ML19093A558

Submit Large Break LOCA-ECCS Reanalysis

ML19093A558
Person / Time
Site:	Surry
Issue date:	08/26/1977
From:	Virginia Electric & Power Co (VEPCO)
To:	Office of Nuclear Reactor Regulation
References
Download: ML19093A558 (59)

Text

LARGE BREAK LOCA-ECCS REANALYSIS SURRY POWER.STATION UNIT #1 NOTICE -

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RECORDS FACILITY BRANCH

RETUR" TO REGULATORY CENTRAL fl,lfS.

ROOM 016 Large Break LOCA-ECCS Reanalysis Surry Power Station Units No. 1 and 2 August 26, 1977.

RETURN TO REGULATORY CENTRAL FliLES ROOM 016 _,- *

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INTRODUCTION The reanalysis of ECCS cooling performance for the postulated large break Loss of Coolant Accident (LOCA) has been required in order to allow operation of Surry Units No. 1 and 2 at higher levels of steam generator tube plugging.

The results of this reanalysis are presented herein.* This reanalysis is in compliance with iocFR50.46(1), Acceptance Criteria for Emergency Core Cooling Systems for Light Water Reactors.

This reanalysis was performed with the October, 1975 version of the Westinghouse Evaluation Model.

The analytical techniques utilized in the reanalysis are in compliance with Appendix K to 10CFR50 and the August 27, 1976 Order for Modification of License.

(See References 2, 13, 14, 15, 16, 17, and 18).

As required by Appendix K of 10CFR50, certain conservative assumptions were made for the LOCA analysis.

The assumptions pertain to the conditions of the reactor and associated safety system equipment at the time that the LOCA occurs, and include such items as the core peaking factors, the contaiment pressure, and the performance of the emergency core *cooling system (ECCS).

All assumptions and initial operating condition input data used in this reanalysis were the same as was used in the previously applicable LOCA-ECCS analysis (see our letter of October 29, 1976-Serial No. 219/082776 as supplemented by our letter of March 4, 1977-Serial No. 219/082776) except for (1) the limiting value of heat flux hot channel factor (changed from 2.0 to 1.85), (2) the minimum allowable value of the RCS flowrate (changed from 88,500 gpm/loop to 79,650 gpm/loop) (3) the number of steam generator tubes assumed to be plugged (changed from 20% to 25%),

(4) the elimination of the +4°F uncertainty applied to the core inlet temper-

ature, (5) the maximum allowable time to flow initiation of the containment spray system changed from 20 seconds to 46 seconds), (6) the use of a flow initiation

• It should be noted that a reanalysis of the small break LOCA is not necessary, and therefore the analysis of this accident subm1tted in our letter of June 6, 1975 (Serial No. 500-s)(8) remains applicable.

time for the inside a~d outside recirculation spray system of 120 seconds and 300 seconds, respectively, and (7) the use of a reduced hot assembly enthalpy rise peaking factor (changed from 1.435 to 1.38).

2.0 DESCRIPTION

OF POSTULATED MAJOR REACTOR COOLANT PIPE RUPTURE (LOSS OF COOLANT ACCIDENT -

LOCA)

A LOCA is the result of a rupture of the Reactor Coolant System (RCS) piping or of any line connected to the system.

The boundary considered for the LOCA as applicable to this connected piping is defined in the FSAR.

Should a major break occur, depressurization of the RCS results in a pressure decrease in the pressurizer.

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.

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 fission product decay heat.

(It should be noted, however, that no credit is taken in the analysis for the ins~rtion of.. control rods to shut down the reactor.)

2.

Injection of borated water provides heat transfer from the core and prevents excessive clad temperatures.

Before the break occurs, the unit 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.

At the beginning of the blowdown phase, the entire RCS contains subcooled liquid which trans-fers heat from the core by forced convection with some fully developed nucleate boiling.

After the break develops, the time to departure from nucleate boiling is calculated, consistent with Appendix K of 10CFR50.

Thereafter the core heat transfer is based on local conditions with transition boiling and forced e

convection of 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.

The heat transfer between the reactor coolant system and the secondary system may be in either direction depending on the relative tereper-atures.

In the case of continued heat addition to the secondary, secondary system pressure increases and the main safety valves may actuate to reduce the pressure.

Make-up to the secondary side is automatically provided by the*

auxiliary feedwater system.

The safety injection signal stops nonnal feed-water flow by closing the main feedwater control valves, trips the main feedwater pumps and initiates emerge~cy feedwater flow by starting the auxiliary feed-water pumps.

The secondary flow aids in the reduction of reactor coolant system 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 K of 10CFR50.

In addition, the reactor coolant pumps are assumed to be tripped at the initializationof the accident and effects of pump coastdown are included in the blowdown analyses.

The water injected by the accumulators cools the core and subsequent operation of the low head safety injection pumps supply water for long term cooling.

After the contents of the refueling water storage tank is emptied-; long

~erm cooling.of the core is accomplished by switching to the recirculation mode of core cooling, in which the spilled borated water is drawn from the containment sump by the low head safety injection pumps and returned to the reactor vessel.

The containment s~ray system and the recirculation spray system operate to return the containment to subatmospheric pressure.

'r The large break LOCA transient is divJded, for analytical purposes, into three phases:

b.lowdown, refill, and reflood.

There are three distinct transients analyzed in each phase, including the thennal-hydraulic transient in the RCS, the pressure and temperature transient within the contai~~ent, and the fuel and clad temperature transient of the hottest fuel rod in the core.

Based on these considera~ions, a system of inter-related computer codes has been developed for the analysis of the LOCA.

The description of the various aspects of the LOCA analysis method-ology is given in WCAP-8339. (Z)

This document describes the major phenomena modeled, the interfaces among the computer codes, and the features of the codes which ensure compliance with the Acceptance Criteria.

The SATAN-VI, wr,.EFLOOD, COCO, and LOCTA-IV codes, which are used in the LOCA analysis, are described in detail in WCAP--8306 (J), WCAP-8171 (S), WCP...P-8326 (6), and WCAP-8305 (4), respe.c-tively.

These codes are *able to assess whether sufficient heat transfer

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-geometry and core amenability *to cooling are preserved during the time spans applicable to the blowdown, refill, and re.flood phases of the LOCA.

e The SATAN-VI computer code analyzes the thermal-hydraulic transient in the RCS during blowdown, and the WREFLOOD computer code is used to calculate this tran-

. s-:!.ent during the refill and reflood phases of the accident.

The COCO* computer code is used to calculate the containment pressure transient during all three phases of the LOCA analyais.

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 RCS and steam gen-erator 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 break mass and energy flow rates that are assumed to be vented to the containment during blowdown.

At the end of the blowdown phase, these data are* transferred

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co the WREFLOOD C:ode, Also at the end of blo,.,rdown, the mass and energy release rates during blowdown are transferred to the COCO code for use in the deter-mination of the containment pressure response during this first phase of the LOCA.

Additional SATA.i.~-VI output data from the end of blowdown, including the core inlet flow rate and enthalpy, the core pressure, and the core power decay transient, are input to the LOCTA-IV code.

With input from the SATAN~vr code, WR.EFLOOD 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.

WREFLOOD also calculates the mass and energy flow rates that are assumed to be vented to the containiaent.

Since the mass flow rate to the containment depends upon the core flooding rate **and the local core pressure, which is a function of the containment backpressure, the WREFLOOD and COCO codes are

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interactively linked.

WREFLOOD is also linked to the LOCTA-IV code in that thermal-hydraulic parameters from WREFLOOD are used by LOCTA-IV in its cal-culation of the fuel temperature.

LOCTA-IV is used througgout the analysis of the LOCA transient to calculate the fuel and clad temperature of the hottest rod in the core.

The input to LOCTA-IV which consists of appropriate thermal-hydraulic output from SATAN-VI and WREFLOOD and the conservatively chosen initial.RCS: operating con-ditions is summarized in Table 1 and Figure 1.

(The axial pbwer shape of* *

.Figure 1 assumed for LOCTA-IV is a cosine curve which has been previously verified to be the shape that produces the maximum peak clad temperature. (?, lZ, 13))

The COCO code, which is also used throughout the LOCA analysis, cal-culates the containment pressure.

Input to COCO is obtained from the mass and energy flow rates assumed to be vented to the containment as calculated by the

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SATAN-VI and WREFLOOD codes.

In addition, conservatively chosen initial con-tainment conditions and an assumed mode of operation for the containment cooling system are input to COCO.

These initial containment conditions and assumed modes of operation are provided in Table 2.

3.0 DISCUSSION OF SIGNIFICANT INPUT Si~nificant changes in the input data used in this reanalysis from those used in the*currently ~pplicable analyses are delineated in Section 1.0 and are discussed in more detail below.

The changes made in this analysis reflect the operational conditions and limit necessary to allow full power operation at a steam generator tube plugging level of up to 25%.* Increased steam generator tube plugging results in a reduction of steam generator heat transfer area, an increase in the steam generator flow resistance, and a reduction in RCS flowrate (with associated in-crease in the core outlet temperat~re and reduction in the core inlet temperature).

The first two impacts have been accommodated in the same manner as was done in all previous submittals.

The last impact (i.e., lower RCS flowrates) has been conservatively incorporated by assuming a flowrate of 79,650 gpm which is 10%

below the flowrate of 88,500 gpm used in all previous* submittals.

The flowra*te used in this analysis (i.e., 79,650 gpm) is applicable up to a steam generator tube plugging level in excess of 40%.

The magnitude of the change in core inlet and outlet temperature is calculated consistent with the reduced flowrate as-silinption of 79,650 gpm. In addition, for conservatism, no allowance for measure-ment uncertainty has been added to the value of TIN used in this analysis.

In order to ensure compliance with 10CFRS0.46 criteria, several changes to the operational limits assumed in the analysis were made.

Specifically,

• The percentage of steam generator tube plugging for each steam generator is assumed to be identical.

This assumption accurately approximates the actual steam generator tube plugging distribution.

The impact of non-symmetric plugging dis-tribution on the LOCA-ECCS analysis has been found to be insignificant (reference our letter of October 29, 1976 - Serial No. 219 /082776).

e the assumed value of the maximum local peaking factor for the hot fuel rod and the maximum enthalpy rise peaking factor for the hot assembly were reduced (i.e.., 2.0 to 1.85 and 1.435 to 1.38, respectively) from those values used in the currently applicable analysis.

These reduced values will become the new operational peaking factor limits for the LOCA-ECCS analysis.

In addition, changes were made to the Consequence Limiting Safeguards System.

Specifically, the containment quench spray flow iniation time of 20 seconds used in the currently applicable analysis was determined to be overly conservative. The flow initiation time was basedonassump-tion that the piping filled instantaneously after pump acceleration and diesel startup.

Based on a conservative estimate of piping fill time, it was determined that a more realistic, but yet still conservative, flow initiation time of 46 seconds for the quench spray subsystem could be assumed.

In addition, updates to flow initia-tion times assumed for Recirculation Spray Subsystem were made.

The Inside Re-circulation Spray Subsystem flow initiation time was changed from 125 seconds to 120 seconds for conservatism.

Further, the Outside Recirculation Spray Subsystem was incorporated into the analysis, since preliminary analyses indicated that the peak clad temperature could reach a maximum after 300 second, the flow initiation time for the Outside Recirculation Spray System.

The flow initiation times for both the inside and outside recirculation sprays do not-take-credit for either pump acceleration or piping fill delays.

4.0 RESULTS Table 3a P!esents the time sequence of events and Table 4a presents the results for the double ended cold leg guillotine break (DECLG) for the Cn=0.4 break size.

The DECLG has been determined to be the limiting. break size and location based on the sensitivity studies reported in WCAP-8356, WCAP-8572, and WCAP-8853 and is applicable to Surry Units 1 and 2 as reported in our letter dated December 15, 1976-Serial No. 219/082776.

Based on all previous LOCA-ECCS submittals for the Surry units, the results obtained with a. Cn=0.4 break size has been demonstrated to be limiti_ng (see references 8 and 19).

Assuming the initial conditions and modes of 1 **

operation presented in Tables l_and 2, the current analyses resulted in a peak clad temperature of 2177°F, a maximum local cladding oxidation level of 7.4 percent, and a total core metal-water reaction of less than 0.3 percent.

The detailed results of the LOCA reanalysis are provided in Tables 3a through 7a and Figures 2a through 18a.

5.0 CONCLUSION

S For breaks up to and including the double ended serverance of a reactor coolant pipe and for the operating conditions specified in Tables 1 and 2, the Emergency Core Cooling System will meet the Acceptance Criteria as presented in 10CFR50.46.

1.

2.

That is:

The calculated peak fuel element clad temperature is below the requirement of 2200°F.

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

3.

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

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-6. 0 REFERENCES e

1.

"Acceptnnce Criteria for Emergency Care Cooling Systems for Light Water CooledNuclcar Power Reactors" 10CFRS0.46 and Appendix K of 10CFR'.50.

Federal Register, Volu111e 39, Number 3, January 4, 1974.

2.

"Westinghouse ECCS Ev2.luation 1:-fodel-Summary" WCAP-8339, Bordeloni F, ~~.,

Massie, H. W., and Zordan, T. A., July 1974.

3.

Bordelon, F. M., et al., "SATAN-VI Program: Comprehensive Space-Time Department Analysis of Loss-of-Coolant, 11 WCAP-8306, June 1974.

4.* Bo1*delon, F. M., et al., "LOCTA-IV Program:*Loss-of-Coolant Transient Analysis," WCAP-3305, June 1974.

5.

Kelly, R. D., et al., "Calculational Hodel for Core Reflooding after a Loss-of-Coolant Accident (WREFLOOD Code)," WCAP-8171, June. 1974.

6.

Bordelon, F. H. and Murphy, E. T., "Containment Pressure lu1.alysis Code (COCO)," WCAP-83~6, June 1974.

7..

Buterbaugh, *T. L., Johnsoni W. J. and Kopelic; S. D., "Westi1:1ghouse ECCS-Plant Sensitivity Studies," WCAP-83~, July 1974.

8..

Letter from C. H. Stallings (Vepco) to K. R. Goller (NRC), Serial No.

SOO-S 1 dated Jtu~e 6, 1975.

9. Letter from C. M. Stallings (Vepco) to B. C. Rusche (NRC):1 Serial Ho.

017/043073, da:ed Hay 14, 1976.

10.

Lettci:s from C. H. Stallings (Vepco) to Bo C. Rusche (NRG): Scr:i.al No.

194~ dated August *18, 1976, and Serfal No. 211, de.tee! August 26, 1976.

11.

Letter from C.. M. Stallings (Vc.pco) to B. C. Rusche (NRC) ~ Se:ri.s.1 No.

260/092276, dated October 19~ 1976.

12.

Buterbaugh; T. L. ~ Julian, H. V., Tome, A. E.,

11We.stinghouse ECCS-*T'm:ee Loop Plant (17 x 17) Sensitivity Studiess II WCA.P-8572 July J-975 (Pro-prietary) and WCAP-8573, July 1975 (Non-Proprietm:y).

13.

Bordelon, F. M., et al.~ "Westinghouse ECCS Evaluatior. Model-Supplemental Information," WCAP-8472, January, 1975.

llf.

Julian, H. V., Tabone, C. J., and Thompson~.C. H., "Westinr;house ECCS-Three Loop Plant (17 x 17) Sensitivity Studies,. WCAP-8~3, September 1976 (Non-Proprietary).

15.

"Westinghouse ECCS C:valuation Hodel-October 1975 Version", WCP.J?-8622 November 1975 (Proprietary) and WCAP-8623, November 1975 (Non-Proprietary).

16.

Letter. from C. Eicheldinger of Westinghouse Electric Corporation to D. F.

Vassallo of the Nuclear Regulatory Conmd.ssion, Letter Number NS-CE-9, dated January 23, 1976.

)*

17.

Letter fr.om C. Eicheldinger of Westinghouse Electric Corpor.ation to V. Stello of the Nuclear Regulatory Commission, Letter Number NS-CE-1163, dated August 13, 1976.

18.

Letter from R. W. Reid (NRC) to W. L. Proffitt, Serial No. 219/032776, dated August 27, 1976.

19.

Letter from C. M. Stallings (Vepco) to B. C. Rusche (NRC), Serial No. 219/

082776, dated October 29, 1976.

20.

Letter from C. M. Stallings (Vepco) to B. C. Rusche (NRC), Serial No. 219/

082776, dated December 15, 1976.

e TABLE 1 INITIAL RCS CONDITIONS Core Power, MWt, 102% of Peak Linear Power, kw/ft, 102% of Peaking Factor Accumulator Water Volume, ft3 Reactor Coolant System Flow, gpm (90% of Thermal Design)

Steam Generator Tube Plugging Level,%

Inlet Temperature, °F, (TAvG=574.4°F)

Temperature of the Fluid in the Upper Head Region of the Reactor Vessel Core Temperature Rise, °F Hot Assembly Radial Peaking Factor Most Limiting Fuel Region Unit 1 Unit 2 Cycle All All 2441 11.55 1.85 1075 (per accumulator) 79,650 (per loop) 25 539°F 100% of THOT 69.1

1. 38 Region 4

4

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TABLE 2 Containment Data (Dry Containment) 1-H:T ma VOLU/*1[

rn n l /a.L COIW JT 10:!S Pressure Tempcrt;ture Rl*JS T Tcm112 ra tu1'c Servi LC: lfo tcr Tempe re: tu re Out:. i clc, Tcr:ipc!ril t:urc

.srrui.Y S YS1TH 1 l~un1ber o-.C rumps Operating r,unout r-1 o\\'1 !(a tc

/\\ctu,thon T*ime

1. %3>:10 6 f-"t 3

9. 3~i [)Si a 90 C> F 40 0 r-35 OF 9 OF 2

3200 gpm

% secs SPR/\\'r' SYSTEM II -

RCCIRCUL/1,TIO!~ SPR1W FRO;-'; PU:<',P Number* ru::,ps Opeta.ting Run.out Flo*.:rate (e:~ch)

/!,ctuuti on T*i m2 lfr:o.t Exchilngcr [ua (pt:i' pump)]

2 3500 spm

• 120 secs

3. 5):1 C6!3TU/!m-°F Servi.ce 1-.'u.te:r flrn*1 (per exchanuoi')

GlOO gpm

,ium::,21' Pu;Jli)S Orei-ut*ing Run out Fl o*.-1rc1 te ( cacli)

/\\ctun ti on T*i me Service WJlcr Flow 2

3500 gpm 300 secs

3. ~) )'. *1 0 G l3 TU/ I! R- ° F G100 gpm

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S"fRUCTURfiL 11((\\T SHU:.S T 11 i cl: 1w ~. s

( i n * )

Coner-etc 6 Cunc:r-ctc 12 Concrete 13 Concro Le 2fJ

- t Concrete 36 Curbon Si..ce 1 o.3*n:;

Concrete 5!1, Carbon Steel 0.50 Concn:Lc 30 Concrete 24-C;:. rbron Steel 0.366 Stainless Steel 0.426 TABLE 2 (CONT'D)

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, 1*1 u11u:1*l.i"lli1.Y 6,972 57,%0

/\\0,/170 10,500 4,~iO ti6, P,S 7

?.5,07S 11, 28fr 167,165 3,399

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.. e TABLE 3a TIME SEQUENCE OF EVENTS START Reactor S. I. Signal Acc. Injection End of Bypass Pump Injection End of Blowdown Bottom of Core Recovery Acc. Empty DECLG CD=0.4 (Sec) 0.0 0.649 2.26 15.8

23. 71 27.26 27.94 37.18 55.46

TABLE 4a RESULTS FOR DECLG-Co=O. 4 Peak Clad Temp, OF 2177 Peak Clad Location, Ft.

10.5 Local Zr/H20 RXN (max), %

7.4 Local Zr/H20 Location, Ft.

9.0 Total Zr/H20 RXN, %

<0.3 Hot Rod Burst Time, sec.

24.2 Hot Rod Burst Location, Ft.

6.0

e TJi.1E (SEC)

37. *18

38. 11 ii.

• • *-:,r 2G

~:,.JO*

4 3. 69 47.0

/i9. 0 '

!33. 2G

!19. 0

67. 11 83.91 l 02. G 1 122. 51 168.01 222.Gl

?.90. 81 400.0 TABLE Sa Reflood Mass and Energy Releases For Limiting Case at 25% Tube Plug-ing (DECLG, Cn=0.4)

TOT/\\L 1-1,'\\SS rLmm,~TE (Lil/SEC) 0.0 er. o

1. 392 33.52

'1~.03 2~i9~.

2~)01.

23G.5 259.5 266.G 272.9 293.3 310.3

• 323. t+

343. 3 TOT/\\L Ul [:l{CiY r:

FLO',*!IUiT[ (10:JLnurt*c) 0.0 0.0 0.0181 OJ13L1B 0.5712 2.7G6

3. OG3

1. 38S l

"")(l')

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• 1. 372 1.338 1.298

1. 165

1. 03G 0.9120

0. 0778

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e T J l*H:

0.0

l. 0 3.0 5.0 7.0 l 0. 0 15.0 20.0 TABLE 6a Broken Loop Accumulator Flow To Containment For Limiting Case At 25%

Tube Plugging (DECLG, Cn=0.4)

(SEC)

I.V1SS FL Cfrii-U\\ T [. -;; ( l.13 m/ S [ C )

'110 7 3738 3232 2891 2639 2355 202'1 1800 23.01 1703 -,

32.27 0.0

~Tor cn~rgy mass flm*.1r2te multiply r.1a.ss flm*li'Dte

. by a constant of 58 13TU/LSl*i.

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"'" Ir> <.D..... <XlOO Change No. 57 To The Technical Specifications Surry Power Station Units No. 1 and 2 August 26, 1977 I

TS 2.1-1 2.0 SAFETY LIMITS AND LIMITING SAFETY SYSTEM SETTINGS 2.1 SAFETY LIMIT, REACTOR CORE Applicability Applies to the limiting combinations of thermal power, Reactor Coolant System pressure, coolant temperature and coolant flow when a reactor is critical.

Objective To maintain the integrity of the fuel cladding.

Specification A.

The combination of reactor thermal power level, coolant pressure, and coolant t*emperature shall not:

1.

Exceed the limits shown in TS Figure.2.1-1 when 90% of design flow from three reactor coolant pumps exist.

2.

Exceed the limits shown in TS Figure 2.1-2 when full flow from two reactor coolant pumps exist and the reactor coolant loop stop valves in the non-operating loop are open.

3.

Exceed the limits shown in TS Figure 2.1-3 when full flow from two reactor coolant pumps exist and the reactor coolant loop

~top valves in the non-operating loop are closed.

I

TS 2.1-3 uniform and non-uniform heat flux distributions.

The local DNB heat flux ratio, defined as the ratio of the heat flux that would cause-DNB at a particular core location to the actual heat flux, is indicative of the margin to DNB.

The minimum value of the DNB ratio (DNBR) during steady state operation, normal operational transients and anticipated transients, is limited to 1. 30.

A DNBR of 1. 30 corresponds to a 95% probability at a 95%

confidence level that DNB will not occur and is chosen as an appr-opriate margin to DNB for all operati'ng conditions. (1)

The curves of TS Figure 2.1-1 which show the allowable power level decreas-ing with increasing temperature at selected pressures for constant flow (three loop operation) represent limits equal to, or more conservative than, the loci of points of thermal power, coolant system average tempera-ture, and coolant system pressure for which the DNB ratio is equal to 1.30 or the average enthalpy at the exit of the core is equal to the sat-uration value.

The area where clad integrity is assured is below these lines.

The temperature limit*s are considerably more conservative than would be required.if they were based upon a minimum DNB ratio of 1.30 alone but are such that the plant conditions required to violate the limits are precluded by the self-actuated safety valves on the steam generators.

/

The three loop operation safety limit curve has been revised to allow for heat flux peaking effects due to fuel densification and to apply to 90%

I of design flow.

The effects of rod* bowing are also considered in the DNBR analyses.

The curves of TS Figures 2.1-2 and 2.1-3 which show the allowable power level decreasing with increasing temperature*at selected pressures for constant flow (two loop operation), represent limits equal to, or more conservative,

J::.,

0 --A

~

0 u f-1

+

'f-1 0

r-F -

N -

r-1..

f,:.1

~

~-

~

~

~

E-l

~

l!)

~ 5 670 660 650 640 630 620 610 600 590 580 570 560 550

.0 TS Figure 2.1-1 10 20 30 40

.50 60 70 80 90 100 110 120 POWER (PERCENT OF RATED)

FIGURE 2.1-1 REACTOR CORE THERMAL & HYDRAULIC SAFETY LIMITS-THREE LOOP OPERATION, 90% DESIGN FLOW

e where I

(b)

High pressurizer pressure - <2385 psig.

(c)

Low pressurizer pressure - >1860 psig.

(d)

Overtemperature ~T TS 2.3-2 6T.::_~T0 [K1 - K2( i + TlS) (T -

T 1 ) +

K3 (P - P') - f(~I)]

+ T 2S

~T0 = Indicated 6T at rated thermal power, °F T = Average coolant temperature, °F T' = 574.4°F P = Pressurizer pressure, psig P' =

Kl =

K2 =

K3 =

Kl =

K2 =

K3 =

K1 =

K2 =

K3 =

2235 psig

/

1.07 0.0095 0.0005 0.951 0.01012 0.000554

1. 026 0.01012 0.000554 for 3-loop ope~ation for 2-loop operation with loop stop valves open in inoperable loop for 2-loop operation with loop stop valves closed in inoperable loop 61 = qt - qb, where qt and qb are the percent power in the top and bottom halves of the core respectively, and qt+ qb is total core power in percent of rated power f(6I) = function of ~I, percent of rated core power as shown in Figure 2.3-1 Ti*= 30 seconds Tz =

4 seconds (e)

Overpower 6T T S 6T<6T

[K4 - K5(

3

) T -

K6 (T - T') -

f (61)]

o 1 + T3S I

TS 2.3-3 where AT0 = Indicated AT at rated thermal ~ower,.°F T = Average coolant temperature, °F T' = Average coolant temperature measured at nominal conditions and rated power, °F K4 = A constant = 1. 07 K5 = 0 for decreasing average temperature A constant, for increaseing average temperature 0. 02/°F K6 = 0 for T < T'

= 0. 0011 for T > T' f(AI) as defined in (d) above, T 3 = 10 seconds (f)

Low reactor coolant loop flow -.:_90% of normal indicated loop flow as measured at elbow taps in each loop (g)

Low reactor coolant pump motor frequency -.:.. 57.5 Hz (h)

Reactor coolant pump under voltage -

> 70% of normal voltage

3.

Other reactor.trip settings (a)

High pressurizer water level -

< 92% of span (b)

Low-low steam generator water level -

> 5% of narrow range instrument span (c)

Low steam generator water level -

> 15% of narrow range instrument span in coincidence with steam/feedwater mismatch flow -

< l.Oxl06 lbs/hr (d)

Turbine trip (e)

Safety injection - Trip settings for Safety Injection are detailed in TS Section 3.7.

e TS 2.3-5 and source range high flux, high setpoint trips provide additional protection against uncontrolled startup excursions.

As power level increases, during startup, these trips are blocked to* prevent unnec-essary plant trips.

The high and low pressurizer pressure reactor trips limit the pressure range in which reactor operation is permitted.

The high pressurizer pressure reactor trip is also a backup to the pressurizer code safety valves for overpressure protection, and is therefore set lower than the set pressure for these valves (2485 psig).

The low pressurizer pressure reactor irip also trips the reactor in the unlikely event of a loss-of-coolant accident.())

The overtemperature ~T reactor trip provides core protection against DNB for all combinations of p?essure, power, coolant temperature, and axial power distribution, provided only that the transient is slow with respect to piping transit delays from the core to the tem-perature detectors (about 3 seconds), and pressure is within the range between high and low pressure reactor trips.

With normal axial power

. (2) distribution, the reactor trip limit, with allowance* for errors, is always below the core safety limit as shown on TS Figure 2.1-1.

If axial peaks are greater than design, as indicated by the difference between top and bot.tom power range nuclear detectors, the reactor limit is automatically reduced. (4 ) (5 )

The overpower and overtemperature protection system setpoints have been revised to include effects of fuel densification on iore safety limits and to apply to 90% of design flow.

The revised setpoints in the Technical Specifications will ensure that the combination of power, temperature, and pressure will not exceed the revised

TS 3.12-2

-and physics data obtained during unit startup and subsequent operation,

"\\

will be permitted.

c.

The shutdown margin with allowance for a stuck control rod assembly shall ~e greater than or equal to 1.77% reactivity under all steady~

state operation conditions, except for physics tests, from zero to full power, including effects of axial power distribution.

The shut-down margin as used here is defined as the amount by which the reactor core would be subcritical at hot shutdown conditions if all control rod assemblies were tripped, assuming 0

(T

~547 F)

~g-that the highest worth control rod assembly remained fully withdrawn, and assuming no changes in xenon, boron, or part-length rod position.

4.

Whenever the reactor is subcritical, except for physics tests, the critical rod position, i.e., the rod position at which criticality would be achieved if the control rod assemblies were withdrawn in normal sequence with no other reactivity changes, shall not be lower than the insertion limit for zero power.

5.

Operation with part length rods shall be restricted such that except during physics tests, the* part. length rod banks are withdrawn from the core at all times.

6.

Insertion limits do not apply during physics tests or during periodic excerise of individual rods..

However, the shu'tdown margin indicated above must be maintained except for the low power physics test to.measure control rod worth and shutdown margin.

For this test the reactor may be critical with all but one full length control rod, expectE:_d to have the highest worth, inserted and part length rods fully withdrawn.

l

e e

FQ(Z)..:. (1.85/P) x K(Z) for P >.5 FQ(Z) ~ (3. 70) x K(Z) for P ~.5 F!HIAssm *. ~ 1. 38/P

~H~ 1.55 (1 + 0.2(1-P)) x T(BU)

TS 3.12-4 where Pis the fraction of rated power at which the core is operating, K(Z) is the function given in Figure 3.12-8, Z is the core height loca-tion of FQ, and T(BU) is the interim thimble cell rod bow penalty on F~ given in TS Figure 3.12..;.9.

2. Prior to exceeding 75% power following each core loading,.and during each effective full power month of operation thereafter, power distri-bution maps using the movable detector system, shall be made to confirm that the hot channel factor limits of this specification are satisfied.

For the purpose of this confirmation:

a.

The measurement of total peaking factor, ~eas, shall be increased by eight percent to account for manufacturing tolerances, measure-ment error, and the effects of rod bow.

The measurement-of enthalpy rise hot channel factor, F!g, and the hot assembly enthalpy rise factor, ~HjAssm., shall be increased by four percent to account for measurement error. If any measured hot channel factor exceeds its limit specified under 3.12.B.1, the reactor power and high neutron flux trip setpo.int shall be reduced until the limits under 3.12.B.1 are met. If the hot channel factors cannot be brought to within the limits FQ.S. l.85 x K(Z), ~H~l.55 x T(BU) and F~HIAssm *

.s.1.38 within 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />, the Overpower 8T and Overtemperature 8T trip setpoints shall be similarly reduced.

/

I

b.

TS 3.12-4a FQ(Z) shall be evaluated for normal (Conditon I) operation of each unit by combining the measured values of Fxy(Z) with the design Co~-

dition I axial peaking factor values, Fz(Z), as listed in TS Table 3.12-lA and TS Table 3.12-lB.

For the purpose of this specification Fxy(Z) shall be determined between 1.5 feet and 10.5 feet elevations of the core exclusive of grid plane regions located at 25.9 +/-3.2 inches, 52.1 +/-3.2 inches, 78.3 +/-3.2 inches, and 104.5 +/-3.2 inches.

The measured values of Fxy(Z) shall be in~reased by nine percent to account for manufacturing tolerances, measurement error, rod bow, xenon redistribution, and any burnup dependent peaking factor increases.

If the results of this evaluation predict that FQ(Z) could potentially violate its limiting values as established in Specification 3.12.B.1, either:

(1) the thermal power and high neutron flux trip setpoint shall be reduced at least 1% for each 1% of the potential violation (for the purpose of this specification, this power level shall be called PTHRESHOLD)' or (2) movable detector surveillance shall be required for operation when the reactor thermal power exceeds PTHRESHOLD.

This sur-veillance shall be performed in accordance with the following:

(a)

The normalized power distribution,-FQ(Z) I ii,DM' from th~-

ble j at core elevation Z shall be measured utilizing at least two thimbles of the movable incore flux system for

e TS 3.12-12 still assure compliance with the shutdown requirement.

The maximum shut-down margin requirement occurs at end of core life and is based on.the value used in the analysis of the hypothetical steam break accident.

The rod insertion limits are based on end of core life conditions.

The shut-down margin for the entire cycle length is established at 1. 77% reactivity.

All other accident analyses with the exception of the chemical and volume control system malfunction analysis are based on 1% reactivity shutdown margin *

. Relative positions of control rod banks are determined by a specified control rod bank*overlap.

This overlap is based on the consideration of axial power shape control.

The specified control rod insertion limits have been revised to limit the potential ejected rod worth in order to account for the effects of fuel densification.

• I The various control rod assemblies (shutdown banks, control banks A, B, C and D and part-length rods) are each to be moved as a bank, that is, with all assemblies in the bank within one step (5/8 inch) of the bank position.

Position indication is provided by two methods:

a digital count of actuating pulses which shows the demand position of the banks and a linear position indicator, Linear Va.riable Differential Transformer, which indicates the actual assembly position.

The position indication accuracy of the Linear Differential Transformer is approximately.:_5% of span

(.:_7.5 inch~s) under steady state conditions.

The relative accuracy of

\\

the linear position indicator is such that, with the most adverse errors, an alarm is actuated if any two assemblies within a bank deviate by more than 14 inches.

In the event that the linear position indicator is not in service, the effects of

e TS 3.12-13 malpositioned control rod assemblies are observable from nuclear-and process information displayed in the Main Control Room and by core thermocouples and in-core movable detectors.

Below SO~ power, no special monitoring is required for malpositioned control rod assemblies with. inoperable rod posi-tion indicators because, even with an unnoticed complete assembly misalign-ment (part-length or full length control rod assembly 12 feet out of align-ment with its bank) operation at 50% steady state power does not result in exceeding core limits.

The specified control rod assembly drop time is consistent with safety analyses that have been performed.

An inop~rable control rod assembly imposes additional demands on the opera-tors.

The permissible number of inoperable control rod assemblies is limited to one in order to limit the magnitude of the operating burden, but such a failure would not prevent dropping of the operable control rod assemblies upon reactor trip.

Two criteria have been chosen as a design basis for fuel performance related to fission gas release, pellet temperature and cladding mechanical properties.

First, the peak value of linear power density must not exceed 21.1 kw/ft for Unit 1 and 20.4 kM/ft for Unit 2.

Second, the minimum DNBR in the core must not be less than 1.30 in normal operation or in short term transients.

In addition to the above, the peak linear power density and the hot assembly enthalpy rise factor must not exceed their limiting values which.result from the large break loss of coolant accident analysis based on the ECCS acceptance criteria limit of 2200°F on peak clad temperature.

This is required to meet, the initial conditions assumed for the loss of coolant accident.

To aid in specifying the limits on power distribution the following hot channel factors are defined.

e TS 3.12-14 FQ(Z), Height Dependent Heat Flux Hot Channel Factor, is defined as the maximum local heat flux on the surface of a fuel rod at core elevation Z divided by the average fuel rod heat flux, allowing for manufacturing tolerances on fuel pellets and rods.

E FQ, Engineering Heat Flux Hot Channel Factor, is defined as the allowance on heat flux required for manufacturing tolerances.

The engineering factor allows for local variations in enrichment, pellet density and diameter, surface area of the fuel rod and eccentricity of the gap between pellet and clad.

Combined statistically the net effect is a factor of 1.03 to be applied to fuel rod surface heat flux.

N FtiH' Nuclear Enthalpy Rise Hot.Channel Factor, is defined as the ratio of the integral of linear power along the rod with the highest integrated power to the average rod power.

~alAssm., HotAssembly Nuclear Enthalpy Rise Factor, is defined as the ratio of the integral of linear power along the assembly with the highest integrated power to the average assembly power.

It should be noted that ~Hand F~HIAssm. are based on integrals and are used as such in the DNB and LOCA cal.cu~ations, respectively.

Local heat fluxes are obtained by using hot channel and adjacent channel explicit power shapes which take into account variations in horizontal (x-y) power shapes throughout the core.

Thus the horizontal power shape at the N

point of maximum heat flux is not necessarily directly related to FtiH*

The results of the loss ot coolant accident analyses are conservative with respect to the. ECCS acceptance criteria as specified in 10 CFR 50.46 using an upper bound envelope of 1.85 times the hot channel factor normalized operating envelope of TS figure 3.12-8.

.e TS 3.12-15 When an Fq measurement is taken, measurement error, manufacturing tolerances, and* the effects of rod bow must be allowed for.

Five percent is the appro-priate allowance for measurement error for a full.core map (240 thimbles monitored) taken with the movable incore detector flux mapping system, three percent is the.appropriate allowance for manufacturing tolerances, and five percent is the appropriate allowance for rod bow.-. These uncertainties are stat:is-tically combined and result in a net increase of 1.08 that is applied to the measured value of Fq.

N In the specified limit of FAR there is an eight percent allowance for un-certainties which means that normal operation of the core is expected to result in F~H~ 1.55(1 + 0.2(1-P)) x T(BU)/1.08 where T(BU) is the interim N

thimble cell rod bow penalty on FAH given in TS Figure 3.12..-9.

The logic behind the larger uncertainty in this case is that (a) normal perturbations in the radial power shape (e.g. rod misalignment) affect ~H, in most cases without necessarily affecting Fq, (b) the operator has a direct influence on Fq through movement of rods, and can limit it to the desired value, he N

has no direct control over FAR, and (c) an error in the predictions for radial power shape, which may be detected during startup physics tests and which may influence Fq can be compensated for by tighter axial control*

Four percent is the appropriate allowance for measurement uncertainty for

--~ff *obtained from a full core map (2:40 thimbles monitored) taken with the movable incore detector flux mapping system.

The value specified for the limit of ~JAssin. is the value used in the LOCA analysis. It has been determined that four percent is the appropriate allowance co be applied for measurement uneertainty.

Measurement of the hot channel factors are required as part of startup physics tests, during each effective full power month of operation,

TS 3.12-15a I

and whenever abnormal power distribution conditions require a reduction of core power to a level based on measured hot channel factors.

The incore map taken following core loading provide~ confirmation of the basic nuclear design bases including proper fuel loading patterns.

The periodic incore mapping provides additional assurance that the nuclear design bases remain inviolate and identify operational anomalies which would, otherwise, affect these bases.

e TS 3.12-17 For normal (Condition I) operation, it may be necessary to perform surveillance to insure that the heat flu~ hot channel factor, Fq(Z).

limit is met.

To determin~ whether and at what power level surveil-

_lance is required, the potential (Condition I) values of Fq(Z) shall be evaluated monthly by combining the measured values of Fxy(Z) ob-tained from the analysis of the monthly incore flux map with the values of the design Cond~tion I axial peaking factors, Fz(Z).

The product of these shall be increased by nine percent to account for measurement uncertainty, manufacturing tolerances, rod bow, radial redistribution of xenon during normal (Condition I) operation, and any burnup depen-dent peaking factor increases.

PTHRESHOLDis defined as the value of rated power minus one percent power for each percent of potential Fq(Z)

I violation. If the potential values of Fq(Z) for normal (Condition I) operation are greater than the Fq(Z) limit, then surveillance shall be performed at all power levels above PTHRESHOLD" Movable incore instrumentation thimbles for surveillance are selected so that the measurements are representative of the peak core power density.

By limiting the core average axial power distribution, the total peaking peaking factor Fq(Z) can be l,imited since all other components remain relatively fixed.

The remaining part of the total power peaking factor can be derived based on incore measurements, i.e., an effective radial peaking factor, R, can be determined as the ratio of the total peaking

TS FIGURE 3.12-7 DELETE

HOT CHANNEL FACTOR NORMALIZED OPERATING ENVELOPE SURRY POWER STATION UNIT NOS. 1 AND 2 TS FIGURE 3.12-8

-~~TI~7-:;;J-?~TTJ:::21 1.0 lc=:=.ittLi*c{:c;*J :~f ***~***i*:;j ; :: =:***:c*-i*:_:::c~.; *** J

'---+--+---+-~--+-----,---+-----'---!--,--i------,---c----;----,---'-1---,-------11

~

l O'

I:<<

~

r,::i N i 0 z N

~

• ~~ii~~-;~~(~iji-~

0 2

4 6

8 10 12 CORE HEIGHT (ft.)

),

.f.

TS 4.10-1 4.10 REACTIVITY ANOMALIES Applicability Applies to potential reactivity anomalies.

Objective To require evaluation of applicable reactivity anomalies within the reactor.

Specification A.

Following a normalization of the computed boron concentration as a function of burnup, the actual boron concentration of the coolant shall I

be compared monthly with the predicted value. If the difference between the observed and predicted steady-state concentrations reaches the equivalent of *one percent in reactivity, an evaluation as to the cause of the discrepancy shall be made and repotted to the Nuclear Regulatory I

Commission per Section 6.6 of these Specifications.

B.

During periods of power operation at greater than 10% of rated power, I

the hot channel factors identified in Section 3.12. shall be determined during each effective full power month of operation using data from limited core maps.

If these factors exceed their limits, an evaluation as to the cause of the anomaly shall be made.

I

TS 4.10-2 DELETED Basis BORON CONCENTRATION To eliminate possible errors in the calculations of the initial reactivity of the core and the reactivity depletion rate, the predicted relation between fuel burnup and the boron concentration necessary to maintain adequate control char-acteristics must be adjusted (normalized) to accurately reflect actual core conditions.

When full power is reached initially, and with the control rod assembly groups in the desired positions, the boron concentration is measured

~

and the predicted curve is adjusted to this point.

As power operation proceeds, the measured boron concentration is compared with the predicted concentration, and the slope of the curve relating burnup and reactivity is compared with that PFedicted.

This process 'of normalization should be completed after about 10% of the total core burnup.

Thereafter, actual boron con.centration can be compared with prediction, and the reactivity status of the core can be continuously evaluated.

Any reactivity anomaly greater than 1% would be unexpected, and its occurrence would be thoroughly investigated and evaluated.

The value of 1% is considered a safe limit since a shutdown margin of at least 1% with the most reactive control rod assembly in the fully withdrawn position is always maintained.

\\

e TS 4.10-3 PEAKING FACTORS A thermal criterion in the reactor core design specified that "no fuel melting during any anticipated normal operating condition" should occur.

Tb meet the above criterion during a thermal overpower of 118% with additional margin for design uncertainties, a steady state maximum linear power is se-lected.

This then is an upper linear power limit determined by the maximum central temperature of the hot pellet.

The peaking factor is a ratio taken between the maximum allowed linear power density in the reactor to the average value over the whole reactor. It is of course the average value that determines the operating power l~vel.

The peaking factor is a constraint which must be met to assure that the peak linear power

• density does not exc7ed the maximum allowed value.

I During normal reactor operation, measur~d peaking factors should be signifi-cantly lower than design limits.

As core burnup progresses, measured designed peaking factors typically decrease.

A determination of these peaking factors J

during each effective full power month of operation is adequate to ensure that core reactivity changes with burnup have not significantly altered peaking factors in an adverse direction.

3.

TS 5.3-2 Reload fuel will be similar in design to the initial core.

The enrich-ment of reload fuel will not exceed 3.60 weight percent of U-235.

4.

Burnable poison rods are incorporated in the initial core.

There are 816 poison rods in the form-of 12 rod clusters, which are located in vacant control rod assembly guide thimbles.

The burnable poison rods consist of pyrex clad with stainless steel.

5.

There are 48 full-length control rod assemblies and 5 part-length con-trol rod assemblies in the reactor core.

The full-length control rod assemblies contain a 144-inch length of silver-indium-cadmium alloy clad with stainless steel. The part-length control rod assemblies contain a 36-inch length of silver~indium-cadmium alloy with the remainder of the stainless steel sheath filled with Al203.

6.

Surry Unit 1, Cycle 4, Surry Unit 2, Cycle 3, and subsequent cores will meet the following criteria at all times during the operation lifetime.

a.

Hot channel factor limits as specified in Section 3.12 shall be met.

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