such as loss of offsite power, are considered. The subcompartments are designed in accordance with General Design Criteria 4 and 50. Short Term Subcompartment Analyses are perioced to ensure that the walls of a subcompartment can maintain their structural integrity dring a short term pressure pulse which accompanies a high energy line pipe rupture within the subcompartment. Structural , adequacy due to pressurization loadings in the reactor vessel region, the pressurizer - ! cubicle, and the steam generator compartment is verified. A double ended rupture of the spray line and a double ended rupture of the surge line are considered in the upper pressurizer cubicle and the lower pressurizer cubicle, respectively. Section 6.2.1.2 of Reference 1 describes the containment Subcompartment Analyses, including the design basis breaks. l There are two major aspects in performing containment integrity and subcompartment-analyses: (1) generating the mass and energy releases and (2) determining the containment and subcompartment environmental response to the releases. Westinghouse does not have containment integrity scope or subcompartment scope for Millstone 3, however, as discussed above, the mass and energy release from LOCA and steamline break which are inputs to the containment and subcompartment analysis were not adversely affected by the upgrade to VANTAGE 5H and the additional changes. , I 00309 7:10 080990 5 122
. l, i I 1 5.4 RADIOLOGICAL CONSEQUENCES OF ACCIDENTS
- 5.4.1 Radioloojeal Consecuences of Extended Discharoe Burnuo
! This section summarizes the impact of the upgrade to VANTAGE 5H fuel for Millstone Unit i 3 on the radiologl cal consequences of accidents. The change in the fuel has the potential ,
- for affecting the accident closes only to the extent radiological source terms are ;
i increased, included in this evaluation is consideration of extended fuel burnup of 45,n0 ( l MWD /MTU for the batch average discharge. This extended burnup has a peak fuel rod j average burnup of < 60,000 MWD /MTU. q The extension of fuel burnup has been shown to have negligible impact on the core - l inventory of radioactive isotopes which are of concern in evaluating the radiological i consequences of accidents (i.e., the short half life noble gases and lodines). This has- ! been documented by Westinghouse in a topical report evaluating the impact of extended burnup en Westinghouse fuel (Reference 8). Independent review performed for the NRC ! (Reference 20) reached the same conclusion. These reviews considered burnups of up to 60.000 MWD /MTU for the lead rod which bounds the fuel design being provided for Millstone Unit 3. i l Also of concern in evaluating the impact of extending fuel burnup on the radiological - consequences of accidents is the fraction of core activity that is assumed to migrate out of the fuel matrix and into the fuel clad gap region and thus available for release in the O event the cladding is breached. The accident analyses for Millstone 3 have used the gap fractions recommended by Regulatory Guides 1.25 and 1.77. As indicated in Reference 8, i these gap fractions remain extremely conservative for extonded burnup fuel. Reference 20 agrees tha' the extended burnup fuel gap fractions for inost isotopes of ccacern do not exceed the gap fractions specified in the Regclatory Guldts. However, Reference 20 claims that for the Fuel Handling Accident (FHA) the 1131 gap fraction could be as high as 12% (an increase of 20% over the Regulatory Guide 1.25 value). This increase is inconsistent with Westinghouse topical report and with specific plant analyses which have conservatively modeled the anticipated fuel management and which have determined a maximum gap fraction for 1131 of just over one percent which is consistent with the
- findings in Reference 8. Although these analyses were not specific to the Millstone plant, they were determined to be applicable. Other documents such as NUREG 1228 (October 1988) place an up9er bound on lodine gap activity of 2% Hence, the Regulatory Guide assumptions remain valid.
As discussed above, there would be no expected increase in the isotopic inventories or in the activity gap fractions assumed in radiological calculations for various accidents. Thus the doses reported in the FSAR remain bounding, , 5.4.2 Radioloalcal Conseauences of Fuel and Plant Chanaes i Tho previous section concluded that extended fuel burnup results in no adverse impact on , offsite doses resulting from the accidents evaluated in Chapter 15 of the FSAR. The 00309 8 10.080990 5 123
E change in fuel design to VANTAGE SH as well as the other changes discussed in Section j 5.0 of this document have also been reviewed for impact on the various accident analyses i' reported in Chapter 15 of the FSAR. Of the events reviewed for this document. an evaluation of the impact on calculated doses was performed for those which have a l radiological consequence analysis reported in the FSAR. These events include LOCA, steam line break, locked rotor and rod ejection. The radiological consequences of a
- LOCA are independent of any change in the analyzed core response. Therefore, although the changes discussed above may affect the core response to a LOCA. there will be no
- subsequent impact on the calculated offsite doses.
1 The effect of the design changes on mass and energy releases resulting from a steam line break was evaluated and is reported in Section 5.3.2. It was concluded that the mass and energy releases for the steam line break contained in the FSAR remain valid. The 1
' results of the locked rotor analysis are reported in Section 5.1.5.3. The analysis determined that the percent of rods in DNB is bounded by the assumptions for fuel failure ! currently used for the accident dose analysis (see Reload Safety Evaluation, Millstone Unit l
3 Cycle 3. Revision 1, May 1989). Finally, for rod ejection, it was determined that the mass and energy release calculat.d in the FSAR remains valid. Therefore, the radiological consequences of the steam line break, locked rotor and rod ejection events are not increased due to the change to VANTAGE SH and other changes discussed above. O. f l 9. i 00309 8:1o 080990 5;124
l l l 5.5 NORMA'. OPERATION CONTROL S STEM OPERATION AND MARGIN TO TRIP
- 5.5.1 Introduction j
Some of the more important parameters in defining the NSSS response to a temperature , l or power transient are the core reactivity coefficients. With the new VANTAGE SH core design there is a potential for changes in these coefficients which may change the response of the NSSS and the control systems to a less stable configuration. In addition, degraded NSSS response to normal operating transients may reduce the available margin to various reactor trips beyond acceptable limits, if either of these conditions occur, control system setpoints can be modified to restore proper NSSS response. Control i systems do not perform any protective function in the FSAR accident analysis. Cc,ntrol system operation is assumed only in cases where their action aggravates the - 4 consequences of an event, or as required to establish initial plant conditions for an analysis. With a few exceptions, the control system setpoints can be modified without affecting the accident analysis, i I An evaluation of the best estimate VANTAGE 5H core reactivity parameters and the subsequent NSSS response, shows that there should be little change from the previous core design. Control system operation should remain stable, and except for the Overtemperature Delta T trip, adequate margin to all reactor trips and ESF actuations l should remain. As always, NUSCO has the option of modifying to the control system l O setpoints. If desired, system performance can be verified by executing a series of plant tests (e.g., step load changes, load rejections, etc.) following installation of the VANTAGE 5H core. Adjustments to the control system setpoints can then be made based j
)
on the results of these tests.
]
l 5.5.2 Method of Analysis The control system evaluations were conducted using a model of the Millstone 3 NSSS with the LOFTRAN computer code. Best estimate values for initial conditions and core l parameters were used to evaluate plant response to those normal condition transients that best demonstrate control system stability or that cause the largest reduction in operating margin. For Millstone, three transients were chosen for this evaluation, The first is a 10% step load increase from 90% power. This transient demonstrates rod control system ' stability at high power and also generates the largest challenge to the high flux reactor trip-and low steamline pressure SI actuation. The second transient is a 10% step load increase from 15% power. This transient demonstrates rod control system stability at low power. The third transient evaluated is the 50% load rejection from 100% power. This transient demonstrates the stability of the steam dump system and also generates 't largest challenge to several of the reactor trips of ESF actuation. i
\
(U 00309 8.10 080990 5 125
'l
l l 1 5.5.3 Results 5.5.3.1 Rod Control System Stability The stability cu the rod control system was examined for N loop operation at 3425 MWt with a full power Tavg of 587.1'F. (Previous studies have shown that the calculated 1 response of the rod control system for N 1 loop operation is similar to the response for N loop operation.) The basis of the evaluation was the plant response to a 10% step load . increase transient from 90% and 15% power for moderator temperature coefficients l (MTC) ranging from + 10 pcm/*F to 40 pcm/'F and rod worths ranging from 1 pcm/ step l to 30 pcmtstep. Actual MTC values should not be greater than about + 5 pcm/'F based on preliminary Cycle 4 information. The results of the evaluation showed adequate stability in the rod control system with the setpoints as specified in revision 2A of the PLS document. It should be noted that the occurrence of low frequency rod stepping may occur during operation with a positive or slightly negative MTC. This is a strong function of the rod control system deadbands and the reactor core kinetics. In order to minimize the possibility of excessive CFIDM wear, either minimize the time spent operating below 50% power or operate with the rod control system in manual early in the fuel cycle when positive MTC conditions are most likely to exist. This precaution should be maintained until the predicted MTC at 15% power is more negative than about 3 pcm/'F. i 5.5.3.2 Steam Dumo Load Relection Controller Stability e The stability of the load rejection steam dump controller was examined for N loop operation at 3425 WMt with a full power Tavg of 587.1'F. (Previous studies have shown that the calculated response of the steam dump system for N 1 loop operation is similar to the response for N loop operation.) The basis for the evaluation is the plant response to a 50% load rejections from 100% with 40% steam dump capacity for moderator temperature coefficients (MTC) ranging from + 10 pcm/'F to 40 pcm/'F and rod worths l ranging from 1 pcm/ step to 30 pcm/ step. Actual MTC values should not be greater than 0 pcm/*F based on Technical Specification requirements and will be closer to 5 pcm/'F bases on the preliminary cycle 4 information. The results of the evaluation showed adequate stability in the steam dump control system With the setpoints as specified in revision 2A of the PLS document. Some instabilities or steam dump / rod control interactions may occur for combinations of large rod worths and ) high (positive or nearly positive) moderator coefficients. The onset of these conditions is I somewhat dependent on the steam dump system setpoints but their occurrence is not expected based on the VANTAGE SH operating region. l 5.5.3.3 Marain To Trio Evaluation , This section ;: c@ies a discussion of the operating margin to the various reactor trips during steady state and normal operating transients with the VANTAGE SH fuel. There are D0309 81D 080990 5126
l ! l l ' 11 automatic reactor trip and 4 ESF actuation systems active during normal operation. . The reactor trips are High Flux, High Flux Rate, Overtemperature AT, Overpower AT, Low I Flow, Low Shaft Speed, High Pressurizer Pressure. Low Pressurizer Pressure, High l Pressurizer Level Low Low Steam Generator Level and High Steam Generator Level l Turbine Trip. The remaining reactor trips are either manually actuated or are not active : i during power operation. The ESF actuations are Low Pressurizer Pressure SI, High Steam Pressure Rate Si and Steam Line Isolation, Low Steam Line Pressure Si and i Steam Line Isolation and High Containment Pressure SI, Steam Line isolation and j Containment Spray. Below is an itemization of the effect the VANTAGE SH fuel will have j on the operating margin of each actuation system. l Hlah Flux ; Current Setpoint: 109% of rated thermal power (80% for N 1 operation) (unchanged by VANTAGE 5H fuel) ; The operating transients that provide the most challenge to the high flux reactor trip are the power increase transients. The most severe of these is the 10% step increase from 90% power (65% power for N 1 loop operation). Computer modeling of this transient j shows that the peak reactor power does not exceed 103% for N loop operation or 78% for N 1 loop operation. This is sufficient margin for the N loop case. For the N 1 loop case, variation in the trip setpoint due to channel calibration error and setpoint drift may result in at least a partial trip. However, since power increase transients are generally much ;
- t slower than assumed here, sufficient margin exists to the high flux reactor trip even for N 1 operation.
Overtemperature aT and Overoower AT Reference Setpoints VANTAGE 5H , OTAT Trip Ki 1.08 1.20 K2 0.01313/' F 0.02456/' F , K3 0.00066/ psi 0.001311/ psi f(al) penalty ; Breakpoints 30%, + 10% 26%, + 3% ; Negative slope 3.6%/% al 3.55%/% al Positive slope 2.0%/% al 1.98%/% al > 1 OPAT Trip l K4 1.09 1.09 , KS 0.02/* F 0.02/' F ' Ke 0.00129/'F 0.00180/' F The margin to the OTAT reactor trip during steady state operation at full power is l' O illustrated in Figure 5.51, it illustrates that the revised setpoints for the VANTAGE SH fuel
- 00309 8.10 080990 5 127 l
't u ,, -- , ,- e , ..n. , - , , --.
.m- -~e.. w n
will provide adequate margin to the OTAT turbine runback actuation (2% below the trip setpoint) for axial offsets less than about 8%. At this value, the low break point and high slope on the f (AI) penalty causes a reduction in the trip setpoint of about 10% in AT. When additional margin reductions for control system deadbands (5%) and Tavg i oscillations (3% to cover process noise such as upper head flow anomaly) are considered, the reactor trip setpoint is reduced to about 102% and a turbine runback may occur with AT at 100%. Insufficient margin to the OTAT trip / runback will exist for operation l with an axial offset near the positive RAOC limit (12% al at 100% power to 35% al at 50 % power) down to about 90% reactor power. At lower power levels, the additional margin gained through the reduction in Tavg is sufficient to offset any margin reduction due to the f (al) penalty. On the other end of the scale, adequate margin will exist for normal steady ; state operation with axial offsets near the negative RAOC limit ( 17% al at 100% power to 35% al at 50% power) because the f(AI) penalty will never become effective. As mentioned, the evaluation given above assumes the overtemperature and overpower AT turbine runback setpoint is 2% below the trip setpoint. Westinghouse previously recommended this setpoint as a means of reducing the number of intermittent runback and trip alarms that were occurring following the RTD bypass elimination. The justification used to make this recommendation remains applicable with the VANTAGE SH fuel. If a
- turbine runback setpoint of 3% below the trip is used, adequate margin to the OTAT turbine runback actuation should exist for axial offsets less than about 7.5%. l The operating transient that produces the most severe challenge to the OTAT and OPAT trips is the large load rejection from full power. For Millstone this transient is a 50%
reductiun in turbine power from 102% power (77% power for N 1 operation ) at the maximum turbine runback rate with 40% steam dump capacity and all control systems operating. Since the power reduction is limited by the runback rate, the minimum margin to the OTAT and OPAT trips for Millstone occurs at the initiation of the transient. ' Therefore, sufficient margin to trip should exist for initial axial offsets less than about 8% at full power and 27% at 75% power. Low Pressurizer Pressure Current Setpoint: 1900 psia with a 10/1 lead / lag compensation l (unchanged by VANTAGE SH) 4 J The limiting transient for the low pressurizer pressure reactor trip is the large lor.d rejection as described above for the OTAT and OPAT reactor trips. This study shows that pressurizer pressure (with lead / lag compensation) stays above 2180 psia during the , transient for N and N 1 operation. Thus, there is sufficient margin to the low pressurizer I pressure reactor trip. l l l e 00309 8.1D 080990 5 128
, - - , -. . - - - - , - - .-y _
. _ _ _ _ . . _ . _ . . . _ _ _ _ _ _ _ _ . _ . _ . _ _ .-. _ _ _ _ _ _~
i I l l t Low Low Steam Generator Water Level Current Setpoint: 18.1% of narrow range l l l (unchanged by VANTAGE 5H) Generally there should be adequate margin to the steam generator low low water level ( l reactor trip (based on field experience from other Westinghouse plants). This oxperience l l covers plant start ups and shut downs, steady state operation and design basis load ' l swings. The problems with level control that occur at low power depend mostly on the i characteristics and response of the feedwater control valves, the setpoints in the , l leedwater control system, and the experience of the operators when the control system is in manual. Changing to VANTAGE 5H fuel does not affect these conditions. ; i The other area of concern is the level response during a large load rejection. While plant i startup test at various sites show there should be sufficient margin for a 50% load _ rejection with runback rates up to 133%/ minute, the results are highly dependent on the . response of the condenser steam dump valves. Oper.ing of the steam dump valves ; mitigates the collapse of the voids in the steam generators and limits the extent of the ! level drop. For Millstone there should be adequate margin to the low low level trip with the , steam dump setpoints in the PLS document and design valve response characteristics. ! This condition is not unique to the VANTAGE SH fuel design nor is it significantly affected i by the change in core parameters, but is presented here for information purposes. ! Low Steam Line Pressure Current Setpoints: 658.6 psig with a 50/5 lead / lag compensation (unchanged by l VANTAGE 5H) , l ' l l The margin to Si and steam line isolation were ovaluated by looking at the 10% step load l increase transient which challenges the low steam pressure setpoint. The transient [ resulted in a minimum steam pressure of 877 psig which corresponds to minimum [ compensated pressure of about 822 psig, This is adequate margin to the low steam l pressure setpoint. Also, the loading rate t,ill nornially be much less than analyzed and ; thus willimprove the margin. ; Others The remaining reactor trips and ESF actuations are not affected by the VANTAGE SH fuel ! design: ! e High Flux Rate e Low Flow - fI e Low Shaft Speed f e High Pressurizer Pressure : p e High Pressurizer Level V e High Steam Generator Water Level Turbine Trip . ; 00309-8 10,080990 5 129 i i
, t a 1 e Low Pressurizer Pressure SI e High Steam Pressure Rate Si and Steam Line Isolation l
- High Containment Pressure SI, Steam Line isolation and Containment Spray ,
i 5.5,4 Qqng.ly.g8Ln l j The VANTAGE 5H core should not introduce any significant differences in control system j response and only minor changes to the margin to reactor trips and ESF actuations.~ Low i frequency rod stepping may occur early in the core life when there is a positive or small ! negative MTC. However, this condition will occur with any core with a positive MTC and is [ not unique to VANTAGE 5H. The duration of plant operation at power and temperature conditions that result in a MTC more positive than about 3 pcm/'F should be minimized ! l or operate with the rod control system in manual. These precautions should minimize the - , additional wear on the CRDM caused by the low frequency rod stopping. l The only reactor trip or ESF actuation that does not have adequate operating margin l during the normal conditions transients is the overtemperature AT trip. This condition is l l the result of the change in the f(Al) penalty in the setpoint calculation and the large axial i offset allowed by the VANTAGE 5H core. Margin to this trip la maintained for axial offsets , less than about 3% at full power, Above this axial offset, the reactor protection system I may start generating turbine runback alarms or actuation signals due to the random ; variations in Tav; that are present, including during the steady state operation. O ai f [ r i l I l 00309 8.10 080990 5 130 ;
II i ( : 5.6 RWST/ ACCUMULATOR g,QBM CONCENTRATION INCREASE EVALUATION { 5.6.1 Introduction 'l i I i The large break Loss of Coolant Accident (LOCA) Analysis for Millstone Unit 3 takes no l credit for control rod insertion. Consequently, to maintain the validity of the analysis, it ! must be demonstrated for each cycle that the core can be maintained suberitical via ! boron addition from the ECCS in the unlikely event of a LOCA 2,3.0 ft2, As a result of the i Cycle 4 core design this requirement is not assured with the present plant design. In r order to achieve adequate design flexibility for future cycles, an increase in the boron i concentration range to 2700-2900 ppm for the RWST and 2600-2900 ppm for the j l accumulators is proposed. The following evaluation is provided to address the impacts of l
- this proposed change. -
5.6.2 Non LOCA Safety Analysis The RWST, accumulators, and the Safety injection System (SIS) are subsystems of the ! ! Emergency Core Cooling System. Upon actuation of the SIS, borated water from the ! RWST is delivered to the reactor coolant system in order to provide adequate core f cooling as well as provide sufficient negative reactivity following steamline break transients ; l to prevent excessive fuel failures, The accumulators are a passive system and provide ; borated water to the RCS when the system pressure drops below approximately 600 psig. . The only non LOCA safety analyses in which boron from the RWST or accumulators is taken credit for, or assumed to be present, are those in which the SIS is actuated. These f analyses are: ; e inadvertent Operation of the Emergency Core Cooling System During Power Operation (FSAR Section 15.5.1) ' e inadvertent Opening of a Steam Generator Relief or Safety Valve Causing a i Depressurization of the Main Steam System (FSAR Section 15.1.4) _
- Steam System Piping Failure (FSAR Section 15.1.5) [
- Feedwater System Pipe Break (FSAR Section 15.2.8) :
Rupture of a Control Rod Drive Mechanism Housing /RCCA Election (FSAR , Section 15.4.8) ! Mass and Energy Release Analysis for Postulated Secondary System Pipe. Ruptures (FSAR Section 6.2) [
}
The effect of an increase in the minimum RWST boron concentration from the current - { O level of 2300-2600 ppm to 2700-2900 ppm on each of these transients for both N and N 1 loop operation is discussed in Section 5,1 Non LOCA Accidents. The effect of an . 00309 8.10/080990 5 131 :
~, a,v, , + *w<- e.e,e- -- - . .
increase in the minimum accumulator boron concentration from the current level of 2200- ' 2600 ppm to 2600-2900 ppm was evaluated for transients where the accumulators were actuated. S.6.2.1 Inadvertent Operation of the Emeroency Core Coolino System Durino Power Ooeration p Spurious actuation of the emergency core cooling system while at power would result in a negative reactivity excursion due to the injected boron from the RWST. The decreasing reactor power causes a drop in the core average temperature and coolant shrinkage. If reactor trip on SIS actuation is assumed not to occur, the reactor will ultimately trip on low pressurizer pressure. DNBR never drops tielow the initial value, if the RWST boron concentration were increased from the current minimum value of 2300 ppm to 2700 ppm - the negative reactivity excursion would occur at a faster rate causing a more rapid drop in the core average temperature and coolant shrinkage. The reactor will trip on low ! pressurizer pressure as before, though at an earlier time in the transient. As before the DNBR will never decrease below the initial value. Thus, the conclusions in the FSAR remain valid for both N and N 1 loop operation. 5.6.2.2 Inadvertent Ooenino of a SG Safety Valve ! An accidental depressurization of the main steam system due to the inadvertent opening i of a steam generator safety or relief valve results in a cooldown of the RCS which, in the presence of a negative moderator temperature coefficient, causes a positive reactivity . excursion. Borated water from the RWST enters the core following actuation of the SIS on 1 low pressurizer pressure or low steamline pressure. The negative reactivity provided by j the 2300 ppm water from the RWST limits the return to power to an acceptable level so ! that the :ninimum DNBR remains above the limiting value. As the transient proceeds and ! moro water from the RWST reaches the RCS, the boron concentration in the RCS gradually increases, ultimately causing the core to become subcritical, if the RWST boron concentration were increased to 2700 ppm more negative reactivity would be available to terminate the return to power sooner and at a reduced peak power level. Thus, the maximum core heat flux reached will be reduced. Additionally, the core would become subcritical earlier in the transient. Thus, the minimum DNBR would be higher than for the case currently analyzed and the conclusions in the FSAR remain valid for both N and N 1 loop operation. l I 5.6.2.3 Steam System Ploino Failure A major rupture of a main steam line results in a rapid cooldown of the RCS which,in the presence of a negative moderator temperature coefficient, causes a positive reactivity [ excursion. Borated water from the RWST enters the core following actuation of the SIS on low steam line pressure or low pressurizer pressure. Borated water from the accumulators will enter the RCS whenever system pressure drops below aproximately 600 psia. The negative reactivity provided by the 2300 ppm water from the RWST and accumulators limits the return to power to an acceptable level so that the minimum DNBR 00309 8:10 080990 5132 l
remains above the limiting value. As the transient proceeds and more water from tfie O RWST and accumulators reaches the RCS. the boron concentration in the RCS gradually increases, ultimately causing the core to become subcritical. Increasing the boron l concentrations would add more negative reactivity to terminate the return to power sooner ~ i and reduce the peak power level. Thus, the maximum core heat flux reached wili be ! reduced. Additionally, the core would become subcritical earlier in the transient. Thus, ! the minimum DNBR would be higher than for the cases currently analyzed and the ; conclusions in the FSAR for both N and N 1 loop operation remain valid. t 5.6.2.4 Feedwater System Pipe Break i Following the rupture of a main feedwater line actuation of the SIS may occur. Although
- boron from the RWST is not required to maintain the reactor in a subcritical condition I following a feedwater line break, the cold SIS water serves to reduce the RCS l l temperatures and pressures. An increase in the minimum RWST boron concentration l l from 2300 ppm to 2700 ppm will increass the negative reactivity insertion rate without !
, affecting the reduction of the RCS temperatures and pressures. Thus, an increase in the : RWST boron concentration to 2700 ppm will have no adverse impact on the feedwater line j t break analysis and the conclusions in the FSAR for both N and N 1 loop operation will i remain valid. l 5.6.2.5 Ruoture of a Control Rod Drive Mechanism Housino/RCCA Election ! Following the ejection of a cc ' trol rod the rapid nuclear power excursion causes the RCS to experience a large prestze rise due to the energy released into the coolant. The RCS i pressure then drops as fluid inventory is lost through the break (2.75 inch diameter) in the control rod housing. As the RCS pressure continues to drop actuation of the SIS on low ' l pressurizer pressure will inject borated water from the RWST into the RCS.' An increase i l In the RWST boron concentration from the current minimum of 2300 ppm to 2700 ppm will , result in more rapid negative reactivity insertion to the core and no interference with the ! core cooling capability. Thus, the conclusions in the FSAR for both N and N 1 loop ; operation remain valid, j i
^
1 5.6.2.6 Mass and Enerov Release Analysis for Postulated Secondarv Steam System Ploe Ruotures l A major rupture of a main steam line results in cooldown of the RCS which, in the ; presence of a negative moderator temperature coefficient, causes a positive reactivity ' excursion. Borated water from the RWST enters the core following actuation of the SIS on ; low steam line pressure, low pressurizer pressure, or HI 1 containment pressure. Borated , l- water from the accumulators will enter the RCS whenever system pressure drops below ! approximately 600 psia. The negative reactivity provided by the borated water from the i RWST and accumulators limits the return to power to an acceptable level so that the ! minimum DNBR remains above the limiting value. Adodionally, by limiting the return to: .i power, the borated water reduces the total energy that is dissipated via steam release O- through the ruptured steam line. As the transient proceeds and more water from the
]
5 l D0309 8:10 080990 . 5 133 2 1
'l i I RWST and accumulators reaches the RCS, the boron concentration in the RCS gradually increases ultimately causing the core to become subcritical. Increasing the RWST and
, I accumulator boron concentrations would add more negative reactivity to terminate the return to power sooner and reduce the peak power level. Thus, the maximum core heat l i
flux reached will be reduced, and the core would become suberitical earlier in the i transient. Over the course of the transient. the reduced peak power and earlier return to l suberiticality would reduce the integral mass and energy releases, as a function of time. ! Therefore, the conclusions in the FSAR for both N and N 1 loop operation remain valid. 5.6.2.7 Conclusion An increase in the minimum RWST boron concentration from 2300 ppm to 2700 ppm along with an increase in accumulator boron concentration from 2200 ppm to 2600 ppm, will have no adverse impact upon the non LOCA accident analyses. The conclusions as stated in the FSAR will remain valid. 5.6.3 FSAR LOCA Analysis (10CFR 50.46) Foi sne full spectrum of LOCA postulated breaks, the ECCS is designed to limit the consequences of an accident to within the acceptance criteria of 10CFR50.46. The l analysis takes credit for pumped safety injection from the RWST and passive injection of l accumulator water to prevent or mitigate the resulting peak clad temperature excursion. I Also post LOCA long term core cooling takes credit for the available water in the RWST ; and accumulators in determining the post LOCA RCS/ sump boron concentration and the I hot leg switchover time to prevent boron precipitation. The effect of an increase in the I boron concentrations to 2700-2900 ppm for the RWST and 2600-2900 ppm for the accumulators on these aspects of the LOCA analysis is discussed below. j 5.6.3.1 Small Break LOCA The small break LOCA analyses described in FSAR Section 15.6.5 were performed with the NOTRUMP Evaluation Model. The analyses assume that the reactor core is brought i to a suberitical condition by the trip reactivity of the control rods. There is no assumption requiring the presence of boron in the ECCS water or the need for negative reactivity provided by the soluble boron. Thus the changes in the RWST and Accumulator boron concentrations do not alter the conclusions of the FSAR small break LOCA analysis. 5.6.3.2 Laroe Break LOCA The large break LOCA analysis described in FSAR Section 15.6.5 was performed with the 1981 Evaluation Model + BASH. The analysis does not take credit for the negative reactivity introduced by the soluble boron in the ECCS water in determining reactor power level during the early phases of a postulated large break LOCA. In addition, no credit is
% ken for the negative reactivity introduced by the control rods. During this time period the j reactor is kept suberitical by the voids present in the core. Thus the changes in the RWST . O l 00309-81 0 080990 5 134
i i O and Accumulator boron concentrations do not alter the conclusions of the FSAR large break LOCA analyses. 5.6.3.3 LOCA Mass and Enerov Releases An increase in the RWST and accumulator boron concentrations, would have no effect on i short term mms and energy release since the analysis does not consider any safety ( injection flow taken from the RWST, The long term mass and energy release calculations ; I do not take credit for the soluble boron present in the safety injection from the RWST ; supplied to the RCS. This is similar to the LOCA analyses assumptions, and therefore an j increase in ECCS water boron concentration, would have no effect on the long term mass and Onorgy releases calculated for Millstone Unit 3. l 5.6.3.4 Lono Term Coolino Post LOCA Shutdown ; l Long term cooling is discussed in FSAR Section 15.6.5. The Westinghouse licensing position for satisfying the requirements of 10CFR50.46 Paragraph (b) Item (5) "Long term ; cooling" is defined in WCAP 8339 (page 4 22). The Westinghouse evaluation model } committment is that the reactor remain shutdown by the borated ECCS water. Since credit I for the control rods is not taken for large break LOCA, the borated ECCS water provided l by the RWST and Accumulators must have a concentration that, when mixed with other i sources of water, will result in the reactor core remaining subcritical assuming all control f O rods out (ARO). The post LOCA RCG/ Sump boron concentration will increase as a result of increasing the minimum Tec' Spec boron concentration of the RWST and Accumulators. Confirmation that this proposed increase provides enough margin to keep l { the core subcritical for long term cooling requirements will be verified through the normal i reload evaluation process for Cycle 4. ; L 5.6.3.5 Lono Term Coolino Boron Precioitation l A discussion on the not leg switchover time is presented in FSAR Section 15.6.5, An [ analysis has been performed to determine the maximum boron concentration in the- i reactor vessel following a hypothetical LOCA. This analysis considered a proposed maximum boric acid concentration of 2900 ppm in the RWST.and accumulators and 2700 1 ppm in the RCS. t The analysis considers the increase in boric acid concentration in the reactor vessel during the long term cooling phase of a LOCA, assuming a conservatively small effective - ; vessel volume. This volume includes only the free volumes of the reactor core and upper : plenum below the bottom of the hot leg nozzles. This assumption conservatively neglects i the mixing of boric acid solution with directly connected volumes, such as the reactor i vessellower plenum. The calculation of boric acid concentration in the reactor vessel i i considers a cold leg break of the reactor coolant system in'which steam is generated in ; the core from decay heat while the boron associated with the boric. acid' solution is j T completely separated from the steam and remdns in the effective vessel volume. Procedures philosophy assumes that it would be very difficult for the operator to = i 00309 8:10 080990 5 135 I
,i
. ~ , _ , - , , , - . . _ . . . . . . _ . . . _ . . _ . . - . _ . - _
l i differentiate between break sizes and locations. Therefore one hot leg switchover time is i used to cover the complete break spectrum. The results of the analysis show that the maximum allowable boric acid concentration of
- 23.53 weight percent which is the boric acid solubility limit less 4 weight percent, will not be exceeded in the vessel if hot leg injection is initiated 9 hours after the inception of a LOCA. The operator should reference this switchover time against the reactor trip /SI actuation signal. The typical time interval between the accident inception and the reactor trip /SI actuation signal is negligible when compared to the switchover time.
5.6.3.6 Rod Election Mass and Enerav Releases for Dose Calculations The Rod Ejection mass and energy releases for Millstone Unit 3 are presented in FSAR Section 15.4.8.4. The increase in the RWST and accumulator boron concentrations will be negligible on tne Rod Ejection Accident analysis. Since the Sl flow taken from the RWST is modeled under similar assumptions as in the large break and small break LOCA analyses, there will be no adverse effect on the FSAR Rod Ejection accident. , 5.6.3.7 LOCA Hydraulic Vessel and Loop Forces The blowdown hydraulle loads resulting from a loss of coolant accident are considered in the FSAR. Section 3.6.2.2.2. Section 3.98.1.4.2.4 Section 3.9N.1.4.3 and Section 3.9N.2.5 ! for Millstone Unit 3. The increase in the RWST and accumulator boron concentrations will have no effect on the LOCA blowdown hydraulic loads since the maximum loads are generated within the first few seconds after break initiation. For this reason the ECCS, including the RWST, is not considered in the LOCA hydraulle forces modeling and thus the increase in RWST and accumulator boron concentrations will have no effect on the results of the LOCA hydraulic forces calculations. 5.6.3.8 C.,onclusions The increase in the RWST boron concentration from a range of 2300-2600 ppm to a range of ~2700-2900 ppm and Accumulator boron concentration from a range of 2200 to 2600 to a range of 2600 to 2900 ppm does not have a negative effect on the FSAR LOCA related analyses. Current margin to the post LOCA shutdown requirement is increased with continued conformance venfied through the normal reload evaluation process. The higher concentrations decrease the allowable time for operator action to initiate hot leg recirculation to 9 houe. The resulting time requirement of 9 hours is still more than , adequate to assure those 90erator actions can be accomplished. Therefore, there is no o adverse effect on FSAR LOCA elated accidents for the proposed boron concentration increases. These calculations are gnlicable to both N and N 1 loop operations. 5.6.4 Post Accident Chemical Environment increasing the current boron concentration range of 2300 to 2600 ppm in the Refueling Water Storage Tank (RWST) to 2700 to 2900 ppm has the effect of decreasing the pH of 00309 81 0 080990 5 136 l -
7 i l the containment spray and of the post accident recirculating core cooling solutions. A
\ decrease in the pH can increase the rate of hydrogen production due to the corrosion of f
I zine and can increase the potential for chloride induced stress corrosion cracking of l stainless steel. ! 5.6.4.1 Post LOCA Sumo and Sorav oH ) The required post LOCA pH range of 7.0 to 7.5 can not be assured with the increased ! boron concentration range unless tiicri is :nn increase in the amount of sodium hydroxide i transferred from the Chemical Additive Tank (CAT) to containment via the sprays. .The . ; CAT volume and NaOH concentration requirements were reevaluated to determine the i changes necessary to ensure that the minimum sump pH will be in the required range, j Following are the revised CAT volume and NaOH concentration requirements: ' Current Revised } CAT Minimum Contained Volume (gal) 18,000 17,760 . . CAT Maximum Contained Volume (gal) 19,000 18,760 ! Minimum NaOH Concentration (weight percent) 2.4 3.4 Maximum NaOH Concentration (weight percent) - : 3.1 4.1 Minimim Sump pH 7.0
'O Maximum Sump pH 7.35 7.0 7.25 l
Maximum Spray pH ~ 8.7 8A f i 5.6.4.2 Radioloalcal Conseauences i The radiological consequences of the large. break LOCA are unaffected by the boron increase. This is based primarily on the fact that the radiological analysis presented in ! Section 15.6 of the FSAR does not utlize containment sprays for fission product removal. l If LOCA radiological calculations are performed which do take credit for spray fission ; product removal, the slight change in spray pH would not affect the assumed removal ; efficiency based on information provided in Rev. 2 to the Standard Review Plan Section 6.5.2. 5.6.4.3 Hydrocon Production Hydrogen produced by the corrosion of aluminum and zine is a function of solution pH. ! Since the NaOH concentration will be adjusted to compensate for the increase in the ..' ; RWST boron concentration such that the design basis pH range for the equilibrium sump i is maintained, there is no adverse impact on the hydrogen production due to corrosion. .. 4 00309 0:10 080990 5 137-L .
5.6.4.4 Eouioment Qualification The primary concerns of equipment qualification are the protection of stainless steel [ components of the Emergency Core Cocling System from chloride induced stress ; corrosion cracking, failures of electrical components required to operate post accident, and failures of containment coatings which could jeopardizo the ECCS by flaking or. ! peeling off and then enter the emergency sump and other flow paths, potentially restricting i the flow of emergency core cooling water. I i Protection of Stainless Steel I To minimize the occurrence of chloride induced stress corrosion cracking of stalniess ; steel, the NRC recommends a solution pH in the range of 7 to 9.5 (Reference 22). With ! the combination of increased RWST boron and the increased NaOH concentration in the CAT, the minimum revised sump solution pH is consistent with this recommendation. ' l (. Electrical Comoonents , , Electrical equipment is tested to determine the ability of component seals to exclude the f containment environment from the interior of the component. To maximize the challenge i to the seal materials, high pH sprays, in the range of 8 to 11, have traditionally been used. t The quench spray pH (short term) and the recirculation pH (long term) with the revised i CAT and RWST solutions remain consistent with those currently specified in the FSAR. - Specifically, FSAR Section 6.2.2.1 specifies sump solution pH in the range of 7 to 7.5 and j spray pH at less than 10.5. Hence, there will be no adverse impact on the electrical components, i Additionally, although Westinghouse equipment is qualified for environments containing up l to 2750 ppm boron, the extension of the boron concentration range to as high as 2900 j ppm boron has been reviewed and it has been determined to have no impact on the ' t applicability of the previously generated test data. Containment Coatinos Coatings are used in the containment to provide corrosion protection for metals and to aid . In decontamination of surfaces during normal operation. ' Like electrical equipment, coatings are tested with a high pH solution to maximize the ' potential deterioration of the coating. Since the solution pH values associated with the revised RWST boron concentration and the revised NaOH concentration in the CAT are i consistent with those in the FSAR, there will be no adverse effect on the containt1ent ' coatings. 1 G; D0309-81 D 080990 - S138 L
5.6.5 Conclusions The proposed increase in the RWST and accumulator allowable boron concentration limits to 2700 to 2900 ppm for the RWST and 2600 to 2900 ppm for the accumulators has i been assessed from a safety standpoint. Based on these results, it is concluded that the ' proposed boron concentration increases will have no adverse impact on the non.LOCA accident analysis, the LOCA analysis or LOCA related design considerations and is thus acceptable for implementation at Millstone Unit 3 beginning with Cycle 4. Confirmation that the boron concentration increase provides enough margin to meet post.LOCA shutdown requirements will be confirmed through the normal reload evaluation process, i 1 1 i O ' 00309 81 0 080990 .5139
i
. 1 5.7 RCCA PARKED POSITION 5.7.1 Introduction f i
The Rod Cluster Control Assembly (RCCA) all rods out (ARO) parked position will be i defined each cycle such that the RCCAs will be allowed to p. ark in the range of 222 to 231 ; i steps (physicallimitation of the fully withdrawn position). There are indications that after j keeping the control rods " parked" at the same position for several cycles, some amount ; l of wear will occur to the control rod cladding. This wear is a result of contact between the ! ! control rod and the control rod guide cards which are located in the upper internals. To ! l avoid chronic wear at the same location on the control rod cladding, the " parked" position f can be changed. This does not stop the wear process, but simply relocates it to a i different part of the control rod. Periodic redefinition of the fully withdrawn " parked" I position will spread the wear over a greater surface area of the control rod clad, thus ! minimizing the probability of complete wear through the cladding at any spot. -i l The following safety related analyses as impacted by RCCA repositioning are presented in j this section: { e Non LOCA Accident Analyses l e LOCA and LOCA Related Analyses l ; e Containment Integrity ,
, e Steam Generator Tube Rupture
- l e Mechanical Equipment Performance Evaluation ;
i 5.7.2 Analyses f 5.7.2.1 Non LOCA Accident Analyses I ! Non LOCA analyses performed for this report included consideration of an increase in the ARO position from 228 to 231 steps. The data used in the safety analyses is sufficiently , l conservative to accommodate the repositioning, Specifically, the RCCA worth, total trip ! reactivity, shutdown margin, and reactivity insertion versus time curve remain valid. The ' increase in rod drop time caused by the change to VANTAGE SH has been been ' l specifically incorporated into the non LOCA analysis, The change in the parked position . does not affect the new rod drop time. 5.7.2.2 LOCA and LOCA Related Analyses ' i l 5.7.2.2.1 Small Break LOCA j The small break LOCA analyses were performed with the NOTRUMP Evaluation Model which assumes the reactor core is brought to a subcritical condition by the trip reactivity ( of the control rods. Since it has been determined that the rod drop time and trip reactivity j have not altered for the slight increase in steps of the ARO position, there will be no effect - O' on the reference FSAR small break LOCA analyses. 1 i D0309 8.1D'o80990 5 140 I _4 -. _ ._ _ __ _
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- 5.7.2.2.2 Laroe Break LOCA f There is no impact on large break LOCA analysis for the slight increased in the ARO i
- position. ;
5.7.2.2.3 Post LOCA and Lono Term Core Coolino ! i The Westinghouse Evaluation Model commitment is that the reactor will remain shutdown i by borated ECCS water residing in the RCS/ sump after a LOCA. Since credit for the ; control rods is not taken for a large break LOCA, the borated ECCS water provided by the . accumulators and the RWST must have a boron concentration, when mixed with other : water sources which will result in the reactor core remaining subcritical assuming ARO. Since there is no credit taken for control rods, there will be no change in the calculated RCS/surnp boron concentration after a postulated LOCA, and there is no adverse effect on ' the post LOCA long-term core cooling requirement as presented in the FSAR. It is conservative to assume no credit for control rod insertion since this renults in a more i reactive condition upon which the evaluation is performed. - i 5.7.2.2.4 Hot Leo Switchover to Prevent Boron Precloitation The switchover time is dependent on power level, RCS, RWST, and accumulator water ; volumes, and boron concentrations, Increasing the steps for the ARO position does not : affect either the power level or the maximum boren concentrations assumed for the RCS, # RWST, and accumulators in the hot leg switchover calculation. Thus, there is no adverse 'i effect on the post LOCA hot leg switchover time present6d in the FSAR. l t 5.7.2.2.5 Rod Election Mass Releases ; i Since the rod drop time doss not change from that determined for VANTAGE SH, the [ change in ARO position does not create an adverse effect on the rod ejection mass - releases as presented in the FSAR. l 5.7.2.2.6 LOCA Hydraulic Forcina Functions > t Based on the current LOCA analyses, the shortest time calculated for tripping the control l rods is about 0.5 seconds after a postulated LOCA. The peak loads (which generally , subside well before 500 milliseconds) generated on the reactor vessel and internals as a l result of a postulated LOCA typically occur between 10 to 50 milliseconds after the break.~ Since the hydraulic forces as a result of a LOCA have peaked and subsided before the . time the control rods are calculated to irlp, the change in ARO position will have no effect i on the structural analysis presented in the FSAR. Oi :
?
D0309 8.10 080990 5 141 t?W+%
I I i 5.7.2.3 Mechanical Per'ormance Evaluatjg.nn ' l The RCCAs are designed to fall via gravity into the reactor core when power is removed : l from the drive mechanisms. The time to fall into the dashpot region of the core is > l- currently assumed to be 2.7 seconds. Mechan!cally, the repositioning will not affect the t operation of the control rod drive mechanisms (CRDM) and as such does not represent a l design or operational issue. Therefore, the conclusions of any analyses which assume a ( l . I 2.7 second rod drop time remain valid for the range of parked positions from 222 tc 231 l steps. l l i l 5.7.3 Conclusions ! l l Changing the Rod Cluster Control Assembly (RCCA) to an all rods out (ARO) parked - , , position which allows the rod to sit in the range of 222 to 231 steps presents no adverse ' effect on the Millstone 3 FSAR LOCA and LOCA related analyses and does not represent a design or operational issue (i.e. conclusions remain valid) for any mechanical performance evaluation which assumes a 2.7 second rod drop time. ; I i L O : i t l i i 5 142 00309 8:10.080990 - I
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I i 5.8 INSTRUMENT UNCERTAINTY - REVISED THERMAL DESIGN PROCEDURE (RTDP)
- Introduction Using the Revised Thermal Design Procedure (RTDP) methodology requires a review of temperature, pressure, power and flow uncertainties used in the safety analysis. This is an NRC requirement to obtain their approval for the new methodology that develops reactor core and DNB limits. For Millstone 3 the uncertainties are calculated based on the installed plant instrumentation or special test equipment and Millstone 3 calibration and calorimetric procedures A report has been written to describe the uncertainty ;
evaluation (Reference 43) which is submitted under separate cover to the NRC in support - of the safety analysis. 1 In addition, the uncertainties developed for implementing RTDP methodology are included in analyses where it is not appropriate to use RTDP, Instead, these events apply uncertainties using the Standard Thermal Design Procedure (STDP). General Method of Uncertaintv Analysis - The method of uncertainty analysis is discussed in detail in Reference 43 and is the same regardless of whether the application to the safety analysis is RTDP or non RTDP methodology. The uncertainty analysis is statistical and combines various uncertainties using the square root of the sum of the squares. The analysis includes uncertainties for I
- k. the method of measurement, for the type of field devices (RTDs, transmitters, special test {
measurements), and for the calibration of the instrumentation. The temperature, pressure, l power and flow uncertainties are used in the development of the reactor core limits and l the DNB limits. The delta T reactor trip setpoints are developed for the technical 1 specifications from the new reactor core limits. I I With respect to the application of uncertainties, the non LOCA design basis events are . 1 divided into two categories; those that employ RTDP and those that use non RTDP (i.e., STDP) methodologies. Generally, RTDP is applied to analyses that are primarily intended ( to address DNB concerns. However, certain DNB events have initial conditions, such as ! hot zero power, that are not fully consistent with RTDP methods and those events continue i to use STDP methodology. Those events that are not DNB related also apply STDP methodology. For each category of events, the uncertainties are applied differently in - developing initial conditions for the analysis. Aoolication of the.RTDP Methodolooy For the RTDP events, the uncertainties are included in the development of the limit DNBR
- (see Section 4.0 of this report) and nominal values are assumed for the initial conditions for RCS pressure, RCS temperature and reactor power. Minimum measured flow (MMF) is a newly defined parameter used for RTDP events.and is equivalent to the thermal design flow (TDF) plus a flow uncertainty. The primary acceptance criteria for the RTDP events are developed using the limit DNBRs as defined in Section 4.0. These limit ;
00309 8.10 080990 5 143
f i i ! l DNBRs are elInt specific since they are based on equipment and procedures at Millstone Unit 3. Acolication of the STOP Methodolooy For the events using STDP methodology, the uncertainties . ire applied consistent with the reference licensing basis analyses found in the Millstone Unit 3 FSAR. The uncertainties j are directly applied to the nominal values for RCS pressure, RCS temperature and power - to define the initial conditions for the non LOCA events. Whether positive or negative ; uncertainties are applied is determined by defining the most conservative direction on a j specific event. Thermal design flow is used for these events. As noted earlier, some i events that use STDP continue to consider DNBR as the primary acceptance criterion,- l though the majority of events in this category consider other acceptance criteria. , Results ' The uncertainty results which are based on the Millstone 3 instrumentation and procedures, are presented in Reference 43. Conclusions Reference 43 documents the pressure, temperature, power and flow uncertainties that have been determined for Millstone 3. More conservative values have been used in the analysis. All FSAR acceptance criteria have been met for N and N 1 loop operation. 9 o0309 8.10 080900 5 144
'I 5.g SAFETY EVALUATION FOR VANTAGE SH AND ASSOCIATED CHANGES l 5.g.1 Backaround !
I i This safety evaluation addresses the upgrade to VANTAGE 5' tuel and the additional changes discussed in Section 5.0 of this report. These changJs fall into four categories. A description of these four categories and a listing of the specific changes in each is given below. 5.g ?.1 Chance in Fuel Tvos from 17x17 Standard to 17x17 VANTAGE 5H The specific features of VANTAGE 5H which represent a change from standard fuel are:
- Zircaloy grids the inconel structural grids used in standard fuel are replaced by 1 Zircaloy grids in VANTAGE 5H (except for the top and bottom grids which remain inconel) ,
o Intermediate Flow Mixing (IFM) grids these are three non structural grids which are added at the mid span locations of the upper portion of the fuel assembly; the use of IFM grids results in a change in the rod drop time from 2.2 to 2.7 l seconds. -
) , Additional information on these features can be found in sections 2.0 and 5.0.2.
Westinghouse has full scope for this change in fuel type and is supplying all necessary analyses and documentation to support licensing the use of VANTAGE 5H for cycle 4. This represents a change in the facility as described in the FSAR. 5.G 1.2 Associated Fuel / Core Related Chanaes l Numerous changes are being made which are closely related to the change in fuel type. l Three of these changes represent physical changes to be made at the plant. These three ) changes are: j 1 e Thimble plug deletion these devices will be removed over several cycles e RCCA parked position change-the full out position of the RCCA banks will vary in the range of 222 231 steps withdrawn e RWST boron concentration increase the boron concentration maintained in the RWST willincrease from the range 2300 2600 ppm to 2700 2900 ppm One of these changes involves a change to the plant procedures: I e Relaxed Axial Offset Control (RAOC) the plant will utilize this method of control , which allows a wider axial offset band compared to Constant Axial Offset Controi which is currently being used i l 0030911:1D 080990 5145 _._, . .-I
I l I . The remainder of the changes in this category are modifications to the safety analysis procedures or inputs: e increase in FAH and F q .
- Reactivity coefficient changes :
- Revised Thermal Design Procedure (RTDP) i
- Revised non RTDP uncertainties i e Reduced shutdown margin {
- Modified OTAT and OPAT trip setpoints !
l A detailed discussion of each of these changes can be found in section 5.0.3. ; l Westinghouse has full scope for each of these changes and is supplying all necessary j licensing documentation to support their implementation. These changes represent modifications to the safety analysis of record as described in the FSAR, changes to ; j licensed design methodology previously employed for Mills' tone 3 and t.;hanges to the [ ! plant protection system. l 5.9.1.3 Plant / Tech Soec Related Chances i The following changes were not necessitated by the change to VANTAGE SH but were requested by Northeast Nuclear Energy Company. '
- Expanded accumulator level band O'f
- Revised pressurizer level program i As a result of this effort. Westinghouse fully supports the implementation of these changes. ,
l These changes are discussed in more detail in section 5.0.4. These changes represent a change to the plant as described in the FSAR. , 5.9.1.4 Additional Assumotions Incoroorated into Safety Analysis ' f The effect of the following changes were incorporated into the safety analysis at the l request of Northeast Nuclear Energy Company. These items are discussed more fully in ! Section 5.0.5. , e increased safety valve drift (pressurizer and steam generator) from + / 1% to l
+ /. 3% :
- Revised auxiliary feedwater flow rates ,
e increase in delay time by 15 seconds for diesel start '
- Negative flux rate trip deletion
- 10% steam generator tube plugging. ;
- Degraged safety injection and charging parameters
- Modified N.1 loop conditions
- Modified safety injection system flow source for the steamline break event
- Modified setpoints (high steam generator ~ level. low steam pressure, low pressurizer pressure, low reactor. coolant flow) i
?
00309 11:10 080900 -5 146 j
l l l l l O These additional assumptions are conservative and require no Tech Spec changes for the analysis to be valid. However, additional work is required before these Tech Spec ! i changes can be implemented at the plant. ! The safety analysis performed demonstrated the acceptability of all of the above changes ! in each of the four categories. Transition core effects and N 1 (3) loop operation were
~
i also addressed in the analysis. l 5.9.2 Licensina Basis The current licensing basis for Millstone 3 is reflected in the Millstone 3 FSAR Amended.' l The changes discussed in Section 5.9.1 represent changes to the plant as described in ! the FSAR, changes in the methodology employed in certain areas of the reference safety l analysis and changes to the input and assumptions used in the reference safety analysis, } These changes must be evaluated to determine if an unreviewed safety question exists. ! Chapter 10 of the Code of Federal regulation, Section 50.59 (10CFR 50.59) allows the ; holder of a alconse authorizing operation of a nuclear power ft.ciltity the capacity to initiate ! certain changes, tests, and experiments not described in the FSAR. Prior NRC approval ! is not necessary to implement the proposed change, test, or experiment if it does not ! involve an unreviewed safety question or a change in the technical specifications incorporated in the license, it is, however, the obligation of the licensee to maintain a record of these changes, tests, or experiments to the facility to the extent that such O changes affect the FSAR.10CFR 50.59 further stipulates that these records shallinclude a written safety evaluation which provides the bases for the determination that the change, . [ test or experiment does not involve an unreviewed safety question. This safety evaluation ; concludes that implementation of the changes discussed in Section 5.9.1 does not constitute an unreviewed safety question. These changes will, however, require changes ! i to Technical Specifications. A license amendment will be required to implement these , ! changes. ! l The conclusion that there is no unreviewed safety question involved is based on the { consideration of the licensing basis and the applicable acceptance criteria'as defined in i the following documents. Final Safety Analysis Report, Millstone Nuclear Power Station, i Unit 3: Chapters 4,6 and 15 (as amended through October 1989) ; ! NUREG 0800 NRC Standard Review Plan ; Technical Specifications, Amendment 51 i ASME Code l Title 10 of the Code of Federal Regulations 50.46,50.59 l l These documents require the following performance and safety requirements to be met: j! O 1, Fuel damage (defined as penetration of the fission product barrier, i.e. the fuel rod > clad) is not expected during Condition I (Normal operation) or Condition 11 (Incidents f 00309 11:10 080990 5 147 :
- + ,y -, -w+ -> , -
.m,e=m wvw
1 I of Moderate Frequency) events, it is not possible, however, to preclude a very small number of rod failures. These will be within the capability of the plant cleanup system and are consistent with the plant design bases.
- 2. The reactor can be brought to a safe state following a Condition lli (Infrequent faults) event with only a small fract!on of fuel rods damaged, although sufficient fuel damage might occur to preclude resumption of operation without considerable outage time.
- 3. The reactor can be brought to a safe state and the core can be kept suberitical with acceptable heat transfer geometry following transients arising from Condition IV (Limiting faults) events.
I To satisfy the above requirements, the following criteria have been established. )
- 1. Departure from Nucleate Boiling (DNB) Design Basis There will be at least a 95% probability that DNB will not occur on the limiting fuel l rods during normal operation ~and operational transients and any transient conditions !
arising from faults of moderate frequency (Condition i or ll events), at a 95% l confidence level. Table 4 2 provides the DNBR limits for four_ and three loop ' operation for VANTAGE SH and standard fuel, i
- 2. Fuel Temperature Design Basis :
During modes of operation associated with Condition I and Condition ll events, there is at least a 95% probability that the peak kw/ft fuel rods will not exceed the UO2 - I melting temperature at the 95% confidence level. The melting temperature of UO2 l Is taken as 4900'F (unirradiated) and 4800*F at end of life. By precluding UO, ! melting, the fuel geometry is preserved and possibly adverse effects of moltra UO2 i on the cladding are eliminated. To preclude center melting and to provide a basis for overpower protection system setpoints, a conservative centerline fuel temperature has been selected as the overpower limit.
- 3. Reactor Coolant System Pressure ,
Peak RCS pressure is not to exceed 110% of the design pressure during Condition ll and lil events. Peak RCS pressure is not to exceed 120% of the design pressure - I during Condition IV events. '
- 4. Loss of Coolant Design Bases (10CFR50.46) j A. Peak cladding temperature :
The calculated maximum fuel element cladding temperature shall not exceed ! 2200'F. l i l j l 00309 11:10.080990 5 148-
._. _ _ _ _ _ _ _._. _ ~~ ~_ _ _ . _ . . . _ . _ - _ ,
i i B. Maximum cladding ' oxidation The calculated total oxidation of the cladding shall nowhere exceed 0.17 times ' the total cladding thickness before oxidation. C. Maximum hydrogen generation . l The calculated total amount of hydrogen generated from the chemical reaction j
, of the cladding with water or steam shall not exceed 0.01 times the hypothetical !
amount that would be generated if all of the metal in the cladding cylinders'- ! surround;ng the fuel, excluding the cladding surrounding the plenum volume, ; were to "sact. D. Coolablo geometry I Calculated changes in core geometry shall be such that the core remains : amenable to cooling. I E. Long term cooling . After any calculated successful initial operation of the ECCS, the calculated . t core temperature shall be maintained at an acceptably low va ue and decay heat shall be removed for the extended period of time required by the long lived i radioactivity remaining in t% core. ; i
- 5. Rupture of an RCCA Drive Mechanism Housing j Rupture of an RCCA drive mechanism housing (RCCA ejection)is classified as a Condition IV event. Limiting criteria for this accident are given in the NRC's :
Regulatory Guide 1.77. However Westinghouse has historically applied the ! following, somewhat more conservative, criteria: ; A. Average fuel pellet enthalpy at the hot spot below 225 cal /gm for unirradiated fuel and 200 cal'gm for irradiated fuel.
- 8. Peak reactor coolant pressure less than that which would cause stresses to !
exceed the faulted condition stress limits. ~ C. Fuel melting is limited to less than 10% of the fuel volume at the hot spot even if the average fuel pellet enthalpy is below the limits of criterlon 1 above, j
- 6. Containment pressure must be maintained below the design pressure. t
- 7. Offsite doses must be maintained below the acceptance limits contained in ,
10CFR100. . l l The above criteria provide adequate safety margin. Meeting these criteria insure the - , margin of safety as defined in the basis of the technical specifications has not been 1 reduced. These design criteria are interpreted as safety limits for the safety evaluation. 1 ? 00309 11:10 080990' 5149
. I
t The compliance to these requirements is demonstrated in this safety evaluttion. 5.9.d Evaluation i t 5.9.3.1 NON LOCA Evaluation Non LOCA safety evaluations were performed to assess the affect of the VANTAGE SH upgrade and additional changes specified in Section 5.0 on the affected transients presented in Section 15 of the FSAR. The non LOCA reload safety evaluation j methodology is described in Reference 4.- This methodology is applicable to VANTAGE 5H transition and full cores. Section 15 of the Millstone Unit 3 FSAR provides information on the analysis codes and methods that are applicable for the non LOCA reanalyses performed for this report. The key safety parameters are documented in Reference 4. Values of these safety parameters which bound both fuel types,17x17 Standard and ' t VANTAGE SH, were assumed in the non LOCA safety analyses. I l Each affected transient was either re analyzed or re evaluated depending on the sensitivity' of the transient to the change in fuel type and additional changes. Additionally, the possible difference in relative behavior of the three loop cases as compared to the fcur- 'l loop cases had to be taken into consideration. Table 5.1.2 shows whether an analysis or evaluation was performed for each affected transient to determine the affect on that transient of the changes being addressed. The assumptions included in the evaluation. ' are discussed in Sections 5.1.3 through 5.1; through 5.1.8. For all transients which were re analyzed or re evaluated, the applicable acceptance criteria were met and therefore the conclusions of the FSAR remain valid. 5.9.3.2 LOCA Evaluations i 5.9.3.2.1 Larae Break LOCA (LBLOCA) Evaluations Large break LOCA (LBLOCA) analyses have been performed fcr Millstone 3 which assess
.[
the affect of the VANTAGE SH upgrade and additional changes specified in Section 5.0 on the results presented in Chapter 15 of the FSAR. ' A detailed discussion of th's analysis is. , provided in Section 5.2.1 of this report. The changes'specified in Section 5.9.1 above were incorporated into the analyses.. The Westinghouse 1981 Evaluation Model + BASH was utilized and a spectrum of cold leg breaks were analyzed. This represents a charge < in the methodology from the current licensing basis. The effect of the transition core l cycles were conservatively evaluated by applying a 50 F penalty to the peak clad ; L temperature (PCT) obtained for the full core of VANTAGE SH. A LBLOCA analysis was also performed for Standard fuel which examined the most limiting single break in order to assure that the most limiting PCT was identified. ' l LBLOCA analysis was performed to address the condition of operation with a loop out of se.a.> ice (3 loop operation). Analysis was pt."ormed which ' examined both a full core of VANTAGE SH and a full core of Standard fuel. Again the Westinghouse 1981 Evaluation 1 00300-11:10.080990 5 150 L . s
. . ._ . _. u.._ ,_. ..
m Modei + BASH was utilized with modifications necessary to account for the loop out of service configuration. An evaluation was performed to address concerns related to.the break occurring in the inactive loop. The results of the LBLOCA analysis demonstrate that for double ended cold leg guillotine breaks in the reactor coolant piping, the Emergency Core Cooling System (ECCS) will meet the acceptance criteria of 10CFR 50.46. This conclusion applies for either a full-core of VANTAGE 5H fuel or Standard fuel as well as transition cores. 5.9.3.2.2 Small Break LOCA (SBLOCA) Anaissis The SBLOCA was analyzed to determine the affect on the FSAR Chapter 15 analysis resulting from the upgrade to VANTAGE 5H fuel and other changes specified in Section 5.0 above. The currently approved NOTRUMP Small Break ECCS Evaluation _Model was utilized for a spectrum of cold leg breaks. This represents a change in methodology _ from the current licensing basis. A detailed discussion of the analysis performed is found in - Section 5.2.2 of this report. - The changes specified in 5.9.1 above were incorporated in ' the analysis. Specific analyses were not performed for Standard fuel since the significant difference in the fuel type with respect to LOCA analysis is the use of a different grid design. Since the grids are not modelled in the SBLOCA analysis the results for Standard fuel would be equivalent to those for VANTAGE SH. Specific SBLOCA analyses for three loop operation. were also not performed since plant conditions for four loop operation were judged to oe bounding. The results of the SBLOCA analysis show that the acceptance criteria of 10CFR50.46 are. met. 5.9.3.2.3 Lono Term Coolino Post LOCA Shutdown . ' Long term cooling is discussed in FSAR Section 15.6.5. - The Westinghouse l'icensing position for satisfying the requirements of 10CFR 50.46 Paragraph (b) Item (5) "Long term cooling" is defined in WCAP 8339. The Westinghouse evaluation model committment is that the reactor remain shutdown by the borated ECCS water. :A discussion of this topic is - ! provided in Section 5.6 of this report. Confirmation that the proposed increase in the - y boron concentration of the RWST and accumulators will provide enough margin to' keep the core subcritical for long term cooling requirements will be verified through the normal reload evaluation process. 5.9.3.2.4 Lono Term Coolino Boron Precioltation A discussion of the hot leg switchover time is presented in FSAR Section 15.6.5. An analysis has been performed t.) determine the maximum boron concentration in the g reactor vessel following a hypothetical LOCA. This analysis considered a proposed _ - maximum boric acid concentration of 2900 ppm in the RWST and accumulators and 2600 q 00309 11:10 080990 5 151 ! m
L i j l
\
l ppm in the RCS. A discussion of this topic is provided in Section 5.6 of this report The- I results of the analysis show that the maximum allowable boric acid concentration of 23.53 l weight percent which is the boric acid solubility limit less 4 weight percent, will not be i exceeded in the vessel if hot leg injection is initiated 9 hours after the inception of a LOCA. ! This is a reduction from the current licensing basis time of 11 hours. The typical time interval between the accident inception and reactor trip /Si actuation signal is negligible ; when compared to the switchover time. 5.9.3.2.5 LOCA Reactor Vessel and Looo Forces ! Section 5.2.5 provides a discussion of the effect of the upgrade to VANTAGE SH and associated changes on the LOCA reactor vessel and loop forces, it has been concluded I that these changes do not increase the consequences of a hypothesized LOCA on the reactor vesselinternals or RCS loop piping considering both fou'r and three loop operation. Thus the reference FSAR analysis is considered bounding,- 5.9.3.3 Mass and Eneroy Release and Containment Inteority , The effect on LOCA mass and energy release caused by the changes discussed in 'i Section 5.9.1.1 through 5.9.1.3 above have been evaluated A detailed discussion of the ' ' evaluation is presented in Section 5.3 of this report. The evaluation concludes that there ; is no impact of these changes on LOCA mass and energy release. i The effects on LOCA mass and energy release of increasing the delay time for dies'el start and degraded safety injection flow were specifically excluded since these would require extensive analysis and were considered out of scope for this effort. Therefore, the diesel start delay cannot be implemented until this effort is completed. Mass and energy release associated with the steam line break event is primarily influenced by the changes in core wide power level. It is therefore not affected by the changes discussed in Section 5.9.1. [ Westinghouse does not have containment integrity and subcompartment analysis scope. However, as discussed above, the mass and energy release data, which are inputs to the containment and subcompartment analyses, were not adversely affected from the upgrade ; to VANTAGE 5H and the additional changes. 5.9.3.4 Radiolooical Consecuences of Accidents l l The I apact on the radiological consequences of accidents has been has been evaluated l takmg into account the changes discussed in section 5.9.1 above. A discussion of this l evaluation is presented in Section 5.4 of this report. The evaluation concludes that there
~
would be no expected increase in the isotopic inventories or in the activity gap fractions, Thus, the doses reported in the FSAR remain bounding.- 0 00309 11:10 080990 5-152 L
1 5.9.3.5 Plant and Systems Evaluation 5.9.3.5.1 Post Accident Chemical Environment Increasing the boron concentration range in the RWST from 2300 2600 ppm to 2700 - 2900 ppm has the effect of decreasing the pH of the containment spray and of the post-accident recirculating core cooling solutions. A decrease in the pH can increase the rate of hydrogen production due to the corrosion of zinc and can increase the potential for chloride induced stre+3s corrosion cracking of_ stainless steel. Since the NaOH concentration will be adjusted to compensate for the increase in the - RWST boron concentration such that the design basis pH range for the equilibrium samo t is maintained, there is no adverse impact on the hydrogen production due to corrosion; To minimize the occurrence of chloride induced stress corrosion _ cracking of stainless steel, the NRC recommends a solution pH in the range of 7 to 9.5 .With the combination: of increased RWST boron and the increased NaOH concentration in the chemical addition J tank (CAT), the revised sump solution pH range is consistent with this recommendation.- Coatings are used in the containment to provide corrosion and chemical reaction protection for metals and concrete and to aid in decontamination of surfaces during normal operation. Coatings are tested _with a high pH solution to.maxiriize the potential deterioration of the coating. Since the solution pH values associated with the revised RWST boron concentration and the revised NaOH con'centration in the CAT are consistent - O with those in the FSAR, there will be no adverse effect on the containment coatings. 5.9.3.6 Instrumentation and Controls System Performance All changes given in Section 5.9.1 have been reviewed for their impact on instrumentation' and controls systems. Those changes in the categories of change in fuel type, associated fuel / core related changes and plant related changes have no impact on the ability of the safety systems to perform their intended function. The changes in the category of 5.9.1.4 may have some impact. The evaluation of the impact of these changes in that category was considered to be outside the scope of this effort. Of the changes in the first three categories, only the change in RWST boron concentration had any impact on I&C systems. This is because of the equipment qualification testing ' parameters for class 1E electrical equipment which is discussed below.. 5.9.3.7 Eauioment Qualification
~
Westinghouse supplied Class 1E electrical equipment is tested to determine the ability of . component seals to exclude the containment environment from the interior of the - component. To maximize the challenge to the seat materials, high pH sprays, in the range of 8 to 11, have traditionally been used. O
. 00309 11:10 080990 '5153
-f 1
5 i The quench spray pH (short term) and the recirculation pH (long term) with the revised -l CAT and RWST solutions are comparable to those currently specified in the FSAR. ; Hence, there will be no adverse impact on the electrical components, j Additionally, although documentation exists to qualify Westinghouse equipment to :
- environments containing up to 2750 ppm boron, the extension of the boron concentration >
range to as high as 2900 ppm boron has been reviewed and it its not considered to have i a negative impact on the applicability of the previously generated test data.- { I 5.9.3.8 Comoonent Performance l The impact on the reactor coolant systems and components performance has been - i reviewed with respect to the impact of the changes discussed in Section 5.9.1 above. Of ; the changes in the first three categories, only two may have any impact in this area: - change in RCCA parked position and thimble plug deletion. The change in RCCA parked position is discussed in Section 5.7. Repositioning is being implemented as a management scheme to mitigate the effects of wear and prolong the life ! of the control rods to perform their functional requirements. These functional requirements include providing negative reactivity, maintaining geometric configuration for l expected duty and maintaining continuous free movement. Mechanically,' the ; repositioning will not affect the operation of the control rod drive mechanisms and as such i does not represent a design or operationalissue. Further, the rod drop time criteria of 2.7 { seconds is met for the range of parked positions from 222 to 231 steps withdrawn. - Therefore, the conclusions of any analyses which assume a 2.7 second rod drop time remain valid. j ( Thimble plug deletion has been reviewed with respect to the manner in which'the core i components will respond to LOCA and seismic forces, it has been determined that ' removal of the thimble plugs is a benefit in this area. NUSCO has indicated that the' thimble tube plugging devices will not be removed all at one time but rather over two or t three cycles. There are no restrictions in placement of the remaining thimble plugs in the I core in this situation. , 5.9.4 Determination of Unreviewed Safety Question The information presented above forms the basis for responding to the. questions posed in 10CFR 50.59 for the determination of an unreviewed safety question.
- 1. The probability of occurrence of an accident or malfunction of equipment important to safety previously evaluated in the safety analysis report increased is not:
increased. - The mechanical design changes associated with VANTAGE SH, the Zircaloy grids and IFM grids, result in the capability for relaxation of analytical input parameters such that increased margin can be' generated without violation of any acceptance 00309-11 10 080990 5 154' - 1 m - - - + r - . _
- criteria. This margin can then be appiled towards relaxation'of operationallimits. i O' such as reduced safety injection flow or. increased steam generator tube plugging, in each case however, the appropriate design and acceptance criteria are met. No new performance requirements are being imposed on any system or component inn order to support the revised analysis assumptionse Subsequently, overall plant integdty remains at least consistent with that established by the original licensing" q basis. Furthermore, the features of VANTAGE 5H which are different from Standard I fuel and the associated fuel / core related changes included as modifications to the l safety analysis procedures or input (see Section 5.9.1) are associated with features ;
used as limits or mitigators to assumed accident scenarios'and not accident: . { initiators. These features affect the consequences of postulated accidents,.not their. probability of occurrence. 4 1 Deletion of the thimble tube plugging devices does not affect the probability of . , occurrence of an accident. The RWST is an accident mitigator rather than initiator; I and as such, increasing its bor'on concentration does not affect the probability of i occurrence of an accident. . RCCA repositioning does not increase the probability'of:. j any fuel related accident since it merely is a change in position not in the system - 1 that controls their operation. l The increase in RWST boron concentration has been evaluated with respect to'its effect on components in a post accident environmentc The corrosion of aluminum, zinc and stainless steel components are affected by the'pH of tnelr environment. The volume and NaOH concentration of the Chemical Addition Tank'have been D t modified to neutralize the effect of the. higher boric acid concentration coming from .! the RWST. There is thus no affect on the performance of components containing
' ]
these materials. The other plant related changes and additional assumptions incorporated into the : 3 safety analysis discussed in Section 5.9.1 also are not initiators.of any accident and-therefore do not affect the probability of occurrence.'These changes rather affect the consequences of an accident.
- 2. The consequences of an accident or malfunction of equipment important to safety i previously evaluated in the safety analysis report are not. increased.' l i
The evaluation of the impact of the proposed changes specified in Section 5.9.1 is l discussed in Section 5.9.3 above. The evaluation addressed a full core of - VANTAGE SH as well as transition cores consisting of VANTAGE 5H and Standard fuel. Four loop and three loop operation also was addressed. ; An evaluation of the , radiological consequences of the changes was included as part of the evaluation ] L and the results are summarized below. L . The change to VANTAGE _5H and the increase in discharge burnup to 45.000 l MWD /MTU has been determined not to affect the radiological source terms used to -
O calculate accident doses. The fuel rod size is not changing so there is no impact l
0030911:1D 080090 5 155 l' ' j u , . .
on the core inventory of radioactive isotopes. Also, the extension of fuel burnup has been shown to have negligible impact on this parameter, The fraction of core activity that is assumed to migrate out of the fue! matrix and into the fuel clad gap . region and thus is available for release in the event the cladding is breached remains well below the limits specified in Regulatory Guides 1.25 and 1,77. The accident parameters which have_ a direct impact on dose release include fuel damage, offsite steam releases and primary to secondary leakage, These _ parameters have been reviewed for the following events: main steamline break, locked rotor, loss of AC power, rod ejection and LOCA; The results of the accident ' analysis shows that these parameters are not adversely affected by the upgrade to VANTAGE SH and the associated changes, Since these parameters are not adversely affected, the radiological consequences of the accidents are not - Increased.
- 3. The possibility of an accident or malfunction of equipmerit important to safety'which is different frons that already evalu'ated in the FSAR is 'not created.
Mechanical evaluations have been performed on the fuel assemblies, RCCAs and. other equipment important to safety to confirm that their function and reliability are - not negatively impacted due to the upgrade to VANTAGE 5H and the additional changes discussed in Section 5.0.1. The removal of thimble tube plugging devices and the change in the RCCA parked position were specifically evaluated and the conclusion was made that there was no negative impact on_ safety as a resu_It of = making these changes. No new accident scenarios, failure mechanisms or limiting single failures are introduced as a result of the fuel transition; The presence of : VANTAGE SH fuel assemblies in the core or the revised analytical assumptions have no adverse effect and'do not challenge the performance of any other safety' i related system. Therefore, the possibility _of a different accident or malfunction of
-equipment important to safety has not been created,-
- 4. The margin of safety as defined in the basis for any technical specification is not reduced.
The margin of safety in the plant licensing basis _which is affected by the upgrade to VANTAGE SH fuel and associated changes discussed in Section 5.0.1 is defined in the BASES to those technical specification identified in Section 5.9. These BASES and the scpporting technical specification values are defined by the accident-analyses which are performed to conservatively bound the operating basis defined by the technical specifications and to demonstrate meeting the regulatory acceptance limits. The upgrade to VANTAGE SH fuel and the other changes discussed in Section 5.0.1 were evaluated against the applicable acceptance criteria. The criteria are discussed in Section 5.9.2 and are summarized below: 9' 00309 11:10,080990 - 5 156 a
t I -- Fuel related criteria: e DNBR greater than safety analysis limit
- PCT less than 2200'F for LOCA :
e fuel centerline. temperature less than 4900'F (BOL),4800'F (EUL) - e average fuel pellet enthalpy less than 200 cal /gm for rod ejectionL e Fuel melting limited to 10%~ for rod ejection' e remainder of 10CFR50.46 criteria (clad oxidation, hydrogen generation o coolable geometry, long term cooling)- RCS oressure boundary related criteria: e pressure less than 110% for Condition 11 and til events e pressure less than 120% for Condition IV~ events Containment oressure: e pressure less than design pressure A significant increase in DNB margin is provided by the IFM grids which are a feature of VANTAGE SH. Additional DNB. margin is created by theLuse of the WRB 2' correlation and the use of the Revised Thermal Design Procedure. This additional, DNB margin can then be used to increase the FAH limit as well as remove thimble tube plugging devices. The non LOCA analyses confirm that the DNB design basis is met both for Standard 17x17 and VANTAGE 5H fuel. The non LOCA analyses confirmed that the acceptance criteria for rod ejection were met. The results also showed that peak RCS pressure remains oelow 110% of-design for all events.- Margin to peak clad temperature (PCT) is obtained primarily through the use of improved LOCA evaluation models (BASH for large break LOCA analysis and - NOTRUMP for small break LOCA analysis). ~Also the Zircaloy grids and IFM grids used in VANTAGE SH provide additional PCT margin. .This. margin can then be. used to implement the changes described in Section 5.0.1.' . The LOCA analysis considered N and N 1 looo operation as well as transition core effects. The results for Standa_rd fuel provided the most limiting PCT. - On balance, the LOCA analysis demonstrates that the PCT acceptance criteria contained in 10CFR50.46 of 2200'F continues to be met as well as the criteria related to clad oxidation and maximum hydrogen generation. LOCA related analyses demonstrate that the margin of safety with respect to blowdown reactor vessel and loop forces is preserved thus satisfying the 10CFR50.46 criteria that the core remain' amendable to cooling after a LOCA. Long term cooling and post LOCA subcriticality concerns are satisfied by increasing the RWST boron concentration. This in turn affects the concern with boron precipitation. The analysis shows that hot leg switchover must be accomplished 9 hours after-accident initiation, which is acceptable, i i 00309 11:10.080990 5 157. !
'I
- i. , t
-f Mass and energy release from LOCA and steamline break are not adversely.
affected by any of the changes discussed in Section 5.0;1, Therefore, the input to l-the containment and subcompartment analysis c'ontinues to.be valid. The ' evaluation of LOCA mass and energy release specifically excludes consideration of L diesel generator start time delay, degraded Si flow and degraded charging flow. [ L in summary, performance of analyses and evaluations for the upgrade to VANTAGE ) 5H and associated changes have confirmed that.the operating envelope defined by - the technical specifications continues to be bounded by the revised analytical basisc which in no case exceeds the acceptance limits. Therefore, the rnargin of safety provided by the analyses in accordance with these acceptance limits.is maintained- , and not reduced. 4
.i 5.9.5 Conclusion-a It is concluded from this evaluation that the changes addressed in Sections 5.9.1.1'. "
through 5.9.1.3 do not constitute an unreviewed safety question as defined in 10CFR50.59l An amendment to the operating license will be required in the form of changes to the Technica! Specifications. The changes addressed in Section.5.9.1.4 are specifically; ; excluded from consideration since they were~ addressed only from the' standpoint of their j i impact on LOCA and non LOCA safety analysis. e1 - In 11 9i l D030911:10.080990 5 158 lL i o I
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- (9 NORNAL FEEDWATER WITH OFFSITE- ,s Q POWER (FOUR LOOP OPERAT!010) )
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. UNIT 3-FItuRE 5.1.411 l PRESSURIZER PRESSURE AfC WATER -'
'V0LINE TRANSIDfTS FOR' LOSS OF NORMAL FEEDWATER WITH LOSS OF 0FF
( SITE. POWER :(Folm LOOP OPERAT19 ( , J
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^ b
' -FIGURE 5.1.4 13l i
MAIN FEEDLINE RUPTURE'WITH 0FF5ITE POWER AVAILABLE -
.(F0UR LOOP OPERATION)'
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- FleURE 5.1.4 13A.
MAIN FEEDLINE' RUPTURE WITH 0FFSITE POWER AVAILASLE j l (THREE LOOP OPERAT10N) R
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. UNIT 3-
- t. . FIGURE 5.1.4-14.
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- UNIT 3 -
FltURE 5.1.4-14A NAIN FEEDLINE RUPTURE WITH >
=1 10FFSITE POWER AVAILABLE- l
=(THREE LOOP OPERATION);'
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FIGURE 5.1.4-15
.NAIN FEEDLINE RUPTURE WITH 0FFSITE POWER AVAILABLEi (FOUR LOOP OPERATION)
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j
't UNIT . 3 =.
FIGURE 5.1'.4-15A E 1
- MAIN FEEDLINE RUPTURE WITH .
' 1 OFFSITE POWER AVAILABLE:
(THREE LOOP OPERATION); ! o I I g 4 .
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=OFFSITE POWER AVAILABLE
'(FOR LOOP'OPERAT1000)
I
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$ 620.
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A 0 13 10 I :102' ).10,I. 't o d-TIME.(SEC) i 1
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f n -r , NILLSTOIE 1RICLEAR POWER STAT 10ll-
'tallt 3
-FIS ME 5;1.4 16A-
' MAIN FIGLINE !MIPflRE WITH :
,0FFSITE POWER AVAILASLE
-(TIREE' LOOP OPERAT10ll); <
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. MILL 8701E leNUAR POWER STATICII
- 18117. 3 '- ;
FIGRE 5.1.4 ' ItAIN FEIDL111E fluPTURE WITH - 0FFSITE POWER AVAILASLE
-(FOlm LOOP 0PERATICII) -
vs n * . u --
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. TIME (SEC)-
l l l .a ;
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- MILLSTOIE lelCLEAR POWER STATION.- >
LallT 3
~
.i. FIG NE:5.1~4 17A.. o! i i ). ' NAlli FEEDLINE RUPTURE WITN ' -
, i 0FFSITE POWER'AVAILASLE,-
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- NILL 8701E IRN2. EAR POWER STATION'
-1881T 3-
- FIB 5tt .$.1.4 18 -
4 PAIN FEEDL!llt. RUPTURE WITH 1 -
, 0FFSITE POWER AVAILA8LE-(FOUR LOOP OPERAT1000)c 8
' J
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- l. OFFSITE POWER AVAILASLE'
.(TIREE LOOP OPERATION); )
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$ 2.5 2.25 E' :
? t 1.75 ; I
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.!.25 ,
1. 0 10 20 30 40: 50 .l 6 0 ' 70 E 80 .90 100 T!NE ISEC) ; 9 h 5 1 NILLSTONE NUCLEAR-POWER'. STATION UNIT 3 FIGURE 5.1.4 19'
-NAIN,FEEDLINE ItUPTURE WITH 0FFSITE POWER AVAILASLE.
-(FOUR' LOOP OPERATION)- '
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- FIGURE 5.1.4-19A~
MAIN FEEDLINE RDPTURE WITH
< 0FF31TE POWER AVAILASLE>.
1 (TMtEE LOOP;0PERATION)'
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. MAIN FEEDLINE RUPTURE WITHOUT OFFSITE POWER (F00RLOOPOPERATION) _
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- UNIT 3 :
! FIGURE'5.1'.4-20A y
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.1 MILLSTONE NUCLEAR POWER' STATION lL UNIT.3
-l
, FIGURE 5.1.4 21. i !
. MAIN FEEDLINE RUPTURE ~WITHOUT-0FFSITE POWER .
,(FOUR LOOP OPERATION)
.]
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. UNIT ' 3 ::
l FIGURE 5.1.'4 21A' MAIN FEEDLINE RUPTURE WITHOUT' .t
-OFFSITE POWER- .
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. UNIT 3 L FIGURE 5.1.4a22A MAIN FEEDLINE RUPTURE WITHOUT OFFSITE' POWER (THREE-LOOPOPERATION).
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UNIT 3 :
- FIGURE 5.1.4-22: .l MAIN FEEDLINE RUPTURE WITHOUT
-0FFSITE POWER -(
- (F0UR'LOOPOPERATION)-
il i t 4
< r s -- - , r ~ -,d.,,s. ,y y , % v' : - - -u
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TINE- 'tTECI' , c
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ir MILLSTONE NUCLEAR POWER STAT!0NL UNIT 3-s* FIGURE 5.1.4 23-p . i
\j -
' MAIN FEEDLINE RUPTURE W!THOUT 0FFSITE POWER = -
(FOUR: LOOP OPERATION)- w , y =c - -
,e-e y ,e+ * = Lm e. t.,s- + - = .--= w+m. .a --n-s - - - - - - - - - - - - - - - - - - - - - - - -
, 1 j
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+
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+
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.h 1:
- MILLSTONE NUCLEAR POWER STATION' Sc UNIT 3-
't FIGURE 5'.l.4 23A MAIN FEEDLINE RUPTURE WITHOUT-'
'0FFSITE1 POWER'.
(THREE-' LOOP OPERAT10N) s
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55)..- w 440.+- g HOT
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E S00. - g COLD - 560. ' 540. -
?
a 120. : i'_ , l , i' i-i3. - 10 3 10' 102_ 10 3- 10 4' T!nE' ~tSEC). s .T)
. MILLSTONE NUCLEAA POWER STATION'
. UNIT 3: '
. 1 FIGURE ~5.1.4 24 MAIN FEEDLINE RUPTURE WITHOUT-
.- 0FFSITE POWER
.(FOUR LOOP.0PERATION)
)
9 ,( '
O a 700. -- 340. " i I"I st0. * - C neo. - 4 g
. 620. HOT 600. "
a a f 580. - l 660. COLO w 540. " W g 520. " Sco. 100 10 1 108 10 3 10' ! i t tit (SEcl ! i
-l i
MILL $ TONE NUCLDR POWER STATION - ! UNIT 3 I FIGURE l.1.4 24A
]
MAINFtt0LINERUPTURENITHOUT l 0Fr$1TE POWER- - !
'(THRtt LOOP OPERATION)- I 9
9 8 4 g e
l E i i i i I I 1 s l i t i i i t
;t03. !
I
, t j 3 1400. '
INTACT LOOPS
,I !
1 l i d
/ .1 1100. -
N \
-c .
Ci 1000. -
'l 2" r r
h too.
\
su i i W too. l l l d I l E 400. l l d 2ec. FAULTED I,00P ( l l
- 0. i 100 gol not go3 go4 '{
f ! rit ($tC1 ;
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i i t l r e i i- : l i 1 l NILLSTONE NUCLEAR POWER STATION ; I UNIT 3 i J FIGURE 5.1.4 25
/ NAIN FEEDLINE RUPTURE WITHOUT j l I . OFFSITE POWER 1 (FOUR LOOP OPERATION) l j
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, (THRtt LOOP OPERATION)- '
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i F t i r MILLSTONE NUCLEAR POWER STATION UNIT 3 }
- FIGURE 5.1.4 26 e 3
MAIN FEEDLINE RUPTURE WITHOUT ' l (s CFFSITE POWER (FOUR LOOP OPERATION) I l N l T 1
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l l c 2.7% d 5 2. 5 < l 2.25 o 2. 1.75 - l : Ie6 1.as g' l. 0 13 20 30 40 50' 60 70' 00 90 100 T!NE 11Ecl i l . > i 6 MILLSTONE NUCLEAR POWER STATION-UNIT 3 FIGURE 5.1.4 26A -l MAIN FEEDLINE RUPTURE WITHOUT OFF$1TE POWER
' (THREE LOOP OPERATION) 1 1
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