ML20116N858

Virginia Power North Anna Units 1 & 2 NSSS Uprating - 2,910 Mwt Safety Evaluation

ML20116N858
Person / Time
Site:	North Anna
Issue date:	12/15/1984
From:	WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
To:
Shared Package
ML20116N845	List: ML20116N841 ML20116N858 ML20116N864 ML20116N875
References
NUDOCS 8505070379
Download: ML20116N858 (215)

Text

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ENCLOSURE 1 VIRGINIA POWER NORTH ANNA UNITS 1 AND 2 NSSS UPRATING -

2910 MWT SAFTETY EVALUATION i

l WESTINGHOUSE ELECTRIC CORPORATION l

l DECEMBER 15, 19S4 8505070379 850502 l PDR ADOCK 05000338 l

P PDR Q.

VIRGINIA POWER NORTH ANNA UNITS 1 AND 2 NSSS UPRATING - 2910 MWT SAFETY EVALUATION TABLE OF CONTENTS PAGE OBJECTIVES xi CONCLUSIONS xit

1.0 INTRODUCTION

1 -1 2.0 COMPARISON OF PARAMETERS 2-1 3.0 ACCIDENT ANALYSES 3-1 3.1 NON-LOCA EVENTS 3-1 3.

1.1 INTRODUCTION

3-1 3.1.2 TRANSIENTS NOT REANALYZED 3-2 3.1.3 NON-LOCA ACCIDENT REANALYSIS 3-6 3.1. 3.1 GENERAL 3-6 3.1.3.2 COMPUTER CODES UTILIZED 3-7 3.1. 3. 3 REANALYZED ACCIDENTS DESCRIPTION 3-9 3.1. 3. 3.1 UNCONTROLLED R00 CLUSTER CONTROL ASSEMBLY BANK 3-9 WITH0RAWAL FROM A SU8 CRITICAL CONDITION 3.1. 3. 3. 2 UNCONTROLLED ROD CLUSTER CONTROL ASSEMBLY BANK 3-11 WITHORAWAL AT POWER

3.1. 3. 3. 3 R00 CLUSTER CONTROL ASSEMBLY MISOPERATION 3-13 1

3.1. 3. 3. 4 PARTIAL LOSS OF FORCEO REACTOR COOLANT FLOW 3-17 l

3.1. 3. 3. 5 STARTUP 0F AN INACTIVE REACTOR COOLANT LOOP 3-18 l

3.1.3.3.6 LOSS OF EXTERNAL ELECTRICAL LOAD AND/0R TURBINE 3-21 TRIP 3.1. 3. 3. 7 EXCESSIVE HEAT REMOVAL DUE TO FECCWATER SYSTEM 3-23 MALFUNCTION l

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4 VIRGINIA POWER NORTH ANNA UNITS 1 AND 2

' NSSS UPRATING - 2910 MWT l SAFETY EVALUATION i

TABLE OF CONTENTS (Continued) i i

PAGE 3.1.3.3.8 EXCESSIVE LOAD INCREASE INCIDENT 3-24

! 3.1. 3. 3. 9 ACCIDENTAL DEPRESSURIZATION OF THE REACTOR 3-27

COOLANT SYSTEM 3.1.3.3.10 SPURIOUS OPERATION OF TNE SAFETY INJECTION SYSTEN 3-28 AT POWER 3.1.3.3.11 COMPLETE LOSS OF FORCEO REACTOR COOLANT FLOW 3-30 l

t 3.1.3.3.12 SINGLE REACTOR COOLANT PUMP LOCKED ROTOR 3-31

$ 3.1.3.3.13 RUPTURE 0F A CONTROL R00 DRIVE MECHANISM HOUSING 3-33 i-(

(R00 CLUSTER CONTROL ASSEMBLY EJECTION) e l 3.1'.3.3.14 SINGLE R00 CLUSTER CONTROL ASSEMBLY WITHORAWAL AT 3-34

, FULL POWER 3.2 LOCA EVENTS 3-44 3.2.1 LARGE BREAK LOSS-OF-COOLANT ACCIDENT 3-44 I

3.2.

1.1 INTRODUCTION

3-44 i 3.2.1.2 ACCIDENT DESCRIPTION 3-45

3. 2.1. 3 ANALYSIS 3-46 3.2.1.4 RESULTS 3-48 l

i 3.2.

1.5 CONCLUSION

S 3-51

,' 3.2.2 SMALL BREAK LOSS-OF-COOLANT ACCIDENT 3-53 3.2.

2.1 INTRODUCTION

3-53 3.2.2.2 ACCIDENT DESCRIPTION 3-53 3.2.2.3 METHOD OF ANALYSIS 3-54 3.2.2.4 RESULTS 3-55 3-58 3.2.

2.5 CONCLUSION

S 3.2.2.6 POST TNI EVALUATION 3-58 2156e:1d/012385 ii

VIRGINIA POWER NORTH ANNA UNITS 1 AND 2 NSSS UPRATING - 2910 MWT SAFETY EVALUATION TABLE OF CONTENTS (Continued) l 1

PAGE 4.0 NS$$ SYSTEMS REVIEW 4-1 4.1 8 ASIS OF EVALUATION 4-1 4.2 SYSTEMS EVALUATION 4-1 4.2.1 REACTOR COOLANT SYSTEM 4-1 4.2.2 RESIDUAL HEAT REMOVAL SYSTEM 4-3 4.2.3 CHEMICAL AND VOLUME CONTROL SYSTEM 4-4 4.2.4 EMERGENCY AND VOLUME CONTROL SYSTEM 4-4 4.

2.5 CONCLUSION

S 4-4 5.0 NSSS COMPONENTS IMPACT 5-1

! 5.1 BASIS FOR EVALUATION 5-1 3.2 EQUIPMENT REVIEWS 5-1 5.2.1 REACTCR VESSEL 5-1 5.2.2 REACTOR INTERNALS 5-3 i

5.2.3 CONTROL R00 ORIVE MECHANISMS 5-4 5.2.4 REACTOR COOLANT PUMPS 5-4 5.2.5 STEAM GENERATORS 5-4 l

i 5.2.6 PRESSURIZER 5-4

( 5-5 5.2.7 REACTOR COOLANT PIPING 5.2.8 LOOP STOP VALVES 5-5 5.2.9 REACTOR COOLANT SYSTEM SUPPORTS 5-5 5.2.10 AUXILIARY SYSTEMS COMPONENTS 5-5

5.3 CONCLUSION

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VIRGINIA POWER NORTH ANNA UNITS 1 AND 2 NSSS UPRATING - 2910 MWT SAFETY EVALUATION TABLE OF CONTENTS (Continued)

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PAGE 6.0 NSSS/80P INTERFACES 6-1

6.1 INTRODUCTION

6-1 6.2 MASS AND ENERGY RELEASE DATA 6-1 6.3 AUXILIARY FEEDWATER SYSTEM 6-1 6.4 RADIATION SOURCE TERMS 6-1 6.5 REACTOR COOLANT SYSTEM PIPING DESIGN DATA 6-2 6.6 COMPONENT COOLING SYSTEM HEAT LOADS 6-2 i 6.7 STEAM SYSTEM DESIGN TRANIENTS 6-3 6.8 TURBINE GENERATOR 6-3 1

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VIRGINIA POWER NORTH ANNA UNITS 1 AND 2 i NSSS UPRATING - 2910 MWT SAFETY EVALUATION LIST OF TABLES TABLE TITLE PAGE 2-1 COMPARISON OF PEACTOR COOLANT SYSTEM PARAMETERS 2-4 3.1 -1

SUMMARY

OF INITIAL CONDITIONS AND COMPUTER CODES 3-39 3.1 -2 LOCKED ROTOR RESULTS 3-42 3.1 -3 PARAMETERS USED IN AND RESULTS OF THE R00 CLUSTER 3-43 CONTROL ASSEMBLY INCIDENT 3.2-1 INITIAL CORE CONDITIONS ASSUMED FOR THE DOUBLE-ENDED 3-63 COLD-LEG GUILLOTINE BREAK (DECLG) 3.2-2 CONTAINMENT DATA 3-64 3.2 TIME SEQUENCE OF EVENTS 3-65 3.2-4 RESULTS FOR DECLG 3-65 3.2-5 REFLOOD MASS AND ENERGY RELEASES DECLG (C=0.4) 3-66 3.2-6 BROKEN LOOP ACCUMULATOR FLOW T0' CONTAINMENT 3-66 1

DECLG (C=0.4) 3.2-7 ASSUMPTIONS AND RESULTS FOR SMALL BREAK 3-67 3.2-8 TIME SEQUENCE OF EVENTS FOR SMALL BREAK 3-68 4.2-1 REACTOR COOLANT SYSTEM DESIGN AND OPERATING 4-6 PARAMETERS 4.2-2 DESIGN 8ASES FOR RESIDUAL HEAT REMOVAL SYSTEM 4-7 OPERATION 4.2-3 CHEMICAL AND VOLUME CONTROL SYSTEM HEAT EXCHANGER 4-8 PARAMETERS FOR 2910 MWT l

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VIRGINIA POWER

NORTH ANNA UNITS 1 AND 2 NSSS UPRATING - 2910 MWT SAFETY EVALUATION LIST OF F16URES FIGURE TITLE 2-1 COMPARISON OF REACTOR COOLANT SYSTEM PARAMETERS 3.1 -1 UNCONTROLLED ROD WITH0RAWAL FROM A SU8 CRITICAL CON 0! TION, NEUTRON FLUX VERSUS TIME i 3.1 -2 UNCONTROLLEO R00 WITHORAWAL FROM A SU8 CRITICAL CONDITION, CORE HEAT FLUX VERSUS TIME l

, 3.1 -3 UNCONTROLLED R00 WITHDRAWAL FROM A SUSCRITICAL CONDITION, TEMPERATURE VERSUS TIME 1

3.1 -4 MINIMUM DN8R VERSUS REACTIVITY INSERTION RATES AT 100% POWER MAXIMUM AND MINIMUM FEEDBACK 3.1-5 MINIMUM DNBR VERSUS REACTIVITY INSERTION RATES AT 60% POWER MAXIMUM AND MINIMUM FEEDBACK 3.1 -6 MINIMUM ONBR VERSUS REACTIVITY INSERTION RATES AT 10% POWER MAXIMUM AND MINIMUM FEEDBACK 3.1 -7 PARTIAL' LOSS OF FLOW, FLOW COASTDOWN VERSUS TIME,

ALL LCOPS OPERATING, ONE PUMP COASTING 00WN 3.1-8 PARTIAL LOSS OF FLOW, FLUX TRANSIENTS VERSUS TIME, ALL LOOPS OPERATING, ONE PUMP COASTING DOWN 3.1 -9 PARTIAL LOSS OF FLOW, ONBR VERSUS TIME, ALL LOOPS OPERATING, ONE PUMP COASTING DOWN 3.1-10 STARTUP 0F AN INACTIVE REACTOR COOLANT LOOP, LOOP STOP VALVES INITIALLY OPEN 3.1-11 LOSS OF LOAD ACCIDENT WITH PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, BEGINNING 0F LIFE 3.1-12 LOSS OF LOAD ACCIDENT WITH PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, BEGINNING 0F LIFE 3.1-13 LOSS OF LOAD, ACCIDENT WITHOUT PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, 8EGINNING OF LIFE

, 3.1-14 LOSS OF LOAD ACCIDENT WITHOUT PRESSURIZER SPRAY l AND POWER OPERATED RELIEF VALVE, BEGINNING OF LIFE 2156e:1d/012385 ,

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VIRGINIA POWER NORTH ANNA UNITS 1 AND 2 i NSSS UPRATING - 2910 WT SAFETY EVALUATION LIST OF FIGURES (Continued) 1 FIGURE TITLE 3.1-15 LOSS OF LOAD ACCIDENT WITH PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, END OF LIFE -

3.1-16 LOSS OF LOAD ACCIDENT WITH PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, END OF LIFE 3.1-17 LOSS OF LOAD ACCIDENT WITHOUT PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, END OF LIFE 3.1-18 LOSS OF LOAD ACCIDENT WITHOUT PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, END OF LIFE 3.1-19 FEE 0 WATER SYSTEM MALFUNCTION, NUCLEAR POWER VERSUS TIME 3.1-20 FEE 0 WATER SYSTEM MALFUNCTION, PRESSURIZER PRESSURE VERSUS TIME 3.1-21 FEE 0 WATER SYSTEM MALFUNCTION, LOOP DELTA-T VERSUS TIME 3.1-22 FEE 0 WATER SYSTEM MALFUNCTION, CORE TAVG VERSUS TIME 3.1-23 FEE 0 WATER SYSTEM MALFUNCTION, STEAM GENERATOR MASS

, (FAULTED LOOP) VERSUS TIME 3.1-24 FEE 0 WATER SYSTEM MALFUNCTION, STEAM GENERATOR MASS l (UNFAULTED LOOP) VERSUS TIME

3.1-25 FEE 0 WATER SYSTEM MALFUNCTION, DN8R VERSUS TIME 3.1-26 EXCESSIVE LOAD INCREASE WITH MANUAL R00 CONTROL, BEGINNING 0F LIFE 3.1-27 EXCESSIVE LOAD INCREASE WITH MANUAL R00 CONTROL, SEGINNING OF LIFE 3.1-28 EXCESSIVE LOAD INCREASE WITH MANUAL R00 CONTROL, l

EN0 0F LIFE 3.1-29 EXCESSIVE LOAD INCREASE WITH MANUAL R00 CONTROL, END OF LIFE 2156e:1d/012385 vii

l VIRGINIA POWER NORTH ANNA UNITS 1 AND 2 NSSS UPRATING - 2910 MWT SAFETY EVALUATION LIST OF FIGURES (Continued) l FIGURE TITLE ,

3.1-30 EXCESSIVE LOAD INCREASE WITH AUTOMATIC R00 CONTROL, BEGINNING OF LIFE 3.1-31 EXCESSIVE LOAD INCREASE WITH AUTOMATIC R00 CONTROL, BEGINNING 0F LIFE 3.1-32 EXCESSIVE LOAD INCREASE WITH AUTOMATIC R00 CONTROL, END OF LIFE 3.1-33 EXCESSIVE LOAD INCREASE WITH AUTOMATIC R00 CONTROL, END OF LIFE 3.1-34 ACCIDENTAL DEPRESSURIZATION, FLUX TRANSIENT 3.1-35 ACCIDENTAL DEPRESSURIZATION, PRESSURIZER PRESSURE TRANSIENT 3.1-36 ACCIDENTAL DEPRESSURIZATION DNBR TRANSIENT 3.1-37 SPURIOUS ACTUATION OF SAFETY INJECTION SYSTEM AT POWER 3.1-38 SPURIOUS ACTUATION OF. SAFETY INJECTION SYSTEM AT POWER 3.1-39 COMPLETE LOSS OF FLOW, CORE FLOW VERSUS TIME i 3.1-40 COMPLETE LOSS OF FLOW, FLUX TRAtlSIENTS VERSUS TIME 3.1-41 COMPLETE LOSS OF FLOW, DNBR VERSUS TIME 3.1-42 ONE LOCKE0 ROTOR, ALL LOOPS OPERATING, PRESSURE i

VERSUS TIME 3.1-43 ONE LOCKED ROTOR, ALL LOOPS OPERATING, CORE FLOW VERSUS TIME 3.1-44 ONE LOCKED ROTOR, ALL LOOPS OPERATING, FLUX TRANSIENTS 2156e:1d/012385 vill l*

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t VIRGINIA POWER NORTH ANNA UNITS 1 AND 2 f NSSS UPRATING - 2910 miT 1 SAFETY EVALUATION I

LIST OF FIGURES (Continued) l b FIGURE TITLE 3.1-45 ONE LOCKED ROTOR, ALL LOOPS OPERATING, CLAD

TEMPERATURE VERSUS TIME

, 3.2-1 PEAKING FACTOR VERSUS CORE HEIGHT, Fg = 2.15 l

, 3.2-2 MASS VELOCITY VERSUS TIME - DECLG 3.2-3 HEAT TRANSFER COEFFICIENT VERSUS TIME - DECLG ,

i 3.2-4 CORE PRESSURE VERSUS TIME - DECLG 3.2-5 BREAK FLOW RATE VERSUS TIME - DECLG

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3.2-6 CORE PRESSURE DROP VERSUS TIME - DECLG 1 3.2-7 PEAK CLA0 TEMPERATURE TRANSIENT - DECLG 3.2-8 FLUID TEMPERATURE VERSUS TIME - DECLG 3.2-9 CORE FLOW VERSUS TIME - TOP AND BOTTOM - DECLG 3

! 3.2-10 REFLOOD TRANSIENT - CORE AND 00WNCOMER WATER LEVEL VERSUS TIME - DECLG 3.2-11 REFLOOO TRANSIENT - CORE INLET VELOCITY VERSUS TIME - DECLG I

3.2-12 ACCUMULATOR FLOW VERSUS TIME, SLOWOOWN - DECLG 3.2-13 PUMPED ECCS FLOW VERSUS TIME, REFLOOO - DECLG 3.2-14 CONTAINMENT PRESSURE VERSUS TIME - DECLG 3.2-15 CORE POWER TRANSIENT - DECLG f 3.2-16 BREAK ENERGY RELEASED TO CONTAINMENT VER$US TIME - DECLG 3.2-17 CONTAINMENT WALL HEAT TRANSFER COEFFICIENT VERSUS TIME - DECLG l

3.2-18 FLUID QUALITY VERSUS TIME - DECLG 3.2-19 SAFETY INJECTION FLOW RATE VERSUS RCS PRESSURE 2156e:1d/012385 ix

VIRGINIA POWER NORTH ANNA UNITS 1 AND 2 NSSS UPRATING - 2910 MWT SAFETY EVALUATION LIST OF FIGURES (Continued)

FIGURE TITLE 3.2-20 CORE POWER VERSUS TIME 3.2-21 CORE, POWER DISTRIBUTION 3.2-22 RCS DEPRESSURIZATION TRANSIENT - 3 INCH SREAK 3.2-23 CORE MIXTURE HEIGHT - 3 INCH SREAK

. 3.2-24 STEAM FLOW VERSUS TIME - 3 INCH BREAK 3.2-25 HOT SPOT FLUID TEMPERATURE VERSUS TIME - 3 INCH BREAK 3.2-26 R00 FILM COEFFICIENT VERSUS TIME - 3 INCH SREAK 3.2-27 CLAD TEMPERATURE TRANSIENT - 3 INCH BREAK 3.2-28 RCS DEPRESSURIZATION TRANSIENT - 4 INCH 8REAK 3.2-29 CORE MIXTURE HEIGHT VERSUS TIME - 4 INCH BREAK 3.2-30 CLAD TEMPERATURE TRANSIENT - 4 !NCH BREAK 4

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OBJECTIVES North Anna Unit Nos.1 and 2 are currently licensed to operate at a total NSSS power of 2787 MWT. This report supports the Virginia Power applica- ,

tion to the Nuclear Regulatory Commission for approval to operate both North Anna units at 2905 MWT NSSS power. A safety evaluation of NSSS designs, operations and analyses has been performed to provide the follow-ing information relevant to that application:

i 1. A description of the proposed change in the licensed power rating of the North Anna units.

2. An assessment of the impact of that change on NSSS equipment

designs, safety analyses, and systems operations.

3. A technical basis for establishing that the proposed increase

in power rating does not involve an unreviewed safety question in accordance with requirements of 10 CFR 50.59.

This report summarizes the results of the safety evaluation performed by Westinghouse, and presents the conclusions based upon it.

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CONCLUSIONS The proposed increase in the licensed power rating of North Anna Units 1 and 2 has been reviewed in detail with respect to its impact on the following aspects of NSSS design and operation: ,

1. The consequences of accidents postulated in the FSAR. ,

2. The capability of systems and equipment to meet design bases specified in the FSAR.

3. The capability of equipment to maintain structural integrity under conditions defined in the FSAR.

4. Definition of NSSS/80P safety related interfaces.

! 5. Operating limits and conditions contained in Technical Specifications

that are impacted by the power rating increase.

' This review has demonstrated that North Anna Units 1 and 2 are capable, in their present design configuration, of operating at the proposed power rating

without violating any of the design criteria or safety limits specified in the

• FSAR for NSSS systems and equipment, providing that the plant is operated in accordance with the Technical Specification changes proposed by Westinghouse.

,' The review has verified that: i

1. The probability of a malfunction of NSSS equipment important to safety i previously evaluated in the FSAR will not be increased at the proposed power rating.

2. The consequences of a malfunction of NSSS equipment important to safety previously evaluated in the FSAR will not be increased at the

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proposed power rating.

) 3. The possibility of a malfunction of NSSS equipment important to safety difforent from any already evaluated in the FSAR is not created by operation at the proposed power rating.

2156e:1d/112004 x11

4. The margin of safety as defined in the bases to any technical specification will not be reduced by operation at the proposed power rating.

Therefore, it has been concluded that operation of North Anna Units 1 and 2 at the increased power rating does not reduce the NSSS safety margins, and does not involve an unreviewed question as defined by 10 CFR 50.59.

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4 SECTION 1 INTRODUCTION Virginia Po'er w is engaged in a program to maximize the electrical output of North Anna Units 1 and 2. The current phase of the program is directed toward i gaining approval from the USNRC to operate the plants at a slightly increased power level. At present, the North Anna units are licensed to operate at an NSSS power rating of 2787 MWt. Virginia Power is applying for an amendment to the operating license that will permit operation of the two units at 2905 MWt, I an increase of 4.55.

As a part of the program to uprate the North Anna Units, Virginia Power authorized Westinghouse to perform a review NSSS systems and equipment designs to verify their capability to meet requirements for operation at 2g05 MWt.

That review was conducted in accordance with groundrules and criteria put j forth in Westinghouse topical report WCAP-10263, A Review Plan for Uprating i the Licensed Power of a Pressurized Water Reactor Power Plant (Reference 1 -1 ) . A sunenary of the major guidelines followed in the NSS$ design review

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follow:

1. Scope of Review f The review encompassed all aspects of the North Anna NSSS design and operation that were impacted by the power increase.

f f 2. Safety Review Acceptance Criteria i

j i NS$$ designs have been reviewed to verify compliance at the increased power rating with licensing criteria and standards currently required

! by the North Anna operating license. In addition, a review has been f

made as defined in 10 CFR 50.5g to identify any potential unreviewed i safety question that might occur as a result of the increased power I rating.

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3. Structural Review Acceptance Criteria The structural design of NSSS equipment was reviewed to assure that compliance has been maintained at the increased power rating with irdustry codes and standards that applied when the equipment was originally built.

4. Functional Capability

' A review has been made to verify that NSSS components and systems will continue to meet functional requirements specified in the FSAR at the increased power rating.

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5. Analytical Techniques Current NRC approved analytical techniques have been used for analyses performed at the increased power rating.

6. Balance of Plant Interfaces i

j Information provided by Westinghouse to other design groups has been reviewed and revised when impacted by the increase in. power rating.

l Although Virginia Power is applying for a license amendment to operate the North Anna units at 2905 MWt, many of the North Anna UFSAR analyses and evaluations have already been perfomed at the engineered safeguards design rating of 2910 MWt. To maintain a consistent basis between information l

i reported here and that reported by reference to the UFSAR, this evaluation of I NSSS capability has also been performed at 2910 MWt. When the term "uprated power" is used in this report, it should be understood to mean 2910 MWt. ,

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1.0 REFERENCES

1 -1. R. H. McFetridge, R. T. Marchese, and R. H. Faas, A Review Plan for Ucratina the Licensed Power of a Pressurized Water Reactor Power 11431. WCAP-10263, January 1983.

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SECTION 2 ,,

COMPARISON OF PARAMETERS At the present time, North Anna Units 1 and 2 are licensed to operate at an NSSS power rating of 2787 MWt. By virtue of a license amendment recently granted by'the USNRC, the two units operate at a core average temperature of 587.5'F. The current amendment application requests approval to operate both units at a power rating of 2905 MWt with a reactor average temperature of 586.8'F. The calculated steam pressure for both units at these conditions is 850 psia. It is estimated that the electrical power output of each unit will be increased by 32 MWe at the uprated power conditions.

Table 2-1 contains a summary of Reactor Coolant System design parameters for current operating conditions of the two units, as well as parameters calculated for both units at the increased power rating. A comparison of the two sets of parameters follows:

1. NSSS Thermal Power The requested license amendment would raise the NSSS thermal power from its current level of 2787 MWt to 2905 MWt, an increase of about 4.55. As explained in Section 1, however, this NSSS evaluation has been performed at a slightly conservative power level of 2910 MWt.

This has been done to maintain consistency between information presented in this report and that already reported in the UFSAR st the engineered safeguards design rating of 2910 MWt.

2. Reactor Flow thd Tube Plugging Based on calorimetric data from North Anna Units 1 and 2, the measured reactor inlet flow rate is 302.100 gpm with 2.8% of the steam generator tubes plugged. If the steam generator tube plugging level were increased from the present level of 2.8% to a level of 7% as assumed for the uprating safety evaluation, the measured flow would

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decrease by just slightly more than 15. As shown in Table 2-1, a reactor flow of 278,400 gpm was assumed for thermal and hydraulic analyses at the uprated power. Even after allowances are made for uncertainties in the calorimetric data used to determine the 302,100 gpm measured reactor inlet flow, and additional allowances are made for the flow reduction caused by the increased steam generator tube plugging that has been assumed, 278,400 gpm is still a conservatively low flow for thermal and hydraulic analyses at the uprated power. ,

3. Reactor Coolant Temperatures i

Reactor coolant temperatures for the 2910 MWt power rating do not dif fer significantly from those for the currvnt 2787 MWt power level. However, the higher power level is reflected by a slightly greater iewperature l rise in the coolant as it passes through the reactor vessel.

4. Steam Pressure and Temperature Operation at 2910 MWt requires an increase in steam generator heat .

transfer rate, which is obtained by increasing the temperature differ-ence between reactor coolant and secondary plant steam. Since reactor coolant temperatures for 2910 MWt operation are nearly the same as those for 2787 MWt operation, it follows that the higher power rating will be obtained at a lower steam temperature. Steam pressure (saturation pres-sure at the steam temperature) will be correspondingly lower.

5. Steam Flow Steam flow at the 2910 MWt conditions has increased over the 2787 MWt conditions roughly in proportion to the thernal power increase. i i

Comparison of the parameters given in Table 2-1 shows that operating conditions l proposed for future 2910 MWt NS$$ operation are not appreciably different from i l

those at which the two North Anna units are currently operating. i l

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i Figure 2-1 is a graphical comparison of the reactor vessel cold leg, hot leg, and. vessel average temperatures as a function of power for both the current 2787 Wt operating conditions and the proposed 2910 MWt operating conditions.

This figure shows that the Reactor Coolant System temperatures for 2910 MWt operation do not differ significantly from those for the presently licensed

' 2787 MWt operation throughout the power range.

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l TABLE 2-1 COMPARISON OF REACTOR COOLANT SYSTEM PARAMETERS l l

Design Conditions Current Uprated License Power NSSS Power, MWt 2787 2910 Reactor Power, MWt 2775 2898 Reactor Coolant Pump Heat, MWt 12 12 Reactor Flow, Total, gpm 285,000 278,400 Reactor Flow, Total, million 1bm/hr 106.3 104.3 Reactor Coolant Pressure, psia 2250 2250 Reactor Coolant Temperature, *F 4 Core Outlet 622.8 624.0 Vessel Outlet 620.1 621.2 C;re Average 591.1 590.4 Vessel Average 587.8 586.8 Vessel / Core Inlet 555.5 552.3 Steam Generator Outlet 555.2 552.0 Steam Generator Steam Temperature 'F 532.0 525.2 Steam Pressure, psia 900 850 Steam Flow. Total, million Ibm /hr 12.2 12.78 )

Zero load Temperature. *F 547 547 l

> Tube Plugging, percent 5 7 Core Bypass, percent 4.5 4.5 Fuel Design- 17x17 std. 17x17 std.

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FIGURE 2-1 NORTH ANNA ! AND !!

REACTOR COOLANT TEMPERATURES

- VS -

1 0F RATED THERMAL LOAD REACTOR VESSEL 620 - f HOT LEG TEF.?IRATURE

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610 - f -

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6CO - /

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/

SEO - /

f f REACTOR VESSEL

/ / AVERAGE TENFERATURE

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• S e. .e- / /

w / /

E / /

/ l f ,

5 ,'

$ E70 - /

560 -

REACTOR VESSEL

" COLD LEG TEMPERATt;RE i 5!0 -

Preccsed I

Current 530 . . . . .

0 20 40 60 80 100

% RATED THERP.AL LCAO 1

SECTION 3 ACCIDENT ANALYSES 3.1 NON-LOCA EVENTS 1

3.

1.1 INTRODUCTION

. The following Safety Analysis evaluates the effects of the power uprating on

) non-LOCA transient events analyzed in Chapter 15 of the North Anna FSAR (Ref.

3.1 -1 ) . In conjunction with the power uprating, the effect of a +6 pcm/*F moderator temperature coefficient was incorporated. In this evaluation, transients which were determined to be sensitive to an increase in core power and a positive moderator temperature coefficient were analyzed consistent with the latest Westinghouse techniques and methods. The design DNB method employed is the Improved Thermal Design Procedure (ITOP) (Ref. 3.1-2) with the WRB-1 DNS correlation (Ref. 3.1-3). For those transients in which the ITDP method was not employed, the W-3 DNS correlation (Ref. 3.1-4, 3.1-5) was used. The results of this study demonstrate that the proposed NSSS power '

increase and positive moderator temperature coefficient can be accommodated with margin to applicable FSAR safety limits. An assessment of the impact of I these changes on the Chapter 15 FSAR transients is presented in this section of the Safety Evaluation.

i For those transients which are DNB linited, nominal values of initial conditions are assumed. The allowances on power, temperature and pressure are l determined on a statistical basis and are included in the limit DNBR, as described in WCAP-8567 (Ref. 3.1-2). This is .the Improved Thersal Design Procedure (ITOP). For transients which are not DNB listited or for wnich the i . Improved Thermal Desi 5 n Procedure is not erployed, initial conditions are l

obtained by adding unximum steady state errors to related values. The j conservative steady state errors assumed it, these analyses are censistent with those described in Chapter 15 of the North Anna FSAR.

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~.

For all accidents which were required to be re-analyzed, it was conservatively assumed that a positive moderator temperature coefficient (MTC) existed at

. full power (although it will be 10 at or above 70 percent power) except where '

it would provide a benefit. The +6 pcm/*F moderator temperature coefficient was assumed to remain constant for variations in temperature with two exceptions" The control rod ejection and rod withdrawal from subcritical analyses were based on a coefficient that was +6 pcm/*F at a zero power (no load Tavg) and which became less positive at higher temperatures. This was necessary because the TWINKLE diffusion theory computer code used in the <

analysis cannot artificially hold the MTC constant with temperature.

The trip setpoints used in the accident analysis for the 2910 MWt power uprating evaluation are the same as those used in the FSAR with the exception of the overtemperature and overpower delta T trip functions. New setpoints for these trip, functions were calculated based on core thermal limits for the uprated power, and are included with proposed Technical Specification changes in Section 7.

In general, analysis was based on the analytical methods, computer codes and assumptions-employed in the FSAR and subsequent safety analyses; any exceptions are noted in the discussion of each accident. All the methods used l in this analysis have been approved by the NRC.

3.1.2 TRANSIENTS NOT REANALYZED For a number cf the Chapter 15 postulated transierts, it was not deemed necessary to perform a reanalysis. Each of these transients are discussed and  !

the justification for not reanalyzing for upratir.g. positive moderator temperature coefficient, and ITDP are provided below.

l - A. Uncontrolled Boron Dilution (FSAR Section 15.2.4) i

( The Boron 011ution Transient is analyzed to ensure that adequate time is available for the operator to terminate the inadvertent addition of unDorated

., J-2156e:1d/120484 . 3-2

t.

makeup water to the RCS before losing all shutdown margin. If a Boron dilution event should occur.while the reactor is at power in manual control, an overpower transient may result. However, the consequences of the overpower

portion of the event are bounded by those of the RCCS bank withdrawal at power j i transient (Ref. 3.1-2).- The uncontrolled Boron dilution during refueling, hot j and cold shutdown were also not analyzed for the uprating since they are

. initiated at zero power conditions (which have not changed).

l -

8. Loss of Normal Feedwater, Loss of Offsite Power (FSAR Sections 15.2.8 15.2.9)

The loss of normal feedwater and loss of offsite power accidents are analyzed to determine the ability of the secondary system to remove decay heat. These events are not sensitive to a positive moderator temperature coefficient since the reactor trip occurs at the beginning of the transient before the reactor coolant system temperature increases significantly. Also, the loss of normal feedwater accident (which may be initiated by a loss of offsite power) i analysis currently in the FSAR assumed the power to be at 102 percent of the safety features design rating (2910 MWt). Therefore, these events were not reanalyzed.

Accidental Depressurization of the Main Steam System (FSAR Section 15.2.13) i @)I1 The Accidental Depressurization of the Main Steam System analysis is performed to demonstrate that there will be no return to criticality after reactor trip for a steam release equivalent to the spurious opening, with failure to clnse, of the largest of any single steam dump, relief, or safety valve. The positive moderator temperature coefficient would lessen the severity of the accident since this is a cooldown event.

Should the reactor be just critical or operating at power at the time of a steam release, the reactor will be tripped by the normal overpower protection when power level reaches a trip point. Following a trip at power, the reactor coolant system contains more stored energy than at no-load, the average coolant temperature is higher than at no-load, and there is appreciable energy stored in the fuel.

i 2156e:1d/112084 21 3-3

~

Thus, the additional stored energy is removed via the cooldown caused by the steam-line break before the no-load conditions of reactor coolant system temperature and shutdown margin assumed in the analyses are reached. After the additional stored energy has been removed, the cooldown and reactivity insertions proceed in the same manner as in the analysis, which assumes f i

no-load condition at time zero. However, since the initial steam-generator l

water. inventory is greatest at no-load, the magnitude and duration of the '

reactor c'oolant system cooldown are less for steam-line breaks occurring at power. lo determine whether the reactor remains subcritical following the break, the most limiting case of the reactor initially subcritical at hot zero power with no decay heat was assumed. This is more conservative than the full power case since it results in minimum stored energy in the system, thereby resulting in most rapid cooldown and the greatest loss of shutdown margin.

The consequences of this accident are less severe than the loss of feedwater incident (Ref. 3.1-1), thus the cooling is adequate to protect the core.

Therefore this accident did not need to be reanalyzed for the uprating, j 0. Minor Secondary System Pipe Breaks (FSAR Section 15.3.2) ,

l Since the major secondary system pipe break is more limiting than a minor secondary system pipe break (Ref. 3.1-1), reanalysis of this accident was not l

performed.

E. Major Secondary System Pipe Rupture (FSAR Section 15.4.2) i

1. Rupture of Main Steam 11ne The steam release arising from a rupture of a main steam pipe would result in an initial increase in steam flow, which decreases as the steam pressure falls. 1he energy removal from the reactor coolant system causes a reduction of coolant temperature and pressure. In the presence of a negative moderator temperature coefficient, the cooldown results in a reduction of core shutdown margin. .Therefore a positive moderator temperature coefficient would lessen the severity of the accident. -

2156e:1d/1'.2084.22 3-4

^

l All the cases presented in the FSAR assume initial hot shutdown conditions at time zero, since this represents the most pessimistic initial condition.

, Should the reactor be just critical or operating at power at the time of steam-line break, the reactor will be tripped by the normal overpower protection system when power level reaches a trip point. Following a trip at power, the reactor coolant system contains more stored energy than at no-load, and there is appreciable energy stored in the fuel. Thus, the additional -

stored en'ergy is removed via the cooldown caused by the steam-load break before the no-load conditions of reactor coolant system temperature and I shutdown margin assumed in the analyses are reached. After the additional stored energy has been removed, the cooldown and reactivity insertions proceed

in the same manner as in the analysis that assumes no-load condition at time zero.

a However, since the initial steam-generator water inventory is greatest at 4

no-load, the magnitude and duration of the reactor coolant system cooldown are less for steam-line breaks occuring at power. Therefore, a reanalysis was not required.

2. Major Rupture of a Main Feedwater Pipe -

A major feedwater line rupture is defined as a break in a feedwater pipe 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.

Further, a break in this location could preclude the subsequent addit. ion of

. auxiliary feedwater to the affected steam generator. (A break upstream of the feedline check valve would affect the nuclear steam supply system only as a loss of feedwater). Since this accident assumed an initial operating condition at 102 percent of the engineered safeguards design rating (2910 MWt), reanalysis was not required.

2156e:1d/112084 23 3-5

3.1.3 NON-LOCA ACCIDENT REANALYSIS 3.1.3.1 GENERAL The reanalyzed accidents were performed using current Westinghouse methodology and computer codes. Table 3.1-1 sumarizes the initial conditions and computer codes used in the analysis.

For those transients which are DNB limited, nominal values of initial conditions are assumed. The allowances on power, temperature and pressure are determined on a statistical basis and are included in the ONBR limit, as described in WCAP-8567 (Ref. 3.1-2). This is the Improved Thermal Design Procedure (ITDP).

For transients which are not DNB limited or for which the Improved Thermal i Design Procedure is not employed, initial conditions are obtained by adding maximum steady-state errors consistent with those described in Chapter 15 of North Anna's FSAR to related values.

The following steady-state errors are considered: ,

Core Power + 2 percent calorimetric error allowance

2. Average RCS temperature + 4'F controller deadband and measurement l

error allowance

! 3. Pressuriser Pressure 3, 30 psi - steady-state fluctuations and

! measurement error allowance

4. Reactor Flow Thermal Design Flow l

Reactor Protection Setpoints and response times used are listed in FSAR l

l Chapter 15.1 and the Technical Specifications with the exception of the l overtemperature delta-T (OT&T) and overpower delta-T (0 PAT) setpoints.

The new OTAT and OPAT setpoints were calculated for the design basis based on the core thermal limits. The results are included in the proposed Technical Specification changes (Section 7).

I t

t 2156e:1d/120384 3-6

l

.3.1.3.2 COMPUTER CODES UTILIZED

-Summaries of the principal computer codes used in the transient analyses are given below.

FACTRAN I .

FACTRAN calculates the transient temperature distribution in a cross-section of a metal clad UO 2 fuel rod and the transient heat flux at the surface of the clad using as input the nuclear power and the time-dependent coolant parameters (pressure, flow, temperature, and density). The code uses a fuel model which simultaneously exhibits the following features:

1. A sufficiently large number of radial space increments to handle fast transients such as rod ejection accidents.

2. Material properties which are functions of temperature and a sophisticated fuel-to-clad gap heat transfer calculation.

3. The ne:essary calculations to handle post-dep.arture from nucleate boiling (DNB) transients: film boiling heat transfer correlations, Zircaloy-water

, reaction, and partial melting of the materials.

I FACTRAN is further discussed in Reference 3.1-6.

l LOFTRAN r

l The LOFTRAN program is used for transient response studies of a pressurized water reactor (PWR) system to specified perturbations in process parameters.

LOFTRAN simulates a multiloop system by a model containing the reactor vessel, hot and cold leg piping, steam generators (tube and shell sides), and the

! pressurizer. The pressurizer heaters, spray, relief, and safety valves are also considered in the program. Point model neutron kinetics, and reactivity effects c? the moderator, fuel, boron, and rods are included. The secondary side of the steam generator utilizes a homogeneous, saturated mixture for the I

i 2156e:1d/112004 25 3-7

thermal transients and a water level correlation for indication and control. l The reactor protection system is simulated to include reactor trips on high neutron flux, overtemperature AT, overpower AT, high and low pressure, low flow, and high pressurizer level. Control systems are also simulated

• including rod control, steam dump, feedwater control, and pressurizer pressure control. The ECCS.-including the accumulators, is also modeled.

LOFTRAN a' Iso has the capability of calculating the transient value of DNBR based on the input from the core limits. The core limits represent the minimum value of DNBR as calculated for typical or thimble cell.

LOFTRAN is further discussed in Reference 3.1-7.

TWINKLE The 1 WINKLE program is a multi-dimensional spatial neutron kinetics code, which was patterned after steady-state codes presently used fer reactor core design. 1he code uses an impitcit finite-difference method to solve the l

two-group transient neutron diffusion equations in one, two, and three dimensions. The code uses six delap a neutron groups and contains a detailed ,

multi-regicn fuel-clad-coolant heat transfer model for calculating pointwise 7._

Doppler and moderator feedback effects. The code handles up to 2000 ss,atial

! points and performs its own steady-state initialization. Aside from basic cross-section data and thermal-hydraulic parameters, the code accepts as input

' basic driving functions such as inlet temperature, pressure, flow, boron

! concentration, control rod motion, and others. Various edits are providea; /

e.g., channelwise power, axial offset, enthalpy, volumetric surge, pointwise power, and fuel temperatures.

The TWINKLE code is used to predict the kinetic behavior of a reactor for transients which cause a major perturbation in the spatial neutron flux j . distribution.

TWINKLE is further described in Reference 3.1-8.

2156e:1d/112084 26-8 1

I PH0ENIX The PH0ENIX (Ref. 3.1-10) code calculates the individual loop flows, core flow and pump speeds as a function of time subsequent to failure of any number of the geactor coolant pumps. The analysis is based on a momentum balance around each reactor coolant loop and across the reactor core. This momentum balance is combined with the continuity equation, a pump momentum balance and the pump characteristics.

THINC -

The THINC code is used to calculate the DN8R during the transient based on system conditions (pressure, temperature, and flow) calculated by LOFTRAN and hot channel heat flux calculated by FACTRAN. The THINC code is described in Reference 3.1-9.

3.1.3.3 REANALYZED ACCIDENTS DESCRIPTION

, Oetails of the reanalysis are presented in the following sections. In all cases, the applicable FSAR acceptance criteria are satisfied.

3.1.3.3.1 UNCONTROLLED ROD CLUSTER CONTROL ASSEMBLY BANK WITHDRAWAL FROM A SUBCRITICAL CONDITION 3.1.3.3.

1.1 INTRODUCTION

A rod cluster control assembly withdrawal accident is defined as an uncontrolled addition of reactivity to the reactor core caused by withdrtwal of rod cluster control assemblies resulting in a power excursion. Such a transient could be caused by a malfunction of the reactor control or control rod drive systems. This could occur with the reactor subcritical, at hot zero power, or at power. The "at power' case-is discussed in Section 3.1.3.3.2.

l

\

1 2156e:1d/112084 27 1 3-9

i.

This analysis is performed to show that in the event of a RCCA withdrawal accident from the s'uberitical or near critical condition, the core and the reactor coolant system are not adversely affected..

The RCCA withdrawal from suberitical transient has been reanalyzed at'the uprated conditions (positive MTC) to demonstrate that the power range high ,

neutron flux reactor trip (Iow setting) provides the necessary protection against f'uel damage.

3.1.3.3.1.2 METHOD 0F ANALYSIS t

The uncontrolled RCCA withdrawal from a subcritical condition accident is analyzed by use of the PHOENIX, TWINKLE, FACTRAN, and THINC digital computer codes. The PHOENIX code is used to provide flow versus time for THINC. The TWINKLE program is a neutron kinetics code which solves the multi-dimensional, two group transient diffusion equation. It replaces the point kinetics code, WIT-6, to perform a one-dimensional (axial) average core neutron kinetic analysis. FACTRAN uses the nuclear power versus time data from TWINKLE to

! calculate the transient heat flux and fuel and clad temperatures. THINC uses the heat flux versus time data from FACTRAN and the flow versus time data from PHOENIX to .41culate the transient DNBR.

The 80L bcron concentration used in TWINKLE reflects a +6 pcm/*F moderator temperature coefficient to yield a high value of peak heat flux.

All assumptions made in this analysis are consistent with those in the North Anna FSAR Chapter 15.

3.1.3.3.1.3 RESULTS Figures 3.1-1 through 3.1-3 show the transient behavior for the indicated reactivity insertion rate with the accident terminated by reactor trip at 35 percent nominal power.

l l

l l

, l l 2156e:1d/112084 28 l l

l 3-10 '

Figure 3.1-1 shows the neutron flux transient. The neutron flux overshoots the full-power nominal value, but this occurs for only a very short time.

Hence, the energy release and the fuel temperature increases are relatively small. The thermal flux response, of interest for DNB considerations, is shown on Figure 3.1-2. The beneficia'1 effect of the inherent thermal lag in i the fuel is evidenced by a peak heat flux less than the full-power nominal value. There is a large margin to DN8 during the transient, since the rod surface heat flux remains below the design value, and there is a high degree of subcooling at all times in the core. Figure 3.1-3 shows the response of the average fuel, cladding, and coolant temperature. The average fuel temperature increases to a value lower than the nominal full-power value.

3.1.3.3.

1.4 CONCLUSION

S In the event of a rod cluster control assembly withdrawal accident from the suberitical condition, the core and the reactor coolant system are not adversely affected, since the combination of thermal power and the coolant temperature results in a ONBR well above the limit value.

} 3.1.3.3.2 UNCONTROLLED R00 CLUSTER CONTROL ASSEMBLY BANK WITHDRAWAL AT POWER 3.1.3.3.

2.1 INTRODUCTION

The purpose of the rod withdrawal at power analysis is to verify the adequacy of the high neutron flux and overtemperature AT reactor trips for core l protection. To assure adequate core protection at 2910 MWt, the Reactor Core Thermal and Hydraulic Safety Limits have been recalculated consistent with the ITDP methods. Based on these new protection limits, the AT setpoint equation constants have been recalculated.

An uncontrolled rod cluster control assembly bank withdrawal at power produces  !

a mismatch in steam flow and core power, resulting in an increase in reactor

. coolant temperature.

A positive MTC would augment the power mismatch and could reduce the margin to DN8.

l 2156e:1d/112084 29 3-11

l 3.1.3.3.2.2 METH00 0F ANALYSIS The transient was analyzed employing the same digital computer codes used for the FSAR and subsequent safety analyses.

This accident is analyzed with the Improved Thermal Design Procedure as .

described in Reference 3.1-2. Initial operating conditions are assumed at values consistent with steady-state N loop operation.

Three different initial conditions, analyzed for both minimum and maximum 8

reactivity feedback conditions, are presented in Section 15.2.2 of the North Anna FSAR:

1. 100 percent power

2. 60 percent power

3. 10 percent power All of the minimum reactivity feedback cases analyzed assumed a constant +6 i pcm/*F MTC. The assumption that a positive MTC exists at full power is conservative since at full power, the moderator coefficient will actually be zero or negative.

3 .1. 3 . 3. 2 . 3 RESULTS Figure 3.1-4 shows the minimum DN8R as a function of reactivity insertion rate from initial full-power operation for the minimum and maximum reactivity feedback. It can be seen that two reactor trip channels provide protection over the whole range of reactivity insertion rates. These are the high-neutron-flux and overtemperature delta-T trip channels. The minimum DN8R is greater than the limit value.

l 1

2156e:1d/012385 32 3-12

Figures 3.1-5 and 3.1-6 show the minimum DNBR as a function of reactivity insertion rate for rod cluster control assembly withdrawal incidents starting at 60 and 10 percent power, respectively. The results are similar to the 100 percent power case, except that as the initial power is decreased, the range over which the overtemperature delta-T trip is effective is increased. In all cases the DNBR is greater than the limit value.

1 -

3.1.3.3.

2.4 CONCLUSION

S

)

The high-neutron-flux and overtemperature delta-T trip channels provide' adequate protection over the entire range of possible reactivity insertion

rates, i.e., the minimum value of DNBR is always greater than the limit value.

3.1.3.3.3 R00 CLUSTER CONTROL ASSEMBLY MISOPERATION 3.1.3.3.

3.1 INTRODUCTION

Rod cluster control assembly (RCCA) misoperation accidents include:

1. One or more dropped RCCAs within the same group,

2. -A dropped RCCA bank, 4

3. Statically misa11gned RCCA. '

Each RCCA has a position indicator channel which displays the position of the assembly. The displays of assembly positions are grouped for the operator's l

convenience. Fully inserted assemblies are further indicated by a rod at bottom signal, which actuates a local alarm and a control room annunciator.

! Group demand position is also indicated.

2156e:1d/112084 31 3-13

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

3.1.3.3.3.2 METHOD OF ANALYSIS

a. One or more dropped RCCAs from the same group.

For evaluation of the dropped RCCA event, the transient s/ stem I response is calculated using the LOFTRAN (Ref. 3.1-7) code. The code simulates the neutron kinetics, reactor coolant system, pressurizer, pressurizer relief and safety valves. The code computes pertinent plant variable's including temperatures, pressures, and power level.

The transient response, nuclear peaking factor analysis, and DN8 design basis confirmation are performed in accordance with the methodology described in Reference 3.1-11.

b. Steady state power distribution are analyzed using the computer codes as described in Table 4.1-2 in the FSAR. The peaking factors are then i

used as input to the THINC code to calculate the DN8R. -

3.1.3.3.3.3 RESULTS

a. One or more Dropped RCCAs Single or multiple dropped RCCAs within the same group result in a negative reactivity insertion which may be detected by the power range negative neutron flux rate trip circuitry. If detected, the reactor is tripped within approximately 2.5 seconds following the drop of the RCCAs. The core is not adversely affected during this period, since power is decreasing rapidly. Following reactor trip, normal shutdown procedures are followed. The operator may manually retrieve the RCCA by following approved operating procedures.

For those dropped RCCAs which do not result in a reactor trip, power may be reestablished either by reactivity feedback or control bank 3

withdrawal. Following a dropped rod event in 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 red control mode of operation as the limiting case.

2156e:1d/112084 32 3-14

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

1 For & dropped RCCA event in the automatic rod control mode, the Rod l Control System detects the drop in power and initiates control bank witiidrawal. Power overshoot may occur due to this action by the automatic rod controller after which the control system will insert the control bank to restore nominal power. Uncertainties in the

~ initial condition are included in the DN8 evaluation as described in

, Reference 3.1-11. In all cases, the minimum DN8R remains above the limit value.

b. Oropped RCCA Bank A dropped RCCA bank typically results in a reactivity insertion greater than 500 pcm which will be detected by the power range negative neutron flux rate trip circuitry. The reactor is tripped within approximately 2.5 seconds following the drop of a RCCA Bank.

The core is not adversely affected during this period, since power is decreasing rapidly. Following reactor trip, normal shutdown procedures are followed to further cool down the plant. Any action

! required of the operator to maintain the plant in a stabilized f condition will be in a time frame in excess of ten minutes following the incident.

c. Statically Misa11gned RCCA l

i i The most severe misalignment situations with respect to'ONBR at significant power levels arise from cases in which one RCCA is fully inserted, or where bank 0 is fully inserted with one RCCA fully withdrawn. h itiple indcpendent alarms, including a bank insertion  !

limit alarm, alert the operator well before the postulated ccaditions

.are approached. The bank can be inserted to its insertion limit with any one assembly fully withdrawn without the DNOR falling below the )

limit value.

• l i

2156e:1d/112084 33 3-15 l

i

The insertion limits in the Technical Specifications may vary from f time to time depending on a number of limiting criteria. It is l preferable, therefor.e. to analyze the misa11gned RCCA case at full power for a position of the control bank as deeply inserted as the criteria on minimum DNBR and power peaking factor will allow. The l full power insertion limits on control bank D must then be chosen to be above that position and will usually be dictated by other c'riteria. Detailed results will vary from cycle to cycle depending on fuel arrangements.

l For this RCCA misalignment, with bank D inserted to its full power i insertion limit and one RCCA fully withdrawn, DNBR does not fall below

! the limit value. This case is analyzed assuming the initial reactor 4 i power, pressure, and RCS temperatures are at their nominal values including uncertainties but with the increasec radial peaking factor associated with the misa11gned RCCA.

l j

DNB calculations have not been performed specifically for RCCAs ,

missing from other banks; however, power shape calculations have been done as required for the RCCA ejection analysis. Inspection of the

! power shapes shows that the DNB and peak kw/ft situation is less

! severe than the bank D case discussed above assuming insertion limits on the other banks equivalent to a bank D full-in insertion limit.

f

! For RCCA misalignments with one RCCA fully inserted, the DNBR does not

( fall below the'11mit value. This case is analyzed assuming the initial reactor power, pressure, and RCS temperatures are at their nominal values, including uncertainties but with the increased radial j peaking factor associated with the misa11gned RCCA.

DN8 does not occur for the RCCA misalignment incident and thus the ability of the primary coolant to remove heat from the fuel rod is not reduced. The peak fuel temperature corresponds to a linear heat generation rate based on the radial peaking factor penalty associated 2156e:1d/112084 34 L

3-16

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

i with the misaligned RCCA and the design axial power distribution. The t

resulting linear heat generation is well below that which would cause fuel melting.

> Following the identification of a RCCA group misalignment condition by )

the operator, the operator is required to take action as required by the plant Technical Specifications and operating instructions.

3.1.3.3.

3.4 CONCLUSION

S ,

i.

For cases of dropped RCCAs or dropped banks, for which the reactor is tripped e r to o 1 a con u nt h DN d ig ba i i met. It is shown for all cases which do not result in reactor trip that the i

DN8R remains greater than the limit value and, therefore, the DNB design basis is met.

l For all cases of any RCCA fully inserted, or bank 0 inserted to its rod insertion limits with any single RCCA in that bank fully withdrawn (static

{ misalignment), the DN8R remains greater than the limit value.

3.1.3.3.4 PARTIAL LOSS OF FORCED REACTOR COOLANT FLOW 3.1.3.3.

4.1 INTRODUCTION

A partial-loss-of-coolant-flow accident can result from a mechanical or electrical failure in a reactor coolant pump, from a fault in the power supply to the pump, or from inadvertent closure of a loop isolation valve.

• This transient was reanalyzed to' determine the effect of the uprating and a positive MTC on the nuclear power transient and the resultant effect on the i minimum DN8R. The effect of the MTC on the nuclear power transient would be limited to the initial stages of the event during which reactor coolant-temperature increases; this increase is terminated shortly after reactor trip.

b I

2156e:1d/112084 35 3-17 I

3.1.3.3.4.2 METHOD OF ANALYSIS This transient is analyzed using three digital computer codes. First LOFTRAN (Ref. 3.1-7) is used to calculate loop flow, core flow, time of reactor trip,

and the nuclear transient. The FACTRAN code (Ref. 3.1-6) is then used to calculate the heat flux transient based on the nuclear power and flows from L0FTRAN. Finally, the THINC code (Ref. 3.1-9) is used to calculate the minimum DNBR based on the heat flux from FACTRAN and flows from LOFTRAN. The WR8-1 DNS correlation is used (Ref. 3.1-3). The transient presented represents the minimum of the typical or thimble cell.

The analysis was performed using the uprated power, a constant +6 pcm/*F MTC, and ITOP as described in Reference 3.1-2. Initial operating conditions are assumed at values consistent with steady-state N loop operation.

3.1.3.3.4.3 RESULTS Figures 3.1-7 through 3.1-9 show the flow coastdown, nuclear power, heat flux, and the minimum DNBR versus time. The ONBR curve does not fall below the limit DNBR value.

3.1.3.3.

4.4 CONCLUSION

S The analysis shows that the DNBR will not decrease below the limit value at any time during the transient. Thus, there will be no cladding damage and no release of fission products to the reactor coolant system.

i 3.1.3.3.5 STARTUP 0F AN INACTIVE REACTOR COOLANT LO0p 3.1.3.3.

5.1 INTRODUCTION

The inadvertent startup of an idle reactor coolant pump with loop stop valves open results in the injection of cold water into the core, causing a rapid reactivity insertion and subsequent power increase.

i 2156e:1d/112084 36 3-18

4 In the case of the plant operated with a reactor coolant loop out of service and the loop stop valves of the loop closed, there is no flow from the reactor vessel and active loops to the inactive loop. With the stop valves in one j

loop closed, the isolated section of the loop would be cooler than the temperature of the active loops. Also, the bcron concentration in the I inactive loop may be less than that in the core and active loops.

4 This transient was reanalyzed to determine the offact of the core uprating.

The positive moderator temperature coefficient would lessen the severity of the accident since the coolant temperature decreases.

i 3.1.3.3.5.2 METHOD OF ANALYSIS This transient is analyzed using three digital computer codes. First, LOFTRAN (Ref. 3.1-7) is used to calculate loop flow, core flow, time of reactor trip, and the nuclear transient following reactor trip. The FACTRAN code (Ref. 3.1-6) is then used to calculate the heat flux transient based on the I

nuclear power and flows from LOFTRAN. Finally, the THINC code (Ref. 3.1-9) is used to calculate the minimum DNBR based on the heat flux from FACTRAN and flows from LOFTRAN. The transient presented represents the minimum of the minimum of the thimble or typical cell.

< The analysis for loop stop valves initially open was performed using the uprated power conditions. Initial conditions are assumed at end-of-life values consistent with steady-state N-1 loop operation with appropriate errors. Following the start of the idle pump, the inactive loop flow reverses and accelerates to its nominal full-flow value. In the analysis, the reactor

' trip occurs on the power range neutron flux exceeding the P-8 setpoint.

The startup of an inactive reactor coolant loop with the loop stop valves j initially closed is analyzed assuming the inactive loop to be at a boron concentration of 0 ppm while the active portion of the system is at 1200 ppe, a conservative value for the required shutdown margin at beginning of life.

2156e:1d/121184 14 .

3 19 L i

3.1.3.3.5.3 RESULTS The results following the startup of the idle pump with loop stop valves initially open are shown in Figure 3.1-10. The minimum DN8R during the transient is never less than the limit value.

Even with the assumption that administrative procedures are violated to the extent that an attempt is made to open the loop stop valves with 0 ppm boron in the inactive loop while the remaining portion of the system is at 1200 ppm I boron, the dilution of the boron in the core is slow. The initial reactivity -

i insertion rate is calculated to be less than 2.6 x 10-5 delta k/sec, which

is within the range of the reactivity insertion rates considered in Section 15.2.2 of the FSAR. For these conditions, the time required for the shutdown margin to be lost and the reactor to become critical is 16.4 min. This

! calculation takes into account the reduced reactor coolant system volume due to the isolated loop. This is ample time for the operator to recognize a high count rate signal and terminate the dilution by turning off the pump in the j inactive loop or by borating to counteract the dilution.

l The reactivity addition at end of life due to an attempt to open stop valves when the inactive loop temperature is less than the core temperature is smaller than the reactivity addition considered in the above beginning-of-life

! case.

3.1.3.3.

5.4 CONCLUSION

S' LOOP STOP VALVES Op[N The transient results show that the core is not adversely affected, i.e.,

there is considerable margin to the limit DNOR.

LOOP STOP VALVES CLOSED 1he redundant interlocks provided in the reactor protection system ensure that the temperature and boron concentration in an isolated loop are brought to equilibrium with the remainder of the system at a slow rate. Should l

2156e:Id/112084 38 l .

3-20

administrative procedures be violated and an attempt be made to open stop valves when the isolated loop temperature or boron concentration is lower than

! that in the core, the rea.tivity addition rate is slow enough to allow the 2

operator to take corrective action before shutdown margin is lost.

3.1.3.3.6 LOSS OF EXTERNAL ELECTRICAL LOAD AND/OR TURBINE TRIP 3.1.3.3.

6.1 INTRODUCTION

't The loss of external electrical load incident is analyzed as a complete and

~

instantaneous loss of steam load on the nuclear steam supply system without automatic steam dump or direct reactor trip. As such, it constitutes one of the most severe transients with respect to overpressurization of the reactor coolant system and steam generator. As shown by the FSAR, this transient does not cause a DNB ratio less than the limit. A reactor trip on high pressurizer pressure would normally terminate the transient prior to any significant reduction in DNS ratio. Also, the protection provided by the overtemperature AT trip would trip the reactor if necessary to prevent a DNB ratio less than the limit.

Two cases, analyzed for both beginning and end-of-life conditions, are presented in Section 15.2.7 of the FSAR:

1. With the operation of the pressurizer spray and the pressurizer power g operated relief valves; and

2. With no credit for pressurizer spray or power operated relief valves j As the MTC is negative at end-of-life, only the beginning-of-life cases included the +6 pcm/*F MTC. The result of a loss of load is a core power level which momentarily exceeds secondary system removal capability causing an increase in core water temperature. The consequences of the reactivity addition due to a positive MTC are increases in both peak nuclear power and pressurizer pressure.

1 2156e:1d/120384 3-21 4

, - . - , , - , . - - , , - e, , - - - -.

(

4 3.1. 3. 3. 6. 2 METHOD OF ANALYSIS This transient war 2nalyzed using the Improved Thermal Design Procedure as described in WCAP-8567 (Ref. 3.1-2). Initial operating conditions are assumed at values consistent with steady-state N loop operation. All other assumptions are consistent with FSAR and subsequent safety analyses. A constant +6 pcm/*F was assumed for beginning-of-life cases while a large (absolute value) negative MTC was assumed for the end-of-life cases.

3.1. 3. 3. 6. 3 RESULTS 4

The system transient response to a total loss of load from 102 percent power at beginning-of-life with pressure control is shown in Figures 3.1-11 and 3.1-12. Peak pressurizer pressure reaches 2522 psia following a reactor trip j on the high pressurizer pressure signal. The minimum DNBR remains well above

] the limit value.

Figures 3.1-13 and 3.1-14 illustrate reactor coolant system response to a loss of load at Beginning-of-Life (BOL) assuming no credit for pressure control.

Peak pressurizer pressure reaches 2554 psia following reactor trip on high pressurizer pressure. The DNBR increases throughout the transient.

Figures 3.1-15 and 3.1-16 show the RCS response for loss of load at End-of-Life (EOL) with pressure control. Peak pressurizer pressure reached during the transient is 2366 psia. The combination of pressurizer spray and pressurizer power operated relief valves and negative moderator temperature coefficient are sufficient to prevent reaching any reactor trip setpoints.

The DNBR increases throughout the transient and never drops below its_ initial value.

5 s L 2156e:1d/012385 42 3-22

I Figures 3.1-17 and 3.1-18 show the RCS response for loss of load at EOL with

[ no credit for pressure control. The reactor trips on a high pressurizer pressure. Maximum pressurizer pressure attained during the transient is 2538 psia which occurs shortly af ter reactor trip. The DNBR increases throughout .

the tran,sient and thus, always exceeds the limit value.

I 3.1.3.3.

6.4 CONCLUSION

S Results of the preliminary analyses show that the plant design is such that a total loss of external electrical load without a direct or immediate reactor trip presents no hazard to the integrity of the RCS or the main steam system.

Pressure relieving devices incorporated in the two systems are adequate to limit the maximum pressures to within design limits.

The integrity of the core is maintained by operation of the Reactor Protection System, i.e., the minimum DNBR is unintained above the limit value. Thus, there will be no cladding damage and no release of fission products.

l

.1 l

3.1.3.3.7 EXCESSIVE HEAT REMOVAL DUE TO FEEDWATER SYSTEM MALFUNCTION 3.1.3.3.

7.1 INTRODUCTION

l l

Reductions in feedwater temperature or additions of excessive feedwater are

Neans of increasing core power above full power.,

\

! This transient was reanalyzed to determine the effect of the uprating on the nuclear power transient and the resulting effect on the minimum DNBR. The accident scenario is somewhat different- from the one presented in the FSAR and ,

i the description is contained in the revised FEAR write-up of.this events (Section 15.2.10).

3.1.3.3.7.2 METHOD OF ANALYSIS The excessive heat removal due to a feedwater system malfunction transient was -

analyzed by use of the digital computer code LOFTRAN (Ref. 3.1-7).

l l

'2156e:1d/120384 3-23

i The FSAR case which was reanalyzed is the excessive feedwater addition due to a control system malfunction or operator error which allows a feedwater control valve to open fully. The assumption is that the reactor is in automatic control at full power with end-of-life conditions consistent with the case presented in the FSAR. The zero load case in the FSAR was not

~

reanalyzed because the core uprating will not affect this transient. The feedwater temperature reduction case in the FSAR was also not reanalyzed  !

~

because it is bounded by the excessive load increase accident.

3.1.3.3.7.3 RESULTS The transient results, Figures 3.1-19 to 3.1-25, show the nuclear power, T,,g loop AT, pressurizer pressure, steam generator water level in faulted i, loop and unfaulted loop, and DNBR associated with the increased thermal load on the reactor. When the steam generator water level in the faulted loop reaches the high-high level setpoint, all feedwater control and isolation valves, and the main feedwater pumps are tripped. This prevents continuous

addition of feedwater to the steam generator. With no incoming feedwater and l steam still being generated, the steam generator water level decreases. When

.the water level reaches the low-low setpoint in an unfaulted loop, the reactor is tripped. This will then initiate a turbine trip. Following feedwater isolation and reactor trip, the plant will approach a stabilized condition at hot standby.

3.1.3.3.

7.4 CONCLUSION

S The transient results show that DNB does not occur at any time during.the l excessive feedwater flow incident; thus, the ability of the primary coolant to l remove heat from the fuel rod is sufficient and cladding damage will not occur.

3~.1.3.3.8 EXCESSIVE LOAD INCREASE INCIDENT' I

l 3.1.3.3.

8.1 INTRODUCTION

j

-l l An excessive load increase incident is defined as a rapid increase'in the steam ficw that causes a power mismatch between.the reactor core power and the j

< steam generator load demand. The reactor control system is designed to 2156e:1d/112084 42 3-24

accommodate a 10 percent load increase or a 5 percent / min 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 the reactor protection system. .

l This accident could result from either an administrative violation, such as j 4

g excessive loading by the operator, or an equipment malfunction in the steam 4 dump control or turbine speed control.  ;

The excessive load increase transients has been reanalyzed at the uprated conditions to demonstrate that the overtemperature delta-T, overpower delta-T, and power range high neutron flux trips provide adequate protection against DNB.

3.1.3.3.8.2 METHOD 0F ANALYSIS This accident is analyzed using the LOFTRAN code. The code simulates the neutron kinetics, reactor coolant system, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables, including temperatures, pressures, and power level.

Four cases are analyzed to demonstrate the plant behavior following a 10 percent step load increase from rated load. These cases are as follows:

1. Manually controlled reactor at beginning-of-life.

l I

2. Manually controlled reactor at end-of-life.

3. Reactor in automatic control at beginning-of-life.

l 4. Reactor in automatic control at end-of-life.

I l-l l

l 2156e:1d/112084 43 3-25

At beginning-of-life, the core is assumed to have a zero moderator temperature I

coefficient of reactivity and therefore, the least inherent transient capability. At end-of-life, the moderator temperature coefficient of reactivity has its highest absolute value. This results in the largest amount  !

of reactivity feedback due to changes in coolant temperature.

All other assumptions are consistent with those in the North Anna FSAR.

3.1.3.3.8.3 RESULTS Figures 3.1-26 through 3.1-29 illustrate the transient with the reactor in the manual control mode. For the beginning-of-life case there is a slight power increase, and the average core temperature shows a small increase. The minimum DNBR remains essentially constant throughout the transient. For the end-of-life, manually controlled case, there is a larger increase in reactor power due to the moderator feedback. A reduction in DNBR is experienced, but remains well above the limit value.

Figures 3.1-30 through 3.1-33 illustrate the transient assuming the reactor is in the automatic control mode. Both the beginning-of-life and the end-of-life cases show large increases in core power. The BOL case shows a slight increase in T,,9 above initial, while the EOL case shows a slight decrease in T,yg below initial. For both BOL and EOL cases, the minimum DNBR remains well above the minimum value.

3.1.3.3.

8.4 CONCLUSION

S In all cases, the minimum DNBR during the transient is greater than the limit value. Also, equilibrium conditions of core power and core average temperature were reached in all four cases.

t 2156e:1d/121184 14 3 26 l

-3.1.3.3.9 ACCIDENTAL DEPRESSURIZATION OF THE REACTOR C0OLANT SYSTEM 3.1.3.3.

9.1 INTRODUCTION

The most severe core conditions resulting from an accidental depressurization of the reactor coolant system are associated with an inadvertent opening of a pressurizer safety valve. Initially, the event results in a rapidly

~

decreasing reactor coolant system pressure until this pressure reaches a value corresponding to the hot-leg saturation pressure. At that time, the pressure decrease is slowed considerably. The pressure continues to decrease, however, throughout the transient. The effect of the pressure decrease would be to decrease the neutron flux via the moderator density feedback, but the reactor control system (if in the automatic mode) functions to maintain the power-  ;

essentially constant throughout the initial stage of the transient.

This transient was reanalyzed to determine the effect of the uprating and I

positive MTC on the nuclear power transient and minimum DN8R.

s l 3.1.3.3.9.2 METHOD OF ANALYSIS

i. The accidental depressurization transient is analyzed with the detailed digital computer code LOFTRAN. The code shulates the neutron kinetics, reactor coolant system, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and steam generator safety valves. The code computes pertinent plant variables, including temperatures, pressures, and power level.

l

' The ITOP was employed in analyzing this transient. Initial operating l

conditions are assumed at values consistent with tt4 steady-state N loop _

operation. All other assumptions are consistent with the FSAR. A constant

+6 pcm/*F MTC was assumed for beginning-of-life conditions.

l l

2156e:1d/112084 45 3-27

3.1.3.3.9.3 RESULTS Figure 3.1-34 illustrates the flux transient following the accident. Reactor trip on overtemperature delta-T occurs as shown in Figure 3.1-34. The l pressure decay' transient following the accident is given in Figure 3.1-35.

The resulting DNBR never goes below the limit value, as shown in Figure 3.1-36.

3.1.3.3.

9.4 CONCLUSION

S The pressurizer low pressure and the overtemperature delta-T reactor protection system signals provide adequate protection against this accident; the minimum DNBR remains in excess of the limit value.

1

{

3.1.3.3.10 SPURIOUS OPERATION OF THE SAFETY INJECTION SYS1EM AT POWER t

3.1.3.3.

10.1 INTRODUCTION

Spurious safety injection system (SIS) operation at power could be caused by operator error or a false electrical actuating signal.

An SIS signal normally results in a reactor-trip followed by a turbine trip.

However, it cannot be assumed that any single fault that actuates the safety injection system will also produce a reactor trip. Therefore, two different

- courses of events are' considered:

. Case A - Trip occurs at the same time spurious injection starts.

Case 8 - The reactor protection system produces a trip later in the transient.

l l -

This accident was reanalyzed to determine the effect of core uprating and l

positive MTC on the nuclear power transient and minimum DN8R.

.The spurious operation of the Safety Injection System accident has been reanalyzed at the uprated conditions to' demonstrate that the low pressure reactor trip provides adequate protection against DN8. ,

2156e:1d/112084 46 I

3-28

3.1.3.3.10.2 METHOD 0F ANALYSIS The spurious operation of the safety injection system is analyzed by using the detailed digital computer program LOFTRAN. The code simulates the neutron kinetics, reactor coolant system, pressurizer, pressurizer relief.tand safety valves,pressurizerspray,steamgenerator,steamgeneratorsafethvalves,and the effect of the safety injection system. The program computes pertinent plant variables, including temperatures, pressures, and power level.

Because of the power and temperature reduction during the transient, operating

. conditions do not approach the core limits. Analysis of several cases show that the results are relatively independent of time to trip.

l A typical transient is presented representing conditions at beginning of core life. Results at end-of-life are similar, except'that moderator feedback effects result in a slower transient.

The ITOP was used in analyzing this transient. Initial operating co1ditions ,

!l are- assumed at values consistent with the steady-state N loop operation. The boron injection into the cold legs of each loop is assumed to be at 12,250 ppm. A constant +6 pcm/*F MTC was assumed for the beginning-of-life conditions and all other assumptions are consistent with the FSAR.

l Four cases were analyzed; 2 maximum feedback cases and 2 minimum feedback

( cases, and for each of these, one case will have reactor trip on SI and the 2

other will not.

3.1.3.3.10.3 RESULTS The transient response is shown in Figures 3.1-37 through 3.1-28. Nuclear power starts decreasing immediately due to boren injection, but steam flow

'does not decrease until 43 sec into the transient, when the-turbine throttle valve goes wide open. The mismatch between load and nuclear power causes 4

2156e:1d/112084 47 3-29

i l

T,, , pressurizer water level, and pressurizer pressure to drop. The low-pressure trip setpoint is reached at 59 sec, and rods start moving into '

the core at 61 sec. y l

After trip, pressures-and temperatures slowly rise, since the turbine is I tripped and the reactor is producing some power due to delayed neutron fissions and decay heat.-

3.1.3.3.

10.4 CONCLUSION

S l

4 Results of the analysis show that spurious safety injection with or without immediate reactor trip presents no hazard to the integrity of the reactor coolant system.

2 The DNBR is never less than the initial value.- Thus, there will be no cladding damage and no release of fission products to the reactor coolant l

system.

If the reactor does not trip immediately, the low-pressure reactor trip will be actuated. This trips the turbine and prevents excess cooldown, thereby expediting recovery from the incident.

j 3.1.3.3.11 COMPLETE LOSS OF FORCE 0 REACTOR COOLANT FLOW l

3.1.3.3.

11.1 INTRODUCTION

As shown in the FSAR, the most severe loss of flow transient is caused by the simultaneous loss of electrical power to all three pumps while the reactor is at full power. This transient was reanalyzed to determine the effect of the uprating, and a positive MTC on the nuclear power transientiand the resultant effect on the minimum DNBR. The effect of the MTC on the nuclear power transient would be limited to the initial stages.of the event during which reactor coolant temperature increases; this increase is terminated shortly after reactor trip..

!=

l 2156e:1d/112084 48 3-30 m

'l 3.1.3.3.11.2 METHOD OF ANALYSIS This transient is analyzed using three digital computer codes. First, LOFTRAN (Ref. 3.1-7) is used to calculate loop flow, core flow, time of reactor trip, and the nuclear transient following reactor trip. The FACTRAN code (Ref.

3.1-6) is then used to calculate the heat flux transient based on the nuclear power and flows from LOFTRAN. Finally, the THINC code (Ref. 3.1-9) is used to calculate the minimum DNBR based on the heat flux from FACTRAN and flows from LOFTRAN. The WRB-1 DNB correlation is used (Ref. 3.1-3). The transient presented represents the minimum of the typical or thimble cell.

The analysis was performed using the uprated power, a constant +6 pcm/*F MTC, and ITOP as described in Reference 3.1-2. ~ Initial operating conditions are assumed at values consistent with steady-state N loop operation.

3.1.3.3.11.3 RESULTS Figures 3.1-39 through 3.1-41 show the flow coastdown, nuclear power and heat flux transients, and the minimum DNBR versus time. The DNBR is greater than j the limit value.

3.1.3.3.

11.4 CONCLUSION

S The analysis performed demonstrated that for a complete loss of forced reactor coolant flow, the DNBR does not decrease below the limit value during the j transient and thus, there is no clad damage or release of fission products to the reactor coolant system.

3.1.3.3.12- SINGLE REACTOR COOLANT PUMP LOCKED ROTOR l

3.1.3.3.

12.1 INTRODUCTION

l The locked rotor incident is analyzed to demonstrate that the peak RCS pressure and clad temperature reached during the transient is less than that which would cause stresses to exceed the faulted condition stress limits and compromise the integrity of the primary coolant system.

2156e:1d/120384 3-31 I

l' i

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

l Analysis was perforned for an instantaneous seizure of a reactor coolant pump i rotor with the reac*or at 102 percent power and three loops operating.

Following the rotor telzure, the RCS temperature rises until shortly after reactor trip.

4 4

3.1.3.3.12.2 NETH00 0F ANALYSIS I

t. .

lheo digital computer codes were used in this analysis. LOFTRAN calculated the loop and core coolant flow, the nuclear power and the peak pressure following ,

i the pump seizure and subsequent reactor trip. The thermal behavior of the fuel located at the core hot spot was investigated using the FACTRAN code l

i based on the core flow and nuclear power calculated by LOFTRAN. Plant setpoints and' errors used were consistent with the FSAR with respect to l initial power and flow fractions, RCS pressure and temperature. To obtain the maximum pressure in the primary side, conservatively high loop pressure drops are added "to the ca1culated pressurizer pressure. A constant +6 pcm/*F MTC was used. This is conservative because the MTC will be zero or negative at f

greater than 70 percent power.

3.1.3.3.12.3 RESULTS Transient response of the reactor coolant system during the analyzed locked l rotor incident is shown in Figures 3.1-42 through 3.1-44. Maximum RCS pressure, maximum clad temperature and amount of zirconium-water reaction are l

listed in Table 3.1-2. Figure 3.1-45 depicts the clad temperature during the j transient.

3.1.3.3.12.4- CONCLUSIONS I Results of the postulated locked rotor event show that the integrity of the j primary coolant system is maintained since the peak RCS pressure during the ,

j transient is less than that which would cause stresses to exceed the faulted condition stress limits. The peak clad surface temperature calculated for the 2156e:1d/120384 3-32

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

hot spot during the worst transient remains considerably less than 2700*F, thus, the core will remain in place and intact with no loss of core cooling capability.

3.1.3.3.13 RUPTURE OF A CONTROL R00 ORIVE MECHANISM HOUSING (R00 CLUSTER CONTROL ASSEM8LY EJECTION) i 3.1.3.3.

13.1 INTRODUCTION

This accident is defined as the mechanical failure of a control rod mechanism pressure housing resulting in the ejection of a rod cluster control assembly and drive shaf t. The consequence of this mechanical failure is a rapid reactivity insertion together with an adverse core power distribution, possibly leading to localized fuel rod damage. This transient is analyzed at full power and hot standby for both beginning and end-of-life conditions.

3.1.3.3.13.2 METHOD OF ANALYSIS The method of analysis used is consistent with the FSAR. Only the BOL cases included a positive MTC. A +6 pcm/*F MTC was used at zero power nominal average temperature, decreasing to a small positive value at full power T-average. This is still a conservative assumption since the moderator coefficient is zero or negative above 70 percent power. The MTC is negative at end-of-life and thus, was not included in those cases.

Table 3.1-3 lists key input parameters used in this analysis.

3.1.3.3.13.3 RESULTS I

Table 3.1-3 sunnarizes the rod ejection preliminary results for all cases.

The peak hot spot clad temperature 2654*F, is reached in the 80L hot zero power case. Maximum fuel temperatures occur in the 80L hot full power case.

2156e:1d/012385 53 3-33

Although the peak hot spot fuel centerline temperature exceeded the melting temperature in both the 80L and E0L hot full power cases, melting was restricted to less than the innermost 10 percent of the fuel pellet.

3.1.3.3.

13.4 CONCLUSION

S I

As fuel and clad temperatures do not exceed the limits specified in the FSAR, there is.no danger of sudden fuel dispersal into the coolant or consequential damage to the primary loop.

3.1.3.3.14 SINGLE R00 CLUSTER CONTROL ASSEMBLY WITH0RAWAL AT FULL POWER 3.1.3.3.

14.1 INTRODUCTION

i No single electrical or mechanical failure in the Rod Control System could cause the accidental withdrawal of a single rod cluster control assembly from the inserted bank at full-power operation. The operator could deliberately withdraw a single rod cluster control assembly in the control bank. In the l extremely unlikely event of simultaneous electrical failures resulting in I single rod cluster control assembly withdrawal, rod deviation and rod control urgent failure would both be displayed on the plant annunciator, and the rod position indicators would indicate the relative positions of the assemblies in the bank. The urgent failure alarm also inhibits automatic rod motion in the group in which it occurs. Withdrawal of a single rod cluster control assembly by operator action, whether deliberate or by a combination of errors, would result in activation of the same alarm and the same visual indications.

In the unlikely event of multiple failures, that result in continuous withdrawal of a single rod cluster control assembly, it is not possible, in all cases, to provide assurance of automatic reactor trip so that core safety limits are not violated. Withdrawal of a single rod cluster control assembly results in both positive reactivity insertion tending to increase core power, and an increase in local power density in the core area " covered" by the rod

. cluster control assembly.

i 2156e:1d/112084 52 3- 34

. 3.1.3.3.14.2 METHOD OF ANALYSIS 1

The power distributions are analyzed using the computer codes as described in Table 15.1-2 in the FSAR. The peaking factors are then used as input to the THINC code to calculate the DN8R.

. Two cases have been analyzed:

i 1. The reactor is in the manual control mode

2. The reactor is in the automatic control mode.

i Soth cases were performed at a core power of 2898 MWt and pressure of 2280 psia.

3.1.3.3.14.3 RESULTS When in the manual control mode, continuous withdrawal of a single rod cluster control assembly results in both an increase'in core power and coolant temperature, and an increase in the local hot-channel factor in the area of the failed rod cluster control assembly. In terms of the overall system

! response, this case is similar to those presented in Section 3.1.3.3.2; however, the increased local power peaking in-the area of the withdrawn rod cluster control assembly results in lower minimum DN8R's than for the withdrawn bank case. Depending on initial bank insertion and location of the

withdrawn rod cluster control assembly, automatic reactor trip may not occur

! sufficiently fast to prevent the minimum core DNBR from falling below the

limit value. Evaluation of this case at the power and coolant conditions at which the overtemperature delta T 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 percent.

2156e:Id/112084 53 ,

3-35 9 w v--,+ ,y-+y. -y --w ----,*=w +.- --g- ,, w -p- y-g .w.-w-,w --y.y- ww gr-a,-wg . w ,w+- -n-ywt, + yww-- r

If the reactor is in the automatic control mode, withdrawal of a single rod cluster control assembly will result in the immobility of the other rod

. cluster control assemblies in the controlling bank. The transient will then proceed as described above. For cases like the above, a trip will ultimately i ensue, although not sufficiently fast in all cases to prevent a minimum DNBR in the core of less than the limit value.

3.1.3.3.

14.4 CONCLUSION

S For the case of one rod cluster control assembly fully withdrawn, with the reactor in the automatic or manual control mode and initially operating at full power'with bank 0 at the insertion limit, an upper bound of the number of fuel rods experiencing a DNBR less than the limit value is 5 percent of the total fuel rods in the core.

/

1 l

l I

i 2156e:1d/112084 54 3-36 F

..m-__. .

3.1 REFERENCES

3.1-1. North Anna FSAR.

3.1-2. Chelemer, H.; Boman, L. H.; Sharp, D. R., " Improved Thermal Design Procedures," WCAP-8567. July 1975.

1' 3.1-3. Notely F. E.; et al., "New Westinghouse Correlation WRB-1 for L Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids,"

WCAP-8762, July 1976.

3.1-4. Tong, L. S., " Critical Heat Fluxes in Rod Bundles, Two Phase Flow and 4 Heat Transfer in Rod Bundles," Annual Winter Meeting ASME, November i 1969, p. 3146, 3.1-5. Tong, L. S., " Boiling Crisis and Critical Heat Flux." AEC Office of j

Information Services, TID-25887, 1972.

3.1-6. Hargrove, H. G., "FACTRAN-A FORTRAN-IV code for Thermal Transients in a U0 Fuel Rod," WCAP-7908, June 1972.

2 3.1-7. Burnett, T. W. T.; et al., "LOFTRAN Code Description," WCAP-7907-A, April 1984.

Risher, D. H., Jr.; Barry, R. F., " TWINKLE-A Multi-Dimensional Neutron 3.1-8.

Kinetics Computer Code," WCAP-8028-A, January 1975.

i i

2156e:1d/112084 55 3-37 l

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

3.1 REFERENCES

(Continued) i l

3.1-9. Hochreiter, L. E.; Chelemer, H.; Chu, P. T., "THINC-IV an Improved Program for Thermal Hydraulic Analysis of Rod Bundle Cores,"

WCAP-7956, June 1973.

3.1-10. Bordelon, F. M., " Calculation of Flow Coastdown After Loss of Reactor Coolant Pump (PHOENIX Code)," WCAP-7969, September 1972.

3.1-11. Morita, T., et al., " Dropped Rod Methodology for Negative Flux Rate Trip Plants," WCAP-10297-A, June 1983.

f l

2156e:1d/112084 56 3-38

TABLE 3.1-1 SmetaaY OF INITIAL ConstTIONS Aus COIWUTit C00E5 USED seactivity Coefficients Assumed .

lattial N555 stederatera Itederatera Thermal Power Camputer Temperature Beastty eutput Assumed SIS Faults Codes Used (Ak/*f) (Ak/ge-cm3) Sepplerb (ftet) Correlatlea Canditlea 11 Uncontrolled K C assently beak with- PWOENIX,1WIINLLE 4 a 10-5(c) -

Lower e v.3 drawal from a subcritical conditten FAtitAII.TNInc uncontrolled BCC assently beak with- LOFTRAll 4 a 10-5(d) 0.43 (d) Lower and 2918 unt-1 drewel at power Upper DCC asse d ly alsalignment TWINC. TURTLE, -3 a 10-5 _ ,,,,, ,,y , ,,,, ,

LOFitAII  ;

Partle) less of forced reacter LOFTRAal,FACTRAal 4 a 10-5 - upper 2910 tes-1 coolant flew Startup of ao teective reacter LeFiaAII.FACTRAll ---

0.43 Lower 1101 W-3 coolant leep . TWIIIC Less of esternal electrical lead LeFiaAN + 6 a 10-5 g,,3 y,,,, ,,g ,,,, y,,,,

and/or turbine trip tower I

!  % 1y one is used la an analysis, f.e., either moderator temperature er moderator density coefficleet.

b j eeference Figure 15.1-5, merth Anne FSAR.

"This corresseeds to A 4 pcm/*F at aere power (se lead T,q) and becomes less positive at higher temperatures.

d+4 pcm/*F is assumed for stalmum feedback ello 0.43 Ak/gn-cm3 is used for meulaum feedback, t

1

risse

idiiro w4 3-M

TASLE 3.1-1 (continued)

SWWIAAT OF IRITIAL CeutlileNS AND CSIPWTER CSOES If5ES Beactivity Coefficients Assumed .

Initial e555 leaderatera Nederatera Thermal power Camputer Temperature pensity Output Assimmed Ong Faults Codes used ( Alt /*F) (Ak/ge-cm3) Sepplerb (ftet) Correlattee Cead1ilos II Encessive heet removal due to LOFieAn -

0.43 Leuer Mit tes-1 feeduster system malfunctlens Encessive lead Secrease LeFTAAN - 0.43 Leuer and upper Mit ime-1 Accidental depresserlastion of LOFitAII +4 a 10-5 _ gg,,,, gyg g,g,y the reacter coeleet system .

Inadvertent operetten of ECCS LOFisAII 4 10-5 0.43 (d) Lauer and M10 IAS-1 during pesier operatlea lipper Condities III Complete less of forced reacter LSFT8All. TNINC. 4 a 10-5 - tipper M10 IAS-1 coolant flew FAC1 Ball

Slogie BCC asseely withdrauel TURTLE. Tulut. -3 a 10-5 - IIA Mit Wee-1 ct full peuer LEspese l

• Ibily one is used la an analysis, i.e.. either anderator temperature or anderator doestty coefficleet.

"Aeference Fleere 15.1-5. Isorth Anne F5Aa.

) and becomes less positive et higher temperatores.

l "Tils 84 correspeeds pcW*F to A 4fed is asW tw staloms pcm/*F .dille at 0.43aere Aa/ peuer (nofor s 15 used lead T,Mfeseeck.

mesleum #

2t %e:1d/120384 3-40 l

~ .

TASLE 3.1-1 (coattneed) 54pstARY OF INITIAL CONOffl0NS ANS CONPUTER Cests USEB Beactivity Coefficleets Assumed .

Initial N555 floderatera floderstora Thersel power Campeter Teauperature Deestty Setput Assemed gm Feelts Codes used ( Ak/*f ) (R/gn-cm 3) Sepple,6 (Itft) Correlatten Condities 11 Slagte roector coolant peep LeFisAN. +6 a 10-5 - upper 2910 use-1 locked roter FACitan emptore of a centrol red influttt.FAcitasp +6 pen /*F SOL - Consisteet 2910 Not mechenlso heeslag (ACCA LEspAas with lower Appilcable ejectlee) llelt shown la Figure 15.1-5. .

I a

anty one is used in an analysts.1.e., either moderator tomtere or moderator doestty coefficient.

I "eeference Flgere 15.1-5. Iterth Anne FSAR.

"This correspeeds te A e6 pca/*F at aere power (ne lead T,q) and becemos less positive at higher temperatores.

3 8+6 pcm/*F is assumed for slainen feedback telle 0.43 Ak/gn-cm 15 used for eenlem feedhock. .

21%e:1d/120384

4 TABLE 3.1-2 0 5

~

LOCKED ROTOR RESULTS 2910 MWt Ucratina Initial power, percent 102 Moderator temperature coefficient, prm/*F +6 Maximum reactor coolant pressure, psia 2722 l

Maximum clad average temperature, 'F 2203 Amount of Zr-water reacted at core hot 1.1 spot, percent weight 4

i i

i l

6 i

i l

i i

2156e:1d/012305 62 3-42

8

\

TABLE 3.1-3 PARAMETERS USED IN AND RESULTS OF THE t

. ROD CONTROL CLUSTER ASSEM8LY EJECTION ACCIDENT BOL 80L EOL EOL HZP HFP HIP HFP Power Level, percent 0 102 0 102 Ejected rod worth, percent K .878 .20 .99 .21 Delayed neutron fraction, percent .52 .52 .43 .43 Feedback reactivity weighting 2.73 1.30 3.55 1.60 Trip reactivity, percent K 2 4 2 4 Fq before rod ejection 2.55 2.55 2.55 2.55 Fq after rod ejection 15.40 7.07 19.20 7.50 MTC, pcm/*F +6 +6 i none none Max fuel ave temp, *F 3746 4038 3113 4115 Clad avg temp, 'F .t' /)) 2654 2200 2197 2397 Max fuel centerline' temp, 'F 4325 4967 3620 4921 Max fuel stored energy, cal /gm 1 61 .2 176.3 128 179 Percent fuel melting, 0 <10 0 <10 t

2156e:1d/121184 15 3-43

1 1

i i

) .

==

EE 10.0!! 55 l 33 .

.. ~. .

g .. ..

~

E t3 1. 0 =.. =

:

as w  :: ..

u m *a d ..

5 =

oc 0.1 : :

• g  :: $

w

' E ..:". .t-

.L

.01 +

10.0 15.0 20.0 25.0 30.0 0.0 5.0 Time (Seconds)

I FISURE 3.1-1 i

I NEUTRDN FLUX VERSUS TIME

' UNCONTROLLED ROD WITHDRAWAL FROM A SUSCRITICAL CONDIT!Dtl i

l I

l

l l

l I i t  !

l l  !

1.0 l l l

^

0.8 - -

W h

.8 0.6 - -

5 C'

W 5 ..

g 0.4 - -

i d W

W ""

w 0.2 - -

8 .

0.0 -

15.0 20.0 25.0 30.0 35.0 40.0 O.0 5.0 - 10.0

' Time (Seconds) l FIGURE 3.1-2 CORE HEAT FLUX VERSUS TIPE UNCONTROLLED ROD WITHDRAWAL FROM A SUSCRITICAL CONDITION

I

l . I i 1000  ; i

~~

~

900 '

C.

~ ~~

800- -

E E

ky

~~

700 - - Fuel U

5 Clad 600 - -

l Coolant '

l  ;

500 l l l O.0 5.0 10.0 15.0 20.0 25.0 30.0 35.0 40.0

' Time (Seconds)

FIGURE 3.1-3 TEMPERATURE VERBUS TIME UNCONTROLLE D ROD WITHDRAldAL FROM A SUSCRITICAL CONDITION

j l

l

\

i

! )

1 2.2 i /

2.1 - -

Minimum Feedback /

---

Maximum Feedback .,

/

2.0 - - l l

- High y Neutron Flux i

! 1.g . . / l

= '- I l $ '

E 1.s - - OTST s e E s,4 i

1.7 - -

1.6 - -

1.5 4 6 8 10-5 2 4 6 8 10~4 2 4 6 81 10-6 2 l

Reactivity Insertion Rate, ak/sec.

l l

FIGURE 3.1-4 MINIMUM DNBR VERSUS REACTIVITY INSERTION RATES AT 100% POWER ,

MAXIMUM AND MINUMUM FEEDB4CK '

1 2.9 -

' 2.7 . Minimum Feedback

---

Maximum Feedback ,

l H'9h -' .

2. 5 - .

Neutron Flux i 2.3 - -

f

=

OTai s

,e " (

i s

• E 2.1 - > o

'. \1 1.9 -

1. 7 - -

1.5  ; ,

6 8 10-3 2 4 6 8 10-5 2 4 6 8 10'4 2 4 I

i Aeactivity Insertion Rate. AVsec.

I FIGURE 3.1-5 MININUM DNBR VERSUS REACTIVITY INSERTION RATES AT 607. POWER MAXIMUM AND MINIMUM FEEDB/4CK .

i I

i 3,9 - - Minimum Feedback <

i

---

Maximum Feedback f.

3.5 - - f

' I Hich 8 I

Neutron Flux  :

i 3.1 I 8 OTAT j  ! 2.7 " " -

e. s

, N/

t

- , y

2. 3 - ~

l

1. 9 - -

l

! 1.5 .

r 2 4 6 810~4 2 4 6 8 10' 2 t i

Reactivity Insertion Rate, ak/sec.

FIGURE 3.1-6

^

MINIMUM DNBR VERSUS REACTIVITY INSERTION RATES AT 10% POWER MAXIMUM AND MINIMUM FEEDBACK i

6 />

s

+%

i 1.2  :  :  :  : ,

1 i

1.0 - ,

CORE FLOW 3d 55 --

t-5

=z 0.8 - -

m

'o LOOP FLOW

$5 g~ 0. 6 . . ..

. vu mm ,

~~

83 0.4 - -

dd W85 8S --

0.2 - -

i l 0.0 -  :  :

0.0 2.0 4.0 6.0 8.0 10.0

(

i i

! Time (Seconds)

FIBURE 3.1-7 ALL LODPS OPERATING, DNE PUMP COASTING DOWN FLOW COASTDOWN VERSUS TIPE PARTIAL LDSS DF FLOW

l l

1 1.2 '?  ;  ;  ; -a -

l 1.0 -

, 3 0.8 . .

Heat Flux 5 '

5

=

la.

o 0.6 .

5

. NEUTRON b FLUX 5 0.4 - - --

's d

0. 2 - -

0.0 -  :  ;  ;

[ 0.0 2.0 4.0 6.0 8.0 10.0 I Time (Seconds)

FISURE 3.1-B ALL LOOPS OPERATING, DNE PUMP COASTING DOWN i

FLUX TRANSIENTS VERSUS TIME i PARTIAL LDSS OF FLOW l

o

~

i t-

I f

I 2.4 -' l l l l 2.2 - -

2.0 - -

ac --

g 1. 8 - -

i

1. 6 - -

1.4 l

1.2 l  :

f 0.0 2.0 4.0 6.0 8.0 10.0 i

Time (Seconds)

FIGURE 3.1-9 ALL LDDPS DPERATING, DNE PUMP CDASTING DOWN DNBR VERSUS TIME PARTIAL LDSS OF FLOW t

I l

1 l

1.2 l

- 1 l 1.0 < -

i g!

[/

.40 -

EU

, g .20 " '

L 0.0 -

_ 1.20 N 1.00 El .800 -

W 5

> .600-E E .

C .400 W .,

E .200 0.0 '

- 620

$ 600 5

g ;E 580

<s w Q 560 85 "g- 540 520 l 2600 j , .

, .~

ar 5 " "

l 2400

, gg .

Ew 2200 l

cg , ..

)

w e EO 2000 " "

E . .,  !

1800 0 5 10 ~15 '20 25 30 Time (Seconds) l l

FIGURE 3.1-10 STARTUP OF AN INACTIVE REACTOR COOLANT LOOP LOOP STOP VALVES INITIALLY OPEN

1. 1.20  :  :  :  :

5 z

@ 1.0< ,

u.

o

. 0.8< -

o E

t;, 0.6< ,

5 3

n.

0.4< ,

l 0.2, ,

w ,

y Y

= 0.0:  : . . .

26u0  :  :  : .

- 2500 ,

G n.

2400< ,

M 2300, ,

5 g 2200, ,

w g 2100, ,

5 N

2000< ,

$ 1900< , -

$ 1800, ,

E 1700 . .

5.00  :  :  :  :

4.50 ,

4.00 ,

3.50 ,

a:

E 3.00, ,

o , ,

2.50, ,

2.00< ,

4 .

1. 50< ,

1.0C  :  :  :

0 20 40 60 80 100 TIME (SECONDS)

FIGURE 3.1-11 LOSS OF-LOAD ACCIDENT, WITH PRESSURIZER SPRAY AND i

POWER 0PERATED RELIEF VALVE, BEGINNING OF LIFE

i

$400.0  : 0  ;  ;

3300.0 < -

p.0 < -

1500 0 < -

1 2000.00< -

g s00.00 < - -

800.00< -

g m.00 < , -

y .00.00< -  ; -

500.00  :  ; "

$40.00  :  :  :

C u 520.00- -

U I<600.00<

W g 580.00- -

E I 4 560.00< -

l 540.00  :  :  :

0 2'O 40 60 80 100 TIME (SECONDS)

, FIGURE 3.1-12 l LOSS OF LOAD ACCIDENT, WITH PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, BEGINNING OF LIFE

3 E 1.20 .  : ,

E E 1.0' # , ,

5

. 0.8< ,

u < ,

5 0.6< -

, p 0.4- -

=

5 0. 2 ' '

d E 0.0 :

L 2600  :  :  :  :

5 2500<

m s <.

~

S 2400- -

w w5 2300 '

,0 2200< ,

c.

= 2100 < >

M g 2000' -

S '

=

1900<

1800  :  :  :  ;

5.00 ^

^

4.50 '

4.00 -

3.50 < ,

m

= 3.00< >

E < >

2.50 ,

i 2.00 .

1.50 ,

, 1.00  :  :  : -

0 20 40 60 80 100 TIME (SECONDS)

FIGURE 3.1-13 LDSS OF LOAD ACCIDENT, WITHOUT PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, BEGINNING OF LIFE l

l

i i

1400.0  :  :  ;

gs300.0 - -

gia00.0 < -

1100.0- -

)

i100000- .

\

t 300.00< -

k gg,g .

g,00.00

[ s00.00 < -

500.00  :  :  :  :

640.00:  ;  ;  ;  ;

C fa 620.00- -

i U

4 SM .00 < -

I

, ,00.00 .

4 I

4 560.00< -

540.00 :  :  : 0  ;

O 20 40 60 80- 100 TIME (SECONDS) l l

FIGURE 3.1-14 '

LOSS OF LOAD ACCIDENT, WITHOUT PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, BEGINNING OF LIFE '

i 1.2 O

~

g i

\

gg .8+

d t

y !5 -

"E g: .4 ..

I n

6 ,,

i 0

2600  ;

q ,,

17.

E 2400 w

g .,

N "

[ 2200 g .,

~ "

g 2000 ..

a w ,,

f '

1800 4

3 t l l ee ,

g  ;

'2 .,

I l 10 20 30 40 50 0

TIME (Seconds) l FIGURE 3.1-15 i LOSS OF LOAD ACCIDENT, WITH PRESSURIZER SPRAY AND FOWER OPERATED RELIEF VALVE, END OF LIFE l

{

l

1400 1300 "

1. 1200 -

5

]

$ 1100 - .

W l --

% 1000 - -

5 m "

y 900 ~

s m "

O 800 - -

f -

700 --

600 i

i 640 620 --

C.

w w

B E 600 --

5 b

- l W '

$ 580 --

i e W

= ..

560 -

540 20 30 23 '50 0 10 i;ME (Secous)

FIGURE 3.1-16' l

. LOSS OF LDA'D ACCIDENT, WITH PRESSURI7ER SPRAY AND l i POWER OPERATED RELIEF VALVE, END OF LIFE l

1.2 n

$ \

l

• 8 *'

ng '\

d

,8 .,

) <

"E h3 .4 o I a w

L 0

2600 d,

2400 "

w g ,,

l M ,

E 2200 5

• 2 ,,

E 2000 "

g g ,,

n.

1800 5

4 '( i 3"

2 'l l

1 O 10 20 30 40 M TIME (Seconds)

FIGURE 3.1-17

- LOSS OF LOAD ACCIDENT, WITHOUT PRESSURIZER SPRAY AND POWER DPERATED RELIEF VALVE, BEGINNING DF LIFE p

F . _ _ - . - - - - ---

l 1400-

^

, 1200 " -

.C I g - -

W "

" 1000 "

- W 5

~

a ..

W N

E "

$ 800 E

600 ,

640 '

- 620 "

w 5

E 600 k-w

' ~

$ 580 "

E

=

W ..

l 560 - -

540 10 20 30 40 SC

, 0 TIME (Seconds)

FISURE 3.1-18 LDSS OF LDAD ACCIDENT, WITHOUT PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE, BEGINNING OF LIFE

l w

i i

• 1.2 ;  ;

,,..w n .

8,y!.Til"ll!",, l , , ,

I

_ l .0_-

a i_. ~

l . 80 - -

b ,

E c' ""

W .60 - -

_E ac "'

. 40 -

E "i -

l .20 . .

l-

.0 .

7 z

50 75 100 125 150 175 0 25 i

TIME (Seconds)

FIGURE 3.1-19 NUCLEAR POWER VERSUS TIPE FEEDWATER SYSTEM MALFUNCTON l

l

)

2500 2400 7 ""

5 2300 - - - -

~

l E"'

M --

g 2200 5

.~ ""

! 2100 e

E w ..

2000 1900 50 75 100 126 150 175 l 0 25 1

TIME -(Seconds)

FIGURE 3.1-20 PRESSURIZER PRESSURE VERSUS TIME FEEDWATER SYSTEM MA:.FUNCTON

I I

M , , , ,

I 70 . .

60 - -

^

w ..

e -

g 50 -

i 4 40 - -

x 30 . .

8 l

20 - -

10 . .

; e i 0  :  :

0 25 50 75 100 125 150 175 TIME (Seconds)

I FIGURE 3.1-21 LOOP DELTA-T VERSUS TIME FEEDWATER SYSTEM MALFUNCTON I

600  :  :  :  :

~~

590 r --

F 580.

~

w k

n " ~

y 570- -

v ..

8 560. .

E W

550 100 125 150 175 0 25 50 75 i

TIME (Seconds)

FIGURE 3.1-22 CORE TAVG VERSUS TIME FEEDWATER SYSTEM MALFUNCTON I

--e -- A - ~-

i l

I l

l l l I i 1.50 , l l l 1.45 1.35 1.25 E

m $

g 1.15 ..

e x

&~

w a 1.05 --

58 /

w a _

E 0.95 ..

Wb

^Qu.

0.85 ..

0.75 . - '

l 1

25 50 75 100 125 150 175 0

TIME (Seconds)

FIGURE 3.1-23 STEAM GENERATOR MASS (FAULTED LOOP) VERSUS TIME FEEDWATER SYSTEM MALFUNCTON i

i

)

-.rw

I 1

l i

e i '

I I 1.25  :

1.15 2

co v, m

" 1.05 . ..

hE

=-

! 25 0.95 --

s3 xS .

5E ""

$s 0.85 0.75 , - * ' '

O 25 50 75 100 125 150 175 TIME (Seconds) i FIGURE 3.1-24 STEAM GENERATOR MASS (UNFAULTED LOOP) VERSUS TIME FEEDWATER SYSTEM MALFUNCTON i

l

l

-. i l

I i

5.0  : .

4.5 t 4.0 -- ,

3.5 --

E 3.0 - .

-2. 5 --

2.0 ..

1.5 -

100 125 150 175 O 25 50 75 i

TIME (Seconds) l l

l 1

FIGURE 3.1-25 DNBR VERSUS TIME l

l FEEDWATER SYSTEM MALFUNCTON*

i

- y - , ,-r - , , -. v-. .w ,.m--

1.20  ::  ;  ;  ;  ;

1

l 5 1.15 . .

i - 1 .10 . .

-ac s~

l 1.05 . .

$ 1.00 . #

.950 . .

~

.900 t  ;  ;,  ;  ;  ;

620.0 <  ;  ;  ;  ;  ;

610.0 . .

C.

e 600.0 5 b

w g 590.0 < N 4

s 580.0 q .

570.0 . .

g

=

560.0 . .

550.0 1  ;  ;  ;  ;  ;

i 0 50 100 150 200 250 300 TIME (SEC) t

! FIGURE 3.1-26 EXCESSIVE LOAD INCREASE WITH MANUAL ROD CONTROL-BEGINNING OF LIFE t _ . _ . . _

2400 ~ ' ' ' '

2300 ,

h 2200 , ,

m 2100 ,

R ,*

2000 , ,

5"

E 19b0 .

1800 -  ; i  ; "

l 71 - ' '

s' e u

70 .

C 69 . .

W 68 ,, .

• T IE 67 66 .

65 , '

3.00 n  ;  ;  ;  ;  ;

2.50 .

~. se 2.00 , ,

. b 1.50 ,

I ,

(

1.00 - -

0 f f 0 0 0 50 100 150 200 250 300 TIME (SEC) ..

FIGURE 3.1-27 EXCESSIVE LDAD INCREASE WITH MANUAL ROD CONTROL BEGINNING OF LIFE

1.20 :  :  :  :  :

.E 1.15 ,

_ 1.10 . .

g 1.05 . .

E8 m 1.00

.950 . .

.900 1 l l l 620.0 -

l l l l  :

i

_ 610.0 . .

4 4

e i

$ 600.0 . .

w k

g 590.0.<

l

, W 580.0 . .

l W

1

= 570.0 .h u 560.0 .

\

550.0 , - *

. L.31'...{ ),.

. g 0 50 100 150 200 250 ~300 TIME (SEC)

FIGURE 3.1-28 EXCESSIVE LOAD INOREASE WITH MANUAL RDD CONTROL END OF LIFE

2300'. .

L 2200 .

5 2100 .

E 2000 i_E I

i 1900 ,

1800 - * = = = =

4 .h *

• 4 e 71 ' ' '

• 70 .

C 69 ,

a W

~

68 ,

T

=*

67 .

66 ,

(

65- ' ' '

! 3.00 e  ;  ;  ;  ;

2.50 . .

,2.00 i

1.50 . .

1.00 - -

l 0 50 100 150 200 250 300 ,

TIME (SEC) .

FIGURE 3.1-29 EXCESSIVE LDAD INCREASE WITH MANUAL ROD CONTROL END OF LIFE

i 1.20 : .

1.15 , ,

E.

~

1.10 . .

l l ^ l

!~ an. - k l

-5 1.05 '

m:

g l

1.00 . .

!!! g -

w g .s50 .

3: .

.900 = * * *

=, . . . . ..

620.0 <-  ;  ;  ;  ;  ;

I 610.0 -

C d

E 600.0 . .

w 5 590.0 W N g

a:

1 580.0

~

r I w 570.0 . .

E w

. g 560.0 < ,

u 550.0 1 l 0 50 100 150 200 250 300 TIME (SEC)

FIGURE 3.1-30 EXCESSIVE LOAD INCREASE WITH AUTOMATIC RCD CONTROL BEGINNING OF LIFE

2400  ;  ; I l l l 2300 .

k-2200 ,

m l_ s 2100 . .

E. 2000 . .

I R e

g 1900 . .

1800 l f

n, . . . . . .

70 , ,

~

g 69 ..

68 T

5 67 .g j 66 3.00

l l l  ;  ;

l l

i 2.50 ,

i g 2.00 , ,

1.50 .

1.00 -

• 0 50 100 150 200 250 300 i

TIME (Seconds)

FIGURE 3.1-31 EXCESSIVE LDAD INCREASE WITH AUTDMATIC ROD CONTROL BEGINNING OF LIFE

1.20 g  ;  ;  ;  ;  ;

E 1.15 . .

A u

@_ 1.10

~l g 1.05

• Y

@ hi 1.00 . .

l. .950 . .

.900 4 i i 1 0 620.0 j -

610.0 . .

C e 600.0 . .

~

1 l g 590.0 , r s

5 w

580.0 . .

w

] 570.0 5 .

W

=

w 560.0 . .

.u 8

550.0 '

; O  ;  ! .

0 50' 100 150 200 250 300 TIME (SEC)

FIGURE 3.1-32

. EXCESSIVE LOAD INCREASE WITH AUTOMATIC ROD CONTROL END OF LIFE

i 2400  ;  ;  ;  ;  ;

4 2300 . .

' g A R

m 2200 . .

b2 2100 . ,

5~

~E g~ 2000 . .

c E 1900 . .

1800 . _ . . . . .

.E y v d y

71 ' = - - '

. . . . a' 70 . .

I 69 . .

C

4 68 l

I E

T 67 "

a E 66 , ,

65 - = . . t 3.00 x 2.50 . .

1 2.00- < .

as R . .

1.50 . .

1;00 W W W

• 0 50 100 150 200 250 300 TIME (SEC)

FIGURE 3.1-33 EXCESSIVE LDAD INCREASE WITH AUTOMATIC ROD CONTROL END OF LIFE

~

1

I

.l.2-- l l l l .

_ 1.0 d

5 5 --

[

o 0.8 -

5 C --

y 0.6 .

5.

E d --

, 0.4 - -

5 d

! E --

l 0.2 -

l 1 0.0 -l l O.0 5.0 10.0 15.0 20.0 25.0 30.0 35.0 40.0 TIME (SECONDS) t FIBURE 3.1-34 l

FLUX TRANSIENT FDR ACCIDENTAL DEPRESSURIZATION

l l

l l

l i

I I I 2400 .

2250 -

3 ' 2000 - -

E d "

5 1750 - -

E

= -'

M 1500 - -

E 8

i

\

x -t

' .1250 --

l J

1000 . - .

. I l

) '

I I I

800 O.0 5.0 10.0 15.0 20.0 25.0 30.0 35.0 TIME (SECONDS)

FIGURE 3.1-35 FRESSURIZER PRESSURE TRANSIENT FOR ACCIDENTAL DEPRESSURIZATION l

r ,, .

t i

1 4

/-

L e ', -

i 4

4.0 . , , , . . .

l 3.5 - -

t I

' 3.0 - -

l

! =

= --

E 2.5 - -

l 2.0 - -

1.5 - -

' ' c 3 1.0 0.0 5.0 10.0 15.0 20.0 25.0 30.0 35.0 40.0 TIME (SECONDS)-

?,I :i FIGURE 3.1-36 l DNBR TRANSIENT FOR ACCIDENTAL DEPRESSURIZATION l

>~- -

, i

'4 '

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

1. 2 -- l l l  :  :

~ -

1.0 -

tE

'd E 2 0.8 - -

W5 e z 5 STEAM FLOW E u. --

u 0.6-- NUCLE R W5 POWER -

2p un 5 E 0.4. -- --

.g ~

z i

0.2- -

L . .

0.0 .

0.0 50 100 150 200 250 300 350 400 l

! Time (Seconds) 600 C "

L 575 .

i<

2

> 550 .

525 500-O 50 100 150- 200 250 300 350 400 ,

Time (Seconds)

FIGURE'3.1-37' SPURIOUS' ACTUATION DF: SAFETY INJECTION SYSTEM AT POWER

1 12D0

,. g 1000

) d ll>

5- 800 - --

4C 2 .

58

~ - 600 " --

E

$ ~

,e 400 - - -.

200 - ~

100 2600 2500 "

2400 "

w "

E 2300 "

g 2200

'2 "

5g 2100

~ o. "

@~ 2000 "

$ 1900 "

g

) o. -

1800 - -

1700 -

l 1600 0 50- 100 150 200 250 300 350 400 I

1 Time (Seconds)

FIGURE 3.1-38 SPURIOUS ~ ACTUATION OF SAFETY INJECTION SYSTEM AT POWER l

I h.

i 1.0000< < .

7 l r i u.

o .00000< -

1' d

5 5 .50000 - -

3 d

W .40000< -

8

.t0000 - -

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0 2 4 6 8 10 TIME (SECONDS) i l

l i

i FIGURE 3.1-39 COMPLETE LOSS OF FLOW CORE FLOW VERUS TIME

l-l l

I

! 9.2000 ~  :  ; ,- ,-

l i

I 3.0000 -

l a 5 HOT CHANNEL g .MO ' '

HEAT FLUX , .

E>

g .40000< -

o E j D 40000< -

NEUTRON FLUX % AVERAGE CHANNEL 7 -

g HEAT FLUX d

.20000< -

0.0  :  ;  ;  ;  ;

O 2 4 6 8 10 i TIME (SECONDS) l l

FIGURE 3.1-40 COMPLETE LOSS OF FLOW FLUX TRANSIENTS l

+

l ... . , - - . _ . _ . . , , ._- , ..., . . - - - - . . - _ . - - . . - . . - - . . . . , . . , _ . . . . . , _ , . . - - _ . . - - . .

l l

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2.0000< -

1.8000 < -

g,gg , , DNBR Limit Typical Cell +  ; ,

Thimble Cell +

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0 2 4 6 8 10 TIME (SECONDS) l FIGURE 3.1-41 COMPLETE LOSS OF FLOW DNBR VERSUS TIME

- . - - - - - - - --- -- - - - - - - - ~ - -

l es 2800 2

2700 S.

y 2600 -

Oi w

E 2500 -

ti l 2400 -

g l 2300 -

u 15

" I I I I I l 2200 I l I 0 1 2 .3 4 5 '8 7 8 9 10 TIME (SECONDS)

FIOLNtE 3.1-42 ALL LOOPS OPetATING, ONE LODED ROTOR PRESSURE VUtSUS TIME

, ,,--<--, - - ,,, --n. - - - - - - . , - - , - - --n-. ,n.- -s,,,--, - - - rew~

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, TIME (SECdNDS) l l

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, l l

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FIGURE 3.1-43 I

ALL LOOPS OPERATING, DNE LDCKED RO. TOR CORE FLOW VERSUS TIME l

l

-m .. - - . - , , .- , -, - - - - ----e. -

1.2000; 0 0 0 0 1

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+ Heat Flux

~

z .6000< -

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w .4000< - -

+ Nuclear Flux

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• 0.0 : i 0 0 0 0 2 4 6 8 10 TIME (seconds) 1 FIGURE 3.1-44 ALL LOOPS OPERATING, ONE LOCKED ROTOR FLUX TRANSIENTS

d 4 >

I 2000.0< * ' '

i

$ 1750.0 - -

I 2 1500.0 " "

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i 188:88" e

f 4 4 l

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e I,

o TIME (SEC)

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FIGURE 3.1-45 ALL LOOPS OPERATING, OE LOCKED ROTOR CLAD TEMPERATURE VERSUS TIE i

i I

{

i

i 3.2 LOCA EVENTS 3.2.1 LARGE BREAK LOSS-OF-COOLANT ACCIDENT 3.2.

1.1 INTRODUCTION

A reanalysis of the Emergency Core Cooling System (ECCS) performance for the postulated large-break LOCA has been performed in compliance with Appendix K to 10 CFR 50. The results of this reanalysis are presented here, and are in compliance with 10 CFR 50.46, " Acceptance Criteria for Emergency Core Cooling Systems for Light Water Reactors." This analysis was performed with the NRC-approved 1981 model with BART version of the Westinghouse LOCA-ECCS evaluation model (Ref. 3.2-1 and 3.2-12). The analytical techniques used are in full compliance with 10 CFR 50, Appendix K.

As required by Appendix K of 10 CFR 50, certain conservative assumptions were 4 made for the LOCA-ECCS analysis. The assumptions pertain to the conditions of the reactor and associated safety system equipment at the time that the LOCA is assumed to occur, and include such items as the core peaking factors, the containment pressure, and the performance of the Emergency Core Cooling System. All assumptions and initial operating conditions used in this reanalysis were the same as those used in the previous LOCA-ECCS analysis (Ref. 3.2-2), with the following exceptions:

n The core power level has been increased to 2898 MWt.

4 1.

2. Seven percent steam generator tube plugging was assumed.

r

3. A thermal design flow of 92,800 gpm per loop was used.

4. The 1981 LOCA-ECCS evaluation model with BART (Ref. 3.2-10 and 3.2-11) was used to perform this analysis.

4 With the above changes incorporated into the analysis, it was found that the assumed heat flux hot-channel factor could be 2.15 and still ensure compliance with the 10 CFR 50.46 acceptance criteria.

i 2156e:1d/112684 3-44

5 i

3.2.1.2 ACCIDENT DESCRIPTION i

A LOCA is the result of a rupture of the reactor coolant system (RCS) piping or of any line connected to the system. The system boundaries considered in the-LOCA analysis are defined in Section 3.6 of the North Anna FSAR.

Sensitivity studies (Ref. 3.2-3) have indicated that a double-ended cold-leg guillotine (DECLG) pipe break is limiting. Should a DECL6 break occur, rapid

.depressurization of the reactor coolant system occurs. The reactor trip

$ signal subsequently occurs when the pressurizer low-pressure trip setpoint is reached. A safety injection system (SIS) signal is actuated when the ,

appropriate setpoint is reached and the high-head safety injection pumps are activated. The actuation and subsequent activation of the Emergency Core Cooling System, which occurs with the SIS signal, assumes the most limiting i single-failure event. These countermeasures will limit the consequences of the accident in two ways:

. 1. Reactor trip and borated water-injection complement void formation in causing rapid reduction of power to a residual level corresponding to l

fission product decay heat. No credit is taken in the analysis for the insertion of control rods to shut down the reactor.

I

2. Injection of borated water provides heat transfer from the core and

, prevents excessive clad temperature.

! 'Before the break occurs, the unit is in an equilibrium condition, i.e., the l

' heat generated in the core is being removed via the secondary system. During t

blowdown, heat from decay, hot internals, and the vessel continue to be l_

j transferred to the reactor coolant system. At the beginning of the blowdown

' phase, the entire reactor coolant system contains subcooled liquid that transfers heat from the core by forced convection with some fully developed nucleate boiling. After the break develops, the time to ON8 is calculated, consistent with Appendix K of 10 CFR 50. Thereafter, the core heat transfer is based on local conditions, with transition boiling and forced convection to steam as the major heat transfer mechanisms. During the refill period, it is assumed that rod-to-rod radiation is the only core heat transfer mechanism.

2156e:1d/112084 62 3-45 l

The heat transfer between the reacter coolant system and the secondary system

may be in either direction, depending on the relative temperatures. For the case of continued heat addition to the secondary side, secondary-side pressure l increases and the main safety valves may actuate to reduce the pressure.

Makeup-to the secondary side is automatically provided by the auxiliary feedwater system. Coincident with the safety injection signal, normal feedwater flow is stopped by closing the main feedwater control valves and tripping the main feedwater pumps. Emergency feedwater flow is initiated by l starting the auxiliary feedwater pumps. The secondary-side flow aids in the reduction of RCS pressure. When the reactor coolant system depressurizes to 600 psia, the accumulators begin to inject borated water into the reactor coolant loops. The conservative assumption is then made that injected accumulator water bypasses the core and goes out through the break until the termination of bypass. This conservatism is again consistent with Appendix X of 10 CFR 50. In addition, the reactor coolant pumps are assumed to be tripped at the initiation of the accident, and effests of pumps coastdown are included in the blowdown analysis.

The water injected by the accumulators cools the core, and subsequent i operation of the low-head safety injection pumps supplies water for long-term cooling. When the reactor water storage tank (RWST) is nearly empty,

! long-term cooling of the core is accomplished by switching to the

recirculating mode of core cooling, in which the spilled borated water is drawn from the containment sump by the low-head safety injection pumps and

, r returned to the reactor vessel.

I l The containment spray system and the recirculation spray system operate to l return the containment environment to subatmospheric pressure.

3.2.1.3 ANALYSIS 1he large-break LOCA transient is divided, for analytical purposes, into three phases: blowdown, refill, and reflood. There are three distinct transients analyzed in each phase, including the thermal-hydraulic transient in the l

2156e:1d/112084 63 3-46

reactor coolant system, the pressure and temperature transient within the containment, and the fuel clad temperature transient of the hottest fuel rod in the core. Based on these considerations, a system of interrelated computer codes has been developed for the analysis.

The description of the various aspects of the LOCA analysis methodology is given in WCAP-8339 (Ref. 3.2-4). This document describes the major phenomena

modeled, the interfaces among the computer codes, and the features of the codes that ensure compliance with 10 CFR 50, Appendix K. The SATAN-VI, COCO, WREFLOOD, BART, and LOCTA-IV codes, which are used in the LOCA analysis, are described in detail in WCAP-8306 (Ref. 3.2-5), WCAP-8326 (Ref. 3.2-6),

WCAP-8171 (Ref. 3.2-7), WCAP-9695 (Ref. 3.2-10), WCAP-10062 (Ref. 3.2-11), and WCAP-8305 (Ref. 3.2-8), respectively. These codes assess whether sufficient heat transfer geometry and core amenability to cooling are preserved during the time spans applicable to the blowdown, refill, and reflood phases of the LOCA. The SATAN-VI computer code analyzes the thermal-hydraulic transient in i

the reactor coolant system during blowdown, and the COCO computer code calculates the containment pressure transient during all three phases of the LOCA analysis. Similarly, the LOCTA-IV computer code is used to compute the thermal transient of the hottest fuel rod during the three phases.

SATAN-VI is used to determine the RCS pressure, enthalpy, and density, as well as the mass and energy flow rates in the reactor coolant system and steam-generator secondary, as a function of time during the blowdown phase of the LOCA. SATAN-VI also calculates the accumulator mass and pressure and the pipe 1 break mass and energy flow rates that are assumed to be vented to the containment during blowdown. At the end of the blowdown, the mass and energy release rates during blowdown are transferred to the COCO code for use in the determination of the containment pressure response during this first phase of the LOCA. Additional SATAN-VI output data from the end blowdown, including

,the core inlet flowrate and enthalpy, the core pressure, add the core power decay transient, are input to the LOCTA-IV code.

1 i

l 2156e:1d/112684 3-47

With input from the SATAN-VI code, WREFLOOD uses a system thermal-hydraulic model to determine the core flooding rate (i.e., the rate at which coolant enters the bottom of the core), the coolant pressure and temperature, and the quench front height during the refill and reflood phases of the LOCA.

WREFL000 also calculates the mass and energy flow rates that are assumed to be vented to the containment. Since the mass flowrate to the containment depends upon the core pressure, which is a function of the containment backpressure, the WREFL000 and C0C0 codes are interactively linked. With the input and 4

boundary conditions from WREFLOOD, the mechanistic core heat transfer model in BART calculates the fluid and heat transfer conditions in the core during reflood.

4 i LOCTA-IV is used throughout the analysis of the LOCA transient to calculate l the fuel and clad temperature of the hottest rod in the core. The input to LOCTA-IV consists of appropriate thermal-hydraulic outputs from SATAN-VI,

) WREFL000 and BART, and conservatively selected initial RCS operating conditions. These initial conditions are summarized in Table 3.2-1 and Figure l 3.2-1. (The axial power shape of Figure 3.2-1 assumed for LOCTA-IV is a ,

l cosine curve that has been previously verified (Ref. 3.2-9) to be the shape

that produces the maximum peak clad temperature.)

.l The COC0 code, which is also used throughout the LOCA analysis, calculates the

containment pressure. Input to C0C0 is obtained from the mass and energy flowrates assumed to be vented to the containment, as calculated by the SATAN-VI and WREFLOOD codes. In addition, conservatively chosen initial f containment conditions and an assumed mode of operation for the containment

! cooling system are input to C0CO. These initial containment conditions and

] assumed modes of operation are provided in Table 3.2-2.

3.2.1.4 RESULTS 4

! Tables 3.2-1 and 3.2-2, and Figure 3.2-1 present the initial conditions and mo' des of operation that were assumed in the analysis. Table 3.2-3 presents

the time sequence of events, and Table 3.2-4 presents the results for the 1

1 1

i 4

)

2156e:1d/112084 65

! 3-48

I double-ended cold-leg guillotine break for the CD = 0.4, 0.6, and 0.8 dis-

! charge coefficients. The double-ended cold-leg guillotine break has been determined to be the limiting break size and location based on the sensitivity studies reported in Reference 3.2-3. The analysis resulted in a limiting peak l clad temperature of 2160.6*F for the CD = 0.4 case, a maximum local cladding l j oxidation level of 6.69%, and a total core metal-water reaction of less than l

0.3%. The detailed results of the LOCA reanalysis are provided in Tables j

j. 3.2-3 through 3.2-6 and Figures 3.2-2a through 3.2-18c. The figures show the l l following:

1. Peaking Factor vs. Core Height - Figure 3.2-1 shows the cosine power-l shape used in the analysis.

2. Mass Velocity - Figures 3.2-2a through 3.2-2c show the mass velocity at the clad burst and hot-spot locations on the hottest fuel rod for the discharge coefficient used.

3. Heat Transfer Coefficient - Figures 3.2-3a through 3.2-3c show the i heat transfer coefficient at the clad burst and hot-spot locations on

, the hottest rod for the discharge coefficient used. The values of

!, heat transfer coefficient that are shown were calculated by the LOCTA-IV code based on equations for heat transfer in the nucleate boiling, transition boiling, film boiling, and steam cooling regimes.

l

4. Core Pressure - Figures 3.2-4a through 3.2-4c show the calculated f pressure in the core for the discharge coefficient used.

! 5. ' Break Flowrate - Figures 3.2-5a through 3.2-5c show the calculated flowrate out of the break for the discharge coefficient used. The j flowrate out of the break is plotted as the sum of flow at both the l pressure vessel end and the reactor coolant pump end of the guillotine j break. l l

l l

2156e:1d/041285 3-49

%- y .

7 6. Core Pressure Drop - Figures 3.2-6a through 3.2-6c show the calculated core pressure drop for the discharge coefficient used. The core pressure drop is interpreted as the pressure immediately before entering the core inlet to the pressure just outside the core outlet.

7. Peak Clad Temperature - Figures 3.2-7a through 3.2-7c show the calculated hot-spot clad temperature transient and the clad temperature transient at the burst location for the discharge coefficient used. The peak clad temperature for the limiting discharge coefficient of 0.4 is 2160.6*F at the 6.75 ft elevation in the core.

1

8. Fluid Temperature - Figures 3.2-8a through 3.2-8c show the calculated fluid temperature for the hot spot and burst locations for the discharge coefficient used.

9. Core Flow - Figures 3.2-9a through 3.2-9c show the calculated core flow, both top and bottom, for the discharge coefficient used.

10. Reflood Transient - Figures 3.2-10a through 3.2-10c show the reactor l pressure vessel downcomer and core water levels for the discharge i

coefficient used. Figures 3.2-11a through 3.2-11c show the core inlet velocity for the discharge coefficient used. -

11. Accumulator Flow - Figures 3.2-12a through 3.2-12c show the calculated flow for the discharge coefficient used. The accumulator delivery during blowdown is discarded until the end of bypass is calculated.

Accumulator flow, however, is established in the refill-reflood calculations. The accumulator flow assumed is the sum of that injected in the intact cold legs.

i

12. Pumped ECCS Flow (Reflood) - Figures 3.2-13a through 3.2-13c show the calculated flow of the emergency core cooling system for the discharge

{

coefficient used.

\

j I

l 2156e:1d/041285 3-50

? - _ _ _. _

double-ended cold-leg guillotine break for the CD = 0.4, 0.6, and 0.8 dis-charge coefficients. The double-ended cold-leg guillotine break has been determined to be the limiting break size and location based on the sensitivity studies reported in Reference 3.2-3. The analysis resulted in a limiting peak clad temperature of 2160.6'F for the CD = 0.4 case, a maximum local cladding oxidation level of 6.69%, and a total core metal-water reaction of less than 0.3%. The detailed results of the LOCA reanalysis are provided in Tables 3.2-3 through 3.2-6 and Figures 3.2-2a through 3.2-18:. The figures show the following:

1. Peaking Factor vs. Core Height - Figure 3.2-1 shows the cosine power I shape used in the analysis. r j 2. Mass Velocity - Figures 3.2-2a through 3.2-2c show the mass velocity at the clad burst and hot-spot locations on the hottest fuel rod for the discharge coefficient used.

3. Heat Transfer coefficient - Figures 3.2-3a through 3.2-3c show the heat transfer coefficient at the clad burst and hot-spot locations on the hottest rod for the discharge coefficient used. The values of

heat transfer coefficient that are shown were calculated by the l

LOCTA-IV code based on equations for heat transfer in the nucleate boiling, transition boiling, film boiling, and steam cooling regimes.

4. Core Pressure - Figures 3.2-4a through 3.2-4c show the calculated pressure in the core for the discharge coefficient used.

, 5. Break Flowrate - Figures 3.2-Sa through 3.2-Sc show the calculated

) flowrate out of the break for the discharge coefficient used. The flowrate out of the break is plotted a' the sum of flow at both the pressure vessel end and the reactor coolant pump end of the guillotine break.

2156e:1d/041285 3-49

. . - - - - - - - _ ,y.-, y _ . _ . _ , , -

, ,y.- .,,y- , y e

I

6. Core Pressure Orop - Figures 3.2-6a through 3.2-6c show the calculated

(

core pressure drop for the discharge coefficient used. The core pressure drop is interpreted as the pressure inmediately before entering .the core inlet to the pressure just outside the core outlet.

7. Peak Clad Temperature - Figures 3.2-7a through 3.2-7c show the l

calculated hot-spot clad temperature transient and the clad

! temperature transient at the burst location for the discharge coefficient used. The peak clad temperature for the limiting i discharge coefficient of 0.4 is 2160.6*F at the 6.75 ft elevation in the core.

8. Fluid Temperature - Figures 3.2-8a through 3.2-8c show the calculated fluid temperature for the hot spot and burst locations for the f

discharge coefficient used.

9. Core Flow - Figures 3.2-9a through 3.2-9c show the calculated core flow, both top and bottom, for the discharge coefficient used.

f

10. Reflood Transient - Figures 3.2-10a through 3.2-10c show the reactor pressure vessel downcomer and core water levels for the discharge coefficient used. Figures 3.2-11a through 3.2-11c show the core inlet velocity for the discharge coefficient used.

l l

11. Accumulator Flow - Figures 3.2-12a through 3.2-12c show the calculated flow for the discharge coefficient used. The accumulator delivery during blowdown is discarded until the end of bypass is calculated.

Accumulator flow, however, is established in the refill-reflood calculations. The accumulator flow assumed is the sum of that injected in the intact cold legs.

I l

. '12. Pumped ECCS Flow (Reflood) - Figures 3.2-13a through 3.2-13c show the calculated flow of the emergency core cooling system for the discharge i coefficient used.

l l

2156e:1d/041285 3-50

13. Containment Pressure - Figures 3.2-14a through 3.2-14c show the calculated pressure transients for the discharge coefficient used.

The analysis of this pressure transient is based on the data given in Tables 3.2-2, 3.2-5, and 3.2-6.

4 I 14. Core Power Transient - Figures 3.2-15a through 3.2-15c show the core power transient calculated by the SATAN-VI code for the discharge coefficient used.

15. Break Energy Release - Figure 3.2-16 shows the break energy released to the containment for the limiting discharge coefficient of 0.4.

16. Containment Wall Heat Transfer - Figure 3.2-17 shows the containment i wall heat transfer coefficient for the limiting discharge coefficient j of 0.4.

17. Fluid Quality - Figures 3.2-18a through 3.2-18c show the fluid quality

• i at the clad burst and hot-spot locations (location of maximum clad temperature) on the hottest fuel rod (hot rod) for the limiting breaks.

3.2.

1.5 CONCLUSION

S For breaks up to and including the double-ended rupture of a reactor coolant l

pipe, and for the operating conditions specified in Tables 3.2-1 and 3.2-2,

the emergency core cooling system will meet the acceptance criteria as

presented in 10 CFR 50.46, as follows

l 1. 1he calculated peak fuel rod clad temperature is below the requirement

! of 2200*F.

! 2. The amount of fuel element cladding that reacts chemically with water or steam does not exceed 1% of the total amount of Zircaloy in the reactor. I l

1 2156e:1d/112084 68 3-51  ;

3. The clad temperature transient is terminated at a time when the core geomatry is still amenable to cooling. The localized cladding oxidation limits of 17% are not exceeded during or after quenching.

4. The core remains amenable to cooling during and af ter the break.

l

, 5. The core temperature is reduced and the long-term decay heat is l

removed for an extended period of time.

,,l:U)

' o f p ': )

l 1

i 2156e:1d/112084 69 3-52

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

j 3.2.2 SMALL BREAK LOSS-OF-COOLANT ACCIDENT 3.2.

2.1 INTRODUCTION

A reanalysis of the Emergeicy Core Cooling System (ECCS) performance for the postulated small-break LOCA has been performed and is presented in this section. The analytical techniques used in this analysis are in compliance

! with Appendix K of 10 CFR 50 and are described in the topical report, l " Westinghouse Emergency Core Cooling System Evaluation Model Summary" (Ref.

2 3.2-4). 1he results of the analysis are shown in Table 3.2-7, and demonstrate l-compliance with 10 CFR 50.46, " Acceptance Criteria for Emergency Core Cooling Systems for Light Water Reactors."

3.2.2.2 ACCIDENT DESCRIPTION j A LOCA is the result of a rupture of the Reactor Coolant System (RCS) piping or of any line connected to the system.- Ruptures of a small cross section l will cause expulsion of the coolant at a rate that can be accomodated by the l charging pumps. These pumps maintain an operational water level in the pressurizer, permitting the operator to execute an orderly shutdown.

The maximum break size for which the normal makeup system can maintain the pressurizer level is obtained by. comparing the calculated flow from the

, Reactor Coolant System through the postulated break against the charging pump f makeup flow at normal RCS pressure, i.e., 2250 psia. A makeup flow rate from one centrifugal charging pump is typically adequate to sustain pressurizer level at 2250 psia for a break through a 3/8-in.-diameter hole. This break results in a loss of approximately 17.5 lb/sec.

For breaks in the range between 3/8-in. diameter and 1-ft2 area, depres- ,

l i surization of the Reactor Coolant System causes fluid to flow to the Reactor Coolant System from the pressurizer, resulting in a pressure and fluid level decrease in the pressurizer. Reactor trip occurs when the pressurizer low-pressure trip setpoint is reached, and the safety injection system (SIS) 2156e:1d/112084 70

._. . _ . _ _ _ _ _ _ _ _._ __._ _ _-53 3

i l

1s actuated when its appropriate setpoint is reached. The consequences of the accident are limited in two ways:

4 l

1. Reactor trip and borated water injection complement void formation in causing rapid reduction of nuclear power to a residual level corresponding to the delayed fission and the fission product decay.

2. Injection of borated water ensures sufficient flooding of the core to i prevent excessive clad temperatures. I Before the break occurs, the plant is in an equilibrium condition, i.e., the heat generated in the core is being removed via the secondary system. During blowdown, heat from decay, hot internals, and the vessel continues to be transferred to the Reactor Coolant System. The heat transfer between the  !

Eeactor Coolant System and the Secondary System may be in either direction,

, depending on the relative temperatures. In the case of continued heat addi-tion to the secondary, system pressure increases and steam dump may occur.

Makeup to the secondary side is automatically provided by the auxiliary feedwater pumps. 1he safety injection signal stops normal feedwater flow by l

closing the main feedwater line isolation valves and initiates emergency feedwater flow by starting auxiliary feedwater pumps. The secondary flow aids in the reduction,3f Reactor Coolant System pressure. Wien the Reactor Coolant j

System depressurizes to 600 psia, the accumulators begin to inject water into the reactor coolant loops. The reactor coolant pumps are assumed to be tripped at the beginning of the accident, and effects of pump coastdown are l included in the blowdown analysis.

j 3.2.2.3 METHOD 0F ANALYSIS 2

for breaks less than 1.0 ft the WFLASH digital computer code (Ref. 3.2-13) is used to calculated the transient depressurization of the Reactor Coolant System as well as to describe the mass and enthalpy of flow through the break.

4 2156e:1d/112084 71 3-54

l l

The WFLASH program used in the analysis of the small-break LOCA is an

, extension of the FLASH-4 code, (Ref. 3.2-14). The WFLASH program permits a detailed spatial representation of the reactor coolant system by using nodalized volumes interconnected by flowpaths. The broken loop is modeled explicitly with the intact loops lumped into a second loop. The transient behavior of the system is determined from the governing conservation equations of mass, energy, and momentum applied throughout the system.

I The use of WFLASH in this analysis involves, among other things, the representation of the reactor core as a heated control volume, with the associated bubble rise model to permit a transient mixture height calculation. The multi-node capacity of the program enables an explicit and i detailed spatial representation of various system components. In particular i it enables a proper calculation of the behavior of the loop seal during a l loss-of-coolant transient.

4 The SIS flow rate delivery considers pumped injection flow, which is depicted i in Figure 3.2-19, as a function of reactor coolant system pressure. This j figure represents injection flow to the reactor coolant system from the SIS I pumps based on performance curves degraded 55 from the design head. The

- safety injection system is assumed to be delivering to t:1e reactor coolant

, system in 25 sec from the generation of the SIS signal. The 25-sec delay includes the time required for diesel startup and loadirig of the safety l

injection pumps onto the emergency buses. The offeet of residual heat removal l (RHR) pump flow is not considered, since the RHR puA, shutoff head is lower than RCS pressure during the time portion of the transient considered here.

Also, minimum safeguards and emergency core cooling system capability and operability have been assumed in this analysis. Peak clad temperature analyses are performed with the LOCTA IV (REf. 3.2-g) code, which uses as input parameters the RCS pressure, fuel rod power history, steam flow past the uncovered part of the core, and mixture height history as generated by the 2156e:1d/1126g4 3-55 .

I l

WFLASH code. The LOCTA IV code is able to assess whether sufficient heat l transfer geometry and core amenability to cooling are preserved during and after the accident, i

, The analysis of the small-break LOCA is performed at 1025 of the uprated core power rating (2898 Mt). Core power decay versus time after scram is provided in Figure 3.2-20. The axial power shape used to perform the small-break analysis is depicted as curve 1 in Figure 3.2-21. This power shape was chosen j because it provides an appropriate distribution of power versus core height, i.e., linear power is maximized in the upper regions of the reactor core (10 to 12 f t). The selected power shape is skewed to the top of the core, with I the peak linear power occurring at the 10-f t core elevation. The linear power for this power shape at levels above 10 ft essentially matches the shape of the generic operating Fg envelope, which is presented as curve 2 in Figure 3.2-21. Therefore, linear power is maximized for the 10-f t core elevation.

Elevations above 10 ft are limiting for the small-break analysis because of

) the uncovering process for a small break. As the core uncovers, the cladding

! in the upper elevation of the core heats up and is sensitive to the linear power at that elevation. The cladding temperatures in the lower elevations of the core, below the two-phase mixture height, remain low. Because of the

! uncovering process, the peak clad temperature occurs above 10 ft.

i 3.2.2.4 RESULTS l In this analysis, results obtained using the WFLASH code are reported for the '

l 3-in. and 4-in. diameter break sizes. These analyses were perforised during

!. the uprating feasibility evaluation to determine the limiting break size and l peak clad temperature. WFLASH results for the 6-in. break have not been calculated but are expected to be less limiting than the 3-in. and 4-in.

l

break results. The 6-in. break produces more rapid depressurization and accumulator activation, which results in core recovery sooner than for the smaller break cases.

! Break sizes of less than 3 in, are not reported because the core does not significantly uncover and the clad temperature remains low. Break sizes l

l 2156e:1d/112684 3-56

n - - - - - - - - -.. , - - . . . ~ .. - . a -~a . . , -_ - 2,- o.a- .-...=-uas a. -s-z.s, - . - ,

4 9

greater than 6'in. are not reported either, since it is expected that depres-l- surization leading to accumulator activation and core recovery will occur sooner for break sizes greater than 6 in than for break sizes less than or i equal to 6 in. (Ref. 3.2-3). This more rapid depressurization and accumulator activation for break sizes larger than 6 in. keeps the resulting peak clad .

temperature lower, even though the break size is larger. This phenomenon, however, does not necessarily result in limiting peak clad temperatures for l break sizes in the range between 3 and 6 in. (see Table 3.2-7) so that the 3-in. break was found to be the limiting break size (i.e., it hts - highest peak clad temperature and greatest amount of local clad oxidatinn'. The 3-in.

i break is limiting because the depressurization is sufficiently '. Low to allow

! significant loss of coolant through the break prior to accumulator activation, )

f l resulting in more ofithe core being uncovered for a longer period of time. i All break sizes analyzed here or reported in Referrace 3.2-15 provided substa ial margin to the limits specified by 10 CFR 50.46.

l The limiting small break was found to be the 3-in, break. This size break i resulted in a peak clad temperature of 174g*F (including the effects of fuel

! densification), 2.27% local clad oxidation, and less than 0.35 total core clad reaction with the coolant. The core geometry remained amenable to cooling i during and after the break, and long-term cooling capability is not jeopardized, as described in Charter 6 of the FSAR. Significant results are  ;

summarized in Tables 3.2-7 and 3.2-8. l Supporting results for the 3-in. break analysis, including the coolant system depressurization, core mixture height, steam flow rate, hot-spot fluid i

temperature, rod film coefficient, and peak clad temperature, are provided in l Figures 3.2-22 through 3.2-27, respectively. Results depicting the RCS depressurization, core mixture heights, and peak clad temperature for the f

l 4-in. break size is provided in Figures 3.2-2g through 3.2-30.

1 1

_ .__-_ _ 2_156e:1_d _/112.684_ . _ ._ __ _._ _ __ -. 3 _5 7

. . . . . ~ - .- . . . - - - - .-- - .- .-

1 3.2.

2.5 CONCLUSION

S ,

For the small-break LOCA, the emergency core cooling system will meet the acceptance criteria as presented in 10 CFR 50.46, as follows:

1. The calculated peak fuel element clad temperature provides margin to

, the requirement of 2200'F.

2. The amount of fuel element cladding that reacts chemically with water or steam does not exceed 1% of the total amount of Zircaloy in the reactor.

3. The clad temperature transient.is terminated at a time when the core l

i- geometry is still amenable to cooling.g The localized cladding oxidation limits of 17% are not exceeded.

i 4. The core remains amenable to cooling during and after the break.

5. The core temperature is reduced and decay heat is removed for an extended period of time, as required by the long-lived radioactivity remaining in the core.

3.2.2.6 POST-TMI EVALUATION i

l While the analyses described in Sections 3.2.2.2 through 3.2.2.5 are valid j within their range, the March 28, 1979, incident at Three Mile Island Unit 2 exposed the need for further evaluation of small breaks under various j conditions.

Analyses of small-break LOCAs, symptoms of inadequate core cooling and required actions to restore core cooling, and analyses of transient and accident scenarios' including operator actions not previously analyzed, have been performed on a generic basis by the Westinghouse Owners' Group. The small-break analyses were reported in WCAP-9600 (Ref. 3.2-15), which was submitted to the Bulletins and Orders Task Force by the Owners' Group on June 29, 1979.

2156e:1d/112084 75 3-58 s

As a continuation of the above report, a series of LOCA analyses for a range of break sizes and a range of time lapses between initiation of break and pump trip applicable to the 2 , 3 , and 4-loop plants has been performed by the Westinghouse Owners' Group. A report sununarizing the results of the analysis of delayed reactor coolant pump trip during small LOCAs for Westinghouse nuclear steam supply systems was submitted (Ref. 3.2-16) to the NRC on Auguit 31, 1979. In the report, maximum peak clad temperatures for each break sir.e

! considered and pump shutoff times have been provided. The report concludes d that if the reactor coolant pumps are tripped before the reactor coolant

system pressure reaches 1250 psia, the resulting peak clad temperatures do not exceed those reported in Section 5.3.1.4 of the FSAR. In addition, it is shown that there is a finite range of break sizes and reactor coolant pump trip times (in all cases at least 10 min after the event) that will result in i peak clad temperatures in excess of 2200*F as calculated with conservative i Appendix K models. In any event, the operator would have at least 10 min to trip the reactor coolant pumps following a small-break LOCA, especially in l

4 light of the conservatism in the calculations. This is appropriate for manual I rather than automatic action, based on the guidelines for termination of reactor coolant pump operation presented in WCAP-9600.

As a result of NRC Generic Letter 83-10c and 10d (Ref. 3.2-17) (issued in February of 1983), which requested that further analyses be performed in order to more clearly discriminate the need for RCP trip, the Westinghouse Owners Group submitted letter 0G-110 (Ref. 3.2-18) dated December 1,1983. This l

document contained detailed Appendix K 58LOCA, 56TR and non-LOCA transient analysis results in support of establishing alternate RCP trip criteria. The intent of the alternate criteria is to preclude RCP trip in events where

! continued RCP operation is desirable while providing adequate discrimination capability to trip the RCPs in the event of Sas11 treak LOCA events. The l

analysis results demonstrated that the available operator action time to trip l

j the RCPs in the event of a 58LOCA, which could result in peak clad temperatures in excess of 2200*F, was reduced to approximately 6 minutes asst. ming conservative Appendix K LOCA models.

7 As a result of 10 minutes of operator action time not being available, NRC Letter 83-10c and 10d allowed for the use of "Most Probable" Best Estimate SBLOCA analyses to demonstrate that at least 10 minutes was available. The additional work was performed by the Westinghouse Owners Group and submitted to the NRC via letter OG-117 (Ref. 3.2-19) dated March 4,1984. This letter transmitted the analysis results assuming various delayed RCP trip times and break sizes using a bounding plant approach. The results denvanstrated that there were ILg, RCP trip times which would result in peak clad temperatures in excess of the Appendix K limit of 2200*F.

As a result.of these Westinghouse Owners Group letters, the alternate RCP trip criteria were incorporated into Revision 1 of the background documents to the Westinghouse Emergency Response Guidelines (ERGS). Choice of the appropriate RCP trip setpoint will be determined on a plant specific basis using a setpoint which results in discrimination capability between SOLOCAs, SGTRs and

,non-LOCAs. The Westinghouse procedural position on RCP trip for 58LOCA continues to be to trip the RCPs promptly upon acknowledgement of the trip setpoint. Specific guidelines are contained in the ERGS which detail when indeed continued RCP operation during a SBLOCA would be warranted, i.e. ICC conditions.

3-60 2156e:1d/112684

3.2 REFERENCES

3.2-1. Letter from J. R. Miller, NRC, to E. P. Rahe, Westinghouse, dated December 1,1981.

3.2-2. Letter f rom W. L. Stewart, Vepco, to H. R. Denton, NRC, dated December 30, 1982 (Serial No. 726).

3.2-3. R. Salvatori, Westinahouse ECCS Plant Sensitivity Studies, WCAP-8356, July 1974.

3.2-4. F. M. Bordelon et. al., Westinahouse ECCS Evaluation Model -

Summary, WCAP-8339, July 1974.

3.2-5. F. M. Bordelon et al., SATAN-VI Procram: Comorehensive $nace-Time Dependent Analysis of Loss-of-Coolant, WCAP-8306, June 1974.

3.2-6. F. M. Bordelon and E. T. Murphy, containment Pressure Ahalysis Code (C0CO), WCAP-8326, June 1974.

3.2-7. R. D. Kelly et al., Calculational Model for Core Refloodina Af ter a loss-of-Coolant Accident (WREFLOOD Code), WCAP-8171, June 1974.

3.2-8. F. M. Bordelon et al., LOCTA-IV Procram: Loss-of-Coolant Transient Analvsis. WCAP-8305, June 1974.

3.2-9. Letter f rom C. M. Stallings, Vepco, to E. 6. Case, NRC, Serial No.

092, dated February 17, 1978.

3.2-10. Young M. Y. et al., BART-A1: A Connuter Code for the Best Estimate Analvsis of Reflood Transients, WCAP-9695, January 1980.

3.2-11. Chiou, J. S. et al., Models for PWR Reflood Calcult , ;na the BART Code, WCAP-10062, December 1981.

2156e:1d/012385 81 3-61

3.2 REFERENCES

(continued) 3.2-12. Letter from C. O. Thomas, NRC to E. P. Rahe, Westinghouse dated December 21, 1983.

3.2-13. V. J. Esposito, K. Kesavan, and 8. A. Maul, WFLASH-A FORTRAN IV Computer Proaram for Simulation of Transients in a Multi-loon PWR, WCAP-8261 Rev.1. July 1974.

3.2-14. T. A. Porsching, J. H. Murphy, J. A. Redfield, and V. C. Davis, FLASH-4: a Fully Imolicit FORTRAN-IV Proaram for the Diaital Simulation of Transients in a Reactor Plant, WAPD-TM-84, Bettis Atomic Power Laboratory, March 1969.

3.2-15. D. L. Paterline, S. Altomare, and C. Apollo, Report on Sma11-8reak Accidents for Westinchouse NSSS Systems. WCAP-9600, October 1979.

3.2-16. Report submitted by Cordell Reed to D. F. Ross on August 15, 1979.

l 3.2-17. U.S. Nuclear Regulatory Commission generic letter no 03-10c,

" Resolution of TMI Action Item II.K.3.5, ' Automatic Trip of Reactor Coolant Pumps,'" February 8,1983.

3.2-18. Westinghouse Owners Group, " Evaluation of Alternate RCP Trip Criteria " Letter from J. Sheppard, Chairman Westinghouse Owners 6'roup to D. Eisenhut, USNRC. OG-110, December 1,1983.

3.2-19. Westinghouse Owners Group " Justification of Manual RCP Trip for Small 8reak Events," Letter from J. Sheppard, Chairman Westinghouse Owners Group to D. Eisenhut, USNRC.06-117, March 9,1984, l

I l

l

_ _ _ _ _ 2156e:1d/121184 14 3-62 -- _.

l l

l TABLE 3.2-1 l INITIAL CORE CONDITIONS ASSUMED FOR THE l

DOUBLE-ENDED COLD-LEG GUILLOTINE BREAK (DECLG) l 1

Calculational Incut Core power (MWt) 102% of 2898 Peak linear power (kW/ft) 102% of 12.225 Heat flux hot-channel factor (Fg) 2.15 Enthalpy rise hot-channel factor (F"H) 1.55 3

Accumulater water volume (ft , each) 1025 Reactor vessel upper head temperature equal to T hot Limitina Fuel Recion and Cycle Cycle Reaion Unit 1 All All regions Unit 2 All All regions l

i 1

l t

l l 2156e:1d/112084 80 3-63

TABLE 3.2-2 CONTAINMENT DATA Net free volume 1.916 x 106 gg3 Ic.itial conditionsa Pressure. 9.6 psia Temperature 90'F RWST temperature 35'F Outside temperature -10'F l

Spray systen#

Number of pumps operating 2 l Runout flowrate (per pump) 2000 gpm Time in which spray is effective 59 see Structural heat sinksa Thickness (in.) Area (ft 2 ), with allowance for uncertainties 6 concrete 8,393 12 concrete 62,271 18 concrete 55,365 24 concrete 11,591 27 concrete 9,404 36 concrete 3,636

.375 steel, 54 concrete 22,039

.375 steel, 54 concrete 28,933

.500 steel, 30 concrete 25,673 26.4 concrete, .25 steel, 120 concrete 12,110

.407 stainless steel 10.527

.371 steel 160,326

.882 steel 9,894

.059 steel 60,875 aSee Section 6.3.3.12 of the FSAR for a detailed breakdown of the containment heat sinks and for justification of the other input parameters used to calculate containment pressure.

i 2156e:1d/112084 81 l >M l

TABLE 3.2-3 TIME SEQUENCE OF EVENTS DECLG DECLG DECLG CD = 0.4 CD = 0.6 CD = 0.8 (sec) (sec) (sec)

I Start 0.0 0.0 0.0 Reactor trip 0.64 0.62 0.61 Safety injection signal 2.12 1,69 1.46 Accumulator injection 15.70 11.90 9.76 Pump injection 27.12 26.69 26.46 ,

End of bypass 31.72 26.65 23.82 End of blowdown 31.72 26.65 23.82 Bottom of core recovery 45.25 39.74 37.03 Accumulator empty 55.94 51.06 48.27 i

TABLE 3.2-4 RESULTS FOR DECLG CD = 0.4 CD = 0.6 CD = 0.8 ,

Peak clad temperature, *F 2160.6 2013.7 1829.7 Peak clad location, ft 6.75 7.25 7.25 Local Zr/H O reaction 2

(max),% 6.69 3.77 1.92 Local Zr/H O location, ft 6.0 6.75 7.25 2

Total Zr/H O reaction, 5 <0.3 <0.3 <0.3 ,

2 Hot-rod burst time, sec 40.70 66.00 73.00 Hot-rod burst location, ft 6.0 6.75 6.775

(

2156e:1d/041285 3-65

TABLE 3.2-5 REFLOOD MASS AND ENERGY RELEASES DECL6 (CD = 0.4)

Total Mass Total Energy Time-(sec) Flow Rate (1b/sec) Flow Rate (105 8tu/sec) 45.252 0.0 0.0 46.427 0.013 0.000171 56.412 87.30 1.059 72.137 209.96 ,

1.382 91.587 261.33 1.457 113.137 270.53 1.413 137.687 294.51 1.406 200.937 317.58 1.360 TABLE 3.2-6 BROKEN LOOP ACCUMULATOR FLOW TO CONTAINMENT DECLG (CD" **'

Time (sec) Mass Flow Ratea (1bm/sec) 0.00 4010.1 1.01 3622.2 3.01 3105.1 5.01 2762.6 7.01 2510.3 10.01 2226.3 15.01 1898.9 20.01 1674.6 25.01 1516.9 29.01 1575.3 "For energy flowrate, multiply mass flow rate by a constant of 59.60 Btu /lbm.

b For energy flowrate at this time, multiply mass flowrate by 54.09 Btu /lbs.

2156e:1d/112084 83 3-66

1.

TABLE 3.2-7 ASSUMPTIONS AND RESULTS FOR SMALL BREAK Break Size (effective diameter) 3 in. 4 in.

, Results Peak clad temperature. *F 1749.4 1517.0 Peak clad location, ft 11.50 11.0 Local Zr/H O reaction (max.), % 2.27 0.55 2

Local Zr/H 0 location, ft 11.00 11.00 2

Total Zr/H O reaction, % <0.3 <0.3 2

Hot-rod burst time, sec Hot-rod burst location, ft No Burst No 8urst Calculation Core power (102% of 2898), MWt 2956 Average core linear power (102% of 5.69), kW/ft 5.80 Peaking factor (at license rating) See Figure 3.2-21 3

Accumulator water volume, ft / accumulator 1025 Fuel region - cycle analyzed Cycle Region i Unit 1 All All Unit 2 ,

All All i

i l

l '2156e:1d/112084 84 3-67

TABLE 3.2-8 TIME SEQUENCE OF EVENTS FOR SMALL 8REAK Time After Start of LOCA (sec)

Break Size (effective diameter) 3 in. 4 in.

Event Start 0.0 0.0 Reactor trip signal 24.72 16.69 Top of core uncovered 526.3 272.0 Accumulator injection begins 1271.8 641.2 Peak clad temperature occurs 1376.5 669.9 Top of core covered 2957.5 2400.6 i

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SECTION 4 NSSS SYSTEMS REVIEW s 4.1 BASIS OF EVALUATION NSSS systems designs have been reviewed to verify that they will remain in compliance with functional requirements specified in the FSAR when the North Anna units are operated at the increased power rating. That review was performed in accordance with the following guidelines:

1. The review encompassed all aspects of the North Anna NSSS design and j operation that were impacted by the power increase.

2. The review was performed in accordance with licensing criteria and standards that currently apply to the North Anna units.

3. Current techniques have been used for those analyses requ red in the 4

course of the NSSS review. These are the criteria put forth in WCAP-10263 that apply to the rtview of the design and operation of NSSS systems and components. 4.2 SYSTEMS EVALUATION 4.2.1 REACTOR COOLANT SYSTEM I Review of the Reactor Coolant System design and operation at the increased

power demonstrated that functional requirements for operation at 2910 MWt are 4

within the original design capability of the system. The following specific 1 aspects of design were reviewed:

1. Operating parameters Reactor Coolant System design and operating parameters for 2910 MWt operation are presented in Table 4.2-1. This table contains the same operating information that is presently given in FSAR Table 5.1-1.

4 2156e:Id/112084 86 4-1

Parameters for uprated power conditions have been compared to current l operating parameters in Section 2 of this report. ) 2. Design Transients i A review of operating transients used in the original plant design for normal, upset, emergency, faulted and test conditions has verified f ! that the original transients remain applicable at the uprated power. j These transients along with the Reactor Coolant System parameters , j contained in Table 2-1 constitute two of the fundamental design inputs j to the primary system equipment evaluations presented later in Section i 4.3. l 3. Reactor Control and Protection Systems Based on Reactor Coolant System operating parameters for the uprated power conditions, studies were performed to assess operating margin and control system capability. The capat,ility of the NSSS control

 !                systems (e.g., rod control, steam dump, pressurizer pressure and level j                  control) was found to be adequate for operation at the uprated power.

l Control system setpoints have been revised for operation at the j uprated power based on these studies. Accident analyses documented in , j Section 3 confirm the capability of the Reactor Protection System for operation at 2910 MWt. Reactor Protection System setpoint revisions are included in Section 7. Technical Spectftcation Revisions. , l! - 4 i . 4. Piping Pressure Losses i lhe maximum calculated surge rate at uprated power conditions is less ) than the flow assumed in the original North Anna design. Therefore, f the surge line is adequately sized for operation at the higher power

               - conditions. Review of the pressurizer safety valve and power operated l

relief valve design requirements indicated that the existing capacity j 4 is adequate for operation at uprated conditions. The existing safety i valve discharge piping, which is sized for the prtstntly installed f I safety valve capacity, is therefore also adequate for operation at the , 3 uprated power. There was no need to evaluate the F'.T0 bypass system or l i l ! 2156e:1d/112004 47 l 4-2

the pressurizer spray line piping since the steady state flows and Reactor Coolant System pressure drops are virtually unaf fected by the

uprating.

4.2.2 RESIDUAL HEAT REMOVAL SYSTEM i Review of Residual Heat Removal System functional design showed that operation at the uprated power is within the capability of the existing system. System  ; l l design parameters for operation at 2910 MWt are given in Table 4.2-2. This i table contains the same data as FSAR Table 5.5-8. The only change has been an { increase in the decay heat generation 20 hours after reactor shutdown to reflect the higher full power rating. Cooldown calculations have shown that the Residual Heat Removal System is capable of reducing reactor coolant ' temperature from 350'F to 140*F over a period of 16 hours at 2910 MWt operation. Cooldown calculations also showed that safe shutdown of the plant l 1s not compromised at 2910 MWt operation if one of the two pumps nr one of the ! two heat exchangers is not operable. Component Cooling System heat loads have been revised based on the cooldown calculations performed at 2910 MWt. I Calculations indicate that reload cycles with core average burnup of 16,500 mfd /MTU per cycle have very little offeet on residual decay heat eenerated by the core during the time period that plant cooldown occurs. Over the time span from one hour post-shutdown to forty-eight hours post-shutdown, decay ^ heat generation rates for cycles of extended burn-up are only a few percent greater than the decay heat generation rates used to establish the cooldown l capability reported above. Those dif forences are within the margir of ! conservatism allowed in the cooldown calculation. As a result, it has been concluded that cocidown capability of the existing residual heat reuval 4 system is adeguate for operation at uprated power with a core averagt burnup i of M.500 IRS /NTU per cycle, i Siece the terrent system functional requirements are met, no hardware changes ! see esgsteed for operation at the increased power rating. Data given in FSAR ! Tames 5.fr-9 and 5.5-10 remain as currently given in the FSAR. 1 l J 21 %e:ld/139484 4_3 I

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

4.2.3 . CHEMICAL AND VOLUME CONTROL SYSTEM d The increased power rating has a minor impact on Chemical and Volume Control l System operation since the temperature of reactor coolant entering the system is slightly lower at 2910 MWt (552.3*f) than at the current 2787 MWt $ conditions (555.5'F). As a result, heat loads on the regenerative, non-regenerative, excess letdown, and seal water heat exchangers are reduced slightly for the higher power conditions. Obviously, the currently licensed

conditions bound the uprated power conditions. Table 4.2-3 contains revised heat exchanger parameters comparable to those contained in FSAR Table 9.3-5.

Component Cooling System heat loads have been revised to reflect revisions in the system heat exchanger parameters. Other than the changes in heat exchanger parameters described above, no f changes in Chemical and' Volume Control System functional design or operation ! are required for operation at the uprated power. 1 4.2.4 EMERGENCY CORE COOLING SYSTEM l ? l The Emergency Core Cooling System is designed to cool the reactor core as well as to provide additional shutdown capability following initiation of certain accident events described in Chapter 15 of the FSAR. Those Chapter 15 accident events that are impacted by the power uprating have been reanalyzed l as described in Section 3 assuming the same Emergency Core Cooling System capability that currently exists in the North Anna plants. Since the results of those reanalyses meet acceptance criteria provided in Chapter 15, it ! follows that the existing Emergency Core Cooling System design is adequate for

operation of the North Anna units at an NSSS power rating of 2910 MWt.

I i l 4.

2.5 CONCLUSION

S ! Review of NSSS systems designs and operating capability has verified that they l will remain in compliance with the functional requirements specified in the FSAR when the North Anna Units are operated at the increased power rating. A l l 2156e:1d/lf2064 89 4-4 l 4J_, - . .-, - , - - - -

1 l few system parameters and setpoints have been revised to reflect operation at

     .the higher power, but those operating conditions have been shown to be within the design capability of the systems as they currently exist.

J l i i 4 l

v.

( l l 2156e:1d/112084 90 l 4-5

TABLE 4.2-1 REACTOR COOLANT SYSTEM DESIGN AND OPERATING PARAMETERS Plant design life, years 40 Nominal operating pressure, psig 2235 Total system volume including pressurizer and surge line, ft 3 9957.2 System liquid volume, including pressurizer water at maximum guaranteed power, ft 3 9380.4 Total nuclear steam supply system heat output at full power, 8tu/hr 9929.3 x 106 Total coolant flow rate, lb/hr 104.3 x 106 System thermal and hydraulic data Reactor vessel Inlet temperature, 'F 552.3 Outlet temperature. *F 621.2 AP, psia (at T = 552.3*F) 52.6 Steam generator Inlet temperature, 'F 621.2 Outlet temperature, 'F 552.0 AP, psia (at T = 552.3*F) 34.6 Design fouling factor 0.000055 Piping AP, psia (at T - 552.3'F) 12.6 ! Reactor coolant pump t Inlet temperature, 'F 552.0 Outlet temperature 'F 552.3 Developed head, psia (at T - 552.3*F) 99.8 Developed head, ft 312 Flow (each), gpm 92,800 Steam pressure at full power, psia 850 Steam flow at full power, Ib/hr (total) 12.78 x 106 feedwater inlet temperature 'F 440 Pressurizer spray rate, maximum, gpm 600 Pressurizer heater capacity, kW 1400 Pressurizer relief tank volume, ft3 1300 f 2156e:Id/112004 91 4-6

1 TABLE 4.2-2 DESIGN BASES FOR RESIDUAL HEAT REMOVAL SYSTEM OPERATION Parameter Value Residual ' heat removal system startup 4 hr after reactor shutdown Reactor coolant system initial pressure, psia approximately 450 Reactor coolant system initial temperature, 'I approximately 350 r Component cooling water design temperature. *F 105 Cooldown time, hours after initiation of residual approximately 16 heat removal system operation Reactor coolant system cold temperature, 'F 140 Decay heat generation at 20 hr after reactor 64 x 10 0 shutdown, 8tu/hr t l 4 2156e:Id/112004 92 4-7

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

TA8LE 4.2-3 CHEMICAL AND VOLUME CONTROL SYSTEM

;                            HEAT EXCHANGER PARAMETERS FOR 2910 MWT Reaenerative Heat Exchancer Number                                                    1 Heat transfer rate at design conditions, 8tu/hr                                    8.85 x 106 Shell side

Design pressure, psig 2485 Design temperature, 'F 650 i Fluid Borated reactor coolant Material Austenitic stainless steel i

Tube side Design pressure, psig 2735 - Design temperature, 'F 650 ! Fluid Sorated reactor coolant ! Material Ai.stenitic stainless steel Shell side (letdown)

Flow Ib/hr 29.820 Inlet temperature, 'F 555.5 Outlet temperature. 'F 291 Tube side (charging)

! Flow, Ib/hr 22,370 Inlet temperature 'F 130 l Outlet temperature. 'F 505

Nonrenenerative Heat Exchanner Number 1 Heat transfer rate at design i corditions, 8tu/hr 16.08 x 100

[ Shell side l Design pressure, psig 150

Design temperature. 'F 250 l Fluid Component cooling water i

Material Carbon steel Tube side ( Design pressure, psig 600 l Design temperature, 'F 400 l Fluid Borated reactor coolant [ Material Austenitic stainless steel 2156e:Id/112004 93 4-8

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

1 TA8LE 4.2-3 (Cont'd) l CHEMICAL AND VOLUME CCNTROL SYSTEM

HEAT EXCHANGER PARAMETERS FOR 2910 MWT t

Nonrenenerative Heat Exchanner (Cont'd)

!         Shell side 9111RE

Flow, Ib/hr 537,100 Inlet temperature. 'F 105

;                 Outlet temperature, 'F                               135 l               Normal l                 Flow, Ib/hr                                          81,250 j                 Inlet temperature. *F                                105 l                 Outlet temperature, 'F                               170 Tube side (letdown)

} E111RE t Flow, Ib/hr 59,700 4 Inlet temperature, 'F 380 < Outlet temperature. 'F 115

                       , Ib/hr                  -                      29,820 l                  Inlet temperature, 'F                                291 i                  Outlet temperature. 'F                               115
                                                                           +

i i & j Excess-Letdown Heat Exchanaer i Number 1

!         Heat transfer rate at design j               conditions. Otu/hr                                      3.27 x 106

Shell side Design pressure, psig 150 Design temperature. 'F 250

Shell side (Cont'd) i Oesign flow, Ib/hr 83,500 i

Inlet temperature. 'F 105 Outlet temperature, 'F 144 fluid Component cooling water

Material Carbon steel 1

I l (' i 21%e:1d/112004 94 4-9

i l TABLE 4.2-3 (Cont'd) CHEMICAL AND VOLUME CONTROL SYSTEM HEAT EXCHANGER PARAMETERS FOR 2910 MWT Excess-Letdown Heat Excl.anaer (Cont'd) Tube side

Design pressure, psig 2485 Design temperature 'F 650 Design flow, Ib/hr 7500 Inlet temperature, 'F 555.5 Outlet temperature. *F 153 ,

Fluid Borated reactor coolant Material Austenitic stainless steel Seal-Water Heat Exchanaer Number 1 1 Heat transfer rate at design conditions Btu /hr 1.37 x 106 i Shell side j Design pressure, psig 150 i Design temperature, 'F 250 j Design flow Ib/hr 180,000 Inlet temperature, 'F 105 , Outlet temperature, 'F 118 , Fluid Cemponent cooling water i Material Carbon steel < lube side Design pressure, psig 150 Oesign temperature, 'F 250 Design flow, Ib/hr 111,600 Inlet temperature, 'F 133 Outlet temperature, 'F 121 Fluid Berated reactor coolant l Material Austenttic stainless steel I i l 2156e:1d/112084 95 ( 4-10 l

't 4 SECTION 5 NSSS CONPONENTS INPACT

    ^ 5.1 BASIS FOR EVALUATION T,1e mechanical design of NSSS equipment has been reviewed to assure that structural integrity of the plant will be maintained under conditions specified in the FSAR when the North Anna units are operated at the uprated power. The review was performed in accordance with the following guidelines:

1. The review encompassed all aspects of the North Anna NSSS equipment mechanical design that were impacted by the power increase.

2. The review was performed in accordance with licensing criteria and standards that currently apply to the North Anna Units.

j 3. Equipment mechanical designs were evaluated against the original

industry design codes and standards to which the equipment was built. <

4. Current techniques have been used for those analyses required in the course of the NSSS equipment review.

These are the criteria put forth in WCAP-10263 as a basis for reviewing the mechanical design of NSSS equipment during an uprating safety evaluation. 5.2 EQUIPNENT REVIEWS ! 5.2.1 REACTOR VESSEL To assess the impact of the uprating on the reactor vessel design and operation, the vessel stress report, and fracture mechanics analyses were I reviewed. This review was based on the following information developed during the review of other components:

1. The original pipe reactions analyzed on the reactor vessel bounded those at uprated conditions, i 1804e:1d/091984 '*

F

2. The original reactor internals interface loads on the vessel bounded i those at uprated conditions, i

i

3. The design transients were not modified except for the plant loading temperatures being increased.

The reactor vessel stress report was reviewed to determine the effect of the uprating on the stresses and usage factors of the ten critical regions of the vessel that have been analyzed. The results of these analyses showed that the i maximum stress intensity ranges on the reactor vessel, including the of f act of . the plant loading transient change, were within code allowable limits. l 1

Reactor vessel fracture mechanics were reviewed on the following basis.

i i 1. The orig'nal end-of-life fluence calculation was based on a core power level that bounded the uprated core power. f 2. The design transients were not modified except for the plant loading

temperatures being increased.  ;

  ,              3. Generic fracture mechanics evaluations have been performed at a power i                    level that bounds the uprated power to evaluate the offeet of large I

steam break, large'LOCA, and small LOCA transients on reactor vessel integrity. These evaluations indicate that the proposed power j increase will not significantly change the results of the reactor j vessel integrity evaluation. Review of the reactor vessel fracture mechanics has indicated that operation ! at the uprated power would have no more significant ef fect on reactor vessel integrity than operation at power conditions for which the plant was originally designed. I i i l Ig04e:1d/091984 5-2

3 5.2.2 REACTOR INTERNALS - 1 Operation at the uprated power conditions can potentially affect the hydraulic  ! loads on reactor internals components, the propensity for flow induced vibration, thermal gradients induced within and between components, rod cluster control assembly scram time, and core bypass flows. For each of these j aspects of design, evaluations and/or analyses were performed to verify that the existing reactor internals design remains in compliance with the current FSAR design requirements when operating at the uprated power. Since the design transients for the uprated power are bounded by those used for the original reactor internals design, only components which are directly influenced by the core radiation heat generation need be structurally evaluated. These components are the baf fle-barrel-former region, the upper f core plate and the lower core plate. Results of the reactor internals evaluations and analyses are sammarized as follows: { ! 1. There is no impact on the core bypass flow. ', 2. Because of the decrease in fluid density at the uprated conditions, l hydraulic lift forces are bounded by those at the conditions used in the original reactor internals design. Therefore, hydraulic loadings on the internals at the uprated power are bounded by those in the original design.

3. The potential for flow induced vibrations is not increased.

4. Stresses and fatigue usage factors for components in the i baffle-barrel-former region of the internals are bounded by the original analysis.

5. Stresses and fatigue usage factors for the upper and lower core plates are bounded by analyses performed for other plants using the same core plate design.

6. There will be a slight decrease in rod cluster control assemb'ly scram time when compared to the original plant design conditions, which remain bounding.

le04e:Id/012305 5-3 l 1

Based on these evaluations and analyses, the North Anna reactor internals remain in compliance with the requirements of all applicable design criteria defined by the FSAR. 5.2.3 CONTROL R00 ORIVE MECHANISMS Review of the control rod drive mechanism design showed that operating conditions for 2910 MWt operation are bounded by the original thermal and structural design analyses. 5.2.4 REACTOR COOLANT PUMPS e Review of the reactor coolant pump design showed that operating conditions for 2910 MWt operation are bounded by the original thernal and structural design analyses. 5.2.5 STEAM GENERATORS Modifications to the model 51' series steam generator stress report have been made on a generic basis using a set of operating parameters that bound con-ditions for the North Anna units at 2910 MWt operation. The ASNE-Boiler and Pressure Vessel Code, Section III,1968 edition was used to determine

acceptable states of stress for the components. The generic evaluation established that the model 51 series steam generator stress report satisfies all applicable ASME Code requirements when updated to the enveloping set of plant operating parameters. An appendix to that report shows that the North Anna operating parameters for 2910 MWt are conservatively bounded by the enveloping parameters of the generic report.

5.2.6 PRESSURIZER l Review of the pressurizer design showed that operating conditions for 2910 MWt ! operation are bounded by the original thermal and structural design analyses. 1904e:1d/091984 5-4

 ~
                             ._-                                    - .n           -.

As indicated in item 4 of Section 4.2.1, review of the Reactor Coolant System , design has established that the existing pressurizer safety valves and power I

operated relief valves are adequate for operation at the uprated conditions.

5.2.7 REACTOR COOLANT PIPING ! A thermal reanalysis was made of the reactor coolant loop, reactor coolant loop bypass, and pressurizer surge line piping for temperatures which envelope the 2910 MWt operating conditions. Result of that analysis showed that piping stresses and support loads at the uprated conditions are within engineering tolerances of the values originally calculated. 5.2.8 LOOP STOP VALVES Review of the main loop stop valves design showed that operating conditions for 2910 MWt operation are bounded by the original thermal and structural , design analyses. 5.2.9 REACTOR COOLANT SYSTEM SUPPORTS Uprating of the North Anna units to 2910 MWt has a negligible impact on reactor coolant loop equipment supports when compared to the original support design. The loop expansion effects due to the changes increase in hot leg / cold leg temperatures are within the tolerance band for shim measuring / machining, and therefore have no effect on the supports design. It was l established in Section 5.2.7 that design loads on the supports would not change. i 5.2.10 - AUXILIARY SYSTEMS COMPONENTS l The only auxiliary systems components impacted by the uprating are the , regenerative, non-regenerative, seal water and excess letdown heat exchangers l in the Chemical and Volume Control System. A review of the structural design ! of these components showed that in each case the conditions used in the

     -original design enveloped those required for operation of the North Anna units i        at 2g10 Inft.

l i I 1804e:1d/091984. 5-5

5.3 CONCLUSION

S l l ) Review of North' Anna equipment designs that are impacted by the power uprating ) has shown that in most cases requirements for operation at the higher power l are enveloped by either the original North Anna design, or by the generic component design. In a' few cases, it has been necessary to perform additional design calculations to verify the capability of a component for operation and

       . compliance with the original design codes and standards at the uprated conditions. In every case, however, it has been shown that the NSSS equipment originally provided for the North Anna units is capable of operation at 2910 MWt without modification.

L i l lt04e:1dI091984 5-6

1 SECTION 6 NSSS/ BOP INTERFACES

6.1 INTRODUCTION

To coordinate the NSSS review with the Balance of Plant (80P) review, a program was established to examine plant design data in those areas where the power uprating could have an impact on the 80P design. This section presents the results of that evaluation. 6.2 MASS AND ENERGY RELEASE DATA The original Loss of Coolant Accident data for containment integrity evaluations was based on an NSSS power rating of 2910 MWt with a steam pressure of 850 psia. This data remains applicable since these are the uprated power conditions. Data for the Main Steamline Break analysis was based on the event occurring at no load conditions, which are not changed by the power uprating. Therefore, this data also remains applicable for the power uprating evaluation. 6.3 AUXILIARY FEE 0 WATER SYSTEM The original Auxiliary Feedwater System design requirements were based on an NSSS rating of 2910 MWt with a steam pressure of 850 psia. This data remains applicable since these are the uprated conditions. 6.4 RADIATION SOURCE TERMS e Bata currently in the FSAR are based on a core power rating of 2900 MWt. These source terms are essentially a function of power level and burnup only. This evaluation proposes a change incore power level to 2898 MWt, and operation such that the batch average fuel burnup at discharge is less than 45,000 MMD/NTU. In reference 6-1, NRC provided approval for operation of - North Anna and Surry units with batch average . fuel burnups not to exceed this 1904e:1d/121184 6-1

value. Vepco had requested extension of batch average burnups 45,000 MWD /MTU in reference 6-2. Reference 6-2 concluded that WCAP-10125, " Extended Burnup Evaluation of Westinghouse Fuel" (Reference 6-3) allows for operation of Westinghouse fuel to such burnups. This conclusion remains valid for operation at the proposed uprated conditions. Reference 6-1 concluded that radiological consequences from operation at extended burnup would be bounded by the existing North Anna FSAR results. This was based upon independent NRC Staff calculations and supplemental information supplied by Vepco via references 6-4 and 6-5. The supplemental information was based upon 4.1 weight percent U-235 fuel, operated at a core power level of 2900 MWt. The NRC approval of reference 6-1 stated that their evaluation assumed there will be no increase in peak linear heat generation rate for the North Anna or Surry units. Vepco has reviewed the corresondence-and concluded that the small increase in linear heat generation rate associated with the proposed uprating will not invalidate the conclusion's of the previous evaluations. It is concluded from these results that the 2 radiological ef fects reported in the UFSAR will remain bounding for the proposed extended burnup operation at a core power of 2898 MWt. 6.5 REACTOR COOLANT SYSTEM PIPING DESIGN DATA A thermal reanalysis of the reactor coolant loop piping was performed for the uprating evaluation (see Section 5.2.7). Results demonstrated that piping loads and thermal displacements at the uprated conditions are within i . ' engineering tolerances of the loads provided to the Architect Engineer during the original North Anna design. Therefore, the original dasign data remain applicable for the uprating condition. 6.6 COMPONENT C0OLING SYSTEM HEAT LOAOS l Heat loads imposed on the Component Cooling System by NSSS equipment have been

;             revised as appropriate for operation at 2910 Mwt. Chemical and Volume Control System heat exchanger heat loads have been revised consistent with parameters I

f i 1804e:1d/121184 6-2

     . _ _             . ..                      L  _         z -      ,    . . - . _:          - . -    .

i

      -given in Table 4.2.3. The increase heat load for these three heat exchangers was less than 15 of the original Component Cooling System capability.

Residual heat exchanger heat loads have been revised to reflect the results of cooldown calculations made during the power uprating evaluation. Decay heat data was provided to-the Architect Engineer as input to the Spent Fuel Pit

• I Cooling System review in the 80P uprating evaluation. Heat load increases on the Residual Heat Removal System and the Spent Fuel Pit Cooling System were 4.5% or less since heat input by those systems is largely dependent on power level.

1 6.7 STEAM SYSTEM DESIGN TRANSIENTS The Steam System Design Transients provided for the original North Anna design i are unchanged, and remain applicable at the uprated power conditions. 6.8 TURBINE GENERATOR f Based on a detailed review, it was determined that the assumptions, analyses, and evaluations (e.g., turbine missiles) performed to verify the operating characteristics and structural integrity of the turbine generator bound the operating conditions for 2905 Mwt, which is the maximum calculated turbine rating. i i 3 I 1904e:1d/012385 6-3

l

6.0 REFERENCES

6-1. Letter from Steven A. Varga, NRC, to W. L. Stewart, Vepco dated April 9, 1984. 6-2. Letter f rom W. L. Stewart, Vepco, to H. R. Denton, NRC, Serial No. 678, dated November 23, 1983. 6-3. P. J. Kersting, et. al., " Extended Burnup Evaluation of Westinghouse Fuel " WCAP-10125, July 1982. 6-4. Letter from B. R. Sylvia, Vepco, to H. R. Denton, NRC, Serial No. 195, dated March 26, 1981. 6-5. Letter from R. H. Leasburg, Vepco, to H. R. Denton, NRC, Serial No. 432, dated July 24, 1981. l i l l l i > I l h

  ' ,-1804e:1d/121184                            6-4
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