Text

Proposed Tech Specs Allowing Use of Reload Fuel Assemblies VANTAGE-5 Design
	ML20079D610
Person / Time
Site:	Farley
Issue date:	07/15/1991
From:	; ALABAMA POWER CO.
To:	;
Shared Package
ML19302E789	List: ML20079D601; ML20079D602; ML20079D603; ML20079D605; ML20079D607; ML20079D610;
References
	NUDOCS 9107180303
	Download: ML20079D610 (623)
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O Attachment 1 Joseph H. f arley Nuclear Plant Units 1 and 2 Request for Technical Specifications Changes VANTAGE-5 Fuel Design Sjtsis for Proposed Chanaes O

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V Joseph M. Farley Nuclear Plant Request for Technical Specifications Changes VANTAGE-5 fuel Design jLasis for Proposed Chanaes Proposed Chanaes The proposed Technical Specifications changes are listed in the attached t able. The proposed changes are based on the operational and core design benefits provided through use of the VANTAGE-5 fuel design in conjunction with impruved computer code methodologies.

EiUL11 in order to implement a long-term fuel management strategy planned by Alabama Power Company for the Joseph M. Farley Nuclear Plant (FNP) Units 1 and 2, it has been decided to use reload fuel assemblies of the Westinghouse 17x17 VANTAGE-5 design. This will require a transition from the current 17x17 LOPAR fueled core to a full VANTAGE-5 fueled core. The transition is expected to be completed by the third reload of VANTAGE-5 O

d fuel for each FNP unit. This long-term strategy includes continuation of the current high energy 18-month fuel cycles with high capacity factors, low leakage loading patterns, and extended fuel burnup. In addition, the VANTAGE-5 analyses assumed power rerate for FNP from 2652 to 2775 MWt where possible; however, the Technical Specifications changes for power rerate are not being pursued in this license amendment request.

The Technical Specifications changes provided in Attachment 2 are based on the use of VANTAGE-5 fuel specific design features and use of improved computer code methodologies. The fuel design features include the smaller diameter Optimized Fuel Assembly (0FA) fuel rods, mid-span zircaloy grids, Intermediate Flow Mixer (IFM) grids, natural uranium oxide (U02) axial blankets, Integral Fuel Burnable Absorbers (IFBA), extended fuel burnup, and Reconstitutable Top Nozzles (RTNs). The RTNs and extended fuel burnup features have previously been licensed at FNP and are currently in use in both Units 1 and 2. Axial blankets are optional starting with the first transition cycle. In addi 'on to the above, the Modified Debris filter Bottom Nozzle (HDFBN) wiii continue to be used. The new computer code methodologies relative to the current FNP safety analyses include the NOTRUMP (small-break LOCA) and improved THINC-IV (thermal-hydraulics) computer codes, as well as the Revised Thermal Design Procedure (RTDP) and the WRB-1 and WRB-2 DNB correlations, in addition, methodologies have been used to allow implementation of either Relaxed Axial Offset Control (RA0C) or the current Constant Axial Offset Control (CAOC), although RAOC is not being implemented at this time. Each Technical Specifications change associated with either the change to VANTAGE-5 fuel or change in o methodology is discussed below.

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Attachment i

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V Page 2 Attachment 2 also provides instructions for incorporating the proposed VANTAGE 5 fuel design Technical Specifications changes. The Unit 2 Technical Specifications changes will be implemented for the first fuel loading of VANTAGE-5 fuel in Cycle 9 of Unit 2 scheduled for late March /early April 1992. The Unit 1 Technical Specifications changes will be implemented for the first fuel loading of VANTAGE-5 fuel in Cycle 12 of Unit 1 scheduled for late September /early October 1992.

The Technical Specifications changes for the core safety limits, reactor trip system instrumentatic. setpoints, and DNB parameters (Technical Specifications Sections 2.1, 2.2 and associated Bases, Section 3/4.2.5, and Bases 3/4.2, 3/4.2.2, 3/4.2.3 and 3/4.2.5) were changed as a result of using DNB margin gained through the use of the VANTAGE-5 IfM grid feature, improved THINC-IV code, WRB-1 and WRB 2 DNB correlations, and RTDP. The use of the VANTAGE-5 fuel design requires the use of the WRB 2 correlation because of its proper treatment of the IFM grids. The WRB-1 correlation is used in the associated analyses for the LOPAR fuel design. The RTDP methodology was used to statistically combine the uncertainties in the DNB correlations with the uncertainties in the plant instrumentation and to provide operational margin to allow greater RCS temperature and pressure uncertainties (Technical Specifications changes to Sections 3/4.2.5 and bs) Bases 3/4.2.5). The DNB temperature, pressure, and flow limits also V include allowances for indicated limits. The Technical Specifications core limits and setpoints changes accommodate the use of higher pealing factors for fbH and F0, deletion of thimble plugs, axial blankets (optianal),

IFBAs, the VARTAGE-5 rod and lattice geometry, mixed fuel type core DNB transition penalty, and future power rerate changes and future higher steam generator tube plugging limit. Since the power rerate and increased tube plugging were assumed in determining the core safety limits and setpoints proposed in this licensing amendment request, plant operation at the current licensed 2652 MWt power rating with the proposed changes results in additional real margin to the DNB limit.

The Technical Specifications change to Table 3.3-2 (Reactor Trip System Instrumentation Response Times) changes the response time for the power range, neutron flux, high negative flux rate trip to "Not Applicable."

This modification will make Technical Specifications response times for ESF and reactor trip functions that are not used as primary protection functions consistent. The high negative flux rate reactor trip function is not taken credit for in any accident analysis, including the Rod Control Cluster Assembly (RCCA) transient (one or more dropped rods).

The Technical Specifications change to Section 3/4.1.3.4 increases the control rod drop time from 2.2 to 2.7 seconds. The VANTAGE-5 IFM grid feature slightly increases the core pressure drop, and the VANTAGE-5 guide thimble inside diameter is slightly reduced compared to the current LOPAR p fuel design. Both of these VANTAGE-5 mechanical design changes result in O

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an increased control rod drop time. Therefore, the safety analyses i performed for the FNP VANTAGE 5 fuel design incorporated an increased i Technical Specifications control rod drop time of 2.7 seconds.

The Technical Specifications change to Section 3/4.1.1.3 is proposed to increase the beginnina of-life Moderator Temperature Coefficient (MTC) limit from 0.5x10-4 Ak/k/of to 0.7x10-4 ok/k/0F. In addition, a linear ramp from 0.7x10-4 Ak/k/0F at 70% power to 0.0x10-4 Ak/k/0F at 100% power is introduced. These changes are proposed to allow greater design flexibility. The et'fects of these changes have been incorporated in the analyses or evaluations of events that are sensitive to this parameter.

The Technical Specifications changes to Heat Flux Hot Channel Factor (FQ ,

Section 3/4.2.2) and Nuclear Enthalpy Hot Channel Factor (FAH, Section 3/4.2.3) are proposed to allow greater design flexibility. The proposed VANTAGE-5 fuel F0 and FAH are 2.45 and 1.65, respectively, at Rated Thermal Power. The LOPAR fuel FQ and FAH are not changed and remain 2.32 and 1.55, respectively, at Rated Thermal Power. The effects of these changes have been incorporated in the analyses or evaluations of events that are sensitive to these parameters.

(~4 The Technical Specifications chmges to Table 3.3-4 are required to

( incorporate the effects of FNP-:,pecific uncertainties associated with plant instrumentation, procedures, and measurement techniques in some ESF allowable values. The changes to the allowable values are proposed to preserve the setpoint values. The changes include reducing the containment pressure high-high-high allowable value from 29 psig to 28.3 psig and reducing the containment pressure high-high allowable value from 18.2 psig to 17.5 psig.

The Technical Specifications change te Table 3.3-5 (ESF Response Times) changes the response time for steam line isolation on high steam flow coincident with Tavg low-low to "Not Applicable." This modification will make Technical Specifications response times for ESF and reactor trip functions that are not used as primary protection functions consistent.

The high steam flow coincident with Tavg low-low ESF function is not taken credit for in any accident analysis including main steam pipe rupture, non LOCA, containment response, or in equipment qualification (superheat) outside of containment.

The Technical Specifications change to Section 3.4.1.2 and the associated Bases is proposed to allow Mode 3 operation with two reactor coolant pumps running and their associated loops operable. Safety analyses, including bank withdrawal accidents which are sensitive to this change, have been reanalyzed or re-evaluated, s The Technical Specifications change to Section 6.9.1.11 is proposed to i change the reporting requirement for the Radial Peaking Factor Limit Report to 30 days after cycle initial criticality. The current requirement is 60

l p Attachment 1 Page 4 y

days prior to criticality. This change is proposed to allow more flexibility after shutdown, in particular for emergency core redesigns.

There are no safety analysis or design criteria that are sensitive to this administrative change.

In addition to the above, changes to the Unit 2 Technical Specifications and Bases are proposed to support removal and replacement of the existing Resistance Temperature Detector (RTD) bypass manifold temperature measurement system with fast response RTDs located in the reactor coolant piping. The proposed changes have previously been reviewed and approved by the NRC for Unit 1. The changes include the allowable values for loss of flow and Tavg low-low [for P-12 (increasing and decreasing) as well as engineered safeguards actuation c;, coincident high steam flow / low-low Tavg]

to include specific margins gained by the setpoint methodology calculations. Additional dynamic compensations were added to the OPAT and OTAT equations to more fully describe the as-installed hardware. Although these dynamic functions are set to have no dynamic effects, they are included to provide complete compatibility with the accident analyses. The OTAT reactor trip response time was increased to account for the new hardware configuration (i.e., RTD element and thermowell versus RTD element only) . The OPAT and OTAT Bases were modified to account for the RTD bypass included with the evaluation for RTD Bypass f

x manifold elimination, Eliminattun for Unit 1, a comprehensive evaluation of the effects of RTD Bypass Elimination for Unit 2 has been completed and no adverse safety implications have been identified.

A significant hazards evaluation (Attachment 3) has been performed to support the Alabama Power Company conclusion that these proposed Technical Specifications changes do not involve significant hazards considerations.

A Safety Assessment and supporting safety analyses (Attachment 4 and Appendices A, B, and C to Attach,aent 4) also have been performed by Westinghouse to support these proposed Technical Specifications changes.

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SUMMARY

AND JU$11FICAtl0N FOR TNE FNP UNITS 1 AND 2 TECHNICAL $PECIFICAfl0N$ CHANGES FC2 VANTAGE *5 FUEL RAgg $?ction Descriotion Justification 22 2.1 Change to core limits, These changes are a result of changes i 2 5,2 8 2.2 and the CTAT and OPAT associated with the VANTAGE 5 fuel, i 2 9,2 10 . Setpoints and the OTAT increased hot channel factore, future l B26 2.2.1 Basis OPAT and Low Flow design and operational considerations, i e25 Attowebte Vetues and the laptementation of RfD sypass 3/4 3 10 3/4.3 Eliminetton for Unit 2.

B21 2.1.1 tests Addition of the W 8 1 This change ceflects the new DNS 8 2-3,5 2 6' 2.2.1 Basis at WB 2 correlations correlations used in analys m .

B 3/4 2 1 3/4.2 Basis These correlations are stoptomented by B 3/4 2 2 the W 3 correlation.

B 3/4 2 4 B 3/4 2 5-3 3/4 4 1 3/4.4 Basis 3/4 1 4 3.1.1.3 Revision of the MtC this change is to allow flealbility d.sring core design. The effect of this increase on the safety analysis has been conaldered.

3/4 1 19 3.1.3.4 Revised rod drop time This change is a resut* of mechanical to less than er equal changes assochted with the VAY'%E 5 to 2.7 seconds fuel. The ef fect of this incrtue on j the safety analysis has been -- -- l considered.

B21 2.1.1 Basis Fgg and Fo(Z) changes These changes are made to give the 3/4 2 4 3/4.2.2 plant core design ffealbility and 3/4 2 7 3/4.2.2 also es a result of ehenges 3/4 2 8 3/4.2.3 associeted with the VANTAGE 5 fuel, f" B 3/4 2 1 3/4.2 Basis The ef f ect of these increases on the safety analysis has been considered.

25 2.2 DNB parameter changes These chonoes are made to account for 3/4 2 14 3/4.2.5 the use of RfDP and to give the plant-3/4 2 15 operating flexibility.

B 3/4 2 5 3/4.2.5 Basis 3/4 3 10 3/4.3.1 OTAT Response Time Accounts for RfD Bypass Elimination in Unit 2.

3/4 3 26 3/4.3.2 contelrunent pressure These changes are made to account for-3/4 3 27 High 2 and High 3 Allow- instrumeat mcertainty to prevent 3/4 3 28 able Vetues and Low Low changes to E$F setpoints; and to Tevg Attowebte Vetues account for RfD Bypass Elimination in Unit 2.

s 3/4 4 2 3.4.1.2 -Change to pu y operable This change allows operational flent-8 3/4 4*1 3/4.4.1 Basis requirement bility. The effects of this change have been conaldered in the safety onetysis.

6 19 6.9.1.11 Radial Peaking Factor This change is made to provide Limit Report adninistrative flexibility.

3/4 3 10 3/4.3.1 PR Neg Rate and This deletion makes these diverse 3/4 3 30 3/4.3.2 Hi Stm flow w/LoLo f mettons consistent with other diverse f avg Respmse Time functions.

!!211 Pegas B 2 4, B 2 5, and 3/4 3-28 are for Unit 2 only.

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l O t Attachment 2 Joseph M. Farley Nuclear Plant Units 1 and 2 l Request for Technical Specifications Changes 1

i Technical Specifications Chanaes l

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A FARLET - UNIT 1 24 AMINDKDff NO. II '

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. REACTOR TRIP SYSTRII INSTRIEEENTAtleft TRIP SETFelflTS

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and f, (af) is a functies of the Indicated difference between top and bettee detectors of the power-ramEe. !

i suelear les chambers alth galms to be selected based on messered lastrument response declag plant startup tests sech that:- ~31 (t) for g beteeen @ percent and O percent, f, ( Af) = 0 (where q, and g, are percent RAftB TREd- le the top and bettee halves of the core respectively, and g, + g, is total TaEmeAL '

POWER to percent of RATED TEERNAL F0VER):

(11) for eneb percent-that the - g,) escoeds percent, the W trip setpoint shall beanteesti-allyredecedbyk. Itude pereofto(g,f its value at RATED TaERNAL F0WEEg and 4

(ill) for each percent that the metalt#de

g. - g,) escoeds percent, the at trip setpoint shall he seismettently reduced by dyercen(t of its value a TED TsRanaL revER.

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E . Nessured K by RTD lastrumentatleeg where

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T . Average temperatore, 'F3 T* Reference T,,,at SATED T m POWER (Collbraties temperature for W Instrumentatlee,

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TABI.E 2.2-1 (Continsed) ,

REACTINt TRIP SYSTEM INSTRtMDfrATION TRIP SETPOINTS 9

POTATION (Centinued) g a

w T, . Time constant ettilred in the rate las controller for 7,,,, t, - 10 secs:

I * ** * . The fonction generated by the lead-lag ceneroller for er dynnele coopensatlear I + t,s T, & t, . Time constants stilfred in the lead-lag controller for ar, Y, . Y, . O seconds:

. I.ag compensator on measured T,,,;

I + t,s t, . Time constant stilfred in the ocasured T,,, lag compensator, t, . O see: ,

h s . 1.aplace transfers operator, see -'; g (2(at) . O for all af. '

Note 3: The channel's nacimum trip point shall not exceed its competed tely point by more than percent. ll Note 4: Pressere value to be determined during initial startop testing. pressure value of f 55 psis to be used prior to determinetton of revised value.

- 3 Note 5: Pressere value to be determined during initial startop testing.

Note 6t The channel's merious trip polnt shall not exceed Its coopsted triy polnt by more tk se pereent. l' t

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2.1 $ArtTY LIMITS i l

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l 2.1.1 Rf ACTOR CORT The restrictions of this safety Limit prevent overheating of the fuel and possible cladding perforation which would result in the release of fission products to the reactor coolant. Overheating of the fuel cladaing is prevente by restricting-fuel operation to within the nucleate boiling regime where the g.rde- e, m heat transfer coefficient is large and the cladding surface , temperature-1s M I"

, j slightly above the coolant saturation temperature. l es emy J Operation above the upper boundary of the nucleate boiling regime could !

result in excessive cladding tenpustures because of the onset of separture from ,

nucleate boiling (DN8) and the resultant sharp reduction in heat transfer ,

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coefficient. . DN8 is not a directly measurable parameter during operation and

,I therefore THERMAL POWER and Reactor Coolant Temperature anti Pressufe have been related to Dh8 tnrough 11- - :--- :-^ --

~ - - -- ' - " - - - - - be e n developed to predict the DNB flux and the location of DN8 for axially uniform and I non-uniform heat flux distributions. The local DN8 heat fluz ratio, DNBR, 8 defined as the retic of the heat flux that sculd cause DNB at a particular cor *

@g location to the local heat flux, is, indicative of the margin to Ohl.

g The a imum value of he DNBR dyrt steady state o eration, noras perational ansients, an nticipated t ansients is lia 'ed to 1.30. is O

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(-Q th t DNB will M occur and is hosen as an propriate marg to DN8 for a y at a g5 perc t confidence 1 e1 *

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included for an tacrease in F N at reduced power based on the expression:

AK F", . . [1+c.3(1P)] e VfWtME 5 Li 6aaO where is the fraction of RATED THERMAL p0WER These limiting heat flux conditions are higher than those calculated for the l range of all control rods fully withdrawn to the maximum allowable control red-

= insertion assaing the axial power imbalance is within the limits of the (Lie 5 C.c WhM i I(ht,ea LaaMt. Cu e,LFL.hs54 FARLEY. UNIT 1 8 2-1 Amindment No. 37

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INSERT D The DNB thermal design criterion is that the probability of DNB not occurring on the most limiting rod is at least 95% (at a 95% confidence level) for any Condition I or 11 event.

In meeting the DNB design criterion, uncertainties in plant operating parameters, nuclear and thermal parameters, fuel fabrication parameters, and computer codes must be considered. As described in the FSAR, the effects of these uncertainties have been statistically combined with the correlation uncertainty. Design limit DNBR values have been determined that satisfy the PNB design criterion.

Additional D.iR margin is maintained by performing the safety analyses to a higher DNBR limit. This margin between the design and safety analysis limit DNBR values is used to offset known DNBR penalties (e.g., rod bow and transition core) and to provide DNBR margin for operating and design flexibility.

1RSERT E

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\ which satisfy the following criteria:

A. The average enthalpy at the vessel exit is less than the enthalpy of saturated liquid (far left line segment in each curve).

B. The minimum DNBR satisfies the DNB design criterion (all the other line segments in each curve). Each curve reflects the most limiting result using either low-parasitic (LOPAR) fuel or VANTAGE-5 fuel. The VANTAGE-5 fuel is analyzed using the WRB-2 correlation with design limit DNBR values of 1.24 and 1.23 for the typical and thimble cells, respectively. The LOPAR fuel is analyzed using the WRB-1 correlation with design limit DNBR values of 1.25 and 1.24 for the typical and thimble cells, respectively.

C. The hot channel exit quality is not greater than the upper limit of the quality range (including the effect of uncertainties) of the DNB correlations. This is not a limiting criterion for this plant.

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2.2 LTM1 TING SAFETY SYSiD: SETT!MCS

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2.2.1 _RE. ACTOR TRIP SYSTEM INSTRWENTAT10N SETPO!WT5 The Reactor Trip 5etpoint Limits specified in Table 2.2-1 are the values et which the Reactor Trips are set for each functional unit. The Trip set-points have been selected te ensure that the reactoi core and reactar coolant systes are prevented from anceeding their safety limits during normal operation and design basis anticipated operational occurrences and to assist the Engi-neered Safety Features Actuation System in sitigating the consequences of accidents. Operation with a trip set less conservative than its Trip Setpoint but within its specified Allowable Value is acceptable en the basis that the difference between each Trip Setpoint and the Allowable Value is equal to er less than the drift allowance assumed for each trip in the safety analyses.

Manual Roseter Trio The Ma.nual Reactor Trip is a redundant channel to the automatic protective instrumentation channels snd provides manal reactor trip capaht11ty.

Power Rance, Neuteen Flux -

[ ') The Power Range, Neutron Flux channel high setpoint provides reactor core protection against reactivity axcursions which are too rapid to be protacted by temperature and pressurs protective cir1:uitry. The low set point provides redundant protection in the power range for a power axcursion beginning from low power. The trip associated with the low setpoint may be manually bypassed when P-10 is active (two of the four power range channels ind;cata a power level of above approximataly 10 percent of RATED THERMAL POWER) and is auto-a natically reinstated when P-10 becomes inactive (three of the four channels 1 indicata a power level below approximately 8 percent of RATED THEIDEL POWER).

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Power Rance, keutron Flux, Mich Rates *

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The Power Range Positive Rata trip provides protection against rapid flux increases which are charactaristic of rod ejection events free any power level. Specifically, this trip complements the Power Range Neutron Flux High

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and Low trips to ensure that the critaria are set for rod ejection from partial

-E \ power.The Power Range Megative kata trip provides protection t.! ensure that the g w -- m 4 :_ w _g + e = = ' M for control rod drop accidenta. At high 2 power a W sultiple rod drop accident could cause local flux peaking a which, when in conjunction with nuclear power being maintained equivalent to turbine power by action of the automatic cod control systas, could cause an unconservative local DNBR to exist. The Power Range Negative Rata trip will provant this free occurring by trippinti the reacter for hu-e4e,1.-ed multiple g

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he cmd'd war hkes for &

pero b *C Mt 4+(p 4A h act)M AAA sf O *<-

Om s>udca,. ik' *'e4u"%b %ct 4.mA +* tecoa.41 re.wu uh .c w* tm% +'. *ve'y-FARLU -uMIT 1 % gu 2-3 AKENIMth7 No. 26

~

L191TTNC SMEN SY53 $t i IMG5 . . - . . .. . . _ . .

I SA.5E5 lattar tH p will > w o-m e t _ -+ M cf 2.; O c n w geir.; e,% :. W euMag normal opentional transiants anc anticipatac trsnsiants when 2 locos art in operation and the Overtanceraturs seita T tMp set point is adjustee to the value specified far all 1 oops ia operation. With the Overtameerature delta T trip set point adjusted to the value specified for 21000 ooerstien.

the p-t tMp at 6CK R.ATED THED%L PWER willi +re-m -tN einis s-wi* M a

$$A 're ge1N ;;1w 1.FduHng normal operational trusients and d anticipata%

transtant.: wita Z loops in operation. .

j Staas Ganentse Vatar Level c " * "E M N C 4 EL5$n c.r R cr'e n is m et The Staas Geerstar Vatar Level Low-Low tHp pewides cars pr'tacion try preventing opension viu the stans generator wetar level below the minimum volume required for adequata hett removal capacity. The specified setdoint previesa allowanca that than will be sufficient wetar inverrtory in the staan generstars at the ties of tHp to allow for starting 6elays of the auxiliary feecmatar systas S ta ma/rt eo a ta r n ow wi sma ten and Lsw S ta es Gee rste r Vata r t. eve l ._ . .

, The Staas/Feedmater Flow Micaatch in esineidenca with a Staem Generstar Lew Watar Level tHp is net used in tm transient and accident analyses but is included in Table 2.2-1 to ansure the functional capability of the specified tHp settings and thersey anhance the overall reliatility of the Rosetor PMtectica Systas. This tHp is redurniant to the 5tsas Generator Vatar Level Low-Low trip. The Stams/Fe+omatar Flow Risaatch portion 'af this tHp is activated wnen the stans flow axceeds the feeewatar flow try greatar than er equal to 1.55 x 10* lbs/ hour. The staam Generstar Low Watar level portion of the trip is activated w6en the urtar level drops below 25 percant, as indicated by the narrow range instrument. These tHp values include sufficient allowa.nce in axcass of normal operating values to preclude spurious tHps but will initiata a Mactor trip befort the staan generators art dry. Therefore, the recuired casacity and starting time requirements of the auxiliary fetewatar punos are reduced and the resulting thersal transient on the Reactor Coolant Systas and steam generators is minimited.

tJnder voltage and Under%cuency - Reartse Caelant Pune Susses The Undervoltage and Urdarfrequency Rearter Coolant Pumo bus trips provide ,

reactor cars protection against CM8 as a result of loss of voltags er under-frecuency to acre than one rtacter coolant pumo. The specified set points assurs a nacter trip signal is generated befort the low flow tHp set point m m -u m i 1 24 u.Excnxi m. zs

t.co5A** be ptr low .s 9 to RAA mivan Coomt0L SYSTEMS . We pew 4 a % nog 7, c./,

N0DDATOR TEMPDATM C05FTICIDrT N'N foWER. N #F4800% 6 T @8Q ,

LIMITING CONDITION FOR OPERAT!0N

'3.1.1.3 The sederator temperature teefficient (NTC) shall bei

a. Less than er equal t's withdrawn. berlaning e cycle a 10** delta UU'F life CSOL). N ewfor 700 "he all rods-y a t eq i R 0 de UU *1 t er A

b. Less sagstive than .4.3 a 10** delta U U 'F for the all rods vithdrava, and of cycle life (80L), AATED TEIRMAL POVER l sendition.

APPLICA3fLTTT Specifiestion 3.1.1.3.a - Nepts 1 and 2* enlyt Specification 3.1.1.3.b - N0 des 1, 2 and 3 onlyt ACTION: '

a. With the WTC more positive than. the limit of 3.1.1.3.s above, operation ta N0 DES 1 and 2 any proceed provided:

1. Centrol red withdrawal limits are established and asistataed sufficient to restore the NTC te withis its limit within 24 hours2.777778e-4 days 0.00667 hours 3.968254e-5 weeks 9.132e-6 months or he la BOT STANDlf vithin the aant 6 hours6.944444e-5 days 0.00167 hours 9.920635e-6 weeks 2.283e-6 months . These withdrawal of Specifiaation limits3.1.3.6.

shall be la addities to the insertion limits O

2.-The control rods are maintained within the withdroval limits established above until a subsequent calculation verifies that the MTC has withdrava seedities. been restored to withia its limit for the all rods

I

.3. A Special Report is prepared and submitted te_ the Counission i I

pursuant to Specification 6.9.2 withis 10 days, describias the value of the maasured artc the interin sentrol red withdrawal '

limits and the predicted average core burnup secessary for.

restering the positive afrC te withis its limit Car the all rods withdrava seedittoa.

b. With the NTc more serative than the limit of 3.1.1.3.b above, be i in 30T S WrDorN within 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months .

With K.,, greater tha'a er equal to 1.0 9 See Special Test Rzeeption 3.10.3

, FARLRT. WIT 1 3/4 1-4 Amendmaat No, U 86 O

REACTIVIT7 CbW?ROL.5Th2Jf 5 RCD DROP TIMI LIMITING CONDITION FOR OPERATION us . ,

, . 3.1.3.4 The individual full length (shutdown and centrol). rod drop time from the fully withdrawn esition (225 to 231 steps. inclusive)* shall l be less thaa or equal t seconds from beginning of decay of stationary gripper coil voltage to ashpot entry with:

(

a. T,,, greater than er equal to 541't, and

b. All reactor coolant pumps operating.

APPLICA$ILITTt MODIS 1 and 2.

ACTION:

a. Vith the drop time of any full length red detersined to exceed the above limit. restore the red drop time to within the above limit prior to proceeding to MODI 1 er 2.

b. Vith the red drop times within limits but determined with 2 reactor coolant pumps operating, operation may proceed provided TEDJLAL POVER is restricted to less than er equal to 66% of RATED TIERMAL POVER.

. SUR7t!LI.A.NC1 R10UIREMDrYS -

4.1.3.4 The red drop time of full length rods shall be demonstrated through sensureeant prior to reactor criticality:

a. For all rods folleving each removal e,f the reactor vessel head,

b. For specifically affected individual rods felleving any maintenance se er modificaties to the control sed drive systes which could affect the drop time of these specific reds, and

c. At,least asce per 18 months.

'The fully withdrawn position used for determining red drop time shall be greater than er equal to the fully withdrawn position used during subse-quest plaat operation.

FARLgT-tMIT 1 3/A 1-19 AMENDNDrf No. 83 O

I l

8C614 3:5?A!!UT N 1.!?'1*$

3/4.2.2 HEAT FLUX HOT CriANNEL FaC*3R F f 2' LIMITI M CONDIT10N FOR OPERATICn 3.2.2 F (Z) sna11 ee 11mitec ey sne following relationsnips:

n Fq (Z) 3 ) [K(Z)) for P > 0.5 % Sg Yadg3g,5 Li ;

Fg (Z) 3 Q (K(Z)) for P y 0.5 for Vedagf. 5 fact 2-.

1 where P . THERMAL POWER Axii,0 inia % FG ER and K(Z) is the f anction cotainec from Figure (3.2-2) for a given core neignt location.

ADPLICA!!LITY: MODE 1 ACTION:

Witn Fg (I) excercing its limit:

a. Recuce THERMAL POWER at least 17. for eacn 15 F witnin 15 minut s anc similarly recuce tne Po.o(Z) exceecs or Range Neutronthe limit Flux-Hign Trip Se points witnin the next nours; PO'nER OPERATION may proctec for up to a total of 72 nours: sutsacuent POWER QPERATION maj proceec provicec tne Overpower celta T Trip Se ;oints nave eeen recucec at least 11 for each 1% FQ (I) exceecs tne limit. The Ov'erpower celta T Trip Setpoint recuction shall be perforsec witn the reactor in at least 9 HOT STANDBY.

D. Icentify and correct the cause of tne out of limit condition prior to increasing THERMAL POWER above the reduced limit requirec ey a, aoove; l THERMAL P0VER may then te increasec provicec F (Z) is cemonstratec thecugn intore napping to be witnin its limit.n ,

I F (g 3 C u 0 [d a D for P > o.5 for LOPAR fud p

Fq005 Ew] CMO3 for P s o.5 Es LOPAR bd FARLEY-UNIT 1 3/4 7 a I.ME CM.ENT NO. Z$ T *.

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s V 3/4 2-7 N4ENDME'NT NO. 26 FARLEY-UNIT 1 l

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3/4.2.3 $UCL!at Eh*maLDV W0i C aNNEL TAC *03 T4 %

LIMITING 00N0li!0N FOR OPERATION ,

N 3.2.3 F A M sna11 t,e limited by the following relationship:

F1 H 1 @ (1

0.3 (1-P)) for Voetast 5 ' bl u) :

THEtKat POWEe enere P = RAiia IntiirW. Nain ,

F4"g s 1.55 C i4 c.3 0-P)',) hr L o~P A R 4 c.l AS8L1 Cal!LITY: M00E 1 A*i10N:

With F"A M exceeciny its limit: ,

4. Reewee THEAMAL POWER to less than 50% of RATED THERKAL POWER witnin 2 neurs anc reswee tne Power Range Newsron F1wa-Mign Trip setpoints to <~~

~

$bt of RATED THERKAL POWER witnin sne next 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months ,

c. Demonstrate inrewgn in core mapping snat F1H is witnin its limit witnin 2a neurs after exceecing the limit or recute THERMAL P0hER t:

less snan 5% of RATED THERKAL POWER witnin the next 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months , anc

c. Icentify anc correct the cause of the out of limit condition prier to increasing TMERKAL POWER atove tha reewcas limit requirac Of g er e, aseve; swesequent POWER OPERAT10N may procese proviceo snat FAM is cemenstratec through in ccre ma: Ding to be witnim its limit at a nominal 50% cf RATED THERKAL POWER ortTP to exceeding this THERMAL P0WER, at a nominal 75% of RATED THERMAL POWER prior to ascencing tnis THERMAL POWER anc witnin 24 nours after attaining 95% -or greater RATED inERKAL POWER.

3/4 2-8 AMEMOMEhi h0. 21, 37.

FARLEY-yh!T 1

POWER DISTRIBUTION LIMITS ~

b,

)

/

ONB PARAMETERS LIMITING CONDITION FOR OPERATION 3.2.5 The following Ot3B related parameters shall be maintained within the limits shown on Table 3.2-1: ,

a. Reactor Ccolant System T,yg ,

b. Pressurizer Pressure

c. Reactor Coolcat System Total Flow Rate ,

APPLICiBILITY: M)DE 1 ACTION:

With any of the above parameters exceeding its limit, reston the parameter to within its limit within 2 hours2.314815e-5 days 5.555556e-4 hours 3.306878e-6 weeks 7.61e-7 months or nduce THERMAL PCWER to less than 5% of RATED THE M L POWER within the next 4 hours4.62963e-5 days 0.00111 hours 6.613757e-6 weeks 1.522e-6 months .

.O T*

U ..

SURVEILLANCE REQUIREMENTS - 4.2.5.1 Each cf the parameters of Table 3.2-1 shall be verified to be within their limits r.t least once per 12 hours1.388889e-4 days 0.00333 hours 1.984127e-5 weeks 4.566e-6 months .

4.2.5.2 TheReactorCoolantSystemtotalflowrateshallbedeterminedtobe within its limits by measurement at least once per 18 months.

9.1.5,3 indic.ded. RCS Clew r.ht, ska.ll k yeg (eg, 4. k wh & accep &4 lid of but osce per- M dags.

6 e

+

AMEMDMENT NO. 26 FARLEY-UNIT 1 3/4 2'-14

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, TAB 12 3.3-2 REACTOR TRIP SYSTEN INSTRUNENTATION RESF0ftSE TIMES FtlNCfl0NAL 12 TIT RESF0ftSE TIME

1. Manual Reactor Trip Not Applicable

2. Fower Range,Iteutron Flux $ 0.5 seconds *

a. Eigh loot Applicable

b. 14v

3. Fever Range, Neutron Flux, Righ Fositive Rate loot Applicable f id

4. Fever Range, Neutron Flux, High Negative Rate _ 0.3 et s*

Intermediate Range, Neutron Flux Not Applicabli 5.

6. Source Range,18eutron Flux Not Applicable a

7. Overtemperature ST $ 6.0 seconds

l j

Overpower af Not Applicable 5 8.

9. Pressuriser Pressure--tow f 2.0 seconds ,

10. Pressuriser Feessure-Bigh 3 2.0 seconds

11. Pressuriser Vater 14 vel--Bigh Not Applicable

15eutron detectors are exempt from response flee testing. Response flee of the neutron flux signal portlen in channel.

Iofthechannelshallbemeasuredfromdetectoroutputorinpu 8

h

%8 t

.hm e -

1 TA8tE 3. (Continued) h ENGINEERED $AFETY FEAftMtE ACTUATION SYSTUt INSTRUNENTAff0N TR ~

Q Att0WRBLE YALUES TRIP SETPOINT_

FtmCTIONAL 1m1T

" 2. CONTAINIENT SPRAY Not App 1fcable h t Appifcable

a. Mensal in1t1at1on Mot App 1Icable Not Applfcable

b. Actomatfe Actuation tog 1c Epsfg Contefament Pressure--Nigh-Nfgh-High .$ 27 psfg

c. T

3. CONTAINNDET 150tATION gg,3]

1

a. Phase 'A" Isolation Not Applicable Not Applicable

1. Manual

{ Not Appifcable Not App 11 cable T 2. Free Safety injection j 5 Automatic Actuation logic

b. Phase "B" Isolation Not App 11 cable Not Appiscable

1. Manual Not Appifcable Not Applicable

. 2. Automatic Actuation logic -

S 27 ps1g 5 .M psIe g 3. CentsInaent Pressvre--

High-High-Hfgh T

g l c. Purge and Exhaust Iseistfon Not App 1tca61e Not Applicable ,

g . 1. 90nnual r

NMI g 2. '/ A A.^.te, d 1 i f.

(u. ww ,

gel %,he@ (rddApput.nAQ.

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O* FARLIT - LWi1T 1 ' 3/4 3-17 AXDOGN! NO. II* II

TABLE 3.'3 5 (C'entinued) ENCINIERID $AFETY FEATtit$ RESPON$t TIMts INITIATINC !!CHAL AND PUNCTION RESPON$t TIME IN $tCON0$

3. Pressuriser Pressure-Lov' I

{ a. Safety Injection (ECCS) i27.0'8'/12.0

b. Reactor Trip (from SI) c.

i2.0 Teodveter Isolation i 32.0

d. Containment Isolation Phase "A" e.

i 17.0 Containment Purge Isolation 3 S.0

f. Auxiliary Feedvater Fueps Not Applicable 3 Service Vater Systes 3 7 7.0' ' ' /8 7.0' ' '

4 Differential Pressure letveen Steam Lines-High

a. Safety Injection (ECCS) 17, b.

                                                                    $ 12.0/22.0

Reactor Trip (from SI)

                                                                    $ 2.0

c. Feedvater Isolation 3 32.0

d. Containment Isolation-Phase "A" e.

3 17.0/27.0 Containment Purge Isolation Not Applicable

f. Auxiliary Feedvater Pumps Not Applicable 3 Service Vater Systes 3 77.0' ' '/8 7.0' ' '

5. Steam Flov in Two Steam Lines.Righ Coincident gedQ vith T,y --Lov.Lov [

a. Steam Line Isolation

                                                                   )11\0                                l

6. Steam Line Pressure-Lov

a. Safety Injection (ICCS) b.

f 12.0/22.0 Reactor Trip (from SI) f 2.0

c. Feedvater Isolation d.

f 32.0 Containment Isolation-Phase "A" f 17.0/27.0

e. Containment Purge Isolation Not Applicable

f. Auxiliary Feedvater Pumps - Not Applicable

g. Service Water Systea

                                                                    $ 77.0/87.0

h. steam Line Isolation $ 7.0 FARLET - UNIT 1 3/A 3 30 AMENDMINT NO.

25. 89 9

REat:0:t COOLaN: SYS EM

                                            "                                                                               lNSEET HOT STACBY                                                                                                          ,

g LIMITING C0CITION FOR OPERATION kN 3.t .2 A tnree 'e acto. colan. Loops stee oe' = sna De OP 'AELE a in opera. on wne the r contr syst is ope tional *r at st t tacto ol ant ops 1 ted be w sna De 0 'ABLE = 6 one ctor tant op in op ation en the ed CD el s 'em is ' sable open the actor

ip Brea -s or ottinc ewn tn rod ve mot / genera e sets. l

1. Reactor Coolant Loop A and its associated steam generator and Reactor Coolant pump, ,,

2. Reactor Coolant Loop B and its associated steam generator and Reactor Coolant pump,

                                                                                                                                                     ~~
 .              3. Reactor Coolant Loop C and its' associated steam generator and Reaetor Coelant' pump ,             ,

APPLIC ABILITY : MODE 3 . ACTION:

a. Witn less than the aeove required iteactor Coolant loops OPERAB E.

restere tne required loops to OPERABLE status witnin 72 hours or be in HOT SHUTDOWN witnin the next 12 hours. INSERT b. tn le snan ree Re a:Or lant ops i opera.'on an the r " g ce *c1 sy em op tiona witn 1 ho open ne Re ter do N Brent *s or 't dew the r driv 'totor nera. r set

c. With no kesctor Coolant loops in operation, suspend all operations involving a reduction in boren concentration of the Reactor Coolant

              . System anc immeciately initiate corrective action to return tne required coolant 1000 to opsration.

4

      $URVIILLANCE REQUIREMENTS                                .

4.4.1.2.1 At least the above required Reactor Coolant pumps, if not in operation, sna11 be determined to be OPERABLE once per 7 days by verifying correct bretter alignments and indicated power availability. 4.4.1.2.2 The recuired Reactor Coolant loop (s) shall be verified to be in operation and circulating Reactor Coolant at least once per 12 hours. 4.4.1.2.3 The required steam generator (s) shall be determined OPERABLE by verifying setendary side water level to be greater than or equal to 10% of wide range indication at least once per 12 hours. .

       'All Reactor Goolant pumps may be de-energized for up to I hour providec (1) no operations are permitted that would cause dilution of the Reactor Coolant System boren concentration, and (2) core outlet tercerature is maintained at least 1U'F celow saturation tercerature.

O, F ARLEY-UMIT 1 3/4 4-2 AMEcMENT c. 25, SE

*~

INSERT H r, . 3.4.1.2 At least two of the Reactor Coolant Loops listed below shall be OPERABLE and in operation when the rod control system is operational or at least two Reactor Coolant Loops listed below shall be ORERABLE with one Reactor Coolant Loop in operation when the rod control system is disabled by opening the Reactor Trip Breakers or shutting down the rod drive motor / generator sets:* INSERT I

b. With only one Reactor Coolant Loop in operation and the rod control system operational, within I hour open the Reactor Trip Breakers or shut down the rod drive motor / generator sets.

O t O 4 e O

2/4.2 PCwit D 5**:rf:CN LIM!*S

                                            ,_                mub) N DNO bity Crkrie BASE 5                                                                                              x w.....n.          ...e........   .... . . ..............             .................,,.,            )

l Tne specif*, cations of tnis set:1on provice assurance of fuel integetty during Concition 1 (Normat Oce" ten) ane 11 (Incidents of Mocerate F eure e r eveats oy: (4) P: 4 u t - u; i ~-' v ZE' '- t": :: : ; :::: : : : ::.: }

i. ..;;] curing normat operap mo in snort term transtants, ano (c) 11m t: n; tne fission gas release, fuel i,<llet temperature anc claccing secnanical pecperties I; witnin assumed casign criteria. In acettion. limiting tne peaa linear power censity curing Concition 1 events provides assurance tnat tre ;

initial concitions assumed for the LOCA analyses are met and tne EC:5 ac:e:tt :e criteria limit of 2200'F is not exceecec. The oefinitions of certain not enannel anc peating factors as usec in 19ese specifications are as follows: Fg(Z) Heat Flux Hot Channel Factor. is defined as the maximum local neat flux on the surf ace of a fuel roc at c6re elevation I diviced by the average fuel roc heat flux, allowing for manufacturing tolerances on fuel pellets anc rocs anc measurement uncertainty. Nuclear Enthalpy Rise Het Channel Factor, is definec as the ratio of Fh' tne integral of itnear power along sne rod witn tne nignest integratec powe* to the average roc power. F,y(Z) Racial Peating Factor, is defined as the ratio of peat power ceasity to average power censity in the neri2 ental plane at core elevaticn 2. 3 / A .2.1 AXiat TLUX DIFFE:tENCE The limits on A11AL FLUX DIFFERENCE (AFD) assure that the ((I) upper counc

elope of 4 rit times the normalized axial peating factor is not exceecec 6 curing eitner noreal operation or in tne event of menon ,redistricution following [ power enanges. i l Target flux cifference is determined at equilibrium xenon ccnci:1 ens. The I full length rocs may oe positionec within the core in accorcance witn snete respective insertion limits and shoule De inserted near their normal position for stency state operation at nigh power levels. The value of the target flux cif ference ootained uncer these concitions divided by the fraction of RATED TetERMAL POWER is tne target flux dif ference at RATID THERKAL POWER for the associated core burnup concitions. Target flux cifferences for etner ThERXAL POWER levels are obtained by multiplying the RATED THERMAL POWER value by the appropriate fractional THERXAL POWER level. The cariccic upcating of the target flux cifference value is necessary to reflect core burnup consicerations. 2.4$ % Y 5

            **) 2 3 2 hr LbF%R FARLEY-UNIT 1                           33/4 2-1                      AMEMDMENT NO 28 73 0         .

j

P74R DISTM3'JTMN LIMITS O BASE 5 - ArIAL FLUY DIFFEREWCE (Continued) Although it is intanded that the plant will be operated with the AFD within the *(5)% target band about the target flux difference, during rapid plant THERMIL PCVER neuctions, enntrol red action will cause the AFD to deviata outside of the target band at reduced THEEvL PCVER levels. This deviation will not affect the menon ndistribution sufficiently to change the envelope of peating factors which may be reached on a subsecuent return to 3 l AATG THERMAL PChER (with the AFD within the target band) proviced the time ' duration of the deviation is Itaited. Accordingly, a 1 hour penalty deviation limit cimulative during the previous 24 hours is previded for operation outside of the target band but within the limits of Figure (3.2-1) while at THER.vL PCVER 1evels between 53 and 93 of RATD THERKAL PCWEA. For THER.VL PCwER 1evels between 15: and 53 of RATE THEMAL PCwtA, deviations of the AFD outsica of the target band are less significant. The penalty of 2 hours actual time reflects this reduces significance. Provisions for monitoring the AFD on an aut.matic basis are derived from the plant precess computar through the AFD Moniter Alars. The computar detarsines the one minuta average of each of the OPERA 2LE excere detector out:uts and provides an alars message ineediataly if the AFD for 2 or more OPERAELE exccre channels are outsNe the target band and the THERvL PC'4A is greater than 9C% of R.ATD THERMAL POnTR. During operation at THERMAL PC43 levels between 50% and SC: and between 15% and !C% AATED THER9L PO41, the

computer outputs an alars massage when the penalty deviation accumulates beyond the limits of I hour and 2 hours, respectively. Figure 8 3/4 2-1 shows a typical acnthly target band. 3/4. 2. 2 and 3/4.2.3 NEAT FLUY HOT CHANNEL FACTOR. NUCLEAf EVTKALpY HOT CMA M EL FACTCA t The limits on heat flux het channel factor, and nuclear enthalpy rise het channel facter ensure that 1) the desi limit on peak local power density [- = : W :.: T O :. N not azteeded and in th(e event of a LOCA the peak fuel clac tuserature vi u not ascoe e 2.200'F ECC3 acceptance critaria limit. Each of these is measurable but will normally only be determined periodically as specified in Specifications 4.2.2 and 4.2.3. This periodic surveillance is sufficient ta insurs that the Itaits are saintained proviced: 3 4 %e, s. Control rods in a single group move together with no individual rod insertion differing by more than + 12 steps, indicated, from the g g$,

         .                 group demand position.

t.nfLr,im b. Control red baats are sequenced with everlapping groups as described 4 "d J in Specification 3.1.3.6.

c. The control rod insertion limits of Specifications 3.1.3.5 and l 3.1.3.6 are maintained.

l d. The axial peror distribution, arsressed in tems of AXIAL FLUX DIFFERENCE, is maintained within the limits. - i l FARLIY-uMIT 1 3 3/4 2-2 ' AMEN: MENT No. 25 e

8 Wit 015 al!UT10N tiwl?S (- 81515

                ....................................e........................................

F g will se maintained within its limits provisej conditions a. t**ow; d. acove a re maintained. The relaxation of F3 M as a function o' THER"AL PChER allows enanges in the radial power shape for all peranssicle l roc insertion limi ts. ! When an F n masurment is taken, an 411mance for both esperimental error anc manuficturing tolerance aust be mace. An allemance of it is approcriate for a full core map taken with the 1ncore detector flus massing system and a 31 allowance is appropriate for manufacturing tolerance. When F jg is measured, experisental error sust be alloed for and at is the appropr i ate a11mance for a full cor.e mag taken with the intore l cetection system. The specified limit for FDE contains an 81 alto.ance f ee uaceat a utiesg=-': n : : p :: --2 ::: :: : r- ::, - 3 p j ;'. . ;; /'. 1 The n allowance is basec on the following consicerations:

a. Aoncemal perturtations in the racial poner shape, such as from roc N

misa11gment, af fect F g g more cirectly than ( , l

c. Altnowgn roc movernent has a direct influence upon limiting Fn to w(thin its limit, such control is not rese11y availatie to Ilms:

F p . anc - r~ c. Errors in creciction for control power shape setected during startu: (' physics tests can be coecensated for in F by restricting axial f1wn i ci st riou tion. This cogensation for F5H s less reacily availacle, j fuel r. coming rec. s the value" Chl ratio. ects is av 'lacle to o et this re ion in th entric marg The gener esign marg. . 1 tota a 9.11 ON8n, rsl etely set any roe penalties ess snan . < for sne st case un occurs at u.nup.af 3 , OPrd/MTU). This ma rgin

l eluces twe 11 ming: o esign 11m NB R of 1. . 1.28

                                         ),                         1 Grie Spa       Coeffic          (t ) of 0.       6 vs. 0.059

3) The - Diffusion efficient .038 vs. . 9 0$R N1. lier of 0. vs . 0.84 5 +t en reeve a O

e O.

                  ,, a ,-- 1                                                                     . 2,a . :                     --1 = . u . a e

e0WIR D157t!!'."l0N LIM!'S LASES-O Tae racial pearing factor Fay (Z),'is measured periodically to provice asettional essurance that tne het channel fger, FQ(Z), rematas within its-limit. The F, limit for RATED TMERMaL P0iiER (F ) as provised in the Aasial Peating Factor # limit report per Specification 6.I.1.11 was determined from espected power control maneues over the full range of burnup conditions in the core. 3 /a .2.4 OUADRaNT PCWER T!LT RATIO , The quadrant power tilt ratic limit assures that the radial power distribution satisfies the sesign values used in the power capability analysis. Racial po.er sistribution measurements art made during startup testing and periodically during power operation. . Tne limit of 1.02, at whien corrective action is requirec, provices Dh5 an: J linear nest generation rate protection witn a.y plane power tilts. i The t.o hour time allowance for operation with a tilt condition greater than l 1.02 out less snan 1.0g is provicec to allow teentification and correction of a l croppec or misalignee control rec. In the event such action sees not correct tne tilt, the margin for uncertainty et F0 is reinstated by recuting the sastre allowes po.er by 3 percent for each percent of tilt in excess of 1.0. For purposes of monitoring QUADRANT POWEd TILT RATIO when one encore setector is s inopersole, the movatie incore detectors are useo to confirm that the normalizac - k@TJ} symmetric power sistritution is consistent with the QUADRANT POWER TILT RAY 10. Tne intore setector rionitoring is sons with a full intore flum mac or two sets of four symmetric taimoles. Tne twc sets of four symmtric taimoles is a uni:ve set of eignt cetector locations. These locations are C-8. E.E. E-11, N 3, m.13,

    .          L.5, L.ll, h-8.
                                                                                     .                        INSERT J 3/a.2.5 DN! PARAPfrIRs                                                                            HEM ThelimitsontheDNSrelatedparametersassurethateacnoftheparametersare maintainee witnin the normal stesey state envelope of operation asspec in the transient and accident analyses. The limits are consistent with the initial FSAR assumotieefAhave been analytically descastrated adequate tol= /                                 =.;)

Irmr = M :.mthrougnout each analyzed transient. C Re 12 hou riodic suheillance e ese paramat t through in reent reae

Nt isNufficient o ensure th the par es are rest d within the limits Mc fell ing load c nges and a r expected .ansient ope ion. The 1 oth D$* periodi. asureme of the RC stal flow r is adequat a detect f1 3,,g" n to racatto and ensur correlation f the flow ication cha 1s with mea ec f' the indi ow will pro a suffician trification.

NJM of sucn th rate on 12 hour ed percent is. .. h INSERT k AMghDxth; no.64 FARE 1-UNIT 1 3 3/4 2 5 gg e

INSERT J The' indicated Tavg value of 580.70F is based on the average of two control ooard readings and an indication uncertainty of 2.50F. The indicated pressure value of 2205 psig is based on the average of two control board readings and an indication uncertainty of 20 psi. The indicated total RCS flow rate is based on one elbow tap measurement from each loop and an uncertainty of 2.4% flow (0.1% flow is included for feedwater venturi fouling). INSERT K The 12-hour surveillance of Tavg and pressurizer pressure through the control board readings are sufficient to ensure that the parameters are restored within their limits following load changes and other expected transient operation. The 18-month surveillance of the total RCS flow rate is a precision measurement that verifies the RCS flow requirement at the beginning of each q fuel cycle and ensures correlation of the flow indication channels with the measured loop flows. The monthly surveillance of the total RCS flow rate l is a reverification of the RCS flow requirement using loop elbow tap measurements'that are correlated to the precision RCS flow measurement at the beginning of the fuel cycle. The 12-hour RCS flow surveillance is a qualitative verification of significant flow degradation using the control board indicators and the loop elbow. tap measurements that are correlated to the precision RCS flow measurement at the beginning of each fuel cycle. 1 O

j 's_. a RtatTOR t00L4WT $YtTIM o-M & D$ det%n C.citrion ggg3 3/4.4.1 RtitT0s t00LtW1 LOOP 5 AWO *00L4WT CintutaT10 ' The plant ti desicaed te oeerste61th all Reactor Coolant loops in operation, and N-W- P'O :::+W eurir.g 411 nomal operattans and anticipated transients. in uLa 1 and 2 with one teactor Coolant 1000 not in eperation this spMification rewires that the plant be in at least HOT $7ANDtv l l within 1 hour. In MODE 3,thrhJreactor coolsnt loops provide sufficient heat reseval capability for removing core heat even in the event of a bank withdrawal I accident; however 4 s' ngle reactor coolant loop provides sufficient decay hea; l rareval capacity if a bank withdrawal shuttint accident can be preventact 1.e., tiy 60WM the Ptd drivt motoP/qeneratoe ! opening the teactor Trip trHkers eP0 x:'nM :: M ; ;; . ::. :M;?; f: kr; sets.- e: : :: e cf W -: L..

         ;; ;u;       t4 : r: W : p t t.: h:;; t: PtV"L! n ;M I MODE 4, a single reactor 'c'oJ1 ant er Ip(R loop provides sofficient heat                                                .

remove capability for removing decay heat; but Thus, single if failurt the Mactor ensiderations

%1 ant leccs rowir' that at least two loops be OPERA 4Lt.

6te no- OPERABLE, this spMification rewires two RHR loops to be opt.' ult. In MODE 5, single failure considerations rewi.'s two R* 1 oops to u OPERAllt. The operation of ene 9eactor Coolant Puerp er one RHP, pump provides adecutte flow to ensurt string, prevent stratification and produce tradual reactivity J O changes during boron concentration rt@ctions i'i the Reactor Coolant $ystes. The reactivity changs rate associatec with boron reduction will, therefore, be within the capability of operator recognition and control. The restrictions on starting a textar Coolant Puns with one er more Aesctor Coolant system cold legs less than er emal tp 310'r am provided to prevent Reactor Coolar.t systes pressure transients, caused by e Part 50. The Reetter Coolant System, will be protacted against everpressure transients and will het ascoed the limits of A&Mridia G ty either (1) restricting the watar vo)ves in the pressuriter and thereby providing a volume for the primary coolant to expand inta, er (2) by fostricting starting of the Reactor Coelaat Pumps to when the secondary veter t;nparature of e6ch star generator is less than 50*F above each of the Reactor Coolant System cold leg tesperst m

                                                                                                                                          '     O
                                                                       - - - - - - - - - - - _ _ _ _ ~ _ . _ _               __

ADMINISTRATIVE CONTROLS

e. Type of container (e.g., LSA, Type A Type B, Large Quantity), and

f. Solidification agent (e.g., tesent, ures formaldehyde).

The radioactive effluent release reports shall include unplanned releases from the site to unrestricted steas of radioactive materials in gaseous and liquid offluents on a quarterly basis. i The radioactive effluent release reports shall include any changes to the PROCESS C0!UROL PROGRAM (FCP) made during the reporting period. i MONTHLY OPERATING REPORT j 6.9.1.10 Routine reports of operating statistics and shutdovn experience, including docueentation of all challenges to the PORV's or safety valves, shall be submitted on a monthly basis to the Commission, pursuant to 10CTR$0.4, l no later than the 15th of each month following the calendar month covered by the report. Any changes to the OFTSITE DOSE CALCULATION MANUAL shall be submitted with the Monthly Operating Report within 90 days in which the change (s) was made effective. In addition, a report of any major changes to the radioactive vaste treateent systees shall be submitted with the Monthly Operating Report for the period in which the change was implementvid. (b M 3 RADIAL PEAKING FACTOR LIMIT REPORT 6.9.1.11 The F limit for Rated Thermal Power (yRTP shall be provided to the Commission, pur,s6 ant to 10CFR50.4, for all core p1'le)r bank "D" a s containing l control rods and all unrodded core planes &4.-atas& M -saya_-triet: M cycle initial criticality. In the event that the limit would be submitted at some other time during core life, it vill be submitted lEEilayi:stlerrim the date the limit would become effective unless otherwise exempted byh he t Commission. Any information needed to support F}TP will be by request from the NRC and need not be included in this report. ( L ,._ _ ANNUAL DIESEL GE NTOR RELIABILITY DATA REr0nT 6.9.1.12 The number of tests (va12d or invalid) and the number of failures to start on demand for each diesel generator shall be submitted to the NRC annually. Th' report shall contain the information identified in Regulatory Position C.3.b of NRC Regulatory Guide 1.108. Revision 1, 1977. FARLEY-UNIT 1 6-19 AMENDMENT NO. I/, 19. 82 O .

J. H. PARLEY NUCLEAR PLMIT Utili 1 TECHNICAL SPECIPICATIONS Remove Pages Insert Pages 2-2 2-2 2-5 2-5 2-8 2-8 2-9 2-9 2-10 2-10 B 2-1 B 2-1 B 2-2 B 2-2 B 2-3 B 2-3 B 2-6 B 2-6 3/4 1-4 3/4 1-4 3/4 1-19 3/4 1-19 3/4 2-4 3/4 2-4 3/4 2-7 3/4 2-7 3/4 2-8 3/4 2-8 3/4 2-14 3/4 2-14 3/4 2-15 3/4 2-15 3/4 3-10 3/4 3-10 3/4 3-26 3/4 3-26 3/4 3-27 3/4 3-27 3/4 3-30 3/4 3-30 0< 3/4 4-2 3/4 4-2 B ?/4 2-1 B 3/4 2-1 B 3/4 2-2 B 3/4 2.2 B 3/4 2-4 B 3/4 2-4 B 3/4 2-5 B 3/4 2-5 B 3/4 4-1 B 3/4 4-1 6-19 6-19 O

l O 670' UNACCEPTABLE 660 OPERATION 2440 psia 650< 640, 2250 pala 630< 620-

      "                                1875 psia 610-        1840 pala O

600-590 ' ACCEPTABLE OPERATION

           $80<

570 - -

0. .1 .2 .3 4 .5 .6 .7 .8 .9 1. 1.1 1.2 POWER (FRACTION OF RATED THERMAL POWER)

Figure 2.1-1 Reactor Core Saf ety Limits Three Loops in Operation O FARLEY - UNIT 1 2-2 AMENDMENT NO.

O O O n TABLE 2.2-1 E

REACTOR TRIP SYSTEM INSTRUMEfGATION TRIP SETPOIPES

 '                                                                                                                               I E                                               TRIP SETPOI?U                            ALLOV/ ELE VALUES                      l y        FUNCTIONAL UNIT Manual Reactor Trip               Not Applicable                           Not Applicable 1.

Low Setpoint - f 25% of RATED Lov Setpoint - f 26% of RATED

2. Power Range, Neutron Flux THERMAL POWER i THERMAL POVER High Setpoint - f 109% of RATED High Setpoint - f 110% of RATED THERMAL POVER THERMAL POVER Power Range, Neutron Flux, f 5% of RATED THERMAL POVER vith $ 5.5% of RATED THERMAL POVER

3. with a time constant 2 2 seconds High Positive Rate a time constant 2 2 seconds Power Range, Neutron Flux, f 5% of RATED THERMAL POWER with f 5.5% of RATED THERMAL POVER 4.

a time constant 2 2 seconds with a time constant 2 2 seconds y High Negative Rate f 25% of RATED THERMAL POVER $ 30% of RATED THERMAL POVER

5. Intermediate Range, Neutron Flux f 10' counts per second f 1.3 I 10' counts per second

5. Source Range, Neutron Flux l

See Note 1 See Note 3

7. Overtemperature aT See Note 2 See Note 6

8. Overpower AT 2 1865 psig 2 1855 psig

9. k% :-urizer Pressure--Lov Pressurizer Pressure--High $ 2385 psig i 2395 psig 10.

f 92% of instrument span f 93% of instrument span E 11. Pressurizer Vater Level--High g 33 > 90% of minimum measured flov > 88.5% of minimum measured flov l

12. Loss of Flav -

per loop

5 per loop
l
Minimum measured flow is 89,290 gpm per loop.

7, 7.. , TAB .-l (Continued) ( ) (v) % ./ n REACTOR TRIP SYSTEM INSTRUMENTATION TRIP SETPOINTS D E NOTATION

  '  Note 1: Overtemperature aT E

Q AT (1 + T,s) f aT, [ K,- K, (1 + T, s ) (T ( l ) - T') + K, (P - P') - fg (aI)]

(1 + T3s) (1 +T,s) 1 + T, s where: AT = Measured AT by RTD instrumentation; AT, = Indicated AT at RATED THERMAL POVER; T = Average temperature, 'F; T' $ 577.2*F (Maximum Reference T,,, at RATED THERMAL POVER); P = Pressurizer pressure, psig; P' = 2235 psig (Nominal RCS operating pressure): Y

I*Ts i . The function generated by the lead-lag controller for T,,, dynamic compensation; 1 + T, s i

T & T, 3

                              = Time constants utilized in the lead-lag controller for          T,,,,  T1 - 30 sec, T,    4 sec; I'T8 4       = The function generated by the lead-lag controller for AT dynamic compensation; 1+Ts 3 T, &T    3
                              - Time constants utilized in the lead-lag controller for ST, T, =T            3
                                                                                                              - O sec; 1
                              = Lag compensator on measured T,,,;

g 1 + T, s t1 g T, = Time constant utilized in the measured T,,, lag co.ipensator, T, = 0 see; x s - Laplace transform operator, sec *; - h 5 Operation with 3 loops Operation with 2 loops Kg = 1.14; Kg = (values blank pending l K, = 0.0250; K, = NRC approval of l K3 - 0.001275; K 3

                                                                                                     - 2 loop operation)                                   l

                                                                     ~1 5   "II""    }                                                                                                t, /

s REACTOR TRIP SYSTEM INSTRUMENTATION TRIP SETPOINTS NOTATION (Continued) , t @ and f 3 (AI) is a function of the indicatei difference between top and bottom detectors of the power-range y nuclear ion chambers; with gains to be selected based on measured instrument response during plant startup - tests such that: (i) for q -q between -39 percent and +13 percent, f (aI) = 0 (where q and q are percent RATED l THERM 1L PODER in the top and bottom halves of the c, ore respectively, a,nd q, ,+ q, is total THERMAL POVER in percent of RATED THERMAL POVER); (11) for each percent that the r.agnitude of (q, - q,) exceeds -39 percent, the ST trip setpoint shall be automatically reduced by 1.92 percent of its value at RATED THERML POVER; and (iii) for each percent that the magnitude of (q, - q,) exceeds +13 percent, the aT trip setpoint shall be autematically reduced by 2.17 percent of its value at RATED THERMAL POVER. Note 2: Overpower at w & ST (1 + T,s) f AT, [K,- K3 ( Ts 3 ) ( 1 ) T - K, (T ( 1 ) - T") - f 3 (aI)] (1 + T3s) 1 +T 3s 1 + T, s 1 + T, s where: AT = Measured aT by RTD instrumentation; aT,= Indicated AT at RAT 13 THERMAL POWER; T = Average temperature, 'F; T" - Reference T,,, at RATED THERMAL POVER (Calibration temperature for ST instrumentation, f 577.2'F); g K, = 1.07; l K 3 = 0.02/'F for increasing average temperature and 0 for decreasing average temperature; h K, = 0. W W T & T > P , K, = 0 h T f P ; l 5 T)s = The function generated by the rate lag controller for T,,, dynamic compensation; 3+ T3 s

q s

                                                            '\_                                              b N                                                       TABLE 2.2-1 (Continued)

E REACTOR TRIP SYSTEM INSTRUMENTATION TRIP SETPOINTS

 @                                                        NOTATION (Continued) l T 3 - Time constant utilized in the rate lag controller for   T,,,,  T3 - 10 see;              ;

I*Ts e - The function generated by the lead-lag controller for ST dynamic compensation; 1+Ts 3 T, &T 3- Time constants utilized in the lead-lag controller for ar, T, - g = 0 see; I

                            - Lag compensator on measured T,,,;

[+ T 68 i T, - Time constant utilized in the measured T,,, lag compensator, T, = 0 sec; i

~*

[ s - Laplace transform operator, sec ; f2(AI) = 0 for all al. Note 3: The channel's maximum trip point shall not exceed its computed trip point by more than 1.8 percent. ] , Note 4: Pressure value to be determined during initial startup testing. Pressure value of f 55 psia to be used prior to determination of revised value. l l Note 5: Pressure value to be determined during initial startup testing. 1 Note 6: The channel's maximum trip point shall not exceed its computed trip point by more than 2.3 percent. l 9 e 1 a 4 r 5 i i

l 2.1 SAFETY LIMITS J BASES 2.1.1 REACTOR CORE The restrictions of this Sefety Limit prevent overheating of the fuel and 4 possible cladding perforation which vould result in the release of fission I products to the reactor coolant. Overheating of the fuel cladding is prevented by restrictins; fuel operation to within the nucleate boiling regime where the heat transfer coefficient is large and the cladding surface temperature is slightly above the coolant saturation temperature, Operation abovr the upper boundary of the nucleate boiling regime could result in excessive cladding temperatures because of the onset of departure from nucleate boiling (DNB) and the resultant sharp reduction in heat transfer coefficient. DNB is not a directly measurable parameter during operation and therefore THERMAL POVER and Reactor Coolant Temperature and Pressure have been related to DNB through correlations which have been developed to predict the l DNb flux and the location of DNB for axially uniform and non-uniform heat flux distributions. The local DNB heat flux ratio, DNBR, defined as the ratio of the heat flux that vould cause DNB at a particular core location to the local heat flux, is indicative of the margin to DNB. The DNB thermal design criterion is that the probability of DNB not occurring on the mort limiting rod is at lesst 95 percent (at a 95 percent confidence level) for any Condition 1 or 11 event. In meeting the DNB design criterion, uncertainties in plant operating parameters, nuclear and thermal parameters, fuel fabrication parameters and computer codes must be considered. As described in the FSAR, the effects of O these uncertainties have been statistically combined with the correlation uncertainty. Design limit DNBR values have been determined that satisfy the DNB design criterion. Additional DNBR margin is maintained by performing the safety analyses to a higher DNBR limit. This margin between the design and safety analysis limit DNBR values is used to offset knovn DNBR penalties (e.g., rod bov and transition core) and to provide DNBR margin for operating and design fhxibility. The curves of Figures 2.1-1 and 2.1-2 show the reactor core safety limits for a range of THERMAL POVER, Reactor Coolant System pressure and average temperature which satisfy the following criteria

a. The average enthalpy at the vessel exit is less than the enthalpy of saturated liquid (far left line segment in each curve).

b. The minimum DNBR satisfies the DNB design criterion (all the other line segments in each curve). Each curve reflects the most limiting result using either lov-parasitic (LOPAR) fuel or VANTAGE 5 fuel. The VANTAGE 5 fuel is analyzed using the VRB-2 correlation with design limit DNBR values of 1.24 and 1.23 for the typical and thimble cells, respectively. The LOPAR fuel is analyzed using the VRB-1 correlation with design limit DNBR values of 1.25 and 1.24 for the typical and thimble cells, respectively,

c. The hot channel exit quality is not greater than the upper limit of O the quality range (including the effect of uncertainties) of the DNB correlations. This is not a limiting criterion for this plant.

FARLEY - UNIT 1 B 2-1 AMENDMENT NO.

_. . _ _ _ . _ . . _ . _ . _ . _ - . _ _ . _ _ __..__m . . _ ._ _._ _ ._ _ _ _ . - _ _ _ SATETY LIMITS BASES ; i

                                -The curves of Figures 2.1-1 and 2.                                          2 are based on the most limiting result uginganenthalpyhotchannel. factor,F),,'of1.65forVANTAGE5fuelandan F ,, of 1.55 for LOPAR fuel and a refere,ce                                          n     cosine with a pgak of 1.55 for axial power shape. An allovance is included for an increase in F ,, at reduced power based on the expression:

F",, = 1.65 [1 + 0.3 (1-P)] for VANa 3E 5 fuel and l

                                               ?",,                1.55 [1 + 0.3 (1-P)] for LOPAR fuel                                                                                      l               ,

where P is the fraction of RATED THERMAL POVER. i These limiting heat flux conditions :.re higher than those calculated for the range of all control rods fully withdrawn to-the maximum allovable control rod. insertion assuming the axial power imbalance is.vithin the limits of the f 3 (delta I) function.of the Overtemperature trip. When the axial power imbalance is.not within the-tolerance, the axial power imbalance effect on the Overtemperature delta T trips vill reduce the setpoints to provide protet' ion copyistent with core safety limits. 2.1.2 REACTOR COOLANT SYSTEM PRESSURE The restriction of this Safety Limit protects the integrity of tb. F.eactor Coolant System from overpressurization and thereby prevents the relesse of radionuclides contained in the reactor coolant from reaching the containment atmosphere. I The reactor pressure vessel, pressurizer and the reactor. coolant system piping and fittings are designed to Section III of.the ASME Code for Nuclear Power Plant which permits a maximum transient pressure of 110% (2735 psig) of '

                    . design pressure. The Safety Limit of 2735 psig is therefore consistent with the
                  - design criteria and associated code-requirements.                                                                                                                                       i The entire Reactor Coolant System is hydrotested at 3107 psig, 125% of.                                                                                                  ..

design pressure, to demonstrate integrity prior to initial operation. ' O

FARLEY UNIT 1 B 2-2 AMENDMENT NO.

9 i-y3----sv ,m-,-+-e-wr,-w,e-se,rwr v--ee--es-evr*,--+e, v. m* , e- +eww ~-wew- .e-w-~s - .- < ww- r -m*,r-m - - , - -+ -+e--:- r--= + - -

2.2 LIMITING SATETY SYSTEM SETTINGS BASES O 2.2.1 REACTOR TRIP SYSTEM INSTRUMENTATION SETPOINTS The Reactor Trip Setpoint Limits specified in Table 2.2-1 are the values at which the Reactor Trips are set for each functional unit. The Trip Setpoints have been selected to ensure that the reactor core and reactor coolant system are prevented from exceeding their safety limits during normal operation and design basis anticipated operational occurrences and to assist the Engineered Safety Features Actuation System in mitigating the consequences of accidents.

operation with a trip set less conservative than its Trip Setpoint but within its specified Allovable Value is acceptable on the basis that the difference between each Trip Setpoint and the Allovable Value is equal to or less than the drift allovance assumed for each trip in the safety analysis. Manual Reactor Trip The Manual Reactor Trip is a redundant channel to the automatic protective instrumentation channels r provides manual reactor trip capability. Power Range, Neutron Flux _ The Power Range, Neutron Flux channel high setpuint provides reactor core protection against reactivity excursions which are too rapid to be protected by temperature and pressure protective circuitry. The lov setpoint provides redundant protection in the power range for a power excursion beginning from lov power. The trip associated with the lov setpoint may be manually bypassed when P-10 is active (two of the four power range channels indicate a power level of above approximately 10 percent of RATED THERHAL POVER) and is automatically tO reinstated when P-10 becomes inactive (three of the four channels indicate a pover level belov approximately 8 percent of RATED THERHAL POVER). Power Range, Neutron Flux High Rates The Power Range Positive Rate trip provides protection against rapid flux increases which are characteristic of rod ejection events f rom any power level. Specifically, this trip complements the Power Range Neutron Flux High and Lov trips to ensure that the criteria are met for rod ejection from partial power. The Power Range Negative Rate trip provides protection to ensure that the DNB design criterion is met for control rod drop accidents. At high power a multiple rod drop accident could cause local flux peaking which, when in conjunction with nuclear power being maintained equivalent to turbine power by action of the automatic rod control system, could cause an unconservative local DNBR to exist. The Power Range Negative Rate trip vill prevent this from occurring by tripping the reactor for multiple dropped rods. No credit was taken for operation of this trip in the accident analyses; however, its functional capability at the specified trip setting is required by this specification to enhance the overall reliability of 'he Reactor Protection System. O FARLEY - UNIT 1 B 2-3 AMENDHENT NO.

LIMITING SAFETY SYSTEM SETTINGS BASES /N - Q 1atter trip vill ensure that the DNB design criterion is met during normal l operational transients and anticipated transients when 2 loops are in operation and the Overtemperature delta T trip setpoint is adjusted to the value specified for all loops in operation. .Vith the Overtemperature delta T trip setpoint adjusted to the value specified for 2 loop operation, the P-8 trip at 66% RATED THERHAL POVER vill ensure that the DNB design criterion is met during l normal operational transients and anticipated transients with 2 loops in operation. Steam Generator Vater Level The Steam Generator Vater Level Lov-Lov trip provides core protection by i preventing oper. tion with the steam generator vater Itvel below the minimum volume required for adequate heat removal capacity. The specified setpoint provides allovance that there vill be sufficient water inventory in the steam generators at the time of trip to allow for starting delays of the auxiliary feedvater system. Steam /Feedvater Flov Hismatch and Lov Steam Generator Vater Level The Steam /Feedvster Flov Hismatch in coincidence with a Steam Generator Lov Vater Level trip is not used in the transient and accident analyses but is included in Table 2.2-1 to ensure the functional capability of the specified trip settings and thereby enhance the overall reliability of the Reactor Protection System. This trip is redundant to the Steam Generator Vater Level Lov-Lov trip. The Steam /Feedvater Flov Hismatch portion of this trip is activated when'the steam flov exceeds the O feedvater flov by greater than or equal to 1.55 x 10 lbs/ hour. The steam Generator b Lov Vater Level portion of the trip is activated when the water level drops belov 25 percent, as indicated by the narrov range instrument. These trip values include sufficient allovance in excess of normal operating values to preclude spurio n tripe but vill initiate a reactor trip before the steam generators are dry. Therefore, the required capacity and starting time requirements of the auxiliary feedvater pumps are reduced and the resulting thermal transler. on the Reactor Coolant System and steam generators is minimized. Undervoltage and Underfrequency - Reactor Coolant Pump Busses The Undervoltage and Underfrequency Reactor Coolant Pump bus trips provide reactor core protection against DNB as a result of loss of voltage or under-frequency to more than one reactor coolant pump. The specified setpoints assure a reacior trip signal is generated before the lov flov trip setpoint f' (- FARLEY - UNIT I B 2-6 AMENDHENT NO.

REACTIVITY CONTROL SYSTEMS MODERATOR TEMPERATURE COEFFICIENT _ l L_IMITING CONDITION FOR OPERATION 3.1.1.3 The moderator temperature coefficient (MTC) shall bet

a. Less than er equal to 0.7 x 10 delta k/k/'F for the all rods withdrevn, beginning of cycle . life (BOL), condition for power levels up to 70% THERMAL POVER vith a linear ramp to O delta k/k/'F at-100%

THERMAL POVER.

b. Less negative than -4.3 x 10-4 delta k/k/'F for the all rods withdrawn, end of cycle life (EOL), RATED THERMAL POVER condition.

APPLICABILITY: Specification 3.1.1.3.a - MODES 1 and 2* only# Specification 3.1.1.3.b - MODES 1, 2 and 3 only# ACTION:

a. Vith the MTC more positive than the limit of 3.1.1.3.a above, operation in MODES 1 and 2 may proceed provided:

1. Control rod withdraval limits are established and maintained sufficient to restore the MTC to within its limit within 24 hours or be in HOT' STANDBY vithin the next 6 hours. These vithdrawal limits shall be in addition to the insertion limits L of Specification 3.1.3.6.

2. The control rods are maintained within the withdrawal limits established above until a' subsequent calculation verifies that the MTC has been restored to within its limit for the all rods withdrawn condition.

3. A Special Report is prepared and submitted to the Commission pursuant to Specification 6.9.2.vithin 10 days, describing the value of the meanured MTC, the interim control rod withdrawal limits and the predicted average core burnup necessary for restoring the prisitive MTC to within its limit for the all rods withdrawn condition.

b. Vith the MTC more ::egative than the limit of 3.1.1.3.b above, be in HOT SHUTD0VN vithin 12 hours.

Vith K,,, greater than or equal to 1.0 0 See Special Test Exception 3.10.3 O

FARLEY - UNIT 1 3/4 1-4 AMENDMENT NO.

l REACTIVITY CONTROL SYSTEMS ROD DROP TIME hd LIMITING CONDITION FOR OPERATION 3.1.3.4 The individual full length (shutdovn and control) rod drop time from the fully withdrawn position (225 to 231 steps, inclusive)* shall be less than or equal to 2.7 seconds from beginning of decay of stetionary gripper l coil voltage to dashpot entry with:

a. T, y greater than or equal to 541'F, and

b. All reactor coolant pumps operating.

APPLICABILITH HODES 1 and 2. ACTION:

a. Vith the drop time of any full length rod determined to exceea the above limit, restore the rod drop time to within the above limit l prior to proceeding to H0DE 1 or 2. l

b. Vith the rod drop times within limits but determined with 2 reactor coolant pumps operating, operation may proceed provided THERHAL POVER is restricted to less than or equal to 66% of RATED THERMAL POVER.

O SURVEILLANCE REQUIREHENTS _ 4.1.3.4 The rod drop time of full length rods shall be demonstrated through measurement prior to reactor criticality:

a. For all rods following each removal of the reactor vessel head,

b. For specifically affected individual rods following any

                      .u ntenance on or modification to the control rod drive system vhach could affect the drop time of those specific rods, and

c. At least once per 18 months.

          *The fully withdrawn position used for determining rod drop time shall be

_ greater than or equal to the fully withdrawn position used during subsequent plant operation. FAPLEY - UNIT 1 3/4 1-19 AMENDHENT NO.

POVER DISTRIBUTION LIMITS 3/4.2.2 HEAT FLUX HOT CilANNEL FACTOR F,(ZJ LIMITING CONDITION FOR OPERATION 3.2.2 F9 (Z) shall be limited by the following relationships: F,(Z) 5 [2.45] [K(Z)] for P > 0.5 for VANTAGE 5 fuel l P F,(Z) f [4.9) lK(Z)] for P $ 0.5 for VANTAGE 5 fuel and l F,(Z) f [2.32] [K(Z)) for P > 0.5 for LOPAR fuel l P F,(Z) f [4.64] {K(Z)] for P f 0.5 for LOPAR fuel l vhere P = THERMAL POVER RATED THERMAL POVER and K(Z) is the function obtained from Figure (3.2-2) for a given core height location. APPLICABILITY: MODE 1 O ACTION: Vith F,(Z) exceedit.g its limits

a. Reduce THERMAL POVER at least 1% for each 1% F,(Z) exceeds the limit within 15 minutes and similarly reduce the Power Range Neutron Flux-High Trip Setpoints within the next 4 hourst POVER OPERATION may proceed for up to a total of 72 hours; subsequent POVER OPERATION may proceed provided the Overpover delta T Trip Setpoints have been reduced at least 1% for each 1% F (Z) exceeds the limit. The Overpower delta T Trip Setpoint reduction siall be performed with the reactor in at least HOT STANDBY.

b. Identify and correct the cause of the out of limit condition prior to increasing THERMAL POVER above the reduced limit required by a above; THERMAL T0VER may then be increased provided F,(Z) is demonstrated through incore mapping to be within its limit FARLEY - UNIT 1 3/4 2-4 AMENDMENT NO.

O 1.2 o.0.1.0 s.o,1.o 12.0,0.933 bo 0.8 u. O - ul N

              .3 4 0.6      -

E cc O . Z 8 0.4

 ,]            x 0.2 0

O 2 4 6 8 10 12 CORE HEIGHT (FEET) Figure 3.2-2 K(Z) Normalized F,(Z) as a Function of Core Height O /ARLEY - UNIT 1 3/4 2-7 AMENDMENT NO.

__ _ . .- _.- _ _ _ _ . _ _ . . _ _ _ _ - - . _ _ _ _ _ _ _ _ _ , _ . _ _ . . _ . _ . _ . . _ = . _ - - _ . _ l POVER DISTRIBUTION LIMITS 3/A.2.3 NUCLEAR ENTHALPY HOT CHANNEL FACTOR - F",, LIMITING CONDITION FOR OPERATION _ l 3.2.3 F",, shall be limited by the following relationships F",, f 1.65 [1 + 0.3 (1-P)] for VANTAGE 5 fuel and l F",,'f 1.55 [1 + 0.3 (1-P)] for LOPAR fuel l l vhere P = THERMAL POVER RATED THERMAL POVER APPLICABILITY:_ MODE 1 ! l l ACTION: ) Vith F",, exceeding its limit:~ l

a. Reduce THERMAL POVER to less than 50% of RATED THERMAL.POVER vithin 2 .

hours and reduce the Pcver Range Neutron Flux-High Trip Setpoints to < ~ 55% of RATED THERMAL POVER vithin the next 4 hours,

b. is within its limit within Demonstrate 24 hours afterthrough exceeding in-core mapping the limit or redueethat F",,THEPEAL POVER to less than O 5% of-RATED THERMAL POVER vithin the next 2 hours, and

c. Identify and correct the cause of the out of limit condition prior to increasing THERMAL POVER above the reduced limit required by above; subsequent POVER OPERATION may proceed provided is that Fg or b, demonstrated through in-core mapping to be within its limit at,,a nominal 50%-of RATED THERMAL POVER prior to exceeding this THERMAL POVER, at a-nominal 75% of RATED THERMAL POWER prior to exceeding this THERMAL POWER 1 and within 24' hours after attaining 95% or greater RATED THERMAL POVER._

h

               = FARLEY - UNIT 1                                                                                                                         3/4 2-8                                             AMENDMENT NO.

s.,..,, .-.s, ,n . . , . . , . . , ~ , . , . . . , - - . , . , . _ . . . . , . . _ , . _ , . , . ,. _ . , . . , - . . . . . - , . , , . - -

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

POVER DISTRIBUTION LIMITS DNB PARAMETERS () LIMITING CONDITION FOR OPERATION _ , , _ _ _ , 3.2.5 The following DNB related parameters shall be maintained within the limits shown on Table 3.2-1:

a. Reactor Coolant System T,,,

b. Pressurizer Pressure

c. Reactor Coolant System Total Plov Rate.

APPLICABILITY: H0DE 1 ACTION: Vith any of the above parameters exceeding its limit, restore the parameter to vithin its limit within 2 hours or reduce THERMAL POVER to less than 5% of RATED THERHAL POVER vithin the next 4 hours. SURVEILLANCE REQUIREMENTS _ _ _ 4.2.5.1 Each of the parameters of Table 3.2-1 shall be verified to be within O their limits at least once per 12 hours. 4.2.5.2 The Reactor Coolant System total flow rate shall be determined to be within its limit by measurement at 1 cast once per 18 months. 4.2.5.3 The indicated RCS flow rate shall be verified to be within the acceptable limit at least once per 31 days. FARLEY - UNIT 1 3/4 2-14 AMENDHENT NO. 1O ._

O O O TABLE 3.2-1 G

DNB PARAMETERS I

E LIMITS y ~ 3 Loops in operation 2 Loops in Operation PARAMETER f 580.7'F (**) l Indicated Reactor Coolant System T,,, 2 2205 psig* (**) l Indicated Pressurizer Pressure 1 267,880 gpm*** (**) l Indicated Reactor Coolant System Total Flow Rate T E 0 E E E

Limit not applicable during either a THERMAL POVER ramp in excess of 5% of RATED THERMAL POVER per minute or a THERMAL POVER step in excess of 10% of RATED THERMAL POVER.

5 ** Values blank pending NRC approval of 2 loop operation. l

                  ***  Value includes a 2.4% flow uncertainty (0.1% feedvater venturi fouling bias included).

0' L TABLE 3.3-2 h , REACTOR TRIP SYSTEM INSTRUNDrfATION RESPONSE TINES i n FUNCTIONAL UNIT RESPONSE TIME j r' E r H 1. Manual Reactor Trip .Not Applicable !' 2. Power Range, Neutron Flux.

a. High .f 0.5 seconds

2

b. Lov Not Applicatie

3. Power Range, Neutron Flux, o High Positive Rate ..Not Applicable l

4. Power Range, Neutron Flux, High Negative Rate' Not Applicable l

w !

g 5. Intermediate Range, Neutron Flux Not Applicable w .6. Source Range, Neutron Flux 'Not Applicable

7. Overtemperature aT f 6.0 seconds *
8. Overpower.AT Not Applicable i

{ 9. Pressurizer Pressure-Lov f 2.0 seconds ,

10. Pressurizer Pressure-High f 2.0 seconds l l [

4 .

11. Pressurizer Vater Level-High Not Applicable i

[ . E i 4 E : a M I'i

Neutron detectors are exempt from response time testing. Response time of the neutron flux signal portion [

1 z of the channel shall be measured from detector output or input of first electronic component in channel.

          ?                                                                                                                              !

1 f I I i

     .             .  . - ~                            .-   ,,             _     ..          _ . ,        -   . _ , . ________.      ___

P (- . m,

                                                            ).                                                                                                                                           [

e s TABLE 3.3-4 (Contim.cd) Si G ENGINEERED SAFETT FEATURE ACTUATION SYSTEM INSTRUEENTATION TRIP SETPOINTS t FUNCTIONAL UNIT TRIP SETPOINT ALLOUABLE VALUES

 ,       2. CONTAINMENT SPRAY

a. Manual Initiation Not Applicable Not Applicable

b. Automatic Actuation Logic Not Applicable Not Applicable

c. Containment Pressure- $ 27 psig $ 28.3 psig l High-High-High

3. CONTAINMENT ISOLATION y a. Phase "A" Isolation

v. '

1. Manual Not Applicable Not Applicable 5 2. From Safety Injection Not Applicable Not Applicable !

Automatic Actuation Logic

b. Phase "B" Isolation

1. Manual Not Applicable Not Applicable

2. Automatic Actuation Logic Not Applicable Not Applicable

3. Containment Pressure- $ 27 psig $ 28.3 psig l High-High-High h c. Purge and Exhaust Isolation z

E 1. Manual Not Applicable Not Applicable M

2. Automatic Actuation Logic Not Applicable Not Applicable l 5

b

O O O

TAELE 3.3 4 (Continued)

ENGINEERED SAFETT FEATURE ACTUATION SYSTEM INSTRUMENTATION TRIP SETPOINTS 4 " TRIP SETPOINI ALLOVABLE VALUES FUNCTIONAL UNIT

4. STEAM LINE ISOLATION Not Applicable Not Applicable

a. Manual Not Applicable Not Applicable

b. Automatic Actuation Logic f 16.2 psig $ 17.5 psig l

c. Containment Pressure--

g High-High e-

                                                        < A function defined as follows:   < A function defined as follows:

d. Steam Flov in Two Steam A op corresponding to 40% of full A op corresponding to 44% of full Y' Lines--Bigh, Coincident steam flow between 0% and 20% load E'

steam flow between 0% and 20% load and then a op increasing linearly si th T--Low-Low and then a op increasing linearly to a op corresponding to 110% of to a op corresponding to 111.5% of full steam flow at full load with full steam flov at full load with T,,, 2 540'F T,,, 2 543*F 2 Sa5 psig 2 575 psig

e. Steam Line Pressure--Lov

5. TURBINE TRIP AND FEED VATER ISOLATION f 75% of narrov range instrument f 76% of narrow range instrument j Steam Generator Vater i

a. span each steam generator span each steam generator EE Level--High-High E E e 4 5

_ . - - = _ _ . - - I TABLE 3.3-5 (Continued) p ENGINEERED SAFETY FEATURES RESPONSE TIMES I 'A INITIATING SIGNAL AND FUNCTION RESPONSE TIME IN SECONDS

3. Pressurizer Pressure-Lov

a. Safety Injection (ECCS) 3 27.0'8'/12.0

b. Reactor Trip (from SI) f2.0 l

c. Teedvater Isolation f 32.08 '

d. Containment Isolation-Phase "A" f 17.0

e. Containment Purge Isolation f 5.0

f. Auxiliary Feedvater Pumps Not Applicable

g. Service Vater System f 77.0/87.0

4. Differential Pressure Between Steam Lines-liigh

a. Safety Injection (ECCS) f 12.0/22.0

b.- Reactor Trip (from SI) f 2.0

c. Feedvater Isolation f 32.0

d. Containment Isolation-Phase "A" f 17.0/27.C i

e. Containment Purge Isolation Not Applicable

f. Auxiliary Feedvater Pumps Not Applicable

[] V g. Service Vater System f 77.0/87.0

5. , Steam Flov in Two Steam Lines-!!!gh coincident with T,y --Lov-Lov

a. Steam Line Isolation Not Applicable l

6. Steam Line Pressure-Lov

a. Safety Injection (ECCS) $ 12.0/22.0'58

b. Reactor Trip (from SI) f 2.0

c. Feedvater Isolation f 32.0

d. Containment Isolation-Phase "A" f 17.0/27.0

e. Containment Purge Isolation Not Applicable

f. Auxiliary Feedvater Pumps Not Applicable

g. Service Vater System j 77.0/87.0

h. Steam Line Isolation f 7.0 n

v FARLEY - UNIT 1 3/4 3-30 AMENDHENT NO.

REACTOR COOLANT SYSTEM HOT STANDBY LIMITING CONDITION FOR OPERATION 3.4.1.2 At least two of the Reactor Coolant Loops listed belov shall be OPERABLE and in operation when the rod control system is operational or at least two Reactor Coolant Loops listed belov shall be OPERABLE vith one ReLetor Coolant Loop in operation when the rod control system is disabled by opening the Reactor Trip Breakers or shutting dovn the rod drive motor / generator sets *

1. Reactor Coolant Loop A and its associated steam generator and Reactor Coolant pump,

2. Reactor Coolant Loop B and its associated steam generator and Reactor Coolant pump,

3. Reactor Coolant Loop C and its associated steam generator and Reactor Coolant pump.

APPLICABILITY: MODE 3 ACTION:

a. Vith less than the above required Reactor Coolant Loops OPERABLE, restore the required loops to OPERABLE status within 72 hours or be in HOT SHUTD0VN vithin the next 12 hours.

b. Vith only one Reactor Coolant Loop in operation and the rod control (i rystem operational, within 1 hour open the Reactor Trip Breakers or shut devn the rod drive motor / generator sets.

c. Vith no Reactor Coolant Loops in operation, suspend all operations involving a reduction in boron concentration of the Reactor Coolant System and immediately initiate corrective action to return the required coolant loop to operation.

SURVEILLANCE REQUIREMENTS 4.4.1.2.1 At least the above required Reactor Coolant pumps, if not in operation, shal.'. be determined to be OPERABLE once per 7 days by verifying correct breaker alignments and indicated power availability. 4.4.1.2.2 The required Reactor Coolant Loop (s) shall be verified to be in operation and circulating Reactor Coolant at least once per 12 hours. 4.4.1.2.3 The required steam generator (s) shall be determined OPERABLE by verifying secondary side water level to be greater than or equal to 10% of vide range indication at least once per 12 hours.

All Reactor Coolant pumps may be de-energized for up to I hour provided (1) no operations are permitted that vould cause dilution of the Reactor Coolant O System boron concentration, and (2) core outlet temperature is maintained at least 10'F belov saturation temperature.

FARLEY - UNIT 1 3/4 4-2 AMENDMENT NO.

3/4.2 POVER DISTRIBUTION LIMITS BASES O~ The specifications of this section provide assurance of fuel integrity during Condition I (Normal Operation) and II (Incidents of Moderate Frequency) events by: (a) meeting the DNB design criterion during normal operation and l in short term transients, and (b) limiting the fission gas release, fuel pellet temperature and cladding mechanical properties to within assumed design ; criteria. In addition, limiting the peak linear pover density during Condition I events provides assurance that the initial conditions assumed for the LOCA analyses are met and the ECCS acceptance criteria limit of 2200'T is not exceeded. The definitions of certain hot channel and peaking factors as used in , these specifications are as follows: 1 F,(Z) 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 and measurement uncertainty. F(H 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. Fg(Z) Radial Peaking Factor, is defined as the ratio of peak power density to average power density in the horizontal plane at core elevation Z. 3/4.2.1 AXIAL FLUX DIFFERENCE The limits on AXIAL FLUX DIFFERENCE (AFD) assure that the F (Z) upper bound envelope of 2.45 for VANTAGE 5 and 2.32 for LOPAR times th,e normalized l axial peaking factor is not exceeded during either normal operation or in the event of xenon redistribution following power changes. Target flux difference is determined at equilibrium xenon ecnditions. The full length rods may be positioned within the core in accordance with their respective insertion limits and should be inserted near their normal position for steady state operation at high power levels. The value of the target flux difference obtained under these conditions divided by the fraction of RATED THERMAL POVER is the target flux difference at RATED THERMAL POVER for the associated core burnup conditions. Target flux differences for other THERMAL POVER levels are obtained by multiplying the RATED THERMAL POVER value by the appropriate fractional THERMAL POVER level. The periodic updating of the target flux difference value is necessary to reflect core burnup considerations. O FARLEY - UNIT 1 b 3/4 2-1 AMENDMENT NO.

l POVEh DISTRIBUTION LIMITS BASES l AXIAL FLUX DIFFERENCE (Continued) l Although it is intended that the plant vill be operated with the AFD vithin the 4(5)% target band about the target flux difference, during rapid plant THERMAL POVER reductions, control rod motion vill cause the AFD to deviate outside of the target bu,d at reduced THERMAL POVER levels. This deviation vill not affect the xenon redistribution sufficiently to change the envelope of peaking factors which may be reached on a subsequent return to RATED THERMAL POVER (with the AFD vithin the target band) provided the time duration of the deviation is limited. Accordingly, a 1 hour penalty deviation limit cumulative during the previous 24 hours is provided for operation outside of the target band but within the limits of Figure (3.2-1) vhile at THERNAL POVER levels between 50% and 90% of RATED THERMAL POVER. For THERMAL POVER levels between 15% and 50% of RATED THERMAL POVER, deviations of the AFD outside of the target band are less significant. The penalty of 2 hours actual time reflects this reduced s'gnificance. Provisions for monitoring the AFD on an automatic basis are derived from the plant process computer through the AFD Monitor Alarm. The computer determines the one minute average of each of the OPERABLE excore detector outputs and provides an alarm message immediately if the AFD for 2 or more OPERABLE excore channels are outside the target band and the THERMAL POVER is greater than 90% of RATED THERMAL POVER. During operation at THERMAL POVER O levels between 50% and 90% and between 15% and 50% RATED THERMAL POVER, the computer outputs an alarm message when the penalty deviation accumulates beyond the limits of I hour and 2 hours, respectively. Figure B 3/4 2-1 shows a typical monthly target band. 3/4.2.2 and 3/4.2.3 IIEAT FLUX Il0T CHANNEL FACTOR, NUCLEAR ENT!!ALPY !!OT CHANNEL FACTOR The limits on heat flux hot channel factor, and nuclear enthalpy rise hot channel factor ensure that 1) the design limit on peak local power density is not exceeded, 2) the DNB design criterion is met, and 3) in the event of a LOCA the peak fuel clad temperature vill not exceed the 2200'F ECCS acceptance criteria limit. Each of these is measurable but vill normally only be determined periodically as specified in Specifications 4.2.2 and 4.2.3. This periodic surveillance is sufficient to insure that the limits are maintained provided:

a. Control rods in a single group move together with no individual rod insertion differing by more than 2 12 steps, indicated, from the group demand position,

b. Control rod banks are sequenced with overlapping groups as described in Specification 3.1.3.6.

g-~g c. The control rod insertion limits of Specifications 3.1.3.5 and ( j 3.1.3.6 are maintained.

d. The axial power distribution, expressed in terms of AXIAL FLUX DIFFERENCE, is maintained within the limits.

FARLEY - UNIT 1 B 3/4 2-2 AMENDMENT NO.

POVER DISTRIB11 TION LIMITS BASES F" through,H vill beare

d. above maintained maintained. within Theitsrelaxation limits providgd of F conditions a.

THERHAL POVER allows changes in the radial power for shape,H as a function of all permissible rod insertien limits. When an F measurement is taken, an allovance for both experimental errorandmanufacturingtolerancemustbemade. An allovance of 5% is appropriate for a full core map taken with the incore detector flux mapping system and a 3% allovance is appropriate for manufacturing tolerance. When F" the appropria,H is measured, te allowance for a experimentalfull core map error takenmust be allowed with the incore for and 4% is detection system. The specified limit for F H contains an 8% allovance for uncertainties. The8%allovanceisbasedonthefollowingconsiderations:

a. Abnormal perturbations in the radial power shape, such as from rod misalignment, affect F",H more directly than F,,

b. Although rod movement has a direct influence upon limiting F to vjthinitslimit,suchcontrclisnot readilyavailabletoISmit F ,H, and

c. Errors in prediction for control power shape detected during startup

( t physics testsThis distribution. can compensation be compensated for for F ,H by restricting axial flux ip F,is less readily available. FARLEY - UNIT 1 B 3/4 2-4 AMENDMENT No.

POVER DISTRIBUTION LIMITS BASES The radial that assurance peaking factor the hot F,h(Z), chann is measured factor F periodically to provide additional The F ( limit for RATED THERHAL POVER (pRTP) a,s pr(Z), remains within its limit.ovided in the Ra limit report per Specification 6'5.1.11 was' determined from expected power control maneuvers over the full range of burnup conditions in the core. 3/4.2.4 OUADRANT POVER TILT RATIO The quadrant power tilt ratio limit assures that the radial power distribution satisfies the design values used in the power capability analysis. Radial power distribution measurements are made during startup testing and periodically during power operation. The limit of 1.02, at which corrective action is required, provides DNB and linear heat generation rate protection with x-y plane power tilts. The two hour time allovance for operation with a tilt condition greater than 1.02 but less than 1.09 is provided to allow identification and correction of a dropped or misaligned control rod. In the event such action does not correct the tilt, the margin for uncertainty on F is reinstated by reducing the maximum allowed power by 3 percent for each perce,nt of tilt in excess of 1.0. For purposes of monitoring OUADRANT POVER TILT RATIO vhen one excore detector is inoperable, the movable incore detectors are used to confirm that the normalized symmetric power distribution is consistent with the OUADRANT POVER TILT RATIO. The incore detector monitoring is done with a full incore flux map or two sets of four symmetric thimbles. The two sets of four symmetric thimbles is a unique set of eight detector locations. These locations are C-8, E-5, E-11, H-3, H-13, L-5, L-11, and N-8. 3/4.2.5 DNB PARAMETERS The limits .on the DNB related pararacters assure that each of the parameters are maintained within the normal steady state envelope of operation assumed in the transient and accident analyses. The limits are consistent with the initial FSAR assumptions and have been analytically demonstrated adequate to meet the DNB design criterion throughout each analyzed transient. The indicated T value of 580.7'F is based on the average of two control board readings an8,In indication uncertainty of 2.5'F. The indicated pressure value of 2205 psig is based on the average of two control board readings and an indication uncertainty of 20 psi. The indicated total RCS flow rate is based on one elbow tap measurement from each loop and en uncertainty of 2.4% flow (0.1% flov is included for feedvater venturi fouling). The 12 hour surveillance of Tavg and pressurizer pressure through the control board readings are sufficient to ensure that the parameters are restored within their limits following load changes and other expected transient operation. The 18 month surveillance of the total RCS flow rate is a precision measurement

    =that verifies the RCS flow requirement at the beginning of each fuel cycle and ensures correlation of the flow indication channels with the measured loop flows. The monthly surveillance of the total RCS flow rate is a reverification of-the RCS flow requirement using loop elbow tap measurements that are correlated to the precision RCS flow measurement at the beginning of the fuel                                                                    -

A cycle. The 12 hour RCS flov surveillance is a qualitative verification of () .significant flow degradation using the control board indicators and the loop elbov tap' measurements that are correlated to the precision RCS flow measurement at the beginning of each fuel cycle. FARLEY - UNIT 1 B 3/4 2-5 AMEPDMENT NO.

3/4.4 REACTOR COOLANT SYSTEM BASES O 3/4.4.1 REACTOR COOLANT LOOPS AND COOLANT CIRCULATION The plant is designed to operate with all Reactor Coolant Loops in operation, and meet the DNB design criterion during all normal operations and l anticipated transients. In H0 DES 1 and 2 vith one Reactor Coolant Loop not in operation this specification requires that the plant be in at least HOT STANDBY vithin 1 hour. In H0DE 3, two Reactor Coolant Loops provide sufficient heat removal l capability for removing core heat even in the event of a bank withdrawal accident; however, a single Reactor Coolant Loop provides sufficient decay heat i removal capacity if a bank withdrawal accident can be prevented: 1.e., by opening the Reactor Trip Breakers or shutting down the rod drive motor / generator sets. l In HODE 4, a single reactor coolant or RHR loop provides sufficient heat removal capability for removing decay heat, but single failure considerations require that at least two loops be OPERABLE. Thus, if the Reactor Coolant Loops are not OPERABLE, this specification requires two RHR loops to be OPERABLE. In H0DF. 5, single failure considerations require two RHR loops to be OPERABLE. The operation of one Reactor Coolant Pump or one RHR pump provides adequate flov to ensure mixing, prevent stratification and produce gradual reactivity changes during boron concentration redections in the Reactor Coolant System. 9 The reactivity change rate associated with boron reduction vill, therefore, be vit.' '.a the capability of operator recognition and control.

   ,             t i restrictions on starting a Reactor Coolant Pump with one or more Reactor Coolant System cold legs less than or equal to 310*F are provided to prevent Reactor Coolant System pressure transients, caused by energy additions from the secondary system, which could exceed the limits of Appendix G to 10 CFR Part 50. The Reactor Coolant System vill be protected against overpressure transients and vill not exceed ti.e limits of Appendix G by either (1) restricting the water volume in the pressurizer and thereby providing a volume for the primary coolant to expand into, or (2) by restricting starting of the 1

Reactor Coolant Pumps to when the secondary water temperature of each steam f generator is less than 50'F above each of the Reactor Coolant System cold leg temperatures. O FARLEY - UNIT 1 B 3/4 4-1 AMENDMENT NO.

ADMINISTRATIVE CONTROLS A) e. Type of container (e.g. , LSA, Type A, Type I, Large Quantity), and (J

f. Solidification agent (e. g., cement, urea formaldehyde).

The radioactive effluent release reports shall include unplanned releaLes from the site to unrestricted areas of radioactive materials in gaseous and liquid effluents on a quarterly basis. The radioactive eff1'ent release reports shall include any changes to the PROCESS CONTROL PROGRAM (N ; n during the reporting period. MONTHL.' GiikATING i #RT o.9.1.10 Routine reports of operating statistics and shutdown experience, including documentation of all challenges to the PORV's or safety valves, shall be submitted on a monthly basis to the Commission, pursuant to 10 CFR 50.4, no later than the 15th of each month following the calendar month covered by the report. Any changes to the OFFSITE DOSE CALCULATION MANUAL shall be submitted with the Monthly Operating Report within 90 days in which the change (s) was made effective. In addition, a report of any major changes to the radioactive vaste treatment systems-shall be submitted with the Monthly Operating Report for the period in which the change was implemented. RADIAL PEAKING FACTOR LIMIT REPORT O) 6.9.1.11 The F limit for RATED THERMAL POVER (FRTP Commission, pursuEnt to 10 CFR 50.4, for all core plaEes) shall bebank containing provided "D" to the control rods and all unrodded core planes no later than 30 days after cycle l initial criticality. In the event that the limit vould be submitted at some other time during core life, it vill be submitted 30 days af ter the date the l limit vould become effective unless otherwise exempted by the Commission. Any information needed to support F%P vill be by request from the NRC and need not be included in this report. ANNUAL DIESEL GENERATOR RELIABILITY DATA REPORT 6.9.1.12 The number of tests (valid or invalid) and the number of failures to start on demand for each diesel generator shall be submitte' to the NRC annually. This report shall contain the information identified in Regulatory Position C.3.b of NRC Regulatory Guide 1.108, Revision 1, 1977. (. U FARLEY - UNIT 1 6-19 AMENDMENT N0. x . w ,.,

O E Unit 2 Marked-up Paaes O O

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Figure 2.1-1 Reactor Core Safety Limits Three Loops in Operation O O e

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stactua Tatt sYsten tiesvatsnorratices Tatt sarrotstes e a s007AT3088 Isote It [0%rteeperatge srl s Ar, lK,-E, I e t,5 R ATN . E, (P - r') - f, ( AI){ u '** (T { ,' Q T ' AT e Measured Af by m.Tb *nsera,menfa+b i l shoret W , . Indicated K at SATED TWERital POWER T = Average temperature. 'F l T' S 577.2*F (teamleen Reference T,,, at RATED TWEENAL POWER) P = Preetcarlser pressere, pelg l 9- ! P' . 2235 pels (IIemlut Sc3 operating pressere) f ! u 1 e t,$ l & 1 + r,5 = The functies geneested by the lead-leg centroller for T,,, dyneele compensatie's ! sed 5CitT T & t, . Ties constants stilised le the lead-lag eestreller for t T,,, t, 30 secs, t, . 4 secs. p p , S 1.aplace transform operator, see. eperntles eith 3 loops

Et a E M operation with 2 loops E a= E, . (valese bleelt pendlag l

, dE W' *" E, . NBC approvel 6f t, . M ~ and f g, . 2 leep operation) y (SF) is a tenettee of the ladicated difference between top and bettee detectors of the po chambe,rs3 eith gales to be selected based en seasured lastrumentwer-range respoetse durin moelear ten uch that:

E.

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         .e TASLE 2.2-1 (Contfaced)

REACTOR TRIP $T57298 IstsTRepeterTATIce TRtr SETPetHTS i SIOTATION coatlowed E (1) Q for g - betweenhpercentand* p"ercent , f (AI) . O TsaastlL in the top and bettee ma o (ves of the, core respec(where g, and g are percent Rafts ! reven In percent of RATED TusanAL rowsa). tively, and g, + g, is total Tatestat : (II) be seteestically reduced by s30 percentpercent, of its value the a tripat RATED setpelet shall 7, i (g, - g,) escoeds . @ (Ill) for each percent that the be seteestically reduced by percent, the R trip setpolet shall percent of Its value at matte Tatastal rotta. i+ Ts3 i te 2. N e m s n.ia, - r, 1 4, \ - n. f , T(jkd-f , i

                                                                                              -f,(an i                                                                                            '

se steeres lus m M -= Indicated N at RATED TWER8tAL PovER - C. T I y NetE . .eerage te.,erature. r (, I , y,s ,T . s y u rt.s y T* . Reference T,,, f SFF.2*F) at RATED TWERflAL POWER (Calibraties temperature for K lastrumentatl es, , n, iwei6 i g R E ,,.ft.D4194/'F

                                             '. 0.02/*F.

for leeressing average temperature and 8 for dec Jer T > T*3 K, O for T f T* I

                      ,             NI                                                                                                                                                                      i e
                                   @ - The functies generated by the rate las centroller for T,,, dyneele tempensatise                                                                                     !

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                                                                                                         ,weasured "T'a9 2                                                                   '

I + t.3 g2 Ti.e coa =ht ut;Itted ia *e ==-sund Tag 15 c peosma.c, p o ,,, . I 9 f INSER.T 0. AT = Mensure d AT Q RTb insfrumulc.%s t

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9 2.1 $AFITV t.1MIT5 . ( nsts 2.1.1 tt ACTOR CDRI "'. The restrictions of this Safety 1.1mit prevent everheating of the fuel and possible cladding perforation which would result in the release of fission products to the reactor coolant. Overheating of the fuel cladding is prevented by restricting fuel operation to within the nucleate boiling regime where the beat transfer coefficient is large and the cladding surface temperature is

  - slightly above the coolant saturation temperature.

Operation above the upper boundary of the nucleate boillag regime could [8eg fAsult 3 nucleate in excessive cladding teeperatures because of the enset of departure from I;m; w .,I 2 boiling (DNB) and the resultant sharp reduction in heat tran:fer arme y coefficient. DNB is not a directly measurable parameter during operation and

tnerefore THIRMAL p0ER and anacter caelant Tareerature and pressure have been I related to Dh5 througK k to 17 r au^4 . ';a 0 0 M ie J e R :..lbeen developed to predict the DNB fluz and the locction of DNB for axialis uniform an

       $ non-uniform heat flux distributions. The local DN8 heat flux ratio, DNBR,                                                                                        gat $ ggt p defined as the ratic of the heat flux that umuld cause DNB St a particular core location to the local heat flux, is indicative of the margin to DNS.                                                                                          ,

O (/ ratio m1D tran valu mts, f the antic R du ed tr staa 1ents att a limit tion, 1.3 . This (' alu erres ds to 5 pe et pro 111ty a 95 p ent to ence vel

           . t DN              11 no     ccur        is            n as          sprep     te mar                  to D           or a sper        ng c           itions y                                                  Ialstti E The curves of Figures 2.11 and 2.12 show the ".d ef ^# :' of THERMAL                                                                                                        .

podR, Raatter Coolant System pesssure and eversee tencerature f".. - . . ~ . j _.____,___m__ , ,- _ _ _ ._

                                                                                          ;_.                3. .           =g                               .. g
           ;;.d tc tu ;^;nd;,, ;' ::t. ;;; li;.id.                                                                                                                           i v_ __ _--_-                   _ ----- -- an enthalpy het channel factor. F a ef S and a reference cosine with a peat of 1.55 for asial gewer shape. Anblevanceis included for an increase in F N at reduced power based en the expression:

AH , 3 IfAWTME 5 Let e.JI F = S [1+0.3(1.P)) 7 where p is the fraction of RATED M.iAAL PodR These limiting heat flux conditions are higher than those calculcted for the . I range of all control rods fully withdrawn to the maxima allowable control red ) insertion assuming the axial power imbalance is within the limits of the , .1 - _.- l i L65 4.r W4c 5 F,1= .ssD+ o.m-P)J for LOPAR bel blM an Fgof155t. LoFKR $.el g

             'ARLEY.UMIT 2                                          821                                                  8'rNNiaN                                                            -

i M*5 *b m f.1-1 a4 Z.l.2. a.rc bueA m & M.) Q%cua ug _ . l

INSERT D The DNB thermal design critarion is that the probability of DNB not occurring'on _the most limiting rod is at least 95% (at a 95% confidence

level)_for any condition I or 11 event. In meeting the DNB design criterion, uncertainties in plant. operating parameters, nuclear and thermal parameters, fuel fabrication parameters, and computer codes.must be considered. As described in the FSAR, the effects of these uncertainties have been statistically combined with the correlation uncertainty. Design limit DNCR values have been determined that satisfy the DNB design criterion. Additional DNBR margin is maintained by performing the safety analyses to a higher DNBR limit. This margin between the design and safety analysis limit DNBR values is used to offset known DNBR penalties (e.g., rod bow and transition core) and to provide DNBR margin for operating and design a flexibility. l INSERT E ; which satisfy the following criteria:

       ~A. The average enthalpy at the vessel exit is less than the enthalpy of saturated liquid (far left line segment in each curve).

B. The minimum DNBR satisfies the DNB-design criterion (all the other line segments in each curve). Each curve reflects the most limiting result using either low-parasitic (LOPAR) fuel or VANTAGE-5_ fuel. The 1 VANTAGE-5 fuel is analyzed using the WRB 2 correlation with design limit DNBR values of 1.24 and 1.23 for the typical and thimble cells, respectively. 'The LOPAR fuel is analyzed using the WRB-1 correlation with design limit DNBR values of 1.25 and 1.24 for the typical and thimble cells,-respectively. C. The hot channel exit quality is not greater than the upper limit of the quality-range-(including the effect of uncertainties) of the DNB correlations. This is not a limiting criterion for this plant.

2. 2 '!MfitNG

                             .        SAFE Y SYSTEM SETTINGS BASES 2.2.1           REACTOR TRIP SYSTEM INSTRUMENTATION SETPo!NTS                                                        ,,

The Reactor Trip setpoint Limits specified in Table 2.2-1 are the values at which the Reactor Trips are sat for each functional unit. The Trip Set-points have been selected to ensun that the rsactor core and reactor coolant systen are prevented free exceeding their safety limits during normal operation and design basis anticipated operational occurrences and to assist the Engi-neered Safety Features Actuation Systee in sitigating the consequences of accioents. Operation with a trip set less conservative than its Trip Setpoint but within its specified Allowable Value is acceptable on the basis that the difference between each Trip setpoint and the Allowable value is equal to or less than the drift allowance assumed for each trip in the safety analyses. Manual Reacter Trip The Manual Reactor Trip is a redundant channel to the automatic protective instrussntation channels and provides manual reactor trip. capability. Power Rance, Neutren Flux p The Power Range, Neutron Flux channel high setpoint provides reactor core Q protection against enactivity excursions which are too rapid to be protected

        \- ', by temperature and pressure protective circuitry. The low set point provides redundant protection in the power range for a power excursion beginning from low power. The trip associated with the low setpoint say be manually bypapac when P-10 is cetive (two of the four power range channels indicate a power level of above approximately 10 percent of RATED THERMAL POWER) and is auto-matica11y reinstated when P-10 becomes inac,tive (three.of the four channels indicate a power level below approximately 8 percent of RATED THERMAL power).
          -         Pewer Rance. Neutron Flux, High Rates
           'lll, E                The Power Range Positive Rata trip presides protaction against rapid flux
           .T       increases which are charactaristic of rod ejection events free any power g     level. Specifically, this trip complements the Power Range Neutron Flux High

C, .and Low trips to ensure that the criteria are met for rod ejection from partial

( power.

O N The Power Rance Necative Rate trd p provides protaction to ensure that tne h l & ' = M R i: ; i n u i n ; ; ;;;; . ; f. M for control red drop accidents. At hign

             *h     power am n:: maultiple roc crop accident could cause local flux peaking
             '      which, wnen in conjunction with nucisar power being maintained equivalent to                                                    '

e tureine power by action of the automatic rod control systas, could causa an 2 unconservative local DNBR to exist. The Power Range Negative Rata trip will 9 prevent this from occurring by tripping the reactor for k Q-g : :-laultiple droceed rods. M in . l No trt.d.tcopoV wasH+ thka.n fee tied o$andon c4is $s hd; gd,howcrspu.heen se$g, acelantand ib Nnc%nd. Acsp.4 + t.m r Wru)

                    +o enhanc.s. the everall 7PO M. chht. Reat.Por bh+.'enys+t                                                           .

TA'kil.Y-UNIT 2 8 2-3

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

LIMITING SaFITY $YSTEM SETTING 5 o-Oc ~ Intemediate and_ $ource Ranee, Nuclear Flux - - I

                   . core The      Intamediate protection      duri and   Source Range, Nuclear Flux trips provide reactor reactor sta          .

These trips provide redundant protee-tion to the low setpo nt trf of the over Range, Neutron Flux ; SourceRangeChannelswilliitiateareactortripatabout10'ghannels. The

                                                                                                                   -counts per-second unless manus 11 blocked when p-6 becomes active.

The Intermediate Range Channels will i itiata a reactor trip at a current level proportional to 1 i approximately

                   -p-10  becomes active. 25 percent at 4ATED THERMAL POWEk unless manually blocke No JNJit was taken for operation of the trips associ-
                   -ated with either the Intermediate or Source Range Channels in the accident
                   ~ analyses; however, their functional capab.111ty at the specified trip setti Reactor protection Systas.is required by this specification to enhan
                                                          ~
                . Overtesoerature AT f                      .

f* "U l(for -all The Overteeperatun delta T trip pr combinations of pressure, power,ovides core prot: coolant temperature and axial power !

             .Ldistribution, provided that the transient is alow with respec,t to piping

, O and

                 -trips. pressure is within the nnge between the High and L This setpoint includes cornetions for changes in density and hesu creacity of water with tesoerature and evnamic compensation for ;iek; ;

L' a t; ar. u ;ha 4;; n;;ar;.t.r; ;;nura lWith normal aa1al power .A distribut<on, this reactar trip limit Is always below the core safety limit as the difference between.topiand bottee power range nu 2.2-1. reactor trip is automatically reduced according to the notations in Table setpoint Coos not require reactor protection systes setpo because the p-8 setpoint and associated trip will prevent DN8 during 2 loop operation exclusive of the Overtemperature delta T setpoint. Two loop operation above the 3 loop p-8 setpoint is permissible after resetting the K1, K2, and K3its'2 to inputs loop to'the Overtemperature solta T channels and raising the p-8 setpoin value. L functions as a High Neutron Flux trip at the reduced power-leve 9 l e bespek, Armowell and Rip etsp.ese he deleys $,wm

                            & e.m +. Rro cutdh Wrub.

\ O - FARLEY-UNIT 2 8 2-4 m

LIMITING SAFETY SYSTEM SETTINGS.- BASES __ Over:ower aT The m ower del. T reactor provides urance of 1 integr ,

                 . no melt      under a         essible av       ower condi           s, limits                        e requir ran       for Overte       ature del T protectio             and provi          a backup                          the Hign Ne        on Flux tr       The    setp   ' t includes         tractions         axial                     pc stributio      changes in         sity and        t capacit         f water wi                         temperatur an        namic com    s& tion for         ing dela       from the c          to the          lo                tempera-ture e         tors. No     edit was        ' n for ope      on of thi           ip in                     the   acident nalyses;       wever, its        nctional c bility a ' e specif                   trip satt                                 is re      red by th    specifica           to enhan       the overa         reliabili                         f  the Reacto       rotection _ stem.

Pressurizer Pressure The Pressurizer High and Low Pressure trips are provided to limit the pressure range in which reactor operation is permitted. The High Pressure

  .,          trip is backed up by the pressurizer code safety valves for RCS overpressure protection, and is therefore set lower than the set pressure for these valves p         *

(2485 psig). The Low Pressure trip provides protection by tripping the reactor V in the event of a loss of reactor coolant prer,ure. Pressurizer Water Level The Pressurizer High Water Level trip ensures protection against Reactor Coolant Systes overpressurization by limiting the water level to a volume sufficient to retain a steam bubble and preveat water relief through the pressurizer safety valves. No credit was taken for operatierf of this trip in the accident analyses; however, its functional capability at the specified trip setting is required by this specification to enhance the everall reliability of the Reactor Protection System. Loss of Flow The Loss of Flow trips provide core protection to prevent DNB in the event of a loss of one or more reactor coolant pumps. Above 10 percent of RATED THERMAL POWER, an automatic reactor trip will occur if the flow in any two loops drop below 90% of nominal full loop flow. . Above 36% (P-8) of RATED THERMAL POWER, automatic reactor trip will occur if the flow in any single loop drops below 90% of nominal full loop flow. This S FARLEY-UNIT 2 B 2-5

       "       ^                                      - - - - _ _ _ .                                               _.

\-' INSERT F

 -   Overeewer AT The Overpower delta T reactor trip provides assurance of fuel integrity (e.g., no fuel pellet melting) under all possible overpower conditions, limits the required range for Overtemperature delta             TheT protection, cetpoint includes                   and provides a backup to the High Neutron Flux trip. corrections for axial                                                 I capacity of water with temperature, and dynamic compensation for transport, thermowell, and RTD response time delays from the core to RTD output indication. No credit was takau for operation of this trip in the accident analysts; however, its functicnal capability at the specified trip setting is required by this specification to enhance the overall reliability of the Reactor P,rotection system.

O e O m

                                                                          - - ~ - - - - - _ _ _ _ _ _ _ _

LIMIT!% SAFETY SYSTEM SETTIES . SASES latter trip will V: : st "- -'-'- '"l during normal e,1erational transients and anticipated transients when 2 loops are in operation and the Overtemperature delta T trip set point is adjusted to the value specified for all loops in operatien. With the Overtemperature delta T trip set point adjusted to the value seecified for 2 loco oceentien. the P 8 trie et M RATFn THERML POWER will'- " "- ---'-'"-! lM 'r a d ic td ow b:M:3duringnormaloperationaltransientsandanticipated transients with 2 loops in operation. ensee A.t ne DW5 de@ c.rMeu ifener _,,,,,,,,,. Steam Generator Water Level . The Steam Generator Water Level Low-Low trip provides core protection by preventing operation with the steas generator water level below the sintaus volume required for edequate heat removal capacity. The specified setpoint provides allowance that there will be sufficient water inventory in the stsaa generators at the time of trip to allow for starting delays of the auxiliary fasewater system. Steam /Feedwater Flow Misaatch and low Steas Generator Water Level O The Steam /Feedwater Flow Missatch in coincidence with a $ teas Generator . Low Water Level trip is not used in the transtant and accident analyses but is included in Table 2.?-1 to ensure the functional capability of the specified trip settings and thereby enhance the overall reliability of the Reactor Protection Systan. This trip is redundant to the Steam Generator Water Level

Loc Low trip. The stems /Feedvater Flu Misaatch portion of this trip is activated when tha steam flow axceeds the feedwater flow by greater than or equal to 1.55 x 10' lbs/ hour. The steam Generator Low Water level portion of the trip is activated when the water level drops below 25 percent, as indicated by the narrow range instrument. These trip values include sufficient allowance in excess of norsel operating values to preclude spurious trips but will initiata a rsaetor trip befors the steam generators art dry. Therefore, the required capacity and starting time esquirements of the auxiliary feedwater pumps are reduced and the resulting thermal transiant on the Reactor Coolant System and staas generators is minimized.

Underveltage and Underfrecuency - Resetor Coolant pusc lusses The Undervoltage and Underfrequency Rosetor Coolant Pimp bus trips provide, reactor core protection against DN8 as a result of loss af voltage er under-frequency to more than one reactor coolant pump. The specified set points assure a reactor trip signal is generated before the low flow trip set point 0. FARLEY-WIT 2 B 2-6 _ _ - _ - _ _ _ _ _ _ _ _ - - - - . - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - ' ~ ~ - ' " ~ ^ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~~

fgpg Men befCMP bth y ko 70 R.DCTIVITY CORTROL SYSTEMS TAtt % t. F*W E2.ud& 6.Un**e r*'af

4. D dida. KlE l '5" d 10 0./.

MODEMTOR TEMPEMTlTRE COEFTICIENT 7gggggg poggg,, , LIMITING CONDITION FOR OPERATION 3.1.1.3 The moderator temperature coefficient (MTC) shall bes

a. Less than or equal to h. x 10** delta k/k/'F for the all rods withdrevn, beginning et cycle life (BOL).. W h. 70: -" Rc.;.L 5$hNu N.

b. Less negative than 4.3 x 10** delta k/k/'T for the all rod l vithdrawn, and of cycle life (EOL), RATED THIRMAL POVER condition.

APPLICABILITT Specification 3.1.1.3.a MODES 1 and 2* onlyt Specification 3.1.1.3.b - MODES 1, 2 and 3 onlyl ACTION:

a. With the MTC acre positive than the limit of 3.1.1.3.c nbove, operation in .10 DES 1 and 2 any proceed provided:

1. Control rod withdraval limits are established and maintained sufficient to restore the MTC to within it6 limit within 24 hours or be in 807 STANDET vithin the next 6 hours. These withdrawal limits shall be in addition to the insertion limits of Specific'a tion 3.1.3.6.

2. The control rods are maintained within the withdrawal limits established above until a subsequent calculation verifies that the MTC has been restored to within its limit for the all rods withdrawn condition.

3. A Special Report is prepared and submitted to the Consission pursuant to Specification 6.9.2 within 10 days, describing the va2ue of the sensured KIC, the interin control rod withdrawal limits and the predicted average core burnup necessary for restoring the positive KIC to within its limit for the all rods withdrawn condition.

b. Vith the MTC acre negative than the limit of 3.1.1.3.b above, be in ROT SEUTDOVN vithin 12 hours.

FeTl"ih K ,, greater than or equal to 1.0 . 9 See Special Test Exception 3.10.3 O U FARLIT-WIT 2 3/4 1-4 Arendment No. H, 80

                      .         . em REACTIVITT CONTROL'5TSTIMS ROD DROF TIME LIMITING CONDTTION FOR OFEAATION
   ,   3.3.3.4 The individual full length (shutdown and control) red drop time from the fully withdrawn esition (225 to 231 steps, inclusive)* shall                                                   l be less than or equal to          seconds free beginning of decay of stationary 8

gripper cell voltage to dashpot entry with: .

s. T,,, greater than or equal to 541'F. and

b. All reactor coolant pumps operating.

AFFLICABILITT MODES 1 and 2. . ACTION:

a. With the drop time of any full length rod determined to exceed the above limit restore the rod drop time to within the above limit prior to proceeding to M001 1 or 2.

b. Vith the rod drop times within limits but Jetermined with 2 reactor coolant pumps operating, operation may proceed provided TatAMAL t0VIR is restricted to less than er equal to 66% of RATED TEDJtAL 70VIA.

StmVEILLANCE REQUIREMENTS 4.1.3.4 The rod drop time of full length rods she.11 be demonstrated through sensurenant prior to reactor criticality ,

s. For all rods following asch removal of the reactor vessel head.

b. For specifically affected individual rods following any maintenance on or modification to the control red drive systen veich could affect the drop time of those specific rods, and

c. At least once per 18 months. .

          'The hlly withdrawn position used for detersining rod drop time shall be graater than er equal to the fully withdrawn position used during subse-quant plant operation.                    ,

PAR 1XT-UNIT 2 3/4 1-19 AMENDMDff NO. 7& O e

                                                                                           ~         ^~   ~~

505'ER' 01S'TRIBUTION Lf51TS 3/L 2.2 HEAT FLUI HOT CHANNEL FA NOR - A m LIMITING CUNDITION FOR OPERATION

                ...- -              ------ - - - = - ...- -. -              ..   ... -      ..   . .......

7-~ 3.2.2 Fn (Z) shall be limited by the following relationships: , Fg (2) , [K(2)1 for P > 0.5 % j forbta{e5Ld Fn (Z) _ @ (K(Z)) for P < 0.5 Sor \lantage S be\ cce l where P . THERMAL POWER RAi w InERna PunER

                                and K(Z) is the function obtained from Figure (3.2 2) for a given core neignt location.

APPLICABILITY: MODE 1 ACTION: Witn Fg(Z) exceeding its limit: i l I

4. Reduce THERMAL POWER at least 15 ?:. each 15 Fo(Z) exceeds the limit

   .                      within 15 minutes and similarly reduce the Power Range Neutron                           l Flux.Hign Trip fetpoints witnin the next 4 hours; POWER OPERATION mav                    l procees for up to a total of,72 hours; subsequent POWER-0PERATION may                    l proceed proviced the Overpower delt4 T Trip Setpoints have been recuced                  !

f) at least 1% for each 1% Fo(Z) exceeds the limit. The Overpower delta T V Trip Setpoint reduction shall be perfomed with the reactor in at least HOT STANDBY. .

b. Identify and cor*ect the cause of the out of limit condition prior to increasing THERMAL POWER above the reduced limit required by a, abcve; THERMAL POWER may then be increasec provided Fg(Z) is demonstratec tnrough incere mapping to be witnin its limit Fqm5 C'f2 E kW3 be Pr 0 5 kr LOPAR bei Fq (,0 5 9.C [.kuo3 br Ps 0.5 kr LoPAR bel FARLEY-UNIT 2 3/4 2 4 AMENDMENT NO. 23 55

 -A b                                      -

o 6

IhlSE RT HERE

~.

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i ... .N i l KI .i .. I i\l i N i i N. ! . . :. N. 1. !\l i i INi I ,, K [y i\' i i \ !.

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            ! in I. NI                                      i 3I                i- i \i- .1:- O . . .m .:iw      ..

i ._\ . .i .y. # y. .iix 3 .. g .

                                                                                                                                          =iu i ..::.N.      ..! :          is .     .;          .tN        i  ..         i4,'                 .i    .Ni:     .  . i....p-i l . i- g': l l \i                                    ..i : l\ :l.:-            i.. ] \! ..~ j ;p .:Nl.:j:ih e                       e.                        m e

a o w e e o e- = e .

                                               =             1210:1 C121*lNYWWON-(ZD1 O                         .

FARLEY-UNIT 2 3/4 2-7 , l

1 0 _7,4SE,R, G

                                                                     '                       b en C

o - O' 5

m e

O w O 0 1

to 4 T l N 0 6 , , , o q e em W y q C '.

     .-                 o       e        o                 o
                                                                                                                          \

e l I e

80W!R D!!TR!!Ut10% t1MITS

     . 3/a.2.3 NUCLEAR ENTHALPV WOTCMaNNELFACTOR.FNH L!MITING CONDITION FOR OPERATION "         +

N 3.2.3 FAH small be limitee by the following relationship: N FA H < @ (1 + 0.3 (1-P)] A for Ud*0e 5 bd 'ad i 2 THERMAL P0WER j wnere P e RAILD InEAnAL Ps=ER ' F,",

1.55 r_l + c.3 (I-P)] hr LOPAR Al i APPLICABILITY: MODE 1 l ACTION: l l

N witn FAH exceecing its limit:

a. Recuce THERMAL POWER to less than 50% of RATED THERMAL POWER within 2 neurs and recute the Power Range heutron Flux High Trip Setpoints to <**

55% of RATED THERMAL POWEA witnin the next 4 hours. N . l

b. Demonstrete through in core mapping that FAH is within its limit within 24 hours af ter exceecir.g the limit or recute THERMAL POWER to less than 5% of RATED THERMAL P0WER witnin the next 2" Acurs, anc

c. Icentify anc correct the cause of the out of limit concition prior to increasing THERMAL POWER above the recucee limit required ey 4 or e, j aeove; suesequent POWER OPERAT10N may proceed provised snat Fa M is

   ,                 cemenstratec through in-core mapping to be within its limit at a
                                                                        ~
   ;                 nominal 50% of RATED THERMAL POWER apier to enceecing this THERMAL POWER, at a nominal 75% of RATED THERMAL POWER prior to exceecing tnis l                 THERMAL POWER anc within 24 hours after attaining 951 tr greater RATED THERMAL POWER.

1 . FARLEY. UNIT 2 3/4 2 8 AMENOMENT NO. 27. 5* e 8

i

POWER DISTRI9tf710N t.!MITS

.)

2 QNB PARAME*ERS. .. 1 t,IMITING CONDITION FOR OPERATION 4 i o 3.2.5 The following DNB related parameters shall be maintained within the limits shown on Table 3.2-1:

s. Recctor Ccolant Systee T,yg

b. Pressurizer Pressure

c. Reactor Coolant Systers Total Flow Rate APOLICABILITY: MODE 1. .

ACTION: With any of the above parameters exceeding its limit, costore the parametar to within its limit within 2 hours or reduce THERMAL POWER to less tr.an 5% cf i RATED THERMAL POWER within the next 4 heurs. aO)m.. w, SURVE!LLsNCE RECUIREMENTS 4.2.5.1 Each of the parameters of Table 3.2-1 shall be verified to be within tncir limits at least once per 12 hours. 4.2.5.2 The Reactor Coolant System total flow rate shall be determined to be

within its limits by measurement at least once per 18 months.

        ,       4 7.s.3 %. isdicM Ret Ctes rek A.at k vedes J. L JA j                      & accep W li d d W ca. et, 31 d 3 3 i
                                                                                                 .   ")

p- .

             "* FARLEY-UNIT 2                                    3/4 2-14

I. x L) : e 8 n E - I a 2 : : : = 1, l .e e 1

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                                                         ' TA81E 3.3-2 o'  -

REAC10R 1 RIP 5) STEM _IN51RUNENTATION RESPONSE TIMES-' ' RESPONSE TIME

  'E FUNC110NAL UNI _T 1                                                                                     Not Appilcable                                                                                       ;

o 1. - Manual Reactor Trip-

2. . Power Range, Neutron Flux < 0.5 seconds *

a. High flot Appiicable

b. Low 9

3. Power Range, Neutron Flux, Not Aprilicable o@pphedh liigh Positive Rate

4. Power Range, Neutron Flux, High Negative Rate \

                                                                                          <0ksechs\

Not Applicable w 5. Intermediate Range, Neutron Flux k Not Applicable

y 6. Source Range, Neutron Flux seconds *

7. Overtemperature ai r Not Appilcable

8. Overpower AT Pressurlier Pressure--Low 1 2.0 seconds 9.

                                                                                           < 2.0 seconds

10. Pressurizer Pressure--Nigh j

Not Applicable

11. Pres *surizer Water level--High

z Response time of the neutron flux signal portion ; Neutron detectors are exempt from response time testing.of the cliannel shall he me.asured ., l e I

I

O O O _TA8tE 3.3-4 (Continued) ENGINEERED SAFETY FEATURE ACTUATION SYSTEM INSTRUNENTATION TRIP SETP0fMTS -

   ?
  $  FUNCTIONAL UNIT TRIP SETPOINT                 ALLOWRRLE VAltES m  2. CONTAINNENT SPRAY

a. :

Manual-Inttlation Not Applicable Not Appifcable 4

b. Automatic Actuation logic Not App 1Icable Not Applicable i

c. Containment Pressure--High-High-High 127 psig 1 psig

3. CONTAINNENT ISOLATION

a. Phase "A" Isolation l $ 1. Manual Not Applicable Not Applicable Z 2. FromSafetyIn3ection Not Applicable Not Applicable m

, Automatic Actuation logic

b. Phase "B" Isolation i

1.- Manual Not Applicable i Not Applicable

2. Automatic Actuation logic Not Applicable Not Appilcable i

3. Containment Pressure-- < 27 psig

High-High-High ~ ~

                                                                                     <         psig
                                                                                                                                                                          ^

c. Purge and Exhaust Isolation -

U*

1. Manual Not Appilcable Not Applicable ,

2. * ^ ':;*-1 llJ#

NW 4 T T ok pyic.U '

                                                                                         *k S PP          li caw)

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m k i . m FA t 1

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=slp-t-a.,e 5 N- N- = g 12 3 = ~r : I a. - s ls v? h 2 ==

g A A .iv4g:-],s.5 5 2 l $ m:t i ' v 48.

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                                        .M o             mm..n ,                                                                   uo.,

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     +                  - . _ _ _ _ _ _ _ _ _ _ _ _ _ _

L 'c O I i . TaSt2 3.3-4 (Cantfema4) i 7 se EIICimene SAF: TTY FEATURE ACTWATISI SYSTWI INSTWWI5ifAfteN TRIP 5::TvetWFS

    . .N     FWeCTiestAL Wr!T                                  TRIP SETPetWT                     Atassasta 74 tats e

i 3 e 6. 4WEILIART m !

     "                                                                                                                                \

a. Aetematic Actantles lagle N.I. W.A.

i , u b. Steam Generator .. Water taveletee-lee. > 171 td astree raags > 162 et sacrow range 1 l instr = ces spee each Testrement op.e e.ch l i' steam generster, ,-. steam generater l c. Onderveltage - act 1 2600 weltb &2640welta 9 I l 4.. S.I. .See I % (all St Setpelats)

u e. Trip of Nele Feedveter Pumps W .A.

                                                                    ,                                     N.A.

2. less .F Fi==

ll a. 4.36 hv Raeegency See moderveltage 1 3255 velts bec,.astange* > 3222 velea hos weleage e l (tees et vettage) - 33418settsbesvoltage* 5

b. 4.36 hv EM.gency See Enderve!! age 1 M75 velts bus weltage* > M 38 velts hos weltage e (Segeaded VAeage) j 33749veltsbusweltage* -

e. sus:1pmenes e Psavend Actanfred srsras seramiscxS . j

  ,  ,          s. Pressoriser Pressore, F-Il               3 200ggg                         $ 20t6 p,eg e                                             .                                      4 a-

b. Ime-Ime T,,,, F-12 (tecreasirs) 544*F -' < 545*F ,

a . (Beereaslag) 543 F

F 544 F h c. Steam Cenerator 149el F-14 54 !

(3ee 5. ebeve) t

~~

d. Reacter TrlF . F-4 N.A. 88. A.

Refer to appropelate relay setting sheet for calibration regseleceents I t

l l

 ,                                 TABLE 3.3-5 (Continued)                              y ENGINEERED SAFETY FEATURES RESPONSE TIMES 1

INITIATING SIGNAL AND FUNCTION RESPONSE 11HF IN SECONDS

3. Pressurizer Pressure-Low

a. SafetyInjection(ECCS) 127.0(I)/12.0(4)

b. Reactor Trip (from SI) < 2.0

c. Feedwater Isolation 5 32.0(6) I

d. Containment Isniation-Phase "A" $ 17.0(4)

e. Containment Purge Isolation 1 5.0 I

f. Auxiliary Feedwater Pumps Not Applicable

g. Service Water System 5 77.054)/87.0(l) c

4. Differential Pressure Between Steam Lines-High

a. SafetyInjection(ECCS) 512.0(4)/22.0(5)

b. Reactor Trip (from SI) < 2.0

c. Feedwater Isolation 1 32.0(6)

O d. Containment Isolation-Phase "A" 517.0(4)/27.0(5) U e. Containment Purge Isolation Not Applicable

f. Auxiliary Feedwater Pumps Not Applicable

g. Service Water System 1 77.0(4)/87.0(5)

5. Steam Flow in Two Steam Lines - High Coincident with f hp(;e,s.

T avg "'l0*

a. Steam Line Isolation h 9.$

6. Steam Line Pressure-Low

a. S#f'4tyInjection(ECCS) 1 12.0(4)/22.0(5)

b. Reactor 1 rip (from SI) $ 2.0

c. Feedwater Isolation < 32.0(6)

d. Containment Isolation-Phase "A" 117.0(4)/27.0(0)

e. Containment Purge isolation Not Applicable

f. Auxiliary Feedwater Pumps Not Applicabla

g. Service Water System 5 77.0(4)/87.0(5)

h. Steam Line Isolation _< 7. 0 A V -

FARLEY-UNIT 2 3/4 3-30

REACTOR COOLANT SYSTEM e-HOT'5TAND!Y lNSEET H LIMITING CONDITION FOR OPERATION O 1.2 All Reactor lant Loop ' tee below 11 be DP r anc in ope:- when the control is oper nel or at t qwo Rea . colant Lo isted be hall be LE with actor Co ' Loop in op on wnen a contro tem is e d by ope the React ip

                               . Breakers             hutting          the rod              motor     /ge           'er    setst

I

1.

Reac'ter Coolant Locp A and its associated steam generator and i Reactor Coolant pump, .

2. Reactor Coolant Loop I and its associated steam generator and Reactor Coolant pug.

3. Reactor Coolant Loop C and its associated steam generator and --

Reactor Coolant pump. APPLICABILITY: MODE 3 ACTION: _ , s. With less than the above reuired Reactor Coolant loops OPERABLE, [ 7 restore the recuired loops to OPERABLE status within 72 hours or be in fjNS6kT NOT 5HUTDOWN within the next. 12 Pours,

b. W1 ess tna h qee Rea clan ops in estion the roc N contro tem ope. ional, w 1 how tor /ge ,

en the ctor k ers_ or t down fed dri ter se.

c. With no Reactor Coo' ant loops in operation, suspend all operations-involving a reduction in boron concentration of the Reactor Coolant System and ispeciately initiate corrective action to return the required coolant loop to operation. '

                                                                                  ^
                                 $URVE!L .ct REQUIREMENTS 4.4.1.2.1          At least the above required Reactor Coolant pumps, if not in operation, shall be determined to be OPERA 8LE once per 7 days by verifying                                                                 ,

correct breaker alignments and indicated power availability. , 4.4.1.2.2 The required Reactor Coolant loop (s) shall be verified to be in i operation and circulating Reactor Coolant at least once per 12 nours. 4.4.1.2.3 - The required steam generator (s) shall be determined 0PERABLE by verifying secondary side water level to be greater than er equal to 101 of wide. range inoication at least once per 12 hours. -

                                 'All Reactor coolant pumps may be de-energized for up to I hour provided (1) no operations are permitted that would cause dilution of the Reactor Coolant System boren concentration, and (2) core outlet tageraturt is maintained at least 10'F below saturation tagerature.

FARLEY-UNIT 2 3/4 4-2 AMINDMENT NO. 58 O

                                                                                                                                                             .m-..r.y.m.,

INSERT H 3.4.1.2 At least two of the Reactor Coolant Loops 14td klow shall be OPERABLE and in operation when the rod control system is operational or at least two Reactor Coolant Loops listed below shall be OPIAABLE with one Reactor Coolant Loop in operatien when the rod control system is disabled by opening the Reactor Trip Breakers or shutting down the rod drive motor / generator sets

i l

l l I 1 INSERT I i

b. With only one Reactor Coolant Loop in operation and the rod control system operational, within I hour open the Reactor Trip Breakers or shut down the rod drive motor / generator sets. 1 O -

U 4 0

3,4,2 Stutt t!s m su 10N (191 5 ) mesy 5 DWB desp c.ederion 3A5E$ The specifications of this section provide assurance of fuel integrity during Concitit- 1 (Nomai Deeration) and 11 (Incioents of Moderate Fre:veaevi events by: (a;je*w----t:: - m r ^ ^ p: =: ;^:::: : : r m : I b; :..t.iouring .armal operation anc in snort term transtants, and Jo) limiting tne fission gas release, fuel pellet temperature and claddini mechanical properties to witnin answned design criteria. In addition, Limiting the peat ! Ifnear power density during Condition ! events provides assurance that the ' initial conditions assened for the LOCA analyses are met and the ICCS acceptares criteria limit of 2200'F is not exceeded. The definitions of certain het channel anc peaking factors as usec in snese specifications are as follows: Fg(I) Heat Flux Hot Channel Factor, is agfined as the maximum local neat flux en the surface of a fuel rod at cc,re elevation Z divided Dy the average l fuel rod nest flux, allowing for manufacturing tolerances on fuel cellets and rods and measurement uncertainty. Ffg Nuclear Enthalpy Ris6 Mot Channel Factor, is defined as the ratio of the integral of linear power along tne rod with the hignest integrated power to the average roe power. F,y(Z) Radial Fdaking Factcr is defined as the ratio of peak power density to average power sensity in the norizontal plane at core elevation 2. 3/a.2.1 AXfiL FLU 2 DIFFERENCE The limits on A1!AL FLUI O!FFERENCE (AFD) assure that the ((Z) upper bound envelope of Qtimes the normalized axial peaking factor is not exceesu curing eitner normal operation or in the event of renon todistribution fo11owing power changes. , , i Target flux difference is determined at equilibrium menon conditions. The full length rocs may be positioned within the core in accordance witn their respective insertion ilmits and should be inserted near their normal position for steady state operation at high power levels. The value of the target flux cifference obtained under these conditions divided by the fr.iction of AATED THERMAL POWER is the target flux difference at RATED THERMAL POWER for the associatec core burnup conditions. Target flux differences for other THERMAL POWER levels are obtained by multiplying the RATED THERMAL PDWER value by the appropriate fractional THERMAL POWER level. The periodic updating of the target flux difference value is necessary to reflect core burnup cons 1Gerations. 2J5forYant od 2.32 Ar y 5LOPAR ,- O FARLEY-UNIT 2 83/4 2-1 AMENDMENT ND. J3 65

                                 - - -           _ ----,.+-.- ,                                             ,      a

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

S M * !!!":!r !:9 .!" "I .. U!E! Q V n:st ewx 3::rEREsCE f:: +d u.e - 1

Altneu within n e gn(l)% it istar;etintances band tnat snethe cout plant will flux target ce eparatan difference wi duringu tne AF: ra ic plant THERMXI. PCVE.R reductions, c:ntrol rce sotien will cause,the AFD := caviata :uts'te of the target banc at reduced "ht?M L 70hEA levels. ~ 5's tev'at an will not affec* ue xenen regis rituti:n sufficient *y ta c .arge d tse envel:ce of seat'ng fact:rs nica may se reae..e: an a suosecuent re u 9 :: RATED ThEVAL PCWER (with the AFD witnin the ta.get tanc) provices the time curation of the deviation is liniud. Ac:ortingly, a 1 hour penalty caviati:n , limit cumulative during the previous la hours'is provided for operation cussite J of us target bane but within the Itaits of Figure (3.2-1) while at THEVAL PCWEA levels between 5C% and 90% of AATED THERMAL POWER. Fcr THERMAL PCWER levels between 15% and 50% of AATED THERMAL PowfA, caviations of the AFD outsica of the. target band are less significant. The penalty e.f 2 hours

ectual time reflecu this recuced sdgnificance.

Provisions for monitaring the AFD on an automatic basis are co-ived ft:m the plant pr: cess c:meuter. thr:ugn :ne AFD Menitor Alarm.- The c:meuter cate-sines the one minuta average of each of the CPERA8LE exc:re detect:r cut:uts anc provides an alars message innesistaly if the AFD for 2 or more CPERAILE ex::re channels : e cuisice tne tar;st band and u s 'MERMAL PC'.ER is teater than IC% of RATED THEPNL POWER. During c:trati n af THEVAL PC'*ER 1 O lev;els between SCE anc 904 and totneen 1!: ano SCE RATED TMEPAL PCWER, t..e ; c:t: uter outsuts an alars message wnen tha penalty caviation scrumulates teyonc , tne lisiu of 1 hour and 2 hours, res;,ectively. Figure 3 3/4 2-1 snows a typical acnthly target band,. . 3/2.2.2 anc 3/4.2.3 W D* F'.UX HOT CFANNEL FACTOR. NUCLD R INTHAlpY WCT TW% M @ t The limits on heat flux hot channel ifactor and nuclear enthalpy rise not channel fee s- easure that 1) the-design / limit $,on peak local power density

              ,;.;; r V ;.: :C ;d not exceeded, andG in ths event of a LOCA the peak fuel clac temmerature wi' 1 not axceedfthe 2200'F ICC3 acceptance critaria limit.

(-ser iccically as specified in specifications 4.2.2 and 4.2.3.EscaThis of these peri:cic is seasuramle surveillance is sufficient to insure that the limits are saintained provi:ec: a.- Control rods in a single g ous sove t:gether with nc individual r:c . 3)tiw 1 insartion diffe-ing by more than : 2.2 ste s, inticated, fr:e tr.e - aus grouc casitad pcsition. 4%a ,

b. Control r d tanks are sectennt with overlacping groups as :es:rd:e:

r.n W ee in Spe:ificatien 3.1.3.6. ,

           ' m d'     e. The control red insertien limits of Specifications 3.1.3."5.anc 3.1 3.5 are saintained.

Ok. d. The axial power distribution, expressed in terms cf AX*AL FL'JX Vg OIFFERENCE, is saintained within tne limits. FARLff-UNIT 2 $ 3/4 2-2 ,,

                 %~                                ,                 ,- ,,           -
                                                                                           , , , , - - . , . _ , . . . _ _                               ..,n_   , . ,+,-.-n     -

POWie 0!$Ta190t10w tju!T5 *-

                                            - 3A$t$

O " " " " " " " ' " " " " ~ ' " ' " " " ' " " " " " " " " " " ~ ~ ~ " " " " " ' ' - F wili 6e maintained within its limits provide conditions a, throwg d. stove are maintained. The relaintion of F as a function of THERMAL p0WER allows chantes in the radial pwer snap or all Armisstole rod insertion limits. When an To measurenant is taken, an allowance for Doth esperimental error and manufacturing tolerance suit be made. An a11swance of it is appropriate for a full core map taken with the intore detector flus mapping system and a 31 allowance is appropriate for annufacturing tolerance.

                                                         - When F[M is measured, esperisontal error must be allemed for and at is the appropriate allesance for a full core na                          en with the intore detection system. The sentified limit for F u                              contains am at allowance fe une .en w 4.s. h .e                      .,. ,,,.;.... ,; .J ,.;; a .11: 7;;.M n l
                                              'O;;.;g.; .                       The 51 allowance is cases on the rollowing consiserations:

4. Abnormal perturbations in the radial poner shape, such as from roe misalignment, af fact Ffg mort directly than F g,

b. Although red movement has a direct influence upon limiting Fo to within its limit, such control is not readily available to limit FjH and

c. Errors in prediction for control power shape detected during startup anysics tests can be cortensated for ij Fo by restricting asial flus sist ribution. This cargensation for Fwis less readily availaole.

Fue) d bowing h ces the va of DNS ret . Credit availabl e o et this action in generic ma n. The g ric desig. rgins, tota ng 3.1% 0 R. cowlete of fset any a bow pena ies (less n 25 for tn- rst case nica occurs t a burnup 33,000 U). Tnis rgin neludes felimi . Design ftDMR 1.30 vs.1. t tal Gr1 acing Ce ficient(ks of 0.046 0.05l 3 in 1 Dif ion Coeff ent of 0.0 vs . 0.05 g a DNB A I tiplier C.865 vs 0.88 pitch r etion

             .                                                                                                                                               l

~..' FALLtY. UNIT ! 8 3/4 f.4 AMENDMtNT h0. 57 o

                 ,A-s

 --- - - ---                                                          - -- ~.. - - -.- -.. - - - - - - -
                                                                                                            -:     ----   ~ ~ -            .            ..

1 00vtt cl$tRTIV'!0N Llw1Ts SA$t5 , The racial peaking factor Fe (2), is measured periodically to previse ascitional assurance that the not thann 1 tor Z remains within its limit. The F limitforRATEDTHERMALF0WER( P ),a Op(ro)v,ised in the Aa8141 peaking Fact limit report per Specification 8. 1.11 was determined from espectes power sentrol maneuvers over the full range of burnup tenditions in the core. o 3 /a .f .a CUCRANT 90 wit TILY EAT 10 The quadrant' Dower tilt ratie limit assures that th6 radial power distribution satisfies the design values used in the power capability analysis. Radial power sistribution measurements are nace during startup testing and periodically during power operation.

1 The limit:of 1.02, at wnici corrective action is required, provides 048 and l linear heat generation rate protection with a.y plane power tilts. I l

The two hour time allowa9ce for operation with a titt condition greater than i 1.02 Dut less tnan 1.0g is provioes to allow identification anc correction of a ' eroppee or miss11gned control red. In the ever.t.such action does not. correct the tilt, the margin for uncertainty on F i ~

          .                     anowee power .y 3 ,ercent for each ,ercen%                                   .s reinstates f siis  in . nesi by   resucin8.tne maximu
                                                                                                                             .f 1.                                !

For purposes of monitoring QUADRANT 70Wtt TILT RATIC when one encore detector is a inoperaole, the movable incore setectors are used to confirm that the normaliaec

                            . symmetric power distribution is consistent witn the 00ADRANT p0WER TILT RATIO.
                        '       The incore detector monitoring is sone with a-full incere flua map or two sets of f our symmetric thimbles. "ne two sets of four sysunstric thimbles is a unique set of eigat estector locations. These locations are C.8, E.5, t.11, N.3. H.13                                               n
                              -L.5, L.11, b 8.

3/a.2.6 DNS PARAMETER $ N5fAr 3 NERE The limits en the Dh8 related parameters assure that each of the parameters are maintained witnin the normal stead transient and accident analyses. yThe state limitsenvelope of operation are consistent withassees in the the initial 1 FSAR asseottoas ama have teen analyticall g= = :.= tnrou.hout .acn ..i,y demonstrated sed transient. adequate

                                                                                                                        ,              tol;.r W ; p              ;

mest The 12 h r periodic evaillance these par rs throu instrument aeout g sufficie to ensure at the par ters are re red within ett limit ; M. fo wing los antes and her expect transient op tien. Th I month perio measur of the t otal flow- te is adeau to detect ow D Cr N #6 egregat and ensu correlatio f the fle 81 cation a els with sueed such t the indi ed percent on will p ide suffiste vertftcat . of rate e 12 tour e is.

O - FARLEY. UNIT t' S 3/4 2 5 pgRg

AMICMEkT W. 52

                                                                                                                                              .-,,n,.,..,~-. ,-%

INSERT J The indicated Tavg value of 580.70F is based on the average of two control board readings and an inditation uncertainty of 2.50f. The indicated pressure value of 2205 psig is based on the average of two control board readings and an indication uncertainty of 20 psi. The indicated total RCS flow rate is based on one elbow tap measurement from each loop and an uncertainty of 2.4% flow (0.1% flow is included for feedwater venturi fouling). INSERT K The 12 hour surveillance of Tavg and pressurizer >ressure through the control board readings are sufficient to ensure t1at the parameters are restored within their limits following load changes and other expected transient operation. The 18 month surveillance of the total RCS flow rate is a precision measurement that verifies the RCS flow requirement at the beginning of each fuel cycle and ensures correlation of the flow indication channels with the measured loop flows. The monthly surveillance of the total RCS flow rate is a reverification of the RCS flow requirement using loop elbow tap pj measurements that are correlated to the precision RCS flow measurement at q the beginning of the fuel cycle. The 12 hour RCS flow surveillance is a qualitative verification of significant flow degradation using the control board indicators and the loop elbow tap measurements that are correlated to the precision RCS flow measurement at the beginning of eacn fuel cycle. im

meet the DNE 3/4.a REAtTOR COOLANT Sv5 TEM 7 dt&*ian C.rNerien BASES j

3/4.4.1 REACTOR COOLANT LOOPS AND COOLANT CIRCULATION

The plant is desiemed to eeerstefwith all Reactor Coolant loops in operation, endic;;a w . ; C e m t.;;iduring all normal operstlens and anticipated transients. In EDL51 ano 2 with one Reactor Coolant loco not in operation this specification requins that the plant be in at least10T ITANDBY within 1 hour. I In PODE 3,hehtl reactor coolant loops' provide sufficknt heat mmovat capability for removing core heat even in the event of a bank withdrawal accident; however, a s< ngle mactor coolant loop provides sufficient decay heat removal capacity if a bank withdrawal accident can be prevented; f.e., by opening the Reactor Trip treekers or shutting down the red drive actor / generator sets.1 =: ; ::-t e n: ;.;l ;;;;;;n; ;;; c; ; m ;nt;;. ;;n;;; 7;;;. ; 1:; :';: :t': : r;;. ' r: p;t t : ';;;; i: 0^tV L! ;; ;" Of :::.

In MODE 4. a single nactor ' coolant or RHR loop provides sufficient heat removal capability for mmoving ocay heat; but single failure considerations

                                                                                                                -       i nouire that at least two loops be OPERABLE. Thus, if the reactor coolant loops                ,

are not OPERAILE, this specification requires two R$ loops to be OPERA 8LE. In MODE 5. single failure considerations reelire two RM loops to be OPERABLE. The operation of one Reactor Coolant Pump or one RNR pump provides adecuate flow to ensure mixing, creyent stratification and produce gradual reactivity O changes during boren concentration retketions in the Reactor Coolant System. The reactivity change rate associated with boron reduction will, therefore, be within the capability of operator recognition and control. , The restrictions on starting a Reactor Coolant Pye with one or more Reactor Coolant System cold legs less than or egual to 310*F are provided to prevent Reactor Coolant System pressurs transients, caused by energy additions frce the secondary systes, which could exceed the limits of Aphendia G to 10CFR Part 50. The Reactor Coolant Systes.will be protected against overpressure transients and will not exceed the limits of Appendix 4 by either (1) restricting the water volur in the pressurizer and thereby providing a volpe for the primary coolant to expand into, or (2) by restricting starting of the Reactor Coelaat Pumps to when the secondary water temperataire of each steam generator is less than 50*F above each of the Reactor Coolant System cold leg temperatures. 9 O

FARLEY. UNIT 2 5 3/4 4 1 AEND(NT NO.58

l i 1 ADMINISTRATIVE C0tfTROLS

e. Type of containar (e.g., LSA, Type A Type 3. Latge Ouantity), and

f. Solidification agent (e.g., cement urea formaldehyde).

The radioactive effluent release reports shall include unplanned' releases from the site to unrestricted areas of radioactive materials in gaseous and liquid effluents on a quarterly basis. The radioactive effluent release reports shall include any changes to the PROCESS C0tfrROL PROGRAM (PCF) made during the reporting period. M0trTHLY OPERATING RIPORT 6.9.1.10 Routine reports of operating statistics and shutdovn experience. including documentation of all challenges to the FORV's or safety valves, shall be submitted on a monthly basis to the Comeission pursuant to 10CTRSO.4 l no later than the 15th of each me- h following the calendar month covered by the report. Any changes to the 0FFSITE DOSE CALCULATION MANUAL shall be submitted with the Monthly operating Report within 90 days in which the chLnge(s) was made effective. In addition, a report of any major ahanges to the radioactive vaste treatment systees shall be submitted with the Monthly Operating Report for the period in which the change was implemented. RADIAL PEAKING FACTOR LIMIT REPORT 6.9.1.11 The F lielt for Rated Thermal Power (rRTP shall be provided to the Commission. purs6 ant to 10CFR50.4, for all_ core plaEes ontaining bank "D" l control rods ano all unrodded core planesIn I-- i " ' = er:e; y cycle initial

  /, D                  criticality. In the evrnt that the limit would be subsitted et some other time V                     during cere life, it vill be submitted; q..

become effective unless otherwise exempted by- dTConsission.

                                                                                             . . , ,,   the date the limit vould Anyinformationneededtosupportp# P villbekbyrequestfromtheNRCandneed not be included in this report.

fm t> clo~y A ANNUAL DIESEL CENERATOR RELI ABILITY DATA REPORT 6.9 1.12 The number of tests (valid or invalid) and the number of failures to start on demand for each diesel generator shall be submitted to the NRC annually. This report saall contain the information identified in Regulatory Position C.3.b of NRC Regulatory Guide 1.108. Revision 1. 1977. FARL5Y. UNIT 2 6 19 AMENDMEttr NO. H,82,74 O 9 a , -- , .- -- -, ,.e.-- .- . ~ - - - , , , -e

-( J. H. FARLEY NUCLEAR PLANT UNIT 2 TECHNICAL SPECITICATIONS Remove Pages Insert Pages 2-2 2-2 2-5 2-5 28 2-8 2-9 2-9 2 10 2-10 B 2-1 B 2-1 B 2-2 B 2-2 B 2-3 B 2-3 B 2-4 B 2-4 B 2-5 B 2-5 B 2-6 B 2-6 3/4 1-4 3/4 1-4 3/4 1-19 3/4 1-19 3/4 2-4 3/4 2-4 3/4 2-7 3/4 2-7 3/4 2-8 3/4 2-8 3/4 2-14 3/4 2-14 3/4 2-15 3/4 2-15 3/4 3-10 3/4 3-10 (r'3 3/4 3-26 3/4 3-26 L,_) 3/4 3-27 3/4 3-27 3/4 3-28 3/4 3-28 3/4 3 3/4 3-30 3/4 4-2 3/4 4-2 B 3/4 2-1 B 3/4 2-1 B 3/4 2-2 B 3/4 2-2 B 3/4 2-4 B 3/4 2-4 B 3/4 2-5 B 3/4 2-5 B 3/4 4-1 B 3/4 4-1 6-19 6-19 tr% j

O 670< UNACCEPTABLE 660 OPERATION 2440 psia 650-640, 2250 psia 630' 2000 pstr { p 620'-

     "                                1875 psla 610-         1840 psia O

600< 590< ACCEPTABLE OPERATION

580-570 . -

0. .1 .2 .3 .4 .5 .S .7 .8 .9 1. 1.1 1.2

,                              POWER (FRACTION OF RATED THERMAL POWER)

Figure 2.1-1 Reactor Core Safety Limits Three Loops in Operatiori O FARLEY - UNIT 2 2-2 AMENDMENT NO.

4

            ,                                                                  TABLE 2.2-1 D

e REACTOR TRIP STSTEM INSTRUMEICATION TRIP SETPOINTS y 8 FUNCTIONAL UNIT. TRIP SETPOIPrr ALLOVAELE VALUES

1. Manual Reactor Trip. Not Applicable Not Applicable

2. Power Range, Neutron Flux Lov Setpoint - f 25% of RATED Low Setpoint - f 26% of RATED THERMAL POVER THERMAL POVER I. High Setpoint - f 109% of RATED- ' High Setpoint - $ 110% of RATED

! THERMAL POVER THERMAL POWER i 3. Power Range, Neutron Flux, f 5% of RATED THERMAL POWER with 's 5.5% of RATED THERAAL POVER { High Positive Rate 'a time constant 1 2 seconds with a time constant 2 2 seconds

4. Power Ran2e, Neutron Flux, f 5% of RATED THERMAL POWER with f 5.5% of RATED THERMAL POWER
High Negative Rate a' time constant 1 2. seconds with a time constant 1 2 seconds i

h y 5. Intermediate Range, Neutron f 25% of RATED THERMAL POWER S 30% of RATED THERMAL POVER v' Flux 5

6. Source Range, Neutron Flux f 10' counts per second f 1.3 I 10 counts per second 7, Overtemperature & See Note 1 See Note 3; 8. Overpower & See Note 2 See Note 6 l; 9. Pressurizer Pressure-Lov 2 1865 psig 2 1855 psig; 10. Pressurizer Pressure-High .f 2385 psig 5 2395 psig g 11. Pressurizer Vater S 92% of instrument span f 93% of instrument span
y Level-High tr

             $        12. Loss of Flov-                       > 90% of minimum seasured flow             > 88.5% of minimum measured flov l E                                                 hrloop*                                    hrloop*

! 8

Minimum measured flow is 89,290 gpa per loop. l

_ . , , - . . . . . - . .,,-c- ,

T h

                                                   . REACTOR TRIP STSTEM INSTRUMENTATION TRIP SETFOINTS NOTATION h

5;

        < Note 1: - Overtemperature E.                                                                                                  >

e j g M (l' + T s) $ E, (K - K 2(1 +T s) . (T ( e g i 1 ) - T') + K 3(P - P') - fg (81)] Q (1 + T3 s) (1 +T,s) ~ 1 + T, s ! w where: M = Measured M by RTD instrumentation; aT, = Indicated M at RATED THERMAL POWER; P T = Average. temperature, 'F; T' f 577.2*F (Maximum Reference T,,, at RATED THERMAL POWER); } P = Pressurizer pressure, psig; [

P' = 2235 psig (Nominal RCS operating pressure); ;: ' i u 1+Ts 3 - The function generated by the. lead-lag controller for dynamic compensation; T,,,

l 1+T8 2

T 1 '& T,- - Time constants utilized in the lead-lag controller for (,,, Tg - 30 sec, T, = 4 see; I

T* * = The function generated by the lead-lag contro11e. for a dynamic compensation;

1 + .T3s ,

I T, &T 3

                                     - Time coastants utilized in the lead-lag cuntrolht for E , T,         = g - O sec; i                                                                                                                                       ,

1

                                     = Lag compensator on measured T,,,;                                                               }

1 +T68 f 1

       @                       T,    = Time constant utilized in the measurels T,,, lag compensator, T,       = 0 see;                 i m
       =                                                                  _t

,' g s - Laplace transform operator, sec ; m l 5 Operation with 3 loopa operation with 2 loops z ;

       ?            Kg =-   1.14;                                                                   K, = (values blank pending      l K, = 0.0230;                                                                    K, = NRC approval of            l i

i- K3 = 'O.001275; K, = 2 loop operation) ] l

REACTOR TRIP SYSTEM INSTRUMENTATION TRIP SETPOIhTS k NOTATION (Continued) .G< and f t (aI) is a function of the indicated difference between top and bottom detecters of the power-range g nuclear ion chambers; with gains to be selected based on measured instrument resoonse during plant startup s tests such that: a (1) for q -q between -39 percent and +13 percent, f (AI) = 0 (where q, and q l THERM 1L PODEh in the top and bottom halves of the c, ore respectively, and q, , are percent RATED POVER in percent of RATED THERMAL POV3R); (ii) for automatically each percent that the magnitude of (q,f its

                                                                        - q,) exceeds -39 percent, the aT trip setpoint shall value at RATED THERMAL POVER; and be                   reduced  by 1.92 percent    o (iii) for each percent that the magnitude t'          (q, - q,) exceeds +13 percent, the & trip setpoint shall be automatically reduced by 2.17 peretnt of its value at RATED THERMAL POVER.

Note 2: Overpover aT AT (1 + T,s) $ AT, [K,- K3 ( Ts 3 ) ( 1 ) T - K, (T ( 1 ) - T") - f,(aI)] ,'o (1 + T3s) 1 +T s 3 1 + r, s 1 + T, s where: aT = Measured at by RTD instrumentation; AT,= Indicated M at RATED THERFAL POVER; T = Average temperature, 'F; T" - Reference T at RATED THERMAL POWER (Calibration temperature for aT instrumentation, f 577.2*F); K, = 1.07; ] g K 3 = 0.02/*F for increasing average temperatute and 0 for decreasing average temperature; O g K, = 0.00165/'F for T > T", K, = 0 for T f T"; } d s g 3 - The function generated by the rate lag controller for T,,, dynamic compensation; 1+T3 s

                     ,- \

O O TABLE 2.2-1 (Continued) REACTOR TRIP SYSTEM INSTRUMENTATION TRIP SETPOINTS ,

n. ..

                  ,                                                      NOTATION (Continued)

E O w T 3 = Time constant utilized in the rate lag controller for T,,,, T3 = 10 sec; I + T* s - The function generated by the lead-lag controller for M dynamic compensation; 1+Ts 3 T, &T3- Time constants utilized in the lead-lag controller for E , T, = T3 - O sec; 1

                                             = Lag compensator on measured T,,,;

1 + T, s T, = Time constant utilized in the measured T,,, lag compensator, T, = 0 see;

~*

s - Laplace transform operator, sec ; u L o f2( AI) = 0 for all M. Note 3: The channel's maximum trip point shall not exceed its computed trip point by more than 1.8 percent. l Note 4: Pressure value to be determined during initial startup testing. Pressure value of f 55 psia to be used prior to determination of revised value. Note 5: Pressure value to be determined during initial startup testing. Note 6: The channel's maximum trip point shall not exceed its computed trip point by more than 2.3 percent. l E e n 5 .

2.1 SAFETY LIMITS BASES 2.1.1 REACTOR CORE The restrictions of this Safety Limit prevent overheating of the fuel and possible cladding perforation which vould result in the release of fission products to the reactor coolant. Overhea*.ing of the fuel cladding is prevented by restricting fuel operation to within the nucleate boiling regime where the heat transfer coefficient is ltrge and the cladding surface temperature is slightly above the coolant saturation temperature. Operation above the upper boundary of the nucleate boiling regime could result in excessive cladding temperatures because of the onset of departure from nucleate boiling (DNB) and the resultant sharp reduction in heat transfer coefficient. DNB is not a directly measurable parameter during operation and therefore THERHAL POVER and Reactor Coolant Temperature and Pressure have been related to DNB through correlations which have been developed to predict the l DNB flux and the location of DNB for axially uniform and non-uniform heat flux distributions. The local DNB heat flux ratio, DNBR, defined as the ratio of the heat flux that vould cause DNB at a particular core location to the local heat flux, is indicative of the margin to DNB. The DNB thermal design criterion is that the probability of DNB not occurring on the most limiting rod is at least 95 percent (at a 95 percent confidence level) for any Condition I or II event. In meeting the DNB design criterion, uncertainties in plant operating parameters, nuclear and thermal parameters, fuel fabrication parameters and computer codes must be considered. As described in the FSAR, the ef fects of O these uncertainties have been statistically combined with the correlation uncertainty. design criterion. Design limit DNBR values have been determined that satisfy the DNB Additional DNBR margin is maintained by performing the safety analyses to a higher DNBR limit. This margin between the design and safety analysis Ifmit DNBR values is used to of fset known DNBR penalties (e.g., rod bov and transition core) and to provide DNBR margin for operating and design flexibility. The curves of Figures 2.1-1 and 2.1-2 show the reactor core safety limits for a range of THERMAL POVER, Reactor Coolant System pressure and average temperature which satisfy the following criteria:

a. The average enthalpy at the vessel exit is less than the enthalpy of saturated liquid (far left line segment in each curve).

b. The minimum DNBR satisfies the DNB design criterion (all the other line segments in each curve). Each curve reflects the most limiting result using either lov-parasitic (LOPAR) fuel or VANTAGE 5 fuel. The VANTAGE 5 fuel is analyzed using the VRB-2 correlation with design limit DNBR values of 1.24 and 1.23 for the typical and thimble cells, respectively. The LOPAR fuel is analyzed using the VRB-1 correlation with design limit-DNBR values of 1.25 and 1.24 for the typical and thimble cells, respectively,

c. The hot channel exit quality is not greater than the upper limit of the quality range (including the effect of uncertainties) of the DNB O correlations This is not a limiting criterion for this plant.

B 2-1 FARLEY - UNIT 2 AMENDMENT NO.

_ ~ _ . . - .__ .- - -__ - - - . _ . - . _ . _ _ - _ _ _ _ _ _ _ _ _ -- SATETY LIMITS BASES f The curves of Figures 2.1-1 and 2.1-2 are based on the most limiting result using an enthalpy hot channel factor. F",,, of 1.65 for VANTAGE 5 fuel and an l

F",, of 1.55 for LOTAR fuel and a reference cosine with a pgak of 1.55 for axial power shape. An allowance is included for un increase in T ,, at reduced power j based on the expression: F",, . 1. 0 [1 4 0.3 (1-0 ] M V M E 5 hel and l F",7 - 1.55 [1 + 0.3 (1-P)] for LOPAR fuel l vhere P is the fraction of RATED THERMAL POVER. These limiting heat flux conditions are higher than those calculated for the range of all control rods fully withdravn to the maximum alloveble control rod insertion assuming the axial power imbalance is within the limits of the f 3 (delta I) function of the Overtemperature trip. When the axial power imbalance is not within the tolerance, the axial pover imbalance effect on the Overtemperature delta T trips vill reduce the setpoints to provide protection consistent with core safety limits. , 2.1.2 REACTOR COOLANT SYSTEM PRESSURE The restriction of this Safety Limit protects the integrity of the Reactor Coolant System from overpressurization and thereby prevents the release of radionuclides contained in the reactor coolant from reaching the containment atmosphere. The reactor pressure vessel, pressurizer and the reactor coolant system D piping and fittings are designed to Section III of the ASME Code for Nuclear Power Plant which permits a maximum transient pressure of 110% (2735 psig) of design pressure. The Safety Limit of 2735 psig is therefore consistent with the design criteria and associated code requirements. The entire Reactor Coolant System is hydrotested at 3107 psig, 125% of design pressure, to demonstrate integrity prior to initial operation. ( FARLEY - UNIT 2 B 2-2 AMENDMENT No.

.-. . _ , _ _ - - - -_- _ = - - - _ - - . ... - . -- - _ - . .- I i 2.2 LIMITING SAFETY SYSTEM SETTINGS BASES 2.2.1 REACTOR TRIP SYSTEM INSTPUMENTATION SETPOINTS The Reactor Trip Setpoint Limits specified in Table 2.2-1 are the values at which the Reactor Trips are set for each functional unit. The Trip Setpoints have been selected to ensure that the reactor core and reactor coolant system l are prevented from exceeding their safety limits during normal operation and I design basis anticipated operational occurrences and to assist the Engineered 1 Safety Features Actuation System in mitigating the consequences of accidents. Operation with a trip set less conservative than its Trip Setooint but within its specified Allovable Value is acceptable on the basis that Me difference between each Trip Setpoint and the Allovable Value is equal to or less than the drift allowance assumed for each trip in the safety analysis. Manual Reactor Trip The Manual Reactor Trip is a redundant channel to the automatic protective instrumentation channels and provides manual reactor trip capability. Power Range, Neutron Flux The Power Range, Neutron flux channel high setpoint provides reactor core protection against reactivity excursions which are too rapid to be protected by temperature and pressure protective circuitry. The lov setp, int provides redundant protection in the power range for a power excursion beginning from lov power. The trip associated with the lov setpoint may be manually bypassed when P-10 is active (two of the four power range channels indicate a power level of O above approximately 10 percent of RATED THERHAL POVER) and is automatically reinstated when P-10 becomes inactive (three of the four channels indicate a power level below approximately 8 percent of RATED THERMAL POVER). Power Range, Neutrcn Flux, High Rates The Power Range Positive Rate trip provides protection against rapid flux increases which are chcracteristic of rod ejection events from ar/ power level. Specifically, this trip complements the Pover Range Neutron Flux High and Lov trips to ensure that the criteria are met for rod ejection from partial pover. The Pover Range Negative 'Pate trip provides protection to ensure that the DNB design criterion is met for control rod drop accidents. At high power a multiple rod drop accident could cause local flux peaking which, when in conjunction with nuclear power being maintained equivalent to turbine power by action of the automatic tod control system, could cause an unconservative local DNBR to exist. The Power Range Negative Rate trip v411 prevent this from occurring by tripping the reactor for multiple dropped rods. No credit was taken for operation of this trip in the accident analyses; however, its functional capability at the specified trip setting is required by this specification to enhance the overall rellat,ility of the Reactor Protection System. O . FARLEY - UNIT 2 B 2-3 AMENDMENT NO.

LIMITING SAFEYY SYSTEM SETTINGS O b 8ASES Intermediate and Source Range, Nuclear Flux The Intermediate and Sour:e Range, Nuclear Flux trips provide reactor c re protection during reactor startup. These trips provide redundant potection to the lov setpoir t trip of the Power Range, Neutron Flux chapnels. The Source Range Channels vill initiate a reactor trip at about 10' counts per second unless manually blocked when P-6 becomes active. The Intermediate Range Channels vill initiate a reactor trip at a current level proportional to approximate?.y 25 percent of RATED THERMAL POVER unless manually blocked when P-10 'oeco:nes active. No credit was taken for operation of the trips associated with either the Intermediate or Source Range Channels in the accident analyaes however, their functional capability at the specified trip settings is required by this specification to enhance the overall reliability of the Reactor Protection System. Overtemperature oT The Overtemperature telta T trip provides core protection to prevent DNB for all combinations of nr:ssure, pover, coolant temperature, and axial power distribution, provided tnat the transient is slov vith respect to piping transit, thermovell, and RTD response time delays from the core to the l temperature detectors (about 4 seconds), and pressure is withi: the range between the High and Lo.r Pressure reactor trips. This setpoint includes corrections for changes in density and heat capacity of water with temperature and dynamit: compensation for tranapert, thermovell, and RTD 5 V) response time delays from the core to RTD output indication. Vith normal axial power distribution, this reactor trip limit is always below the core safety limit as shown in Figure 2.1-1. If axial peaks are greater than design, as indicated oy the difference between top and bottom power range nuclear detectors, the reactor tr!p is automatically r-duced according to the notations in Tab *te 2.2-1. Operation with e. reactor coolant loop out of service below the 3 loop P-8.setpoint does not require reactor protection system setpoint modi'ication because the P-8 setpoint and associated trip vill prevent DNB during 2 loop opers. tion exclusive of the Overtemperature delta T setpoint. Two loop operation above the 3 loop P-8 setpoint is permissible after resetting the K1, K2, and K3 inputs to the Overtemperature delta T channels and raising the P.8 setpoint to its 2 loop value. In this mode of operation, the P-3 interlock and trip functions as a High Neutron Flux trip at the reduced pcver level. O FARLEY - UNIT 2 B 2-4 AMENDMENT NO.

LIMITING SAFETY SYSTEM SETTINGS / t 4 BASES V Overpover AT The Overpover delta T reactor trip provides assurance of fuel integrity (e.g., no fuel pellet melting) under all possible overpover conditions, l limits the required range for Overtemperature delte T protection, and provides a backup to the High Neutron Flux trip. The setpoint includes corrections for axial power distribution, changes in density and heat capacity of water with temperature, and dynamic compensation for transport, thermovell, and RTD response time delays from the core to RTD output indica' ion. No credit was taken for operation of this trip in the accident analyses; however, its functice? capability at the specified trip setting is required by this specifici lon to enhance the overall reliability of the Reactor Protection System. Pressurizer Pressure The Pressurizer High and Lov Pressure trips are provided to limit the pressure range in which reactor operation is permitted. The High Pressure trip is backed up by the pressurizer code safety valves for RCS overpressure protection, and is therefore set lover than the set pressure for these valves (2485 psig). The Lov Pressure trip provides protection by tripping the reactor in the event of a loss of reactor coolant pressure. Pressurizer Vater Level The Pressurizer High Vater Level trip ensures protection against Reactor Coolant System overpressurization by limiting the water level to a volume sufficient to retain a steam bubble and prevent water relief through the pressurizer safety valves. No credit was taken for operation of this trip in the accident analyses: however, its functional capability at the specified trip setting is required by this specification to enhance the overall reliability of the Reactor Protection System. Loss of Flov The Loss of Flov trips provide core protection to prevent DNB in the event of a loss of one or more reactor coolant pumps. Above 10 percent of RATED THERMAL POVER, an automatic reactor trip vill occur if the flov in any two loops drop belov 90% of nominal full loop flov. Above 36% (P-8) of RATED THERMAL POVER, automatic reactor trip vill occur if the flow in any single loop drops belov 90% of nominal full loop flov. This

FARLEY - UNIT 2 B 2-5 AMENDMENT NO.

LIMITING SAFETY SYSTEM SETTINGS BASES n ._,

latter trip vill ensure that the DNB design criterion is met during normal l operational transients and anticipated transients when 2 loops are in operation and the Overtemperature delta T trip setpoint is adjusted to the value specified for all loops in operation. Vith the Overtemperature delta T trip setpoint adjusted to the value specified for 2 loop operation, the P-8 trip at 66% RATED THERHAL POVER vill ensure that the DNB design criterion is met during l normal operational transients and anticipated transients with 2 loops in operation. Steam Generator Vater Level The Steam Generator Vater Level Lov-Lov trip provides core protection by preventing operation with the steam generator vater level below the minimum volume required for adequate heat removal capacity. The specified setpoint provides allovance that there vill be sufficient vater inventory in the steam generators at the time of trip to allow for starting delays of the auxiliary feedvater system. Steam /Feedvater Flov Hismatch and Lov Steam Generator Vater Level The Steam /Feedvater Flov Hismatch in coincidence with a steam Generator Lov Vater Level trip is not used in the transient and accident analyses but is included in Table 2.2-1 to ensure the functional capability of the specified trip settings and thereby enhance the overall reliability of the Reactor Protection System. This trip is redundant to the Steam Generator Vater Level Lov-Lov trip. The Steam /Feedvater Flow Hismatch portion of this trip is activated when,the steam flov exceeds the feedvater flov by greater than or equal to 1.55 x 10 lbs/ hour. The Steam Generator (V} Lov Vater Level partion of the trip is activated when the water level drops t-lov 25 percent, as indicated by the narrow range instrument. These trip values incluoc ! sufficient allovance in excess of normal operating values to preclude spurious tr.fs but vil) initiate a reactor trip before the steam generators are dry. Therefore, the required capacity end starting time-requirements of the auxiliary feedvater pumps are reduced and the resulting thermal transient on the Reactor Coolant System and steam generators is minimized. Undervoltage and Underfrequency - Reactor Coolant Pump Busses The Undervoltage and Underfrequency Reactor Coolant Pump bus trips provide reactor core protection against DNB as a result of loss of voltage or under-frequency to more than one reactor coolant pump. The specified setpoints assure a reactor trip signal is generated before the lov flow trip setpoint ( C FARLEY - UNIT 2 B 2-6 AMENDHENT NO.

i l l REACTIVITY CONTROL SYSTEMS H0DERATOR TEMPERATlTRE COEFFICIENT

LIMITING CONDITION FOR OPERATION i l 3.1.1.3 The moderator temperature coefficient (HTC) shall bet

a. Less than or equal to 0.7 x 10 delta k/k/'F for the all rods withdravn, beginning of cycle life (BOL), condition for power levels up to 70% THERHAL POVER vith a linear ramp to O delta k/k/*F at 100%

THERHAL POVER.

b. Less negative than -4.3 x 10 delta k/k/'F for the all rods withdrawn, end of cycle life (EOL), RATED THERHAL POVER condition.

APPLICABILITY: Specification 3.1.1.3.a - H0 DES 1 and 2* only# Specification 3.1.1.3.b - H0 DES 1, 2 and 3 onlyt ACTION:

a. Vith the HTC more positive than the limit of 3.1.1.3.a above, operation in H0 DES 1 and 2 may proceed provided:

1. Control rod withdraval limits are established and raaintained sufficient to restore the HTC to within its limit within 24

,o hours or be in HOT STANDBY vithin the next 6 hours. These lj vithdraval limits shall be in addition to the insertion limits of Specification 3.1.3.6.

2. The control rods are maintained within the withdrawal limits established above until a subsequent calculation verifies that the HTC has been restored to vithin its limit for the all rods withdrawn condition.

3. A Special Report is prepared and submitted to the Commission pursuant to Specificatian 6.9.2 vithin 10 days, describing the value of the measured H10, the interim control rod vithdraval limits and the predicted average core burnup necessary for restoring the positive HTC to within its limit for the all rods withdrawn condi:fon.

b. Vith the HTC more negative than the limit of 3.1.1.3.b above, be in HOT SHUTD0VN vithin 12 hours.

Vith K,,, greater than or equal to 1.0 W See Special Test Exception 3.10.3 O

FARLEY - UNIT 2 3/4 1-4 AMENDHENT NO.

REACTIVITY CONTROL SYSTEMS ROD DROP TIME

      )    LIMITING CONDITION FOR OPERATION                                                                                                                                            l 3.1.3.4 The individual full length (shutdovn and control) rod drop time                                                                                                     l from the fully withdravn position (225 to 231 steps, inclusive)* shall be less                                                                                              ;

than or equal to 2.7 seconds frem beginning of decay of stationary gripper l l coil voltage to dashpot entry with: 1

a. T,,, greater than or equal to 541er, and

b. All reactor coolant pumps operating.

APPLICABILITY: MODES 1 and 2. ACTION:

a. Vith the drop time of any full length rod determined to exceed the above limit, restore the rod drop time to within the above limit prior to proceeding to MODE 1 or 2.

b. Vith the rod drop times within limits but determined with 2 reactor coolant pumps operating, operation may proceed provided THERMAL POVER is restricted to less than or equal to 66% of RATED THERMAL POVER.

sv % f )

 %J SURVEILLANCE REQUIREMENTS 4.1.3.4 The rod drop time of full length rods shall be demonstrated through measurement prior to reactor criticality:

a. For all rods following each removal of the reactor vessel head,

b. .For specifically affceted individual rods follov'ng any maintenance on or modification to the control rod drive system which could affect the drop time of those specific rods, and

c. At least once per 18 months.

           *The fully withdrawn position used for determining rod drop time shall be greater than or equal to the fully withdravn position used during subsequent plant operation.

FARLEY - UNIT 2 3/4 1-19 AMENDMENT NO. l L

POVER DISTRIBUTION LIMITS 3/4.2.2 HEAT FLUX HOT CHANNEL FACTOR - F,(ZJ LIMITING CONDITION TOR OPERATION 3.2.2 F,(Z) shall be limited by the following relationships: F,(2) $ [2.45] [K(Z)] for P > 0.5 f or VANTAGE 5 f uel l P F,(Z) f [4.9) (K(Z)] for P $ 0.5 f or VANTAGE 5 fuel and l F,(2) f [2.32] [K(Z)] for P > 0.5 for LOPAR fuel l P F,(Z) f [4.64] [K(Z)] for P f 0.5 for LOPAR fuel l vhere P - THERHAL POVER RATED THERHAL POWER and K(Z) is the function obtair.ed from Figure (3.2-2) for a given core height location. APPLICABILITY: H0DE 1 ACTION: Vith F,(Z) exceeding its limit:

a. Reduce THERHAL POVER at least 1% for each 1% F,(Z) exceeds the limit within 15 minutes and similarly reduce the Power Range Neutron Flux-High Trip Setpoints within the next 4 hours; POVER OPERATION may proceed for up to a total of 72 hours; subsequent POVER OPERATION may proceed provided the Overpower delta T Trip Setpoints have been reduced at least 1% for each 1% F The Overpover delta T TripSetpointreductionsEa(2)exceedsthelimit.

ll be performed with the reactor in at least HOT STANDBY.

b. Identify and correct the cause of the out of limit condition prior to increasing THERHAL POVER above the reduced limit required by a, above; THERHAL POVER may then be increased provided F,(Z) is demonstrated through incore mapping to be within its limit O FARLL"I - UNIT 2 3/4 2-4 AMENDHENT NO.

1.2 o.0.1.0 s.o,1.o 12.o, o.933

bo 0.8 L O N a 4 0.6 2 cc O . Z

                $0.4 0.2 0

O 2 4 6 8 10 12 CORE HEIGHT (FEET) Figure 3.2.2 K(2) Normalized F,(Z) as a } metion of Core Height FARLEY - UNIT 2 3/4 2-7 AMENDMENT NO.

POVER DISTRIBUTION LIMITS i 3/4.2.3 NUCLEAR ENTHALPY HOT CHANNEL FAf#OR - T" o LIMITING CONDITION FOR OPERATION 3.2.3 T" a shall be limited by the following relationship F" g 3 1.65 [1 + 0.3 (1-P)) for VANTAGE 5 fuel and l l F" o $ 1.55 [1 + 0.3 (1-P)] for LOPAR fuel vhere P = THERMAL POVER RATED THERMAL POVLR APPLICABILITY: MODE 1 ACTION: Vith F"g exceeding its limits

a. Reduce THERMAL POVER to less than 50% of RATED THERMAL POVER vithin 2 hours and reduce the Pover Range Neutron Flux-High Trip Setpoints to <-

55% of RATED THERMAL POVER vithin the next 4 hours.

b. Demonstrate through in-core mapping that F"o is within its limit within O 24 hours after exceeding the limit or reduce THERMAL POVER to less than 5% of RATED THERMAL POVER vithin the next 2 hours, and

c. Identify and correct the cause of the out of limit condition prior to increasing THERMAL POVER above the reduced limit required by a or b, above; subsequent POVER OPERATION may p'coceed provided that F"g is demonstrated through in-core mapping to be within its limit at a nominal 50% of RATED THERMAL POVER prior to exceeding this THERMAL POVER, at a nominal 75% of RATED THERMAL POVER prior to exceeding this THERMAL POVER [

and within 24 hours after attaining 95% or greater RATED THERMAL POVER. O FARLEY - UNIT 2 3/4 2-8 AMENDMENT NO.

                                                                                                                                     - . - . .      _ . . .                        .. . _ _ m     _
                        .M      f" 4.

fc .

y. i .POVER DISTRIBUTION LIMITS r, ; ,

7 y gwy DNB PAPMETERS' LIMITING CONDITION l FOR OPERATION , g - - m 3.2.$ _. The fv116ving DNB related parameters-shall be maintained within the limits shown on-Table 3.2-1:

a. -Reactor Coolant System T,y w b. Pressurizer Pressure ,

[' c.^ Reactor Coola t System Total Flov Rate. L R fAPPL)CABILITY: H0DE l' EACTION:: '" ~

                  -Vith'any of the above parameters exceeding !ts limit, restore the parameter to-
                 -vithin its limit within 2 hours oc reJuce THERMAL F0VEh to less than 5% of RATED
                 -THERMAL POVER within the next-4 hours.

_ : SURVEILLANCE RE0biREMENTS

p. '

                                                                            - rameters of Table 3.Z-1 shall be verified to be within q f (4.2;5.1    their. lieni            Each   ts r aof'*h t . Luet . ,ee_per 12 hours.

4.2.5.2 *2he Rear ia . Coolant System total, flow rate-shall be determined to-be ivithin its. limit , measurement at least once per 18 months.

4.?.543 Th d adicated RCS flow rate shall be verified to be within the ecceptable limit-:at lea r once per 31 days.

w 3 4 -FARLEY - UNIT 2 3/4 2-14 AMENDHENT No. 67 f

   ,                       -            , ,                     .,,;,-..s.,     nn.       a w,, - - . . . . ,          . . , , _ , .           -n,,  ,- + . . , . , - ,..,n.ne..-e          -, ,, v ,.   -n,, &

y. - -

y y . ;;; _. c d - JTABLE .i.2-1. t ~

                                                                 . DNB PARAMETERS -

w

I

                                                                     -LIMITS h

H PARAME""2R 3 Loops in Operation: 2 Loops in or ration: 9 Indicated Reactor l Coolant System T,,,; .4580.7'F' (**) -l Indicated Pressurizer Pressure 2 2 05'psig*' ~(**). l: Indicated reactor Coolant. System 1 267,880 gpa***' (**)' _l a

             -Total Flow Rate                                                                                                                                                                                               !

1, t I'

1 ta i 2 , I' G t

i .

i ! i > ' $

Limit not-applicable during either a THERMAL POWER ramp in excess of 5% of RATED THERMAL POWER per- ,

     $             ~ minute or a THERMAL P0FSR step'in' excess of.10% of RATED THERMAL POWER.-

Values blank pending-NRC approval of 2 loop operation.

    .4
     ~z o        *** Value' includes'a 2.4% flow uncertainty (0.1% feedvater venturi fouling bias included).

l-l L a* _ . .._ u- . _ _ _ _ ._ _ . _ _ . _ _ _ _ _ __-______________..____._._______________.___.__m - m____m__m.

(.

                                                        %)                                                 Y.]

TABLE 3.3-2 5 g REACTOR TRIP SYSTEM INSTRUMEffrATION RESPONSE TIMES 5 FUNCTION /.L UNIT RESPONSE TIME' E

1. Manual' Reactor Trip Not Applicable

2. Power Range, Neutron Flux

a. High 5 0.5 seconds *

b. Lov Not Applicable

2. Power Range, Neutron Flux, High Positive Rate Not Applicable

4. Power Range, Neutron Flux, High Negative Rate Not Applicable l

{ 5. Intermediate Range, Neutron Flux Not Applicable u 6. Source Range, Neutron Flux Not Applicable L 0

7. Overtemperature Er 5 6.0 seconds

l

8. Overpover Er Not Applicable

9. Pressurizer Pressuta -Lov j 2.0 seconds

10. Pressurizer Pressure-High f 2.0 seconds l 11. Pressurizer Vater Level-High Not Applicable E

E E h

Neutron detectors are exempt from response time testing. Response-time of the neutron flux signal portion 2 of the channel shall be measured from detector output or input of first electronic compone,t in channel.

s

                                                                                  ~~
    .o                              ,

04 TABLE 3.3-4 (Continued). 1 O

g. ENGINEERED'SAFETT FEATtiPE ACTUATION SYSTEN INSTRUMEMATION TRIP SETPOIM S-5;

f FUNCTIONAL UNIT ' TRIP SETPOINT ALLOVABLE VALUES h 2. C0ffrAINMENT SPRAT H 1 m a. Manual Initiation Not Applicable Hot Applicable

b. Automatic Actuation Logic Net Applicable Not Applicable

c. Containment Pressure- . $ 27 psig ,y 28.3 psig l High-High-High

       ". C0PrrAINilEfff ISOLATION

a. Phase "A" Isolation.

 $                1. Manual                                         Not Applicatle Not Applicable w                2. From Safety Injection'                         Not Applicable Not Applicable-Automatic Actuation Logic h

b. Phase "B" Isolation

1. Manual Not Applicable Not Applicable

2. Automatic'Actuatien Logic Not Applicable Not Applicable ,

3. Containment Pressure-- f 27 psir f.28.3 psig .l High-High-High

c. Purge and Exhaust 1 solation i 5 1. Manual Not Applicable Not Applicable 5

ih 2. Autostatic Actuation Logic Not Applkeble Not Applicable l 4 5 i I

p ,- , 0 N %) _

                                                        . TABLE 3.3-4 (Continued)

C . ENGINEERED SAFETT FE/.TURE ACTUATION SYSTEM INSTRUMENTATION TRIP SETPOINIS i FUNCTIONAL UNI

TRIP SETPOINT ALLOVABLE VALUES

 "   4. STEAM LINE ISOLATION u

a. Manual Not Applicable Not Applicable

b. Automatic Actuation Not Applicable Not Applicable Logic

c. Containment Pressure- f 16.2 psig f 17.5 psig i i High-Righ

d. Steam Flov in Tvc Steam f A function defined as follovs: $ A function defined as follows:

w Lines-High, Coincident A op corresponding to 40% of full A op corresponding.to 44% of full 2 steam flov between 0% and 20% load steam flow between 0% and 20% load

vi th T"'-Low-Lov and then a op increasing linearly and then a op increasing linearly Y to a op corresponding to 110% of to a op corresponding to 111.5% of

 "                                                  full steam flov at full load with    full steam flow at full load with T,,,    1 543'F                      T,,,   1 540'F                                                                                         l i          e. Steam Line Pressure-Lov               2 585 psig                           1 575 psig i

l 5. TURBT.NE TRIP AND FEED VATER ISOLATION

a. Steam Generator Vater 5 75% of narrov range instrument f.76% of narrov range instrument Level-Figh-High span each steam generator span each steam generator E

e M 5

Gjlf - f-

                                                                ,LT                                           '

ENGINEERED SAFETY FEATURE ACTUATION ~ SYSTEM INSTRUMENTATION TRIP-SETPOINr5

E FIMCTIONAL UNIT' TRIP SETPOINT. -ALLOVABLE VALUES

  ' ti -      .    -         .
    ,         6. AUXILIARY FEEDVATER.

E s a. Automatic Actuation

N.A. N.A. H Logic , u !

                 'b. Steam Generator Vater                   1.17; of. narrow range instrument                   -2 162 of' narrow range instru=ed Level-Low-Lov                         ' span each steam generator'                         ' span each steam generator

c. Undervoltage - RCP 2'2680 volts' 1 2640 volts

                 -d., S.I.                                    See 1 above.(all SI Setpoints)

e. Trip of Main Feedvater N.A. N.A.

Pumps- , s w :

,   i        ~7. LOSS OF POVER-y             a.- 4.16 kv Emergency Bus                   1 3255: volts bus voltage *.                         2 3222 volts bus ec1 age *                                                            ,

y Undervoltage (Loss of 5 3418 volts bus voltage

Voltage) "

i

b. 4.16 kv Emergency Bus ~ 2 3675 volts bus voltage

l'3638 volts bus voltage
l' Undervoltage (Degraded 5 3749' volts bus voltage
i Voltage) i-

8. ENGINEERED SAFETT FEATURE ACTUATION SYSTEM INTERLOCKS

a. Pressurizer Pressure, f 2000 psig $ 2010 psig '

P-11 , E . . m b. Low-Low T (Increasind,P-12 544*F f 5'>7'F ; h (Decreasing) e 543*F -1 540*F l 4 ; z c. Steam Generator Level,' (See 5. above)' *

    ?                 P-14 l                  d. Reactor Trip. P-4                       N.A.                                                 11. A.

Refer to appropriate. relay setting sheet calibration. requirements.

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

l TABLE 3.3-5 (Continued) ENGINEERED SAFETY-FEATURES RESPONSE TIMES q - Q) INITIATING SIGNAL AND FUNCTION RESPONSE TIME'IN SECONDS _ 3 ._ - - Pressurizer Pressure-Lov

a. Safety Injection (ECCS) f 27.0'*'/12.0 ,

b .- Reactor Trip (from SI) f2.0 l

c. 'Feedvater Isolation f 32.0

d. Containment Isolation-Phase "A" 3 17.0

e. Containment = Purge Isolation f 5.0

f. _ Auxiliary Feedvater Pumps- No' Applicable lg. Service Water System' _f 77.03 /87.0'*'

1 4.-- -Differential Pressure Between Steam Lines-High .I

a. . ; Safety Injection-(ECCS) f 12.0/22.0

b. Reactor Trip-(from SI). 'f 2.0

c. Feedvater Isolation- f 32.0 l

d. Containment-Isolation-Phase "A" f 17.0'/27.08

',, e. . Containment Purge Isolation Not Applicable l f ~fA Auxiliary Feedvater Pumps. Not Applicable j

          /            c g. .         Service WateriSystem                               < 77.0'*8 /87.0'53

5'. Steam Flov'in'Two Steam Lines-High coincident vith T --Low-Lov-

                       -a.            Steam Line Isolation                               Not Applicable                    l

6. - Steam Line Pressure-Low-

a. Safety Injection (ECCS) -f 12.0/22.0 ,

b. . Reactor Trip-(from SI) . f 2.0

c. - FeedvateriIsolation. f 32.0

                       -d.            Containment Isolation-Phase "A"                    f 17.0/27.0

e.. . Containment Purge-Isolation Not Applicable

f. LAuxiliary Feedwater Pumps Not-Applicable

                       'g.            Service Vater-Syetem                               f 77.0/87.0

h. Steam Line Isolotion f 7.0 O FARLET - UNIT ' 2.-- 3/4 3-30 AMENDMENT NO.

b u REACTOR COOLANT SYSTEM HOT STANDBY LIMITING CONDITION FOR OPERATION 3.4.1.2 -At least two of the Reacts ;oolant Loops listed belov shall be OPERABLE and in operation when the roo control system is operational or at least two Reactor. Coolant Loops listed belov shall be OPERABLE vith one Reactor Coolmot Loop in operation when the rod control system is disabled by opening the Reactor Trip. Breakers or shutting down the rod drive motor / generator sets:*

1. Reactor Coolant Loop _A and its associated steam generator and Reactor Coolant pump,

2. Reactor Coolant Loop B and its associated steam generator and Reactor Coolant pump,

3. Reactor Coolant Loop C and its associated steam generator and Reactor Coolant pump.

APPLICABILITY: H0DE 3 ACTION:

a. Vith less-than the above required Reactor Coolant Loops OPERABLE, restore the requirs.d loops to OPERABLE status within 72 hours or be in HOT.SHUTD0VN vithin the next 12 hours.

     .O          b. ..Vith only one Reactor Coolant Loop.in operation and the rod control V-               system operational, within 1 hour open the Reactor Trip Breakers or
       ~

shut dcyn the rod drive motor / generator sets.

c. . Vith no Reactor. Coolant Loops in operation, suspend all operations involving a reduction in boron concentration of the Reactor Coolant System-and immediately initiate corrective action to return the required coolant loop to operation.

SURVEILLANCE REQUIREMENTS-4.4.1.2.1 At least the above required Reactor Coolant pumps, if not in operation, shall-be determined to be CPERABLE once~per 7 days oy verifying correct breaker alignments and indicated power availability. The required Reactor Coolant Loop (s) shall be verified to be in

         -4".4.1.2.2
~

operation and circulating Reactor Coolant at least once per 12 hours. 24'.4.1.2.3 The~ required steam generator (s) shall be determined OPERABLE by

verifying secondary side water le 'l to be greater than or equal to 10% of vids range indication at least once pt.

hours.

All Reactor Coolant pumps-may be de-energized for up to I hour--provided (1) no operations are permitted.that would cause dilution of the Reactor Coolant System boron concentration, and (2) core outlet temperature is maintained at

' OQ, least~10'F belov saturation temperature. FARLEY - UNIT 2 3/4 4-2 AMENDMENT NO.

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

L 3/4.2 VO'JER DISTRIBUTION LIMITS .

            ' BASES

~

~.

(~L .;during _ The specifications of this section' provide assurance of fuel integrity Condition'I (Normal Operation) and II (Incidents of Moderate Frequency)

events bys: (a) meeting the DNB design criterion during normal operation and l in shcrt tera transients,-and (b) limiting the fission gas release, fuel pellet

             . temperature and cladding mechanical properties to within assumed design                                          i criteria. In addition, limiting the peak linear-power density during Condition
            ;I. events-provides assurance that the initial conditions assumed for the LOCA analyses are met and the ECCS acceptance criteria _ limit of 2200'F is not exceeded.

The definitions of certain hot channel and peaking factors as used in these specificatiens are as follows:

F,(Z) 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 and measurement uncertainty. l F(H 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.

Fy(Z) Rajial Peaking Factor, is defined as the ratio of peak power density to average power density in the horizontal plane at core elevation Z.

[VV 3 /4.' 2.1 -AXIAL FLUX DIFFERENCE ~ The limits'on AXIAL FLUX DIFFERENCE (AFD) assure that the F (Z) upper bound envelope of 2.45 for VANTAGE 5 and 2.32 for LOPAR times th,e normalized- l axial peaking factor is'not exceeded during either normal operation or in the event of xenon redistribution following' power changes. ~

                     -Target flux difference is determined at equilibrium xenon conditions. The full-length rods may be positioned-vithin the core in accordance with their respective insertion' limits and should be inserted near their normal position
            .for steady state operation-at high power levels. The'value_of the target flux l difference obtained-under these conditions divided by the fraction of RATED THERMAL POVER is the istget flux difference at RATED THERMAL F0VER for the
             -associated core buri.. . :onditions. Target flux differences for other THERMAL POWER levels are obtained by multiplying the RATED THERMAL POVER value by-the appropriate. fractional THERMAL POVER level. The periodic updating of the
            ; target flux difference value is necessary to reflect core burnup considerations.

'N

d. -

FARLEY-- UNIT 2 B 3/4 2-1 AMENDMENT NO.

POVER DISTRIBUTION LIMITS p y BASES AXIAL FLUX DIFFERENCE (Continued) Although it is intended that the plant vill be operated with the AFD-vithin the +(5)% target band about the target flux difference, during rapid plant THERMAL POVER reductions, control rod motion vill cause the AFD to deviate outside of the target band at reduced THERMAL POVER levels. This deviation vill not affect the xenon redistribution sufficiently to change the envelope of peaking factors which may be reached on a subsequent return to RATED THERMAL POVER (with the AFD vithin the target band) provided the time duration of the deviation is limited. Accordingly, a 1 hour penalty deviation limit cumulative during the previous 24 hours is provided for operation outside of the tarFet bans but within the limits of Figure (3.2-1) while at THERMAL POVER levels between 50% and 90% of KATED THERMAL POVER. For THERMAL POVER levels between 15% and-50% of RATED THERMAL POWER, deviations of the f.FD outside of the target band are less aignificant. The penalty of 2 hours actual time reflects this reduced significance. Provisions for monitoring the AFD on an automatic basis are derived from the plant process computer through the AFD Monitor Alarm. The computer determines the one minute average of each of the OFERABLE excore detector outputs and provides an alarm message immediately if the AFD for 2 or more OPERABLE excore channels are outside the target band and the THERMAL POVER is greater than 90% of RATED THERMAL POWER. During operation at THERMAL POVER 1evels between 50% and 90% and between 15% and 50% RATED THERHAL POVER, the q y computer outputs an alarm message when the penalty deviation accumulates beyond the limits of I hour and 2 hours, respectively. Figure B 3/4 2-1 shows a typical monthly target band. 3/4.2.2 an6 3/4.2.3 HEAT FLUX HOT CHANNEL FACTOR, NUCLEAR ENTHALPY HOT CHANNEL FACTOR The limits on heat flux hot channel factor, and nuclear enthalpy rise hot channel factor ensure that 1) the design limit on peak local power density is not exceeded, 2) the DNB design criterion is met, and 3) in the event of a LOCA the peak fuel clad temperature v!11 not exceed the 2200'F ECCS acceptance criteria limit. Each of these is measurau. but vill normally only be determined periodically as specified in Specifications 4.2.2 and 4.2.3. This periodic surveillance is sufficient to insure that the limits a e maintained provided:

a. Control rods in a single group move together with no individual rod insertion Jiffering by more than + 12 steps, indicated, from the

group demand position.

b. Control rod banks are sequenced .rith overlapping groups as described in Specification 3.1.3.6.

c. The centrol rod insertion limits of Specifications 3.1.3.5 and 3.1.3.6 O

V _ are maintained.

d. The axial power distribution, expressed in terms of AXIAL FLUX

DIFFERENCE, is maintained within the limits. FARLEY - UNIT 2 B 3/4 2-2 AMENDMENT NO.

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

    'POVER DISTRIBUTION LIMITS                                                                                                                          :

BASES F" H vill be maintained within its limits providgd conditions a. through,d. above are maintained. The relaxation of F

            ' THERMAL POVER' allows chenges in the radial power                                              forshape,H          as a function of all permissible rod. insertion limits..

When an F measurement is taken, an allovance for both m perimental error and manuiacturing tolerence must be made. An allowance ~of 5% is appropriate for a full core map taken with the incore detector flux mapping system and a 3% allovance.is appropriate for manufacturing tolerance. When F" H is measured, experimental error must be allowed for and 4% is. taken with the incore the appropriate detection system.s11ovance for a full The specified core limit formap' F H contains an 8% allovance for uncertainties. The 8% allovance is based on ,the following considerations:

a. Abnormal perturbations in the radial power shape, such as from rod misalignment, affect F",H more directly than F,,

b. Although rod movement has a direct influence upon-limiting F to

                     .vjthinitslimit,suchcontrolisnotreadilyavailableto1$mit F ;H, and

c. -Litors in prediction for control pover shape detected during startup by restricting exial flux physics tests: This distribution. can be compensated compensation f orfor ip F,is Icss readily available.

F ,H 1 LO FARLEY - UNIT 2 B 3/4 2-4 AMENDMENT NO.

POVER DISTRIBUTION LIMITS BASES a (Z), is measured periodically to provide additional 1

'")  -

The radialthat assurance peaking the hotfactor channF,El factor, F (Z), remains within its limit. The F limit for RATED THERHAL POVER (FRTP) as provided in the Radial Peaking Factor " limit report per Specification 6 3.1.11 was determined from expected power control maneuvers over the full range of burnup conditions in the core. 3/4.2.4 OUADRANT POVER TILT RATIO The quadrant power tilt ratio limit assures that the radial power distribution satisfies the design values used in the power capability analysis. Radial power distribution measurements are made during startup testing and periodically during power operation. The limit of 1.02, at which corrective action is required, provides DNB and

          > linear heat generation rate protection with x-y plane power tilts.

The two hour time allovance for operation vish a tilt condition greater than 1.02 but less than 1.09 is provided to allow identification and correction of a dropped or misaligned control rod. In the event such action does not correct the tilt, the margin for uncertainty on F is reinstated by reducing the maximum allowed power by 3 percent for each perce,nt of tilt in excess of 1.0.

         'For purposes of monitoring QUADRANT POVER TILT RATIO vhen one excore detector is inoperable, the movable incore detectors are used to confirm that the normalized symmetric power distribution is consistent with the QUADRANT POVER TILT RATIO.

The incore detector monitoring is done with a full incore flux map or two sets of~four symmetric thimbles. The two sets of four symmetric thlmbles is a unique set of eight detector locations. These locations are C-8, E-5, E-11, H-3, H-13, (')N u L-5, L-11, and N-8. 3/4.2.5 DNB PARAMETERS The limits on the DNB related parameters assure that each of the parameters are maintained within the normal steady state envelope of operation assumed in the transient and accident analyses. The limits-are consistent with the initial FSAR assumptions and have been analytically demonstrated adequate to meet the DNB design criterion throughout each analyzed transient. .The indicated T value of 580.7"F is based on the average of two control board readings and,In indication uncertainty of 2.5'F. The indicated pressure value of 2205 psig is based on the average of two control board res, dings and an indication uncertainty of 20 psi. The indicated total RCS flow rate is based on one elbow tap measurement from each loop and an uncertainty of 2.4% flov (0.1% flow is included for feedvater venturi fouling). The V hour surveillance of Tavg and pressurizer pressure through the control board readings are sufficient to ensure that the parameters are restored within their limits following load changes and other expected transient operation. The 18 month surveillance of the total RCS flow rate is a precision measutement that verifies the RCS flow requirement at the beginning of each fuel cycle and ensures correlation of the flow indication channels with the measured loop flows. The monthly surveillance of the total RCS flow rate is a reverification of the RCS flow requirement using loop elbow tap measurements that are correlated.to the precision RCS flov measurement at tha beginning of the fuel (9, cycle. The 12 hour RCS flov surveillance is a qualitative verification of V significant flow degradation using the control board indicators and the loop elbow tap reasurements that are correlated to the precision RCS flow measurement at the &;dnning of each fuel cycle. FARLLY - UNIT 2 B 3/4 2-5 AMENDHENT NO.

       - R.

3/4.4 REACTOR COOLANT SYSTEM p BASES 3/4.4.1 REACTOR COOLANT LOOPS AND COOLANT CIRCULATION The plant is designed to operate with all Reactor Coolant Loops in operation, and meet the DNB des!gn criterion during all normal operations and l anticipated transients. In MODES 1 and 2 vith one Reactor Coolant Loop not in operation this specification requires that the plant be in at least HOT STANDBY vithin 1 hour. In MODE 3, two Reactor Coolant Loops provide sufficient heat removal l capability for removing core neat even in the event of a bank withdraval accident; however, a single Reactor Coolant Loop provides sufficient decay heat removal capacity if a bank withdrawal accident can be prevented; i.e., by opening the Reactor Trip Breakers or shutting dovn the rod drive motor / generator sets. l In H0DE 4, a single reactor coolant or RHR loop provides sufficient heat removal capability for removing decay heat, but single failure considerations require that at leas' tvo loops be OPERABLE. Thus, if the Reactor Zo>lant Loops are not OPERABLE, this specification requires two RHR loops to be Ol'_RABLE. In MODE 5, single failure considerations require two RHR loops to be OPERABLE. The operation of one Reactor Coolant Pump or one RHR pump provides adequate flow to ensure mixing, prevent stratification and produce gradual reactivity /G changes during boron concentration reductions in the Reactor Coolant System. V The reactivity enange rate associated with boron reduction vill, therefore, be within the capability of operator recognition and control. The restrictions on starting a Reactor Coolant Pump vith one or more Reactor Coolant System cold legs less than or equal to 310'F are prwided to prevent. Reactor Coolant System pressure transients, caused by energy additions from *he secondary system, which could exceed the limits of Appendix G to 10 CFR Part 50. The Reactor Coolant System vill be protected against overpressure tran;ients and vill not exceed the limits of Appendix G by either (1) restricting the water volume in the pressurizer and thereby providing a volume for the primary coolant to expand into, or (2) by restricting starting of the Reactor Coolant Pumps to when the secondary water temperature of each steam generator is less than 50*F above each of the Reactor Coolant System cold leg temperatures. O FARLEY - UNIT 2 B 3/4 4-1 AMENDMENT NO.

ADMINISTRATIVE CONTROLS . _.

e. Type of container (e.g., LSA, Type A, Type B, Large Quantity), and

f. Solidification agent (e. g., cement, urea formaldehyde).

The radioactive effluent release reports shall include unplanned releases from the site to unrestricted areas of radioactive materials in gaseous and liquid effluents on a quarterly basis. The radioactive efiluent release reports shall include any changes to the PROCESS CONTROL PROGRAM (PCP) made during the reporting period. HONTHLY OPERATING REPORT 6.9.1.10 Routine reports of opetr*ing statistics and shutdown experience,

      -including documentation of all N .11enges to the PORV's or safety ve?ves, shall be submitted on a monthly bauls to ne Commission, pursuant to 10 CFR 50.4, no later than the 15th of each month following the calendar month covered by the report.

l Any changes to the OFFSITE DOSE CALCULATION HANUAL shall be submitted with the .

      -Monthly Operating Report within 90 days in which the change (s) was made effective.        !

In addition, a report of any major changes to the radioactive vaste treatment systems shall be submitted with the Monthly Opere ing Report for the period in which the change was implemented. f3 _ RADIAL PEAKING FACTOR LIMIT REPORT b 6.9.1.11 The F limit for RATED THERMAL POVER (FRTP) shall be provided to the Commission, pursulnt to 10 CFR 50.4, for all core plafies containing bank "D" control rods and all unrodded core planes no later than 30 days after cycle l initial criticality. In the event that the limit would be submitted at some other time during core life, it vill be submitted 30 days after the date the l limit-would become effective unless otherwise exempted by the Commission. Any information needed to support FRTP

vill be by request from the NRC and need not be included in this report.

ANNUAL DIESEL GENERATOR RELIABII.ITY DATA REPORT 6.9.1.12- The number of tests (valid or invalid) arid the number of failures to start on demand for each diesel generator shall be submitted to the NRC annually. This report shall contain the information identified in Regulatory Position C.3.b of NRC Regulatory Guide 1.108, Revision 1, 1977. b FARLEY - UNIT 2 6-19 AMENDHENT NO.

O Attachment 3 Joseph H. Farley Nuclear-Plant Units 1 and 2 Req'iest for Technical Specifications Changes 1 _ Sionificant Hazards Evaluation O 4

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

i s } Attachment 3 Significant Hazards Evaluation for the Joseph M. Farley Nuclear Plant Units 1 and 2 Transition to Westinghouse 17x17 VANTAGE-5 Fuel Assemblies 10 CFR 50.92 Evaluation Pursuant-tc 10 CFR 50.92, each application for amendment to an operating license .nust be reviewed to determine if the proposed change involves a

      -significant hazards consideration. The amendment, as defined below, describing the-Technical-Specifications changes associated with                                                     ;

implementation of VANTAGE-5 fuel assemblies has been- reviewed and deemed not to involve significant hazards considerations. The basis for this-determinatiorofollows. Proposed Chanaes The proposed amendment involves the following Technical Specificacions changes. The DNB parameters for RCS Tava, pressurizer pressure, and RCS

      -flow-have been modified. The revised RCS flow rate (267,880 gpm) includes a 2.4% calculated flow uncertainty, which -includes a 0.1% uncertainty for feedwater venturi. fouling. The revised flow rate is also reflected in the

p. low'RCS loo) flow trip setpoint. The RCS Tava and pressurizer pressure Q -

have been c1anged to-include additional' calculation uncertainty (6.00F vs.

      .4.00F and 50 psi vs. 30 psi). In addition, the DNB parameters have been                                            ,

adjusted to make them the indicated values by including indication uncertainty, Modifications to'the OPAT/0 TAT equations resulted from a revision to the core safety limits to reflect the use of VANTAGE-5 fuel,'the Revised

      ? Thermal Design Procedure (RTDP), increased peaking factors, and related
      ' analysis assumptions. These include a reduction in the K1 term (constant in the 0 TAT equation set equel to l.14) and modifications to K2 and K3.to afford protection for the revised core limits.                       In addition, K4 constant

in.the_OPAT equation set equal to 1.07) and K6 were also revised Rod drop (.

time was__ increased from 2.2 to 2.7 seconds to- account for the physical changes associated with' inclusion of VANTAGE-5 fuel. The moderator temperature coefficient: limit was changed to 0.7 x 10-4 ok/k/0F Lp to 70%- i rated thermal power with a ramp from 0.7 x 10-4 Ak/k/0F-to 0.0 x 10 -

Ak/k/0F from 70% to 100% pc ser. Spacifications for the F Q(Z) and the K(Z) curves were modified to reflect the analytical assumptions and to provide design flexibility. In advition, to account for the transition it whb.h VANTAGE-5 and LOPAR (low-rarasitic) assemblies will both be in the car 9, separate Fag and Fg parameters have been provided. For LOPAR fuel, tnt current FAH and FQ have been rotained. These are consistent wit.. 'o i safety analyses assumptions.

Modifications ;to allowable values (based on uncertainty calculations). for P d certain Engineered Safety Features actuation system t-ip setpoints

       .(i.e., Steamline Isolation on Containment- Pressure High High and

Attachment 3

   -O- Page 2 Containment Spray and Phase B Isolation on Containment Pressure High High-High) were necest.ary so that the trip setpoints would not be altered. Allowable values for OT AT/0 PAT trip functions were also modified.

Additional modifications to the f( A1) limits and associated AT trip setpoint reduction are also included to be consistent with the acciJent ar.alyses assumptions. In addition, several bases were changed to reflect the use of the WRB-1 and WRB-2 DNB correlations in the accident analysis. A modification to the response time for high negative flux rate reactor trip was made to reflect that this reactor trip function is no longer used as a primary protective function. In addition, a modification to the Engineered Safety Features (ESF) response time for steam line isolation on high steam flow in two steam lines coincident with low-low Tavg was made to reflect that this initiating signal is no longer used as a primary protective function. Response times for both of these functions have been changed to "Not Applicable" in the Technical Specifications. Analytical results have allowed for the relaxation of the requirement of three (3) RCS pumps and associated loops being operable in Mode 3 to two (2) being operable. An administrative change to the reporting requirements for the Radial Peaking Factor Limit Report bn been proposed to now be 30 days after criticality. Backaround In order to implement the long-term fuel management st6 ategy planned by Alabama Power Company for the Joseph M. Farley Nuclear Plant Units 1 and 2, Westinghouse VANTAGE-5 fuel has been selected for future reloads. This strategy includes the implementation of :urrent high energy 18-month fuel cycles with high capacity factors, low leakage loading patterns, and extended burnup. Westinghouse VANTAGE-5 fuel has been designed to accommodate these operating characteristics by inclusion of specific design features and by use of improved methodologies previously apprond by the NRC. These design features include Intermediate Flow Mixer grids, Modified Debris Filter Bottom Nozzles, zircaloy gri6, natural uranium dioxide (U02) axial blankets, Inteoral Fuel Burnable Absorber, extended burnup, and reconstitutable top nozzles. Each of these features is in support of the safe, efficient fuel management plan for the Farley Units. Extended burnup, reconstitutable top nozzles, and Modified Debris Filter Bottom Nozzles are already in use with L0 PAR fuel at FNP Units 1 and 2. Margin supplied by the VANTAGE-5 fuel-design and by the impruved methodologies have been applied to obtain design flexibility through revised core design parameters that are reflected in the proposed Technical Specifications changes (Technical Specification Sections 3.1.1.3, 3/4.2.2 and 3/4.2.3). These methodologies include use of the RTDP, the WRB-1 and WRB-2 DNB O

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

n ' Attachment 3 Page 3 s

                                                           ~
                             - correlations, the improved THIhC-IV modeling method, and 'the NOTRUMP.

computer code for small-break LOCA. Each feature and methodology is discussed below.- Intermediate Flow Mixer Grids (IFM) - The IFM grids promote flow mixing - i within the assembly and provide increased DNB margin. IFMs also offer

                             - reduced _ rod bow and improved seismic stability of the- assembly.

Axial Blankets - Axial blankets consist of six inches-of natural U02

                             . pellets (instead of enriched uranium) within the fuel rods at each end of the -fuel stack,- which reduces neutron leakage and improves uranium utilization. Axial blankets are optional, starting with the first
     +                          transition cycle.
                             - Inteoral Fuel Burnable Absorber (IFBA)              The advantage provided by IFBA'is improved neutron utilization, since the neutron absorber material is a boride ccating on the: fuel pellets themselves. This reduces the need for individual absorber rods which-displace water molecules that act as neutron moderators;to promote fissions.
                                                 ~

Egtended Burnuo ~- Longer fuel cycles result in increased rod growth and

    - Q                         production of fission product gas which increases rod internal pressure.

V By increasing the fuel = rod plenum and providing additional space between the fuel rod:and nozzle, the effects of-increased internal rod pressure and rod growth resulting.from longer cycles- and extended burnup can be accommodated.- This feature has:previously been licensed at Farley and is currently.in use-at.both units. NOTRUMP - The NOTRUMP. computer code Lis a one dimensional general network code-incorporating several advanced-features, including calculation of

                             ' thermal non-equilibrium in all fluid volumes, ficw -regime dependent. drift flux calculations with counter-current flooding limitations, and regime-dependent heat transfer correlations. NOTRUMP is used in-
                             - small.-break LOCA applications.
                             - Reconstitutable Too Nozzle (RTN) -. The-assembl, RTN-provides the capability to replace damaged fuel-rods. This feature avoids the discharging of an
                                ' entire' assembly for-minor-fuel damage.- This feature has previously been 1.icensed at Farley and-is currently in use at-both units.

Modified Debris Filter Bottom Nozzle (MDFBN1 - The assembly MDFBN is used to~ reduce.the possibility of fuelerod damage dae.to debris-induced fretting. .MDFBN has replaced the-relatively larger holes of the LOPAR-design with -a new pattern of smaller holes. This feature has been licensed and is in~ use .at' both Farle/ units. lircalov Grids'- Zircaloy grids are used in VANTAGE-5 assemblies in place of the inconel grids used in LOPAR designs, except for the top and bottom k - grids. This change is made for reduced parasitic neutron absorption. 4

         .,.---...*---r,--.                    -             ,        -                  + .,,,.-,--e--              ----www      w

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

Attachment 3 '

 ,x(        Page 4 Revised Thermal Desian Procedure (RTDP) - With this methodology, uncertainties in plant instrumentation.that monitors operating parameters, nuclear and; thermal parameters, fuel fabrication parameters,- and DNB correlation predictions are considered statistically to obtain DNBR uncertainty factors. Since the parameter uncertainties are considered in               l determining-the design limit DNBR values, the plant safety analyses are                i performed using values of input parameters without uncertainties. The RTDP
          - methodology was used to provide additional ope. rating margin for the l proposed: Technical Specifications Sections 2.1.1,2.2.1,3/4.2.2,and 3/4.2.3'and associated Bases changes.

WRB The LOPAR fuel DNBR analyses use the WRB-1 DNB correlation. This

          .DNB correlation takes credit for the significant impravement in the accuracy of the critical heat flux predictions over previous DNB correlations. The WRB-1 correlation is incorporated into the Technical Specifications Bases -Sections 2.1.l', 3/4.2, and 3/4.4. The W-3 DNB correlation is used where the primary DNB correlation is not applicable.

Eh2 - This DNB correlation is used for fuel assemblies utilizing IFM ,

          - grids. It takes credit for :,ignificant improvement in the accuracy of the           9 critical. heat flux predie.tions over previous NB correlations as well as           l p        for the reduced grid-to-grid spacing of the VANTAGE-5 fuel ,sembly mixing
 g~       vanes. The WRB-2 DNB correlation is incorporated into the fechnical Specifications Bases Sections 2.1.1, 3/4.2, and 3/4.4. The W-3 DNB correlation-is used where the primary DNB correlation is not applicable.
          -1MPROVED THINC .ly - This improved thermal-hydraulic design modeling method improves the accuracy of the analyses. THINC-IV is used for both LOPAR and VANTAGE-5 fuel DNBR analyses. The improved THINC-IV modeling method was used in the supporting analyses for the proposed changes.to the Technical Specifications Sections' 2.1.1,.3/4.2.2, and 3/4,2.3 and associated Bases.

O 4 S w ..-. - - - - ,

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

Attachment-3' l Page 5_ i 1 l Analysis ~ The proposed Technical Sptcifications changes reflect the operational and core design benefits provided through use of the VANTAGE-5 fuel design. The )roposed changes have been evaluated with respect to plant safety and to tle impact on accident analyses. The implementation of VANTAGE-5 fuel also-includes. increased peaking factors and various other pit.nt design margins. The following evaluations of the proposed Technical Specifications changes have been divided into nine individual groups of changes and each group will be independently assessed relative to the criteria of 10 CFR 50.92. Reactor Core Safety Limits. Re2ctor Trio System Instrumentation Setooints. and-DNB Parameters The IFM grids, the improved THINC-IV thermal-hydraulic design modeling method, the Revised Thermal Design Procedure, and- the WRB-1 and WRB-2 DNB correlations change the basis for determining DNBRs. As such, these items provide '.he basis for proposed changes to the following Technical Specifications. Dl A._ Technical Specification Figure 2.1-1 --The revised DNB methods allow V for revision of the Reactor Core Safety Limits. B. Technical Specification Table 2,2-1.- The revised reactor core safety limit lines allow for changes in the Overtemperature AT and Overpower AT Reactor Trip System Instrumentation Setpoints. Specifically, changes to constants 1K , K 2 , K 3 , K4 , and K6, and the fL(AI) function are-proposed. In addition, setpoints and allowable va' ues were determined using the bounding set of operating parameters and instrument uncertainties associated with either Resistance Temperature

                    ' Detector (RTD) bypass manifolds installed or with RTD bypass manifolds eliminated.

C. Technical Specification Bases 2.1.1 and 7.2.l'- Changes in the bases reflect the use of the revised methods ana correlations and the RIDP. The VANTAGE-5' fuel is analyzed using the WRB-2 DNB correlation with design limit DNBR values of 1.24 and 1.23 for the typical and thimble , cells, respectively. The LOPAR fuel is analyzed using the WRB-1

                    - correlation with design limit DNBR values of 1.25 and 1.24 for the typical and thimble cells, respectively.

D. Technical Specification 3/4.2.5 and Table 3.2-1 The application of revised DNB methods and correlations and the Farley-specific uncertainties in plant operating parameters obtained with the RTDP methodology result in the revision of DNB-related parameters in Technical Specification 3/4.2.5. Therefore, changes are proposed for O limiting values of Reactor-Coolant System Tava, Pressurizer Pressure, and Reactor Coolant System Flow, including allowances for indication uncertainties.

l l (3 t/ AGachment 2 Page 6 E. Technical Specification Bases 3/4.2,3/4.2.2,3/4.2.3,3/4.2.5,and 3/4.4.1 - These bases sections were revised to incorporate changes in l the DNB design basis as a result of the new DNB correlations and RIDP ! methodology use. The above proposed changes will provide additional design flexibility. Specifically, the proposed Technical Specifications will accommodate higher design peaking factors (ftH and fg), fuel rod bow, thimble plug deletion, and transition core DNBR penalty The transients that have DNB as a limiting design basis criterton and those that cssume reactor trips on Overtemperature AT and Overpower AT werc reanalyzed assuming a VANTAGE-5 fuel transition. The safety analyses assumed the transition DNB cffects from LOPAR to a full core of VANTAGE-5 fuel. The VANTAGE-5 fuel, which includes the IfM grid design feature, was generically approved by the NRC following review of the Westinghouse WCAP-10444-P-A. The safety analyses also utilized the NRC-approved RTDP metnodology (WCAP-ll397-P A), the WRB-1 DNB correlation (WCAP 8762-P-A), the WRB-2 DNB correlation (WCAP-10444-P-A), and the improved THINC-IV model (WCAP-12330 P). Both the WRB-1 and WRB 2 DNB correlations have a DNBR n limit of 1.17. However, use of the farley RIDP instrument uncertainty calculations (WCAP-12769 and 'dCAP 12771) resulted in the farley-specific (V) DNBR design limits presented in item C above. With the RTDP methodology, uncertainties in plant instrumentation that monitors operating parameters, nuclear and thermal parameters, fuel fabrication parameters, computer codes, and DNB correlation predictions are considered statistically to obtain DNBR uncertainty factors. Based on the DNBR uncertainty factors, RTDP design limit DNBR values were determined such that there is at least a 95% probability, at a 95% confidence level, that DNB will not occur on the most limiting fuel rad during rormal operation and operational transients (Condition 1) and during transient conditions arising from faults of moderate frequency (Conditior- 11 events). The uncertainties in the plant instrumentation that monitors operating parameters (pressurizer pressure, primary coolant temperature, reactor power, and reactor coolant system flow) were evaluated for farley assuming two different primary coolant temperature measurement configurations. One configuration is the primary coolant temperature measurement with R W located in bypass manifolds. The other configuration is with the rid instrumentation relocated directly in primary loop thermowells with the bypass manifold removed. In the DNBR analyses using the RTDP methodology, a set of plant operating parameter uncertainties was used which is sounding for operation with either configuration. Likewise, the Technical Specifications DNB limits in 3/4.2.5 were determined to be valid for either plant configuration, The safety analyses 6nd results support the proposed Technical () c Specifications changes and show that the DNB design criterion is met.

A tachment 3 l Based on the information_ presented above and the analyses performed for the VANTAGE-5 fuel transition, the following conclusions can be reached with respect to 10 CFR 50.92:

1. The proposed safety limits, reactor trip setpoints, and DNB-related l parameters Technical Specifications changes do not increase the l probability or consequences of an accident previously evaluated in the FSAR. The core safety limits, trip setpoints, and DNB parameters were determined using NRC-reviewed and approved DNB methodologies, namely RTDP, improved THINC-IV model, and the WRB-1 and WRB-2 DNB correlations. No new performance requirements are being imposea on any system or componcnt in order to support the revised DNBR analysis assumptions. Overall plant integrity is not reduced. The DN8 sensitive transients were reanalyzed. The DNB design criterion continues to be met. None of these changes directly initiate an accident;.therefore, the probability of a accident has not increased.

The acceptance criteria for the reanalyzeo analyses with these revised DNB parameters continue to be met; therefore, the consequences of accidents previously evaluated in the FSAR are not significantly _ changed.- All dose consequences have been evaluated for these changes and all acceptance limits continue to be met.

2. The proposed safety limits, reactor trip setpoints, and DNB-related parameters Technical Specifications changes do not create the possibility of a new or different kind of accident than.any accident already- evaluated in.the FSAR. No new accident scenarios, failure mechanisms, or limiting single failures are introduced as a result of the proposed changes. The proposed Technical Specifications changes have no adverse effects on any_ safety-related system and do not challenge-the performance or integrity of any safety-related system.

The DNB design criterion continues to be met. Therefore, the possibility of a new or different kind of accident is not created.

3. The ;roposed Technical Specifications changes do not involve a significant reduction in a margin of safety. The change in the DNBR design-limits are associated with the use of NRC-approved methodologies (RTDP and WRB-1 and WRB 2 DNB correlations) and the improved THINC-IV model. In addition, the VANTAGE-5 fuel design, including IFM grids, uses the WRB-2 correlation and has been generically approved by the NRC. The DNB design criterion remains unchanged, even with the changes in DNBR design limit values. Therefore, the new DNBR design limit values associated with the DNB methodology and correlation changes, upon which the Technical Specifications changes are based, do not result in a significant reduction in the margin of safety because the DNB design criterion continues to be met.

Based upon the preceding information, it has been determined that these proposed changes to the Technical Specifications for core limits, Reactor O' Trotection System setpoints', and DNB parameters do not involve a significant-hazards consideration as defined by 10 CFR 50.92 (c).

1 l f A Attachment 3 l Page 8 l () 1 l Reactor Trio System Instrumentation Resoonse Times 1 Modification to Technical Specifica* e.is Table 3.3-2 (Reactor Trip System Instrumentation Response Times) for item 4 (power range, neutron flux, high negative flux rate) is proposed. The proposed change involves a modification from the current value of .5 seconds to a new entry of "Not Applicable " This modification will make Technical Soecifications response times for the power range, neutron flux, high negative flux rate Reactor Protection System function consistent with other ESF and reactor trip functions that are not used as primary protection functions. The high negative flux rate reactor trip function is not taken credit for in any accident analysis, including the rod control cluster assembly (RCCA) transients (one or more dropped rods). Based on the NRC-approved methodology of WCAP-ll394-P-A, " Methodology for the Analysis of the Dropped Rod Event," no credit is taken for any direct reactor trip due to dropped RCCA(s) or for automatic power reduction due to dropped RCCA(s). This methodology has been employed in the Joseph M. Farley Nuclear Plant VANTAGE-5 fuel analysis. Farley-specific analyses have demonstrated that the DNB acceptance criteria are met for all DNB transients / events, including one or more dropped rods, without taking credit for the negative flux rate trip as the primary protection function. As such, this trip is considered to be a diverse reactor protection system (RPS) feature. p/ U Therefore, the change in response time from .5 seconds to "llot Applicable" will have no effect on any of the previously analyzed accidents. Based on the information presented above, the following conclusions can ce reached with respect to 10 CFR 50.92:

1. The deletion of response time for reactot trip on high negative flux rate does not significantly increase the probability or consequences of an accident previously evaluated in the FSAR. This function provides ro primary protection for any transient in the FSAR. The power range high negative flux rate reactor trip is considered to be a diverse RPS feature. No new performance requirements are being imposed on any system or component. Consequently, overall plant integrity is not reduced. This change has no effect on any dose calculations.

Therefore, the probability or consequences of an accident will not increase.

2. The reactor trip response time change to "Not Applicable" for high negative flux rate does not create the possibility of a new or different kind of accident from any previously evaluated in the FSAR.

This response time is not an initiator for any transient. No new accident scenarios, failure mechanisms, or limiting single failures are introduced as a result of this modification. The response time change does not challenge or prevent the performance of any safety-related system during plant transients. Therefore, the possibility of a new or f~\ /; different kind of accident is not created. %)

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

Attachment 3'

 -(,,

p) Page 9

3. This_ change does not -involve a significant reduction in the margin of safety. All. primary trip functions and ESE actuations are unaffected by the change in this response time. In addition, the other Technical Specification surveillance testing requirements (e.g., channel calibration) associated with the-power range high negative flux rate V9- ~ reactor trip-are not affected by this' change. Therefore, the deletion of the response time does not effect the results of any accident analysis, and the margin of safety is maintained and not significantly reduced.

Based upon the preceding information, it has been detennined that the-response time change for high negative flux rate reactor trip to "Not Applicable" does not involve a significant hazards consideration as defined in 10 CFR 50.92(c). O V

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          ,                        .           - , -  . , , . , .            m , ..m       .. ,,, -   -~            , - - r  -y v , c    y-

/O- Attachment 3 d Page 10 Increase in Shutdown and Control Rod Droo Time Technical Specification 3/4.1.3.4 specifies the allowable shutdown and control rod drop time. This time is being changed from 2.2 seconds to 2.7 seconds. The increase is to account for a slightly higher pressure drop across the VANTAGE-5 fuel assembly due to the IFM grids and for the slightly smaller guide thimble diameter. The revised rod drop time was used for accident and transient reanalyses and evaluations. The results of these analyses and evaluations are that all the safety criteria and previously defined acceptance limits contir,ue to be met. The 0.5 second increase in rod drop time allows for the slight increase in rod dron time expected from VANTAGE-5 fuel. The required verification (surveillance test) of rod drop time will still be performed. The revised Technical Specifications limit is consistent with the value used for the accident and transient analyses. Based on the results of those analyses, the following conclusions can be reached regarding 10 CFR 50.92:

1. The increase in rod drop time will not result in a significant increase in the probability or consequences of any accident previously evaluated in the FSAR, since the increased rod drop time has been accounted for

,. s        in all accident analyses and the same surveillance requirements will- be

( ) used to detect inoperable rods and increased rod drop times. The V effects of increased rod drop time on all applicable analyses and dose calculations have been evaluated or analyzed, and all the acceptance limits continue to be met.

2. The possibility of a new or different type of accident is not involved because the increase in rod drop time used in the analyses and in the proposed Technical Specifications is consistent with the design of the VANTAGE-5 fuel and does not indicate any new or different failure mechanism.

3. The effects of the increased rod drop time have been included in all applicable analyses and evaluations of accidents and transients. These analyses demonstrated that the plant will remain within previously accepted limits; therefore, the increase in the allowable rod drop time does not result in a significant reduct:on in a margin of safety.

Based upon the preceding information, it has been determined that the revision to the allowabta rod drop time does not involve a significant hazards consideration as defined in 10 CFR 50.92 (c). O

i- Attachment 3

#     Page 11 Increased Positive Moderator Temperature Coefficient (PMTC)

A revision to Technical Specification 3.1.1.3 has been made to allow greater design flexibility. The moderator temperature coefficient limit was changed to 0.7x10-4 Ak/k/of up to 70% rated thermal power with a ramp from 0.7x10-9 Ak/k/0F to 0.0 x 10-4 Ak/k/0F from 70% to 100% rated thermal power. Safety analyses which are sensitive to the change have been reanaly..ed or re.-evaluated using conservative values of moderator temperature coefficient, which are determined on an event specific basis, to bound the use of increased PHTC. All analyses and evaluations yielded acceptable results. Based on new evaluations and analyses, the following conclusions can be reached with respect to 10 CFR 50.92:

1. The proposed Technical Specifications changes with respect to increased moderator temperature coefficient do not involve a significant increase in the probability or consequences of an accident previously evaluated in the Farley FSAR. The mechanical design changes associated with VANTAGE-5 fuel result in the capability for relaxation of analytical s input parameters, such that increased margin can be generated without

    )      violation of any acceptance criteria. This margin can then t'e applied towards relaxation of operational limits such as moderator temperature coefficient. Ir all cases, the appropriate design and acceptance cr"eria are met      No new performance requirements are being imposed on any sy? p'n r temponent in order to support the revised analysis assumptions. Subsequently, overall plant integrity is not reduced.

Therefore, the probability of an accident has not significantly increased. The radiological consequences of an accident previously evaluated in the FSAR are not increased due to the proposed increase in PMTC. Evaluations have confirmed that the doses remain within previously approved acceptable limits, as well as those defined by 10 CFR 100. Therefore, the radiologicai consequences to the public resulting from any accident previously evaluated in the FSAR has not significantly increased.

2. The Technical Specifications changes with respect to increased moderator temperature coefficient do not create the possibility of a new or different kind of accident from any previously evaluated in the FSAR. No new accident scenarios, failure mechanism, or limiting single failures are introduced as a result of the fuel transition. Neither the presence of VANTAGE-5 fuel assemblies in the core nor the revised analytical assumptions, including increased PMTC, create new challenges to the performance of any safety-related system. Therefore, the r~ possibility of a new or different kind of accident is not created.

Atigshment 3 O Fage 12

3. The Technical Specifications related to PMTC changes do not involve a significant reduction in the margin of safety. The margin of safety for fuel-related parameters, including increased PMTC, associated with the VANTAGE-5 transition are defined in the Bases of the Technical Specifications. These Bases and the supporting Technical Specifications values are defined by the accident analyses which are performed to conservatively bound the operating conditions defined by the Technical Specifications and to demonstrate meeting the regulatory acceptance limits. Performance of analyses and evaluations for the VANTAGE-5 fuel transition with increased PMTC has confirmed that the operating envelope defined by the Technical Specifications will be bounded by the revised analytical basis, which in no c.ase exceeds the acceptance limits. Therefore, the margin of safety provided by the analyses in accordance with these acceptance limits is maintained and is not significantly reduced.

Based upon the preceding information, it has been determined that the proposed increased PMTC does not involve a significant hazards consideration as defined in 10 CFR 50.92 (c). O O

m Attachment 3 Page 13 -- () Jncreased FgdAp. and Modification of Kf 7) Curve Proposed changes to Technical Specifications 3/4.2.2 and 3/4.2.3 include separate FO's (Heat Flux Hot Channel Factor) and FAH's (Nuclear Enthalpy Hot ChanneT Factor) for the VANTAGE-5 and LOPAR fuels. The current limits for LOPAR fuel have been retained. The proposed changes are: LOPAR FQ (Z) s 2.J1 (K(7)) for P > .5 P s 4.64 (K(Z)) for P s .5 FEg i 1.55 (1+.3(1-P)) VANTAGE-5 FQ (Z) s L_4J (K(Z)) for P > .5 P 1 4.9 (K(Z)) for P s .5 ( FHH s 1.65 (1 + 3(1-P)) . In addition, Technical Specifications Figure 3.2-2 (Normalized FQ(Z) versus Core Height) will be changed to eliminate the K(Z) third line segment and to modify the second line segment. These modifications are supported by accident analyses results. These include large-break LOCA, small-break LOCA, and all affected non-LOCA accidents. All acceptance criteria continue to be met. Bas 3d on the information presented above, the following conclusions can be rt: ached with respect to 10 CFR 50.92:

1. The FQ , FAH, and K(Z) curve Technical Specifications changes do not involve a significant increase in the probability or consequences of an accident previously evaluated in the Farley FSAR. The mechanical design changes associated with VANTAGE-5 fuel and the improved methodologies result in the capability for relaxation of analytical input parameters, such that increased margin can be generated without violation of any acceptance criteria. This margin can then be applied towards relaxation of operational limits such as higher F (Z) Q and FaH for the VANTAGE-5 assemblies, in each case, however, the appropriate design and acceptance criteria are met. No new performance K ~ requirements are being imposed on any system or component in order to

Attachment 3 , 1O Page-14 support the revised analysis assumptions. Subsequently, overall plant , integrity is not reduced. Furthermore, the parameter changes are associated with features _used as limits or mitigators to assumed accident scenarios and are not accident initiators. Therefore, the probability of'an accident has not significantly increased. The radiological consequences of.an accident previously evaluated in the FSAR are not' increased due to these Technical Specifications changes. Evaluations have confirmed that the doses remain within previously approved acceptable limits as well as those defined by 10 CFR 100. Therefore, M radiological consequences to_ the public resulting from any '.ccident previously evaluated in the FSAR have not , significantly increased. p -2. The-FQ , FAH, and K(Z) curve Technical Specifications changes do not create the possibility of a new or different kind of accident from any previously evaluated in the FSAR. No new accident scenarios, failure mechanisms, or limiting single failures are introduced as a result of the fuel transition.- The presence of VANTAGE-5 fuel assemblies in_the core or_ the revised analytical assumptions have no adverse effect and do not challenge _the performance of any other safety-related system. A ' Therefore, the possibility of a new or different kind of accident is i (/ ' bot created.

3. The FQ , FAH, _ and K(Z) curve Technical Specifications changes -do not -

involve a significant reduction in the margin of safety. The margin of safety for fuel-related parameters associated with the VANTAGE-5 transition-are defined in the bases to the Technical Specifications. These' bases and the supporting Technical Specifications values are defined by'the. accident analyses which are performed to conservatively

bound the operating conditions defined by the Technical Specifications

'- and to demonstrate that the regulatory acceptance limits are met. F Performance of analyses and evaluations for the proporad inclusion of . separate-FQ's and FAH's for VANTAGE-5 and LOPAR fuel types and j modification of the K(Z) curve have confirmed that the operating H envelope defined by the Technical _ Specifications continues to be l bounded by the revised analytical basis, which in-no case exceeds the h acceptance limits. Therefore, the margin of safety provided by the analyses in- accordance with these acceptance limits is maintained and is not significantly reduced. Based upon the preceding information, it has been determined that the

proposed changes to the Technical Specifications of increased FQ 's and

     'FAH's for VANTAGE-5 fuel and modification of the K(Z) curve do not involve a significant hazards consideration as defined in 10 CFR 50.92 (c).

IO l _ . ..

Attachment 3 pd Page 15 Enoineered Safety Features (ESF) Allowable Values Some changes to the ESF allowable values in Technical Specifications Table 3.3-4 are required. These changes incorporate the effects of Farley-specific uncertainties associated with plant instrumentation, procedures, and measurement techniques. Adjustment of the allowable values was required to ensure that the allowable values were not overly conservat-ive with respect to the associated setpoints. These slight adjustments in the allowable values will not affect any accident analyses since the ESF setpoints are unchanged. The following allowable values were changed: Containment Pressure High-High-High (containment spray and phase "B" isolation) and Containment Pressure High-High (steamline isolation). Based on the inforr:ation presented above, the following conclusions can be reached with respect to .0 CFR 50.92:

1. The ESF allowable ve f ue changes do not significantly increase the probability or conse suences of an accident previously evaluated in the FSAR. These allowable values are not input parameters to any trans ent in the FSAR. No new performance requirements are being imposed on .ny system or component. Consequently, overall plant integrity is not reduced. These changes have no effect on any dose calculations.

]C Therefore, the probability or consequences of an accident will not increase.

2. The ESF allowable value changes do not create the possibility of a new or different kind of accident from any previously evaluated in the FSAR. These values are not an initiator for any transient. No new accident scenarios, failure mechanisms, or limiting single failures are introduced as a result of these changes. The allowable value changes do not challenge or prevent the performance of. any safety-related system during plant transients. Therefore, the possibility of a new or different kind of accident is not created.

3. These changes do not involve a significant reduction in the margin of safety. All trip setpoints associated with these changes have been preserved. Therefore, the small changes to the allowable values do not effect the results of any accident analysis. Therefore, the margin of safety is maintained and not significantly reduced.

Based upon the preceding information, it has been determined that the ESF allowable value changes do not involve a significant hazards consideration as defined by 10 CFR 50.92 (c).

Attachment 3 !n) Page 16 Enoineered Safety Features (ESF) Restonse Timts Modification to Technical Specifications Table 3.3-5 (ESF Response Times) for item 5a (steam line isolation on high steam flow coincident with low-low Tavg) is proposed. The proposed change involves a modification from the current Unit I value of 11 seconds and the Unit 2 value of 9 seconds to a new entry of "Not Applicable." The modification will make the Technical Specifications resr:ase time for the high steam flow coincident with low-low Tavg ESF function consistent with other ESF and reactor trip functions that are not used as primary protective functions. The high steam flow coincident with low low Tavg ESF function is not taken credit for in any accident analysis, including main steam pipe rupture, non-LOCA, containment response, or in equipment qualification (superheat) outside of containment. Protection for these events is provided by other protection signals. Main steam line isolation on high steam flow in two steam lines coincident with low-low Tavg is provided as a diverse signal that provides no primary protection for cny event. Protection for main steam pipe breaks is provided by the overpower protection, overtemperature delta-T, and low pressurizer pressure reactor trip functions, and the low steam line pressure, high steam line differential pressure, low pressurizer pressure, and the High-1 containment pressure ESF safety injection functions. (" - Primary main steam line isolation protection is provided by the low steam ( line pressure and the High-2 containment pressure ESF functions. Therefore, the change in the Unit I and Unit 2 response times from 11 seconds and 9 seconds, espectively, to "Not Applicable" will have no effect on any of the previously analyzed accidents. Based on the information presented above, the following conclusions can be reached with respect to 10 CFR 50.92:

1. The ESF response time change for this steam line isolation function does not significantly increase the probability or consequences ci an accident previously evaluated in the FSAR. This function provides no primary protection for any transient in the FSAR. No new performance requirements are being imposed on any system or component.

Consequently, overall plant integrity is not reduced and dose calculations are not affected. Therefore, the probability or consequences of an accident will not increase.

2. The ESF respon:e time change to "Not Applicable" for the high steam flow in coincidence with low-low Tavg function does not create the possibility of a new or different kind of accident from any previously evaluated in the FSAR. This response time is not an initiator for any transient. No new accident scenarios, failure mechanisms, or limiting single failures are introduced as a result of this modification. The response time change does not challenge or prevent the performance of any safety-related system during plant transients. Therefore, the f, possibility of a new or different kind of accident is not created.

l

p-Attachment 3 ( Page_17 ! 3. This change does not involve a significant reduction in the margin of i safety. All primary reactor trip functions and ESF actuations are unaffected by the change in this ESF response time. In addition, the other Technical Specification surveillance testing requirements (e.g., periodic channel checks and calibrations) associated with high steam flow coincident with low-low Tavg are not affected by this change; therefore, performance cf these surveillance tests will continue to demonstrate operability. Therefore, the change to the response time does not effect the results of any accident analysis, and the margin of safety is maintained and not significantly reduced. Based upon the preceding information, it has been determined that the ESF response time change to "Not Applicable" for steam line isolation on high steam flow in two steam lines coincident with Tavg low low does not involve L a significant hazards consideration as defined in 10 CFR 50.92(c). ( LJ l i l l l l l l b G

,w Attachment 3 Page 18 (L,) Reactor Coolant locos and Coolant Circulation (Mode 31 A revision to Technical Specification 3.4.1.2 and the associated Bases'is included to allow greater operational flexibility. The revision requires that only two reactor coolant pumps (RCPs) and their associated loops be operable in Mode 3. Safety analyses, including bank withdrawal accidents which are sensitive to this change, have been reanalyzed or re-evaluated. All analyses yielded acceptable results and all applicable acceptance criteria continue to be met. Based on the information presented above the following conclusions can be reached with respect to 10 CFR 50.92:

1. The Technical Specifications change with respect to Mode 3 operation requiring only two RCPs operable does not involve a significant increase in the probability or consequences of an accident previously evaluated in the Farley FSAR. Capability for relaxation of analytical input parameters has been demonstrated by the acceptable analytical results without violation of any acceptance criteria. This can allow a reduction in the number of RCPs required for Mode 3. In all cases, the appropriate design and acceptance criteria are met. No new performance requirements are being imposed on any system or component in order to o)

(" support the revised analysis assumptions. Subsequently, overall plant integrity is not reduced. Therefore, the probability of an accident has not significantly increased. The radiological consequences of an accident previously evaluated in the FSAR are not increased due to this proposed Technical Specifications change. Evaluations have confirmed that the doses remain within previously approved, acceptable limits as well as those defined by 10 CFR 100. Therefore, the radiological consequences to the public resulting from any accident previously evaluated in the FSAR has not significantly increased.

2. The Technical Specifications changes with respect to requiring only two RCPs in Mode 3 does not create the possibility of a new or different kind of accident from any previously evaluated in the FSAR. No new accident scenarios, failure mechanism, or limiting single failures are introduced as a result of this change. No new challenges to safety systems have been identified because of this proposed change.

Therefore, the possibility of a new or different kind of accident is not created.

3. The proposed Technical Specifications change related to requiring only two RCPs in Mode 5 does not involve a significant reduction in the margin of safety. The margin of safety associated with this change is defined in the bases to the Technical Specifications. These bases and O the supporting Technical Specifications are defined by the accident d analyses which are performed to conservatively bound the operating

                                                         .-           =.

g- Attachment 3 ( Page 19 x conditions defined by the Technical Specifications and to demonstrate meeting the regulatory acceptance limits. Performance of analyses and evaluations have confirmed that the operating envelope defined by the Technical Specifications continues to be bounded by the revised analytical basis, which in c.o case exceeds the acceptance limits. Therefore, the margin of safety provided by the analyses in accordance with these acceptance limits is maintained and is not significantly reduced. Based upon the preceding information, it has been determined that the proposed changes to the Technical Specifications requiring only two RCPs available in Mode 3 do not involve a significant hazards consideration as defined in 10 CFR 50.92 (c).

v i

l i -

'O 'C/ Attachment 3 Page 20 Administrative Chance for Radial Peakina Factor LimitDortina The current reporting requirement for radial peaking factor limit reporting in Section 6.9.1.11 of the Farley Technical Specifications is 60 days prior to criticality. To allow for more flexibility after shutdown, in particular for emergency core redesigns, this reporting requirement has been revised to 30 days after criticality. There are no safety analysis or design criteria that are sensitive to this administrative change, and the same reporting requirements will be continued to be met for the 30-day after criticality reporting change. Based on tha information presented above, the following conclusions can be reached with respect to 10 CFR 50.92:

1. Since the reporting requirements are an administrative change, this change does not involve a significant increase in the probability or consequences of an accident previously evaluated in the Farley FSAR.

With the exception of the reporting schedule, no other changes to the reporting requirements are made. )

                                                                                -     l

2. This administrative change in reporting requirements does not create

.O the possibility of a new or different kind of accident from any V previously evaluated in the Farley FSAR, since only the time for , reporting requirements has changed. Therefore, the possibility of a i new or different kind of accident is not created.

3. This administrative change is not safety-related and, therefore, does not involve any reduction in the margin of safety.

Based on the preceding information, it has been determined that the

n. proposed change in reporting times for the radial peaking factors does net involve a significant hazards consideration as defined in 10 CFR 50.92 (c).

Conclusion The package of Technical Specifications changes related to the transition to VANTAGE-5 fuel have been shown to not involve a significant hazard consideration, and all requirements of 10 CFR 50.92 (c) are met. un

fl Attachment 3 () Page 21 Significant Hazards Consideration l for the Joseph M. Farley Nuclear Plant Unit 2 ; Resistance Temperature Detector Bypass Removal l 10 CFR 50.92 Evaluation Pursuant to the requirements in 10 CFR 50.92, each application for , amendment to an operating license must be reviewed to determine if the modification involves a significant hazard. The proposed amendment stpports removal and replacement of the existing Resistance Temperature Detector (RTD) bypass manifold temperature measurement system in Farley Nuclear Plant Unit 2 with fast response RTDs located in the reactor coolant piping. Proposed Chances The proposed amendment involves the following Technical Specifications changes similar to those previously implemented for Unit 1. Allowable values for loss of flow and Tavg low-low (for P-12 (increasing and decreasing) _ as well as engineering safeguards actuation on coincident high steam flow and low-low Tavg) have been modified to include specific margins _(s'u) gained by tne setpoint methodology calculations. Additional dynamic compensations were added to the OPAT and OTAT equations to more fully describe the as-ir, stalled hardware. Although these dynamic functions are set to have no dynamic effects, they are included to provide complete compatibility with the accident analyses. The OTAT reactor trip response time was increased to account for the new hardware configuration (i.e., RTD l element and well versus RTD element only). The OPAT and OTAT Bases were modified to account for the RTD bypass manifold elimination. The bases for i the amendment as described in WCAP-12614, Revision 2, "RTD Bypass Elimination Licensing Report for J. M. Farley Nuclear Plant Units 1 and 2," have been reviewed and deemed not to involve a significant hazard based on the following evaluation. l Backaround i The proposed amendment involves removing and replacing the existing RTD bypass manifold system with fast response RTDs located in the Unit 2 reactor coolant hot leg and cold leg piping. The original RTD bypass , _ system utilized an arrangement which directs a sample of the RCS flow from l the main coolant piping to an independent temperature measurement manifold. l With the proposed system, the hot leg temperature measurement on each loop l will be accomplished with three fast response, narrow range, dual element l RTDs mounted in thermowells. To accomplish the sampling function of the O RTD bypass manifold system and minimize the need for additional tot leg ( piping penetrations, the thermowells will be located within the three existing RTD bypass manifold scoops wherever possible. If plant

O Alt _achment 3 () Page 22 interferences preclude the placement of a thermowell in a scoop, then the scoop will be ct.pped and a new penetration made to accommodate the thermowell. These three RTDs accomplish a sampling function to meas are the average hot leg temperature which is used to calculate the reactor c)olant loop differential temperature (AT) and average temperature (Tavg). Cold leg temperature is taken by means of one fast response, narrow range, dual-element RTD located in each cold leg at the discharge of the reactor coolant pump (RCP). This RTD will replace the cold leg RTDs located in the bypass manifold. This RTD will measure the cold leg temperature which is used to calculate reactor coolant loop AT and Tavg. To assure that the functional capability and inherent reliability provided in the original Reactor Protection System design are not compromised, a Median Signal Felector (MSS) is implemented in the Reactor Control System for receiving eactor coolant temperature information. The function of the signal selectu is to elimir. ate the potential for a control and protection system interaction mechanism involving the Reactor Control System and the thermal overpower and overtemperature protective functions in accordance with the requirements of the Institute of Electrical and Electronics Engineers Standard, IEEE Standard 279-1971, " Criteria for Protection Systems for Nuclear Power Generating Stations," Section 4.7. Installation of a fast response system which measures loop temperature via thermowell-mounted RTDs protruding into the main reactor coolant flow will eliminate the bypass piping network and operating obstacles associated with the bypass system (such as leakage through valves, flanges, etc., and radiation exposure during maintenance). Analysis Conformance of the proposed amendments to the standards for a determination of no significant hazard as defined in 10 CFR 50.92 is shown in the following:

1. The use of fast response RTDs does not involve a significant increase in the probability or consequences of any accident previously evaluated. The non LOCA and LOCA accidents were reviewed in WCAP-12614, Rev. 2, verifying that the variations in uncertainty associated with certain reactor trip functions were acceptable, it is concluded that an increase in RCS temperature uncertainty can be accommodated by margins in the safety analyses to acceptance criteria limits and allocation of generic DNB margin. Evaluation of the modification to the Reactor Coolant System boundary has also been performed, and no degradation in integrity is involved. In addition, the recent analyses and evaluations coaducted to support the transition (q

 'j          to VANTAGE-5 fuel at Farley were performed with a bounding set of plant operating parameter uncertainties that bound plant operation with or

    - --         -.           --        -   -        -   --.       .   .     - .= -      -   ~.

. Attachment 3

, Page 23

             .without the RTD bypass manifold temperature system. Therefore, the L               potential effects of RTD Bypass Elimination (RTDBE)-on plant operatM' with VANTAGE-5 fuel have been considered and found not to involve a significant increase in the probability or consequenc;;s of any accident previously evaluated.

2. The use of fast response RTDs does not create the possibility of a new
or different kind of accident from any accident previously evaluated.

L The three-dual-element hot leg RTDs and one dual-element cold leg RTD L will' utilize-the existing penetrations, whenever possible, into the RCS piping from the bypass system with only slight modifications. Caps and-welds sealing the crossover leg bypass return nozzle and piping, as well as the modification and welding for-the existing penetrations, will be qualified in accordance. with the ASME code,- thus precluding the

              . possibility. for a new or different kind of accident.

The function of the AT/Tavg protection channels is not changed because of the bypass elimination. The newly installed fast response RTDs perform-the'same function in both Thot and Tcold applications. The three Th ot signals are electronically averaged. Dual-element RTDs are L installed in the hot and <:old legs. Should one RTD element fail, the L spare element can be connected. In addition, the average Thot signal can be electronically _ biased to a two RTD average should one dual RTD fail. ~The measured temperature values will still serve as input to two-out-of-three voting logic for protection functions. The Median (- - Signal Selector (MSS) will eliminate the potential for control and , L protection interactions for all AT/Tavg applications. The basi:, for the instrumentation and control design meets the criteria of applicable IEEE standards, regulatory guides, and general design criteria, in that such principals as electrical separation, seismic and environmental ! qualification, and single failure criteria are satisfied. Therefore, I there is no possibility of a new or different kind of accident as a-result of the instrumentation aspects of RTD Bypass Elimination. I 3. The effect of RTD Bypass Elimination,-response time, setp61nt - uncertainty, temperature, and flow measurement ~ uncertainty does not involve a significant reduction in a margin of safety. These changes have been evaluated and compared to the acceptance limits with respect to the fuel, RCS pressure boundary, and containment. 'All acceptance limits continue to be met. The evaluation of the effect of these variables on non-LOCA and LOCA transients has verified that plant operation will be maintained within the bounds of safe, analyzed conditions as defined in the FSAR with the revised Technical Specifications, and the conclusions presented in the FSAR remain valid. As such, there is no significant reduction in the margin of safety for operation of J. M. Farley Nuclear Plant Unit 2 with RTD Bypass ! Elimination; i

3- -- I l l-l.. p AttachEent 3

     !    Page 24 l                                                                                      !

[pnclusion As discussed in WCAP-12614, Revision 2, a comprehensive evaluation of the

effects of RTD Bypass Elimination (RTDBE) has been completed, and no l adverse safety implications have been identified. Based on the preceding l information,-it has been determined that the proposed changes to the Unit 2 1 Tei..'ical Specifications for RTD Bypass Elimination do not involve a significant hazard consideration as defined in 10 CFR 50.92 (c). G O

O , Attachment 4 1 1 Joseph M. Farley Nuclear Plant Units 1 and 2 l Request for Technical Specifications Changes ! Safety Assessment O . O _

r TABLE OF CONTENTS I lection lLtl2 EASA

1.0 INTRODUCTION

AND CONCLUSIONS 2 7 2.0 MECHANICAL EVALUATION 3.0 NUCLEAR EVALUATION 18 4.0 THERMAL AND HYDRAULIC EVALUATION 20 5.0 ACCIDENT EVALUATION 26 6.0

SUMMARY

OF TECHNICAL SPECIFICATIONS CHANGES 49

7.0 REFERENCES

52 LIST OF TABLES Table No. Ittle Eine 2-1 Comparison of 17x17 LOPAR and 17x17 VANTAGE 5 Fuel Assembly- 9 Design Parameters 4-1 Thermal and Hydraulic Design Parameters for FNP Units 1 and 2 22 6-1 Summary and Justification for FNP Units 1 and 2 Technical 51 Specifications Changes for VANTAGE 5 Fuel LIST OF FIGURES Ficure No. Title P_iLqg 2.1 17x17 VANTAGE 5/LOPAR Fuel Assembly Comparison 11 O 4544F/687F910515:50 1

l

1.0 INTRODUCTION

AND CONCLUSIONS The hieph H. Farley Nuclear Plant (FNP) Units 1 and 2 are currently operating with a Westinghouse 17x17 low-parasitic (LOPAR) fueled core. For subsequent cycles, it is planned to refuel and operate the FNP Units 1 and 2 with the Westinghouse VANTAGE 5 improved fuel design. As a result, future core loadings would range from approximately 60% LOPAR, and 40% VANTAGE 5 transition cores to eventually an all VANTAGE 5 fueled core. The VANTAGE 5 fuel assembly is designed as a modification to the current 17x17 LOPAR (standard fuel) and the Optimized Fuel Assembly (OFA) designs, Reference 1. The VANTAGE 5 design features were conceptually packaged to be licensed as a single entity. This was accomplished via the NRC review and approval of the

    " VANTAGE 5 Fuel Assembly Reference Core Report," HCAP-10444-P-A, Reference 2.

The initial irradiation of a fuel region containing all the VANTAGE 5 design features occurred in the Callaway Plant in November 1987. The Callaway VANTAGE 5 licensing submittal was made to the NRC on March 31, 1987 (ULNRC-1470, Docket No. 50-483). NRC approval was received in October 1987. Several of the VANTAGE 5 design features, such as axial blankets. A V reconstitutable top nozzles, extended burnup modified fuel assemblies and integral fuel burnable absorbers have been successfully licensed as individual design features and are currently operating in Hestinghouse plants. The FNP Units 1 and 2 will be operating in reload Cycles 12 and 9, respectively, with VANTAGE 5 fuel containing the following features: Integral Fuel Burnable Absorbers (IFBAs), Intermediate Flow Mixers (IFM) grids Reconstitutable Top Nozzles (RTN), and fuel assemblies modified for extended burnup. The RTN and the fuel assemblies modified for extended burnup are currently operating in LOPAR fuel in both FNP Units 1 and 2. In addition, both the FNP Units 1 and 2 fuel assemblies are currently operating with the modified Debris Filter Bottom Nozzle (HDFBN) which will also be included with the VANTAGE 5 fuel assemblies. Starting with the first transition cycle (i.e., FNP Unit 1. Cycle 12 and FNP Unit 2, Cycle 9), axial blankets are optional. A brief summary of the VANTAGE 5 design features and the major advantages of the improved fuel drsign are given below. These features and figures illustrating the VANTAGE 5 design are presented in more detail in Section 2.0, 4544F/687F910513:50 2

Intearal Fuel Burnable Absorber (IFBA) - The IFBA features a thin boride

 ,Q  coating on the fuel pellet surface in the central core portion of the enriched kl  002 pellets. In a typical reload core, approximately thirty percent of the fuel rods in the feed region are expected to include IFBAs. IFBAs provide power peaking and moderator temperature coefficient control.

Intaunadiate Flqw_tilger_(IFM) Grid - Three IFH grids located between the four upper most zircaloy grids provide increased DNB margin. Increased margin permits an increase la the design basis F AH and Fg . Reconstitutable 100 Nozzle (RTN) - A mechanical disconnect feature facilitates the top nozzle removal. Changes in the design of both the top and bottom nozzles increase burnup margins by providing aaaitional plenum space and room for fuel rod growth. Extended Burnuo - The VANTAGE 5 fuel design will be capable of achieving extended burnups. The basis for designing to extended burnup is contained in the approved Hestinghouse topical WCAP-10125-P-A, Reference 3. The FNP has previously used this topical report in the design basis of the plant. (G3 Axial Blankets - The axial blanket consists of a nominal six inches of natural UO2 pellets at each end of the fuel stack to reduce neutron leakage axially and to improve uranium utilization. For VANTAGE 5 reload cores, low leakage loading patterns (burned radial blankets on_the core periphery) are shown to further improve uranium utilization and provide additional pressurized thermal shock margin. In addition to the above VANTAGE 5 design features, the transition design includes the'0FA design features, Reference 1, with the reduced fuel rod diameter and zircaloy mixing vane grids. These design features result in an improved water to uranium ratio and reduced parasitic neutron absorption. This submittal is to serve as a reference safety evaluation / analysis report for the region-by-region reload transition from the present FNP LOPAR fueled . core to an all VANTAGE 5 fueled core. The submittal examines the differences 4544F/687F910513:50 3

between the VANTAGE 5 and LOPAR fuel assembly designs and evaluates the effect ("} of these differences on the core performance during the transition to an all V VANTAGE 5 core. The VANTAGE 5 core evaluation / analyses were performed at a l core thermal power level of 2775 MHt (except for the large break LOCA, steam j generator tube rupture and containment mass and energy release analyses) for the FNP Units 1 and 2 with the following conservative assumptions made in the safety evaluations: a full power F AH of 1.65 for the VANTAGE 5 fuel and 1.55 for the LOPAR fuel, and 1.70 both for the VANTAGE 5 fuel non-LOCA analyses and the small break LOCA analysis, an increase in the maximum gF to 2.45 for VANTAGE 5 and 2.32 for LOPAR fuel, a peak 20% plant total steam generator tube plugging for both Units 1 and 2, and a core bypass flow of 7.1% l with thimble plugs removed. The analysis assumption of core bypass flow with thimble plugs removed is conservative for operation with thimble plugs and/or Het Annular Burnable Absorber (HABA) rods. Also, a Positive Moderator Temperature Coefficient (PHTC) of +7 pcm/ degree F from 0% to 70% power and decreasing linearly to O pcm/ degree F at 1001 power was used. The standard reload design methods described in Reference 4 will be used as a basic reference document in support of future FNP Units 1 and 2 Reload Safety

'  Evaluations (RSE) with VANTAGE 5 fuel reloads. Sections 2.0 through 5.0 summarize the Mechanical, Nuclear Thermal and Hydraulic, and Accident Analyses / Evaluations, respectively. Section 6.0 gives a summary of the technical specifications changes needed.

Consistent with the Hestinghouse standard reload methodology, Reference 4, parameters are chosen to maximize the applicability of the safety evaluations for future cycles. The objective of subsequent cycle specific RSEs will be to verify that applicable safety limits are satisfied based on the reference evaluation / analyses established in this submittal. In order to demonstrate early performance of the VANTAGE 5 design product features in a commercial reactor, four VANTAGE 5 demonstration assemblies (17x17) were loaded into the V. C. Summer Cycle 2 core and began power production in December of 1984. These assemblies completed one cycle of irradiation in October of 1985 with an average burnup of 11,357 MHD/HTU. } Post-irradiation examinations showed all 4 demonstration assemblies were of good mechanical integrity. No mechanical damage or wear was evident on any of 4544F/687F910513:50 4

l 1 the VANTAGE 5-components. Likewise, the IFH grids on the VANTAGE 5 ) A demonstration assemblies had no effect on the adjacent fuel assemblies. All V four demonstration assemblies were reinserted into V. C. Summer for a second cycle of irradiation. This cycle was completed in March of 1987, at which time the demonstration assemblies achieved an average burnup of about 30,000 I HWD/HTU. The observed behavior of the four assemblies at the end of 2 cycles of irradiation was as good as that observed at the end of the first cycle of irradiation. The four assemblies were reinserted for a third cycle of irradiation which was completed in November 1988 (E0C burnup 46,000 MHD/HTU). The observed behavior of the four assemblies was as good as that observed at the end of the first and second cycles of irradiation. In addition to V. C. Summer, individual VANTAGE 5 product features have been demonstrated at other nuclear plants. IFBA demonstration fuel rods have been irradiated in Turkey Point Units 3 and 4 for two reactor cycles. Unit 4 ; contained 112 fuel rods equally distributed in four demonstration assemblies, j The IFBA coating performed well with no loss of coating integrity or adherence. The IFH grid feature has been demonstrated at McGuire Unit 1. The (] demonstration assembly at McGuire was irradiated for three reactor cycles and \d showed good mechanical integrity. The following plants are currently operating with full regions of VANTAGE 5 fuel assemblies: Callaway, V. C. Summer, Shearon Harris, Diablo Canyon, i Byron /Braidwood and D. C. Cook. The results of the evaluation / analysis described herein lead to the following conclusions:

1. The Hestinghouse VANTAGE 5 reload fuel assemblies for the FNP Units 1 and 2 are mechanically compatible with the current LOPAR fuel assemblies, control rods, and reactor internals interfaces. The j VANTAGE 5/LOPAR fuel assemblies satisfy the current design bases for I the FNP Units 1 and 2.

I l

LJ : 1 4544F/687F910513:50 5 l

l

2. The structural integrity of the 17xl? VANTAGE 5 fuel assembly design for seismic /LOCA loadings has been evaluated for both FNP Units 1 and

2. Evaluation of the 17x17 VANTAGE 5 fuel assembly component stresses and grid impact forces due to postulated faulted condition accidents verified that the VANTAGE 5 fuel assembly design is structurally acceptable.

3. Changes in the nuclear characteristics due to the transition from LOPAR to VANTAGE 5 fuel will be typical of the normal cycle-to-cycle variations experienced as loading patterns change.

4. The reload VANTAGE 5 fuel assemblies are hydraulically compatible with the LOPAR fuel assemblies from previous cycles of operation.

5. The core design and safety analyses results documented in this report show the core's capability for operating safely for the FNP Units 1 and 2 thermal power of 2775 Hwt (2652 Het for the large break LOCA) ,

with F AH of 1.65 for the VANTAGE 5 fai and 1.55 for the LOPAR . fuel, F gof 2.45 for the VANTAGE 5 fuel, Fgof 2.32 for LOPAR fuel O and steam generator tube plugging levels up to a peak of 201 in each steam generator.

6. The implementation of VANTAGE 5 fuel at the FNP Units 1 and 2 does ne involve the addition of any design features that would affect the radiological source terms, and thus, the radiological consequences of accidents are not affected.

7. The previously reviewed licensing basis continues to be met when the FNP Units 1 and 2 are reloaded with VANTAGE 5 fuel. Plant operating limitations riven in the Technical Specifications will be satisfied with the prcposed changes noted in Attachment 2 of this submittal.

This report serves as a reference upon which to base Hestinghouse reload safety evaluations for future FNP reloads with VANTAGE 5 fuel. O 4544F/687F910513:50 6

1 2.0 MECHANICAL EVALUATION A b Introduction and Summat, This section evaluates the mechanical design and the compatibility of the 17x17 VANTAGE 5 fuel assembly with the current low-parasitic (LOPAR) fuel assemblies during the transition through mixed-fueled cores to an all VANTAGE 5 cort. The VANTAGE 5 fuel assembly has been designed to be compatible with the LOPAR fuel assemblies, reactor internals interfaces, fuel handling ! equipment, and refueling equipment. The VANTAGE 5 design is intended to replace and be compatible with cores containing fuel of the LOPAR design. The VANTAGE 5 design dimensions are essentially equivalent to the current FNP Units 1 and 2 LOPAR fuel assembly design from an exterior assembly envelope and reactor internals interface standpoint. References in this section are made to HCAP-10444-P-A, " VANTAGE 5 Fuel Assembly Reference Core Report," Reference 2, and to HCAP-9500-A, " Reference Core Report 17x17 Optimized Fuel l Assembly," Reference 1. The significant new mechanical features of the VANTAGE 5 design relative to the previous LOPAR fuel design for both FNP Units 1 and 2 include the l following: o Integral Fuel Burnable Absorber (IFBA) o Intermediate Flow Mixer (IFM) Grids ! o Reconstitutable Top Nozzle (RTN) l o Extended burnup capability including slightly longer fuel rods o Axial blankets o Replacement of six intermediate inconel grids with zircaloy grids o Reduction in fuel rod, guide thimble and instrumentation tube diameter l l 0 Redesigned fuel rod bottom end plug to facilitate reconstitution capability o Snag-resistant inconel grids (top and bottom) l o Modified Debris Filter Bottom Nozzle (HDFBN) l 4544F/687F910513:50 7

The RTN, MDFBN, redesigned fuel rod bottom end plug, snag-resistant grid, and ,f m the fuel assembly extended burnup modification have been introduced previously () in both FNP Units 1 and 2. These features will continue to be utilized in the VANTAGE 5 design. Table 2-1 provides a comparison of the LOPAR and VANTAGE 5

    . fuel assembly design parameters.

Fuel Rod Performance Fuel rod design evaluations for FNP Units 1 and 2 were performed using the NRC approved models, References 5 and 6, and the extended burnup design methods in Reference 3. Fuel rod performance for all FNP fuel is shown to satisfy the fuel rod design basis on a region by region basis. These same bases are applicable to all fuel roc designs, including the LOPAR and VANTAGE 5 fuel designs, with the only difference being that the VANTAGE 5 fuel is designed to operate with a higher F 3g limit. The design bases for Hestinghouse VANTAGE 5 fuel are discussed in Reference 2. There is no effect from a fuel rod design standpoint having fuel with more than one type of geometry simultaneously residing in the core during the (q) transition cycles. The mechanical fuel rod design evaluation for'each region incorporates all appropriate design teatures of the region, including any changes to the fuel rod or pellet geometry from that of previous fuel regions, for example, the presence of axial blankets or changes in the fuel rod and plenum length. Analysis of IFBA rods includes any geometry changes necessary to model the presence of the burnable absorber, and conservatively models the gas release from the zirconium diboride pellet coating. Fuel performance evaluations are completed for each fuel region to demonstrate that the design criteria will be satisfied for all fuel rod types in the core under the planned operating conditions. Any changes from the plant operating conditions originally evaluated for the mechanical design of a fuel region (for extmple an increase in the peaking factors) are addressed for all affected fuel regions as part of the reload safety evaluation process when the plant change is to be implemented. O V j l 4544F/687F910513:50 8

TABLE 2-1 p ' Comparison of 17x17 LOPAR -s'y and 17x17 VANTAGE 5 Fuel Assembly Design Parameters-l 17xl? 17x17 ; PARAMETER LOPAR DESIGN VANTAGE 5 DESIGN Fuel Assembly Length, in. 159.975 159.975 Fuel Rod Length, in. - 152.200 152.285 Fuel Assembly Height, lbs. 1467 1366 Assembly Envelope, in. 8.426 8.426 l Compatible with Core Internals Yes Yes Fuel Rod Pitch, in. 0.496 0.496 Number of fuel Rods / Assembly 264 264 Number / Guide Thimble Tubes / Assembly 24 24

,f g Number / Instrumentation Tube / Assembly          1                       1
\v)  Fuel Tube Material                                Zircaloy 4              Zircaloy 4 Fuel Rod Clad 00., in.                            0.374                   0.360 Fuel Rod Clad Thickness, in.                      0.0225                  0.0225 Fuel / Clad Gap. mil.                             3.25                    3.10 Fuel Pellet Diameter, in.                         0.3225                  0.3088*

Fuel Pellet-Length, in. 0.387 0.370

     *Does not include IFBA coating thickness.

m 4544F/687F910513:50 9

l

J Grid Assemblies 1 't b The top and bottom inconel (non-mixing vane) grids of the VANTAGE 5 fuel assemblies are similar in design to the inconel grids of the typical LOPAR l fuel assembly design used for FNP Units 1 and 2. The differences are: 1) the spring and dimple heights have been modified to accommodate the reduced , diameter fuel rod, 2) the top grid spring force has been reduced to minimize l rod bow, 3) the VANTAGE 5 top grid uses type 304L stainless steel sleeves l itstead of 304 stainless steel sleeves used for the LOPAR top grid, 4) the top and bottom grids have a snag-resistant design which minimizes assembly itteractions during core loading / unloading, 5) the top and bottom grids have ! dimples which are rotated 90* to minimize fuel rod fretting and dimple cocking, and 6) the top ar4d bottom grid heights have been increased to i 1.522 inches. The snag-resistant grid design was introduced in iNP Units 1, and 2 during Cycles 10 and 7, respectively. Rotated dimples and 1.522 inch grid heights were introduced in FNP Units 1 and 2 during Cycle 10 and 8, respectively. These features will continue to be utilized. The six intermediate (mixing vane) grids are made of Zircaloy-4 material rather than Inconel which is currently used in the LOPAR design. The IFH grids shown in Figure 2.1 are located in the three oppermost spans between the zircaloy mixing vane structural grids and incorporate a similar mixing vane array. Their prime function is mid-span flow mixing in the hottest fJel assembly spans. Each IFH grid cell contains four dimples which are designed to prevent mid-span channel closure in the spans containing IFHs and fuel rod contact with the mixing vanes. This simplified cell arrangement allows short grid cells so that the IFH grid can accomplish its flow mixing objective with minimal pressure drop. The IFH grids are not intended to be structural members. The IFH grids do, however, share the loads of the structural grids during faulted loading, and as such, contribute to enhance the load carrying capability of the VANTAGE 5 fuel assembly. The outer strap configuration of the IFH grid wts designed to be similar to current fuel designs to preclude grid hang-up and damage during fuel handli g. Additionally, the grid envelope is smaller which further [ minimizes the potential for damage and reduces calculated forces during seismic /LOCA events. 4544F/687F910513:50 10

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            . rNP urnt 1 - Region 13 cad Unit 2 - Region to oes.gn configuration
  "                                                                                                                            17x17 VANTAGE 5 / LOPAP JOSEPH W. FARLEY FUEL ASSEMBLY COMPARISON Alabama Power A                                                   :nsa; O       '< E 2.1 11

i a i For seismic and LOCA events, grid impact loads for both strue.tural and IFH j grids have been calculated for FNP Units 1 and 2. Impact testing has also

V. been performed to determine dynamic crush strength of each grid type. A comparison of calculated grid impact loads with'the grid strength indicates ) that a coolable geometry (i.e., structural integrity) will be maintained for j both structural and IFH grids subsequent to the combined seismic and LOCA l events. j

    .                                                                                  1 Reconstitutable Too Nozzle and Modified Debris Filter Bottom Nozzle            !
                                                                                       )

The RTN for the VANTAGE 5 fuel assembly differs from the typical LOPAR welded l top nozzle design in two ways: (a) a groove is provided in each thimble adapter plate guide tube-thru-hole to facilitate removal, and (b) the adapter plate thickness is reduced to provide additional axial space for fuel rod growth (extended burnup capability). The RTN feature was previously introduced in FNP Units 1 and 2 during Cycles 9 and 6 respectively. The RTN height was reduced for FNP Unit 1. Cycle 10 and for Unit 2. Cycle 7 to accommodate extended burnup capability. The RTN will continue to be utilized in FNP L11ts 1 and 2. Q) Q. Similar to the standard LOPAR RTN design, the VANTAGE 5 RTN design includes a stainless steel nozzle insert that is mechanically connected to the top nozzle adapter plate by means of a pre-formed circumferential bulge near the top of the insert. The insert engages a mating groove in the wall of the adapter plate thimble tube thru-hole. The insert has 4 equally spaced axial slots which allow the insert to deflect radially at the elevation of the bulge, thus permitting the installation or removal of the nozzle. The insert bulge is-positively held in the adapter plate mating groove by placing a lock tube with a' uniform ID identical to that of the thimble tube into the insert. The lock tube is secured in place by two means. First, a top flare creates a tight fit. Second, six non-yielding projections on the OD which interface with the concave side of the insert preclude escape during core component transfer. The full complement of these joints comprises the structural connection (reconstitutable design feature) between the top nozzle and the remainder of

  . f') the VANTAGE 5 fuel assembly. The nozzle insert-to-adapter plate bulge joints replace the uppermost grid sleeve-to-adapter plate welded joints found in i.

4544Fio87F910513:50 12

l typical #LOPAR fuel: assembly designs. The nozzle insert-to-thimble _ tube-multiple 4-lobe bulge joint located in the lower portion of the insert N represents the structural connection between the insert and the remainder of , the fuel assembly below the elevation of the insert. The uppermost grid.

                   -sleeve is connected to the thimble tube by similar_4-lobe bulge joints.

To remove the top nozzle, a tool is first inserted through the lock tube and expanded radially to engage the bottom edge of the tube. An axial force is then exerted oli the tool which overr 1es the local lock tube deformations 'and withdraws the lock. tube from the insert. After the lock tubes have been withdrawn, the nozzle is removed by raising it off the upper slotted ends of the _ nozzle inserts which deflect inwardly under the axial lift load.

With the. top nozzle removed, direct access is provided for fuel rod examination or replacement. Reconstitution is completed by the remounting of the nozzle and the insertion of new lock tubes. The design bases and - evaluation of the RTN are.given in Section 2.3.2 in Reference 2.

                   .The VANTAGE 5 design will include the use-of the DFBN (initial use was FNP Unit 1; Cycle 10) to reduce the possibility of fuel rod damage due to
                    ~ debris-induced fretting. For-the DFBN design the relatively large flow holes in the-standard bottom nozzle _are replaced with a new pattern of smaller flow
                    . holes. The holes are sized to minimize passage of debris particles large enough to;cause damage while providing hydraulic equivalency to the previous
                    ' design and continued structural integrity. Recently both in the FNP Unit 1
                     ~ Cycle 11, and in the FNP Unit 2, Cycle 8 the DFBN was modified by adding a reinforcing-ski _rt_to enhance-_the nozzle reliability during postulated adverse handling conditions during_ refueling.

L Axial Blanket

                     'Although noted as a new mechanical feature of the VANTAGE-5 design and

i. licensed in Reference 2,-axial blankets have been.and are currently operating:

                     ,in Westinghouse. plants to reduce neutron leakage and improve: fuel F                       utilization. A description and design application of this feature are

'LO L ! -4544F/687F910513:50- 13 L

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

I contained in Reference 2. Section 3.0. The axial blankets utilize a chamfered q pellet physically different from the enriched pellet in the fuel stack to help C/ prevent accidental minng with the enriched pellet during manufacturing. ' Axial blankets are optional starting with the first transition cycle to VANTAGE 5 fuel. Mechanical comoatibility of Fuel As51mjdica Based on the evaluation of the VANTAGE 5/LOPAR design differences and hydraulic test results. References 1 and 2, it is concluded that the twe designs are mechanically compatible with each other. The VANTAGE 5 fur A l mechanical design bases remain unchanged from that used for the LOPAR fuel j assemblies. Furthermore, the effect of the reduceJ VANTAGE 5 fuel atsamblies l I weight, as compared to the LOPAR fuel assembly, has been evaluated for the reactor internals, fuel handling and refueling interfaces and were found accephble. Rod Bow O V It is predicted that the 17x17 VANTAGE 5 rod bow magnitudes, like'those of the Westinghouse OFA fuel, will be within the bounds of existing 17x17 LOPAR assembly rod bow data. The current NRC approved methodology for comparing rod bow for two different fuel assembly designs is given in Reference 7. The magnitude of fuel rod bow for the VANTAGE 5 design is taken to be the same as that for the LOPAR design in the 20 inch spans (between mixing vane gr' h) . Because of the additional support provided by the IFH grids in the DNB limiting spans, the effective span is reduced to approximately 10 inches. As such, there is much less rod bow in the spans adjacent to an IFM grid when compared to the LOPAR rod bow in the 20 inch spans between the mixing vane grids. Rod bow in fuel rods containing IFBAs is not expected to differ in magnitude or frequency from that currently observed in Westinghouse LOPAR fuel rods under similar operating conditions. No indications of abncemal rod bow have (] been observed on visual or dimensional inspections performed on the test IFBA V rods. Rod growth measurements were also within predicted bounds. 4544F/687F910513:50 '4

l EuftLRod_ Rear . (3 V Fuel rod wear is dependent on both the support conditions and the flow environment to which the fuel rod is subjected. Due to the current LOPAR and l VANTAGE 5 fuel assembly designs employing different grid designs, there is an unequal axial pressure distribution between the assemblies. Crossflow resulting from this unequal pressure distribution was evaluated to determine l the induced rod vibration and subsequent wear. Hydraulic tests (Reference 2 Appendix A.1.4) were performed to verify hydraulic compatibility of the L0rAR and VANTAGE 5 designs. The VANTAGE 5 fuel assembly was flow tested adjacent to a Westinghouse 17x17 0FA, since vibration test results indicated that the j crossflow effects produced by this fuel assembly combination would have the i most detrimental effect on fuel rod wear. Results of the wear inspection and analysis discussed in Reference 2, Appendix A.1.4, revealed that the VANTAGE 5 fuel assembly wear character'stic was similar to that of the 17x17 0FA when both sets of data were normalized to the test duration time. It was ccncluded that the VANTAGE 5 fuel rod wear would f3 be less than the maximum wear depth established. Reference 8 for the 17x17 l U OFA at EOL. l l Instrumentation Tube The VANTAGE 5 instrumentation tube is made from material identical to that used in the LOPAR fuel assembly design. Dimensionally, the VANTAGE 5 instrumentation tube ID has an 0.008 inch diametral decrease as compared to

the LOPAR design. This decrease still allows sufficient clearance for the l flux thimble (max OD = 0.300 inch' to traverse the instrumentacion tube without binding. Instrumentation tubes indicative of the VANTAGE 5 design l have been used successfully in other Westinghouse designed plants having similar incore instrumentah on.

l l l n b l 4544F/687F910513:50 15 l

Seismic /LOCA _ Imp.get on fuel Assemblin l An evaluation of the VANTAGE 5 fuel assembly structural integrity considering the lateral effects of LOCA and seismic loadings has been performed using time-history numerical techniques uased on the Farley plant-specific Safe Shutdown Earthquake (SSE). I The VANTAGE 5 fuel assembly is structurally equivalent to the LOPAR fuel

     -design. The main differences between the two designs are six zircaloy grids replacing inconel mid-grids-in the- LOPAR design, three additional intermediate flow mixers, and optimized ft ~1 rods. The load bearing capability for the                                    >

zircaloy~ grids and the intermi iiate flow mixers under-the faulted condition loadings has been analyzed. .D results indicated that 17x17 VANTAGE 5 grid- ) loads (for Zircaloy and flow mixing grids) are below the allowable grid I strengths. Based on the grid' load results. .the 17x17 VANTAGE 5 zircaloy grid is capable-of maintaining the' core coolable geometry under the SSE and asymmetric pipe rupture transients in either homogeneous or transition core operation. The 17x17 VANTAGE 5 fuel assembly is structurally acceptable for both'FNP Units 1 and 2. This'is also true for. a transition core composed of both VANTAGE 5 and LOPAR fuel assembiy core configur tions. The grids will not buckle due to the combined impact loads of a seismic and LOCA event. The coolable geometry requirement.is met. The stresses in the fuel assemb?y components resulting from seismic and LOCA induced deflections are within acceptable limits. Core components The FNP Units 1 and 2 contain plugging devices which utilize thimble plugs of two different designs. The initial supply of plugging devices included 0.434 inch diameter thimble plugs whici are compatible with the LOPAR fuel assembly design only. Use of this design of plugging device after implementation of VANTAGE 5 into the FNP Units I and 2 will be restricted to LOPAR fuel assembly locations. O . 4544F/687F910513:50. 16

  - - _ . _ _ . _ _ . . . _ . . . . _ _ _ .      . _ _ _ u _ __. . _ _. _ _ . . _ _                            _

Hore recently, the FNP Units 1 and 2 obtained " dually compatible" plugging devices containing 0.424 inch thimble plugs. The " dually compatible" pl*.:gging s devices are compatible with both LOPAR and VANTAGE 5 fuel assembly designs and, as such, have no restrictions on placement, or location, during the transition to a full VANTAGE 5 core. Tran ient and accident analyses have been performed which support operation of the VANTAGE 5 fuel assembly with or without thimble plugs. Secondary source assemblies for both the FNP Units 1 and 2 contain 0.434 inch diameter thimble plugs and are compatible only with the LOPAR fuel assembly design. Use of these particular secondary sour:e assemblies after implementation of VANTAGE 5 into the FNP Units 1 and 2, will be restricted to LOPAR fuel assembly locations. Upon transition to a full VANTAGE 5 core, secondary sources compatible with the smaller diameter thimbles will be used, or sourceless startup will be implemented. The thimble tubes in both LOPAR and VANTAGE 5 fuel assemblies provide sufficient clearance for insertion of control rods to assure proper operation and control of the reactor. Thus, the current RCCA designs are acceptable for

     '    use during and subsequent to implementation of VANTAGE 5 into the FNP Units 1 and 2.

O i l 4544F/687F910513:50 17

3.0 NUCLEAR EVALUATION q V The evaluation of the transition and equilibrium cycle VANTAGE 5 cores presented in Reference 2, as well as the FNP specific transition core evaluation, demonstrate that the impact of implementing VANTAGE 5 does not cause a significant change to the physics characteristics of the FNP cores beyond the normal range of variations seen from cycle-to-cycle. The methods and core models used in the FNP reload transition core evaluations are described in Reference 2, 4, 9, and 10. These licensed methods and models have been used for Farley and other previous Westinghouse reload designs using the OFA and VANTAGE 5 fuel. No changes to the nuclear design philosophy, methods, or models are necessary because of the transition to VANTAGE 5 fuel. For the nuclear design area, the following FNP Units 1 and 2 Technical Specifications changes are proposed:

1) Increased F gg limit. The higher limit for the VANTAGE 5 fuel serves to O

increase nuclear design flexibility and allows loading patterns with reduced leakage which in turn will allow longer cycles.

2) Increased gF limit. The higher Fg limit for the VANTAGE 5 fuel will provide greater flexibility with regard to accommodating the axially heterogeneous cores (axial blankets and part length burnable absorbers).

3) Increased Positive HTC limit. The increased Positive HTC limit serves to increase nuclear design flexibility and allows loading patterns with fewer numbers of burnable absorbers due to the decreased reactivity holddown required at the beginning-of-life.

Power distributions and peaking factors show slight changes as a result of the incorporation of reduced length burnable absorbers, axial blankets, and increased peaking factors limits, in addition to the normal var,iations experienced with different loading patterns. The usual methods of enrichmeni variation and burnable absorber usage can be employed in the transition and full VANTAGE 5 cores to ensure compliance with the peaking factor Technical Specifications. 4544F/687F910513:50 18

The RCCA reactivity worth versus rod insertion used in the FSAR Chapter 15 accident analyses conservatively bounds the effects of axial blankets and O reduced length burnable absorbers as well as cycle-to-cycle diff'rences. No changes to the rod insertion limits were required as a result of the VANTAGE 5 fuel. The key safety parameters evaluated for the Farley reactor as it transitions to an all VANTAGE 5 core show little change relative to the range of parameters experienced for the all LOPAR fuel core. The changes in values of the key safety parameters are typical of the normal cycle-to-cycle vcriations experienced as loading patterns change. As is current practice, each reload core design will be evaluated to assure that design and safety limits are satisfied according to the reload methodology. The design and safety limits will be documented in each cycle specific Reload Safety Evaluation which serves as a basis for any significant changes which may require a future NRC review. O O 4544F/687F910513:50 19

2 4.0 THERHAL AND HYDRAULIC EVALUATION This section describes the calculational methods used for the thermal-hydraulic analysis, the DNB performance, and the hydraulic compatibility during the transition from mixed-fuel cores to an all VANTAGE 5 core. The Westinghouse transition core DNB mathodology is given in References 1 and 11 and has been approved by the NRC via Reference 12. Using this methodology, transition cores are analyzed as if the entire core consisted of one assembly type (full LOPAR or full VANTAGE 5), and the resultant DNBR values are reduced by the appropriate transition core penalty. The LOPAR and the VANTAGE 5 fuel assemblies were shown to be hydraulically compatible in Reference 2. The DNBR analyses for FNP Units 1 and 2 were based on paramnters which conservatively bound the licensing values. ' Table 4-1 summarizes the pertinent thermal and hydraulic design parameters used in the analyses as well as the licensing values. The improved THINC-IV PHR design modeling method, Reference 13, was used for the DNBR analyses of the VANTAGE 5 and LOPAR fuel. For high core power density applications, the improved model yields more conservative values of minimum DNBR than the present model, Reference 14. No changes to the basic THINC-IV models and correlations were made for the improved core modeling scheme. The DNBR analyses of the VANTAGE 5 fuel and LOPAR fuel are baseo on the Revised Thermal Design Procedure (RTDP), Reference 15. The primary DNB correlation used for the LOPAR fuel is the HRB-1 DNB correlation which is described in Reference

16. Tht. VANTAGE 5 fuel DNBR analyses use the WRB-2 DNB correlation which is described in Reference 2. hie WRB-2 DNB e.>rrelation takes credit for the reduced grid-to-grid spacing of the VANTAGE 5 fuel assembly mixing vane grids resulting from the use of the Intermediate Flow Mixer (IFH) grids. Both the HRB-1 and WRB-2 DNB correlations have a correlation limit of 1.17.

O G 4544F/687F910513:50 20

I The H-3 DNB correlation, References 17 and 18, is used for both fuel types where the primary DNB correlations are not applicable. The HRB-1 and HRB-2 DNB V correlations were developed based on mixing vane data and, therefore, are only applicable in the heated rod spans above the first mixing vane grid. The l H-3 DNB correlation, which does nn'. take credit for mixing vane grids, is used to calculate DNBR values in the heated region below the first mixing vane grid. In j addition, the H-3 DNB correlation is applied in the analysis of accident conditions where the system pressure is below the range of the primary correlations. For system pressures in the range of 500 psia to 1000 psia, the H-3 DNB correlation limit is 1.45, Reference 19. For system pressures greater that. 1000 psia, the H-3 ONB correlation limit is 1.30. A cold wall factor, Reference 20, is applied to the H-3 DNB correlation to account for the presence of the unheated thimble surfaces. Also, a 0.BB multiplier is applied to the H-3 DNB correlation to account for the 17x17 fuel rod diameter effect, Reference 21. Hith the RTDP methodology, uncertainties in plant operating parameters, auclear ! and thermal parameters, fuel fabrication parameters, computer codes and DNB correlation predictions are considered statistically to obtain DNBR uncertainty factors. Based on the DNBR uncertainty factors, RTDP design limit DNBR values are determined such that there is at least a 95 percent probability at a 95 percent confidence level that DNB will not occur on the most limiting fuel rod during normal operation and operational transients and during transient conditions arising from faults of moderate frequency (Condition I and II events as defined in ANSI N18.2). Uncertainties in the plant operating parameters (pressurizer pressure, primary coolant temperature, reactor power, and reactor coolaca system flow) have been evaluated for the FNP Units 1 and 2 with Resistance Temperature Detector (RTD) bypass loops, Reference 22, and for the RTD bypass loops eliminated, Reference

23. In the DNBR analyses with RTDP, a set of plant operating parameter uncertainties were used which are bounding for operation with RTD bypass loops or for RTD bypass loops eliminated. Since the parameter uncertainties are considered in determining the RTDP design limit DNBR values, the plant safety analyses are performed using input parameters at their nominal values.

O 4544F/6B7F910513:50 21

TABLE 4-1 THERf%L AND HYDRAULIC DESIGN PARAMETERS

                           \                                                                                                                              FDR INP l! NITS 1 AND 2 Thermal and Hvdraulie Desian Parameters                                                 Analysit ParamJ,igts                                             Licensino Param11til (Using RTDP)

Reactor Core Heat Output. W t 277E' 2652 6 Reactor Core Heat Output. 10 BTU /hr 9469 9051 Heat Generated in Fuel. %" 97.4 97.4 Pressurizer Pressure. Nominal, psia 2250 2250 F

g. Nuclear Enthalpy Rise Hot Channel Factor (LOPAR) 1.60(1+.3(1-P)' 1.55(1+.3(1-P))

(V-5) 1.70(1+.3(1-P)? 1.65(1+.3(1-P)) Minimum DNBR at Nominal Conditions Typical Flow Channel (LOPAR) 2.36 >2.36 (V-5) 2.36 >2.36 Thimble (Cold Wall) Flow Channel (LOPAR) 2.26 >2.26 (V-5) 2.23 >2.23 Design Limit DNBR Typical Flow Channel (LOPAR) 1.25 1.25 (V-5) 1.24 1.24 Thimble (Cold Wall) Flow Channel (LOPAR) 1.24 1.24 [ (V-5) 1.23 1.23 DNB Correlation *** (LOPAR) VRB-1 VRB-1 (V-5) WRB-2 WRB-2

See Section 5 for the LOCA analysis value.

                                                                                      "      Fraction of core heat generated in the fuel rod; used in rod linear power and heat flux calculations.
                                                                                      ""     The W-3 correlation is used for conditions outside the range of applicability of the primary DNB correlation.

1 l l 4544F/687F910513:50 22 i

TABLE 4-1 (continued) THERMAL AND HYDRAULIC DESIGN PARAMETERS O FOR FNP UNITS 1 AND 2 ENpinal Coplant Conditionj ADilysis Param_tLign" Licentina Parameltn" Vessel Minimum Heasured Flow Rate (including Sypass), 0 10 lbm/hr 100.1(a) 101.5(b) gpm 263,400 " " 267,880 "

Vessel Thermal Design Flow Rate (including Sypass),

10 lbm/hr 98.1 99.3 gpm 258,000 261,600 Core Flow Rate (excluding Sypass, based on Thermal Design Flow) f 10 lbm/hr 91.1 92.2 gpm 239,680 243,030 Fuel Assembly Flow Area for heat Transfer, f t * (LOPAR) 41.55 41.55 (V-5) 44.04 44.04 Core Inlet Mass Velocity. 10 lbm/hr-ft (Based en TDF)+ (LOPAR) 2.19 2.22 (V-5) 2,07 ' 0' t. i

Analysis flow rates are based on 20% steam generator tube plugging.

     - "    Licensing flow rates are based on 15% average /20% peak steam generator tube plugging.

l

      '"    Value includes a 2.4% flow uncertainty (0.1% feedwater venturi fouling bias included).
      "" Value includes a 2.1% flow uncertainty.

(a) Inlet temperature = 541.8*F (b) Inlet temperature = 543.8'F (

      + Assumes all LOPAR or VANTAGE 5 Core 4544F/687F910513:50                                  23

TABLE 4-1 (continued) THERML AND HYDRAULIC DESIGN PARAMETERS [O FOR FNP UNIT $ 1 AND 2 Thermal and Hydraulie Desion Parameters Analvsis Paramete,ts Listnsina Partm_gi ns (Based on TDF) Nominal Vessel / Core Inlet Temperature, 'F 541.1 543.1 Vessel Average Temperature, 'F $77.2 577.2 Core Average Temperature, 'F 581.8 581.5 Vessel Outlet Temperature. 'F 613.3 611.3 Average Temperature Rise in Vessel, 'F 72.2 68.2 Average Temperature Rise in Core, 'F 77.0 72.9 Heat Transfar Active Heat Transfer Surf ace Area, f t * (LOPAR) 48,598 48,$98 (V-5) 46.779 46,779 Average Heat Flum, BTU /hr-f t (LOPAR) 189,820 181,410 (V-5) 197,200 188,460 Average Linear Power, kw/ft" 5.45 5.20 Peak Linear Power (LOPAR) 12.63 12.07 for Normal Operation, Kw/f t " (V-5) 13.34 12.75 Temperature Limit for Prevention of Centerline Melt, 'F 4,700 4,700

  +   Assumes all LOPAR or VANTAGE 5 core.

Based on densified active fuel length.

  +* Based on 2.32 F peaking f actor for LOPAR and 2.45 F peaking f actor for VANTAGE 5.

4544F/687F910513:50 24

The RTDP design limit DNBR values are 1.24 and 1.23 for the typical and thimble cells respectively for VANTAGE 5 fuel, and 1.25 and 1.24 for the typical and thimble cells respectively for LOPAR fuel. In addition to the above considerations, plant specific DNBR margin was maintained by performing the safety analyses to DNBR limits higher than the design limit DNBR values. A fraction of the available DNBR margin is utilized to accommodate the transition core penalty. For VANTAGE 5 fuel, this transition core penalty is a function of the number of VANTAGE 5 fuel assemblies in the core as given in Reference 24. There is no transition core penalty for the LOPAR fuel. Additional margin is used to offset the rod bow DNBR penalty. Based on Reference 7, the fuel rod bow DNBR penalty is less than 1.57. for both LOPAR and VANTAGE 5 fuel in the 20 inch grid spans. No rod bow penalty is required in the 10 inch grid spans of the VANTAGE 5 fuel. The remaining DNBR margin, after consideration of these penalties, is available for operating and design flexibility. ; Q The option of thimble plug removal has been included in all of the DNBR analyses performed for the VANTAGE 5 and LOPAR fuel. The primary effect of thimble plug removal is an increase in the core bypass flow. This increased core bypass flow is reflected in the core flow rates and the DNBR values presented in Table 4-1. The Standard Thermal Design Procedure (STDP) is used for those analysis where the RTDP methodology is not applicable. In the STDP method, the parameters used in DNBR analyses are treated in a conservative way to give the lowest minimum DNBR. Sufficient DNBR mar 31n to cover appropriate DNBR penalties is preserved whenever the STDP is used. The fuel temperatures used in safety analysis calculations for the VANTAGE 5 and LOPAR fuel were calculated with the improved fuel performance code, Reference 6. This code was used to perform both dc:ign and safety calculations. These fuel temperatures were used as initial conditions for LOCA and non-LOCA transients. 4544F/6B7F910403:50 25

w. --- - .v- - . , - - -- .

                                                                                  .w w

i r 5.0 -ACCIDENT EVALUATION 5.1 Non-LOCA Accidents k . l This section' addresses the effects of tne VANTAGE 5 design features and the safety analysis a sumptions for the FNP Units 1 and 2 non-LOCA accident analyses.- ; 5.1.1 VANTAGE 5 Design Features The design features that were considered-in the non-LOCA safety analysis : include the following: o Fuel Rod Dimensions ! o Intermediate Flow Hixer (IFM) Grids

                          -o       Axial Blankets-o     Integral Fuel Burnable Absorbers (IFBA)-                                                                           ,

o~ <Reconstitutible Top Nozzle (RTN)-

o Modified Debris Filter Bottom Nozzle (MDFBN) '

o Extended Burnup (} o Zircaloy. Grids , A brief description of.each of these features and its treatment in the non-LOCA safety analysis follows. I Fuel Rod Dimensions The fuel rod dimensions which determine the temperature versus linear power den ~sity relationship include rod diameter, pellet diameter, initial > pellet-to-clad gap size, and stack height. The fuel rod temperature and : geometry' assumptions used in the non-LOCA safety analysis bound both the LOPAR

      -              and VANTAGE 5 fuel, f
                                                                                                                                                    .I p

N

    - (s i

_,4544F/687F910403:50 26 _ _ - ...~ . _ - _ _ -.._._.. _ .,.. _ .. _ _ _ _ _ _ _ _ _ . _ .. ,-...___

Intermediate _finw Mixer (IFM) Grids O v The IFH grid feature of the VANTAGE 5 fuel design provides an increase in the DNB margin over LOPAR fuel. As a result, the safety analysis limit DNBR values contain significant DNB margin for the VANTAGE 5 fuel (see Section 4.0). The IFM grid feature of the VANTAGE 5 fuel design increases the core pressure drop, reduces the guide thimble I.D., the guide thimble 0,0. and results in an increased control rod drop time. The rod drop time to dashpot increases from 2.2 seconds for LOPAR fuel to 2.7 seconds for VANTAGE 5 fuel. The increased rod drop time primarily affects the fast reactivity transients, all of which were explicitly analyzed for this report. The increased rod drop time was modeled in ail the analyzed events and the remaining transients have been evaluated. Axial Blankets and Integral Fuel Burnable Absorbers (IFBAs) Axial blankets reduce power at the top and bottom of the fuel rod, which increases axial power peaking at the center of the rod. This effect is offset by the presence of part length IFBAs which tend to even the power distribution. The net effect on the axial power shape is a function of the number and configuration of the IFBAs in the core and the time in core life. The effects of axial blankets and IFBAs on the safety analysis parameters are taken into accour.t during the reload design process. The axial power di.itribution assumptions in the safety analysis kinetics calculations are sufficiently bounding to accommodate the presence of axial blankets and IFBAs in FNP Units 1 and 2. Reconstitutable Too Nozzle (RTN) and Modified Debris Filter Bottom-_ Nozzle (HDFBtil RTNs and HDFBNs have been used extensively in Westinghouse designs and are currently being used in both FNP Units 1 and 2. Analyses and tests have been performed that confirm the hydraulic compatibility of these particular components to existing designs and confirmed that these components do not affect any parameter important to the non-LOCA safety analysis. O 4544F/687F910513:50 27

Extended Burnup D d The VANTAGE 5 fuel assemblies are designed for extended burnup capability by reducing the thickness of both the top and bottom nozzle end plates, thereby decreasing the height of the bottom nozzle and increasing the length of the fuel rod. The effects of these extended burnup features have been accounted for in the non-LOCA safety analysis. The extended burnup features are currently in use in the FNP Units I and 2. l Zircalov Gridl Zircaloy structural grids have replaced inconel grids in the VANTAGE 5 fuel assemblies with the exception of the top and bottom grids, which remain Inconel. The effects of zircaloy grids have been accounted for in the non-LOCA safety analysis. 5.1.2 Safety Analysis /.ssumptions p The following core related features and assumptions have been considered in l A/ the events analyzed for this report and, as necessary, evaluated for all other non-LOCA events comprising the FNP Units 1 and 2 licensing basis, o Revised Thermal Design Procedure (RTDP) for Appropriate DNB Events o Revised non-RTDP Uncertainties for RCS Temperature and Pressure o Increased Beginning-of-Life Positive Moderator Temperature Coefficient of +7 pcm/*F Up to 70% Rated Thermal Power (RTP), Ramping to O pcm/*F at 100% RTP o Removal of Thimble Plugs o Increased Power Distribution Peaking Factors (F AH and Fg) o Modified Overtemperature and Overpower AT Reactor Trip Setpoints The effect of the following additional items on the results of the safety analysis has been specifically determined. These additional assumptions are conservative and require no technical specification changes for the analysis O 4544F/687F910508:50 28

to be valid. However, additional work is required before these changes can r] become part of the licensing basis and subsequently implemented at the FNP units. o Increased Core Thermal Power o Reduced Thermal Design Flow o 20% Steam Generator Tube Plugging o Relaxed Axial Offset Control (RAOC) o Increased End-of-Life Moderator Density Coefficient o Increased Pressurizer and Steam Generator Safety Valve Uncertainty o Increased Overpower /0vertemperature AT Reactor Trip Response Time A discussion of each of these items is given below. Revised Thermal Desian Procedure (RTDP) for Anorooriate DNB Events The Revised Thermal Design Procedure (RTOP), described in Refer nce 15, was utilized to meet the DNB design basis. Conser9 tive uncertainties in the plant operating parameters were statistically incorporated in the design limit DNBR value as discussed in Section 4.2. Since the parameter uncertainties are considered in determining the design DNBR value, the associated plant safety analyses are performed using nominal initial conditions. Revised Non-RTDP Uncertaintigs for RCS Temocrature and Pressure The pressurizer pressure uncertainty has been increased from 130 psi to l150 psi. The RCS temperature uncertainty has been increased from 14 *F to 16*F. The core power uncertainty is unchanged and remains 2.0% RTP. Increased Beainnina-of-Life Positive Moderator Temocrature Coefficient (PMTC) A positive moderator temperature coefficient (PMTC) of +7 pcm/*F from 0% to 70% RTP and decreasing linearly to O pcm/*F at 1001 RTP was incorporated into the safety analysis. In general, the analyses presented are based on a ! +7 pcm/*F MTC, which is assumed to remain constant for variations in temperature. Exceptions are the Rod Ejection and Rod Hithdrawal from L 4544F/687F910515:50 29

Subtritical events which are based on a +7 pcm/*F at zero power nominal average temperature, but due to moderator feedback effects modeled in THINKLE ( diffusion theory code become less positive as the transient temperature increases. Removal of Thimble Plug Thimble plug removal affects the core pressure drop and increases core bypass flow. These effects have been conservatively incorporated into the non-LOCA analyses performed for this report. The analyses are applicable for either thc thimble plugs in or removed. Increased Power Distribution Peaking Factors (FAH and FQ ) The full power F 3g assumed in the non-LOCA events analyzed for DNBR is 1.60 for LOPAR fuel and 1.70 for VANTAGE 5 fuel. See Section 4.2 of this report for a description of how peaking factors are applied in the analysis. The maximumgF assumed in non-LOCA safety analyses has been increased from 2.32 to 2.50. Modified Overtemoerature and ,0veroower AT Reactor Trio Setunijlis The implementation of the VANTAGE 5 fuel; the inclusion of more conservative uncertainty values for RCS temperature and pre:sure; and changes in hot channel factors, steam generator tube plugging level, and RCS flow t.ause the DNB core limits to change. With the core limit change, the overpower and overtemperature AT reactor trip setpoints are changed in the analysis. The values of the setpoints used in the safety analysis are given in Appendix A. The revised technical specification values are given in Attachment 2. Increased Core Thtrinal Power An increase in the nominal core thermal power from 2652 MHt to 2775 MHt was considered in the non-LOCA safety analyses for this report. The non-LOCA safety analyses performed at 2775 MHt will conservatively bound the current rated core thermal power level of 2652 MHt. l 4544F/687F910515:50 30

a Reduced Thermal Desian Flag

                                    ~
         - A decrease in~ the _ RCS thermal design flow from 251,600 gpm to 258,000 gpm was considered in the non-LOCA safety analyses performed for this report. The non-LOCA safety analyses performed at 258,000 gpm will conservatively bound
          -the current thermal design flow of 261,600 gpm.

20% Steam Generator Tube Pluaging

          'All-of the events analyzed for this report have incorporated modeling              i assumptions that bound up to a maximum of 20% steam generator tube plugging (or the hydraulic equivalent of plugs and sleeves) in each steam generator.
                                                  ~

It is-assumed that~'no single steam generator exceeds 20% tube plugging. Relaxed Axial' Offset Control (RAOC) The non-LOCA safety analyses performed for this report are applicable to RAOC operation with a +10/-15% Axial Flux Difference (AFD) band-at 100% Rated Thermal. Power (RTP) and conservatively bound the current plant constant axial 4 N offset control (CAOC) operation with a 15% AFD band. Increased End-of-Life Moderator Qgnsity Coefficient

          - In order to accommodate longer fuel cycles and increased fuel burnup, a maximum moderator density coefficient of 0.50 ok/g/cc (corresponding to end-       ,

of-life full power condition) was conservatively incorporated into the safety ! i analyses. performed for this; report. 5 inCIRAlpd Pressurizer and Steam Generator Safety Valve Uncertainty The non-LOCA' safety analyses performed for this report include revisions in the treatment of the pressurizer safety valves. In those events that may challenge the peak reactor coolant system pressure limit, the pressurizer , safety valve setpoint included a 1% uncertainty. Additionally, the flow

through the pressurizer safety valves was modeled with a 3% accumulation. i'.e., the flow ramps from zero to full rated flow over the range of 2525 to 4544F/687F910515:50 31

2601 psia. Previously, the pressurizer safety valves had been modeled as opening at a pressurizer pressure of 2500 psia, with full rated flow being reached at 2575 psia. For FNP Units 1 and 2, a change in the pressurizer safety valve setpoint primarily affects the Loss of External Electrical Load and/or Turbine Trip and the Single Reactor Coolant Pump Locked Rotor events. The non-LOCA safety analyses performed for this report modeled the steam , generator safety valves as opening and being full open at 1167 psia or 6% l above the steam generator design pressure of 1100 psia. This 6% can be considered to include allowances for valve accumulation and uncertainty. (Previously, the steam generator safety valves had been modeled as being full open at 1133 psia or 3% above the design pressure.) for FNP Units 1 and 2, a change in the steam generator safety valve setpoint primarily affects the Loss of External Electrical Load and/or Turbine Trip events. Other events also , affected, but to a lesser degree, are the Uncontrolled RCCA Hithdrawal at Power, Rupture of a feedline, Loss of Normal Feedwater, and Loss of All AC Power to the Station Auxiliaries events. I Increased Overocyer/0vertemoerature AT Reattor Trip _Rc3ponse Time The total time delay of the overtemperature and overpower AT reactor trips assumed in the non-LOCA analyses is 8 seconds. This corresponds to the total delay from the time the temperature differences in the loop exceeds the trip setpoint until the rod cluster assemblies are free to fall into the core. Included in this 8 second delay is a 6 second first order lag incorporated into the determination of the time at which the overtemperature and overtemperature AT setpoints are reached. Specifically, the lag function includes allowances for RTD and thermowell time response and RTD bypass loop fluid transport delay (if appropriate). The remaining 2 seconds account for the delay from the time at which the trip signal is initiated until the rod cluster assemblies are free to drop into the core. 4544F/687F910515:50 32 __ _ _ _ _ _ - -_ ~.

5.1.3 Non-LOCA Safety Evaluation Methodology O The non-LOCA reload safcty evaluation methodology is described in Reference 4. This methodology confirms that, if a core configuration is bounded by existing safety analyses, then the applicable safety criteria are satisfied. The methodology systematically identifies both parameter changes on a cycle-by-cycle basis which may violate existing safety analysis assumptions and the transients which require evaluation. This methodology is applicable to the evaluation of VANTAGE 5 transition and full cores. Any required evaluation identified by the reload methodology is one of two types. If the identified parameter is only slightly out of bounds, or if the transient is relatively insensitive to the parameter, a simple evaluation may be made which conservatively evaluates the magnitude of the effect and explains why the actual analysis of the event does not have to be repeated. Alternatively, should the deviation be large and/or expected to have a significant or not easily auantifiable effect on the transients, analyses are required. The analysis approach will utilize Westinghouse codes and methods which have been accepted by the NRC and have been used in previous submittals to the NRC. These methods are those which have been presented to the NRC for a specific plant or as part of reference Safety Analysis Reports or reports for NRC approval. The analysis methods and cc, des are described in Appendix A. The key safety parectcrs are documented in Reference 4. Values of these safety parameters which bouad both fuel types (LOPAR and VANTAGE 5) were assumed in the non-LOCA safety analyses. For subsequent fuel reloads, the key safety parameters will be evaluated to determine if violation of these bounding values exist. An evaluation of the affected accidents will take place as described in Reference 4. The following transients affected by the VANTAGE 5 fuel design features or modified safety analysis assumptions were reanalyzed:

1. Uncontrolled rod cluster control assembly bank withdrawal from a subcritical condition (see Section 15.2.1 of Appendix A) i 4544F/6C7F910515:50 33 1

2. Uncontrolled rod cluster control assembly bank withdrawal at power fl (see Section 15.2.2 of Appendix A) l J 1

3. Rod cluster control assembly misalignment (see Section 15.2.3 of Appendix A)

4. Uncontrolled boron dilution (see Section 15.2.4 of Appendix A)

5. Partial loss of forced reactor coolant flow (see Section 15.2.5 of Appendix A)

6. Startup of an inactive reactor coolant loop (see Section 15.2.6 of Appendix A)

7. Loss of external electrichi load and/or turbine trip (see Section 15.2.7 of Appendix A)

8. Loss of normal feedwater (see Section 15.2.8 of Appendix A) +

  !                                                                                                 i

9. Loss of all offsite power to the station auxiliaries (see Section i 15.2.9 of Appendix A)

10. Excessive heat removal due to feedwater system malfunctions (see 1 Section 15.2.10 of Appendix A)

11. Excessive load increase incident (see Section 15.2.11 of Appendix A)

12. Accidental depressurization of th,e reactor coolant system (see Section 15.2.12 of Appendix A)

13. Accidental depressurization of the main steam system (see Sectien 15.2.13 of Appendix A)

14. Inadvertent operation of the emergency core cooling system during power operation (see Section 15.2.14 of Appendix A) 4544F/687F910515.50 34

15. Complete loss of forced reactor coolant flow (see Section 15.3.4 of N Appendix A)

(G 4

16. Single rod cluster control assembly withdrawal at full power (see Section 15.3.6 of Appendix A) i

17. Rupture of a steamline (see Section 15.4.2.1 of Appendix A)

18. Single reactor coolant pump locked rotor (see Section '5.4.4 of ,

Appendix A) 1

19. Rupture of a control rod drive mecnanism housing (RCCA Ejection) (see Section 15.4.6 of Appendix A) '

The following non-LOCA accidents were evaluated with respect to the VANTAGE 5 fuel design features or modified safety analysis assumptions:

1. Hinor secondary system pipe breaks (see Section 15.3.2 of Appendix A)

O '

2. Inadvertent loading of a fuel assembly into an improper position (see Section 15.3.3 of Appendix A)

3. Rupture of a main feedwater line (see Section 15.4.2.2 of Appendix A) 5.1.4 Steamline Break Mass and Energy Releases 5.1.4.1 Mass and Energy Releases Inside Containment

                                 -The steam generator mass and energy release data inside containment are generated to determine the containment pressure and temperature response following a steamline break event. As discussed in Section 6.2.1.2.11 of the FSAR, the mass and energy releases are calculated for a variety of power levels, break sizes, and single failure assumptions. An evaluation of the current licensing basis steamline break inside containment transient analysis revealed many conservative assumptions with respect to the actual FNP design.

4544F/687F910515:50 35

1 I l 1 l The evaluation considered the combined effect of the VANTAGE 5 fuel, associated licensing basis changes, and conservative analysis assumptions on the mass and energy release data. The evaluation concluded that the current steamline break inside containment mass and energy releases for FNP Units 1 and 2 remain valid for the VANTAGE 5 fuel upgrade and associated licensing basis changes. 5.1.4.2 Mass and Energy Releases Outside Containment The steam generator mass and energy release data outside containment are j generated to ensure that the equipment environmental qualification limits are j met following a steamline break. The outside containment analyses were ! performed generically (Reference 36) such that the analyses bound several l Westinghouse 3 loop plants. These analyses are conservative with respect to the actual FNP design, including the changes discussed in Sect',on 5.1.1 and those analysis assumptions being incorporated into the FNP USits 1 and 2 licensing basis. Therefore, it is concluded that the current steamline break i outside containment mass and energy releases for FNP Units 1 and 2 remain O

 % ./

valid for the VANTAGE 5 fuel upgrade and associated licensing bNis changes. 5.1.5 Conclusions Descriptions of the non-LOCA accidents analyzed for this report, mathod of l analysis, results, and conclusions are contained in Appendix A. Appendix A J conforms to the format of the FNP Units 1 and 2 FSAR. It was found that the appropriate safety criteria were met for each of the transients analyzed. Based on the plant operating limitations given in the Technical Specifications and the proposed technical specification changes given in Section 6.0 of this report the results show that the transition from LOPAR to VANTAGE 5 fuel can be accommodated with margin to the applicable FSAR safety limits. 1 5.2 LOCA Accidents This section addresses the effects of the VANTAGE 5 design features and modified safety analysis assumptions for the FNP Units 1 and 2 LOCA analyses. 4544F/687F910515:50 36

_ ._ . =, __ . .- . _ . 5.2.1 Large Break LOCA n v 5.2.1.1 Description of Analysis / Assumptions for 17X17 VANTAGE 5 fuel l

The large break loss-of-coolant accident (LOCA) analysis for FNP Units 1 and 2, i applicable to a full core of VANTAGE 5 fuel assemblies, was performed to develop specific peaking factor limits for the Farley Units. This is consistent with the methodology employed in the Reference Core Report for 17x17 VANTAGE 5, Reference 2. The Westinghouse 1981 Evaluation Model with I BASH, References 25 and 26, was utilized and a spectrum of cold leg breaks was j analyzed for FNP Un'ts 1 and 2 limiting conditions that bound nominal l operating conditic.is. Note that the downflow barrel / baffle configuration of I Farley Unit 2 a; used to perform the spectrum, followed with the limiting break size analyzed for the upflow barrel / baffle cordiguration of Farley Unit 1 (see Reference 27 for additional discussion regarding upflow versus downflow barrel / baffle). Other pertinent large break LOCA analysis l assumptions include: O o A core thermal power of 2652 MHt o 20% steam generator tubes plugged in each of three steam generators (i.e. uniform among the loops) o A total core peaking factor, F , gof 2.45 for VANTAGE 5 fuel and 2.32 for LOPAR fuel o AF AH of 1.65 for VANTAGE 5 fuel and 1.55 for LOPAR fuel o A thermal design flow of 86,000 gpm per loop o A 8% (HHSI/LHSI) pump degradation o A 10 gpm charging /(HHSI) pump flow imbalance o Thimble Plug Removal o Fuel temperatures and pressures based on the improved fuel thermal model, Reference 6 o Containment mini-purge automatic isolation o A RCS temperature operating band of 6*F o A 15 second diesel generator start time o A two line segment K(z) curve o A RCS pressure uncertainty 60 psi o 102% calorimetric uncertainty for nuclear power 4544F/687F910515:50 37

__ . _ _ _ . . . . _ - _ . _ . . _ _ _ . _ _ _ . ~ . . t s5.2.1.2 Method of Analysis The. methods used to analyze the large break LOCA accident for FNP Units 1 and 2 for-VANTAGE 5 fuel, including computer codes used and assumptions, are i described in detail in Appendix B Section 15.4. f 5.2.1.3 Results : The results of this analysis, including tabular and plotted results of the-break' spectrum analyzed, are provided in Appendix B, Section 15.4, which has ; been prepared using the NRC Standard Format and Content Guide, Regulatory Guide 1.70,-Revision 2.for accidents applicable to FNP Units I and 2. Reference 25 states three restrictions related to the use of the 1981

         ' Evaluation Model_(EM) a BASH calculational model. The application of these restrictions to the plant specific large break LOCA analysis was addressed with the following conclusions:                                                                                   ;

FNP Units 1 and 2 are neither Upper Head Injection.(UHI) nor Upper Plenum , Injection'(UPI) plants'so restriction 1 does not apply.

The FNP Units:1- and 2 plant specific large break LOCA analysis considered both minimum and. maximum ECCS cases to address restriction 2. The CD - 0.4 Double Ended Cold Leg Guillotine =(DECLG) break with minimum ECCS flows was found _ to result in the most limiting consequences.

         'Concerning the third restriction, consistent with the conclusions of Reference 29, a chopped cosine power shape was used in the large break LOCA analysis'for FNP_ Units 1 and 2.                                                                                  .

t 5.2.1.4 Conclusions i The l'arge break-LOCA analysis performed for the FNP Units 1 and 2 plants has demonstrated:that for breaks up.to a double-ended severance of the reactor coolant piping, the_ Emergency Core Cooling System'(ECCS) will meet the : acceptance criteria o_f Title 10 CFR Part 50 Section 46. That is: 4544F/687F91b515:50: 38

1. The calculated peak cladding. temperature will remain below the required

{ 2200*F.

2. The amount of fuel cladding that reacts chemically with the water or steam to generate hydrogen does not exceed 1% of the hypothetical amount

                                                       -that_would be generated if all the zirconium metal in the cladding cylinders surrounding the fuel, excluding the cladding surrounding the plenum volume, were to react.

3. The localized cladding oxidation limit of 17 percent is not exceeded during or after quenching.

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

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

In-addition, the contribution of the Zircaloy grids to the overall Zr-H O 2 reaction-has been investigated, although it is not required as per item 3

                                                -above. The 1981 Evaluation Model with BASH, Reference 25, which explicitly models the grids, confirms that the grid temperatures are such that the contribution to the total core wide Zr-H 2O reaction is insignificant. In
                                                . fact, recent LOCA analyses have shown that a core with Zircaloy grids has less total.Zircaloy reacting than the compr.rable core with Inconel grids. This-results because the Zircaloy grids and the cladding in the vicinity of the grids remain at lower temperatures than the corresponding case with Inconel grids,.thus reducing the rate of the Zr-H 2O reaction.

A summary of the large break LOCA cases analyzed is shown in Table 15.4.1-1 of Appendix B. -The time sequence of events for all large breaks analyzed is shown in Table 15.4.1-5 or Appendix B. The large break LOCA results for all the breaks analyzed is shown in Table 15.4.1-6 of Appendix B. O 4544F/687F910515:50 39

The large break LOCA analysis for FNP Units 1 and 2 assuming a full core of VANTAGE 5 fuel, utilizing the 1981 EH with BASH calculational model, resulted O in a Peak Cladding Temperature (PCT) of 1965'F for Unit 1 and 2073'F for Unit 2, for the limiting C 0.4 D DECLG break at a total peaking factor, Fg , of 2.45. The maximum local metal-water reaction was 4.96% for Unit 1 and 6.59% for Unit 2 and the total core wide metal-water reaction was less than 1.0% for all cases analyzed. Further, the clad temperature transients reached a maximum at a time when the core geometry was still amenable to cooling. To account for a 16*F initial operating temperature uncertainty, the effects of containment mini-purge automatic isolation, and steam generator tube collapse; the PCT is increased by 8, 3, and 6'F respectively, or 17*F total. This then brings the PCT to 1982*F and 2090*F for FNP Units 1 and 2, respectively. The effect of the transition core cycles is conservatively evaluated to be at most 50'F higher in calculated peak cladding temperature (Reference 2) which would yield a transition core PCT of 2032'F and 2140*F for FNP Units 1 and 2, O respectively. The transition core penalty can be accommodated by the margin to the 10 CFR 50.46, 4200*F limit. It can be concluded from the results contained in Appendix B, Section 15.4 that the large break LOCA ECCS analysis for the FNP Units 1 and 2 plants remains in compliance with the requirements of 10CFR50.46 including consideration for transition core configurations. 5.2.2 Small Break LOCA 5.2.2.1 Description of Analysis and Assumptions for 17X17 VANTAGE 5 Fuel Consistent with the logic presented in Section 5.2.1.1 for large break LOCAs, the small break loss-of-coolant accident (LOCA) was analyzed assuming a full core of VANTAGE 5 fuel to determine the peak cladding temperature. As with the large break LOCA, the methodology employed in HCAP-10444-P-A, Reference 2, for transitior.ing from Westinghouse 17x17 LOPAR fuel to 17xl? VANTAGE 5 fuel was applied to the transition to VANTAGE 5 fuel for the FNP Units 1 and 2. 4544F/687F910E'.o:50 40

i The currently approved NOTRUMP small break LOCA ECCS Evaluation Model, References 30 and 31, was utilized for a spectrum of cold leg breaks. v Appendix B, Section 15.3, includes a full description of the analysis and assumptions utili:*ac for the VANTAGE 5 ECCS small break LOCA analysis. Pertinent assumptions l'u ' %

o A core thermal power of 2 HP.

o 20% steam generator tubes pt ' h eich of three steam generators (i .e. , uniform among the loop i o Atotalcorepeakingfactor,yor4.50 o aF AH of 1.70 o A thermal design flow of 86,0C0 gpm per loop o A 8% ECCS (HHSI/LHSI) pump degradation o A 10 gpm charging /(HHSI) pump flow imbalance o Thimble plug removal o Fuel temperatures and pressures based on the improvd f mi thermal model, Reference 6 o A RCS temperature operating band of 6*F

 /7        o    A 15 second diesel generator start time o    A two line segment K(Z) curve o    A RCS pressure uncertainty of 60 psi o    102% calorimetric uncertainty for nuclear power o   Increased rod drop time of 2.7 seconds The e st limiting small break LOCA was computed from a spectrum of breaks for the upflow barrel / baffle configuration of FNP Unit 1. This break size was subsequently analyzed for the downflow barrel / baffle configuration of FNP Unit 2 (see Reference 27 for additional information regarding upflow versus downflow barrel / baffle).

Sensitivity studies performed using the NOTRUMP Small Break Evaluation Model have demonstrated that VANTAGE 5 fuel is more limiting than 0FA fuel in the calculated ECCS performance. Similar studies using the HFLASH Small Break Evaluation Model have previously shown the limiting nature of 0FA when compared to LOPAR fuel small break results. For the small break LOCA, the effect of the fuel difference is most pronounced during core uncovery periods and, therefore, shows up predominantly in the LOCTA-IV calculation in the evaluation model analysis. Consequently, the orevious conclusion drawn from 1 1 4544F/687F910515:50 41

the NFLASH studies, regarding the the fuel difference, may be extended to the

      " '"""" ''" " "  ""'~ '"'  '"' ' """"       ' ""
 'C)  that a small break LOCA analysis for a full core of Westinghouse 1"ivl7 VANTAGE 5 fuel is bounding, remains valid. On this basis, only VANTAGE 5 fuel was analyzed, since it is the most limiting of the two types of fuel (17x17 1.0 PAR and 17x17 VANTAGE 5) that would reside in the core for FNP Units 1 and 2.

5.2.2.2 Method of Analysis The methods of analysis, including codes used and assumptions, are described in detail in Appendix B, Section 15.3. 5.2.2.3 Results The results of this analysis, including tabular and plotted results of the break spectrum analyzed, are provided in Appendix B, Section 15.3. 5.2.2.4 Conclusions - The VANTAGE 5 small break LOCA analysis for FNP Units 1 and 2, utilizing the currently approved NOTRUMP Evaluation Model resulted in a peak cladding temperature (PC1) of 1805*F for Farley Unit 1 and 1711*F for Farley Unit 2 for the 3-inch diameter cold leg break. These results include a 20*F PCT increase due to T avg uncertainty. The analysis assumed a limiting small break power shape consistent with a F (Z) envelope of 2.50 at the core midplane elevation 9 and 2.333 at the top of the core. The maximum local metal-water reaction is 2.367., and the total core m;tal-water reaction is less than 1.0 percent for all cases analyzed, corresponding to less than 1.0 percent hyh, en generation. The clad temperature transients turn around at a time when the core geometry is still amenable to cooling. The limiting small break LOCA analysis presented in Appendix B, Section 15.3, shows that one high head safety injection charging pump, together with the accumulators, previde sufficient core flooding to keep the calculated peak

     -clad temperature + *11 below the required limit of 10 CFR 50.46 for the FNP Units 1 and 2.      A can also be seen that the ECCS analysis remains in 4544F/687F910515:50                          42 l

i _

compliance with all other requirements of 10 CFR 50.46 and the peak cladding temperature results are below the peak cladding temperatures calculated for 'g w) the large break LOCA. Adequate protection is therefore afforded by the ECCS in the event of a small break LOCA. 5.2.3 Transition Core Effects on Large Break LOCA and Small Break LOCA When assessing the effect of transition cores on the large break LOCA analysis, it must be determined whether the transition core can have a gr6ater calculated peac cladding temperature (PCT) than either a complete core of the 17x17 LOPAR assembly design or a complete core of the 17x17 VAN' AGE 5 design. For a given peaking factor, the only mechanism available to cause a transition core to have a greater-calculated PCT than a full core of either fuel is the possibility of flow redistribution due to fuel assembly hydraulic resistance mismatch. Hydraulic resistance mismatch will exist only for a transition core and is the only unique difference between a complete core of eithe: fuel type and the transition cere. C 5.2.3.1 Large Break LOCA !w. The large break LOCA analysis was performed with a full core of VANTAGE 5 fuel and conservatively applies the blowdown transient results to transition cores. The VANTAGE 5 differs hydraulically from the 17x17 LOPAR assembly design it replaces. The difference in the total assembly hydraulic resistance between the two designs is approximately 10% higher for VANTAGE 5. An evaluation of hydraulic mismatch of approximately 10% showed an insignificant effect on blowdown cooling during a LOCA. The SATAN-VI computer code models the crossflows between the average core flow channel (N-1 fuel assemblies) and the hot assembly flow channel (3ne fuel assembly) during blowdown. To better understand the transition core large break LOCA blowdown transient phenomena, conservative blowdown fuel clad heatup calculations have been performed to determine the clad temperature effect on the new fuel design for mixed core configurations. The effect was determined by reducing the axial flow in the hot assembly at the appropriate elevations to simulate the O effects of the transition core hydraulic resistance mismatch. In addition, G 4544F/687F910515:50 43

the Westinghouse. blowdown evaluation model was modified to account for grid p heat transfer anhancement during blowdown for this evaluation. The results of this evaluation have shown that no peak cladding temperature penalty is observed during blowdown for the mwd core. Therefore, it is not necessary to perform a blowdown calculation for the VANTAGE 5 transition core configuration because the evaluation model blowdown calculation performed for the full VANTAGE 5 core is conservative and bounding. Since the overall resistance of the two types of fuel is estenth '- the same, only the crossflows during core reflood due to Intermediate Flow Mixing grids need be evaluated. The LOCA analysis uses the BASH computer code to calculate the reflood transient, Reference 25, which utilizes the BART code, Reference 32. A detailed description of the BASH code is given in Appendix B. Fuel assembly design specif!c analyses have been performed with a version of the BART computer code, which accurately models mixed core configurations during reflood. Westinghouse transition core designs, including specific 17X17 LOPAR to VANTAGE 5 transition core cases, were analyzed. For this case, BART modeled both fuel assembly types and predicted the reduction in axial flow f rates at the appropriate elevations. As expected, the increase in hydraulic resistance for the VANTAGE 5 assembly was shown to produce a reduction in ~ reflood steam flow rate for the VANTAGE 5 fuel at mixing vane grid elevations for transition core configurations. This reduction in steam flow rate is partially offset by the fuel grid heat transfer enhancement predicted by the BART code during reflood. The various fuel assembly specific transition core analyses performed resulted in peak cladding temperature increases of up to 50'F for core axial elevations that bound the location of t S PCT. Therefore, the maximum PCT penalty possible for VANTAGE 5 fuel residing in a transition core i' .:0*F, Reference 2. Once a full core of VANTAGE 5 fuel is achieved the large break LOCA analysis will apply without the transition core penalty. 5.2.3.2 Small Break LOCA The NOTRUMP computtr code, Reference 31, is used to model the core hydraulics during a small break LOCA event. Only one core flow channel is modeled in the NOTRUMP computer code, since the core flow rate during a small break LOCA is relatively slow, providing enough time to maintain flow equilibrium between 4544F/587F910515:50 44

fuel assemblies (i.e., no crossflow). Therefore, hydraulic resistance p) mismatch is not a factor for small break LOCA. Thus, it is, not necessary perform a small break LOCA evaluation for transition cores, and it is to sufficient to reference the small break LOCA for the complete core of the JANTAGE 5 fuel design, as bounding for all transition cycles. In addition, all the small break LOCA related analyses discussed above have been analyzed or evaluated to include a control rod drop time of 2.7 seconds as is required for the 17X17 VANTAGE 5 fuel. 5.2.4 Post-LOCA Long-Term Core Cooling - ECCS Flows, Core Subcriticality and Switchover of the ECCS to Hot Leg Recirculation The implementation of VANTAGE 5 fuel at the FNP Units 1 and 2 does not affect the assumptions for decay heat, core reactivity or boron concentration for sources of water residing in the containment sump Post-LOCA. Thus, these lic. sing requirements associated with LOCA are in:1gnificantly affected by the impismentation of VANTAGE 5 fuel. Although the Reactor Coolant System total water volume increases due to the smaller diameter VANTAGE 5 fuel rod, the increase in water volume has a negligible effect on the computed sump mean boron concentration. Additionally, Westinghouse, during the specific reload cycle design, performs an independent check on core subtriticality for each fuel cycle operated at FNP Units 1 and 2. 5.2.5 Meam Generator Tube Rupture The steam generator tube rupture (SGTR) analysis performed for the FNP Units 1 and 2 FSAR was used to evaluate the radiological consequences resulting from an SGTR accident. The major factors that affect the resultant offsite radiation doses are the amount of fuel defects (level of reactor coolant coni. amination), the primary to secondary mass transfer through the ruptured tube, and the steam released from the ruptured steam generator to the atmosphere. An evaluation has been performed to determine the effect on these factors due to the change to VANTAGE 5 fuel.

 %J 4544F/687F910515:50                        45

l The SGTR analysis for FNP Units 1 and 2 is based on the conservative I g assumption of 1% fuel defects, and this assumption will not be affected by the ( change to VANTAGE 5 fuel. The primary to secondary break flow and the steam release to the atmosphere are primarily dependent upon the full power i operating conditions for the reactor coolant system and the secondary system. The core is not explicitly modeled in the SGTR analysis. Since the full power operating conditions for 2652 MHt are the same for LOPAR and VANTAGE 5 fuel, the transition to VANTAGE 5 fuel will not affect the amount of break flow or steam release to the atmosphere for an SGTR. On this basis, it is concluded that the change to VANTAGE 5 fuel will not affect the Farley SGTR analysis, and thus, the conclusions in the FSAR remain valid with the change to VANTAGE 5 fuel. 5.2.6 Containment Mass and Energy Release Analyses The containment mass and energy releases for containment integrity analyses are documented in the FNP Units 1 and 2 FSAR, Chapter 6.2. The analyses were performed to support a core power of 2652 MHt. These analyses consider both long and short term mass and energy releases for a range of postulated loss of U coolant accidents. The introduction of VANTAGE 5 fuel will not affect the results of the containment mass and energy release analyses. This is due to a number of reasons; the most significant of which is a result of the combination of a reduced pellet size, a reduced pellet to clad gap, and lower pellet and clad volumes that lead to an overall reduction in the stored energy in the core. The slight increase in the initial reactor coolant system water mass resulting from the smaller rods is insignificant when compared to the total mass released during the transient. Therefore, the transition to the VANTAGE 5 fuel will not increase the results of mass and energy analyses. The increase in reactor coolant system temperature and pressure uncertainties can also potentially affect the results of the mass and energy analyses. For the short term analyses, the initial few seconds of the transient are the most important. A decrease in the coolant temperatures, which would result by subtracting the temperature uncertainty from the initial coolant temperatures, 4544F/687F910515:50 46 __. - _.~ . _ _ _. __ .

would result in an increase in break flow rates and higher releases due to the

 ,     increased coolant density. By assuming the maximum RCS pressure plus

(,/ uncertainty the maximum increase in break flow is determined. For the long term case the energy stored in the coolant prior to the event is important, so increasing tha coolant temperatures and pressure would be conservative and result in hi s t energy releases. In the current design basis analyses presented in the Farley FSAR an uncertainty of 4*F and 30 psi were originally included. Increasing these uncertainties to i 6*F and 2 50 psi will result in a negligible increase in the break mass flow rates, much less than 1%, and an equally small increase in the energy released. These small increases are easily accommodated by the margin inherent in the input assumptions and initial conditions used in the analyses. Based on the discussion presented above it is concluded that the use of VANTAGE 5 fuel and the small increase in temperature and pressure uncertainties will not affect the results of the LOCA mass and energy releases analyses performed for FNP Units 1 and 2. 5.3 LOCA Hydraulic Forces Analysis This section addresses the vessel LOCA hydraulic forcing function analysis for the 17x17 VANTAGE 5 reload for FNP Units 1 and 2. 5.3.1 Introduction The purpose of the LOCA hydraulic forces evaluation was to demonstrate the

     -    applicability of the forces ualysis of Reference 33 to the current VANTAGE 5 reload. The evaluation considered the presence of 17x17 VANTAGE 5 fuel, an uprated power of 2T/5 MHt, and a peak steam generator tube plugging level of 20%. The hydraulic forces analysis of Reference 33 was originally performed for the FNP Unit I upflow conversion.

4544F/687F910515:50 47

5.3.2 Methodology

V The method of evaluation iniolved the utilization of sensitivities to determine an expected increase in the postulated hydraulic forces and show that the margin inherent in the Weltinghouse evaluation model assumptions are sufficient to offset the above postulated increase in LOCA hydraulic forces. The analysis of Reference 33 used the MULTIFLEX code (Reference 34) to calculate the thermal-hydraulics. The LATFORC code (Reference 34) utilizes the MULTIFLEX pressure distribution in the downcomer annulus region to determine the lateral forcing functions on the reactor vessel, core barrel and thermal shield. FORCE 2 (Reference 34) uses the pressure transient in the reactor vessel as calculated by MULTIFLEX to calculate the vertical forces on the vessel internals and core components. 5.3.3 Results The original analysis of Reference 33 used the MULTIFLEX code (Reference 34) to calculate the the: mal hydraulics of the Reactor Coolant System due to postulated 144 in' reactor vessel inlet and outlet nozzle pipe ruptures. l 'v The maximum calculated break area of Reference 35 was found to be 100 in'. Since the maximum calculated actual break area was at least 44 in' less than the analyzed value, this margin was used in part, to offset the effects of the VANTAGE 5 fuel, uprated power, and increased steam generator tube plugging level on the magnitude of the LOCA hydraulic forcing functions. By use of this margin, the original analysis of the LOCA hydraulic forces documented in reference 33 was shown to remain applicable to the FNP Units 1 and 2. O 4544F/687F910515:50 48

l. 6.0

SUMMARY

OF TECHNICAL SPECIFICATIONS CHANGES 73 l- U 6.1 Introduction t Table 6-1 presents a list of the technical specifications _ changes and the justification for the changes. The changes noted in Table 6-1 are l given in the proposed technical specifications change pages in Attachment 2. 6.2 Reactor Trip and Engineered Safety Features Setpoints for Technical Specifications. The following reactor trip and engineered safety features setpoints have been reviewed for the VANTAGE 5 safety analysis using the latest safety limits and the Westinghouse statistical setpoint methodology.

1. Overtemperature delta-T, reactor trip: 2. Overpower delta-T, reactor trip

! f] 3. Loss of flow, reactor trip l V 4. Tavg Low-Low, steam line isolation (coincident with steam flow in L two steam lines - High) and P-12 ! 5. Pressurizer pressure Low, reactor' trip

6. Pressurizer pressure Low, safety injection

7. Containment pressure High, safety injection

8. Containment pressure High-High, steam line isolation

9. Containment pressure High-High-High, containment spray The Trip Setpoints for overtemperature delta-T reactor trip and overpower delta-T reactor trip were changed for the new VANTAGE 5 core limits and uncertainty calculations. The Allowable Values for overtemperature delta-T reactor trip, overpower delta-T reactor trip, loss of flow reactor trip, Tavg Low-Low, containment pressure High-High steam line isolation, and containment pressure High-High-High containment spray were changed for the new VANTAGE 5 core limits, the uncertainty calculations and RTD Bypass Elimination. These changes are i

l O 4544F/687F910515:50 49 i

found in:the technical specifications changes for the VANTAGE 5 fuel analysis. The Allowable Values for pressurizer pressure low reactor (>y

-f~

trip, pressurizer pressure low safety injection, and containment pressure high safety injection are acceptable and did not change, [' n.s) t i l 4544F/687F910515:50 50 l

TABLE 6 1

SUMMARY

AND JUSTIFICATION FOR 1HE FWP Uhl1$ 1 AND 2 TECHhlCAL SPECIFICATIONS CHANGES FOR VANTAGE 5 FUEL ( btgt $ectim Description Justification 22 2.1 Change to core limits, these changes are a result of changes 2 5,2 8 2.2 and the OTAT ard OPAT associated with the VANT AGF-5 f uel, 2 9,2 10 Setpointe and the 0161 increased hot channet factors, future B24 2.2.1 Basis OPAT and Low flow design and operational considerations, B25 Attowable Values and the implementation of RfD Bypass 3/4 3 10 3/4.3 Elimination for Unit 2. B 2-1 2.1.1 Basis Addition of the WRB 1 This change reflects the new Dh8 8 2 3,8 2 6 2.2.1 Basis and WRB 2 correlations correlations used in analyses. B 3/4 2 1 3/4.2 Basis , these correlations are supplemented by B 3/4 2 2 the W-3 corret: tion. B 3/4 2-4 8 3/4 2 5 8 3/4 4 1 3/4.4 Bas's 3/4 1-4 3.1.1.3 Revision of the MTC This change is to attow flexibility durths core design. 1he effect of this increase on the safety analysis has been considered. 3/4 1 19 3.1.3.4 Revi.ed rod drop time This change is a result of mechanical to less than or equal changes associated with the VANT AGE 5 to 2.7 seconds f ue t . The effect of this increase on the safety analysis has been considered. B 2-1 2.1.1 Basis Fgg ord Fo(Z) changes these changes are made to give the 3/4 2-4 3/4.2.2 plant core design flexibility and 3/4 2-7 3/4.2.2 also as a result of changes g associated with the VANTAGE-5 fuet. 3/4 2 8 3/4.2.3 B 3/4 2 1 3/4.2 Basis The effect of these increases on the safety analysis has been considered. 2-5 2.2 Dh8 parameter changes these changes are made to account for 3/4 2 14 3/4.2.5 the use of RTDP and to give the plant 3/4 2-15 operating f;enibility. B 3/4 2 5 3/4.2.5 Basis 3/4 3-10 3/4.3.1 OTAT Response ilme accounts for RTD Bypass Elimination in Unit 2. 3/4 3-26 3/4.3.2 contairvnent pressure These changes are made to account for 3/4 3-27 High 2 and High 3 Astow, instrunent uncertainty to prevent 3/4 3 28 able Values and Low Low changes to ESF setpoints; and to Teve Attowable Values account for RfD Bypass Elimination in Unit 2. 3/4 4-2 3.4.1.2 Change to pump operable this change attows operational fleri-B 3/4 4 1 3/4.4.1 Basis requirement bility. The effects of this change have been considered in the cafety analysis. 6-19 6.9.1.11 Radial Peaking Factor This change is made to provide Limit Report adninistrative flexibility. 3/4 3-10 3/4.3.1 PR heg Rate and this deletion makes these diverse 3/4 3-30 3/4.3.2 Hi Stm Flow w/Lolo functions consistent with other diverse favg Response Time functions. Pages B 2 4, r 2-5, and 3/4 3 28 are for Unit 2 only.

7.0 REFERENCES

1. Davidson, S. L. and Iorii, J. A., " Reference Core Report - 17x17 Optimized Fe ' Assembly," HCAP-9500-A, May 1982.

2. Davidson, S. L. and Kramer, H. R., (Eds.) " Reference Core Report VANTAGE 5 Fuel Assembly," HCAP-10444-P-A, September 1985.

3. Davidson, S. L. (Ed.), et al., " Extended Burnup Evaluation of Westinghouse Fuel," HCAP-10125-P-A, December 1985.

4. Davidson, S. L. (Ed.), et al., "Hestinghouse Reload Safety Evaluation Methodology," HCAP-9272-P-A, July 1985.

5. Miller, J. V., " Improved Analytical Models Used in Westinghouse Fuel Rod Design Computations," HCAP-8720, October 1976.

6. Heiner, R. A., et al., " Improved Fuel Performance Models for Westinghouse fuel Rod Design and Safety Evaluations," HCAP-10851-P-A, August 1988.

7. Skaritka, J. (Ed.), " Fuel Rod Bow Evaluation," HCAP-8691, Revision 1, July 1979.

8. Davidson, S. L., and Iorii, J. A. (Eds.), " Verification Testing and Analyses of the 17x17 Optimized Fuel Assembly," HCAP-9401-P-A, August 1981.

9. Nguyen, T. Q. , et al . , " Qualification of the PH0ENIX-P-ANC Nuclear Design System for Pressurized Water Reactor Cores, HCAP-ll596-A-A, June 1988.

10. Davidson, S. L. (Ed.), et al., "ANC: Westinghouse Advanced Nodal Computer Code," HCAP-10965-P-A, September 1986.

11. Letter from E. P. Rahe (H) to Miller (NRC) dated March 19, 1982, NS-EPR-2573, HCAP-9500 and WCAPS-9401/9402 NRC SER Mixed Core Compatibility Items.

4544F/687F910515:50 52

i

12. Letter from C. O. Thomas (NRC) to Rahe (H) " Supplemental Acceptance No.

N 2 for Referencing Topical Report WCAP-9500," January 1983. I b l

13. Friedland, A.J., and Ray, S., " Improved THINC IV Modeling for PHR Core Design," HCAP-12330-P, August 1989,

14. Hochreiter, L. E., and Chelemer, H., " Application of the THINC-IV Program to PHR Desiga," HCAP-8054 (Proprietary) and WCAP-8195 (Non-proprietary).

September.1973.

15. Friedland, A. J. and Ray, S., " Revised Thermal Design Procedure,"

HCAP-11397-P-A, April 1989.

16. Motley, F. E. , et al . , "New Westinghouse Correlation HRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids," HCAP-8762-P, July 1984.

17. Tong, L. S., " Critical Heat Fluxes in Rod Bundles, Two Phase Flow and A Heat Transfer in Rod Bundles," Annual Hinter Meeting ASME, November 1968,

       .p. 3146.

18. Tong, L. S. , " Boiling Crisis and Critical Heat Flux. . . ." AEC Office of Information Services, TID-25887, 1972.

19. Letter from A. C. Thadani (NRC) to H. J. Johnson (Hestinghouse), Jan. 31, 1989,

Subject:

Acceptance for Referencing of Licensing Topical Report, WCAP-9226-P/HCAP-9227-NP, " Reactor Core Response to Excessive Secondary Steam Releases."

20. Motley, F. E., and Cadek, F. F., "DNB Test Results for R-Grid Thimble Cold Hall Cells," HCAP-7695-L Addendum 1, October 1972.

21. Hill, K. H. , Motley, F. E. , Cadek, F. F. , and Henzel, A. H. , 'Effect of 17x17 Fuel Assembly Geometry on DNB," HCAP-8296, March 1974.

22. Moomau, H. H., "Hestinghouse Revised Thermcl Design Procedure Instrument Uncertainty Methodology for Alabama Power Farley 1 & 2 Nuclear Power Stations (For RTD Bypass Loop)," HCAP-12769 (Proprietary), March 1991.

4544F/687F910515:50 53

~ .

I i 1 I

23. Moomau, H. H., " Westinghouse Revised Thermal Design Procedure Instrument j

 ,q       . Uncertainty Methodology for Alabama Power Farley 1 & 2 Nuclear Power      l k/        Stations (For RTD Bypass Loop Elimination)," HCAP-12771 (Proprietary),     I March 1991.-                                                               I l

l

24. Schueren, P. and McAtee, K. R., " Extension of Methodology for Ca Pulating '

Transition Core DNBR Penalties," HCAP-il837-P-A, January 1990, l

25. Kabadi, J.N., et al., "The 1981 Version of the Westinghouse ECCS ;

Evaluation Model Using the BASH Code," HCAP-10266-P-A Revision with Addenda (Proprietary), March 1987. l I

26. Eicheldinger, C., "Hestinghouse ECCS Evaluation Model - 1981 Version," ;

HCAP-9220-P-A (Proprietary), HCAP-9221 (Non-Proprietary), Revision 1, February 1982. l 27. Johnson, N.J. and Thompson, C.M., " Westinghouse Emergency Core Cooling System Evaluation Model Modified October 1975 Version," HCAP-9168 (Proprietary) and NCAP-9150 (Non-Proprietary), August 1977. l 28. Chiou, J.S., et al., "Models For PHR Reflood Calculations Using the BART Code," HCAP-10062 (Proprietary), March 1982.

29. Besspiata, J.J., et al., "The 1981 Version of the Westinghouse ECCS Evaluation Model Using the BASH Code, Power Shape Sensitivity Studies,"

HCAP-10266-P-A Revision 2 Addendum 1 (Proprietary), December 15, 1987.

30. Lee, N., et al., " Westinghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code," HCAP-10054-P-A (Proprietary) and WCAP-10081-A (Non-Proprietary), August 1985.

31. Meyer, P.E., "NOTRUMP - A Nodal Transient Small Break and General Network Code," HCAP-10079-P-A (Proprietary) and NCAP-10080-A (Non-Proprietary),

August 1985. j 32. Young, M.Y., et al., "BART-A1: A Computer Code for the Best Estimate

   '~      Analysis of Reflood Transients," HCAP-9561-P-A (Proprietary), March 1984.

l 4544F/687F910515:50 54

33. J. C.' Miller, "Upflow Modification Safety Evaluation Report -

Revision 1," ALA-82-678, December 17, 1982, i

34. Takeuchi, K., et al . , -"MULTIFLEX 1.0 - A FORTRAN IV Computer Program for Analyzing Thermal-Hydraulic Structure System Dynamics", HCAP-870P-P-A (Proprietary) and NCAP-8709-A (Non-Proprietary), September 1977.

35. Bogard, W. T., " Dynamic Analysis of Reactor Pressure Vessal for Postulated Loss-Of-Coolant Accidents; Joseph M. Farley Urits 1 and 2,"

WCAP-8749, October 1976.

36. Butler, J. C., and Love, D. S., "Steamline Break Mass / Energy Releases for Equipment Environmental Qualification Outside Containment," WCAP-10961-P, October 1985.

I l 4544F/687F910515:50 55

 -i 4

ADnendix A Joseph M. Farley Nuclear Plant Units 1 and 2 Request for Technical Specifications Changes Non-LOCA Accident Analysis l 0

   ,-is                                          FNP-RTSR-15
 ;     e L) -

15.0 ACCIDENT ANALYSES TABLE OF CONTENTS

                                                                                                   .P.,AE A.1 INTRODUCTION                                                                        A-15.1-1 A.2 NON-LOCA ACCIDENT ANALYSES                                                          A-15.1-3 15.1     ACCIDENT ANALYSIS                                                              A-15.1-4 15.1.1     Optimi ation of Control    Systems............                     A-15.1-4 15.1.2     Initial Power Conditions Assumed in. . . . . . . .                 A-15.1-4 the Accident Analyses 15.1.2.1   Power Distribution.........................                        A-15.1-5 j -sg             15.1.3     Trip Points and Time Delays to frip. . . . . . . .                 A-15.1-6 Assumed in Accident Analyses

(- / 15.1.4 Instrumentation Drift and Calorimetric..... A-15.1-6 Errors 15.1.5 Rod Cluster Control Assembly Insertion. . . . . A-15.1-7 Characteristics 15.1.6 Reactivity Coefficients.................... A-15.1-8 15.1.7 Fission Product Inventories....... ........ A-15.1-8 15.1.8 Residual Decay Heat...... ................. A-15.1-8 15.1.9 tc..iputer Codes Utilized.................... A-15.1-8 15.1.9.1 FACTRAN Compute r Code . . . . . . . . . . . . . . . . . . . . . . A-15.1-9 15.1.9.2 LOFTRAN Computer Code...................... A-15.1-9 15.1.9.3 ANC Computer Code...... ................... A-15.1-10 15.1.9.4 TWINKLE Computer Code...................... A-15.1-10 15.1.9.5 THINC Computer Code........................ A-15.1-ll

                  '15.1.10    References. .................. ............                        A-15.1-12 V( O -

A-1 m _ _

15.2 CONDITION II - INCIDENTS Of MODERATE FREQUENCY........ A-15.2-1

 . (J 15.2.1     Uncontrolled RCCA Bank Withdrawal from.....                                            A-15.2-2 a Subcritical Condition 15.2.1.1   Identification of Causes and Accident. . . . . . A-15.2-2 Description 15.2.1.2   Analysis of Effects and Consequences.. ... . .                                         A-15.2-4 15.2.1.3   Conclusions................................                                            A-15.2-7 15.2.2     Uncontrolled RCCA Bank Withdrawal at Power.                                            A-15.2-8 15.2.2.1   Identification of Causes and Accident......                                            A-15.2-8 Description 15.2.2.2   Analysis of Effects and Consequences.......                                            A-15.2-10 15.2.2.3   C o n cl u s i o n s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A-15.2-12 15.2.3     RCCA Misalignment..........................                                            A-15.2-12 15.2.3~1   Identification of Causes and Accident......                                            A-15.2-12 Description
              !5.2.3.2   Analysis of Effects and Consequences....... 'A-15.2-14 15.2.3.3   Co n cl u s i on s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A-15.2-18 15.2.4     Uncontrolled Boron Dilution................                                            A-15.2-18 15.2.4.1   Identification of Causes and Accident......                                            A-15.2-18 Description 15.2.4.2   Analysis of Effects and Consequences.......                                            A-15.2-19 15.2.4.3   C o n c l u s i o n s . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . . . A-15.2-24 15.2.5     Partial loss of Forced Reactor Coolant Flow A-15.2-25 15.2.5.1   Identification of Causes and Accident...... A-15.2-25 Description 15.2.5.2   Analysis of Effects and Consequences. . . . . . .                                      A-15.2-26 15.'2.5.3  Conclusions...............                          ................                   A-15.2 '

15.2.6 .Startup of an Inactive Reactor Coolant Loop A-15.2-28 15.2.6.1 Identification of Causes and Accident...... A-15.2-28 Description L 15.2.6.2 Analysis of Effects and Consequences....... A-15.2-29 15.2.6.3 Conclusions................................ A-15.2-30 g

 'V L

A-2

15.2.7 Loss of. External Electrical Load and/or. .. . A-15.2-31

                         . Turbine Trip

{}.

'- '          15.2.7.1     Identification of Causes and Accident......                                        A-15.2-31 Description 15.2.7.2    Analysis of Effects and Consequences.......                                         A-15.2-32 15.2.7.3    C on cl u s i o n s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A-15.2-36 7

15.2.8 Loss of Normal Feedwater................... A-15.2-37 15.2.8.1 Identification of Causes and Accident...... A-15.2-37 Description 15.2.8.2 Analysis of Effects and Consequences....... A-15.2-39 15.2.8.3 Conclusions................................ A-15.2-41 15.2.9 Loss of All AC Power to the Station........ A-15.2-42 Auxiliaries 15.2.9.1 Identification of Causes and Accident...... A-15.2-42 Description 15.2.9.2 Analysis of Effects and Consequences. ..... . a-15.2-44

             -15.2.9.3    Conclusions................................                                         A-15.2-46 15.2.10     Excessive Heat Removal Due to Feedwater....                                         A-15.2-46
  ~

System Malfunctions 15.2.10.1 Identification of Causes and Accident...... A-15.2-46 Description 15.2.10.2 Analysis of Effects and Consequences....... A-15.2-47 15.2.10.3 Conclusions................................ A-15.2-51 15.2.11 Excessive Load Increase Incident........... A-15.2-51 15.2.11.1' Identification of Causes and Accident...... A-15.2-51 Description 15.2.11.2 Analysis of Effects and Consequences. .. .... A-15.2-53 15.2.11.3 Conclusions................................ A-15.2-55 15.2.12 Accidental Depressurization of the RCS..... A-15.2-55 15.2.12.1 Identification of Causes and Accident...... A-15.2-55 Description 15.2.12.2 Analysis of Effects and Consequences. . . . . . . . A-15.2-56 15.2.12.3 Conclusions................................ A-15.2-57

A-3

_ _ _ -- _. - . _ m _ 15.2.13 Accidental Depressurization of the......... A-15.2-57 Main Steam-System

 ]\'            15.2.13.1  Identification of Causes and Accident......        A-15.2-57
                          -Description 15.2.13.2 Analysis of Effects and Consequences.......         A-15.2-59 15.2.13.3  Conclusions................................        A-15.2-61 15.2.14    Inadvertent Operation of ECCS During.......        A-15.2-61 Power Operation 15.2.14.1  Identification of Causes and Accident......        A-15.2-61 Description 15.2.14.2 Analysis of Effects and Consequences.......         A-15.2-63 15.2.14.3  Conclusions................................        A-15.2-65 15.2,15    References.................................        A-15.2-66 15.3    CONDITION 111 - INFREQUENT FAULTS.....................        A-15.3-1 15.3.1-    Loss of Reactor Coolant from Sma11.........        A-15.3-1 Ruptured Pipes or from Cracks in large IN                      Pipes Which Actuate Emergency Core
 %)                        Cooling System
               ~15.3.2     Minor Secondary System Pipe Breaks.........       A-15.3-2 15.3.2.1   Identification of Causes and Accident......       A-15.3-2 Description 15.3.2.2-  Analysis of Effects and Consequences.......       A-15.3-2 15.3.2.3   Conclusions................................       A-15.3-2
               -15.3.3     Inadvertent Loading of a Fuel..............       A-15.3-2
                          - Assembly into an Improper Position 15.3.4     Complete Loss of Forced Reactor Coolant....       A-15.3-3 Flow 15.3.4.1   Identification of Causes and Accident. . . . . . A-15.3-3 Description 15.3.4.2   Method of Analysis.........................       A-15.3-4 15.3.4.3   Results....................................       A-15.3-5 15.3.4.4   Conclusions.......... .....................       A-15.3-5 O

A-4 y

(T -15.3.5 Waste Gas Decay _ Tank Rupture............ .. A-15.3-5 15.3.6. Single RCCA Withdrawal at Full Power....... A-15.3-6

15.3.6.1 Accident Description....................... A-15.3-6 15.3.6.2 Conclusions................................ A-15.3-8 15.3.7 References................................. A-15.3-9 15.4 CONDITION IV - LIMITING FAULTS........................ A-15.4-1 15.4.1 - Major Reactor Coolant System Pipe.......... A-15.4-1 Ruptures (Loss-of-Coolant Accident) 15.4.2 Major Secondary System Pipe Rupture. . . . . . . . A-15.4-2 15.4.2.1 Rupture of Main Steamline.................. A-15.4-2 15.4.2.1.1 Identific.ation of Causes and Accident...... A-15.4-2 Description 15.4.2.1.2 Analysis of Effects and Consequences....... A-15.4-5 15.4.2.1.3 Conclusion................................. A-15.4-ll D 15.4.2.2- Major Rupture of a Main Feedwater Pipe..... A-15.4-12 (Y 15.4.4 Single Reactor Coolant Pump Locked Rotor.. . A-15.4-14 15.4.4.1 Identification of Causes and Accident...... A-15.4-14 , Description

              !!.4.4.2'      Analysis of Effects and Consequences.......                                      A-15.4-15
 -            15.4.4.3 C o n c l u :, i o n t . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . A-15.4-19 i              15.4.5         Fuel Handling Accident.....................                                      A-15.4-19 15.4.6         Rupture of. a Control Rod Drive. . . . . . . . . . . .                           A-15.4-20 Mechanism (CRDM) Housing (RCCA Ejection) p              15.4.6.1-       Identification of Causes and Accident.....                                      A-15.4-20 Description 15.4.6.2       Analysis of Effects and Consequences......                                       A-15.4-24 l              15.4.6.3       Conclusions...............................                                       A-15.4-31 15.4.7         References..............                  .................                      A-15.4-32 O

A-5

[h L %.) FNP-RTSR-15 LIST OF TABLES 15.1-1 Nuclear Steam Supply System Power Ratings 15.1-2A Summary of Initial Conditions and Computer Codes Used

                            -15.1-2B            Nominal Values of Pertinent Plant Parameters Used in the VANTAGE 5 Accident Analyses 15.1-3            Trip Points and Time Delays to Trip Assumed in Accident Analyses 15.2-1            Time Sequence of Events for Condition II Events 15.2-2            Summary of Boron Oilution Analysis Results and Analysis Assumptions 15.3-1            Time Sequence of Events for Condition III Events 15'.4-5           Time Sequence of Events for Condition IV Events
                            -15.4-12            Parameters Used in the Analysis of the Rod ~ Cluster Control Assembly Ejection Accident 15.4-25           Summary of Results for the Locked Rotor Transient O

A-6

[l w/ FNP-RTSR-15 LIST OF FIGURES 15,1-1A Overpower and Overtemperature AT Protection 15.1-1B Overtemperature AT Setpoint F(AI) Penalty l 15,1-2 Rod Position Versus Time on Reactor Trip l 15,1-3 Normalized RCCA Reactivity Worth Versus Rod Insertion ! 15,1-4 Normalized RCCA Bank Reactivity Worth Versus Time After Trip l 15.1-5 Doppler Power Coefficient Used in Accident Analysis l [ 15.1-6 1979 ANS Decay Heat h 15.1-7 Fuel Rod Cross-Section I 15.2-1 Startup from Subcritical for 3-Loop,17 x 17 Plant, Nuclear Power ) l 15.2-2 Startup from Subtritical for 3-Loop,17 x 17 Plant, FACTRAN Heat Flux

l. 15,2-3 Uncontrolled Rod Withdrawal from a Subtritical Condition -

Temperature Versus Time l 15.2-4 Transient Response for Uncontrolled Rod Withdrawal from Full Power Terminated by High Neutron Flux Trip 15.2-5 Transient Response for Uncontrolled Rod Withdrawal from Full Power Terminated by High Neutron Flux Trip V A-7 l

15.2-6 Transient Response _ for Uncontrolled Rod Withdrawal from Full Power l M Terminated by Overtemperature AT Trip -() ~ i?.2-7 Transient Re~sponse for Uncontrolled Rod Withdrawal from Full Power Terminated.by Overtemperature AT Trip . l 1 15.2-8 Effect of Reactivity Insertion Rate on Minimum DNBR for a Rod Withdrawal Accident from 100-Percent Power 15.279 Effect of Reactivity Insertion Rate on Minimum DNBR for a Rod Withdrawal Accident from 60-Percent Power 15,2-10 Effect of Reactivity Insertion Rate on Minimum DNBR for a Rod Withdrawal Accident from a 10-Percent Power 15.2-11 Transient Response to a Dropped Rod Cluster Control Assembly 15.2-12 All Loops initially Operating, One Loop Coasting Down - Flow Coastdowns Versus Time 15.2-13 All Loops initially Operating, One Loop Coasting Down - Nuclear Power Versus Time 15.2-14 All loops Initially Operating, One Loop Coasting Down - Pressurizer Pressure Versus Time 15.2-15 All Loops Initially Operating, One Loop Coasting Down - Average Channel-Heat Flux Versus Time 15.2-16 All Loc)s Initially Operating, One Loop Coasting Down - Hot Channel Heat Flux Versus Time 15.2'17 All Loops Initially Operating, One Loop Coasting Down - DNBR Versus Time O A-8

                                    .          _._,   , . _ _.~  ,_          ._   __ .

             ..-.u..e          36 (   ,A4.. _AJ-             -    4 -

m - d 15.2-18A Startup of an inactive Reactor Coolant Loop - i/ x 17 Core " { ~G 15.2-18B Startup of an Inactive Reactor Coolant loop - 17 x 17 Core 15.2-18C Startup of an Inactive Reactor Coolant Loop - 17 x 17 Core 15.2-19A Loss of Load Accident With Pressurizer Spray and Power-Operated Relief Valve - BOL 15.2-19E Loss of Load Accident With Pressurizer Spray and Power-0perated Relief Valve - BOL 15.2-19C Loss of Load Accident With Pressurizer Spray and Power-Operated Relief Valve - BOL 15.2-20 Loss of Load Accident With Pressurizer Spray and Power-0perated Relief Valve - BOL

!G    15.2-21A        Loss of Load Accident With Pressurizer Spray and Power-0perated Relief Valve - E0L 15.2-218        Loss of Load Accident With Pressurizer Spray and Power-0perated Relief Valve - EOL 15.2-21C        Loss of Load Accident With Pressurizer Spray and Power-0perated Relief Valve - E0L 15.2-22         Loss of Load Accident With Pressurizer Spray and Power-0perated-Relief Valve - E0L L15.2-23A        Loss of Load Accident Without Pressurizer Spray and Power-0perated Relief Valve - BOL 15.2-23B        Loss of Load Accident Without Pressurizer Spray and Power-0perated Relief Valve - BOL O

A-9

i i I I 15.2-23C Loss'of Load' Accident Without Pressurizer Spray and Power-Operated

  }I a

Relief Valve - BOL 15.2-24 -Loss of Load Accident Without Pressurizer Spray and Power-Operated Relief Valve'- BOL 15.2-25A Loss of Load Accident Without Pressurizer Spray'and Power-Operated Relief Valve - EOL 15.2-25B Loss of Load Accident Without Pressurizer Spray and Power-0perated Relief Valve - E0L j

.)

i 15.2-25C Loss of Load Accident Without Pressurizer Spray and Power-Operated Relief Valve - EOL 15.2 Loss of Load Accident Without Pressurizer Spray and Power-Operated Relief. Valve - E0L 15.2-27A Loss of Normal Feedwater - Pressurizer Pressure and Level Versus d[' \ Time 1 l 15.2-278 Loss of Normal Feedwater - Nuclear Power and Core Heat Flux Versus Time l 15.2-27C Loss of Normal Feedwater - Loop 1, 2, and 3 Hot and Cold Leg L Temperature Versus Time l l-l 15.2-270 Loss of Normal Feedwater - Steam Generator Pressure and Mass-Versus Time g 15.2-27E Loss of All AC Power to the Station Auxiliaries - Pressurizer Pressure and Level Versus Time 15.2-27F Loss of All AC Power to the Station Auxiliaries - Nuclear Power and Core Heat Flux Versus Time A-10 t l-

15.2-27G. Loss of All AC Power to the Station Auxiliaries - Loop 1, 2, and 3

                   -Hot and Cold Leg Temperature Versus Time U                                                                                       j 15.2-27H-   Loss of All AC Power to the Station Auxiliaries - Steam Generator       <

Pressure and Mass Versus Time i 15.2-2GA Feedwater Control Valve Malfunction - Nuclear Power and Core Heat , Flux Versus Time 1 l 15.2-28B Feedwater Control Valve M:lfunction - Pressurizer Pressure and- ) Loop AT Versus Time 15.2-28C Feedwater Control Valve Malfunction - Core Average Temperature and . DNBR Versus Time 15.-2-29 Excessive Load Increase Without Control Action - BOL - DNBR, I Nuclear Power, and Pressurizer Pressure'as a Function of Time jQ 15.2-30 . Excessive Load increase Without Control Action - BOL - T avg and a V Pressurizer Volume as a Function of Time 15.2-31 Excessive Load Increase Without Control Action - E0L - DNBR,-

                  ~ Nuclear Power, and Pressurizer Pressure as a Function of-Time 15.2-32     Excessive Load Increase Without Control Action - EOL - T avg   and Pressurizer Volume as a Function of Time 15.2-33     Excessive Load Increase With Reactor Control - BOL -- DNBR, Nuclear Power, and Pressurizer Pressure as a Function of Time 15.2-34     Excessive Load increase With Reactor Control - BOL - T avg   and Pressurizer Volume as a Function of Time 15.2-35     Excessive Load increase With Reactor Control - E0L - DNBR, Nuclear Power, and Pressurizer Pressure as a Function of Time
   ?

A-ll

15.2-36 Excessive-Load increase With Reactor Control - E0L - Tavg and

      ,br            Pressurizer Volume as a Function of Time
    -V.

15.2-37 Accidental RCS Depressurization - Nuclear Power Versus Time 15.2-38 Accidental RCS Depressurization - T avg and Pressurizer Pressure Versus Time t 15.2-39 Accidental RCS Depressurization - (NBR Versus Time 15.2-40A Variation of K,7f With Core Temperature t-15.2-40B Doppler Power Feedback l 15.2-41 Safety Injection Flow Rate Versus RCS Pressure 15.2-42A Steamline Brr.ak Transient Credible Break, failed Open Valve [ O, 15.2-42B Steamline Break Trensient Credible Break, Failed Open Valve Q l 15.2-42C Steamline Break Transient Credible Break, Failed Open Valve l' [

           -15,2-43  Spurious Actuation of the SI System - Pressurizer Pressure and l

Nuclear Power Versus Time l l 15.2-44 Spurious Actuation of the SI System - T avg and DNBR Versus Time l-15.2-45 Spurious Actuation of the SI System - Steam Flow and Pressurizer l' Water Volume Versus Time ! 15.3-20' All- Loops Operat ing, All Loops Coasting Down - Core Flow -Versus l Time 15.3-21 All Loops Operating, All Loops Coasting Down - Nuclear Power Versus Time A-12 l l

15.3-22 All Loops Operating, All loops Coasting Down - Pressurizer M Pressure Versus Time &S 15.3-23 All Loops Operating, All Loops Coasting Down - Average Channel Heat Flux Versus Time 15.3-24 All Loops Operating, All Loops Coasting Down - Hot Channel' Heat Fluy Versus Time i 15.3-25 All Loops Operating, All Loops Coasting Down - DNBR Versus Time j

     -IL.4-28A     Steamline Break Transient With Offsite Power, 1.061 Square Feet,  4 Double-Ended Rupture                                              I i

15. 4- 28B Steamline Break Transient With Offsite Power, 1.061 Square Feet, Double-Ended Rupture 15.4-28C Steamline Break Transient With Offsite Power,1.061 :e Feet,

(]- Double-Ended Rupture V 15.4-29A Steamline Break Transient With Offsite Power, 1.061 Square Feet, Double-Ended Rupture 15.4-29B Steemline Break Transient With Offsite Power, 1.061 Square Feet, Double-Ended Rupture 15.4-29C Cteamline Break Transient With Offsite Power, 1.061 Square Feet, Doutle-Ended Rupture 15.4-30A Steamline Break Transient Without Offsite Power, 1.061 Square Feet, Double-Ended Rupture 15.4-30B Steamline Break Transient Without Offsite Power, 1.061 Square Feet, Double-Ended Rupture e A-13

15.4-30C Steamline Break Transient Without Offsite Power, 1.06) Square f eet, Double-Ended Rupture 15.4-31A Steamline Break Transient Without Offsite Power, 1.061 Squara Feet. Double-Ended Rupture 15.4-31B Steamline Break Transient Without Offsite Power, 1.061 Square feet Double-Ended Rupture 15.4-31C Steamline Break Transient Without Offsite rawer, 1.061 Square Feet, Double-Ended Rupture 15.4-33 All Loaps initially Operating, One Locked Rotor 15.4-34 All Loops initially Operating, One Locked Rotor 15.4-35 All Loops initially Operating, One Locked Rotor 15.4-36 All loops initially Operating, One Locked Rotor 15.4-37 All Loops initially Operating, One Locked Rotor 15.4-38 All Loops initially Operating, One Locked Rotor 15.4-40 Nuclear Power Transient BOL Hot-Full-Power Rod Ejection Accident 15.4-41 Hotspot fuel and Clad Temperature Versus Time - BOL Hot-Full-Power Rod Ejection Accident 15.4-42 Nuclear Power Transient E0L Hot-Zero-Power Rod Ejection ccident 15.4-43 Hotspot Fuel and Clad Temperature Versus Time - EOL Hot-Zero-Power Rod Ejection Accident A-14

1

 ;                                               Appendix A Nan-LOCA Accident Analysis f

A.1 INTRODU.C1103 Appendix A addresses the impact of the complete transition of the Joseph M. Farley Nuclear Plant (FNP) Units 1 and 2 from Westinghouse 17x17 LOPAR fuel to Westinghouse 17x17 VANTAGE 5 fuel on the FSAR Chapter 15 non-LOCA accident , analyses. Section 15.1.9 of this report discusses the methods used for the accident evaluation. FNP's licensing basis includes the analyses or evaluations of the non-LOCA accidents as listed below. The analyses or evaluations of all non-LOCA accidents performed to determine the impact of the VANTAGE 5 fuel transition are documented in this report. The specific design features associated with the VANTAGE 5 fuel and the modified safety analysis assumptions considered in the non-LOCA safety analyses are described in the Safety Assessment (Attachment 4). ACCIDENTS ANALYZED The following transients affected by the VANTAGE 5 fuel design features or modified safety analysis assumptions were reanalyzed.

1. Uncontrolled rod cluster control assembly bank withdrawal from a subcritical condition (see Section 15.2.1)

2. Uncontrolled rod cluster control asseinbly bank withdrawal at power (see Section 15.2.2)

3. Rod cluster control assembly misalignment (see Section 15.2.3)

4. Uncontrolled boron dilt. tion (see Section 15.2.4)

5. Partial loss of forced reactor coolant flow (see Section 15.2.5)

6. Startup of an inactive reactor coolant loop (see Section 15.2.6)

A-15.1-1

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

l

7. Loss of external- electrical load and/or turbine trip (see Section 15.2.7)

8. Loss of normal feedwater (see Section 15.2.8)

9. Lnu of all offsite power to the station auxiliaries (see 9 ) j 15.2.9)

10. Excessive heat removal oue to feedwater system malfunctions (see i Section 15.2.10) l I

11. Excessive load increase incident (see Section 15.2.11) l i

i

12. Accidental depressurization of the reactor coolant system (see l l

Section 15.2.12)

13. Accidental depressurization of the main steam system (see Section 15.2.13)

14. Inadvertent operation of the emergency core cooling system :

during power operation (see Section 15.2.14)

15. Complete loss of forced reactor coolant flow (see Section 15.3.4)

16. Single rod cluster control assembly withdrawal at full power (see Section 15.3.6)

17. Rupture of a steamline (see Section 15.4.2.1)

18. Single reactor coolant pump locked rotor (see Section 15.4.4) ,

19. Rupture of a control rod drive mechanism housing (RCCA Ejection)

(see Section 15.4.6) A-15.1-2

For each of these events (items 1 through 19), an analysis description em similar in form to the FNP FSAR is presented. All applicable VANTAGE 5 b fuel features and safety analysis assumptions, as discussed elsewhere in this report, have been incorporated inte these analyses. ACCIDENTS EVALVATQ The following non-LOCA accidents were evaluated with respect to the VANTAGE 5 fuel design features or modified safety analysis assumptions.

1. Minor secondary system pipe breaks (see Section 15.3.2)

2. Inar ertent loading of a fuel assembly into an improper position (see Section 15.3.3)

3. Rupture of a main feedwater line (see Section 15.4.2.2)

(3 For each of these events (items 1 through 3), an evaluation of the effect of the VANTAGE 5 fuel design features and safety analysis assumptions, as described elsewhere in this report, is tresented. A.2 NON-LOCA ACCIDENT ANALYSIS The remainder of Appendix B will follow the general format of the FNP FSAR. However, Section 15.1.9 has been revised to incorporate revised methodology. Specifically, analyses using the BLK0VT, MARVEL, and PHOENIX codes have been reanalyted with the LOFTRAN code. Analyses using the LEOPARD and TURTLE codes have been reanalyzed using the ANC code. The Uncontrolled RCCA Bank Withdrawal from a Subcritical event which was previously analyzed using the WIT-6 code was reanalyzed using the multidimensional spatial neutronids TWINKLE code. Refer to the appropriate sections for a discussion of each of the previously ,Q listed accidents. (> A-15.1-3

15.1 ACCIDEl4T ANALYSES kne of the VAf1TAGE 5 fuel design features or modified safety analysis assumptions as described elsewhere in this report affect the classification of transients discussed in this section. 15.1.1 OPTIMlZAT10fl 0F C0f4 TROL SYSTEMS fione of the VANTAGE 5 fuel design features or modified safety analysis assumptions as discussed elsewhere in this report affect this section. 15.1.2 IfllTIAL POWER C0!1DIT10NS ASSUMED lil THE ACCIDEf4T Af1ALYSES Table 15.1-1 lists the principal power rating values assumed in analyses performed in this chapter. The power rating values listed in Table 15.1-1 are based on the design nuclear steam supply system (NSSS) thermal power output which includes the thermal power generated by the reactor coolant pumps (RCPs). Although the initial conditions encloyed in tbe Acciuent analyses are conservative to bound a possible future plant uprating (2785 MWt), the licensing value of IISSS power rating remains 2660 MWt. For most accidents which are Df4B limited, nominal values of

ne initial conditions are as:umed. The uncertainty allowances on power, temperature, pressure, and RCS flow are included on a statistical basis and are included in the limit DNBR value, as described in Reference 1. This procedure is known as the Revised Thermal Design Procedure (RTDP). For accidents analyses which are not DNB limited, or for which RTDP is not employed, the initial conditions are obtained by applying the maximum steady-state errors to rated values (this procedure is commonly known as Standard Thermal Design Procedure or STDP).

The following sutdy-state errors are considered in the analyses. A. Core power - 2 percent allowance for calorimetric error (note that . this error is conservatively applied in the positive direction in non-LOCA aceident analysesj. A-15.1-4

l 1 B. Averaae RIS temperature - 16'F allowance dead band and system m9asurement error. C ! C. Pressurizer pressure - 150 psi allowance for steady-state l

             -fluctuations and measurement errors.                                                      l i

Accidents employing RTDP assume a minimum measured flow (MMF); accidents employing STDP assume a thermal design flow (TDF). In addition to being the flow used in the DNB analysis for RTDP methodology, the MMF is specified in the Technical Specifications as the flow that must be confirmed or exceeded by the flow measurements obtained during plant startup. The TDF equals the MMF minus the plant flow measurement uncertainty. Table 15.1-2A summarizes the initial conditions and computer codes used in the accident analyses. The values of other pertinent plant parameters used in the accident analyses are given in Table 15.1-28. 15.1.2.1 Power Distribution V The transient response of the reactor is dependent on the initial power distribution. The nuclear design of the reactor core minimizes adverse power distribution through the placement of fuel assemblies, control rods, and by operation instructions. The-power distribution may be characterized by the radial peaking factor F3g and the total peaking factor Fg. The peaking factor limits are given in the Technical Specifications. For transients that are DNB limited, the radial peaking factor is of importance. The radial peaking factor increases with decreasing power level due to rod insertion. This increase in F 3g is included in the core limits illustrated on Figure 15.1-1A. All transients that may be DNB limited are assumed to begin with an FAH consistent with_(or greater than) the initial power level defined in the Technical Specifications. For transients that may be overpower limited, the total peaking factor Fg is of importance. The value of Fg may increase with decreasing power level so that the full power hot spot heat flux is' not exceeded, i.e., Fg x Power A-15.1-5

1 l equals the design hot spot heat fl ux. All the non-LOCA transients that may be overpower limited and are analyzed as part of this report assume an initial hot full power Fg of 2.5. However, the licensing-basis value of Fg will be 2.45 for VANTAGE 5 fuel, while retaining the current Fg of 2.32 for LOPAR fuel. 15.1.3 TRIP POINTS AND TIME DELAYS TO TRIP ASSUMED IN ACCIDENT ANALYSES Limiting trip setpoints assumed in accident analyses and the time delay assumed for each trip function are given in Table 15.' . Y A reactor' trip signal acts to open two trip breakers connected in series, which feeds power to the control rod drive mechanisms. The loss of power to the mechanism coils causes the mechanisms to release the RCCAs, which then l fall by gravity into the core. There are various instrumentation delays associated with each trip function, including delays in signal actuation, in opening the trip breakers, and in the release of the rods by the mechanisms. The total delay to trip is defined as the time delay from when the monitored O G parameter reaches the trip setpoint till the rods are free and begin to fall. Table 15.1-3 refers to the overtemperature AT (OTAT) and the overpower AT (0 PAT) reactor trip setpoints shown on Figure 15.1-1A. ' These trip setpoints bound the transition cores and a full core of VANTAGE 5 fuel. The associated OTAT f(41) penalty is shown on Figure 15.1-1B. For all the reactor trips, the difference between the trip setpoints assumed in the analysis and the nominal trip setpoints account for instrumentation channel error and setpoint error. The plant Technical Specifications specify the nominal trip setpoints. The calibration of protection system channels and the periodic determination of instrument response times are in accordance with the plant-Technical Specifications. 15.1.4 INSTRUMENTATION DRIFT AND CALORIMETRIC ERRORS , The VANTAGE 5 fuel design features, the modified safety analysis assumptions, and the application of new methodologies (i.e., RTDP, WRB-1, and WRB-2) as A-15.1-6

discussed elsewhere in this report (with respect to the changes associated with the instrument uncertainties for the NSSS control parameters of power, pressure, temperature, and flow) are covered in WCAP-12769, " Westinghouse l Revised Thermal Design Procedure Instrument Uncertainty Methodology for j Alabama Power Farley Nuclear Plant Units 1 and 2 (For RTD Bypass Loops)," l and WCAP-12771, " Westinghouse Revised Thermal Design Procedure Instrument l Uncertainty Methodology for Alabams Power Farley Nuclear Plant Units 1 and 2 . (For RTD Bypass Loop Elimination)". 15.1.5 R0D CLUSTER CONTROL ASSEMBLY INSERTION CHARACTERISTICS The negative reactivity insertion following a reactor trip is a function of the acceleration of the rod cluster control assemblies (RCCAs) and the variation in rod worth as a function of rod position. With respect to accident analyses, the critical parameter is the time from the start of RCCA insertion up to the dashpot entry or approximately 85 percent of the rod cluster travel. For the accident analyses, the insertion time to dashpot entry is conservatively assumed to be 2.7 seconds. The RCCA position versus time is shown on Figure 15.1-2. , Figure 15.1-3 shows the fraction of total negative reactivity insertion , versus normalized rod insertion. This curve has been conservatively selected to bound future Farley reloads which can include axial blankets of natural uranium. Additionally, there is inherent conservatism in the use of this curve in that its basis is a bottom-skewed axial power distribution. For cases other than those associated with axial xenon oscillations, significant negative reactivity would be inserted earlier due to the more favorable axial power _ distribution existing prior to reactor trip. The normalized RCCA negative reactivity insertion versus time used in the-safety analysis is shown on Figure 15.1-4. The curve shown on this figure results from the combination of Figure 15.1-2 and Figure 15.1-3. The transient analyses presented in this Appendix assume a total negative reactivity insertion of 4.8% ak/k following a trip. Both the trip reactivity and reactivity insertion rate are verified to be conservative with respect to the core design as part of the reload design process (Reference 7). A-15.1-7

I For analyses requiring the use of a dimensional diffusion theory code (TWINKLE, Reference 2), the negative reactivity insertion resulting from reactor trip is calculated directly by the reactor kinetic code and is not separable from other reactivity feedback effects, in this case, the RCCA position versus time of Figure 15.1-2 is modeled in the code. 15.1.6 REACTIVITY COEFFICIENTS The transient response of the reactor coolant system (RCS) is dependent on reactivity feedback effects, in particular the moderator temperature coefficient and the Doppler power coefficient. Depending upon event-specific characteristics, conservatism may dictate the use of either large or small reactivity coefficient values (see Figure 15.1-5). Justification for the use of conservatively large versus small reactivity coefficient values is treated on an event-by-event basis. The values used are given in Table 15.1-2. 15.1.7 FISSION PRODUCT INVENTORIES O The VANTAGE 5 fuel design features and the modified safety analysis assumptions as discussed elsewhere in this report affect the fission product inventories and are addressed in Appendix C of this report. 1 15.1.8 RESIDUAL DECAY HEAT The fission product contribution to decay heat assumed in the non-LOCA ar,alyses is the ANS 1979 decay heat model (Reference 6) increased by two standard deviations for conservatism. Figure 15.1-6 presents this curve as a function of time after shutdown. 15.1.9 COMPUTER CODES UTIL12ED Summary descriptions of the principal computer codes used in the non-LOCA transient analyses are given below. Table 15.1-2 lists the codes used in the analysis of each transient. A-15.1-8

15.1.9.1 FACTRAN Computer Code FACTRAN (Reference 3) calculates the transient temperature distribution in a cross section of a metal clad V02 fuel rod (LOPAR or VANTAGE 5. see Figure 15.1-7) and the transient heat flux at the surface of the : lad, using as input the nuclear power and the time-dependent coolant parameters (pressure, flow, temperature, density). The code uses a fuel model which simultaneously contains the following features. A. A sufficiently large number of radial space increments to handle fast transients such as a rod ejection accident. B. Material properties which are functions of temperature and a sophisticated fuel-to-clad gap heat transfer calculation. C. The necessary calculations to handle post-DNB transients: film boiling heat transfer correlations; zircaloy-water reaction; and partial melting of the fuel. The effects of IFBA are implicitly included in the fuel rod model by appropriately modifying the initial fuel temperatures. FACTRAN is further discussed in Reference 3. 15.1.9.2 LOFTRAN ComDuter Codj Transient response studies of a pressurized water reactor (PWR) system to specified perturbations in process parameters use the LOFTRAN (Reference 4) program. The LOFTRAN program models all three reactor coolant loops. This code 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 valves, _and safety valves are also considered in the program. LOFTRAN also includes a point neutron kinetics model and reactivity effects of the moderator, fuel, boron, and rods. The secondary side of the steam generator uses a homogeneous, saturated mixture for the thermal transients and a water level correlation for indication and control. The code simulates the reactor protection system A-15.1-9

which includes reactor trips on high neutron flux, OTAT, OPAT, high p and low pressurizer pressure, low flow, and high pressurizer water level. d 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 is a versatile program that is suited to both evaluation and control studies as well as parameter sizing. L0fTRAN can also calculate the transient value of DNBR based on the input from the core thermal safety limits. The core limits represent the minimum value of DNBR as calculated for typical or thimble cell. LOFTRAN is further discussed in Reference 4. 15,1.9.3 ANC Computer Code ANC is an advanced nodal code capable of tw-dimensional and three-dimensional neutronics calculations. ANC is the reference model for all safety analysis calculations, power distributions, peaking factors, critical baron concentrations, control rod worths, reactivity coefficients, etc. In addition, three-dimensional ANC validates one-dimensional and two-dimensional results and provides information about radial (x-y) peaking factors as a function of axial position. It can calculate discrete pin powers from nodal information as well. ANC is further discussed in Reference 5. 15.1.9.4 TWINKLE Comouter Code The TWINKLE (Reference 2) program is a multidimensional spatial neutron kinetics code. The code uses an implicit finite-difference method to solve the two-group transient neutron diffusion equations in one, two, and three dimensions. The code uses six delayed neutron groups and contains a detailed multiregion fuel-clad-coolant heat transfer model for calculating pointwise i Doppler and moderator feedback effects. The code handles up to 2000 spatial-points and performs its own steady-state initialization. Aside from basic-A-15.1-10

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. The code provides various output edits, e.g., channelwise power, axial offset, enthalpy, I volumetric surge, pointwise power and fuel temperatures. The TWINKLE code predicts the k m tic behavior of a reactor for transients which cause a major perturbation in the spatia' neutron flux distribution. I I TWINKLE is further described in Reference 2. I l 15.1.9.5 THINC Computer Code j The THINC computer program performs thermal-hydraulic calculations. This code calculates coolant density, mass velocity, enthalpy, void fractions,

      . static pressure, and departure from nucleate boiling ratio (DNBR)                 l distributions along flow channels within a reactor core. The THINC code is further described in Section 4.0 of Safety Assessment (see Attachment 4).

O A-15.1-ll

i I h 15.1.10 REFERENCES

 ./~')

v

1. A. J. Friedland and S. Ray, " Revised Thermal Design Procedure,"

WCAP-11397-P-A, April 1989.

2. D. H. Risher, Jr. and R. F. Barry, " TWINKLE -- A Multi-Dimensional Neutron Kinetics Computer Code," WCAP-7979-P-A (Proprietary) and WCAP-8028-A (Nonoroorietary), January 1975.

3. H. G. Hargrove, "FACTPAN -- A FORTRAN-IV Code for Thermal Transients in a 002Fuel Rod," WCAP-/?O8-A, December 1989.

4. T. W. T. Burnett, et al., "LOFTRAN Code Description, "WCAP-7907-P-A (Proprietary) and WCAP-7907-A (Nonoroprietarvi, April 1984.

(}

5. Y. S. Liu, et al., "ANC-A Westinghouse Advanced Nodal Computer Code,"

WCAP-10965-P-A (ProDrietary) and WCAP-10966-A (Nonoroorietarvi, December 1985. l

6. ' ANSI /ANS-5.1-1979, " Decay Heat Power In Light Water Reactors,"

August 29, 1979.

                       -7.       Davidson, S. L. (Ed..) et al., " Westinghouse Reload Safety Evaluation             ,

Methodology," WCAP-9272-P-A, July 1985. P. b. A-15.1-12

/D Table 15.1-1 V Nuclear Steam Supply System Power Ratinas item Ratina (MWt) Design core thermal power (MWt) 2652 l Thermal power generated by the react.r 8 (a) coolant pumps (nominal) Engineered Safety Features (ESF) design rating 2775 (b) Thermal power generated by the reactor 10 (a) coolant pumps (ESF, nominal) Thermal power generated by the reactor 15 (a) coolant pumps (ESF. maximum)

~~

(a) Nominal pump heat is considered to be 8 MWt for the rated NSSS power ti 2652 MWt and 10 MWt for the uprated NSS5 power of 2785 MWt. The non-LOCA analyses assume a conservative maximum of 15 MWt for those transients in which larger values of pump heat are conservative. For transients in which pump heat would provide a transient benefit, no (zero) pump heat is assumed. (b) Although the non-LOCA events analyzed for the VANTAGE 5 fuel upgrade assume the ESF powc. rating, the licensing value remains 2652 MWt (excluding reactor coolant pump heat). I) c A-15.1-13

O O Table 15.1-2A (3heet I of 3) O SUttiARY OF INITIAL CONDITIONS AND COMPUTER CODES USED Reactivity Coefficients Asse m d Moderator Moderator Initial NSSS Computer Temperature Density Thermal Power Output 3 Faults Codes Utilized (ptm/*F) fak/am/cm 1 DoDDier Assumed (MWt) Condition II Uncontrolled RCCA bank TWINKLE +7.0 --- Coefficient is 0 (subcritical) withdrawal from'a FACTRAN consistent with a (e) subtritical condition THINC defcct of -900 pcm Uncontrolled RCCA bank LOFTRAN +7.0 0.50 Lower and upper withdrawal at power (see Figure 15.1-5) 279,167jE'f.) and 2790 RCCA misalignment THINC, ANC --- --- --- 2775(b,f) LOFIRAN , Uncontrolled boron N/A N/A N/A N/A 0 and 2785(a) 1 dilution u, Partial loss of forced LOFTRAN, +7.0 --- Upper (see 2790(C'f)

reactor coolant flow FACTRAN, THINC F igure 15.1--5)

Startup of an inactive LOFTRAN, --- 0.50 Lower (see 1726.7(a,e) reactor coolant pump FACTRAN, THINC Figure 15.1-5) Loss of external LOFTRAN +7.0 0.50 Lower and upper 2790(C'I) electrical load (see Figure 15.1-5) and/or turbine trip loss of normal feedwater LOFTRAN +7.0 --- Upper (see 2790(C) figure 15.1-5) Loss of all AC power to LOFTRAN +7.0 --- Upper (see 2790(C'*) the station auxiliaries Figure 15.1-5)

O O O Table 15.1-2A (Sheet 2 of 3) SLHMARY OF INITIAL CONDITIONS AND COMPUTER CODES USED Reactivity Coefficients Assuced Moderator Moderator Initial MSSS Computer Temperature Density Thermal Power Output Faults 3 Codes Utilized focm/*F) - (ak/mi:/cm 1 Doppler Assumad (MWt) t Condition 11 Excessive heat removal LOFTRAN --- 0.50 tower (see 0 and 2785(**f) : , due to feedwater system Figure 15.1-5) mal functions Excessive load increase LOFTRAN 0.0 (*, below) 0.50 Upper and lower 2785(a.f) (see Figure 15.1-5) l Accidental depres- LOFTRAN +7.0 --- Lower (see 2790(C'I) surization of the RCS Figure 15.1-5) Accidental depres- LOFTRAN Function of moderatcr density See Figure 15.2-408 0 (subtritiul) i 1 y surization of the (see Section 15.2.13 and (e) i j ; main steam system Figure 15.2-40A)

Inadvertent operation LOFTRAN +7.0 ---

Lower (see 2785(*'f) i w of the ECCS during figure 15.1-5) t power operation ! i 0 i Condition III l Complete loss of forced LOFTRAN +7.0 --- Upper (see 2790(C) j reactor coolant flow FACTRAN, THINC Figure 15.1-5) ! i l . Single RCCA withdrawal ANC, THINC --- --- - - -- 2775(b) l t at full power i I

                  * - More limiting than +7.0 pcm/*F r

i

i O O O Table 15.1-2A (Sheet 3 of 3) j

SUMMARY

OF INITIAL CONDITIONS AND COMPUTER CODES USED : !' Reactivity Coefficients Assumed Moderator Moderator Initial NSSS Computer Temperature Density Thermal Power Output Faults 3 Codes Utilized focm/*F) fak/om/cm 1 Doppler Assumed (MWt) , i Condition IV I Major secondary system LOFTRAN Function of moderator censity See Figure 15.2-40B 0 (subtritical) .; pipe rupture up to and THINC (see Section 15.4.2.1 and (e) including double-ended Figure 15.2-40A) ; , rupture of a steam pipe Reactor coolant pump t0FTRAN, + 7 . 0' --- Upper (see 2846(d.e) ; shaft seizure FACTRAN, Figure 15.1-5) '

(locked rotor) THINC Spectrum of RCCA ejection TWINKLE, Refer to Coefficient is 0 and 2775(b e) accidents FACTRAN, Section 15.4.6.3 consistent with a

, THINC defect of -900 pcm ;

   ?                                                                                                                                                                                  I

,' G ! l (a) Nominal pump heat of 10 MWt is assumed i 4 (b) No pump heat (core thermal power) assumed (c) Maximum pump heat of 15 MWt is assumed (d) Uprated NSSS power with maximum pump heat increased by 2% , (e) Standard Thermal Design Procedure (STDP) with a Thermal Design Flow (TDF) of 86,000 gpm/ loop assumed ' (f) Revised Thermal Design Procedure (RTDP) with a dinimum Measured Flow (PtiF) of 37,800 gpm/ loop assumed ;

                                      -     +. .                 .             .-    . _ _ _ _ _ _ _ _ _ _ _ _ _  _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ . _ _ . _ _ _ . _ _ _ _ _   _ _

h Table 15.1-20 ; Nominal Values of Pertinent Plant Parameters Used in the VANTAGE 5 Accident Analyses STDP RTDP Parameter 19h2 lill42 NSSS Thermal Output (includes 2785* 2785 10 MWt generated by RCPs, MWt) Steam Generator Tube Plugging (%) 20 20 Core inlet Temperature (*F) 541.1 541.8 Vessel Average Temperature ('F) 577.2* 577.2 Pressurizer Pressure (psia) 2250* 2250 Reactor Coolant Flow, Loop (gpm) 86,000 (a) 87,800 (b) Steam Flow at 2775 MWt, Total (lb/hr) 12,200.000 12,200,000 Steam Pressure at Steam Generator 744 744 Outlet (psia) : Maximum Steam Moisture Content (%) 0.25 0.25 Feedwater Temperature at Steam 441.0 441.0 Generator Inlet, (*F). Average Core Heat Flux (Btu /hr-ft2 ) LOPAR Fuel- 189,818.7 189,818.7 VANTAGE 5 Fuel 197,200.5 197,200.5 (a) Thermal Design Flow (TDF) assumed in the non-LOCA analysis (b) Minimum Measured Flow (MMF) assumed in the non-LOCA analysis. Based on 86000 gpm/ loop TDF and a conservative 2.1% flow uncertainty.-

 * - Does not include uncertainties. See the appropriate accident sections.

A-15.1-17

Table 15.1-3 TRIP POINTS AND TlHE DELAYS TO TRIP ASSUMED IN ACCIDENT ANALYSES Limiting Trip Time Delay Trio function Point Assumed in Analyses (seconds) Power range high neutron flux, 118 percent 0.5 high setting Power range high neutron flux, 35 percent 0.5 low setting l i High neutron flux, P-8 74 percent 0.5 l OTAT Variable (See Figure 15,1-1A & B) 8.0(a) OPAT Variable (See figure 15.1-1A) N/A(b) I High pressurizer pressure 2425 psig 2.0 Low pressurizer pressure 1825 psig 2.0 Low reactor coolant flow (from loop flow detectors) 85% loop flow 1.0 RCP undervoltage trip (c) 1.5 Low-low steam generator level 0% of narrow range level span 2.0 High-High steam generator level 100% of narrow range level span 7.0(d) trip of the feedwater pumps and closure of feedwater system 100% of narrow range level span 2.5(8) valves and turbine trip Reactor trip (following N/A 1.0 turbine trip) (a) - Total time delay from the time the temperature difference in the coolant ' loops exceeds the trip setpoint until the rods are free to fall. Valid for operation with and without RTD bypass loops. (b) - Note, no event reanalyzed for VANTAGE 5 trips on OPAT. (c) _A specific undervoltage setpoint was not assumed in the safety analysis. (d) - From time setpoint is reached to feedwater isolation. (e) - From time setpoint is reached to turbine trip. Note: The positive and negative flux rate trips and low steam generator water level coincident with steam /feedwater flow mismatch are not explicitly .C modeled in-the non-LOCA transient analyses. A-15.1-18

LOCUS OF CONDITIONS WHERE DNBR = LIMITING VALUE

       ,                  es-                            %                         %

N \ ' ee , -

.u[

s ,

                                  \1775 psia                                                                             '-
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                   - - - - - OVERPOWER AT TaiP LINE JOSEPH M. FARLEY                                      OVERPOWER AND OVERTEMPERATURE af PROTECTION Alabama Power                                                    wucutAn etAur

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JOSEPH M. FARLEY I OVERTEMPERATURE a1 SETPOINT F(al) PENALTY r~'3 Alabama Power HUCLEAR PLANT ' 's ,/ UNIT 1 AND UNIT 2 FIGURE 15.1-18 A-15.1-20

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4 l 1 JOSEPH M. FARLEY DOPPLER POWER COEFFICIENT USED IN ACCIDENT l , ANALY$l$ Alabama Power nuetzAn etANT (V~h UNIT 1 AND UNIT 2 FIGURE 15.1-5 A-15.1-24

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t l I i I l l ' l JOSEPH M. FARLEY FUIL ROD CROS$ SECTION { Alabama Power NUCLEAR PLANT I L UNIT 1 AND UNff 2 FIGURE 15.1-7 A-15.1-26

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l l l l 15.2 CONDITION 11 - INCIDENTS OF MODERATE FRE0VENCY These faults result at worst in reactor shutdown with the plant being capable of returning to operation. By definition, these faults (or events) do not propagate to cause a more serious fault, i.e., a Condition 111 or Condition IV fault. In addition, Condition 11 events are not expected to result in a fuel rod failures or reactor coolant system (RCS) overpressurization. For the purposes of this report, the following faults I have been grouped into this category. 1 A. Uncontrolled rod clus.ter control assembly (RCCA) bank withdrawal from a suberitical condition B. I Vncontrolled RCCA bank ~ withdrawal at power C. RCCA misalignment i D. Uncontrolled boron dilution E. Partial loss of forced reactor coolant flow F. Startup of an inactive reactor coolant loop G. Loss of external electrical load and/or turbine trip H. Loss of normal feedwater

l. Loss of offsite power to the station auxiliaries J. Excessive heat removal due to feedwater system malfunctions K .' Excessive load increase L. Accidental depressurization of the RCS s

A-15.2-1

M. Accidental depressurization of the main steam system N. Inadvertent operation of the emergency core cooling system (ECCS) during power operation The Farley des;gn incorporates a solid-state reactor protection system. Reference 1 describes the techniques used to evaluate the reliability of the L- relay protection logic and demonstrates that the likelihood of no trip following initiation of a Condition 11 event is extremely small (2 x 10-7 for random component failures). The solid-state reactor protection system design has been evaluated by the same methods as those used to evaluate the relay protection system design, and the same order of magnitudo of reliability hat been demonstrated. 15.2.1 @ , TOLLED RCCA BANK WITHDRAWAL FROM A SUBCRITICAL CONDITION 15.2.1 1 Mff,9 fication of Causes and Accident Descriotion An RCCA withdrawal accident is defined as an uncontrolled addition of reactivity to the reactor core caused by withdrawal of RCCA banks resulting n a power excursion. While the occurrence of a transient of this type is highly unlikely, such a transient could be caused by a malfunction of the reactor control or the control rod drive system. This could occur with the reactor either suberitical, at hot zero power, or at power. The "at power" case is discussed in Section 15.?.2. Withdrawal of an RCCA bank adds reactivity at a prescribed and controlled rate to bring the reactor from a subtritical condition to a low power level during startup. Although the initial startup procedure uses the method of bcron dilution, the normal startup is with RCCA bank withdrawal. RCCA bank movement can cause much faster changes in reactivity than can be made by changing boron concentration (see Subsection 15.2.4, Uncontrolled Boron Dilution). O A-15.2-2

The RCCA drive mechanisms are wired-into preselected bank configurations that l are not altered during core life. These circuits prevent RCCAs from being withdrawn in other than their respective banks. Power supplied to the rod banks is controlled so that no more than two banks can be withdrawn at any time and in their proper withdrawal sequence. The RCCA drive mechanisms are of the magnetic latch type; coil actuation is sequenced to provide variable speet travel. The analysis of the maximum reactivity insertion rate includes the assumption of the simultaneous withdrawal of the two sequential banks having the maximum combined worth at maximum speed. The neutron flux response to a continuous reactivity insertion is characterized by a very fast flux increase terminated by the reactivity feedback effect of the negative Doppler coefficient. This self limitation of l the power burst is of primary importance since it limits the power to a tolerable level during the delay time for protective action. Should a continuous control rod assembly withdrawal event occur, the following automatic features of the reactor protection system are available to terminate the transient. D) L A. Source Ranoe Hiah Neutron Flux Reactor Trio The source range high neutron flux reactor trip is actuated when either of two independent source range channels indicates a neutron flux level above a preselected manually adjustable setpoint. This trip function may be manually bypassed when either intermediate range flux channel indicates a flux level above a specified level. It is automatically reinstated when both intermediate range channels indicate a flux level below a specified level. B. Intermediate Ranae Hiah Neutron Flux Reactor Trio The intermediate range high neutron flux reactor trip is actuated when either of two independent intermediate range channels indicates a flux level above a preselected manually adjustable setpoint. This trip function may be mar.ually bypassed when two of the four power range channels give readings above approximately 10% of full power and is automatically reinstated when three of the four channels indicate a power below this value. A-15.2-3

C. Power Ranae Hiah Neutron flu'x Reactor Trio (Low Settina) p- The power range high neutron flux reactor trip (low setting) is actuated when two-out-of-four power range channels indicate a power level above approximately 25% of full power. This trip function may j be manually bypassed when two of the four power range channels indicate a power level above approximately 10% of full power and is , automatically reinstated when three of the four channels indicate a , l power level below this value. l D. Power Ranae Hiah Neutron Flux Reactor Trio (Hiah Settina) r The power range high neutron flux reactor trip (high setting) is actuated when two-out-of-four power range channels indicate a power level above a preset setpoint (typically, 109% power). This trip function is always active. E. Hiah Positive' Nuclear Flux Rate Reag_ tor Trio The high nuclear flux rate reactor trip is actuated when the positive rate of change of neutron flux on two-out-of-four nuclear power range channels indicates a rate above the preset..setpoint. This trip function is always active.

In addition, control rod stops on high intermediate range flux (one out of two) and high power range flux (one out of four) serve to cease rod withdrawal and prevent the need to actuate the intermediate range.. flux trip and the power range flux trip, respectively. 15.2.1.2 Analysis of Effects and Consecuences 15.2.1.2.1 Method of Analysis The analysis of the uncontrolled RCCA bank withdrawal from subtritical accident _is performed in three stages. First, a spatial neutron kinetics computer code, TWINKLE (Reference 2), is used to calculate the core average nuclear power transient, including the various core feedback effects, i.e., Doppler and moderator reactivity. FACTRAN (Reference ) uses the average nuclear power calculated by TWINKLE and performs a fuel rod transient heat A-15.2-4 )

transfer calculation to determine the average heat flux and temperature transients. Finally, the average heat flux calculated by FACTRAN is used in j THINC for transient DNBR calculations, in order to give conservative results for a startup accident, the following assumptions are made. i A, Since-the magnitude of the power peak reached during the initial part of the transient for any given rate of reactivity insertion is strongly dependent on the Doppler power reactivity coefficient, a conservatively low (absolute magnitude) value for the Doppler power defect is used (900 pcm). Note, although this value of Doppler power defect is larger than that given in Figure 15.1-!~, it is still a conservatively low value for the Farley units. B. The contribution of the moderator reactivity coefficient is negligib!e during the initial part of the transient because the heat transfer time constant between the fuel and the moderator is much l longer than the neutron flux response time constant, However, after the initial neutron flux peak, the moderator temperature reactivity coefficiert affects the succeeding rate of power increase. The analysis assumes a moderator temperature coefficient which is

                          +7 pcm/?F at the zero power nominal temperature.

C. =The analysis assumes the reactor to be at hot zero power nominal temperature of 547'F. This assumption is more conservative than that of a lower initial' system temperature (i.e., shutdown conditions). The higher initial system temperature yields a larger fuel-te-water heat transfer coefficient, a larger specific heat of the water and-fuel, and a less-negative (smaller absolute magnitude) Doppler coefficient. The less-negative Doppler coefficient reduces

                          'he Doppler feedback effect, thereby increasing the neutron flux peak. The high neutron flux peak combined with a high fuel specific heat and larger heat transfer coefficie..t yields a larger peak heat fl ux . The analysis assumes the initial effective multiplication A-15.2-5

O factor (keff) to be 1.0 since this results in the maxiuum neutron flux peak. D. l Reactor trip is assumed to be initiated by power range high neutron flux -(low setting). The most adverse combination of instrumentation error, setpoint error, delay for trip signal actuation, and delay for control rod assembly release is taken into account. The analysis assumes a 10 percent increase in the power range flux trip setpoint (low setting), raising it from the' nominal value of l 25 percent to a value of 35 percent; no credit is taken for the source arid intermediate range protection. Figure 15.2-1 shows that j the rise in nuclear flux is so rapid that the effect of error in the trip setpoint on the actual time at which the rods release is negligible. In addition, the total reactor trip reactivity is based on the-hssumption that the highest worth rod cluster control assembly is stuck in its fully withdrawn position. (See Subsection 15.1.5 for RCCA insertion characteristics.) V E. The maximum positive reactivity insertion rate assumed is greater than that for the simultaneous withdrawal of the two sequential control banks having the greatest con.bined worth at a conservative speed (45 in./ min, which corresponds to 72 steps / min). F. The DNB analysis assumes the most-limiting axial and radial power shapes possible during the fuel cycle associated with having the two highest combined worth banks in their high worth position. G. The analysis assumes the initial power level to be below the power level expected for any shutdown condition (10-9 fraction of nominal power). The combination of highest reactivity insertion rate and low-initial power produces the highest peak heat flux. H. The analysis assumes two RCPs to be in operation (Mode 3 Technical Specification allowed operation). This is conservative with respect V to the DNB transient. A-15.2-6

1. The accident analysis employs the STDP methodology. The use of STDP stipulates that the RCS flow rate will be based on a fraction of the Thermal-Design Flow for two RCPs operating and that the RCS pressure is 50 psi below nominal. Since the event is analyzed from hot zero power, the steady-state non-RTDP uncertainties on core power and RCS ,

average temperature are not considered in defining the initial conditio1s. l l 15.2.1.2.2 Results Figures 15.2-1 throtah 15.2-3 show the transient behavior for the indicated reactivity insertion rate, with the accident terminated by the reactor trip at 35 percent of nominal power. The rate is greater than that calculated for ! the two. highest worth sequential control banks, with both assumed to be in their highest incremental worth- region. ,

      -Figure 15.2-1 shows the neutron flux transient. The neutron fl,ux overshoots the full power nominal value.for a very short period of time; therefore, the energy release and fuel temperature increase are relatively small. The tharmal flux response, of interest for the DNB considerations, is shown in Figure 15.2-2. The beneficial effect of the inherent thermal lag in the fuel is evidenced oy a peak heat flux of much less than the nominal full power value. Figure-15.2-3 shows the transient response of the hot spot average fuel and clad inner temperatures. Note the hot spot average fuel temperature increases, but remains below the nominal full power value. The minimum DNBR remains above the safety analysis limit value at all times.

Table 15.2-1 presents the calculated sequence of events. After reactor trip, the plant returns to a stable condition. The plant rray subsequently be cooled down further by following normal shutdown procedures. 15.2.1.3 Conclusions In the event of an RCCA withdrawal accident from the subcritical condition, the core and the RCS are not adversely affected since the combination of thermal O power and coolant temperature result in a DNBR greater than the limit value. V Thus, no fuel or clad damage is predicted as a result of this transient. A-15.2-7

15.2.2 UNCONTROLLED RCCA BANK WITHDRAWAL AT POWER w 15.2.2.1 Identification of Causes and Accident Description An uncontrolled RCCA withdrawal at power results in an increase in core heat flux. Since the heat extraction from the steam generator lags behind the core power generation until the steam generator pressure reaches the relief or safety valve setpoint, there is a net increase in reactor coolant temperature. Unless terminated by manual or automatic action, the power mismatch and resultant coolant temperature rise could eventually result in DNB. Therefore, to avert damage to the fuel clad, the reactor protection system is designed to terminate any such transient before the DNBR falls below the safety analysis limit value. The automatic features of the reactor protection system which prevent core damage in an RCCA bank withdrawal incident at power include the following. n Power range neutron flux instrumentation actuates a reactor trip on A. neutron flux if two-out-of-four channels exceed an overpower setpoint. B. Reactor trip actuates if any two-out-of-three AT channals exceed an OTAT setpoint. This _setpoint is automatically varied with axial power distribution, coolant average temperature, and coolant aver &ge pressure to protect against DNB. C. Reactor trip actuates if any two-aut-of-three AT channels exceed an OPAT setpoint. This setpoint is automatically varied with coolant average temperature so that the allowable heat generation rate (kW/ft) is not exceeded. D. A high pressurizer pressure reactor trip, actuated from any two-out-of-three pressure channels, is set at a fixed point. This set pressure is less than the set pressure for the pressurizer safety valves. r A-15.2-8

E.- A high pressurizer water level reactor trip actuates if any two-out-of-three level channels exceed a fixed setpoint. Besides the above-listed reactor trips, there are the following RCCA withdrawal blocks. A. High neutron flux (one out of four) B. OPAT (two out of three) 1 C. OTAT (two out of three) The manner in which the combination of overpower and overtemperature AT trips provide protection over the full range of RCS conditions is described in FSAR Chapter 7. This includes a plot (also shown as figures 15.1-1A and 15.1-18) presenting allowable reactor coolant loop average temperature and AT for the design power distribution and flow as a function.of i primary coolant pressure. The boundaries of operation defined by the overpower AT trip and the overtemperature AT are represented as

     " protection lines" on this diagram. The protection lines are drawn to include all adverse instrumentation and setpoint errors so that under nominal conditions a trip would occur well within the area bounded by these lines.

l

The utility of this diagram is in the fact that the limit imposed by a given L DNBR can be represented as a line. The DNB lines represent the locus of condittons for which the DNBR equals the safety analysis limit value. All points below and to the left of a DNB-line for a given pressure have a DNBR greater than the limit. The diagram shows that DNB is prevented for all l cases if the area enclosed with the maximum protection lines is not traversed ! by the applicable DNBR line at any point. The area of permissible operation (power, pressure, and temperature) is bounded by the combination of reactor trips: high neutron flux (fixed setpoint); high pressure (fixed setpoint); low pressure (fixed setpoint); and O Q overpower and overtemperature AT (variable setpoints). A-15.2-9

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

l The purpose of this analysis is to demonstrate the manner in which the protection functions described above actuate for various combinations of reactivity insertion rates and initial conditions. Insertion rate and initial conditions determine which trip function occurs first. 15.2.2.2 Analysis of Effects and Consecuences 15.2.2.2.1 Method of Analysis The rod withdrawal at power event is analyzed with the LOFTRAN computer code (Reference 4). The program simulates the neutron kinetics, RCS, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generators, and main steam safety valves. The program computes pertinent plant variables including temperatures, pressures, and power level. LOFTRAN uses the core protection, illustrated in Figures 15.1-1A and 15.1-1B, as input to determine the minimum DNBR during the transient. For an uncontrolled RCCA bank withdrawal at power accident, the analysis assumes the following conservative assumptions. A. This accident is analyzed with the Revised Thermal Design Procedure as described in WCAP-ll397-P-A (Reference 5). Therefore, initial reactor power, pressure, and RCS tempcratures are assumed to be at their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in Reference 5. B. For reactivity coefficients, two cases are analywad.

1. Minimum Reactivity Feedback A +7 pcm/*F moderator temperature coefficient and a least-negative Doppler only power coefficient form the basis of the beginning-of-life minimum reactivity feedback assumption (we Figure 15.1-5).

2. Maximum Reactivity Feedback A conservath ely large positive moderator density coefficient of 0.5 6k/gm/cc (corresponding to a large negative moderator temperature A-15.2-10

coefficient) and a most-negative Doppler only power coefficient form the basis of the end-of-life maximum reactivity feedback assumption (see Figure 15.1-5). C. The reactor trip on high neutron flux is assumed to be actuated at a conservative value of 118% of nominal full power. The AT trips include all adverse instrumentation and setpoint errors, while the delays for the trip signal actuation are assumed at their maximum values. D. The RCCA trip insertion characteristic is based on the assumption that the highest-worth assembly is stuck in its fully withdrawn position. E. The maximum positive reactivity insertion rate is greater than that which would be obtained from the simultaneous withdrawal of the two control rod banks having the maximum combined worth at a conservative speed (45 in/ min, which corresponds to 72 steps / min). d2 The effect of RCCA movement on the axial core power distribution is accounted for by causing a decrease in overtemperature AT trip setpoint proportional to a decrease in margin to DNB (see Figure 15.1-1B). 15.2.2.2.2 Results Figures 15.2-4 and 15.2-5 show the transient response for a rapid RCCA bank withdrawal incident starting from full power with minimum feedback. Reactor trip on high neutron flux occurs shortly after the start of the accident. Because of the rapid reactor trip with respect to the thermal time constants of the plant, small changes ir. T avg and pressure result, and margin to DNB is maintained. The transient response for a slow RCCA bank withdrawal from full power with minimum feedback is shown in Figures 15.2-6 and 15.2-7. Reactor trip on overtemperature AT occurs after a longer period and the rise in O temperature and pressure is consequently larger than for rapid RCCA bank A-15.2-ll

    ,      withdrawal.      Again, the minimum ONBR is greater than the safety analysis
       /   limit value.

Figure 15.2-B shows the minimum DNBR as a function of reactivity insertion , rate from initial full power operation for both-minimum and maximum reactivity feedback. It can be seen that the two reactor trip functions (high neutron flux and overtemperature AT) provide DNB protection over the whole range of reactivity insertion rates. The minimum DNBR is never ! less than the safety analysis limit value. Figures 15.2-9 and 15.2-10 show the minimum DNBR as a function of. reactivity insertion rate for RCCA bank withdrawal incidents starting at 60% and 10% power, respectively. The results are similar to the 100% power case; however, as the initial power decreases, the range over which the overtemperature AT trip is effective is increased. In neither-case does the DNBR fall below the safety analysis limit value. The calculated sequence of events for this accident is shown on ! Table 15.2-1. With the reactor tripped, the plant eventually returns to a stable condition. Tha plant may subsequently be cooled down further by following normal plant shutdown procedures. 15.2.2.3 Conclusions The high neutron flux and 0 TAT trip functions provide adequate protection over the entire range of possible reactivity insertion rates (i.e., the minimum value of DNBR is always larger than the safety analysis limit value for all fuel types). Therefore, the conclusions presented in the FSAR remain valid. 15.2.3 RCCA MISALIGNMENT 15.2.3.1 Identification of Causes and Accident Descriotion RCCA misalignment accidents include the following. A. One or more dropped RCCAs within the same group. A-15.2-12

-;)

i V

           .B. A dropped RCCA bank C. A statically misaligned RCCA Each RCCA has a position indicator channel which displays the position of the assembly in a display grouping that is convenient to the operator. Fully inserted assemblies are also indicated by a rod at bottom signal which actuates a local alarm and a control room annunciator.       Group demand position is'also indicated.

RCCAs move in preselected banks, and the banks always move in the same preselected sequence. Each bank of RCCAs consists of two groups. The rods comprising a group operate in parallel through multiplexing thyristors. The two groups in a bank move sequentially such that the first group is always within one step of the.second _ group in the bank. A definite schedule of actuation (or deactuation) of the stationary gripper, movable gripper, and lift. coils-of the control rod drive mechanism withdraws the RCCA held by the l () mechanism. Mechanical failures are in the direction of insertion or immobility. A dropped RCCA, or RCCA bank is detected by: A. -Sudden drop in the core power level as seen by the nuclear L instrumentation system; i B. Asymmetric power distribution as seen on out-of-core neutron detectors or core exit thermocouples; L C. Rod at bottom signal; I l D. Rod. deviation alarm; L L l E. Rod position indication. O v i A-15.2-13 l

W - Misaligned RCCAs are_ detected by: A. Asymmetric power distribution as seen on out-of-core neutron detectors or core exit thermocouples; B. Rod deviation alarm; l C. Rod position indicators. i The resolution of the. rod position indicator channel is 5 percent of. span )

    -( 7.2 in.). Deviation of any RCCA from its group by twice this distance (10 percent of span or 14.4 in.) will not cause power distributions worse       ;

I than the design limits. The deviation alarm alerts the operator to rod deviation with respect to-the group position in excess of 5 percent of span, j If the rod deviation alarm is not operable, the operator is required to log the RCCA positions in a prescribed time sequence to confirm alignment. l 1 O ig -If one or more rod position indicator channels- is out of service, the operator must follow detailed operating instructions to ensure the alignment of the nonindicated RCCAs. These operating instructions require selected pairs of-core exit thermocouples to be monitored in a prescribed time l sequence and following significant motion of the nonindicated assemblies. l

     -The operating instructions also call for the use of moveable incore neutron
    -detectors to confirt core exit thermocouple indication of assembly i

misalignment. 15.2.3.2 Analysis of Effects and Consecuences 15.2.3.2.1 Method of Analysis A. One or More Dropped RCCAs from the Same Group The LOFTRAN computer code (Reference 4) calculates the transient I system response for the evaluation of the dropped RCCA event. The (/ code simulates the neutron kinetics, RCS, pressurizer, pressurizer A-15.2-14

relief and safety valves, pressurizer spray, steam generator, and n steam generator safety valves. The code computes pertinent plant (} ' variables including temperatures, pressures, and power level. Transient reactor statepoints (temperature, pressure and power) are calculated by LOFTRAN and nuclear models are used to obtain a hot channel factor consistent with the pricary system conditions and , reactor power. By incorporating the primary conditions from the transient analysis and the hot channel factor from the nuclear analysis, the DNB design bas. is shown to be met using the THINC code. The transient response analysis, nuclear peaking factor analysis, and performance of the DNB design basis confirmation are a performed in accordance with the methodology described in Reference 6. Ncte that the analysis does not take credit for the negative flux rate reactor trip. B. Dropped RCCA Bank A(j~

           .A dropped RCCA bank results'in a symmetric power change in the core. As discussed in Reference 6, assumptions made for the dropped RCCA(s) analysis provide a boundir.g analysis for the dropped RCCA bank.

C. Statically Hisaligned RCCA Steady-state power distributions are analyzed using appropriate nuclear physics computer codes (see Section 4.3 of the Safety Assessment). The peaking factors are then used as input to.the THINC code to calculate the DNBR. The analysis examines the-case of the worst rod withdrawn from bank D inserted at the insertion limit with the. reactor initially at full power. The analysis assumes this incident to. occur at beginning of life since this results in the minimum value of the moderator temperature coefficient (least negative). This assumption maximizes the power rise and minimizes the tendency of the large moderator temperature coefficient (most

 'Q         negative) to flatten the power distribution.

A-15.2-15

15.2.3.2.2 Results A. One or More Dropped RCCAs. Single or multiple dropped RCCAs within the same group result in a , negative reactivity insertion. The core is not adversely affected during this' period, since power is decreasing rapidly. Either ; reactivity feedback or control bank withdrawal will reestablish power. i l Following a dropped rod event in manual rod control, the plant will l establish a new equilibrium condition. Without control system interaction, a new equilibrium is achieved at a reduced power level and reduced primary temperature. Thus, the automatic rod control mode of operation. is the limiting case. For a dropped RCCA event in the automatic rod control mode, the rod control system detects the drop in power and initiates control bank (% Power over3 hoot may occur due to this action by the Q withdrawal. automatic rod controller after which the control system will insert the control bank to restore nomiaal power. The Farley design uses a dual controller which-limits the power overshoct to a maximum of 2 percent. Figure 15.2-11 shews a typical transient response to a dropped RCCA (or RCCAs) in the automatic rod control mcde. In all cases, the minimum DNBR remains above the safety-analysis limit value. Following plant stabilizatice the operator may manually retrieve the RCCA by following approved operating procedures. B. Oropped RCCA Bank A dropped RCCA bank results h a negative reactivity insertion greater than 500 pcm. The core is not adversely affected dur4ng the insertion period, since power is decreasing rapidly. The transient will proceed as described in Part A; however, the return to power will be less due to b the greater worth of the entire bank. The power transient for a dropped A-15.2-16

                                                                                                          .l l

1 RCCAbinklissymmetric. Following plant stabilization, normal I

   .               _ A ,te y are: followed.

C. Stadt i .saligned RCCA The most-severe misalignment situations with respect to DNBR at signi icant power levels arise from cases in which one RCCA is fully inserted or where bank D is fully inserted with one RCCA fully withdrawn. Multiple independent alarms, including a bank insertion limit alarm, alert the operator well before the transient approaches the q postulated' conditions. The bank can be inserted to its insertion-limit !

                  - with any one assembly fully withdrawn without the DNBR falling below the                !

safety analysis limit value.

                   -The insertion limits in the Technical Specifications may vary from time to time depending on several limiting criteria. The full-power                          )

insertion limits on control bank D must be chosen to be above that ]

                   . position.which meets the riinimum DNBR and peaking factors. The full-power insertion limit'is usually dictated by other criteria. Detailed results will vary from '.ycle to cycle depending on fuel arrangements.

For'this RCCA misalignment, with bank D inserted to:its full-power

                   -insertion' limit _ , d one. RCCA fully withdrawn, DNBR does not: fall below the safety ana'.y' sis limit value. The analysis of this case assumes that
                   ?the; initial _ reactor power, pressure,.and RCS temperature are at the
                   = nominal values, with the increased radial peaking factor associated with zthe misaligned RCCA.

For RCCA misalignment with one RCCA fully inserted, the DNBR does not l fall below the safety analysis limit value. The analysis of this case U assumes that initial reactor power, pressure, and RCS temperatures are at the nominal values, with the increased radial peaking factor-associated-with~the misaligned RCCA. p y.. DNB does not occur for the RCCA misalignment incident; thus, there is no k/ reduction in the ability of the primary coolant to remove heat from the A-15.2-17 1

G V fuel rod. The peak fuel temperature corresponds to a linear heat generation rate based on the radial peaking factor penalty associated with the misaligned RCCA and the design axial power distribution. The resulting linear heat generation rate is well below that which would cause fuel melting. After identifying an RCCA group misalignment condition, the operator must take action as required by the plant Technical Specifications and operating instructions. 15.2.3.3 Conclusions For cases of dropped RCCAs or droppei banks, the DNBR remains greater than the safety analysis limit value; +ta ; fore, the DNB design criterion is met.

   ~ For all cases of any RCCA fully ince:;ed. or bank D inserted to its rod insertion limits with any single RCCA in that bank fully withdrawn (static misalignment), the DNBR remains greater than the safety analysis lirr.it value.

\._ 15.2.4 UNCONTROLLED BORON DILUTION 15.2.4.1 Identification of Causes and Accident Descriotion Reactivity can be added to-the core by-feeding primary grade water into the RCS via the reactor makeup portion of the chemical and volume control system (CVCS). Boron dilution is a manual operation under strict administrative controls with procedures calling for a limit on the rate and duration of dilution. A boric acid blend. system is provided to permit the operator to match the baron concentration of reactor coolar.t makeup water during normal charging to that in the RCS. The CVCS is designed to limit, even under various postulated failure modes, the potential rate of dilution to a value which, after. indication through alarms and instrumentation, provides the operator. sufficient time to correct the situation in a safe and orderly manner. A-15.2-18

                  ^+wrr-                      vvr   r--

The opening of the primary water makeup control valves provides makeup to the ( RCS which can dilute the reactor coolant. Inadvertent dilution from this source can be readily terminated by closing the control valve. In order for makeup water to be added to the RCS at pressure, at least one charging pump must be running in addition to ,'rimary makeup water pump. The rate of addition of unborated makcep water to tne RCS when it is not at pressure is limited by the capacity of the primary water makeup pumps. The maximurn addition rate in this case is 300 gal / min with both puicps running. The 300 gal / min reactor makeup water delivery rate is based on a pressure drop calculation comparing the pump curves with the system resistance curve. This is a maximum delivery based on the unit piping layout. Normally, only one primary water supply pump is operating. The boric acid from the boric acid tank is blended with primary grade water in the blender and the composition is determined by the preset flow rates of boric acid and primary grade water on the control board. In order to dilute, two separate operations are required. The operator rust switch from the r- automatic makeup mode to the dilute mode, and the start button mast be (' depressed. Omitting either step would prevent dilution. Information on the status of the reactor coolant makeup is continuously available to the operator. Lights are provided on the control board to indicate the operating condition of the pumps in the CVCS. Alarms are actuated to warn the operator if boric acid or makeup water flow rates deviate from preset values as a result of system malfunction. The signals initiating these alarms will also cause the closure of control valves terminating the addition to the RCS. 15.2.4.2 Analysis of Effects and Consecuences 15.2.4.2.1 Method of Analysis and Results Plant operation during refueling, cold and hot shutdown, startup, and power operation is considered in this analysis. Table 15.2-1 contains the time sequence of events for this accident. Table 15.2-2 presents results of the A-15.2-19

boron dilution analysis for power, startup, and refueling operations. Also included in this table are pertinent analysis assumptions. Perfect mixing is assumed in the analysis. This assumption results in a conservative rate of RCS boron dilution. 15.2.4.2.2 Dilution During Refueling During refueling. the following assumptions are made. A. One residual heat removal pump is operating to ensure continuous mixing in the reactor vessel. B. The seal injection water supply to the reactor coolant pumps is isolated. C. The valves on the suction side of the charging pumps are adjusted for addition of concentrated boric acid. D. The boron concentration in the refueling watt' is assumed to be 2200 ppm (see *, next page), corresponding to a shutdown margin of at least 5% Ak with all RCCAs in; periodic sampling ensures that this concentration is maintained. E. Two neutron sources are installed in the core, and the soarce range detectors outside the reactor vessel are active and provide an audible count rate. During initial core loading, BF3 detectors are installed inside the reactor vessel and are connected to instrumentation giving audible count rates to provide direct monitoring of the core. A minimum water volume of 3290 ft 3 in the RCS is considered. This corresponds to the volume necessary to fill the reactor vessel above the nozzles to ensure mixing via the residual heat removal (RHR) loop. A maximum dilution flow of 300 gal / min, limited by the capacity of the two primary water makeup pumps, and uniform mixing are assumed. O A-15.2-20

I o Q The operator has prompt and definite indication of any boron dilution from the audible count rate instrumentation. The high count rate alarm is l actuated in the reactor containment. In acdition, a high scurce range flux level alarm is actuated in the control room. The count rate increase is l proportional to the subcritical multiplication factor. ! I For dilution during refueling, the boron concentration must be reduced from greater than 2200 ppm

to approximately 1750 ppm before the reactor will go !

critical. This would take at least 18 minutes, Within this time, the I operator must recognize the high count rate signal and isolate the primary water makeup source by closing any one of several valves and stopping the reactor makeup water pumps. i l 15.2.4.2.3' Dilution During Shutdown A plant-specific evaluation of the baron dilution event during plant shutdown (hot and cold) was performed. This evaluation is based upon the operating () procedure outlined in Reference 7. The operating procedure is based ugnn a generic boron dilution analysis assuming active RCS and RHR volumes which are conservative with respect to the Farley units. Additionally, the operating procedure accommodates mid-loop cold shutdown operation. The operating procedure is applicable for maximum dilution flow rates up to 300 gpm and minimum RHR flow rates of 1000 gpm. Current plant procedures require one reactor makeup water pump to be secured when no reactor coolant pumps are running, limiting the dilution flow rate to less than 150 gpm. In the. event of a boron dilution accident during plant shutdown, use of the operating procedure provides the plant operator with sufficient information to maintain an appropriate boron concentration to conservatively assure at least 15 minutes will be available for operator action prior to the reactor reaching a critical condition.

      *- The minimum RWST boron concentration is 2300 ppm. The 2200 ppm value 7,        bounds the case of an initial boron concentration of 2300 ppm. See I  (U         Table 15.2-2.

A-15.2-21

O 15.2.4.2.4 Dilution During Startup O l In this mode, the plant is being taken from one long-term mode of operation, hot standby, to another, power. Typically, the plant is maintained in the 1 startup mode only for the purpose of startup testing at the beginning of each cycle. During this mode of operation, rod control is in manual. All normal actions required to change power level, either up or down, require operator initiation, Conditions assumed for the analysis are: A. The dilution flow is the maximum capacity of the two primary water I makeup pumps, 300 gpm;

                                                      .                               1 B. A minimum RCS water volume of 7627 fte, corresponding to the active RCS volume minus the pressurizer;                                           ;

C. An initial boron concentration of 2100 ppm (see Table 15.2-2), corresponding to a critical iiot zero power condition, rods to insertion limits, and no xenon; and D. A critical boron concentration of 1800 ppm following reactor trip. This represents the maximum boron concentration at which the core can obtain critical conditions with all control rods inserted (less the most-reactive RCCA stuck out of the core), at hot zero power conditions. The - 300 ppm change from the initial condition noted above is a conservative minimum value. The startup mode of operation is a transitory operational mode in which the operator intentionally dilutes and withdraws control rods to take the plant critical. During this mode, the plant is in manual control with the operator required to maintain a high awareness of the plant status. For a normal approach to criticality, the operator must manually initiate a limited dilution and subsequently manually withdraw the control rods, a process that takes several hours. The Technical Specifications require that the operator determine the estimated critical position of the control rods prior to approaching criticality, thus ensuring that the reactor does not go critical A-15.2-22 L

with the control rods below the insertion limits. Once critical, the power escalation must be sufficiently slow to allow the operator to manually block the source range reactor trip after receiving P-6 from the intermediate range (nominally at 105eps). Too fast of a power escalation (due to an unknown dilution) would result in reaching P-6 unexpectedly, leaving insufficient time to manually block the source range reactor trip, and the reactor would immediately shut down. However, in the event of an unplanned approach to criticality or dilution during power escalation while in the startup mode, the plant status is such that minimal impact will result. The-plant will slowly escalate in power until the power range high neutron flux low setpoint is reached and a reactor trip occurs. From the time of reactor trip, there is greater than 15 minutes (22.9 minutes calculated) available for operator action prior to return to criticality. 15.2.4.2.5 Dilution at Power In this mode, the plant may be operated in either automatic or manual rod centrol. Conditions assumed for this analysis are the following. A. With the units at power and the RCS at pressure, the dilution rate is limited by the capacity of the charging flow control valve. Although only one charging pump is normally in operation, the analysis is performed assuming the dilution flow is the maximum capacity of two charging pumps at power operation conditions. Although the dilution flow rate is less, a conservatively large dilution flow rate of 300 gpm is assumed in this analysis. This flow rate is the maximum deliverable dilution flow rate and can be assumed to include seal injection water. B. A minimum RCS water volume of 76273ft , corresponding to the active j RCS volume minus the pressurizer. I C. An initial boron concentration of 2100 ppm, corresponding to a critical, hot full power condition, with the rods at their insertion limits. l l A-15.2-23

l D. A critical boron concentration of 1800 ppm following reactor trip. This i

 'QlO                     represents-the_ maximum boron concentration at which the core can obtain critical conditions at-hot zero power conditions with all control rods
                      - inserted (less the most-reactive RCCA stuck out of the core). No credit                                                                                                                        1 is taken for_ xenon. The.300 ppm change from the initial condition noted                                                                                                                    )

above'is a conservative minimum value. l With-the reactor.in automatic rod control, the power and temperature increase from the boron' dilution results in insertion of the control rods and a Edecrease in available. shutdown margin. The rod insertion limit alarms (Low and' Low-Low settings) alert the operator at least 15 minutes (20.9 minutes calculated) prior to criticality. This is sutficient time to deter.aine the ! cause ofidilution,-isolate the reactor makeup source, and initiate boration before the_ available shutdown margin is lost. . With the reactor in manual control, a rod stop alarm is initiated 3 percent

             =below the overtemperature AT reactor trip setpoint, which would alert the operator.;'If no operator action is taken, however, the power and temperature rise will cause the reactor to reach the overtemperature AT trip setpoint                                                                                                                              4 resulting_in a reactor trip. The-boron dilution transient in this case is                                                                                                                             '

essentially-the equivalent to an uncontrolled RCCA bank withdrawal at power. The maximum reactivity insertion rate-for a boron dilution-is conservatively estimated'to be 3.5 pcm/sec, which is-within the range of insertion rates ' analyzed; Thus, _ the_ effects of dilution prior ~ to reactor trip are bounded by - - the uncontrolled-RCCA bank withdrawal at power 1 analysis (FSAR

             ~Section~15.2.2).                         Following reactor trip, there is greater than 15 minutes                                                                                                   -?
             . (20.1 minutes . calculated) prior- to criticality. This is sufficient time for the operator to determine the cause of dilutior, isolate the reactor water smakeup source, and initiate boration before the available shut down margin is
             -lost.

15.2.4.3 _ Conclusions' Because- of the procedures involved in the dilution process, an erroneous dilution is considered incredible. Nevertheless, if an unintentional-oilution of boron in the reactor coolant does occur, numerous alarms and A-15.2-24 e 4-e. u w .m a ,_ . - , .% r, , ,,,www w..... v. , , - . . ~,# ,.,-m.. -c+ .. w...,... w .._v--, , w., ,,-s,&.g , -r v yy,.

 .f i      indications are available to alert the operator to the condition. The maximum reactivity addition due to the dilution is slow enough to allow the operator sufficient time to determine the cause of the addition and take corrective action before shutdown margin is lost.

15.2.5 PARTIAL LOSS OF FORCED REACTOR COOLANT FLOW 15.2.5.1 Identification of Causes and Accident Description A partial loss of coolant flow accident can result from a mechanical or electrical failure in a reactor coolant pump, or from a fault in the power supply'to the pump. If the reactor is at power at the time of the accident, the immediate effect of loss of coolant flow is a rapid increase in the coolant temperature. This increase could result in DNB with subsequent fuel damage if the reactor is not tripped promptly. ! O Q ~ Normal power for the reactor coolant pumps is supplied through separate buses j from a transformer connected to the generator. Wher. a generator trip occurs, L the buses are automatically transferred to a transformer supplied from external power lines so that the pumps will continue to provide forced j coolant flow to the core. Following any turbine trip where there are no i electrical faults which require tripping the generator from the network, the generator remains connected to the network for approximately 30 seconds after reactor trip before any transfer is made. Since each pump is on a separate bus, a single bus fault will not result in the loss of more than one pump. The simultaneous loss of power to all the reactor coolant pumps-(Section 15.3.4) is a highly unlikely event. The necessary protection against a partial loss of coolant flow accident is provided by the low pr_imary coolant flow reactor trip that is actuated by two-out-of-three low flow signals in any reactor coolant loop. Above ! approximately 35% power (Permissive 8, see FSAR Table 7.2.2), low flow in any m loop will actuate a reactor trip. Between approximately 10% power l (Permissive 7) and the power level corresponding to Permissive 8, low flow in L A-15.2-25

jq any_ two _ loops-will actuate a reactor trip. Reactor trip on low flow is O . blocked below Permissive 7. A reactor trip signal. from the pump breaker position is provided as a backup to the low flow signal. When operating above Permissive 7, a breaker open signal from any two pumps will actuate a reactor trip. Above Permissive 8, a

       - breaker open signal is required from only one pump to actuate a reactor trip. Reactor trip on reactor coolant pump breakers open is blocked below 1        Permissive 7.

15.2.5.2 enalysis of Effects and Conseouences 15.2.5.2.1 Method of Analysis The following case has been analyzed: All loops initially operating, one loop coasting down.

 ~ i';   This- transient is analyzed by three digital computer codes. First, the LOFTRAN (Reference 4) code is used to calculate the loop and core flow transients. The LOFTRAN code-is also used to' calculate the time of reactor trip, based on the calculated flows, the primary system pressure and temperature transients,-and the nuclear power transient. The FACTRAN (Reference 3) code is then used to calculate _ the heat flux transient based on the nucleer power and flow from LOFTRAN. Finally, the THINC code is used to calculate t'te minimum DNBR during the transient based on the heat flux from FACTRAN and the flow from LOFTRAN.          The DNBR transient presented represents the minimum of the typical and thimble cells.

15.2.5.2.2~ Initial Conditions The accident is. analyzed using the Revised Thermal Design Procedure. Initial core power, reactor coolant temperature and pressure are assumed to be at their nominal values consistent with steady-state full power operation. Uncertainties in initial conditions are included in the limit departure from L nucleate. boiling ratio (DNBR) as described in WCAP-ll397 (Reference 5). A-15.2-26

          -                                                        -                  g

l l 4 (3 M 15.2.5.2.3 Reactivity Coefficients A conservatively large absolute value of the Doppler only power coefficient , is used. The total integrated Doppler reactivity from 0 to 100% power is assumed to be -0.016 Ak. The most-positive moderator temperature coefficient (+7 pcm/*F) is assumed since this results in the maximum core power and hot spot heat flux during the initial part of the transient when the minimum DNBR is reached. 15.2.5.2.4 Flow Coastdown The flow coastdown 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 purnp momentum balance, and the as-built pump characteristics and is based on high estimates of system ,

 /^N  pressure losses.

V 15.2.5.2.5 Result 3 Figures 15 ?-12 through 15.2-17 . iow the transient respona for the loss of one reactor coolant pump with three loops initially in ope ation. The

     . figures include trends of the core flow, loop flow, nuclear power, and core heat flux coastdowns. The reactor is tripped on a low loop flow signal.

Figure 15.2-17 shows that the DNBR is alway; greater than the safety analysis limit value. For the case analyzed, since DNB does not occur, the ability of the primary coolant to remove heat from the fuel rod is noi. significantly reduced. Thus, the average fuel and clad temperatures do not increase far above the respective initial values. The calculated sequence of events is shown in Table 15.2-1. The affected reactor coolant pump will continue to coast down, and the core flow will reach a new equilibrium value corresponding to the number of pumps still in A-15.2-27

                                                                                       ~      _

I n, (,l - operation (two RCPs). With the reactor tripped, a stable plant condition , will eventually be attained. Normal plant shutdown may tnen proceed. 15.2.5.3 Conclusions

         . The analysis shows that the DNBR will not decrease below the safety analysis limit value at_any time during the transient. Thus, no fuel or clad damage is predicted, and all applicable acceptance criteria are met.

15.2.6 STARTUP OF AN INACTIVE REACTOR COOLANT LOOP 15.2.6.1 Identification qf_Causes and Accident Descriotion If the plant operates with one reactor coolant pump (RCP) out of service, there is reverse flow through the inactive loop due to the pressure o difference a:ross the reactor vessel, Tho cold leg temperature of the inactive loop is identical to the cold leg temperature of the active loops (the reactor core inlet temperature). If the reactor is operated at power 4 and assuming the secondary side of the steam generator in the in>-tive loop is not isolated, there is a temperature drop across the steam generator in

              +he inactive loop and, with the reverse flow, the hot leg temperature of the inactive loop is lower than the reactor core inlet temperature.

Administrative = procedures require that the unit be brought to a load of less than 25 percent of full power prior to starting a pump in an inactive loop in order tc bring the inactive loop hot leg temperature closer to the core inlet temperattre. Starting an idle reactor coolant pump without bringing the inactive ioop hot leg temperature close to the core inlet temperature would result in the injection of cold water into the core and would cause a rapid reactivity iasertion and subsequent power increase. If the startup of an inactive RCP accident occurs, the transient is terminated automatically by a reactor trip when the power range neutron flux U exceeds the P-8 setpoint. The P-8 setpoint will remain active until the flow A-15.2-28

 . M.

(J . in the inactive loop exceeds the low flow setpoint. If the flow in the inactive. loop exceeds the low flow setpoint before the neutron flux exceeds I the P-8 setpoint, the reactor will be tripped when the power range neutron flux exceeds the high setpoint. l 15.2.6.2 Analysis of Effects and Consecuences I 15.2.6.2.1 Method of Analysis l The transient is analyzed using three digital computer codes. The LOFTRAN cor.iputer code (Reference 4) is used to calculate the loop and core flow, nuclear power, core pressure and temperature transients following the startup 1 of an idle pump. FACTRAN (Reference 3) is used to calculate the core heat flux transient based on core flow and nuclear power from LOFTRAN. The THINC code (see FSAR Section 4.4.3.1) is then used to calculata +

ONBR during the transient based on system conditions (i.e., pressure,
re, and flow) calculated by LOFTRAN and heat fluxes calculated by FACTRAn.

Assumptions made in the analysis are as follows. A. Initial conditions of maximum core pcwer, maximun reactor coolant average temperature, and minimum reactor coolant pressure resulting in the minimum initial margin to DNB. These values are consistent with the maximum steady-state power level permitted with two loops in operation, including appropriate allowances for calibration and instrument errors. The high initial power gives the greatest temperature difference between the core inlet temperature and the inactive lcop het leg temperature. B. Following the start of the idle pump, the inactive loop flow reverses and accelerates to its nominal full-flow value. C. A conservatively large (absolute value) negative moderator , coefficient is associated with the end of life (E0L) is used. A-15.2-29

O D. A conservatively low (absolute value) negative Doppler power coefficient is used. E. The initial reactor coolant loop flows are at the appropriate values for one pump out of service and two pumps running < F. The reactor trip is assumed to occur on the power range neutron flux exceeding the P-8 setpoint. The P-8 setpoint is conservatively assumed to be 74 percent of rated power, which is a sufficient allowance for nuclear instrumentatien errors. 15.2.6.2.2 Results The results following the startup of an idle pump with the above listed assumptions are shown in Figures 15.2-18A through 15.2-18C. As is shown in these curves, during the first part of the transient, the increase in core s flow with cooler water results in an increase in nuclear power and a decrease in the core average temperature. The minimum DNBR during the traasient is greater than the safety analysis limit value. The calculated sequence of events for this accident is shown in Table 15.2-1. The transient results illustrated in Figures 15.2-18A through 15.2-18C indicate that a stabilized plant condition, with thi reactor tripped, is rapidly approached. Plant cooldown may subsequently be achieved by following normal shutdown procedures, 15.2.6.3 Conclusions The transient results show that the core is not advarsely affected. There is considerable margin to the safety analysis limit DNBR. Thus, there will be no cladding damage and no release of fission products to the RCS. O A-15.2-30

15.2.7 LOSS OF EXTERNAL ELECTRICAL LDAD AND/0R TURBINE TRIP O 15.2.7.1 Identif, cation of Causes and Accident Description A major load loss on the plant can result from either a loss of external electrical load ur from a turbine trip, A loss of external electrical load may result from an abnormal variation in network frequency or other adverse network operating condition. For either case, offsite power is available for the continued operation of plant components such as the reactor coolant pumps. The case of lors of all AC power (station blackout) is presented in Section 15.2.9. For a loss of external electrical load without subsequent turbine trip, no direct reactor trip signal would be generated. The station is designed to accept a step loss of load vrom 100% to 50% load without actuating a reactor trip with all NSSS control systems in automatic (reactor control system, pressurizer pressure and level, steam generator water level control, and steam dumps). The automatic steam dump system, with 40% dump capacity to the condenser, together with the r.ractor control system, is able to accommodate the load rejection. Reactor power is reduced to a new equilibrium value consistent with the capability of the rod control system. The pressurizer power-operated relief valves may be actuated but the pressurizer safety valves and the steam generator safety valves do not lift for the 50% load rejection with steam dump. For t. turbine or generator trip, such as would result from a loss of condenser vacuum, the reactor would be tripped directly (unless it is below permissive P-9, approximately 50% power) from a signal derived from the turbine autostop oil pressure and/or turbine stop valves. The automatic steam (* ump system accommodates the excess steam generation. Reactor cociant temperatures and pressure do not significantly increase if the steam dumo system and pressurizer pressure control system are functioning properly, if the turbine condenser was not available, the excess steam generation would be dumped to the atmosphere, and main feedwater flow would be lost. For this situation, steam generator level would be maintained by the auxiliary feedwater system to ensure adequate residual and decay heat removal. A-15.2-31

In the 9 vent the steam dump valves fail to open following a large loss cf load, the steam generator safety valves may lift and tne reactor may be tripped by the high pressurizer pressure signal, the high p essurizer water ) level signal, the overtemperature AT signal, or the low-low steam j generator water level signal. The steam generator shell-side pressure and l reactor coolant temperatures will increase rapidly. The pressurizer safety valves and steam generator safety valves are sized to protect the RCS and l steam generator against overpressure for all load losses without assuming the operation of the steam dump system, pressurizer spray, pressurizer power-operated relief valves, automatic RCCA control, or the direct reactor trip on turbine trip. The steam generator safety valve capacity is sized to remove the steam flow at the engineered safeguards design rating (105% of steam flow at rated power) from the steam gen Tator without exceeding 110% of the steam system desigu pressure. The pressurizer safety valve capacity is sized based on a complete loss of heat sink with the plant initially operating at the maximum 1 calculated turbine load along with operation of the steam generator safety valves. The pressurizer safety valves are then able to maintain the RCS pressure within 110% of the RCS design pressure without a direct reactor trip on turbine trip action. The Farley Reactor Protection System and primary and secondary system designs preclude overpressurization. Reference 8 provides a more complete discussion of overpressure protection. 15.2.7.2 Analysis of' Effects and Consecuences 15.2.7,2.1 Method of Analysis In this analysis, the behavior of;the unit is evaluated for a complete loss of steam load from full power without a direct reactor trip. This assumption is made to show the adequacy of the pressure-relieving devices and to demonstrate core protection margins; it delays reactor trip until conditions

in the RCS result in a trip due to other signals. Thus, the analysis assumes A-15.2-32 I

 .+m.          -
                   ,,we       .,-..c.-                      -.,m...%-- . ,,w,--,,r..,.,,.,,%,y,-,,.,,,,,,m_.,..,                  .._,,.,,m.,_,. . . -, ,,,,,,,,,-.--r-  --,,,v , . , , - , . . - - . . _ , -,v..-,y-~,4+- - +,,-[

1 a worst-case transient, in addition, no credit is taken for steam dump. Main feedwater flow is terminated at the time of turbine trip, with no credit taken for auxiliary feedwater (except for long-term recovery) to mitigate the consequences of the transient. The total loss of load transient is analyzed with the LOFTRAN (Reference 4) computer code. The program simulates the neutron kinetics, RCS, pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generatot s, and main steam safety valves. The program computes pertinent plant variables including temperatures, pressures, and power level.

       "sjor assumptions are summarized below.

A. The accident is analyzed using the Revised Thermal Design Procedure. Initial core power, reactor coolant temperature, and pressure are assumed to be at their nominal values consistent with ste&dy-state full power Operat ion. Uncertainties in initial conditions are included in the departure from nucleate boiling ratio (DNBR) limit as described in WCAP-ll397 (Reference 5). B. The total loss of load transient is analyzed with both maximum and minimum reactivity feedback. The maximum feedback (EOL) cases assume a large (absolute value) negative moderat.or temperature coefficient and the most-negative Doppler power coefficient. The minimum feedback (BOL) cases asmme a positive moderator temperature coefficient ',+7 pcm/'F) and the least-negative Doppler coefficient. C. From the stardpoint of the maximum pressures attained, it is conservative to assume that the reactor is in manual rod control. If the reactor were in automatic rod control, the control rod banks would move prior to trip and reduce the severity of the transient. D. The loss of load event is analyzed both with and without pressurizer pressure control. The pressurizer PORVs and sprays are assumed A-15.2-33

i f. ( . operable for the cases with pressure control. The cases with pressure control minimize the increase in primary pressure which is conservative for the DNBR transient. The cases without pressure control maximize the pressure increase which is conservative for the RCS overpressurization criterion. E. Main feedwater flow to the steam generators is assumed to be lost at the time of turbine trip. No credit is taken for auxiliary feodwater flow since a stabiliztd oiznt cond tion will be reached before auxiliary feedwater inittstior, is ,iormally assumed to occur. F. Only th( overtemperature 47, high pressurizer pressure, and low-low steam generator water level reactor trips are assumed operable for the purposes of tMs analysis. No credit is taken for a reactor trip on high pressurizer level or the direct reactor trip on turbine trip, b V G. No credit is taken for the operation of the steam dump. system or steam generator power-operated relief valves. This assumption maximizes secondary pressure, The main steam safety valves are assumed to lift and be full open at 6% above the steam generator design pressure. This 6% can be considered to include allowances for safety valve setpoint uncertainty and accumulation. H. The pressurizer safety valve setpoint includes an additional 1% uncertainty. An additional 3% is applied to allow for valve accumulation. For those cases which are analyzed primarily for DNBR (pressurizer control cases), the uncertainty is applied in the negative direction and reduces the setpoint. For those cases which are analyzed primarily for peak RCS pressure, the uncertainty is appited in the positive direction and increases the setpoint pressure, in these cases, the pressurizer safety valves begin to open at 2525 psia and are full open at 2601 psia. O A-15.2-34

i 15.2.7.2.2 Results Four cases were analyzed for a total loss of load from 100 percent of 2790 MWt (NSSS power): a) minimum feedback with pressure control, b) maximura feedback with pressure control, c) maximum feedback without pressure control, and d) minimum feedback without pressure control. The calculated seq 9ence of events for the four cases is presented in Table 15.2~1. Case A: Figures 15.2-19A through 15.2-20 show the transient response for the total loss of steam load event under BOL conditions, including a +7 pcm/'F moderator temperature coefficient, with pressure control. The reactor is tripped on overtemperature AT. The neutron flux increases until the l reactor is tripped, and although the DNBR value decreases below the initial i value, it remains well above the safety analysis limit throughout the entire l transient. The pressurizer relief valves and sprays maintain primary pressure below 110% of the design value. The main steam safety valves are l also actuated and maintain secondary pressure below 110% of the design value. Case B: Figures 15.2-21A through 15.2-22 show the transient response for the total loss of steam load event under EOL conditions, assuming a conservatively large positive moderator density coefficient of 0.5 6k/gm/cc Morresponding to a large negative moderator temperature coefficient) and a most-negative Doppler only power coefficient (see Figure 15.1-5), with pressure control. The reactor-is tripped on overtemperature AT. The ONBR increases throughout the transient and never drops below the initial value. The pressurizer relief valves and sprays maintain primary pressure below 110% of the design value. The pressurizer pressure remains below the safety valve setpoint during the transient. The main steam safety valves are also actuated and maintain secondary pressure below 110% of the design-value. Case C: p Figures 15.2-23A through 15.2-24 show the transient response for the total V loss of steam load event under BOL conditions, including a +7 pcm/*F A-15.2-35

O moderator temperature coefficient, without pressure control. The reactor is tripped on high pressurizer pressure. The neutron flux remains essentially constant at full power until the reactor is tripped, and the DNBR remains above the initial value for the duration of the transient. The pressurizer safety valves are actuated and maintain primary pressure below 110% of the design value. The main steam safety valves are also actuated and maintain secondary pressure below 110% of the design value. Case D: Figures 15.2-25A through 15.2-26 show the transient response for the total loss of steam load event under E0L conditions, assuming a conservatively large positive moderator density coefficient of 0.5 6k/gm/cc (corresponding to a large negative moderator temperature coefficient) and a

most-negative Doppler only power coefficient (see Figure 15.1-5), without i

pressure control. The reactor is tripped on high pressurizer pressure. The

DNBR increases throughout the transient and never drops below the initial l value. The pressurizer safety valves are actuated and maintain primary pressure below 110% of the design value. The main steam safety valves are l

also actuated and maintain secondary pressure below 110% of the design value. 15.2.7.3 Conclusions The results of this analysis show that the plant design is such that a total l loss of external electrical load without a direct reactor trio presents no ! hazard to the integrity of the RCS or the main steam system. All of the l applicable acceptance criteria are met. The minimum DNBR for each case is greater than the s Tety analysis limit value. The peak primary and secondary l pressures remain below 110% of design at all times. l l t l !O A-15.2-36

l O 15.2.8 LOSS Of NORMAL FEEDWATER 15.2.8.1 IdentificiLtion of Causes and Accident DescriDtion A loss of normal feedwater (from pump failures, valve malfunctions, or loss of offsite AC power) results in a reduction in capability of the secondary system to remove the heat generated in the reactor core. If the reLctor was not tripped during this accident, core damage would possibly occur from a sudden loss cf heat sink. If an alternative supply of feedwater was not supplied to the plant, residual heat following reactor trip would heat the primary system water to the point where water relief from the pressurizer would occur. Significant loss of water from the RCS could conceivably lead to core damage. Since the plant is tripped well before the steam generator heat transfer capability is reduced, the primary system variables never approach a DNB condition. The following provide the necessary protection against a loss of normal feedwater. A. Reactor trip on low-low water icvel in any steam generator. B. Reactor trip on steam flow-feedwater flow mismatch coincident with low steam gu orator water level in any loop. C. Two motor-driven auxiliary feedwater pumps (350 gpm each) whie.h start automatically on any of the following:

1. Low-low level in any steam generator

2. Trip of both main feedwater pumps

3. Any safety injection signal O

A-15.2-37

D (U

4. Loss of offsite power (automatic transfer to diesel generators).

The motor-driven auxiliary feedwater pumps can also be started inanually from the control room. D. One turbine-driven auxiliary feedwater pump (700 gpm) which starts automatically on any one of the following:

1. Low-low level in any two steam generators

2. Undervoltage on any two reactor coolant pump buses The turbine-driven auxiliary feedwater pumps can also be started manually from the control room.

( The motor-driven AFW pumps are connected to vital buses which are powered by diesel generators if a loss of offsite power occurs. The turbine-driven pump utilizes steam from the secondary system. The controls are designed to start both types of pumps within 60 seconds even if a loss of all AC power occurs simultaneously with loss of normal feedwater. The AFW pumps are normally aligned to take suction from the condensate storage tank for delivery to the steam generators. A backup source of water for the pumps is provided by the safety-related portion of the service water system (see FSAR Section 6.5). The Reactor Protection System and AFW system design ensure that reactor trip and AFW flow will occur following any loss of normal feedwater. The analysis shows that _ following a loss of normal feedwater, the AFW system is capable of removing the stored and residual heat thus preventing

   .overpressurization of the RCS, overpressurization of the secondary side, or uncovery of the reactor core. Consequently, the plant is able to return to a safe condition.

O A-15.2-38

p i 15.2.8.2 Ana'vsis of Effects and Consecuences 15.2.8.2.1 Method of Analysis A detailed analysis using the LOFTRAN (Reference 4) computer code is performed in order to determine the plant transient following a loss of normal feedwater. The code describes the core neutron kinetics, RCS including natural circulation, pressurizer, pressurizer PORVs and sprays, steam generators, main steam safety valves, and the auxiliary feedwater system, and computes pertinent variables, including the pressurizer pressure, pressurizer water level, and reactor coolant average temperature. The following assumptions are made in the analysis. A. Reactor trip occurs on steam generator low-low water level at 0.0% of narrow range span. O

 .(          B.      The plant is initially operating at 102% of the NSSS design rating (2790 MWt). A conservatively large reactor coolant pump heat of 15 MWt is assumed.

C. Core residual heat generation is based on the 1979 version of ANS 5.1 (Reference 9). ANSI /ANS-5.1-1979 is a conservative representation of the decay energy release rates. Long-term operation at the initial power level preceding the trip is assumed. D. The cuxiliary feedwater system is actuated by the low-low steam l generator water level signal. Sixty seconds following this signal auxiliary feedwater flow begins. E. The worst single failure in the auxiliary feedwater system occurs (turbine-driven pump) and one motor-driven pump is assumed to be unavailable. The auxiliary feedwater system is assumed to supply a h total of 350 gpm to two steam generators from the available motor-driven pump. A-15.2-39

l I (3 The pressurizer sprays and PORVs are assumed operable. This V F. maximizes the peak transient pressurizer water volume. G. Secondary splem steam relief is achieved through the self-actuated main steam safety valves. Note that steam relief will, in fact, be through the steam generator power-operated relief valves or condenser dump valves for most cases of loss of normal feedwater. : However, for the sake of analysis, these have been assumed unavailable. H. The main steam safety valves are assumed to lift and be full open at 6% above the steam generator design prtssure. This 6% can be considered to include allowances for safety valve setpoint uncertainty and accumulation.

1. The initial reactor coolant average temperature is 6.0'F higher o than the nominal value which is comprised of the uncertainty on h '

nominal temperature. The initial pressurizer pressure uncertainty is 50 psi. J. The auxiliary feedwater line purge volume is conservatively assumed to be the maximum value for either unit of 128 ft 3 and an initial AFW enthalpy of 90.9 Btu /lbm is assumed. The assumptions detailed above are designed to minimize the heat removal capability of the secondary system and to maximize the potential for water relief from the RCS by maximizing the expansion of the primary system. For the loss of normal feedwater transient, the reactor coolant volumetric flow remains at its nominal value. The reactor coolant pumps may be manually tripped at some later time to reduce the heat addition to the RCS. 15.2.8.2.2 Results Figures 15.2-27A'and 15.2-270 show plant parameters following a loss of

             \                             normal feedwater with the assumptions listed in the previous subsection.

A-15.2-0

O Following the reactor and turbine trip from full load, the water level in the steam generators will fall due to reduction of the steam generator void fractiot, and because steam flow through the safety valves continues so as to dissipate the stored and generated heat. One minute following the initiation of the low-low level trip, the motor-driven AFW pump automatically starts; consequently, reducing the rate at which the steam generator water level is decreasing. The capacity of the motor-driven AFW pump is such that the water level in the steam generators does not recede below the lowest level at which sufficient heat transfer area is available to dissipate core residual heat without water relief through the RCS pressurizer relief or safety valves. From Figure 15.2-27A it can be seen that at no time is there water relief from the pressurizer, if the auxiliary feedwater delivered is greater than that of one motor-driven pump, or the initial reactor power is less than 102% of the NSSS rating, or the steam generator water level in one or more steam

   '  generators is above the conservatively low 0% narrow range span level assumed s

for the low-low steam generator setpoint, the results for this transient will be bounded by the analysis presented. The calculated sequence of events for this accident is listed in lable 15.2-1. 15,2.8.3 Conclusion _t Results of the analysis show that a loss of normal feedwater does not adversely affect the core, the RCS, or the steam system since the AfW capacity is such that the reactor coolant water is not relieved from the pressurizer relief or safety valves. O A-15.2-41

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

i 15.2.9 LOSS OF-ALL AC POWER _TO THE STATION AUXILIARIES ; 15.2.9.1 Identification of Causes and Accident Description f A complete loss of nonemergency AC power will result in a loss of power to i

    .the plant auxiliaries, i.e., the reactor coolant pumps, condensate pumps, etc. The loss of power may be caused by a complete loss of the offsite grid
    'accompanie'd by a turbine generator trip or by a loss of the onsite AC                                            j distribution system. The events following a loss of AC power with turbine and reactor trip-are described in the sequence listed below.-                                                    ,

A. The emergency diesel generators will start on a loss of voltage on the plant emergency buses and begin to supply plant vital loads. B. Plant vital instruments are supplied by emergency power sources. C. As the steam system pressure rises following the trip, the steam system power-operated relief valves are automatically opened to the atmosphere. Steam dump to the condenser is assumed not to be availabl e. - 'If_ the porar-operated relief valves are not available, ! the self-actuated main steam safety valves will lift to dissipate the sensible heat of the fuel and coolant plus the residual heat produced in the reactor. D. As the no-load temperature is approached, the steam system power-operated relief valves (or.the self-actuated safety valves, if , the power-operated relief valves are not available) are used to= dissipate the residual heat and to maintain the plant at the hot standby condition. The following provide the necessary protection against a loss of all AC power. A. Reactor trip on low-low water level .in any steam generator. A-15.2-42

B. Reactor trip on steam flow-feedwater flow mismatch coincident with low steam generator water level in any loop. C. Two motor-driven auxiliary feedwater pumps that are started on: i

1. Low-low level in any steam generator

2. Trip of both main feedwater pumps

3. Any safety injection signal

4. Loss of offsite power (automatic transfer to diesel generators)

5. Manual actuation.

D. One turbine-driven auxiliary feedwater pump that is started on:

1. Low-low level in any two steam generators

2. Undervoltage on any two reactor coolant pump buses 1

3. Manual actuation.

The auxiliary-feedwater (AFW) system is initiated as discussed in the loss of normal #eedwater analysis (Section'15.2.8). The turbine-driven pump utilizes steam fror,the secondary cystem and exhausts it to the atmosphere. The motor-driven AFW pumps are supplied by power from the diesel generators. The AFW pumps-are normally aligned to take suction from the condensate storage tank for delivery to the steam generators. A backup source of water for the pumps is provided by the safety-related portion of the service water system (see FSAR Section 6.5). The Reactor Protection-System and AFW system design ensure that reactor trip and AFW flow will occur following any loss of normal feedwater. > Following the loss of power to the reactor coolant pumps (RCPs), coolant flow ( is necessary for core cooling and the removal of residual and decay heat. A-15.2-43

l Heat removal is maintained by natural circulation in the RCS loops. Following the RCP coastdr vn, the natural circulation capability of the RCS will remove decay heat from the core, aided by the AFW flow in the secondary system. Demonstrating that acceptable results can be obtained for this event proves that the resultant natural circulation flow in the RCS is adequate to remove decay heat from the core. The first few seconds after the loss of AC power to the RCPs will closely resemble a simulation of the complete loss of flow event (Section 15.3.4, where it is demonstrated that the DNB design basis is satisfied). Therefore, the DNB aspects for the station blackout event are not explicitly evaluated i in this analysis. The analysis shows that following a loss of all AC power l to the station auxiliaries, RCS natural circulation and the AFW system are capable of removing the stored and residual heat; consequently, preven.ing- I overpressurization of the RCS, overpressurization of the secondary side or uncovery of the reactor cure. The plant is therefore able to return to a safe condition, i 15.2.9.2 Analysis of Ef cts and Consecuence.s 15.2.9.2.1 Method of Analysis A detailed analysis using the 1.0FTRAN (Reference 4) computer code is performed in order to determine the plant transient following a loss of all AC power. The code describes the core neutron-kinetics, RCS including natural cirt 'ation, pressurizer, pressurizer PORVs and sprays, steam i generators, main steam safety valves, and the auxiliary feedwater system, and computes pertinent variables, including the pressurizer pressure, pressurizer

water level, ed reactor coolant average temperature.

The major assumptions used in this analysis are identical to those used in [ the loss of normal feedwater analysis (Subsection 15.2.8) with the following exceptions. I r A. No credit is taken for the innediate insertion of the control rods O as a result of the loss of AC power. A-15.2-44 l l e . _ _ _ _ _ _ _ . _ _ _ _ . _ _ _ _ _ . _ _ _ _ _ . _ _ _ _ _ _ ,

         ..   . - ~ . - . .                   . . . - - . - - -
       \

B. Power is assumed to be lost to the RCPs following tne start of rod motion. This assumption results in the maximum amount of stored energy in the RCS, C. A heat transfer coefficient in the steam generators associated with RCS. natural circulation is assumed following-the RCP coastdown. D. The RCS flow coastdown 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                                                    I

balance, the as-built pump charactaristics and high estimates of system pressure losses, j E. The initial reactor coolant average temperature is 6*F lowe.

than the nominal value which is comprised of the-uncertainty on ; nominal temperature. ! Plant characteristics and initial conditions are further discussed in , Section 15.1. Consistent with the loss of normal feedwater analysis, the most-limiting single failure occurs in the AFW system. i

                           -15.2.9.2.2                   Results Figures 15.2-27E through 15.2-27H show plant parameters following a loss of offsite power with the assumptions listed above.

The first few seconds after the loss of AC power to the RCPs will closely resemble a simulation.'of the. complete loss of flow incident, i.e., core damage due to rapidly increasing core temperatures is prevented by the , reactor trip on the-low-low SG _ water level signal. After the reactor trip, e stored and-residual heat must be removed to prevent damage to either. the RCS or the core. ' e 1.0FTRAN code results show that the natural circulation flow i.i -- available is sufficient to provide adequate core decay heat removal following reactor trip and RCP coastdown. A-15.2-45

__ _ _. ~ ___ _ _ _ _ _ . _ _ _ _ _ __. _ . . . _ _ _ _ __ .. The capacity of the motor-driven AFW pump is such that the water level in the steam generators does not recede below the lowest level at which sufficient heat transfer area is available to establish enough natural circulation flow in order to dissipate core residual heat without water release through the _RCS relief or safety valves. From Figure 15.2-27E, it can be seen that at no time is there water relief from the pressurizar. The calculated sequence of events for this accident is listed in Table R2-1. As shown in Figures 15.2-27E through 15.2.27H, the plant appr. ..as a stabilized condition followino reactor trip, pump coastdown, and auxiliary feedwater initiation. 15.2.9.3 Conclusions Results of the analysis show that, for the loss of offsite power to the - station auxiliaries event, all safety criteria are met. The DNBR transient is bounded by the complete loss of flow event (Section 15.3.4) and remains above the safety analysis limit value. AFW capacity is sufficient to prevent water relief through the pressurizer relief and safety valves; this assures that the RCS is not overpressurized. Analysis of the nhtural circulation capability of the RCS demonstrates that sufficient long-term heat removal capability exists following reactor coolant pump coastdown to prevent fuel or clad damage. b 15.2.10 EXCESSIVE HEAT REMOVAL DUE TO FEEDWATER SYSTEM MALFUNCTIONS

            -15.2.10.1                         Identification of Causes and Accident Descriotion Reductions in feedwater temperature or excessive feedwater additions are means of increasing co e power above full power. Such transients are attenuated by the thermal capacity of the. secondary plant and of the RCS.

The overpower /overtemperature protections (neutron high flux, overtemperature AT, and overpower AT trips) prevent any power increase that could lead to a DNBR that is less than the safoty analysis i limit value. A-15.2-46

An example of excessive feedwater flow would be a full e-control valve due to a feedwater control system mi error. At power, this excess flow causes a n-due to increased suScooling in the ste r 6'gg60* 5 no-load conditions, the addition n' g@ g p t so RCS temperature and thus a re d (0 sd #sd negative moderator temn' gc\ <c p# 0 excessive feedwate p t* g \ ge g level trip. OM d 4 E4 04 d @* M # o#o%p 5sf g R s9 g.@ S' c d9 ss ' @ ,6'g

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                                                                *so 4pe' ,e*                  pc        *,pd@+-@S 15.2.10.2..                                   g9s                 d      9e       cp co # oe          *19e+*\ g +\              s* g,*

The excessive sd g@' @ @ g 9" #,s.Met g6 g teSee6 to analyzed with ti to o es gc multiloop system, @ <e g 6# g gae ggt T g6 5 safety valves, pret *D

                                                 ,<0         C"     og 0 p@                                         q#* g valves. The code co,                         ,,te* gpc"        gg                                      39 35    g         c pressures, and power i                        o@* ge                    40                    p td      ge64       g<
                                                    **                                   s                             ce\)g The system is analyzed to       ,n n                     p 1.s

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g* <e64g@NS a feedwater system malfunct V. .- @ <

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                                                                                         * @ # g. - 05 low-pressure heater bypass v,                               N6 gs e'>                              se5

p g g (gotg g g *@ gog6 g e, \g trip of the heater drain pump ,95 , 9 ' @* q feedwater addition due to a coni get 4

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i d allows a feedwater control valve ge o9* 94 t*6 64* gog# g No qeti gp 62

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l I An example of excessive feedwater flow would be a full opening of a feedwater control valve due to a feedwater control syste.n malfunction or an operator error. At power, this excess flow causes a greater load demand on the RCS due to increased subcooling in the steam generator. With the piant at no-load conditions, the addition of cold feedwater may cause a decrease in RCS temperature and thus a reactivity insertion due to the effects of the negative moderator temperature coefficient of r; activity. Continuous excessive feedwater addition is prevented by the steam generator high-high l 1evel trip. q A second example of excess heat removal is the transient associated with the accidental opening of the low pressure heater bypass valve which diverts flow around the low pressure feedwater heaters. The function of this valve is to maintain net positive suction head on the main feedwater pump in the event 1 that the heater drain pump flow is lost; e.g., following a large load decrease. At power, this increased subcooling will create a greater load demand on the RC5. 15.2.10.2 Analysis of Effects and Consecuences 15.2.10.2.1 Method of Analysis The excessive heat removal due to a feedwater system malfunction transient is analyzed with-the LOFTRAN (Referenec 4) computer code. This code simulates a multiloop system, neutron kinetics, the pressurizer, pressurizer relief and safety valves, pressurizer spray, steam generator, and main steam safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. The system is analyzed to demonstrate acceptable consequences in the event of a feedwater system malfunction. Feedwater temperature reduction due to low-pressure heater bypass valve actuation in conjunction with an inadvertent trip of the heater drain pump is' considered. Additionally, excessive feedwater addition due to a control system malfunction or operator error that s allows a feedwater control valve to open fully is considered. A-15.2-47

i

 <--                                                                                                                   l (T,)   Two excessive feedwater flow cases are analyzed as follows.                                                    l A. Accidental opening of one feedwater control valve with the reactor just critical at zero-load conditions assuming a conservatively large moderator density coefficient characteristic of EOL conditions.

B. Accidental opening of one feedwater control valve with the reactor in automatic rod control at full power. l The reactivity insertion rate following a feedwater system malfunction is calculated with the following assumptions. A. This accident is analyzed with the Revised Thermal Design Procedure as described in WCAP-ll397-P-A (Reference 5). Therefore, initial reactor power, pressure, and RCS temperatures are assumed to be at their nominal values. Uncertainties in initial conditions are (O

  ,_j included in the limit DNBR described in Reference 5.

B. For the feedwater control valve accident at full power, one feedwater control valve is assumed to malfunction resulting in a step increase to 184% of nominal feedwater flow to one steam generator. C. For the feedwater control valve accident at zero-load condition, a feedwater valve malfunction occurs that results in an increase in flow to one steam generator from zero to the nominal full-load value for one steam generator. D. For the zero-load condition, feedwater temperature is at a conservatively low value of 32*F. E. The initial water level in all the steam generitors is at a conservatively low level. O . A-15.2-48

i F. .No-credit is taken for the heat capacity of the RCS and steam ! generator thick metal in attenuating the resulting plant cooldown. 1 G. No credit is taken for the heat capacity of the steam and water in the unaffected steam generators. H. The feedwater flow resulting from a fully open control valve is terminated by the steam generator high-high water level signal that closes all feedwater main control and feedwater control-bypass

                             . valves, and indirectly closes all feedwater isolation valves, and trips the main feedwater pumps and turbine generator.

Normal reactor control systems and engineered safety systems (e.g., SI) are , not required.to function. The reactor. protection system may actuate to trip the reactor due to an overpower condition. No single active failure in any ; systerr or component required for mitigation will adversely affect the. consequences of this event. The steam-generator overfill protection system meets the. requirements of Generic letter 89-19. , 15.2.10.2.2 Results l Opening of a: low-pressure heater bypass valve and trip of the heater drain , pumps causes a' reduction in the feedwater temperature which increases the thermal load on the primary system. The reduction in the feedwater temperature is less than 60'F, resulting in an increase in the heat load on.the primary system of less than 10 percent of full power. The increased- , thermal load-due to the opening of the low-pressure heater bypass valve would ' resultiin a transient very similar (but of reduced magnitude) to the Excessive Load Increase incident presented in Section 15.2.11. Thus,-the results of this event are bounded by the Excessive Load-Increase event and, ,

.-              .therefore, not presented here.                                                                                   j In t'he case of an accidental full opening of one 'feedwater control valve with.

the reactor at zero power and the above-mentioned assumptions, the' maximum-O reactivity insertion rate is less than the maximum reactivity insertion rate ! A-15.2-49 . . , ,,- -. ,-. . a ....-.-. . . . _ - _ _ . , _ . - . _ . - - _ _ - . . - . - . _ , -- . :

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

presented in Section 15.2.1, Uncontrolled Control Rod Assembly Withdrawal from a Suberitical Condition, and its analysis is, therefore, covered by that of the latter, it should be noted that if the incident occurs with the unit just cr'.tical at no-load, the reactor may be tripped by the power range high neutro's flux trip (low setting) set at approximately 25%. The full-power case (EOL maximum reactivity feedback with automatic rod control) gives the largest reactivity feedback and results in the greatest power increase. A turbine trip, which results in a reactor trip, is actuated when the steam generator water level in the affected steam generator reaches the high-high level setpoint. Assuming the reactor to be in manual rod control results .n a slightly less-severe transient. The rod control system is, however, not required to function for this event. For all cases of excessive feedwater flow, continuous addition of cold feedwater is prevented by automatic closure of all feedwater control valves, closure of all feedwater bypass valves, a trip of the feedwater pumps, and a turbine trip on high-high steam generator water level, in addition, the feedwater isolation valves will automatically close upon receipt of the feedwater pump trip signal. Following turbine trip, the reactor will automatically be tripped, either directly due to the turbine trip or due to one of the reactor trip signals discussed in Section 15.2.7 (Loss of External Electrical Load). If the reactor were in automatic-control, the control rods would be inserted at the maximum. rate following the turbine trip, and the resulting transient would not be' limiting in terms of peak RCS pressure. Transient results (see Figures 15.2-28A through 15.2.28C) show the core heat flux, pressurizer pressure, core average temperature, and DNBR, as well as the increase in nuclear power and loop AT associated with the increased thermal load on the reactor. Steam generator water level rises until the feedwater addition is term 9 ned as a result of the high-high steam generator water level trip. The DNBR does not drop below the safety analysis limit value at any time, A-15.2-50 ;

l Since the power level rises during this event, the fuel temperature will also rise until the reactor trip occurs. The core heat flux lags behind the neutron flux due to the fuel rod thermal time constant and, as a result, the peak core heat flux value does not exceed 118% of nominal. Thus, the peak fuel melting temperature will remain well below the fuel melting point. The calculated sequence of events is shown in Table 15.2-1. The transient results show that DNB does not occur at any time during the feedwater flow increase transient; thus, the ability of the primary coolant to remove heat from the fuel rods is not reduced. Therefore, the fuel cladding temperature does not rise significantly above its initial value during the transient. 15.2.10.3 Conciusions The decrease in feedwater temperature transient due to an opening of the low-pressure heater bypass valve is less severe than the excessive load increase event (see Section 15.2.11). Based on the results presented in Section 15.2.11, the applicable acceptance criteria for the decrease in () feedwater temperature event have been met. For the excessive feedwater addition at power transient, the results show , that the DNB ratios encountered are above the safety analysis limit value; hence, no fuel damage is predicted. Additionally, it has been shown that the reactivity insertion rate which occurs at no-load conditions following an excessive feedwater addition is less than the maximum value considered in the analysis of the rod withdrawal from a subcritical condition event. 15.2.11 EXCESSIVE LOAD INCREASE INCIDENT 15.2.11.1 Identification of Cause and Accident Descrip_ tion An excessive load increase incident is defined as a rapid increase in the steam flow that causes a power mismatch between the reactor core power and O the steam generator load demand. The reactor control system is designed to A-15.2-51

l O accommodate a 10% step-load increase or a 5% per minute ramp-load increase in the range of 15 to 100% of full power, taking credit for all control systems in automatic. Any loading rate in excess of thesc values may cause a reactor trip actuated by the reactor protection system.

      -This accident could result from either an administrative violation such as excessive loading by the 6perator or an equipment malfunction in the steam dump control or turbine speed control. For excessive loading by the operator or by system demand, the turbine load limiter keeps the maximum turbine load                                                                    !

at 100% rated load. During power operation, steam dump to the condenser is controlled by comparing the RCS temperature (median Tavg) to a reference temperature based

on turbine power, where a high temperature difference in conjunction with a loss of load or turbine trip indicates a need for steam dump. A single - controller or control signal malfunction does not cause steam dump valves to ! I open. Interlocks are provided to block the opening of the valves unless a large turbine load decrease or a turbine trip has occurred. In addition, the reference temperature and loss of load signals are developed by independent sensors. l l Protection against an excessive load increase accident-is provided by the following reactor protection system signals. A. Overtemperature AT B. Overpower AT C. Power range high neutron flux D. Low pressurizer pressure ) i A-15.2-52 1 _ . . . __ _._____._~-_._. _._ -

i 15.2.11.2 Analysis of Effects and Consecuences 15.2.11.2.1 Method of Analysis Four cases are analyzed to demonstrate the plant behavior following a 10% step-load increase from rated load. These cases are as follows. A. Reactor in manual control with BOL (minimum moderator) reactivity feedback B. Reactor in manual control with EOL (maximum moderator) reactivity feedback C. Reactor in automatic control with BOL (minimum moderator) reactivity feedback D. Reactor in automatic control with E0L (maximum moderator) reactivity feedback This accident is analyzed using the LOFTRAN (Reference 4) computer code. The code simulates the neutron kinetics, RCS, pressurizer, pressurizer relief and safety valves, pressurizer spray, feedwater system, steam generator, and steam generator safety valves. The code computes pertinent plant variables including temperatures, pressures, and power level. At BOL, minimum moderator feedback cases, the core has the least-negative moderator temperature coefficient of reactivity and the least-negative Doppler only power coefficient curve, therefore, the least-inherent transient response capability. Since a positive moderator temperature coefficient would provide a transient benefit, a zero moderator temperature coefficient was assumed in the minimum feedback cases. For the EOL maximum moderator feedback cases, the moderator temperature coefficient of reactivity has its most-negative value and the most-negative Doppler only power coefficient curve. This results in the largest amount of reactivity feedback due to changes in coolant temperature. A-15.2-53

Normal reactor control systems and engineered safety systems are not required to function. A conservative limit on the turbine valve opening is assumed. The analysis does not take credit for the operation of the pressurizer heaters. The cases which assume automatic rod control are analyzed to ensure that the worst case is presented. The automatic function is not required. The reactor protection system is assumed to be operable; however, reactor trip is not encountered for most cases due to the error allowances assumed in the setpoints. No single active failure in any system or component required for mitigation will adversely affect the consequences of this accident. This accident is analyzed with the Revised Thermal Design Procedure as described in WCAP-ll397-P-A (Reference 5). Initial reactor power, RCS pressure and temperature are assumed to be at their nominal values. Uncertainties in initial conditions are included in the limit DNBR as described in Reference 5. 15.2.11.2.2 Results Figures 15.2-29 through 15.2-32 illustrate the transient with the reactor in the manual rod control mode. As expected, for the BOL case, there is a slight power increase and the average core temperature shows a large decrease. This results in a DNBR which increases (after a slight decrez.se) above its initial value. For the E0L manually controlled case, there is a much larger increase in reactor power due to the moderator feedback. A reduction in DNBR is experienced but DNBR remains above the safety analysis , limit value. Figures 15.2-33 through 15.2-36 illustrate the transient assuming the reactor is in the automatic rod control mode. Both the BOL and E0L cases show that core power increases, thereby reducing the rate of decrease in coolant average temperature and pressurizer pressure. For both of these cases, the minimum DNBR remains above the safety analysis limit value. The calculated time sequence of events for the excessive load increase incident is shown on Table 15.2-1. Note that a reactor trip signal was not generated for any of the four cases. A-15.2-54

I 15.2.11.3 conclusions i It has been demonstrated that for an excessive load increase, the minimum DilBR during the transient will not go tielow the safety analysis limit value. Following the load increase, the plant rapidly reaches a stabilized condition. 15.2.12 ACCIDEllTA1. DEPRESSURIZAT10!10F THE RCS 15.2.12.1 Identification of Causes and Accident Descrintion An accidental depressurization of the reactor coolant system (RCS) could occur as a result of an inadvertent opening of a pressurizer relief or safety valve. Since a pressurizer safety velve is sized to relieve approximately twice the steam flowrate of a relief valve and will allow a much more rapid depressurization upon opening, the most-severe core conditions resulting from an accidental depressurization of the RCS are associated with an inadvertent opening of a pressurizer safety valve. Initially, the event results in a rapidly decreasing RCS pressure, which could reach hot leg saturation conditions without reactor protection system intervention. If saturated conditions are reached, the rate of depressurization is slowed considerably. However, the pressure continues to decrease throughout the event. The effect of the pressure decrease is to increase power via the moderator density feedback. However, if the plant is in the automatic mode, the rod control system functions to maintain the power essentially constant throughout the initial stages of the transient. The average coolant temperature remains approximately the same, but the pressurizer level increases until reactor trip because of the decreased reactor coolant density. The reactor may be tripped by the following reactor protection system sign 's. A. Pressurizer low pressure B. Overtemperature AT I A-15.2-55

4 [ 15.2.12.2 Analvsis of-Effects and Consecuences 15.2.12.2.1 Method of Analysis The accidental depressurization transient is analyzed by using the detailed digital computer code LOFTRAN (Reference 4). This code simulates the neutron kinetics, RCS, 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. In crder to produce cunservative results in calculating the DNBR during the transient, the following assumptions are made. A. The accident is analyzed using the Revised Thermal Design Procedure. Initial core power, reactor coolant temperature, and pressure are assumed to be at their nominal values consistent with steady-state full power operation. Uncertainties in initial conditions are l included in the departure from nucleate boiling ratio (DNBR) limit as described in WCAP-ll397 (Reference 5). B. A +7 pcm/*F (most positive) moderator temperature coefficient of reactivity is assumed in order to provide a conservatively high amount of positive reactivity feedback due to changes in the moderator temperature. C. A small (absolute vdue) Doppler coefficient of reactivity is assumed, such that the resultant amount of negative feedback is conservatively low in order to maximize any power increase due to moderator feedback. D. The spatial effect of voids resulting from local or subcooled boiling is not considered in the analysis with respect to reactivity feedback or core power shape. In fact, it should be noted, the power peaking factors are kept constant at their design values, while the void (O) formation and resulting core feedback effects would result in A-15.2-56

1 considerable flattening of the power distribution, Although this would significantly increase the calculated DNBR, no credit i- e en for this effect. 15.2.12.2.2 Results The system response to an inadvertent opening of a pressurizer safety valve is shown in Figures 15.2-37 through 15.2-39. Figure 15.2-37 illustrates the nuclear power transient following the depressurization. Nuclear power increases slowly until reactor trip occurs on OT?.T. The pressure decay and core averrge temperature transients following the accident are given in Figure 15.2-38. The DNBR decreases initially, but increases rapidly following the reactor trip as shown in Figi're 15.2-39. The DNBR remains above the safety analysis limit value throughout the transient. The calculated sequence of events is shown in Table 15.2-1. 15.2.12.3 Conclusions The results of the analysis show that the pressurizer low pressure and 0 TAT reactor protection system signals provide adequate protection against the RCS depressurization event. Thus, there will be no cladding damage or release of fission products to the RCS. 15.2.13 ACCIDENTAL DEPRESSURIZATION OF THE MAIN STEAM SYSTEM 15.2.13.1 Identification of Causes and Accident Descriotion The most-severe core conditions resulting from en accidental depressurization of the main steam system, which is classified as an ANS Condition 11 event, are associated with an inadvertent opening of a single steam dump, relief, or safety valve. The analysis of the rupture of a main steam pipe, which is classified as an ANS Condition I:/ event, is discussed in Section 15.4. O A-15.2-57

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

The steam released as a consequence of this accident results in an initial V increase in steam flow that decreases during the accident as the steam pressure falls _The energy removal from the RCS causes- a reduction of coolard temperature and pressure. In the presence of a negative moderator temperature coefficient, the cooldown results in an insertion of positive reactivity and subsequent reduction of core shutdown margin. The analysis is performed to demonstrate that the following criterion is satisfied: assuming a stuck RCCA and a single failure in the engineered safety features (ESF), the limit DNBR value will be met after reactor trip for a steam release equivalent to the spurious opening, with failure to close, of the largest of any single steam dump, power-operated relief, or safety valve. The following systems provide the necessary mitigation of an accidental - depressurization of the main steam system. A. ECCS safety injection system (SIS) actuation from any of the following:

1. Two-out-of-three low pressurizer pressure signals, and

           -2.      iigh steamline differential pressure.

B. The overpower reactor trips (neutron flux and AT) and the-reactor trip which occurs in conjunction with receipt of the. safety injection (SI) signal. C. Redundant isolation of the main feedwater lines is provided because sustained -high feedwater flow would cause additional cooldown. Therefore, a safety injection signal will rapidly close all feedwater main control and bypass control valves, trip the feedwater main pumps, and indirectly close the feedwater isolation valves (2/2 steam generator feedwater pump tripped). n v A-15.2-58

   .15.2.13.2 Analysis of Effects and Conseauencrts 15.2.13.2.1    Method of Analysis 4

The following analyses of a secondary system steam release are performed. A. A full plant digital simulation using LOFTRAN (Reference 4) to I determine.RCS temperature and pressure during cooldown B. An analysis to ascertain that the reactor does not exceed the limit DNBR value The following conditions are assumed to exist at the time of a secondary 1 system depressurization incident. ' A. E0L shutdown margin at no-load, equilibrium xenon conditions, and ) with the most-reactive assembly stuck in its fully withdrawn -1 position. Operation of-RCCA banks during core burnup is restricted in such a way that addition of positive reactivity induced by a secondary system break accident wM' not lead to a more adverse condition than the case analyzed. B. A negative moderator temperature coefficient corresponding to the E0L rodded core with the most-reactive RCCA in the fully withdrawn position. The variation of the coefficient with temperature and pressure is included, The ke rr-versus-temperature curve at 1000. psia corresponding to the negative moderator temperature coefficient plus the Doppler temperature effect used is shown on Figure 15.2-40A. The effect of power generation in the core on overall reactivity is shown in Figurc 15.2-408. C. Minimum capability for the injection of conantrated boric acid solution corresponding to the most-restrictive iingle failure in the SIS. The injection curve assumed is shown in Figere 15.2-41. This corresponds to the flow delivered by one charging pueo delivering its full contents to the cold leg header. No credit Las been taken A-15.2-59 .

for the low concentration boric acid that must be swept from the safety injection lines downstream of the refueling water storage tank (RWST) isolation valves prior to the delivery of high-concentration boric acid (2300 ppm from the RWST) to the reactor coolant loops. D. The case studied is an initial total steam flow of 224.3 lb/sec at 1004 psia from one steam generator with offsite power available. This is the maximum capacity of any single steam dump or safety valve. Initial hot-standby conditions with minimum required shutdown margin at no-load T avg are assumed since this represents the most-conservative initial conditions. E. In computing the steam flow, the Moody Curve for fl/D - 0 is used. F. Perfect moisture separation in the steam generator is assumed. G. A boric acid sblution of 0 ppm is assumed in the boron injection tank (BIT). 15.2.13.2.2 Results Since it is postulated that all of the conditions described above occur simu7taneously, the resuits presented are a conservative indication of the events +. hat would occur assuming a secondary system steam release. Figures 15.2-42A through 15.2-42C show the transients resulting from a steam release having an initial steam flow of 224.3 lb/sec at 1004 psia with steam release from one safety valve. The assumed steam release is the maximum capacity of any single steam dump or safety valve. In this case, SI is initiated automatically by low pressurizer pressure. Operation of one centrifugal charging pump is considered. Boron solution at 2300 ppm enters the RCS from the RWST providing sufficient negative reactivity to prevent core famage. The calculated transient is quite conservative with respect to cooldown since no credit is taken for the energy stored in the system metal i other than that of the fuel elements or the energy stored in the other steam 1 A-15.2-60

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

generators. Since the transient occurs over a period of about 5 minutes, the j Q neglected stored energy is likely to have a significant effect in slowing the cooldown. 1 Following blowdown of the faulted steam generator, the plant can be brought l to a stabilized hot-standby condition through control of auxiliary feedwater I flow and SI flow, as described by plant operating procedures. The operating procedures would call for operator action to limit reactor pressure and pressurizer level by terminating safety injection flow, and to control steam i generator level and RCS coolant temperature using the auxiliary feedwater I system. Any action required of the operator to maintain the plant in a stabilized condition will be in a time frame in excess of 10 minutes j following Si actuation. The calculated time sequence of events for this accident is listed in l Table 15.2-1. 15.2.13.3 Conclusions . f) The analysis has shown that the critericn stated earlier in this section is satisfied. For an accidental depressurization of the main steam system, the DNB design basis is met. This case is less limiting than the rupture of a main steam pipe case presented in Section 15.4, 15.2.14 INADVERTENT OPERATION OF ECCS DURING POWER OPERATION 15.2.14.1 Identification of Causes and Accident Descriotion Spurious safety injection system operation at power could be caused by operator error, test sequence error, or a false electrical actuation signal. A spurious signal initiated after the logic circuitry in one solid-state protection system train for any of the following ESF functions could cause this incident by actuating the ESF equipment associated with the affected train. A-15.2-61

A. High containment pressure { B, low pressurizer pressure C. High steamline differentie" 1ressure D. Low steamline pressure following the actuation signal, the suction of the coolant charging pumps diverts from the volume control tank to the refueling water storage tank.

Simutaneously, the valves isolating the boron injection tank (BIT)-from the charging pumps and the valves isolating the BIT from the injection header automatically open and the normal charging line isolation valves close. The charging pumps force the borated water from the RWST through the pump discharge header, the BIT, the injection line, and inte the cold leg of each loop. The passive accumulator tank safety injection and low head system are available. However, they do not provide flow when the RCS is at. normal

/7 pressure. An SI signal normally results in a direct reactor trip and a. turbine trip. However, any single fault that actuates the ECCS will not necessarily produce a reactor trip. If an SI signal generates a reactor trip, the operator-should determine if the signal is spurious. If the SI. signal is determined to be spurious, the operator should terminate SI and maintain the plant in the hot-standby condition as determined by appropriate recovery procedures. If repair of the ESF actuation system instrumentation is necessary, future plant operation will be in accordance with the Technical Specifications. If-the reactor protection system does not produce an immediate trip as a result of the spurious SI signal, the reactor experiences a negative reactivity excursion due to the injected boron, which causes a decrease in reactor power. The power mismatch causes a drop in T avg and consequent coolant shrinkage. The pressurizer pressure and water level decrease, Load decreases due to the effect of reduced steam pressure on load after the turbine throttle valve is fully open. If automatic rod control is used, A-IS.2-62

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

/7 there effects will lessen until the rods have moved out of the core. The O transient is eventually terminated by the reactor protection system low pressurizer pressure trip or by manual trip. The time ~to trip is affected by initial operating conditions. These initial conditions include the core burnup history which affects initial boron concentration, rate of change of boron concentration, and Doppler and moderator coefficients. Recovery is-made in the same manner as described for the case where the SI signal results directly in a reactor trip. The only difference is the lower T avg and pressure associated with the power mismatch during the transient. The time at which reactor trip occurs is of no concern for this transient. At lower loads, coolant contraction will be slower which will result in a longer time to trip.

  '15.2.14.2 Analysis of Effects and Consecuences (3

O-15.2.14.2.1 Method of Analysis The spurious operation of the ECCS is analyzed using the LOFTRAN computer code (Reference 4). The code simulates the neutron kinetics, RCS, pressurizer, pressurizer relief and safety valves, pressurizer spray, feedwater system, steam generator, and steam generator safety valves. The code 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 has shown_that the results are relatively independent of time to trip. A typical transient is presented representing conditions at BOL with minimum reactivity feedback. Results at E0L with maximum reactivity feedback are similar except that moderator feedback effects result in a slower transient. The analysis assumes zero injection line purge volume for calculational A-15.2-63

simplicity; thus, the boration transient begins immediately when the appropriate valves open. The assumptions are as follows. A. Initial Operating Conditions This accident is analyzed with the Revised Thermal Design Procedure as described in WCAP-ll397-P-A (Reference 5). Initial reactor power, RCS pressure and temperature are assumed to be at the nominal full power values. Uncertainties in initial conditions are included in the limit DNBR as described in Reference 5. B. Moderator and Doppler Coefficients of Reactivity The analysis assumes a positive (+7 pcm/*F) moderator temperature coefficient. A low absolute value Doppler power coefficient is used. O C. Reactor Control The reactor is in manual rod control. D. Pressurizer Heaters Pressurizer heaters are inoperable. This assumption yields a higher rate of pressure decrease. E. Boron Injection , At the initiation of the event, two charging pumps inject borated - water into the cold leg of each loop. F. Turbine Load Turbine load is constant until the governor drives the throttle valve wide open, ar.d then drops as steam pressure drops. A-15.2-64

G. Reactor. Trip r I \ - Reactor trip is initiated by low pressurizer pressure at 1840 psia. 15.2.14.2.2 Results The transier ' nse is shown in Figures 15.2-43 through 15.2-45. Table 15.".. s .e calculated sequence of events. Nuclear power starts decreasing i.. < r due to boron injection, but steam flow does not decreate untit lat .n the transient when the turbine throttle valve is wide I open. The misnatch between load and nuclear power causes Tavg, pressurizer water lovel, and ;ressurizer pressure to drop. The reactor trips and control rods start moving into the core when the pressurizer pressure reaches the pressurizer low pressure trip setpoint. The DNBR increases throughout the l transient. 15.2.14.3 Conclusions , 1 D Results of the analysis show that spurious ECCS. operation without immediate reactor trip does not present any hazard to the integrity of the RCS. : DNBR is never less than the initial value. Thus, there will be no cladding damage and no release of fission products to the RCS. If the reactor does not trip immediately, the low pressurizer pressure reactor trip actuat es later. This trips the turbine and prevents excess cooldown, which expedites recovery from the incident. f3 O A-15.2-65

1 15.2.15 REFERENCES

1. W. C. Gangloff, "An Evaluation of Anticipated Operational Transients in Westinghouse Pressurized-Water Reactors," WCAP-7468,.May 1971.

2. D. H. _ Risher, Jr. and R. F. Barry, " TWINKLE ~ ~- A Multi-Dimensional Neutron Kinetics Computer Code," WCAP-7979-P-A (Proprietary) and i WCAP-8028 A (Nonoroorietary),- January 1975. .

3. H. G Hargrove,7"FACTRAN -- A FORTRAN-IV Code for Thermal Transients in )

a _UO2 Fuel-Rod," WCAP-7908-A, December 1989.

4. T.-W; T. Burnett, et al., "LOFTRAN Code Descriptica, "WCAP-7907-P-A

                              -(Proprietary) and WCAP-7907-A (Nonoroprietary),_ April 1984.

1

5. A. J. Friedla_nd and S, Ray, " Revised Therrrl Design Procedure,"

( 'WCAP-11397-P-A, April 1989.

                     ' 6.      R. L. Haessler, et al..,_ " Methodology for the Analysis of the Dropped Rod-Event,;" WCAP-11394-P-A-(Proorietary):and WCAP-11395-A (Nonoroorietary),

January;1990.

                       -7.   : Letter from D.--E. McKinnon to L. K. Mathews,=" Operating Procedure for Mode 4/5 Boron Dilution," 90AP*-G-0041, July 6, 1990, with attachments.                                                  -

8. M.-A. Mangan, " Overpressure Protection for Westinghouse Pressurized-Water Reactors," WCAP-7769 Rev. 1, June 1972.

                    ' 9.       ANSI /ANS-5.1-1979, " Decay Heat Power In Light Water-Reactors,"
                             -August 29, 1979.

A-15.2-66

 ,---~er...+       ., , :..-,-      .
                                         .n..,...-.s,_-_.---~   -._.-:.,v..n,..,.n.. --+r--, nn.--.. , . . .,.mn-.- rw .---wn,,,-~.,-,----...--,rw,--.w

             . .: n .            .                   _-           .                             -
 ,3 t;., i
         \

l i l TT Table 15.2-1 (Sheet 1 of 7) V Time Seouence of Events for Condition II Events Accident Event Time (sec) l Uncontrolled RCCA Initiation of uncontrolled rod 0.0 Withdrawal From A withdrawal (78.75 pcm/sec) Subcritical Condition reagtivityinsertionratefrom 10- fraction of nominal power Power range high neutron flux 9.6 low setpoint reached Peak nuclear power occurs 9.7 Rods begin to fall into core 10.1 Peak heat flux occurs 11.4 Peak average clad temperature 12.0 occurs Peak average fuel temperature 12.5 occurs Uncontrolled RCCA Bank With-drawal At Power (Full Power With Minimum Feedback)

1. Case A Initiation of uncontrolled 0.0 RCCA withdrawal at maximum insertion rate (110 pcm/sec)

Power range high neutron flux high setpoint reached 1.0 Rods begin to fall into core 1.5 Minimum DNBR occurs 2.6

2. Case B Initiation of uncontrolled RCCA 0.0 withdrawal at a small reactivity insertion rate (0.6 pcm/sec)

Overtemperature AT reactor 57.7 trip setpoint initiated Rods begin to fall inte core 59.7 j O Minimum DNBR occurs 60.2 A-15.2-67

Table 15.2-1 (Sheet 2 of 7) Time Seauence of Events for Condition 11 Events Accident Event Time (sec) Uncontrolled Boron Dilution

1. Dilution During Refueling Dilution begins O Shutdown margin lost (if dilution continues) >1100

2. Dilution During Startup Power range - low setpoint reactor trip due to dilution 0 Shutdown margin lost (if dilution continues) >900

3. Dilution During Full-Power Operation

a. Automatic Reactor Control Operator .eceives low-low rod insertion limit alarm due to dilution 0 Shutdown margin lost (if dilution continues) >900

b. Manual Reactor Control Reactor trip on OTAT due to dilution 0 Shutdown margin is lost (if dilution cor.tinues) >900 Partial Loss Of Forced Reactor Coolant Flow All Pumps initially One pump begins coasting down 0.0 In Operation, One Pump Coasting Down Low flow reactor trip 1.6 Rods begin to drop 3.1 Minimum DNBR occurs 3.9 0

A-15.2-68

Table 15.2-1 (Sheet 3 of 7) q Time Seouence of Events for Condition 11 Events -( ,r-Accident Event Time (sec) l Startup Of An inactive Reactor Coolant Loop At An incorrect Temperature Initiation of pump startup 0.0 Power reaches P-8 trip setpoint 6.7 Rods begin to drop 7.2 Minimum DNBR occurs 8.0 l Loss Of External ' l Electrical Load l With Pressurizer Pressure loss of electrical load 0.0 Control (BOL) Overtemperature AT 9.7 reactor trip setpoint reached A-Rods begin to drop 11.7 .{Wj-Minimum DNBR occurs 13.5 Peak pressurizer pressure 13.5 occurs . Initiation of steam release 14.0 from SG safety valves With Pressurizer Pressure Loss of electrical-load 0.0 Control (E0L) Overtemperature AT 9.9 reactor trip setpoint reached Peak pressurizer pressure 10.0 occurs Rods begin to drop 11.9 Initiation of steam release 14.5 from SG safety valves Minimum DNBR occurs -(a) O l/ s (a) - DNBR does not decrease below its initial value. A-15.2-69 _ _ _ - ~

i Table 15.2-1 (Sheet 4 of 7) ( ) , Time Seouence of Events for Condition 11 Events Accident Event Time (sec) Loss Of External Electrical Load Without Pressurizer Loss of electrical load 0.0 Pressure Control (BOL) High pressurizer pressure 5.9 l reactor trip setpoint reached ) Rods begin to drop 7.9 Peak pressurizer pressure 10.0 occurs l Initiation of steam release 14.0 i from SG safety valves Minimum DNBR occurs (a) Without Pressurizer loss of electrical load 0.0 (%]' _- Control (EOL) High pressurizer pressure 5.8 reactor trip setpoint reached Rods begin to drop 7.8 Peak pressurizer pressure 9.0 occurs Initiation of steam release 14.0 from SG safety valves Minimum DNBR occurs (a)

         -(a) - DNBR does not decrease below its initial value.

O A-15.2-70

Table 15.2-1 (Sheet 5 of 7) Time Secuency of Events for Cgndition 11 Events Accident Event lime (seci Loss Of N::rmal Feedwater M .i feedwater flow stops 10.0 Low-Low SG water level 53.5 reactor trip Rods begin to drop 55.5 Peak water level in 59.0 pressurizer occurs One motor-driven pump begins 115.5 to deliver auxiliary feedwater Cold auxiliary feedwater is 443.0 delivered to two steam generators Core decay heat decreases 4576.0 to auxiliary feedwater Loss Of Offsite Power To The Station Auxiliaries AC power is lost and 10.0 main feedwater flow stops Low-Low SG water level 53.7 reactor trip Rods begin to drop 55.7 Reactor coolant pumps 57.7 begin to coastdown One motor-driven pump begins 115.7 to deliver auxiliary feedwatet Cold auxiliary feedwater is 443.0 delivered to two steam generators Core decay heat decreases 2212.0 to auxiliary feedwater heat removal capacity Peak water level in 2268,0 pressurizer occurs I A-15.2-71

Table 15.2-1 (Sheet 6 of 7) p-Time Scovence of Events for Csadition II Events y

                   -Accident              .

Event Time (sec) ,

                                                                                      -l 4

Excessive feedwater flow l At Full Power One main feedwater control 0.0 valves fails fully open l High-high SG water level 50.0 1 signal generated-Minimum DNBR occurs 51.5 Turbine trip occurs due to 52.5 high-high SG water level , Reactor trip due to turbine 53.5 j trip (rod motion begins) ; r Feedwater control valves 57.0 fully closed j C)

'(/

Excessive Load increase Manual Reactor 10% step-load increase 0.0 Control (BOL) Equilibrium. conditions 140.0 reached (approx, time)

       ' Manual Reactor                    10% step-load increase 0.0 Control-(EOL)

Equilibrium conditions 75.0 ; reached Iapprox, time) l Automatic Reactor 10% step-load increase 0.0 Control- (BOL) Equilibrium conditions 250.0 reached (approx, time) Automatic reactor 10% step-load increase 0.0 control (E0L) Equilibrium conditions 75.0 ! reached (approx. time) 1 O G A-15.2-72

y Table 15.2-1 (Sheet 7 of 7) Time Seouence of Events for Condition 11 Events-Accident' Event Time (sec) Accidental Depressurization Of The RCS

                                           !nadvertent opening of one          0.0 RCS safety valve Overtemperature AT reactor          14.5.

trip setpoint reached Rods begin to drop 16.5 Minimum DNBR occurs 17.0 l Accidental Depressurization 0f The Main Steam System Inadvertent opening of one 0.0 main steam safety or relief I valve Borated water from the 252.7 RWST reaches the core j \ Pressurizer empties 253.7 Criticality reached 445.7 Inadvertent Operation of ECCS During Power Operation SI pumps begin injecting 0.0 borated water low pressurizer pressure 46.4 reactor trip setpoint reached Rods begin to drop 48.4 Minimum DNBR occurs (a) (a) - DNBR does not decrease below.its initial value. A-15.2-73

Table 15.2-2 ('}

SUMMARY

OF BORON DILUTION ANALYSIS RESULTS AND ANALYSIS ASSUMPTIONS Dilution Flow Active Volume Operator Action Mode of Ooeration Rate-(aom) (cubic feet) Time (minutes) l Power Operation Auto Rod Control 300 7626.9 20.9 Manual Rod Control- 300 7626.9 20.1 Startup 300 7626.9 22.9 Refueling' 300 3290.0 18.4 OTHER IMPORTANT ANALYSIS ASSUMPTIONS Assumed Assumed . Average 't Initial Baron Critical Boron ' Core Coolant Mode of Operation Conc. (Dom) Conc. (com) Temperature (*F) Power Operation Auto Rod Control 2100 1800 583.2 Manual Rod Control 2100 1800 583.2 Startup 2100 1800 554.5 Refueling 2200 ;750 140.0 O,' v A-15.2-74

                    .          . - . . - ,             ... ..         .- . . . . . ..,. ~ . . . ..... - -. - .             .. -     . - ~ - .              ,     ,.,n I

u J

   ;-j .          .
   -s           -

i 10 3 .

                                                     .                                                                                                         I I                                                                                                         I        ,

<- 1 1

g. i 1

1 W 10 0 + o .$

                        -g                           I U-E w

we' ft

                        .f.         10 1 f oc

~h/h -

                       .g 5

d i

                                                                                                                                                                      -t J

l 1-10'2 0 5 10 15 20 is le

                                                                                                              . TIME (SEC) t t
                                                                                                                                                                       ?

4 J STARTUP FROM SUBCRITICAL FOR 3 LOOP. 17 X 17 !; O ^ iebeme Power A ~OSEPHM.FARLEY ec'E UNIT 1 AND UNIT 2

                                                                                                                 <.~T                                                  1 FIGURE-15.2-1 A-15.2-75 s ,.---                        ,,.,----~..a--,           .w,-,-                              ..,_n.,,,,,.,c       -

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

i f 4 9 i 9+

                                                                                                                                                .i       1
                     . _T . e +                                                      REACTIVITY INSERTION                                                '

5 RATE - 78'.75 x 10-5 A/,/Sec ; g.- - u , i C ,t+ t s C u-

                                .5-
                    ~g.
                    'D            4-                                                                                                            i
                    ~ x.                                                                                                                               ,

a

                                .e l

A),; a 7 w- .2-i I

                                .I                                                                                                                     ,

I

                              -0.                                                                                                              '
;                                    0~                       5'        10            15           20              25             10 TIME.(SEC) i t

i STARTUP FROM SUBCRITICAL FOR 3. LOOP, 17 X 17-JOSEPH M. FARL2Y f Alabama Power NUCLEAR PLANT PLANT FACTRAN HEAT-FLUX I

.~

UNR I AND UNR 2 FIGURE.15.2-2 l A-15.2-76 I w -

2400. O 2200' REACTIVITY INSERTION g , , , , , . RATE - 78.75 x lY S AK/Sec 2 1800. 5 W 1600. N N 1400. n a 1200. 1000, e 800. s 600. ; 400. O. f. . 10. 15, 20. 25. 30. 800 REACTIVITY INSERTION { #5 ' RATE - 78.75 x 10-5 AK/Sec [ bw 700-t D 650-N O d 600-em E 550-

c 500 0 5 to 15 20 25 30 TIME (SEC)

JOSEPH M. FARLEY UNCONTROLLED ROO WITHDRAWAL FROM A SUSCRITICAL CONDITION TEMPERATURE VERSUS TIME Alabama Power O NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 15.2-3 A-15.2-77

                                  .1.6 1
   -\

d 5 I~ REACTIVITY-INe'cRTION RATE 1:0 X 10-5 aK/Sec 1,2 J-1-

                        ,                                                                                                                                                     t-g -~                                                                                                                                               ;

e, .s. 6 [

          -                                                                                                                                                                  4
                          =            4 3
                        .g            .2-00  -

1--

                                                                        -2          3               4          5           6         7          8    9     10 TIME          (SEC) 3 E00_   .

7 _ s00. > 1

                          $2400.                                                                                                                                              l u

g 2300, ' 0 - N 2200. g

                           - 2100.

E L 2000. L 1900.

9. !

800 0. 1. ~ -

2. 3. 4 5. - 6. - 7. 8, 10.

TIME (SEC) l 6 TRANSIENT RE3PONSE F0P UNCONTROLLED R00 . i. JOSE?H M. FARLEY WITHDRAWAL F*n0M FULL POWER TERMINATED BY HIGH ! th l Alabama L Powel NUCLEAR PLANT UNIT 1 AND UNIT 2 ' A-15.2-78

            *       .m.                       .,.---.-.4        ..m-,      - 3a.-       , . - . . - -   ,.es.   +       .4    ,v-      r,y +
                                                                                                                                                            - c y vv --, r

l

       ~:                    620 C     610-REACTIVITY INSERTION-

,~

                       $                                                        RATE - 110 X 10-5 AK/Sec S                                                                                                                   '

w 600- !

                            . IY O '

m i m 5sr , l c , Y. ;$70- < \

                       <                                                                                                                  >l e                                                                                                                     .l cc 8-     560<

550 l 0 '.1 2. 3 4 5 6 7 8 9 10 1 TIME (SEC) l

4 -

g]- 1 i

3. -

E 2.5 E . 2.

1. 5 ,

t

                                   'O       1   2             3    4        5       6        7                8    -9        10

TIME (SEC) m TRANSIENT RESPONSE FOR UNCONTROL;ED R00 -

JOSEPH M. FARLEY WITHDRAWAt. FROM FULL POWER TERillNATED B fs Alabama Power NUCLEAR PLANT i UNIT 1 AND UNIT 2 FIGURE 15.2-5 A-15.2-79

1.6 REACTIVITY INSERTION R TE = 0.6 X 10-5 4K/Sec 51.2 a: 1 v 5 .a 5

      $         .6-E w         .

M

      *'         .z

0. yo 0

20 30 40 30 60 TIME (SEC) 2600.

2500.

       <c E

2400. w

       $ 2300.              _

g E 2200. 5

        % 2100.             '

E m E 2000. ! 5 ! 1900. 1800.O. jo. 20. s - 60, 70. T}oMk (SEC) i TRANSIENT RESPONSE FOR UNCONTROLLED R00 JOSEPH M. FARLEY WITHORAWAL FROM Fut.L POWER TERMINATED BY O Atebeme Power ~ect ' ~1 UNIT 1 AND UNTT 2 FIGURE 15.2 6 A-15.2-80

620 C REACTIVITY INSERTION 610-g RATE - 0.6 X 10-5 .K/Sec e S , g 600 M

        &    590-ig                                                                                                            i b

W 530 z 570 u 8- 560 550 10 20 3o 50 60 7o TIME (SEC) 4 l 3.5-i 3.

           =    2.5-2.

1.5 i

1. , 30 60 7o 10 20 to TIME (SEC)

                                                                                                            <<E, ,,,

so e N .,AeLev =t<=R=~gL ,e eece.T WER TERMINATED By O 41eseme ge.e, NUCLEAR PLAPg UNIT 1 AND UNIT 2 eVERTEMPEM7URE 4T FIGURE 15.2-7 A-15.2-81

y - gw

             .V.                                                        -;,4h-,
                                                                        \ - !s
                                                                          .                                                .C m

cr' o B w m 2.06 0

              ~4                   2.04 -

m

                *1 2.02 -

2-l'

                                   '1.98 -

c z <- 1.96 - Minimum Feedback -r zco aog ., a i-3

            ;h]                 g o

1.94 - J zaF oi High Flux g g .; g 1.92 - OTAT -

m. E ,

  - lS . ag@m u                    $    1.9 -                                                    Maximum Feedback z

8E; 4"R om 1.88 - [ 1.86 - OTAT g g -e 2"% ! 5"= 1*84 _

                ,89 2 0

. 3 o 952 1.82 - l 5* ~; e o-A E- 1.8 - ' E5 ' ' s e i i i i "g i

         ?                                          2o                      4o                60
         ~

x me 3o 300 do Pg 5 REACTIVITY. INSERTION RATE, pcm/sec 2 l 3 5 I

                               )ll I

O l 0 i 0 1 k k c c a a b b d d e a e e e F F m m u u m m i i 0 i x 8 n a i M M c e s

m c p x u , l E F 0 T 1{ i k hg 6 AR i N H O I T i R E S N I 0 Y O i 4 TI V I T C A i E R

i 0 2 T A T O i 0 7 6 5 4 3 2 3 2 9 8 1 2 2 2 1 2 2 1 1

                                      " 5 rE 5r
                                                                    " a9 R =53 ~                  59 E E-    a5h hI g 5"E
       > e rc em$

Tg SeE98"E9* 5 5gCE" MUS O : e DOk,

                     '                 zEEm ,95 cs ;zaEa" n5Ern
                                                                                 -                      wh'"
                                            > ' . w 6 &"

l ; i' , illlll f

M' - p C7" 5 w 3. 4 29 - e i ! ' 18 - Maximum Feedback 2.7 - [ czc 16 - - zco ec ao a m L

     $hh zay 5

x 15 - P E- = E OTAT 7 $ E 14 - T 3[@R u E 2.3 - Minimus Feedback . E E'; 4%A T W,a 12 - i no E"?

          -*                                                            High Flux gg9a 11 -                                                                                               !

2; "E~ I E

              $3                        -2          i    i                   i         e     i      i      i       i     i   i
   -          $E m                            o         20                           40            60            80        100 P          [3                                                 REACTIVITY INSERTI0i1 RATE, pcm/sec m          m w

UE

              -M   *
            .Eg 2x 2

sa

G l I j [* w j j I j ~

                             -                               g
                                                                 ,      \#                                                         -

t 1.0 -- g680 - j-09 5 NO

                            -                               E                                                                                   ,
              $l0.8 y

4 640 - i E0.7 - 0.6 - hSM E , DJ {500 1.2 2400

                  , 1.1    -                                E g                                           g2300

,r . yl1.0 - _ w' % l t. \- tl0.9 gg l2200 , I 2100 - I 0 .8 a 5 l vN c 0,7 .. 2000 - 1 E

0 .6 - w 1900 - E I I I I l l 0.a ,, 0 90 100 150 200 0 50 100 150 200 . TIMC (SECONDS) TIME (SECONDS)

                                                                                                                                               .f-i JOSEPH M. FARLEY          TRANS!ENT RESPONSC TO A DROPPED ROD CLU O          Atenema re er A                       NeCc      < r UNIT 1 AND UNIT 2 FIGURE 15.2-11 A-15.2 85

l

.1 3

                  .1      l*

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                         ..!                                                                                                                           l-
 .A                  C         'j                                                                                                                           '

U- E i t . .~ . 8 .! f b lt g .- a

                   -s W     .4-                                                                                                                         .

t3 : 5 .z. M g- '. ' 0. 0 1 .2 3 4 5 6 7 8 9 10 i TIME -tSEC1 I* ALL LOOPS INIT! ALLY OPERATING, ONE LOOP COASTING: JOSEPH M. FARL.EY DOWN - FLOW C0ASTDOWNS VERSUS TIME .

    .          Alabama Power ,                           NUCLEAR PL. ANT                                                                                 F UNIT 1 AND UNIT 2                                                       g4g                     l A-15,2-86
              -..+_..:.-.. . . . . - - -.. 2 - _ :       . - :.: .__2 . ._. a. = ;- . . . = . _ , ._.....;...                                        ,

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l 1

                 ..a-                                                                                                 l C

G < g .. z La O ! u * (~'s, D t l ( / $ .t

n

! O cr l g 4' d E

                  .2-0.

0 1 2 3 4 5 6 7 8 9 10 T 1 t1E ISCC1 l l l l l JOSEPH M. FARLEY ALL LOOP 5 IN111 ALLY OPERATING. ONE LOOP COAST!NG' Alabama Power NUCLEAR PLANT

UNIT 1 AND UNIT 2 FIGURE 15.2-13 A-15.2-87 i l

O i i 1 . 21 ' 0 .

6 E

w 1 1 1 . ; 1 5; d g 1100. i , O i # w l' i

    $ 1000.+                                                                                                                                      l C

E 1900.< 1000. 5. ti . 7 8. 9. 10. D. 1. 2. 3. 4 Tine (SEC) , i JOSEPH M. FARLEY ALL LOOPS INITIALLY OPERATING. ONE LOOP COASTIEG'

                                                                      ~~                                                  '       $

Alabama Power Q NUCLEAR PLAMT UNrf 1 AND UNIT 2 FIGURE 15.2-14 A-15.2-88

p / i

    -l
                                                                                                ! l l

l Il 1 i 1.- i

1. :

C.x g ..

i 5

I 8 .1 AVERAGE u i N CHANNEL r,T 't 5 l {

      /         ..

(_/ g 5 w

i 1 W 4' l w \ m i I g

                .2'                                                                          l I

O, i 0 1 2 3 4 5 6 7 8 9 10 TIME ISEC) l 1 JOSEPH M. FARLEY ALLLOOPSINITIALLYOPERATING,ONELOOPCOASTINGf Alabama Power NUCLEAR PLANT (~') k UNIT 1 AND UNIT 2 i FIGURE 15.2-15 i A-15.2-89

l j i l I i

i i

                                                                                                         , l i

i,

                'S      :                                                                              !

g i 5

i i l l s*E' HOT CHANNEL.

u ,w.

-                g LJ                b    *'                                                                        i     :

s ! d .d< s j g ; 6 I i

                    .2-                                                                               !

i l

0. l 0 1 2 3 4 5 6 7 8 9 10 ,

T ! r1E iSEC) i I s I ALL LOOPS INITI ALLY OPERATING, ONE LOOP COASTING-JOSEPH M. FARLEY f- "" " " " (s. Alabama Power NUCLEAR PLANT t i UNU 1 AND UNR 2 FIGURE 15.2-16 i A-15.2-90

\v I s i i

1. -

a. l - t t t , l

2. ?

M :

                                                                                                            ' i co                                                                                                    i z

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I 1.4 i

O 1 2 3 4 5 6 7 8 9 10 TIME (SEC) l l ALL LOOPS INITI Atty OPERAT]NG. ONE LOOP COASTING JOSEPH M. FARLEY , DowN . DNBR VERSUS TIME f'; Alabama Power NUCLEAR PLANT l UNU 1 AND UNU 2 , FIGURE 15.2-17 i i e A-15.2-91

1.4 t a

'w.J 31.2         '

5 Il 1. b N 'g. E. 5 .6-d E y 4 W O .2' C. --

                                   ,a 10            15               20          25            30 TIME        (SEC) 3.2        .
  ,~

i

  %.x 2.8    -

I 2.6 2.4 E 2.2-E 2. 1.s-1.6-1.4 1.2-10, 2.5 5. 7.5 10- 12.5 is. 17.5 20. 22.5 25. TIME (SEC) JOSEPH M. FARLEY STARTUP OF AN INACTIVE REACTOR COOLANT LOOP 17 I 17 CORE (~) V Alabama Power NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 15,2 18A A-15.2-92

1.4 r8 *

!,    )
  'v'            g 1.2-
                 =

5

                 =     1.

S W .- g E .6 - a o. E '4 5 B

.i*

0, 0 5 10 15 20 25 30 TIME (SEC) 2600.

 .Oj                2500.

[ .e.-c _ g

m a 2400, w

             @ 2300.

w E 2200. 5 N

             ;;; 2100.

b; m E! a 2000. i 1900.

                        '0. 5,   10.           15.          20.           25,          30.

TIME (SEC) JOSF.PH M. FARLEY STARTUP OF AN INACTIVE REACTOR COOLANT LOOP r~N 17 K 17 CORE v l Alabama Power NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 15.2-18B A-15.2-93

I 2.2 E' 3

<t.

1.8- { z 1.6- , 5

          =      1.4 3                                                                                                                        i b      1.2-                                                                                                              ,

e 1. 3 3 .s. d j t .6 8 .5

                   .2-O. 0      5    10               15                                      20                  25         30 TIME                            (SEC) 620

( 610-C

600-w

              $ 590-y 580 --

5 [ 570 0

              $ 560 5

g 550-8 j 540-e 530-520 0 5 10 35 20 25 30 TIME (SEC) A JOSEPH M. FARLEY STARTUP OF AN INACTIVE REACTOR COOLANT LOOP , Alabama Power O NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 15.2-18C A-15.2-94

i

1 1.4 ( l u/ 1.2-7 1 z 6 u 5 u.

                                                                  .6 62 k                   4-E w
                                                                  .2-0' 0    to                     20   30                           40                      50      60                            70                                      80     90     100 TIME          (SEC) 2600.                                                                                                                                                                                        .

[] f i 2.-. 2500. 1 t.s g - 2400. I u a

                                                 $             2300.

u

n. ,

2200. g u E 2100. Pn ' t0 g 2000.- j i 1900. l 800 0. 10. 20. 30. '0.

50. 60. 10, 80. 90. 100.

TIM 8! (SEC) JOSEPH M. FARLEY LOS$ OF LOAD ACCIDfNT WITH PRES $URIZER $ PRAY AND es POWER-OPERATED RELIEF VALVE BOL Alabama Power NUCLEAR PLAMT 3 (]Y - UNW 1 AND UNR 2 FIGURE 15.2-19A A-15.2-95

b N_-) f

5. ,

4.5 4.. 3.5 E 3. E ,/' + J 2.5- {W Z. 1.5-

1. d 10 20 30 40 50 60 70 80 90 100 TIME (SEC)

JOSEPH M. FARLEY LOSS OF LOAD ACCIDENT WiiH PRES $URIZER SP

(m ~-) 1 Alabama Power NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 15.2-19B A-15.2-96

O O e \. . ._ O

                                                                              -    O
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i bk ke_) ._ _O n r 1 O N _O-I i i i i ;- 0 0 0 W 0 n O v O n O o 5 m - 4 0 o o g) g3 e (d o) 3801V83dW31 WV31S JOSEPH M. FARLEY LOSS OF LOAD ACCIDENT klTH PRESSURIZER SPRAY AND POWER OPERATED RELIEF VALVE BOL (~~~'3 Alabama Power NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 15.2-19C A-15.2-97

7 1800. LN E 1600. [ w 3 1400. S 1 l 5 1200. Q

s 5

U 1000. I l E

           ?;                                                                                        >

w E 800. -

                   '0. 10. 20. 30. 40. 50. 60.      70. 80. 90,     100.

TIME (SEC) 700

,/ m I                680 (s_s)                                                                                                ,

C 660 S O 640- , N 620-E 600-u 8 580< i 560-0 10 20 30 40 50 60 70 80 90 100 TIME (SEC) , LOSS Of LOAD ACCIDENT VITH PRE 55URIZER SPRAY AND. JOSEPH M. FARLEY ,/ N. POWER-OPERATED RELIEF VALVE BOL V Alabama Power NUCLEAR PLANT UNU 1 AND UNR 2 FIGURE 15.2-20 A-15,2-98

1 1 1.4

                          '     1.2 5                                                                                                                                          i I

1.. w o E, . 1 , N 6 i g 4 w ' 4 r J R .2 l C i 0, 0 to 20 30 40 50 60 70 80 90 100 - TIME (SEC)

2600. O

'i

_ 2500. ; g ; w { 2400.. i w 5 m 2300, b, m w. E' 2200. l

                     .u w

N

                       ~
                            .2100.

m w 2000.. - , E I I 1900.. , 1800.O. -10. 20. 30. 40. 50. 60. 70. 80. 90. 100* : TIME (SEC) j t 5 LOSS OF LOAD ACCIDENT WITH PRES $URIZER SPRAY AN0i JOSEPH M. FARLEY POWER-OPERATED RELIEF VALVE . EOL t

 *O         Alabama Power                                        NUctsAn PLAur                                                                              !

UNU 1 AND UNU 2 FIGURE 15.2-21A > l A-15.2-99

   --.-.-..a.._.......--.                 - - - - .. - _ - , -                 . _ .._ -.- . - . - .-. _ -.-. - . - - ...- ... . . - . - ..

t s P r

                                                                                                                                                                           !        I 4

4.5 4 - 4 3.$4 3.' i 2.s l L 2..

1.5 j'

                                                  .1 0 10      20           30             40        50           60             70     80       90      100 TIME         (SEC) i l

t-. t t i JOSEPH M. FARLEY LOS$ OF LOAD ACCIDENT WITH PRES $UR!!ER SPRAY AN01 POWER OPERATED RELIEF VALVE EOL Alabama Power l: O NUCl.EAM PLM(T FIGURE 15.2.21B A-15.2 100

 .        -.. = - - . - - . . . . . . -            - . . . -  -n-._-                   . _ . . - - . - .           . . . . . . - . - .      .-.-.-,__.a.----..-...

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I i i i i l-o , O 8 8 ? o o o o o o n G $ g (d o) 3801VU3dW31 WV31S JOSEPH M' FARLEY loss 0F L W ACCIDENT WITH N $$URllER S M U AND 1 POWER OPERATED RELIEF VALVE EOL Alabama Power g nuetrAn etAnT

k. UNIT 1 AND UNIT 2 FIGURE 15.2-21C A.15.2 101

1800. S [ 1600.4

                                             .            i.00.                                                                                                                                                       ,-

O ; a . 1200. 2 5 i U 1000. E m ! i d A E 800. 600.O. _10.- 20. 30. 40. 50.- 60, 70.- 80. 90. 100. TIME - (SEC) 600 i d 590- ; i

                                             -                                                                                                                                                                     l
                                             "-               580                                                                                                                                                  !
                                          -c d

g 570- _ I W :i

                                          ' Y.                560-l
                                             <                                                                                                                                                                   t w

5 v

                                                              $50-540 c                                                                                                                                              l t         to -   20    30-              40        50                    60     70-              80                   90-    100               ,

TIME (SEC) .

i JOSEPH M. FARLEY LOSSOFLOADACCIDENTWITHPRESSURIZERSPRAYAN0!

                                                                                                                            ~~- -                                                        -"

l0 l

                          ^tebem e ee er A                                                c' UNIT 1 AND UNIT 2
                                                                                                            - >~1                                                                                               >

FIGURE 15.2-22 A 15.2-102

- ,

r. - - . - + w . + , - --,-r.-we, * , ew.-+-e,-w. m s. ~,.-541-, 4 w.w.. -.-,-.3%s.---- ,.m.e--~ ,,,,*wne. -.%,.ww-a'-w-+,vu 4,3, w r e r- = . - - -

k m CT* n. PRESSURIZER PRESSURE (PSIA) NUCLEAR POWER (FRAC OF HOMINAL) h m N N N N a N N N * -*

a a N SA &
y e O O O O O O * * * * # *
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                                                                                -~'~ ~ ~ "

l l lO 1 5 4.5 4 3.5-E '- s O 2.5-

2. 1 1.5 10 10 20 30 40 50 60 70 to go too TIME (SEC)

I JOSEPH M. FARLEY LOSS OF LOAD ACCIDENT WITHOUT PRESSURIZER SPRAY L EF VAL M . M L Alabama Power NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 15.2-23B 1 I A-IS.2-104

I

                    -   .                                                     _                   g                                                   l l     )

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l O i 6 N m. I I- O i i l i I o o o o o o o N W D v m N - D D D D D D D (.d)3tf01VB3dW31WV31S o JOSEPH M. FARLEY LOS$ OF LQAD ACCIDENT WITHOLIT PRES $URIZER SPRAY Alabama Power NUCLMR Pt. ANT (~)s

  '-                              UNIT 1 AND UNIT 2 FIGURE 15.2-23C A 15.2 105

1800. lO _ 1600. Y b w 1400. E s> ; g 1200. Y s ' cc 1000. E 3 '

         $   800.

u o. 600.D. 10. 20. 30. 60. 50. 60 70. 50. 90* 100' TIME (SEC) 600

'O       _

5-1 8 ; S 580 a. 5* 170 2 w g 560 550 l 540 0 10 20 30 'O 50 60 70 30 90 100 TIME (SEC) , JOSEPH M. FARLEY LOSS OF LOAD ACCIDENT WITHOUT PRESSURIZER $ PRAY l R- P ELIEF VALVE - M L Alabama Power NUCLEAR PLANT UNIT 1 AND UNf7 2 FIGURE 15.2-24 i A-15.2-106

1.4 s

    -]                                                                                                 j 3    1,g, M

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            $  2100.+

d ce

            @  2000.<

ce !

            "  1900.                                                                                !

l 1800.O. 10. 20. 30. 40. 50. 60. 70. 80. 90. 100. TIME (SEC) t LOSS OF LOAD ACCIDENT WITHOUT PRE 55URIZER SPRAY JOSEPH M. FARLEY AND POWER OPERATED RELIEF VALVE EOL i (3 Alabama Power NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 15.2-25A A-15.2-107

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                                                                                                          . TIME             (MD                                                             i s

i I I i JOSEPH M. FARLEY LOS$OFLOADACCIDENTWITHOUTPRES$URIZERSPRAYh AND POWER.0PERATED RELIEF VALVE EOL b] - Alabama Power NUCLIAR PLANT UNIT 1 AND UNIT 2 FIGURE 15.2-258 l l + A 15.2-108

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                  'L>

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   'm                 g                 0      10     20          30           40          50          60        70                                             80 s                 ;f                                                                                                                                                                90                   100 M                 ~3                                                                 TIME (SEC)
                     -y f* M
                          =

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                      . D. 20, 30. 40     50. 60,             70,   H.                        90,                   tu, 600 i

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                                'suc1";;" in'esch5cniin!JlltlJatsstarzoa s O   aie8ema Power g            UNIT 1 AND UNIT 2 FIGURE 15.2-26 A-15.2-110

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                                                                                                                     .. T!nE           ISEC) f

[0$$ OF #cAMAL FCEDWATER - PRCS$URIZER PRES $URE I

JOSEPH M. FARLEY ** Ltvtt vtRsus nnt ' '

Alabama Power 6 NUCLEAR PUNT A UNrr1 AND UNIT 2 FIGURE 15.2-27A A 15.2-11]

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       ;                                 NUC1. EAR PLANT UNIT 1 AND UNIT 2 FIGURE 15.2-27B l

A-15.2-ll2

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                           ,                                                            STEAM GENERATORS RECEIVING o-                                                              ' AUXILIARY FEEDWATER h .bE*0$
                        .M a                                                                                       v                                                    .,
                        . g .JE-05; W
                                                ,-    STEAM GENERATOR NOT. RECEIVING                                                                                      i.

c AUXILIARY FEEDWATER t

                                     ,g, g
                                         - e'                                                                                              -

10 C 10 1 10 E- 10 1 . TIME tSEC1 .; JOSEPH M. FARLEY LOSS OF NORMAL FEEDWATER - STEAM GEN PRESSURE AND MASS VERSUS TIME ( Alabama- Power NUCLEAR PLANT 3 UNR 1 AND UNU 2 - FIGURE 15.2-27D A-15.2-114

                                                                                                                                   --,       + -

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          $   1800 7 1600.                                                                                                                         10' 10 0  10  1 102                  10 3 Tine        tSECi LOSS OF ALL AC POWER TO THE STATION AUXILIARIES JOSEPH M. FARI.EY n                                                      PRESSURIZER PRESSURE AND LEVEL VER$US TIME t

,y. i Alabama Power NUCLEAR PLANT UNIT 1 AND UNIT 2 FIGURE 15.2-27E A-15.2-Il5

f m C7* m l a CORE llEAT FLUX (FRAC OF NOH) NUCLEAR POWER (FRAC 0F NOM) g

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LOSS OF ALL AC POWER TO THE STATION AUXILIARIES JOSEPH M. FARLEY LOOP 1, 2, AND 3 HOT AND COLD LEG TEMPERATURE Alabama Power NUCLEAR PL.Am VER$US TIME UNIT 1 AND UNIT 2 FIGURE 15.2-27G A-15.2-Il7

1 It00. i (y $- Q 1400. 1- ! 1200,t N 1000,

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                                   .2E+0s-STEAM GENERATOR NOT RECEIVING AUXILIARY FEEDWATER.
                                                                                                                        'f                                                  I O.

10 0 gol 302 3o3 10' TIME (SEC1 LOSS OF ALL AC POWER TO THE STATION AUXILIARIES'- JOSEPH M. FAR12Y A STEAri GENERATOR PRESSURE AND MASS VERSUS TIME-O Alabama Power g NUC1. EAR PLANT UNR1 AND UNE2 FIGURE 15.2-27H A-15.2-118 4

       ...__.___._______m.                         . _ . _ _ _    _ . . . _ . . _      _ . . . _ _ _ _ . - . _ _ _ . . _ _ . _ _ .

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1 E 7. i EXCESSIVE LOAD INCREASE WITHOUT CONTROL ACTION -

JOSEPH M. FARL.EY BOL - DNBR. NUCLEAR PCWER, AND PRESSURIZER i Alabama Power NUCLEAR PLANT PRESSURE AS A FUNCTION OF TIME UNIT 1 AND UNIT 2 FIGURE 15.2-29 A-15.2-122

I

  /^g .                    1200.                                                                                                                                                      I

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            >                                                                                                                                                                         \

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            $               700.<

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           .E             580 r

5 w 570-

           - t.o                                                                                                                                                                 3 5-g           560.-

Eo 550 v , l. i 540 0 -- 5 0 100 150 200 250 300 350 400 ! TIME (SEC) 4

                                                                    .                        EXCESSIVE LOAD INCREASE WITHOUT CONTROL ACTION -

JOSEPH M. FARLEY BOL - TAVG AND PRE 55URIZER VOLUME AS A FUNCTIONi ; O Alabama Power NUCLEAR PLANT OF TIME j ~L) UNIT 1 AND UNIT 2 g A-15.2-123

3. 2.5-i l N 2. m g ? I o 1.5< i 200 25) 300 150 400 1'O 50 100 l' 0 T ttE iiE..

               - 1.4
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u. m o i b2J o 1. [ l v - %\ m & z o l N t D .8 y , 0 50 100 ;50 200 250 300 350 400 i T 4 . c. =SECi 9 w e 2400. m m m w E _ 2300. m< I dm \ E m 2200. I. m m w m C. . O. 50. 100. 150. 200. 250. 300. 350. 400. . T 1 f1E iiE 1 EXCESSIVE LOAD INCREASE WITHOUT CONTROL ACTION --! JOSEPH M. FARW EOL - DNBR, NUCLEAR POWER, AND PRESSURIZER ! Alabama Power NUCLEAR PLANT PRESSURE AS A FUNCTION OF TIME , UNIT 1 AND UNIT 2 FI;URE 15.2-31 A-15.2-124

1 120). ' 10 - i100. w 100). B W 90).4 ' 2 l w B0). 9 a h 700. E 609. O. 50. 100. 150. 200. 250. 209. 350. 400.

                                        ?!ME        'IEC*

610 l 0 C v 600 I

590 ! w 580 N . 5

570 560 W

        <C I

g 550 8 540 0 50 100 150 200 250 300 350 400 , TIME '5EC, i EXCESSIVE LOAD INCREASE WITH00T CONTROL ACTION - JOSEPH M. FARLEY EOL - TAvG AMD PRE 55uRIIER V0WME AS A FUNCTION Alabama Power " nue: EAR PLANT O UNIT 1 AND UNIT 2

   ,                                                                   FIGURE 15.2-32
                                                                                                                                                  \

A-15.2-125

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TIME i3E;j i EXCES$1VE LOAD INCREASE WITH REACTC2 CONTROL g JOSEPH M. FARLEY BOL DNBR. NUCLEAR POWER, AND PRES $URIZER i p Alabama Power NUCLEAR PLANT PRESSURE M A FUNCTION OF TM UNIT 1 AND UNIT 2 FIGURE 15.2-33 A-15.2-126

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

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                                  .0. 50,                  100,         150.         200.        'E50.         300.        350.        400, i ! f*E         IIEC) i 620                                                                                                                      ,

p* 610- .[ Ox l-Y m' 600 f t== i 5 w -590 7- ; i

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5-w. 570- [ alC .- w- 560 ct: ! o u i 550-540 300 350 400 0 50 10 0 ~ 150 200 250 4) T*HE (SECi .. l i 1 i EXCESSIVE LOAD INCREASE WITH REACTOR CONTROL - JOSEPH M. FA9 LEY BOL - TAVG AND PRESSURIZER VOLUME AS A FUNCTION OF TIME 4 Alab'ama Power NUCLEAR PLANT i UNIT 1 AND UNIT 2 FIGURE 15.2-34 A-15.2-127

l 2,5 O V - 2. ( I 1.5-l 1 1. 0 50 100 150 200 250 300 150 40c . TIME (3E*1 l.4 J I 5- 1.2< at 8 *~ I5 o g- i.. P

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T!rE (S E C ) EXCESSIVE LOAD INCREASE WITH REACTOR CONTROL - i JOSEPH M. FARL.EY EOL DN8R. NUCLEAR POWER. AND PRES $URIZER ! Alabama Power ' # ^ " " O Nect. EAR PI. ANT e~,1, ~e e~,1. FIGURE 15.2-35 A-15.2-128

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     .f.                       1200.

n w j 1100, w t b g 1000. , 5 r 900. l 3 cz: _- , d 800.

                          .E S                                                                                                                                   i i
                            $-  700, E

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15.3 CONDITION 111 - INFRE0VM1 FAULTS O By definitici, Condition 111 occurrences are faults which may occur infrequently during the life of the plant. They will be accommodated with the failure of only a small fraction of the fuel rods although sufficient fuel damage might occur to preclude resumption of the operation for a considerable outage time. The release of radioactivity will not be sufficient to interrupt or restrict public use of these areas beyond the ex';1usion radius. A Condition 111 f ault will not, by itself, generate a Condition IV fault or result in a consecuential loss of function of the reactor coolant system (RCS) or containment barriers. For the purposes of this report, the following faults have been grouped into this category. A. Loss of reactor coolant from small ruptured pipes or from cracks in large pipes which actuate the emergency core cooling system (ECCS) (This group of events is described in Appendix B of the Safety Assessment.) B. Minor secondary system pipe breaks C. Inadvertent loading of a fuel assembly into an iraproper position D. Complete loss of forced reactor coolant flow E. Waste gas decay tank rupture - this is a radiological event (The radiological evaluation of VANTAGE 5 fuel is described in Appendix C of the Safety Assessment.) F. Single rod cluster control assembly (RCCA) withdrawal at full power 15.3.1 LOSS Of REACTOR COOLANT FROM SMALL RUPTURED PIPES OR FROM CRACKS IN LARGE PIPES WillCH ACTUA1E EMERGENCY CORE COOLING SYSTEM Evaluation or analysis of this group of events is described in Appendix B of the Safety Assessment. A-15.3-1

15.3.2 HINOR SECONDARY SYSTEM PIPE BREAKS 15.3.2.1 Identification of Causes..ard Accident Descriotion included in this grouping of ruptures of secondary system lines which would result in steam release rates equivalent to a 6-inch diamete- or smaller break. - 15.3.2.2 Analysis of Effects and Conseauences Minor secondary system pipe breaks must be accommodated with the failure of only a small fraction of the fuel elements in the reactor. Since the resultc of analyses presented in Section 15.4.2 for major secondary system pipe rupture also meet this criteria, separate analysis for minor breaks is not required. The analysis of the more probable accidental opening of a secondary system steam dump, relief, or safety valve is presented in Section 15.2.13 and is illustrative of a pipe break equivalent in size to a single valve opening. , 15.3.2.3 Conclusions The analysis presented in Section 15.4.2.1 demonstrates the consequences of a minor secondary pipe break is acceptable since a departure from nucleate boiling ratio (DNBR) of less than the limit value does not occur even for the

     . more critical major secondary system pipe break.

15.3,3 INADVERTENT LOADING 0F A FUEL ASSEMBLY INTO AN IMPROPER POSITION Fuel type does not affect the ability of the in-core instrumentation to

. detect the inadvertent loading and subsequent operation with a fuel assembly

! in an improper position; therefore, the conclusions presented in FSAR i Section 15.3.3 remain valid. t-A-15.3-2 l .

15.3.4 COMPLETE LOSS OF FORCED REACTOR COOLANT FLOW O 15.3.4.1 Identification of Causes and Accident Descriotion A conplete loss of forced rer.ctor coolant flow may result from a simultaneous loss of electrical supplies to all reactor coolant pumps. If the reactor is at power at the time of th9 accident, the immediate effect of loss of coolant flow is a rapid increase in the coolant temperature. This increase could result in DNB with subsequent fuel damage if the reactor was not trip.ned promptly. Normal power for the pumps is supplied through separate buses from a transformer connected to the generator. When a generator trip occurs, the buses are automatically transferred to a transformer supplied from external power lines so that the pumps will continue to provide forced coolant flow to the core. Following any turbine trip where there are no electrical faults which require tripping the generator from the network, the generator remains connected to the network for approximately 30 seconds after reactor trip before any transfer is made. , The following signals provide the necessary protection against a complete loss of flow accident. A. Undervoltage or underfrequency on reactor coolant pump power supply buses B. Low reactor coolant loop flow C. Pump circuit breaker opening The reactor trip on reactor coolant pump undervoltage is provided to protect against conditions which can cause a loss of voltage to all reactor coolant pumps, i.e., loss of offsite power. This function is blocked below approximately 10 percent power (Permissive P-7). See FSAR Table 7.2.2 for a definition of permissive setpoints. O v A-15.3-3

The reactor coolant pump underfrequency function is provided to trip the reactor for an underfrequency condition resulting from frequency disturbances i on the power grid. The reactor coolant pump underfrequency reactor trip function is blocked below P-7. In addition, the underfrequency function will

     ;n all reactor coolant pump breakers whenever an underfrequency condition occurs (no P-7 or P-8 interlock) to ensure adequate RCP pump coastdown.

The reactor trip on low primary coolant loop flow is provided to protect against loss of flow conditions which affect oni,' one reactor coolant loop. It also serves as a backup to the undervoltage a1d underfrequency trips. This function is generated by two-out-of-three l >w flow signals per reactor coolant loop. Above Permissive P-8, low flow in any loop will actuate a reactor trip. Between approximately 10 percent oower (Peniiissive P-7) and the power level corresponding to Permissive P-8, low flow in any two loops will actuate a reactor trip. A reactor trip from pump breaker position is provided as a backup to the low flow signal. Similar to the low flow trip, above P-8, a breaker open signal from any pump will actuate a reactor trip, and between P-7 and P-8, a breaker open signal from any two pumps will actuate a reactor trip. Reactor trip on reactor coolant pump breakers open is blocked below Permissive 7. 15.3.4.2 Method of Analysis This transient is analyzed by three digital computer codes. First, the LOFTRAN (Reference 1) code is used to calculate the loop and core flow transients, the nuclear power transient, and the primary system pressure and temperature transients. The FACTRAN (Reference 2) code is then used to calculate the heat flux transient based on the nuclear power and flow from LOFTRAN. Finally, the THINC code is used to calculate the DNBR during the transient based on the heat flux from FACTRAN and the flow from LOFTRAN. The DNBR transient presented represents the minimum of the typical and thimble cell s . One case has been analyzed: Loss of all three reactor coolant pumps with three loops in operation. I A-15.3-4

                                           -          - - - - - -                 _J

The method of analysis and the assumptions made regarding initial operating conditions and reactivity coefficients are identical to those discussed in Section 15.2, except that following the loss of power supply to all pumps at power, a reactor trip is actuated by either reactor coolant pump bus undervoltage or underfrequency. Note, with respect to the reactivity coefficient assumptions, the analysis conservatively assumes a moderator temperature coefficient of +7 pcm/*f. The accident is analyzed using the Revised Thermal Design Procedure, initial core power, reactor coolant temperature, and pressure t.re assumed to be at their nominal values consistent with steady-state full-power operation. Uncertainties in initial conditions are included in the limit departure from nucleate boiling ratio (DNSR) as described in WCAP-ll397 (Reference 3). This event is assumed to be initiated by an undervoltage disturbance. 15.3.4.3 Results The calculated sequence of ever<ts , shown in Table 15.3-1. Figures 15.3-20 through 15.3-25 show the transient response for the loss of power to all reactor coolant pumps for the limitirg loss of flow event. The reactor is assumed to be tripped on an undervoltage signal. The DNBR-versus-time plot represents the limiting cell for three-loop coastdown. 15.3.4.4 Conclusions The analysis performed has demonstrated that for the complete loss of forced reactor coolant flow, the DNBR does not decrease below the safety analysis limit value at any time durine the transient. Thus, no fuel or clad damage is predicted, and all applicable acceptance criteria are met. 15.3.5 WASTE GAS DECAY TANK RUPTURE This Condition 111 event is not analyzed as part of the non-LOCA safety evaluation. The radiological assessment of VANTAGE 5 fuel is described in Appendix C of the Safety Assessment. O A-15.3-5

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

15.3.6 SINGLE RCCA WITHDRAWAL AT FULL POWER O V 15.3,6.1 Accident Description No single electrical or mechanical failure in the rod control system could cause the accidental withdrawal of a single rod cluster control assembly (RCCA) from the inserted bank at full-power operation. The event analyzed must result from multiple wiring failures, multiple significant operator errors, or subsequent and repeated operator disregard of event indication. The probability of such a combination of conditions is low such that the limiting consequences may include slight fuel damage. Each bank of RCCAs in the system is divided into two groups of four mechanisms each. The rods comprising a group operate in parallel through multiplexing thyristors. The two groups in a bank move sequentially such that the first group is always within one step of the second group in the bank. A definite schedule of actuation (or deactuation) of the stationary gripper, moveable gripper, and lift coils associated with the four RCCA4 of a rod group are driven in parallel. Any single failure which would cause withdrawal would affect a minimum of one group, or four RCCAs. Mechanical failures are in the direction of insertion or immobility. Note, the operator can deliberately withdraw a single RCCA in a control or shutdown bank since this feature -is necessary in order to retrieve an assembly should one be accidentally dropped. In the unlikely event of simultaneous electrical failures which could result in single RCCA withdrawal, the plant annunciator will display both the rod deviation and rod control urgent failure, and the rod position indicators will indicate the relative positions of the RCCAs in the bank. The urgent failure alarm also inhibits automatic rod motion in the group in which it occurs. Withdrawal of a single RCCA by operator action, whether deliberate or by a combination of errors, would result in activation of the same alarm and the same visual indication. The OTAT reactor trip provides automatic protection for this event, although due to the increase in local power density, it is not possible to always provide assurance that the core safety limits will not be exceeded. A-15.3-6

I l 15.3.6.1.1 Method of Analysis O Power distributions are analyzed using appropriate nuclear physics computer J codes. The peaking factors are then used as input to the THINC code to calcul:te the DNBR. The analysis examines the case of the worst rod withdrawn from bank D, inserted at the insertion limit with the reactor j initially at full power. 15.3.6.1.2 Results I The analyr~ the single rod withdrawal event considers the following two casos, i A. If the reactor is in the manual rod control mode, continuous withdrawal of a single RCCA 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 withdrawing PCCA. In terms of the overall system response, this case is similar to those presented in Section 15.2.2; however, the increased local power peaking in the area of the withdrawn RCCA results in lowcr DNBRs than for the withdrawn bank cases. Depending on initial bank insertion and location of the withdrawn RCCA, automatic reactor trip may not occur quickly enough to prevent the minimum DNBR from falling below the safety analysis limit value. Evaluation of this case at the power i and coolant conditions at which the OTAT trip would trip the plant shows that an upper limit for the number of rods with a DNBR less than the safety analysis limit value is 5 percent. B. If the reactor is in the automatic rod control mode, the multiple failures that result in the withdrawal of a single RCCA cause immobility of the other RCCAs in the controlling bank. The transient will then proceed in the same manner as case A, described above. For such cases, a reactor trip will ultimately ensue, although not quickly enough in all cases to prevent a minimum DNBR in the core of less than the safety analysis limit value. Following reactor trip, normal shutdown procedures are followed. l A-15.3-7 ! l

15.3.6.2 Conclusions for the case of the accidental withdrawal of a single RCCA, with the reactor in the automatic or manual control mode and initially operating at full power with bank D at the insertion limit, an upper bound of the number of fuel rodt experiencing DNBR is 5 percent of the total fuel rods in the core. For both cases discussed, the indicators and alarms mentioned would function to alert the operator to the malfunction before DNB would occur. For C1:9 B, discussed above, the insertion limit alarms (low and low-low alarms) would serve in this regard. { O t A-15.3-8

i 15.

3.7 REFERENCES

(}

1. T. W. T. Burnett, et al., "LOFTRAN Code Description, "WCAP-7907-P-A (Proorietary) and WCAP-7907-A (Nonoronrietary), April 1984.

2. H. G. Hargrove, "FACTRAN - A FORTRAN-IV Code for Thermal Transiehts in a U0 2Fuel Rod," WCAP-7908-A, Dece:nber 1989.

t

3. A. J. Friedland and S. Ray, " Revised Thermal Desit /rocedure,"

VCAP-11397-P-A, April 1989. l l l l t f n l A-15.3-9 1

i O TABLE 15.3-1 TIME SEQUENCE Of EVENTS FOR CONDITION 111 EVENTS Accident {ycal Time (sec) Complete loss of forced reactor coolant flow All pumps in Reactor coolant pump 0.0 operation, all undervoltage trip setpoint pumps coasting reached, all pumps lose down power and begin coasting down Rods begin to drop 1.5 Minimum DNBR occurs 3,5 O A-15.3-10

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15.4 CONDITION IV - LIMITING FAULTS O Condition IV occurrences are faults that are not expected to occur, however they are postulated because their consequences could potentially release significant amounts of radioactive material. Condition IV events represent the most drastic occurrences that the plant must be designed for and consequently represent the limiting design cases. Condition IV faults must not cause a fission product release to the environment resulting in an undue risk to public health and safety in excess of the requirements of 10 CFR 100. A single Condition lY fault must not cause a consequential loss ; of the required functions of systems needed to cope with the fault including those of the Emergency Core Cooling System (ECCS) and containment, f l For the purposes of this report, the following faults have been classified in this category.- ' A. Major rupture of pipes containing reactor coolant up to and including double-ended rupture of the largest pipe in the Reactor Coolant System (RCS) (Loss Of Coolant Accident (LOCA)) (Discussed in Appendix B of the Safety Assessment.) B. Major secondary system pipe rupture up to and including double-ended rupture of a steam pipe C. Steam generator tube rupture D. - Single reactor coolant pump locked rotor E. Fuel handling accident (FHA) F. Rupture of a control rod mechanism housing (Rod Cluster Control Assembly (RCCA) ejection) 15.4.1 MAJOR REACTOR COOLANT SYSTEM PIPE RUPTURES (LOSS-Of-COOLANT ACCIDENT) h Evaluation or analysis of this class of events is considered in Appendix B of the Safety Assessment. A-15.4-1

15.4.2 MAJOR SECONDARY SYSTEM PIPE RUPlVRE O Two major secondary system pipe ruptures are presented in this section: (1) rupture of a main steamline (reanalyzed) and (2) rupture of a main feedwater pipe (evaluated). 15.4.2.1 Ruoture af a Main Steamline , 15.4.2.1.1 Identification of Causes and Accident Description The steam release arising from a rupture of a main steamline will result in an initial increase in steam flow that decreases during the accident as the steam pressure falls. The energy removal from the RCS 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. If the most-reactive RCCA is assumed stuck in its fully withdrawn position after reactor trip, there is an increased possibility that the core will become critical and return to power. A return to power following a O steam pipe rupture is a potential problem mainly because of the high power peaking factors that would exist assuming the most-reactive RCCA to be stuck in its fully withdrawn position. The core is ultimately shut down by the boric acid injection delivered by the ECCS. For a double-ended rupture of a main steamline, the radiation releases must remain within the requirements of 10 CFR Part 100. These requirements represent the ANSI N18.2 criteria for Condition IV events, " Limiting Faults." The criteria are conservatively met by demonstrating that the DNB design basis is met, a criterion typically applied to Condition II events. Therefore, a main steam pipe rupture is analyzed to demonstrate that the following criteria are satisfied.

1. Assuming a stuck RCCA (with or without offsite power) and assuming a single failure in the engineered safety features (ESF), there is no consequential damage to the primary system and the core must remain in place and intact. Radiation doses do not exceed the guidelines cf 10 CFR 100.

A-15.4-2

(3 2. Although DNB and possible cladding perforation following a steam v} pipe rupture are not necessarily unacceptable, the rupture of a major steamline analysis, in fact, demonstrates that no DNB occurs for any rupture assuming the most-reactive assembly is stuck in its fully withdrawn position. The rupture of a major steamline is the most-limiting cooldown transient. It is analyzed at zero power with no decay heat since decay heat would retard l I the cooldown, thus reducing the return to power. A detailed discussion of this transient with the most-limiting break size (a double-ended rupture) is presented below. The following functions provide the necessary protection against a steam pipe rupture. l A. Safety injection system actuation from any of the following: [D 'V

1. Two-out-of-three low pressurizer pressure signals,

2. High steamline differential pressure,

3. Low main steamline pressure in two-out-of-three steamlines,

4. Two-out-of-three high-1 containment pressure signals.

B. The overpower reactor trips and the reactor trip occurring in conjunction with receipt of the Safety injection (SI) signal. C. Redundant isolation of the main feedwater lines: to prevent sustained high feedwater flow which would cause additional cooldown. Therefore, in addition to the normal control action which will close the main feedwater valves, a safety injection signal will rapidly close all feedwater control valves, trip the main feedwater pumps, and indirectly close the feedwater isolation valves that 9 backup the control valves. In addition, trip of the steam generator (G A-15.4-3

i feedwater pumps results in automatic closure of the respective pump discharge isolation valve. ; D. Trip of the fast-acting Main Steamline Isolation Valves (MSIVs, l designed to close in less than 5 seconds) and Main Steamline Isolation Bypass Valves (MSIBVs, designed to close in less taan 10 seconds) after receipt of an ECCS or main steamline isolation signal on:

1. High steam flow in two-out-of-three main steamlines (one-of-two l per line) coincident with two-out-of-three low-low T avg signals, 2.- Low steamline pressure signal on two-out-of-toree steamlines, l

3. Two-out-of-three high-high (hi-2) containment pressure signals. )

l for breaks downstream of the isolation valves, closure of all valves will

completely terminate the blowdown. For any break, in any location, no more that one steam generator would experience an uncontrolled blowdown even if one of the isolation valves fails to close. Circuit design assures that the MSIBVs are closed whenever the MSIVs are closed. Following a steamline break, only one steam generator can blow down completely. Each main steamline is provided with two isolation valves located outside of the containment immediately downstream of the steamline safety valves. The isolation valves are signal-actuated valves which close to prevent flow in the normal (forward) flow direction. The valves on all three steamlines will be driven closed and isolate the other steam gP,nerators . Thus, only one steam generator can blowdown, minimizing the-potential steam release and resultant RCS cooldown. In addition, the remaining two steam generators.will still be available for dissipation of any decay heat after the initial transient is over. In the case of loss of offsite power, this heat is removed to the atmosphere via the atmospheric dump valves which have been sized to handle this situation. A-15.4-4

Steam flow is measured by monitoring the pressure difference between pressure Q b taps located in the steam drum and downstream of the integral flow restrictor nozzles. The effective throat diameter of the flow restrictor nozzles of 14 inches is considerably smaller than the diameter of the main steam pipe. These restrictors are located in the steam generators outlet nozzle and serve to limit the maximum steam flow for any break at any location. l 15.4.2.1.2 Analysis of Effects and Consequences l 15.4.2.1.2.1 Method of Analysis The analysis of the steam pipe rupture has been performed to determine: A. The core heat flux and RCS temperature and pressure resulting from the cooldown following the steamline break using the LOFTRAN code

                                 .(Reference 1),

B. The thermal and hydraulic behavior of the core following a steamline break using the detailed thermal and hydraulic digital computer code, THINC, to determine if DNB occurs for the core conditions computed in (A) above. The following conditions were assumed to exist at the time of a main steamline break accident. A. End-of-Life (EOL) shutdown margin at no-load, equilibrium xenon conditions, and the most-reactive assembly stuck in its fully withdrawn position. Operation of the control rod banks during core burnup is restricted in such a way that addition of positive reactivity in a steamline break accident will not lead to a more-adverse condition that the case analyzed. B. The negative moderator coefficient corresponding to the EOL rodded core with the most-reactive rod in the fully withdrawn position. Tne variation of the coefficient with temperature and pressure has been included. The keff versus coolant average temperature at A-15.4-5 .-- - -_-.--..- _ - -- .- ...-.-._. - ~. . . . - . -. . . - . -

/] 1000 psia corresponding to the negative moderator temperature coefficient plus the Doppler temperature effect used is shown in Figure 15.2-40A. The effect of power generation in the core on overall reactivity is shown in Figure 15.2-40B. The core properties associated with the sector nearest the affected steam generator and those associated with the remaining sectors were conservatively combined to obtain average core properties for reactivity feedback calculations. Further, it was conservatively assumed that the core power distribution was uniform. These two conditions cause underprediction of the reactivity feedback in the high-power region near the stuck rod. To verify the conservatism of this method, the reactivity as well as the power distribution was checked. These core analyses considered the Doppler reactivity for the high fuel temperature near the stuck RCCA, moderator feedback from the high water enthalpy near the stuck RCCA, power redistribution and nonuniform core inlet temperature effects. For cases in which steam generation occurs in the high flux regions of the core, the effect of void formation was also-included. It was determined that the reactivity employed in the kinetics analysis was always larger than the true reactivity verifying conservatism; 1.e., underprediction of negative reactivity feedback from power generation. C. Minimum capability for injection of high concentration boric acid (2300 ppm) solution corresponding to the most-restrictive single failure in the ECCS. The 2300 ppm boron solution corresponds to the minimum boron concentration in the Refueling Water Storage Tank (RWST). A boric acid solution of 0 ppm is assumed in the Boron Injection Tank (BIT). -No credit has been taken for the low concentration of boric acid that must be' swept from the ECCS lines downstream of tne RWST isolation valves prior to the delivery of the concentrated boric acid (2300 ppm from the RWST) to the reactor coolant loops. A-15,4-6

The safety injection curve _ assumed is shown in Figure 15.2-41. The flow corresponds to that delivered by one charging pump delivering 4 - . full- flow to the cold leg header. The variation of the mass flowrate due to' water density changes is included in the calculations, as is the variation in flowrate in the ECCS due to changes in the RCS pressure. The ECCS flow calculation includes the line losses as well as the safety injection pump head curve. The modeling of the ECCS in LOFTRAN is described in Reference 1. The boric acid solution from the ECCS is assumed to be uniformly )i

     ,                                                              delivered to the three reactor coolant loops. The boron in the loops is then delivered to the inlet plenum where the coolant (and boron). from each loop is mixed and delivered to the core. The                                                                                                                         j calculation assumes the boric acid is mixed with and diluted 'by the .                                                                                                                 1 water flowing in _the RCS prior to entering the core. The                                                                                                                               !

concentration after mixing depends on .the relative flowrates of the i RCS:and-the ECCS. The stuck RCCA is assumed to be conservatively

                                                                  --located in the core sector near the faulted steam generator.                          -

U ~ For the case where offsite power is assumed, the-sequence of events ; in'the ECCS is the following. After the generation of the SI signal j (appropriate delays for instrumentation, logic, and signal transport ;

                                                                  ' included),' the appropriate valves begin- to operate' and the high-head injection pump starts. In 27 seconds, the valves are assumed to be in the final position and the pump is assumed to be at full ' speed                                                                                                                 a
                                                                 - and to be drawing suction from the RWST. The 27 seconds can be                                                                                                                          )

assumed to include 2 seconds for electronic delay,10 seconds for . the RWST valve to open, and 15 seconds for.the VCT valve to close. Theivolume containing the low concentration (0 ppm assumed) barated water.is swept from the ECCS before the 2300 ppm baron reaches the. core. This delay in' 2300 ppm boron solution reaching the core is inherent 1y included in the LOFTRAN modeling. In cases-where offsite power is'not available, an additional second delay is assumed to start the diesels and to reenergize the ESF electrical buses. That is, after a total of 42 seconds A-15.4-7 w w' w-a gE . s-'W'd4-4N-'F9' Pf'd-e'--N'r owP-+-- - -ww1e VW a' ese, r Ve-hws 9 Wea w w w- 1et en4wvW1f' - 'e 294==5-SMe e-'- emg s--Tmy*-e=gtwg'gu--*=v=g-ggf-meew m -at eqqr' w per pq-WW we w gule w

I following the time an SI setpoint is reached at the sensor, the ECCS is assumed to be capable of delivering flow to the RCS. LJ D. The design value of the steam generator heat transfer coefficient includes a tube fouling factor allowances. To maximize primary-to-secondary heat transfer, 0% tube plugging is assumed. E. Since the steam generators are provided with integra; flow restrictors with a 1.061 ft2 throat area, any rupture with a break greater than 1.061 ft 2, regardless of the location, would have the same effect on the nuclear steam supply system as the 1.061 ft2 break. The following two cases have been considered in determining the core power and RCS transients.

1. Complete severance of a pipe, with the plant initially at no-load conditions, and full reactor coolant flow with offsite power available, and

 's    2. Complete severance of a pipe with the plant initially at no-load L           conditions with offsite power unavailable. Loss of offsite power ..

results in coolant pump coastdown. F. Power peaking factors corresponding to one stuck RCCA and nonuniform core inlet coolant temperatures are determined at EOL. The coldest core inlet temperatures are assumed to occur in the sector with the stuck rod. The power peaking factors account for the effect of the local void in the region of the stuck control assembly during the return-to-power phase following the steamline break. This void, in conjunction with the large negative moderator coefficient, partially offsets the effect of the stuck assembly. The power peaking factors depend on the core power, operating history, temperature, pressure, and flow, and thus are different for each case studied. Both cases assume initial hot-standby conditions at event initiation since this represents the most-pessimistic initial condition. A-15.4-8

a. - - - _ _ _ _ _ _ _ _ . . __

Should the reactor be just critical or operating at power at the time of a steamline break, the reactor will be tripped by the normal O overpower protection when the power level or AT reaths a trip setpoint. Following a trip at power, the RCS 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. Thus, the additional stored energy is removed via the cooldown caused by the steamline break before the no-load conditions of RCS temperature and shutdo;;n margin assumed in the analyses are reached. After the additianal stored energy has been removed, the cooldown and reactivity insertions proceed in the same manner as in the analysis which assunes no-load condition at time zero. In addition, sirce the initial steam generator water inventory is greatest at no-load, the magnitude and duration of the RCS cooldown are less for steamline breaks occurring at power. G. In computing the steam flow during a steamline break, the Moody Curve (Reference 2) for f1/0 - 0 is used. The Moody multiplier is I with a discharge at dry saturated steam conditions. H. Perfect moisture separation in the steam generator is assumed. The assumption leads to conservative results since, in fact, considerable water would be discharged. Water carryover would reduce the magnitude of the temperature decrease in the core. I. The maximum feedwater flow is assumed. Increasing the feedwater flowrate aggravates cooldown accidents like steamline rupture. All main and auxiliary feedwater pumps are assumed to be operating at full capacity when the rupture occurs. J. The effect of heat transferred from thick metal in the pressurizer and reactor vessel upper head is not included in the cases analyzed. Studies previously performed show that the heat transferred from these sources is a net benefit in DNBR and RCS energy when the effect of the extra heat on reactivity and peak power is considered. A-15.4-9

i 15.4.2.1.2.2 Epsults The time sequence of events for postulated steamline rupture accidents with and without offsite power are presented in Table 15.4-5. The results presented are a conservative indication of the events that would occur assuming a steamline rupture since it is postulated that all of the conditions described in the prior section occur simultaneously, figures 15.4-28A through 15.4-29C show the mattor coolant system transients and core heat fiux following a main steam pipe rupture. Offsite power is The assumed to De available such that full reactor coolant flow exists. transient shown assumes an uncontrolled steam release from only one steam generator. As can be seen, the core attains criticality with RCCAs inserted (with the design shutdown margin assuming one stuck RCCA) before boric acid solution at 2300 ppm enters the RCS from the ECCS which is drawing from the RWST. The delay time consists of the time to receive and actuate the safety injection signal, to start the high-head safety injection (HHSI) pumps, and to completely align valve trains in the ECCS lines. The HHSI pumps are then ready to deliver flow. At this stage, a further delay is incurred before 2300 ppm boron solution can be injected to the RCS due to the low concentration solution being swept from the Si lines. Should a partial loss of offsite power occur such that power is lost to the ESF functions, an additional Si delay of 15 seconds would occur while the diesel generators start up and reenergize the EST buses. Allowing for these delays, a peak core power well below the nominal full-power value is attained. Should the core be critical at near zero power when the rupture occurs, the initiation of the 51 sigt.al by high steamline differential pressure, low stesmline pressure, or high containment pressure will trip the reactor. Steam release from more than one steam generator will be prevented by automatic closure of the isolatioa valves in the steamlines by low steamline pressure, a high steam flow signal in coincidence with low-low RCS h temperature, or high-high containment pressure. The MSIVs are designed to be A-15.4-10

fully closed in less than 5 seconds after receipt of a closure signal, while

                         ,( 7     the MSIBVs are designed to be fully closed in less than 10 seconds after receipt of a closure signal. Complete steamline isolation occurs when both the'MSIVs and MSIBVs are fully closed. This analysis conservatively assumed that both the MSIVs and MSIBVs close within 10 seconds.

Figures 15.4-30A through 5.4-31C show the responses of the salient parameters for the case discussed above with a total loss of offsite power at the time of the rupture. This assumption results in a coastdown of the reactor coolant pumps. In this case, the core power increases at a slower rate and reaches a lower peak value than in the cases in which offsite power is available to the reactor coolant pumps. The ability of the emptying steam generator to extract heat from the RCS is reduced by the decreased flow in the RCS. It should be noted that following a steamline break, only one steam generator blows down completely. Thus, the remaining steam generators are still available for dissipation of decay heat after the initial transient is-over. ( In case of a loss of offsite power, this heat is removed to the atmosphere via the steamline safety valves. Following blowdown of the faulted steam generator, the plant can be brought to a stabilized hot-standby condition through control of the auxiliary feedwater flow and safety injection flow as described by plant operating procedures. The operating procedures would call for operator action to limit RCS pressure and pressurizer level by terminating safety injection flow and to control steam ' generator level and RCS coolant temperature using the auxiliary feedwater system. Any action required of the operator to maintain the plant in a stabilized condition will be in a time frame in excess of 10 minutes following safety injection. 15.4.2.1.3 Conclusion-A DNB analysis was performed for the above cases. It was found that the DNB design basis is met. O A-15.4-11 .

                             ,;p '  ;
                                      ' [* '***~ [  *g       -i(kMW                              f l

15.4.2.2 MAJOR RUPTURE OF A MAIN FEEDWATER PIPE The major feedline rupture accident is described in Section 15.4.2.2 of the

FSAR. This event is defined as a break in a feedwater line large enough to prevent the addition of sufficient feedwater to the steam generators to maintain shell-side fluid inventory in the steam generators. A feedwater line rupture reduces the ability to remove the . aerated by the core from the reactor coolant system (RCS). The I, ition IV acceptance criteria are conservatively met by demonstratin3 . hat the auxiliary feedwater system heat removal capability is suf ficient to prevent substantial RCS overpressurizatica and maintain sufficient liquid in the RCS to keep the ,

reactor core covered. Typically, this is demonstrated by verifying the maximum hot leg temperature is below that which would cause boiling to occur. The feedline rupture event was not reanalyzed for the VANTAGE 5 fuel effort since none of the VANTAGE 5 design features discussed in Safety Assessment Section 5 affect the feedline break analysis. The primary effects of the VANTAGE 5 fuel on the non-LOCA analysis are due to increased rod drop time due to the smaller diameter thimble tubes and increased heat flux resuiting from smaller diameter fuel rods. Increased rod drop time is primarily a concern in the fast transients where slightly longer scram times can make a large difference in the limiting core conditions. The feedline break transient is a slow transient that tests the ability of the auxiliary I feedwater systs. 13 *~mue core decay heat and is not a H et-f Bf .ncreases in a;rdeel rod drop times. ll Some of the modified safety analysis assumptions described in Safety 2

          -messment Section 5, could potentially affect the feedline break analysis.

As discussed in Section 5, the following analysis assumptions must be - evaluated for any non-LOCA event not explicitly analyzed for the VANTAGE 5 fuel:

1. removal of thimble plugs,

2. modified overtemperature and overpower AT trip setpoints, O

A-15.4-12

O beginning-of-life moderator temperature coefficient of +7 pcm/*F 3. up to 70% rated thermal power (RTP), ramping to zero at 100% RTP,

4. increased power distribution peaking factors (FAH and Fg ),

S. revised thermal design procedure (RTDP), and

6. revised non-RTDP uncertainties (temperature and pressure).

An evaluation of the effect of these safety analysis assumptions on the feedline break event was performed. The conclusions of the evaluat;on are that most of the safety analysis assumptions (assumptions 1 through 5) do not affect the feedline break event, in the case of the revised non-RTDP uncertainties (assumption 6), the licensing-basis analysis assum u uncertainties greater than the revised uncertainties. O A-15.4-13

(( 15.4.4 SINGLE REACTOR COOLANT PUMP LOCKED ROTOR 15.4.4.1 Identification of Causes and Accident Descriotion The accident postulated is an instantaneous seizure of a reactor coolant pump rotor (such as discussed in FSAR Subsection 5.5.1) or the sudden break of the shaf t of the reactor coolant pump. Flow through the affected reactor coolant loop is rapidly reduced, leading to initiation of a reactor trip on a low Reactor Coolant System (RCS) flow signal. Following initiation of the reactor trip, heat stored in the fuel rods continues to be transferred to the coolant causing the coolant to expand. At the same time, heat transfer to the shell side of the steam generators is reduced, first because the reduced flow results in a decreased tube-side film coefficient and then because the reactor coolant in the tubes cools down while the shell-side temperature increases (turbine steam flow is reduced to zero upon plant trip due to turbine trip on reactor trip). The rapid (S expansion of the coolant in the reactor core, combined with reduced heat transfer in the steam generators, causes an insurge into the pressurizer and a pressure increase throughout the RCS. The insurge into the pressurizer compresses the steam volume, actuates the automatic spray system, opens the pows-operated relief valves, and opens the pressurizer safety valves, in that sequence. The two power-operated relief valves are designed for reliable operation and would be expected to function properly during the accident. However, for conservatism, their pressure-reducing effect as well as the pressure-reducing effect of the spray is not included in the analysis. The consequences of a locked rotor (i.e., an instantaneous seizure of a pump shaft) are very similar to those of a pump shaft break. The initial rate of the reduction in coolant flow is slightly greater for the locked rotor event. However, with a broken shaft, the impeller could conceivably be free to spin in the reverse direction. The effect of reverse spinning is to decrease the steady-state core flow when compared to the locked retor scenario. Only one analysis hn been performed, and it represents the most-limiting condition for the locked rotor and pump shaft break accidents. O) A-15.4-14

f] 15.4.4.2 Analysis of Effects and Conseouences

 \J 15.4.4.2.1    Method of Analysis Two digital computer codes are used to analyze this transient. The LOFTRAN code (Reference 1) is used to calculate the resulting loop and core flow transients following the pump seizure, the time of reactor trip based on the loop flow transients, the nuclear power following reactor trip, and the peak RCS pressure. The thermal behavior of the fuel located at the core hot spot is investigated using the FACTRAN code (Reference 3) which uses the core f' +

and the nuclear power values calculated by LOFTRAN. The FACTRAN code includes a film boiling heat transfer coefficient. One case is analyzed: I One locked rotor / shaft break with three loops in operation. The accident is evaluated without offsite power available. Power is assumed to be. lost to the unaffected pumps two seconds after rod motion following a reactor trip on low RCS flow. 15.4.4.2.2 Initial Operating Conditions At the beginning of the postulated locked rotor accident, the plant is assumed.to be operating under the most-adverse steady-state operating conditions. These include the maximum steady-state power level, pressure, and coolant average temperature. The reactivity coefficients assumed in the analysis (see Table 1:i.1-2A) include a conservative moderator temperature coefficient of +7 pcm/*F and a conservatively large (absolute value) of the Doppler-only pcwer coefficient. The total integrated Doppler reactivity from 0 to 100% power is assumed to be -0.016 Ak. For this analysis, the curve of trip reactivity versus time (Figure 15.1-4) was used with a 4.8% Ak trip reactivity which includes the most-reactive RCCA stuck out of the core. O V A-15.4-15 e , , . , ~ _w < --

                                                                                     ,- . , , , - - - .ge.,m-

For the peak pressure evaluation, the initial pressure is conservatively r,,\ Q estimated as 50 psi above the nominal pressure of 2250 psia to allow for errors in the pressurizer pressure measurement and control channels. This is done to obtain the highest possible rise in the coolant pressure during the transient. To obtain the maximum pressure in the primary side, conservatively high loop pressure drops are added to the calculated pressurizer pressure. The pressure response shown in Figure 15.4-33 is at the point in the Reactor Coolant System having the maximum pressure (i.e., the outlet of the faulted loop's RCP). For a conservative analysis of fuel rod behavior, the hot spct evaluation assumes that DNB occurs at the initiation of the transient and continues throughout the event. This assumption reduces heat transfer to the coolant and results in conservatively high hot spot temperatures. The reactor coolant flow coastdown 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, g) and the as-built pump characteristics and is based on high estimates of system oressure losses. 15.4.4.2.3 Evaluation of the Pressure Transient After pump seizure, the neutron flux is rapidly reduced by control rod insertion. Rod motion is assumed to begin one second after the flow in the affected loop reaches 85 percent of nominal flow. No credit is taken for the pressure-reducing effect of the pressurizer relief valves, pressurizer spray, steam dump or controlled feedwater flow after plant trip. Although these systems are expected to function and would result in a lower peak pressure, an additional degree of conservatism is provided by ignoring their effect. The pressurizer safety valves are actuated at 2525 psia. This includes 1% uncertainty over the nominal setpoint of 2500 psia. Additionally, the flow through the pressurizer safety valves was modeled with 3% accumulation, i.e., the flow ramps from zero to full rated flow (steam relief of 34.80 ft3 /s) over the range of 2525 to 2601 psia. A-15.4-16

15.4.4.2.4 Evaluation of DNB in the Core During the Accident O For this accident, DNB is assumed to occur in the core and therefore, an evaluation of the consequences with respect to fuel rod thermal transients is performed. Results obtained from analysis of this " hot spot" condition represent the upper limit with respect to clad temperature and zirconium water reaction. In the evaluation, the rod power at the hot spot is assumed to be 2.5 at the initial core power level. 15.4.4.2.5 Film Boiling Coefficient The film boiling coefficient is calculated in the FACTRAN code using the Bishop-Sandberg-Tong film boiling correlation (Reference 3). The fluid properties are evaluated at film temperature (average between wall and bulk temperatures). The program calculates the film coefficient at every time step based upon the actual heat transfer conditions at the time. The neutron flux, system pressure, bulk density, and mass flow rate as a function of time are used as program input. O For this analysis, the initial values of the pressure and the bulk density are used throughout the transient since they are the most conservative with respect to the clad temperature response, As indicated earlier, DNB was assumed to start at the beginning of the accident. 15.4.4.2.6 Fuel Clad Gap Coefficient The magnitude and time dependence of the heat transfer coefficient between fuel and clad (gap coefficient) has a pronounced influence on the thermal results. The larger the value of the gap coefficient, the more heat is transferred between pellet and clad. For the initial portion of the transient, a high gap coefficient produces higher clad temperatures since the heat stored and generated in the fuel redistributes itself in the cooler cladding. This effect is reversed when the clad temperature exceeds the pellet temperature in cases where a zirconium-steam reaction is present. Based on investigations on the effect of the gap coefficient upon the maximum clad temperature during the transient, the gap coefficient was assumed to A-15.4-17

increase from a steady-state value consistent with initial fuel temperatures to 10,000 Btu /hr-ft 2_.F at the initiation of the transient. Thus, the large amount of energy stored in the fuel is released to the clad at the initiation of the transient. 15.4.4.2.7 Zirconium-Steam Reaction The zirconium-steam reaction can become significant above 1800*F (clad temperature). The Baker-Just parabolic rate equation (Reference 3) shown below is used to define the rate of the zirconium-steam reaction. dN 2) - 33.3 x 106 exp -[ fT l where: w - amount Zr reacted, mg/cm 2 _) t - time, sec T - temperature, 'K The reaction heat is 1510 cal /gm. The effect of zirconium-steam reaction is included in the calculation of the

         " hot spot" clad temperature transient.

15.4.4.2.8 Results The calculated sequence of events is shown in Table 15.4-5. The transient results without offsite power available are shown in Figures 15.4-33 through 15.4-38. The peak Reactor Coolant System pressure reached during the transient is less than that which would cause stresses to exceed the faulted condition stress limits. Also, the peak clad surface temperature is considerably less em than 2700*F. It should be noted that the clad temperature was ( '

A-15.4-18

fN conservatively calculated assuming that DNB occurs at the initiation of the

 '9 -transient. The results of these calculations (peak pressure, peak clad temperature, and zirconium-steam reaction) are also summarized in
    - Table IG.4-25.

15.4.4.3 Conclusions The analysis has shown the following. A. Since the peak Reactor Coolant System pressure reached during the transient is less than that which would cause stresses to exceed the faulted condition stress limits, the integrity of the primary coolant system is not endangered. B. Since the peak clad surface temperature calculated for the hot spot during the worst transient remains considerably less than 2700*F and the amount of zirconium-water reaction is small, the core will remain [] in place and intact with no loss of core cooling capability. As 15.4.5 FUEL HANDLING ACCIDENT (FHA) Evaluation or analysis of this event is considered in Appendix C of the Safety Assessment. t A-15.4-19 i

15.4.6 RUPTURE OF A CONTROL ROD' DRIVE MECHANISM (CRDM) HOUSING (RCCA

N EJECTION) 15.4.6,1 Identification of Causes and Accident Description This accident-is defined as a mechanical failure of. a control rod drive mechanism pressure housing resulting in the ejection of an RCCA and drive shaft. The consequence of th-is mechanical failure is a rapid positive reactivity insertion together with an adverse core power distribution, )

possibly leading to localized fuel rod damaoe. j l 15.4.6.1.1- Design Precautions and Prc,tection ) Certain features in Westinghouse pressurized water reactors (PWRs) are intended to preclude the possibility of a rod ejection accident or to limit the-consequences if the accident was to occur. These include a sound, conservative mechanical design of the rod housings, together with a thorough ! quality control (testing) program during assembly, and a nuclear design which lessens the potential RCCA ejection worth and minimizes the number of assemblies asserted at high power levels. 15.4.6.1.1.1 Mechanical Desian The mechanical design is discussed in FSAR Section 4.2, Mechanical design and quality control procedures intended to preclude the possibility of a rod cluster control assembly to be rapidly ejected from the core are listed below. A. Each full-length control rod drive mechanism housing is completely assembled and shop' tested at 4100 psi. B. The mechanism housings are individually hydrotested as they are attached to the head adapters in the reactor vessel head and checked during the hydrotest of the completed Reactor Coolant System. O A-15.4-20

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

[ C. Stress levels in the mechanism are not affected by anticipated system transients at power or by the thermal movement of the coolant loops. Moments induced by the_ design earthquake can be accepted within the allo.,able primary working stress ranges specified in the ASME Code, Section 111, for Class I components. D. The latch mechanism housing and rod travel housing are each a single . langth of forged type-304 stainless steel. This material exhibits excellent notch toughness at all temperatures which will be encountered. A significant margin of strength in the elastic range, together with the large energy absorption capability in the plastic range, gives additional assurance that the gross failure of the housing will not occur. The joints between the latch mechanism housing and rod travel housing are threaded joints and reinforced by canopy-type rod welds. Administrative procedures ~ O TQ require periodic inspections of these welds, f l 15.4.6.1.1.2 Nuclear Desion Even if a rupture of a rod cluster control assembly drive mechanism housing is postulated, the operation of a plant utilizing chemical shim is such that the-severity of an ejected rod cluster centrol assembly is limited. In general, the reactor is operated with the rod cluster control assemblies inserted only far enough to permit load follow. Reactivity changes caused by the core depletion and xenon transients are comnensated by boron changes. Further, the location and' grouping of control rod banks are selected during

  - the nuclear design to lessen the severity of a rod cluster control assembly ejection accident. Therefore, sh.;uld a rod cluster control assembly be ejected from its normal position during full-power operation, only a minor reactivity excursion, at worst, could be expected to occur.

However, it may be occasionally desirable to operate with larger-than-normal insertions. For this reason, a rod insertion limit is defined as a function A-15.4-21 w p

of power level. Operation with the rod cluster control assemblies above the insertion limit guarantees adequate shutdown capability and acceptable power distribution. The position of all rod cluster control assemblies is continuously indicated in the control room. An alarm will occur if a bank of rod cluster control assemblies approaches its insertion limit or if one control rod assembly deviates from its bank. There are low and low-low level insertion alarm circuits for each bank. In addition, bank positions and the low-low limit are provided by the Rod Insertion Limit recorder. Operating instructions require boration at the low level alarm and emergency boration at the low-low level alarm. 15.4.6.1.1.3 Reactor Protection The Reactor Protection System response to a rod ejection accident is described in Reference 4. The protection for this accident is provided by the power range high neutron flux trip (high and low setting) and high rate of neutron flux increase trip. These protection functions are described in FSAR Section 7.2. 15.4.6.1.1.4 Effects on Adiacent Housing Disregarding the remote possibility of the occurrence of a rod cluster control assembly housing failure, investigations have shown that the failure of a housing due to longitudinal or circumferential tracking will not cause damage to adjacent housings leading to increased severity of the initial accident. 15.4.6.1.2 Limiting Criteria Due to the extremely low probability of a rod cluster control assembly ejection accident, some fuel damage could be considered an acceptable consequence. Comprehensive studies of the threshold of fuel failure and of the threshold of significant conversion of the fuel thermal energy to mechanical energy A-15.4-22

O have been carried out as part of the SPERT project by the Idaho Nuclear Corporation (Reference 5). Extensive tests of UO2 zirconium clad fuel rods l representative of-those present in pressurized water reactor-type cores have _; demonstrated failure thresholds in the range of 240 to 257 cal /gm. However, other rods of a slightly different design design exhibited failure as low as 225 cal /gm. These results differ significantly from the TREAT (Reference 6) results which indicated a failure threshold of 280 cal /gm. Limited results have indicated that this threshold decreases 10 percent with fuel burnup. .j The clad failure mechanism appears to be melting for unirradiated (zero burnup) rods and brittle fracture for irradiated rods. The conversion ratio of thermal to mechanical energy is also important. This ratic becomes marginally detectaole above 300 cal /gm for unirradiated rods and 200 cal /gm ! for irradiated rods; catastrophic failure (large fuel dispersal, large l

      - pressure rise), even for irradiated rods, did not occur below 300 cal /gm.

l The real physical limits of this accident are that the rod ejection event and any consequential damage to either the core or the Reactor Coolant System must not prevent long-term core cooling and any offsite dose consequences

must be within the guidelines of 10 CFR 100. More-specific and restrictive criteria are applied to ensure fuel dispersal in the coolant, gross lattice distortion, or severe shock waves will not occur. In view of the above experimental results, the conclusions of WCAP-7588, Rev. 1-A (Reference 7) and Reference 10, the limiting criteria are: A. Average fuel pellet enthalpy at the hot spot must be maintained below 225 cal /gm for unirradiated and 200 cal /gm for irradiated fuel, 4 B. Peak reactor coolant pressure must be less than that which could cause RCS stresses to exceed the faulted-condition stress limits, C. Fuel melting is limited to less than 10 percent of the fuel volume at the hot spot even if the average fuel pellet enthalpy is below the limits of Criterion A. B A-15.4-23

l [U- ' 15.4.6.2 Analysis of Effects and Consecuences 15.4.6.2.1 Method of Analysis The ca'lculation of the RCCA ejection transient is performed in two stages:

                                                    .first, an average core channel calculation and then a hot region calculation. The average core calculation uses spatial neutron-kinetics methods to determine the average power generation as a function of +ime
                                                   . including the various total core feedback effects, i.e., Doppler       .;tivity and moderator reactivity. Enthalpy and temperature transients at the hot spot are then determined by multiplying the average core energy generation by
                                                    'the hot channel factor and. performing a fuel rod transient heat transfer c:lculation. The power distribution calculated without feedback is conservctively assumed to persist throughout the transient. Reference 8 provides a detailed discussion of the method of analysis.

15.4.6.2.1.1 Averaae Core Analysis [J3 The average core-transient analysis uses the spatial-kinetics computer code, TWINKLE-(Reference 9). This code solves the two-group neutron diffusion theory kinetic. equation in one, two or three spatial dimensions (rectangular coordinates) for six delayed neutron groups and up to 2000 spatial points. The computer code includes a detailed multiregion, transient fuel-clad-coolant heat transfer model for calculation of pointwise Doppler and moderator feedback effects. This analysis eses the code as a one-dimensional axial kinetics code since it allows a more-realistic representation of the spatial effects of axial moderator feedback and RCCA movement. However, since the radial dimension is missing, it is still necessary to employ very conservative methods (described below) of calculating the ejected rod worth and hot channel factor. Further description of TWINKLE appears in Subsection 15.1.9 and Reference 9. A-15.4-24

i l 15.4.6.2.1.2 Hot Soot Analysis

  /7 t t

U The average core energy addition, calculated as described above, is , multiplied by the appropriate hot channel factors. The hot spot analysis l uses the detailed fuel and clad transient heat transfer computer code, FACTRAN (Reference 3). This computer code calculates the transient i temperature distribution in a cross section of a metal clad UO2 fuel rod, i and the heat flux at the surface of the rod, using as input the nuclear power i versus time and local coolant conditions. The zirconium-water reaction is explicitly represented, and all material properties are represented as functions of temperature. A parabolic radial power distribution is assumed within the fuel rod. FACTRAN uses the Dittus-Boelter or Jens-Lottes correlation (Reference 3) to determine the film heat-transfer before DNB, and the Bishop-Sandberg-Tong enrrelation (Reference 3) to determine the film boiling coefficient after DNB. The DNB heat flux is not calculated; instead, the code is forced into DNB by specifying a conservative DNB heat flux. The code can calculate the [ gap heat transfer coefficient; however, it is adjusted to force the full-power, steady-state temperature distribution to agree with fuel heat transfer design codes. Further description of FACTRAN can be found in Reference 3 and Subsection 15.1.9. 15.4.6.2.1.3 System Overoressure Analysis Because the safety limits for fuel damage specified earlier are not exceeded during the transient, there is little likelihood of fuel dispersal into the coolant. Therefore, the basis of the pressure surge calculation may be conventional heat transfer from the fuel and prompt heat generation in the cool ant. The pressure surge is calculated by first performing the fuel heat transfer calculation to determine the average and hot spot heat flux versus time. A THINC calculation uses this heat flux data to oetermine the volume surge. Finally, the volume surge is simulated in a plant transient computer code. .-( This code calculates the pressure transient taking into account fluid A-15.4-25

transport in'the RCS'and heat transfer to the steam generators. No credit is taken for the possible pressure reduction caused by the assumed failure of ' the control rod pressure housing. 15.4.6.2.2 Calculaticn of Basic Parameters

  -Input parameters for the analysis are conservatively selected on the basis of
  -values calculated for this type of core.       The discussion of the more-important parameters is presented below.       Table 15.4-12 presents the parameters used in this analysis.

15.4.6.2.2.1 Eiected Rod Worths and Hot Channel Factors The values for ejected rod worths and hot channel factors are calculated using either three-dimensional static methods or a synthesis of one-dimensional and two-dimensional calculations. Standard nuclear design codes are used in the analysis. No credit is taken for the flux flattening effects of reactivity feedback. The calculation is performed for the maximum p 'd allowed bank insertion at a given power level, as determined by the rod insertion limits. The analysis assumes adverse xenon distributions to provide worst-case results. The total transient hot channel factor, F g, is then obtained by combining the axial and radial factors, even though the axial peaks are not coincident under the conditions of the calculation. The ejected rod worth and hot channel factors include appropriate margins to account for any calculational uncertainties, including an allowance for nuclear power peaking due to fuel densification. 15.4.6.2.2.2 Reactivity Feedback Weiahtino Factors

  'The largest temperature rises, and hence the largest reactivity feedbacks, occur in channels where the power is higher than average. Since the weight of a region is dependent on flux, these regions have high weights. This means that the reactivity feedback is larger than that indicated by a simple single-channel analysis. Physics calculations have been performed for temperature changes with a flat temperature distribution and with o large A-15.4-26
                                                                                 .- ~

number of axial and radial temperature distributions. The analysis compares

       ~j  reactivity changes and determines effective weighting factors. These (V  weighting factors take the form of multipliers which, when applied to single-channel feedbacks, correct them to effective whole-core feedbacks for the appropriate flux shape. Axial weighting is not necessary because this analysis employs a one-dimensional (axial) spatial kinetics method. In addition, no weighting is applied to the moderator feedback. The analysis applies a conservative radial weighting factor to the transient fuel temperature to obtain an effective fuel temperature as a function of time accounting for the missing spatial dimension. These weighting factors have also been shown to be conservative compared to three-dimensional analysis (Reference 7).

15.4.6.2.2.3 Moderator and Donoler Coefficient The critical-boron concentrations at the beginning of life and end of life are adjusted in the nuclear code in order to obtain moderator density coefficient curves which are conservative when compared to the actual design conditions for the plant. As discussed above, these results do not have any weighting factor applied to them. The resulting moderator temperature coeffic'ient is at least +7 pcm/*F at the appropriate zero- or full-power nominal average temperature for the beginning-of-life cases.

          -The Doppler reactivity defect is determined as a function of power level using a one-dimensional steady-state computer code with a Doppler weighting factor of 1.0. The Doppler weighting factor will increase under accident conditions, as discussed above.

15.4.6.2.2.4 Delayed Neutron Fraction. B Calculations of-the effedive delayed neutron fraction (Beff) typically yield values no less than 0.70 percent at BOL and 0.50 percent at E0L. The ejected rod accident is sensitive to B if the ejected rod worth is equal to or greater than Beff, as in the zero-power transients. In order to allow for future fuel cycle flexibility, conservative estimates of B of 0.54 percent at BOL and 0.44 percent at EOL are used in the analysis. A-15.4-27

                                                     .-.._s,.

15.4.6.2.k.5 Trio Reactivity insertion The trip reactivity insertion is assumed to be 4.8% Ap from hot full power and 1.77% Ap from hot zero power, including the effect of one stuck RCCA. These values are also reduced by the ejected rod. The shutdown reactivity is simulated by dropping a rod of the required worth into the core. The start of rod uotion occurred 0.5 second after reaching the power range high neutron flux trip setpoint. It is assumed that insertion to dashpot does not occur until 2.7 seconds after the rods begin to fall. This time to full insertion, together with the 0.5-second trip delay, overestimates the time for significant insertion of shutdown reactivity into the core. The choice of such a conservative insertion rate means that there is over 1 second after reaching the trip point before significant shutdown reactivity is inserted into the core. This is a significant conservatism for hot-full-power accidents. The minimum design shutdown marq:n available for this plant at hot zero power (HZP) may only occur at end of life in the equilibrium cycle. This value includes an allowance for the worst stuck rod, an adverse xenon distribution, conservative Doppler and moderator defects, and an allowance for calculational uncertainties. Physics calculations have shown that two stuck ROCAs (one of which is the worst ejected rod) reduce the shutdown margin by about an additional 1% Ap. Therefore, following a reactor trip resulting from an RCCA ejection accident, the reactor will be subtritical when the core returns to HZP. . 15.4.6.2.3 Results The calculated sequence of events is shown in Table 15.4-5. The values of the parameters use.' in the analysis, as well as the results of the analysis, are presented in Table 15.4-12 and are discussed in the following subsections. O l A-15.4-28

j 15.4.6.2.3.1 Becinnina of Life. Full Power This case assumed Control Bank D inserted to its insertion limit. The worst ejected rod worth and hot channel factor were conservatively calculated to be 0.20% Ap and 6.0, respectively. The hot spot average fuel pellet enthalpy was 175.0 cal /g. The peak hot spot fuel center temperature reached the BOL melt temperature of 4900*F. However, fuel melting was well below the limiting criterion of 10 percent of the pellet volume at the hot spot. 15.4.6.2.3.2 Beainnina of Life. Zero Power For this condition, Control Bank D was assumed to be fully ir.serted and Bank C was assumed to be at its insertion limit. The worst ejected rod is in Control Bank D and has a worth of 0.75% Ap and a hot channel factor of 13.0. The hot spot average fuel pellet enthalpy was 139.4 cal /g. The peak hot spot fuel center temperature was 3935'F. 15.4.6.2.3.3 End of Life. Fell Power The analysis assumed Control Bank D at its insertion limit. The ejected rod worth and hot channel factor were conservatively calculated to be 0.21% Ap and 7.0, respectively. This resulted in a hot spot average fuel pellet enthalpy of 165.2 cal /g. The peak hot spot fuel center temperature reached melting, conservatively assumed at 4800*F; however, fuel melting was well below the limiting criterion of 10 percent of the pellet volume at the hot spot. 15.4.6.2.3.4 End of Life. Zero Power The analysis of this case assumes the ejected rod worth and hot channel t : tor by assuming Control Bank D to be fully inserted and Bank C at its insertion limit. The worst ejected rod is in Control Bank D and has a worth of 0.85% Ap and a hot channel factor of 18.0. The hot spot average fuel pellet enthalpy was 129.4 cal /g. The peak fuel center temperature was 3578'F. O A-15.4-29

L x _

 ,j r^            JA summary of the cases presented in Subsections 15.4.6.2.3.1 through
'bI                  15;4:6.2.3.4',.above, is given in Table.15.4-12.             The nuclear power and hot _                          '

spot" temperatures for two. cases (BOL, full power and E0L, zero power) are

               - [ presented 1inFigures'15.4440 through 15.4-43.
                                                               ~

15.4 6.2.3.5 -Fission Product Release

                  . The analysis conservatively ass.:me; for all rods entering DNB that fission products are released:from thr gap. 'In all cases considered,.less than

_ ;10 percent'of:the rods entered DNB based on a detailed three-dimensional ~

THINC analysis (Reference 7). Although' the analysis predicts _ limited fuel melting at theihot spot-for.the full-power cases, in practice, melting is not 1ikely since the analysis conservatively assumed that the hot spots before "and after ejection were-coincident. j

                   ~15.4.6;2.3.6- Pressure Surce.                                                                                    ,
                   - A' detailed calculation of the pressure surge for an ejection rod worth of
           '         1) dollar at beginning of-life, hot full power, indicates that' thel peak

pressure _does.not exceed that which would cause reactor pressure vessel ;

stress to' exceed the faulted-condition-stress limits. Since the severity of.

                    .thepresentcanalysis'does not exceed the worst-case analysis, the accident forsthis plant will not result in an excessive pressure rise or further
                    -adv'erse effects. to-- the-- RCS.

_ 15.4.6.2.3.7' Lattice oeformations A large. temperature gradient exists in the region of the_ hot spot. Since the ifuel' rods are free to move in the vertical direction, differential expansion between' separate rods:cannot produce distortion. However, the temperature gradients across . individual rods 'may produce a- differential expansion tending

                   .to bow the midpoint'of the rod toward the' hotter side of.the rod.                         Physics 1 .-

calculations indicate that this bowing results in a negative reactivity insertion. In practice, significant bowing is not expected since the

                   ~ structural rigidity of the core is more than sufficient to withstand the A-15.4-30 l
                    - -              .-.-- - . . . -                       .-     .._.- - -.             . -   - . . . .           =

 -p forces produced.          Boiling in the hot spot region produces a net flow away from that_ region; however, the fuel releases heat to the water slowly, and it is considered inconceivable that cross flow is sufficient to produce                        I significant lattice forces. Even if massive and rapid boiling tufficient to distort the lattice is hypothetically postulated, the large void fraction in the hot spot region produces a reduction-in the. total core moderator-to-fuel               ,

ratin. Therefore, the net effect is negative feedback which leads to the conclusion that no conceivable mechanism exists for a net positive feedback ! resulting from lattice deformation. In fact, a small negative feedaack may result. The effect is conservatively ignored in the analysis. 15.4.6.3 Conclusions Despite the conservative assumptions, the analyses indicate that the described fuel limits are not exceeded. It is concluded that there is no danger of sudden fuel dispersal into the coolant. Since the peak pressure A does not exceed that which would cause stresses to exceed the faulted-condition stress limits, it is also concluded that there is no danger of further consequential damage to the RCS. The analyses have demonstrated that the fuel rods entering DNB are iess than 10 percent of the fuel rods in the core; therefore, the assumption of 10 parcent of the fuel rods in the core entering DNB for the fission product release calculation is conservative. O A-15.4-31

O 15.

4.7 REFERENCES

1. T. W. T. Burnett, et al . , "LOFTRAN Code Description, "WCAP-7907-P- A (Proprietary) and WCAP-7907-A (Nonoroprietary), April 1984.

2. Moody. F. S., " Transactions of the ASME," Journal of Heat Transfer, p 134, February 1965.

3. Hargrove, H. G. , "FACTRAN -- A FORTRAN-IV Code for Thermal Transients in Fuel Rod," WCAP-7908-A, December 1989.

a U02

4. Burnett, T. W. T., " Reactor Protection System Diversity in Westinghouse Pressurized Water Reactors," WCAP-7306, April 1969.

5. Taxebius, T. G. , ed. , " Annual Report - Spert Project, October 1968 -

September 1969," IN-1370, Idaho Nuclear Corporation, June 1970, t

6. Liimatainen, R. C and Testa, F. J., " Studies in TREAT of Zircaloy 2-Clad, U0 -Core 2 Simulated Fuel Elements," ANL-7225, p 177, November 1966.

7. Risher, D. H., Jr., "An Evaluation of the Rod Ejection Accident in Westinghouse Pressurized-Water Reactors Using Spatial Kinetics Methods,"

WCAP-7588. Rev. 1-A, December 1971.

8. Nuclear Regulatory Commission, Directorate of Regulatory Standards,

            " Assumptions Used for Evaluating a Control Rod Ejection Accident for Pressurized-Water Reactors," Reaulatory Guide 1.77, May 1974.

9. Risher, D. H., Jr. and Barry, R. F., " TWINKLE -- A Multi-Dimensional Neutron Kinetics Computer Code," WCAP-7979-P-A (Proprietary) and WCAP-8028-A (Nonoroprietaryl, January 1975.

A-15.4-32

i I 4 l

  /' N L.)                                                                                 i
       -10. Letter 'from W. J. Johnson of Westinghouse Electric Corporation to Mr. R. C. Jones of the Nuclear Regulatory Commission, letter Number       i NS-NRC-89-3466, "Use of 2700'F PCT Acceptance Limit in Non-LOCA        'I Accidents," - October 23, 1989.                                          l
 =                                                                                   1
                                                                                   .1 l

l l I l 1 O - J

l O A-15.4-33

Table 15.4-5 (Sheet 1 of 3) Time Sequence of Events for Condition IV Events Accident Event TIME (s) Major Steam Pipe Rupture . A. With Offsite Power Steamline ruptures 0.0 Pressurizer empties 22.4 Criticality attained 40.0 Borated water from the RWST 57.4 reaches the core B. Without Offsite Power Steamline ruptures 0.0 Pressurizer empties 35.8 Criticality attained 54.6 Borated water from the RWST 77.6 reaches the core Reactor Coolant Pump Shaft Seizure (Locked Rotor / Broken Shaft) Rotor on one pump locks or 0.0 the shaft breaks low flow reactor tr'.p setpoint 0.03 reached Rods begin to drop 1.03 Remaining pumps lose power 3.03 and begin coasting down Maximum RCS pressure occurs 3.1 Maximum clad temperature occurs 3.8 O A-15.4-34

     ~ . .

Table 15.4-5 (Sheet 2 of 3) Time Sequence of Events for Condition IV Events Accident Event TIME (s) RCCA Ejection Accident A. Beginning of Life, Zero Power Initiation of rod eje: ion .0.0 Power range high neutron flux 0.30 low setpoint reached Peak nuclear power occurs 0.36 Rods begin to fall into core 0.80 Peak clad average temperature 2.54 occurs Peak heat flux occurs 2.55 Peak fuel average temperature 2.80 occurs B. Beginning of Life, Full Power Initiation of rod ejection 0.0 Power range high neutron flux high setpoint reached 0.05 Peak nuclear power occurs 0.13 Rods begin to fall into core 0.55 Peak fuel average temperature 2.46 occurs Peak clad average temperature 2.56 occurs Peak heat flux occurs 2.57 4 O A-15.4-35

( ) Table 15,4-5 (Sheet 3 of 3) Time Sequence of Events for Condition IV Events Accident Event TIME (s) RCCA Ejection Accident (con't) C. End of Life, Zero Power initiation of rod ejection 0.0 Power range high neutron flux 0.19 low setpoint reached Peak nuclear power occurs 0.23 Rods begin to fall into core 0.69 Peak clad average temperature 1.40 occurs Peak heat flux occurs 1.41 Peak fuel average temperature 1.66 3 occurs Q) D. End of Life, Full Power initiation of rod ejection 0.0 Power range high neutron flux high setpoint reached 0.05 Peak nuclear oower occurs 0.13 Rods begin to fall into core 0.55 Peak fuel average temperature occurs 2.19 Peak clad average temperature occurs 2.30 Peak heat flux occurs 2.31 lQ NY A-15.4-36

Table 15.4-12 Parameters Used in the Analysis of the Rod Cluster Control Assembly Ljection Accident HZP HFP HZP HFP Time in tife Beainnino Beainnina [nd n End 0 102 0 102 Power level (%)

Ejected rod worth (% op) 0.75 0.20 0.85 0.21 Delayed neutron fraction (%) 0.54 0.54 0.44 0.44 Doppler feedback reactivity weighting 2.07 1.30 2.755 1.30 Trip reactivity (% Ap) 1.52 4.8 1.62 4.8 Fg before rod ejection -

2.50 - 2.50 13.0 6.0 18.0 7.0 Fg after rod ejection Number of operational pumps 2 3 2 3 Maximum fuel pellet average 3316 4014 3111 3823 temperature at the hot spot (*F) Maximum fuel center temperature at the 3935 >4900 3578 >4800 hot spot (*F) Maximum fuel stored energy at the hot spot (cal /g) 139.4 175.0 129.4 165.2 Fuel melt (percent) 0 < 10 0 < 10

     * - Power Level is percent of 2775 MWt.

O A-15.4-37

I O Table 15.4-25

SUMMARY

OF RESULTS FOR THE LOCKFD ROTOR TRANSIENT 3 Loops Initially Criteria Ooeratina. One Locked Rotor Maximum RCS Pressure 2705 psia Maximum Clad Temperature 2205'F at Core Hot Spot O Zr-H2O Reaction at 1.2 wt. % Core Hot Spot O I A-15.4-38

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m O Aonendix B Joseph H. Farley Nuclear Plant Units 1 and 2 l Request for Technical Specifications Changes l LOCA Accident Analysis O O

q l APPENDIX B TABLE OF CONTfNTS j 1 (3 4

   \_ /  TABLE OF C0NTENTS.................................................                                                                   . ..      .it i

LIST OF TABLES..................................... ......... .... . ...iii LIST OF FIGURES..................................................... ......iv 15.3 CONDITION !!! - INFREQUENT EVENTS.................................. 8 1 15.3.1 LOSS OF REACTOR COOLANT FROM SMALL RUPTURED PIPES OR FROM CRACKS IN LARGE PIPES WHICH ACTUATES THE EMERGENCY CORE COOLING SYSTEM......................... ..B.) 15.3.1.1 Identification of Causes and Accident Description.....B-1 15.3.1.2 Analysis of Effects and Consequences...... ...........B 4 15.3.1.2.1 Method of Analysis.............................B-4 15.3.1.2.2 Results........................................B-6 A' w/ 15.3.1.2.2.1 Limiting Break Case.....................B 6 15.3.1.2.2.2 Additional Break Cases....... ... ......B 7 15.3.1.3 Conclusions...........................................B 8 R E F E R EN C E S , S EC T I ON 15. 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . B 9 T AB L E S , S E C T I ON 15 . 3 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . B - 10 F I GUR E S , S E CT ION 15. 3 . . . . . . , . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . B - 15 ii 9

i l APPENDIX B LIST OF TABLES TABLE 15.3 1 SAFETY INJECTION FLOW RATE............. ....... . . . . . . . B 10 l TABLE 15.3 2 PLANT INPUT PARAMETERS USED IN SMALL BREAK LOCA ANALYSIS FOR 17 X 17 VANTAGE-5 FUEL. . . . . . . . . . . . . . . . . .B 12 TABLE 15.3-3 SMALL BREAK LOSS OF COOLANT ACCIDENT CALCULATION...........E-13 l TABLE.15.3-4 TIME SEQUENCE OF EVENTS FOR CONDITION !!! EVENTS....... ...B 14 W-. p., y y , .yvm.m, y.%s., .,,o._ , y,e. , , ,, ,..-5---- . . - . . . . . _ , , ,

APPENDIX B LIST OF FIGURES /~N \s_,) FIGURE 15.3 1 CODE INTERFACE DESCRIPTION FOR SMALL BREAK MODEL. .. ...B 15 FIGURE 15.3-2 SMALL BREAK HOT ROD POWER SHAPE............ . . .. ......B-16 FIGURE 15.3-3 SMALL BREAK SAFETY INJECTION FLOWS...... ........ ..... .B-17 FIGURE 15.3-4 RCS DEPRESSURIZATION TRANSIENT UNIT 1 (3-INCH BREAK). ..B 18 FIGURE 15.3 5 CORE MIXTURE UNIT 1 (3-INCH BREAK)...... ... . .. ...B 19

         ' FIGURE 15.3-6    CLAD AVERAGE TEMPERATURE UNIT 1 (3 lNCH BREAK).. .. .. ..B-20 FIGURE 15.3-7     CORE EXIT STEAM FLOW UNIT 1 (3-INCH BREAK).... ......... 8 21
         - FIGURE-15.3-8    HEAT TRANSFER COEFFICIENT-HOT ASSY UNIT 1 (3 INCH BREAK).B-22 FIGURE 15.3-9     FLUID TEMPERATURE-HOT SPOT UNIT 1 (3-INCH BREAK) . . . . . . . . .B 23 FIGURE 1,5.3 10 COLD LEG BREAK MASS FLOW UNIT 1 (3-INCH BREAK)...........B-24

(}) , FIGURE 15.3-11 'ECCS PUMPED SAFETY INJECTION UNIT 1 (3 INCH BREAK).......B-25 FIGURE 15.3-12 RCS DEPRESSURIZATION TRANSIENT UNIT 2 (*J-INCH BREAK). . .B-26 FIGURE 15.3-13 CORE MIXTURE UNIT 2_(3-INCH BREAK)..................... .B-27 FIGURE 15.3-14 CLAD AVERAGE TEMPERATURE UNIT 2 (3-INCH BREAK)...........B 28 r;GURE 15.3-15 CORE EXIT STEAM FLOW UNIT 2 (3-INCH BREAK)........ ......B 29 FIGURE 15.3-16 HEAT TRANSF.ER COEFFICIENT-HOT ASSY UNIT 2 (3-INCH BREAK).B-30 FIGURE 15.3 17 FLUID TEMPERATURE-H0T SPOT UNIT 2 (3-INCH BREAK)... .....B-31 FIGURE 15.3-18 COLD LEG BREAK MASS F.0W UNIT 2 (3-INCH BREAK)...........B-32 FIGURE 15.3-19 ECCS PUMPED SAFETY INJECTION UNIT 2 (3-INCH BREAK).......B 33 iv

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APPENDIX B LIST OF FIGURES (cont) O FIGURE 15.3-20 RCS*DEPRESSURIZATION TRANSIENT UNIT 1 (2-INCH BREAK). .B-34 FIGURE 15.3-21 CORE MIXTURE UNIT 1 (2-INCH BREAK)... . .. . . .B-35 FIGURE 15.3-22 CLAD AVERAGE TEMPERATURE UtilT 1 (2-INCH BREAK) . . . . . .. .B-36 FIGURE 15.3-23 RCS DEPRESSURIZATION TRANSIENT UNIT 1 (4-INCH BREAK)... .B-37 FIGURE 15.3-24 CORE MIXTURE UNIT 1 (4-INCH BREAK)............. .... . .B 3B FIGURE 15.3 25 CLAD AVERAGE TEMPERATURE UNIT 1 (4-INCH BREAK)........ .B-39 FIGURE 15.3-26 RCS DEPRESSURIZATION TRANSIENT UNIT 1 (6 INCH BREAK).....B-40 FIGURE 15.3-27 CORE MIXTURE UNIT 1 (6-!NCH BREAK).......................B-41 FIGURE 15.3-2B CLAD AVERAGE TEMPERATURE UNIT 1 (6 INCH BREAK).. ........B-42 , O V O

APPENDIX B 15.3 CONDITION 111 - INFREQUENT EVENTS By definition, Condition 111 occurrences are faults which may occur very infrequently during the life of the plant. They will be accommodated with the failure of only a small fraction of the fuel rods although sufficient fuel damage might occur to preclude resumption of the operation for a considerable outage time. The release of radioactivity will not be sufficient to interrupt or restrict public use of these areas beyond the exclusion radius. A Condition 111 fault will not, by itself, generate a Condition IV fault or result in a consequential loss of function of the reactor coolant system (RCS) or containment barriers. For the purposes of this report, the following faults have been grouped into this category: A. l.oss of reactor coolant from small ruptured pipes or from cracks in large pipes which actuate the emergency core cooling system (ECCS). B. Minor secondary system pipe break. C. Inadvertent loading of a fuel assembly into an improper . position. D. Complete loss of forced reactor coolant flow. E. Waste gas decay tank rupture. F. Single rod cluster control assembly (RCCA) withdrawal at full power. 15.3.1 LOSS OF REACTOR COOLANT FROM SMALL RUPTURED PIPES OR FROM CRACKS IN LARGE PIPES WHICH ACTUATES THE EMERGENCY CORE COOLING SYSTEM This section presents results of the small break loss-of-coolant accident (LOCA) in conformance with 10 CFR 50.46 (Reference 1) and Appendix K of 10 CFR 50. 15.3.1.1 Identification of Causes and Accident Description A loss-of-coolant accident is defined as rupture of the RCS piping or of any line connected to the system. Ruptures of small cross-sections will cause expulsion of the coolant at a rate which can be accommodated by the high head safety injection pumps and which would maintain an operational B-1

ApPEN0lx B f: water level in the pressurizer permitting the operator to execute an orJerly shutdown. The coolant which would be released to the containment contains the fission products existing in it. , The maximum break size for which the normal makeup system can maintain the pressurizer ~ level can be obtained by comparing the calculated flow from the RCS break.against the high head safety injection makeup flow at normal RCS pressure, i.e., 2250 psia. The makeup flowrate from one high head safety , injection pump is adequate to sustain pressurizer level at 2250 psia for a , break as large as 0,277-in diameter. This break results in a loss of approximately 13.3 lbm/sec. A small break, as considered in this section, is defined as a rupture of the RCS piping with_a cross-sectional area less than 1.0 ft2 , in which the ; normally operating charging system flow is not sufficient to sustain pressurizer level and pressure. For small break LOCAs, the most limiting single active failure is the one , that results in the minimum ECCS flow delivered to the RCS. This has been. determined to be the loss of an emergency power train which results in the

                             - loss of one complete train of ECCS components. This means that credit can be taken for only one hi_gh head safety injection pump, and one RHR (low head)

O 3 pump. During the small break transient, one ECCS train is assumed to start and deliver flow through the injection lines (one for each' loop). For break sizes smaller than the diameter of the injectioa lines (about 5.2-inches), one branch injection line is assumed-to spill to RCS backpressure. If, however, the break is postulated to be larger than the diameter of an injection line, one branch injection line is assumed to spill to cortainment

                            - backpressure.
                             ' Should a small break LOCA occur, depressurization of the Reactor Coolant System causes fluid to flow into the loops from the pressurizer resulting in a pressure and level decrease in the pressurizer. The reactor trip signal subsequently occurs when-the pressurizer low-pressure trip setpoint _is reached.- Loss-Of-Offsite-Power (LOOP) is assumed to occur coincident with reactor trip. A-safety. injection signal is generated when_the appropriate-
                             - setpoint (pressurizer low-low-pressure) is reached.- After the safety injection signal .is generated, an additional 27 second delay ensues. This delay conservatively models the 2 second instrumentation delay, the full 15 second diesel generator start time, plus the up to 10 seconds necessary to                                                              ,

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I APPENDIX B align the appropriate valves and increase the pumps to full speed. These _j countermeasures will limit the consequences of the accident in two ways: A. Reactor * ad borated water injection supplement void formation in causing ..pid reduction of nuclear power to a residual le ;l corresponding to the delayed fission and fission proouct decay. No credit is taken in the LOCA analysis for the boron content of the injection water. Howevar, an average RCS/ sump mixed boron concentration is calculated to ensure that the post-LOCA core remains subcritical. in addition, in the small break LOCA analysis, crer is taken for the insertion of Rod Cluster Control Asst, ies (RCCAs) subsequent to the reactor trip signal, dile assuming the mcst reactive RrCA is stuck in the full out position B. Injetion of borated water ensures sufficient flooding of the core to prevent excessive clad temperatures. fi+fere ihn bruk occurs, the plant is assumed to be in an eoutlibr%.- condition, i. , lat te n p *ert d in d e m a ;; L 4 ,9 r mm ed i> 'E secondary system. Duri..; the earlier paIt of t'nb 'smhil break tansient, 7 T effect af the break flow is not strong enough to overcome the flow main.at.ed by the reactor coolant pumps through the core as the pumps coast down following LOOP. Upward flow through the core is maintained. However, the core flow is not sufficient to prevent a partial core uncovery. Subsequently. the ECCS provides sufficient core flow to cover the core. During bluwdown, heat from fission product decay, hnt internals, and the vestel continues to be transferred to the RCS. The heat transfer between the RCS and the secondary system may be in eithic direction depending on the relative temperatures, in this case, centinued heat adcition to the secondary results iri increased secondary system pressure which leads to steam relief via the atmospheric relief valve and/or safety valves. Makeup to the secondary is automatically provided by the auxiliary feedwater pumps. The safety injection signal isolates normal feedwater flow by closing the main feedwater control and bypass valves. Loss of offsite power, assumed concurrent with reactor trip, initiates auxiliary feedvater flow by starting the auxiliary feedwater pumps. The secondary flow aid in the reduction of FO

   'O                                                B-3 I

APPENDIX B

    ,,     RCS pressure.      Also due to the loss of offsi.te power assumption, the reactor
 '(*   )   coolant pumps are assumed to be tripped at the time of reactor trip during the accident and the effects of pump coastdown are included in the blowdown analyses.

When the RCS depressurizes to approximately 600 psia, the cold leg accumulators becin to inject borated water into the reactor coolant loops. However, the vessel mixture level starts to increase to cover the fuel with ECCS pumped injection before the accumulator injection for most breaks. 15.3.1.2 Analysis of Effects and Consequences 15.3.1.2.1 Method of Analysis 2 For small breaks (less than 1.0 ft ) the NOTRUMP digital computer code (References 2 and 3) is employed to calculate the transient depressurization of the Reactor Coolant System as well as to describe the mass and energy of the fluid flow through the break. The NOTRUMP computer code is a state-of-the-art one-dimensional general network code incorporating a numb,er of advanced features. Among these are calculat' ion of thermal non-equilibrium in all . fluid volumes, flow regime-dependent drif t flux calculations with

    /l    counter-current flooding limitations, mixture level tracking logic in

multiple-stacked fluid nodes and regime-depcndent heat transfer correlations. The NOTRUMP small break LOCA emergency core cooling system (ECCS) evaluation model was developed to determine the Rr response to design basis small bre&k LOCAs, and to address NRC concerns expressed in NUREG-0611

         -(Reference 4), " Generic Evaluation of Feedwater Trantients and Small Break Loss-of-C)olant Accidents in Westinghouse-Designed Operating Plants".

The reactor coolant system model is nodalized into volumes interconnected by flowpaths. The broken loop is modeled explicitly, while the intact loops (3 lumped inte a seccad loop. Transient behavior of the system is determined from tne governing conservation equations of mass, energy, and momentum. The multinode capability of the program enables explicit, detailed spatial representation of various system components which, among other capabilities, enables a proper calculation of the behavior of the loop seal during a loss-of-coolant accident. The reactor core is represented as heated control volumes with associated phase separation models to permit transient mixture height calculations. Detailed descriptions of the NOTRUMP code and the evaluation model are provided in References 2 and 3.

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APPENDIX B Peak clad temperature calculations are performed with the LOCTA-IV code yK .(Reference 5) using the .NOTRUMP calculated core pressure, fuel rod power

    'LA         hr Cory. uncovered core steam flow and mixture heights as. boundary conditions (set Figureil5.3-1). Figure 15.3-2 depicts the hot rod axial powersshape used to perform the small brean LOCA analysis. This shape was chosen because '

it represents a distribution with power concentrated in the upper regions of the core. Such a distribution is' limiting for small-break LOC?,s because it minimizos coolant level swell, while maximizing vapor superheating and fuel rod heat generation at the uncovered elevations. The small break LOCA analysis assumes the core continues to operate at full power until the control' rods are completely inserted. However, for conservatism, it is assumed that the most reactive RCCA does not insert.

After the small break LOCA is initiated, reactor trip occurs due to a low pressurizer: pressure signal (1840 psia). Soon after the reactor trip signal

               -is generated, the safety injection actuation signal is generated due to a low-
              -pressurizer pressure (1715 psia). Safety injection systems consist of gas pressurized accumulator tanks and pumped injection systems. The small break LOCA analysis-assumed nominal accumulator water volume with a cover gas pressu're of 600 psia. (the minimum pressure allowed by the Technical                                 i Specifications). Minimum emergency core cooling system availability is as'sumed for the analysis at the maximum RWST temperature. Assumed pumped II           safety injection characteristics as a function of RCS pressure used as boundary conditions in the analysis are shown in Figure 15.3-3 and in Table
                       ~

15 . 3 - 1.- The safety injecticn- flow' rates presented are based on pump performance curves degraded 8 percent from the design head and an assumed

              'cha'rging system branch line imbalance of 10 gpm. The effect of flow from the
              .RHR pumps is not considered in the small break LOCA analyses (except for the six inch break = size) since their shutoff head is lower than the RCS pressure during the. time portion of the transient considered here.                             For the six inch break case,. the RHR flows were necessary to overcome a second minor core uncovery. Safety injection is delayed 27 seconds after the occurrence of:the
              ' low pressure condition. This accounts for signal initiation (2 seconds),

diesel generator startup and emergency power bus. loading consistent with the assumed loss of offsite power coincident with reactor trip (15 seconds) as well as the delay involved in aligning the valves and bringing the pumps up to speed (10 seconds). The small break LOCA analysis also conservatively

              .asiumed that the rod drop time is 4.0 seconds which bounds the actual value of 2.7 seconds.

[ t

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APPENDIX B a

          /                  OnLthe secondary side, a' main feedwater isolation signal is assumed to be                                   ,
                           -generated.on LOOP with a two second signal delay and a fi.ve second valve closure-time'. The auxiliary feedwater pumps are assumed to start and deliver full: flow-(one turbine driven pump and one motor driven pump) 60 seconds The auxiliary feedwater enthalpy is assumed to be
                           .after signal initiation.

that of the main'feedwater until after an additional plant specific feedwater-purge time (165 seconds for Unit 1, 147 seconds for Unit 2) has elapsed. , 15.3.1.2.2 Results t L

                           - 15.3.1. 2. 2. l'   Limiting Break Case

This section presents results of'the limiting small break LOCA analysis (as "

determir.ed by the highest calculated peak. clad temperature) from a range of break sizes, and core barrel-baffle region configurations (upflow 'for Unit I and- downflow for Unit 2). NUREG-0737 (Reference.6), Section II.K.3.31, requir'ed a plant-specific small break LOCA analysis using an Evaluation Model revised per Section II.K.3.30. In accordance with NRC Generic Letter 83-35 , (Reference ~7), generic ~ analyses using NOTRUMP (References 2 and 3) were i performed!and are presented in WCAP-lil45.(Peference 8). Those results 4 demonstrate that in i comparison- of cold leg, hot leg and pbmp suction leg

                           -break' location:;, the cold leg break location is limiting. The limiting break was-found to be a' 3-inch diameter cold leg break for the upflow barrel baffle
                           - configuration. . A list of input assumptions used in the analyses .is provided in Table'15.3-2. The results of a four break spectrum analysis performed for the upflow configuration as well as a 1imiting break size analysis performed for the downflow configuration, are summarized in Table 15.3-3, while:the key' transient event. times are listed in Table 15.3-4. The peak clad temperature in small break LOCA is-largely a function of the' depth of core uncovery which'
                           -in turn is dependent on the overall mass -inventory and ultimately the primary _

side pressure. Since the break size is the predominant determiner of the primary pressure transient and the barrel-baffle configuration has no direct i W impact on these parameters, the limiting break size is not expected to change-

                           . for the downflow configuration (Unit 2). An additional sensitivity calculation confirmed that limiting break *ize did not shift between upflow and downflow barrel-baffle configuration.

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APPENDIX B ; Figures:15.3-4 througn 11 show the following parameters, respectively, for

    'h      the. limiting three-inch break transient with upflow configuration, while Figures 15.3-12 through 19 consider the downflow configuration limiting            ,

transient. ' RCS pre ;,

                   -1   Core mixture level,                                                      ,
                   -    Hot rod clad average temperature,
                   -    Core outlet steam flow rate,
                   -    Hot assembly rod surface heat transfer coefficient, Hot spot fluid temperature,
                   -    Cold leg break mass flow rate, and
                   -    Safety injection mass flow rate.

During the initial period of the small break transient the effect of the break flow- rate is not strong enough to overcome the flow rate maintained by

          'the reactor coolant pumps as the pumps coast down folicwir.g LGOP. Normal upward _ flow-is maintained through- the core and core heat is adequately removed. At_ the low heat generation rates following reactor trip the fuel rods continue to be well cooled as long as the core is covered by-a two-phase .

4 imixture-level. From the clad temperature transients for the limiting break-f~} (3-in break) calculations shown in Figures 15.3-6 and 15.3-14, it is seen TV 'that the peak _ clad temperature occurs near the time when the core is'most

          ' deeply. uncovered and the . top of the core ~ is steam cooled. This time is accompanied by th6: highest vapor superheating above the mixture level, The peak-clad temperature attained for Unit I during the transient was 1805'F (including a-20'F pen. ty for increased T avo uncertainty). For Unit 2, the 3-inch diameter cold leg-breat case yielded a' peak clad-temperature-of 1711'F (also including the 20*F pent.lty mentioned above). At the L            time the transient was terminated, the safety injection flow rate that was delivered to the RCS exceeded the mass flow rate out the break in each case.

I EA lthough the core mixture levm has not yet covered the entire core in Unit 2 (see Figure 15.3-13), there is no longer a concern of exceeding the 10 CFR L -50.46 criteria'since the RCS pretsure is gradually decaying and there is_a L net mass inventory gain. The decreasing RCS pressure results in greater ! - safety injection flow as well as reduced break- flow. As the RCS insentory continues to gradually increase, the core mixture level will continue to b inc.* ease and the fuel clad temperatures will continue to decline. 1 i L' Additionally, only one core channel is modeled in the NOTRUMP computer code since the core flow rate during a small break LOCA is relatively slow. This g,7 u w.

           . .              --          - . . . -     -  . - -     -        - - - - .=      - - - . .

APPENDIX'B-c _

               - provides enough time to maintain flow equilibrium between fuel assemblies
   ' h- <

L(i.e.. no crossflow). ;Therefore, hydraulic resistance mismatch is not a factor for small-break LOCA and it is not necessary to perform a small break

                -LOCA~ evaluation for transition cores, and it is sufficient to reference the small break LOCA for the complete core of the VANTAGE-5 fuel design as bounding for all transition cycles.
   ^
               -15.3.1.2.2.2 Additiunal -Break Cases Studies documented _in Reference 3 dett.rmined that the limiting small_-break-size occurred for breaks less than 10 inches in diameter. To insure that the
                .3-inch ' diameter break was limiting, calculations were run with breaks of- 2     -

inches, 4 inches and 6 inches for the upflow configuration. To insure that the upflow configuration was limiting, calculations were performed at the~ limiting break size for the downflow configuration. The results of these

               -calculations are shown in the Results Table 15.3-3, and the Sequence of Events' Table 15.3-4.

For all cases-analyzed, plots of the following transient parameters are presented:

                          - RCS pressure (Q}

N - Core mixture level

                          -< Hot rod. clad average temperature                                            -

These plots are shown:in Figures 15.3-20 through 22 for the 2-inch break, r Figuresol5.3-23 through 25 for the' 4-inch break, and Figures 15.3-25 through 28 for the 6-inch break. As seen in Table -15.3-3 the peak clad _ temperatures

                 -in all cases were calculated to be less than that for the 3-inch break, upflow configuration.-

15.3.l'.3 Conclusions-I Analyses presented in this section show that the high head and low head

                ~ safety injection of the' Emergency Core Cooling System, together with the accumulators, provide sufficient core flooding to t eep the calculated peak clad temperatures below the required limit of-10 CFR 50.46. Hence adequate protection is afforded by the Emergency Core O ol k.g System in the event of a
                .small break loss-of-coolant: accident.

B-8 1

                                  .n.~.

APPENDIX B ( REFERENCES, SECTION 15.3 wJ

1. " Acceptance Criteria for Emergency Core Cooling Systems f Water Cooled Nuclear Power Reactors" 10 CFR 50.46 and Appendix K of 10 iR 50.

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

2. Meyer, P. E., "NOTRUMP - A nodal Transient Small Break and General Network Code," WCAP-10079-P-A, (Proprietary), August 1985.

3. Lee, N. et al., " Westinghouse Small Break ECCS Evaluation Model Using The NOTRUMP Code," WCAP-10054-P-A, (Proprietary), August 1985.

4. " Generic Evaltation of Feedwater Transients and Small Break Loss-of-Coolant Accidents in Westinghouse - Designed Operating Plants,"

NUREG-0611, January 1980.

5. Bordelon, F. M., et al., "LOCTA-IV Program: Loss-of-Coolant Transient Analysis," WCAP-8305, June 1974, WCAP-8301, (Proprietary), June 1974.

6. " Clarification of THI Action Plan Requirements," NUREG-0737, November

/7 1980, h ~

7. NRC Generic Letter 83-35 from D. G. Eisenhut, " Clarification of TMI Action Plan item !!.K.3.31", November 2, 1983.

8. Rupprecht, S. D., et al., "Wettinghouse Small Break LOCA ECCS Evaluation Model Generic Study With the NOTRUMP Code;" WCAP-lll45-P-A, (Proprietary), October 1986.

(M 'k j B-9

, TABLE 15.3-1 k -- SAFETY INJECTION FLOW RATE . RCS PRESSURE RHR FLOW RATE SI FLOW RATE SI FLOW RATE (spill to O psig) (spill to O psi )9 (spill to RCS) [ psia] [lb/sec) [lb/sec) [lb/sec) 15 354.5 47.64 47.64 25 332.6 47.48 47.54 35 310.0 47.32 47.44 45- 286.9 47.15 47.34 55 262.9 46.99 47.23 65 237.8 46.83 47.13 IN 75 211.4 46.67 47.03 85 183.8 46.51 46.93 95 153.8 46.34 .46.83 105 121.0 46.18 46.73 115 85.0 46.02 46.62 125 43.4 45.. 46.52 135 0.0 45.70 46.42 (cont.) E-10 (3 A;

fm, . TABLE 15.3-1 (cont.) V SAFETY INJECTION FLOW RATE

RCS PRESSURE RHR TLOW RATE S1 FLOW RATE St FLOW RATE (spill to O psig) (spill to 0 psig) (spill to RCS)

(psia) (1b/sec) (lb/sec) [lb/sec) 215 0.0 4o.40 45.61 415- 0.0 41.00 43.51 615 0.0 37.49 41.29 815 0.0 33.73 38.98 1015 0.0 28.95 36.61 . 1215 0.0 23.21 34.18 7 (. 1415 0.0 17.07 31.43 1615 0.0 10.31 28.53 1815 0,0 2,14 25.17 2015 0.0 0.00 21.45 2215 0.0 0.00 16.00 2415- 0.0 0.00 0.79 2615 0.0 0.00 0.00 This table assumes flow from one high head safety injection (SI) pump and one low head (RHR) pump. B-ll

   /3-V

TABLE 15.3-2

b) PLANT INPUT PARAMETERS USED IN SMALL BREAK LOCA ANALYSIS FOR 17 X 17 VANTAGE-5 FUEL Core Power 102% of 2775 MWt Total Core Peaking Factor (Fq) 2.50 Enthalpy Risc Peaking Factor (FAH) 1.70 Steam Generator Tube Plugging Level 20% (peak uniform) A:cumulator Conditions: Cover Gas Pressure 600 psia Water Volume- 1025 ft3 Total Volume 1450 ft 3 RCS Initial Conditions: Loop Temperatures Consistent

            -With Tavg rrogram Setpoint of,            577.20F Pressure                                  2310 psia              .

Vessel Flowrate 258000 gpm l Reactor Trip Signal 1840 psia Safety Injection _ Signal 1715 psia Safety Injection Delay Time 27 seconds Rod Drop Time 4.0 seconds MFW Isolation Time Delay Time 2.0 seconds l Valve Closure Time 5.0 seconds l L ls l l l

       ]                                         B-12 l

l l

S. mig ,-~ (J '

    \                                                      ~-)                                                                                          .QI TABLE 15.3-3                                                                                         ,

L SMALL BREAK LOSS OF COOLANT ACCIDENT CALCULATION Results Parameter Unit 1 Unit 2 Case A Case B Case C Case D Case E 2-inch .3-inch 4-inch G-inch 3-inch Upflow Upflow Upflow Upflow Downflow Peak Clad Temperature (*f) 1071 1805* 1458 1771 1711 [ Elevation (ft) 11.5 11.75 11.5 11.25 11.5 T' Zr/ll 20 Cumulative Rt. action C Local Maximum (%) 0.08 2.36 C 31 0.88 1.52 Elevation (ft) 11.5 11.75 11.5 11.25 11.75 Total Core (%) <l.0 <l.0 <l.0 <l.0 <l.0 Rod Burst Hone None None None N,one

These PCTs include a 20*F penalty due to an increased T avg uncertainty of 1 6*f.

_ _ _ _ _ _-___-.______________.__.._.________._._.____.___._m

       )                                                       j                                                                                                     _)

TABLE 15.3-4 TlHE SEQUENCE Of EVENTS FOR CONDITION 111 EVENTS Small Break Loss of Coolant Accident Time (s) Event Unit 1 Unit 2 Case A Case B Case C Case D Casc E 2-inch 3-inch 4-inch 6-inch 3-inch Upflow Upflow Upflow Upflow. Downflow Break occurs 0.0 0.0 0.0 0.0 0.0 Reactor trip signal 36.34 15.37 9.38 6.11 15.31 Safety injection signal 51.30 26.14 !9.05 8.87 25.85

 ?     Start of safety injection delivery         78.30       S3.14         46.05                                                                   35.87    52.85

% Start of auxiliary feedwater delivery 06.34 75.37 59.38 66.11 75.31 l Loop seal venting (initial) 2006.1 436.7 186.0 95.82 424.8 l Loop seal core uncovery 1055.1 475.0 227.4 115.2 464.8 Loop seal core recovery 1072.0 !.10. 2 262.6 130.4 486.7 Boil-off core uncovery 1474.6 7. '6. 9 419.3 162.0 821.0 Accumulator injection begins (1) 1121.4 612.6 261.3 1,166.9 Peak clad temperature occurs 3598.8 1304.3 717.4 320.7 1243.2 Tcp of core recovered (2) 2062.7 1798.7 394.8 (2)

Si flow rate exceeds break flow rate 4929.8 1779.8 (3) 2023.5 1875.2 l l (1) System pressure never drops below the accumulator cutin pressure (600 psia). (2) Although the core is not yet covered in these cases, Si flow exceeds break flow and the clad - temperature transient is over. (3) Although Si flow has not yet matched break flow, the core is covered, the clad temperature transient has ended, and the total RCS mass is increasing, b

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         -.                                                                                       (DOWNFLCW)-

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JOSEPH M. FAHLEY UNIT 2 (DOWNFLOW) (l 'v' Alabama Power

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FIGURE 15.3-21 i.

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2400. 2200. 2000. 1800. E G 1600. E 1400. 3 l 1200. . 5 5 1000.

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E

600. - 400. 200.-- - - - --

( .- 15 . 500. 1000. 1500. 2000. 2500, i- TIME (SEC) 1 RCS DEPRESSURIZATION TRANSIENT A JosalPH M. FARulY UNIT 1 (UPFLOW) Alabama Power wucts.An m .(6-INCH BREAK) UNff 1 AND UNf72 FIGURE 15.3-26 B-40 l l-

      !    t 54.

52. I I 50. i I 28. N

                -           1                      ill .

26. y i

                                               ;iy {j b

a w 24.

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I0 N l' - le D % w a

22. _,. . , _ __.______. _ _ _ ^77 x )

,s , r 20- y I) i O j 8 18. l 16. 9

14. j 12.

h

        !           10 D.               500.              1000.            1500. 2000.             2500.

TIME (SEC) CORE MIXTURE JOSEPH E FARLEY UNIT 1 (UPFLOW) zm Alabama Power NUCt. EAR Pt. ANT (6-INCH BREAK) (,,) UIC 1 AND UNIT 2 FIGURE 15.3-27 B-41

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

O ! 1900.

                                                                                                                            >l 1600.

l 1 1400. I _ 1200. b d - 3 1000. ) E O & r d 800. I !. 6r0. i l-dd 400.. 200'O. 200. 400. 600. 800. 1000. 1200. TIME (5) CLAD AVERAGE TEMPERATURE - HOT ROI p JosarH u m LEY UNIT 1 (UPFLOW) i Alabama Power wucuunmurr (6-INCH BREAK) FIGURE 15.3-28 8-42

l I APPENDIX B 15.4z CONDITION IV-LIMITING FAULTS

 'd        Condition.IV occurrences are faults which are not expected to take place during the lifetime of Joseph M. Farley Nuclear Plant Unit 1 or 2, but are                                           '

postulated because their consequences would include the potential for the release of significant amounts of radioactive material. They_are the most drastic'occurrances which must be designed against and thus_ represent limiting design cases. Condition IV faults are not-to cause a fission product release to the environment resulting in an undue risk to public health and safety in excess of guideline values of 10 CFR 100. A single Condition IV fault is not to cause a cor. sequential loss of required functions of systems needed to cope with the fault, including those of the emergency core cooling system (ECCS) and the containment. For the purposes of this report, the following faults have been classified in this category:

1. Major rupture of pipes containing reactor coolant up to and including double-ended rupture of the largest pipe in the reactor ,

coolant systen (RCS) (Loss-of-Coolant Accident (LOCA)). q 2. Major secondary system pipe rupture up to and including double-ended lg rupture (rupture of a steam pipe).

3. Steam generator tube rupture.

4. Single reactor coolant pump locked rotor.

5. Fuel handling accident (FHA),

6. Rupture of a control rod mechanism housing (rod cluster control assembly (RCCA) ejection).

15.4.1 MAJOR REACTOR COOLANT SYSTEM PIPE RUPTURES (LARGE BREAK LOCA) 15.4.1.1 Identification of Causes and Frecuency Classification A loss-of-coolant accident (LOCA) is the result of a pipe rupture of the reactor coolant system (RCS) pressure boundary. For the analyses reported here, a major pipe break (large break) is defined as a rupture with a total cross-sectional area equal to or greater than 1.0 ft2. This event is-considered an American Nuclear Society (ANS) Condition IV event, a limiting ( fault, in that it is not expected to occur during the lifetirne of Joseph M. Farley Nuclear Plant Unit 1 or 2, but is postulated as a conservative design basis. B-43

APPENDIX B For large break LOCAs, the most limiting single failure is the one which produces the lowest containment pressure. The lowest containment pressure

   -t would be obtained only if all containment spray pumps and fan coolers operated subsequent to the postulated LOCA. Therefore, for the purposes of large break LOCA analyses, the most limiting single failure would only be the los: cf one RHR pump with full operation of the spray pumps and fan coolers.

However, the large break- LOCA analyses conservatively assume both maxirnum containment safeguards-(lowest containment pressure) and minimum ECCS safeguards (the loss of one complete train of ECCS components which includes one RHR pump and one HHSI pump), which results in the minimum delivered ECCS flow available to the RCS. The Acceptance Criteria for the LOCA are described in 10 CFR 50.46 (Reference

1) as follows:

A. The calculated peak fuel element clad temperature does not exceed the requirement of 2200*F. B. The amount of fuel element cladding that reacts chemically with water or steam to generate hydrogen, does not exceed 1 percent of the total amount of Zircaloy in the fuel rod cladding. C. The clad temperature transient is terminated at a time when the core geometry is still amenable to cooling. The localized cladding oxidation limit of 17 percent is not exceeded during or after quenching. D. The core remains amenable to cooling during and after the break. E. The core temperature is reduced and decay heat is removed for an extended period of time, as required by the long-lived radioactivity j remaining in the core. These criteria were established to provide a significant margin in emergency l core cooling system (ECCS) performance following a LOCA. WASH-1400 (USNRC 1975) (Reference 2) presents a study in regards to the probability of occurrence of RCS pipe ruptures. 15.4.1.2 Seauence of Events and Systems Goerations Thould a major break occur, depressurization of the RCS results in a pressure decrease in the pressurizer. Loss-Of-Offsite Power (LOOP) is assumed coincident with tie occurrence of the break. .The reactor trip signal subsequently occurs when the pressurizer low pressure trip setpoint is reached. A :afety injection signal is generated when the appropriate l B-44

APPENDIX B setpoint (high containment pressure or low pressurizer pressure) is reached. Thest tountermeh,;ures will limit the consequences of the accident in two g ways: A. Reactor trip and borated water injection supplement void formation in causing rapid reduction of power to the residual level corresponding to fission product decay heat. No credit is taken in the LOCA analysis for the boron content of the injection water. However, an average RCS/ sump mixed boron concentration is calculated to ensure that the post-LOCA core remains subcritical. In addition, the insertion of control rods to shut down the reactor is neglected in the large break analysis. B. Injection of borated water provides for heat transfer from the core and prevents excessive clad temperatures. In the present Westinghouse design, the most limiti g large break single failure is the loss of one RHR (low head) pump. As previously stated in

                        -Section 15.4.1.1, credit could be taken for two high head safety injection pumps and one low head pump for a large break. However, the specific large.

break analysis for Farley has only taken credit for one high head pump and one low head pump. This assumption is consistent with the current procedure for large break analyses. For the large break analysis, one ECCS train, including one high head safety injection pump and one RHR (low head) pump, startt and delivers flow through the injection lines (one for each loop) with one branch injection line spilling to the containment backpressure. However, both Emergency Diesel Generators (EDGs) are assumed to start in the modeling of the containment l fan coolers and spray pumps. Modeling full containment heat removal systems ll operation is required by Branch Technical Position CSB 6-1 and is [ conservative for the large break LOCA. To minimize delivery to the_ reactor, the branch line chosen to spill is selected as the one with the minimum resistance. In addition, both the high ! head safety injection pump and the RHR pump performance curves were degraded by 8% and a 10 gpm flow imbalance was assumed for the high head safety l- injection pumps. In the large break ECCS analysis presented here, single failure is conservatively accounted for via the loss of an ECCS train, the spilling of the minimum resistance injection line, and by assuming all containment spray pumps and fan coolers are available. Therefore, the analysis assumed one high head pump, one RHR pump, two containment spray pumps, and four fan coolers are operating. B-45

APPENDIX B 15.4.1.3 -Descriotion of Laroe Break LOCA Transient The time sequence of events following a large break LOCA is presented in Figure 15.4-1A, 15.4-1B, and Table 15.4.1-5. The safety injection 1 performance, as modeled for the various large break LOCA cases, is presented in Figures 15.4-2A and 15.4-2B. 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 fission product decay, hot internals and the vessel, l continues to be transferred to the reactor coolant. At the beginning of the blowdown phase, the entire RCS contains subcooled liquid which transfers 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 10 CFR 50 (Reference 1). Thereafter, the core heat transfer is unstable, with both nucloate boiling and film boiling occurring. As the core becomes uncovered, both transition boiling and forced convection are considered as the dominant core heat transfer mechanisms. Heat transfer due to radiation is also considered. The heat transfer between the RCS and the secondary system may be in either direction, depending on the relative temperatures. In the case of continued heat addition to the secondary system, the secondary system pressure increases and the main steam safety valves may actuate to limit the pressure. Makeup water to the secondary side is automatically provided by the auxiliary feedwater system. The safety injection signal actuates a feedwater isolation signal which isolates main feedwater flow by closing the main feedwater control and bypass valves and also initiates auxiliary feedwater flow by starting the auxiliary feedwater pumps. The secondary flow aids in the reduction of RCS pressure. When the RCS depressurizes to 600 psia, the accumulators begin to inject borated water into the reactor coolant loops. The conservative assumption is

 .      made that all of the accumulator water injected during the bypass period is

[ subtracted from the RCS after the bypass period terminates (called l-end-of-bypass). End-of-bypass occurs when the expulsion or entrainment != mechanisms responsible for the bypassing are calculated not to be effective. ( -This conservatism is again consistent with Appendix K of 10 CFR 50. Since-L LOOP is assumed, the reactor coolant pumps are assumed to trip at the j: inception of the accident. The effects of pump coastdown are included in the blowdown analysis. The blowdown phase of the transient ends when the RCS pressure (initially assumed at 2310 psia) falls to a value approaching that of the containment atmosphere. Prior to or at the end of the blowdown, termination of bypass B-46

I APPENDIX B occurs and refill of the reactor vessel lower plenum begins. Refill is completed when emergency core cooling water has filled the lower plenum of j O the reactor vessel, which is bounded by the bottom of the fuel rods (called bottom of core (B0C) recovery time). The reflood phase of the transient is defined as the time period lasting from BOC recovery until the reactor vessel has been filled with water to the extent that the core temperature rise has been terminated. From the latter stage of blowdown and then the beginning of reflood, the safety inject Pbn accumulator tanks rapidly discharge borated cooling water into the RCS, thus contributing to the filling of the reactor vessel downcomer. The downcomer water elevation head provides the driving force required for the reflooding of the reactor core. The RHR (low head) and high head safety injection pumps aid in the filling of the downcomer and subsequently supply water to maintain a full downcomer and complete the reflooding process. Continued operation of the ECCS pumps supplies water during long-term cooling. Core temperatures have been reduced to long-term steady state levels associated with dissipation of residual heat generation. After the water level of the refueling water storage tank (RWST) reaches a minimum , allowable value, coolant for long-term cooling of the core is obtained by switching to the cold isg recirculation phase of operation. Spilled borated p water is drawn from the engineered safety features (ESF) containment sumps by (%) the RHR (low head) pumps and returned to the RCS cold legs. The containment spray pumps are manually aligned to the containment emergency sumps and continue to operate to further reduce containment pressure and temperature. Increased volume due to the smaller diameter VANTAGE 5 fuel rods will not significantly impact the total RCS volume. Approximately 11 hours after the initiation of the LOCA, the ECCS is realigned to inject water to the RCS hot legs in order to control the boric acid concentration in the reactor vessel. Long-term cooling includes long-term criticality control. Criticality control is achieved by determining the RWST and accumulator concentrations necessary to maintain subtriticality without credit for RCCA insertion. The necessary RWST and accumulator concentrations are a function of each core design and are verified each cycle. The current Technical Specifications range is 2300 to 2500 ppm boron for the RWST and 2200 to 2500 ppm for the accumulators. 15.4.1.4 Analysis of Effects and Consecuences 15.4.1.4.1 Method of Analysis ( The requirements of an acceptable ECCS evaluation model are presented in Appendix K of 10 CFR 50 (Reference 1). B-47

APPENDIX B 15.4.1.4.2 Large Break LOCA Evaluation Model The analysis of a large break LOCA transient is divided into three phases: (1) blowdown, (2) refill, and (3) reflood. There are three distinct transients analyzed in each phase, including the thermal-hydraulic transient in the RCS, the pressure and temperature transient within the containment, and the fuel and 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 of the LOCA. A description of the various aspects of the LOCA analysis methodology is given by Bordelon, Massie, and Zordan (1974) (Reference 3). 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, WREFLOOD, BASH and LOCBART codes, which are used in the LOCA analysis, are described in detail by Bordelon et al. (1974) (Reference 4); Kelly et al. (1974) (Reference 5); Young et al. (1987) (Reference 6); and Bordelon et al. (1974) (Reference 3). Code modifications are specified in References 7, 8, 9, and 10. These codes assess the core heat transfer geometry and determine if the core remains amenable-to cooling through and subsequent to the blowdown, refill, and reflood phases of the LOCA. SATAN-VI calculates the thermal-hydraulic transient, including the RCS pressure, enthalpy, density, and the mass and energy flow rates in the RCS, as well as steam generator energy transfer between the primary and secondary systems-as a function of time during the blowdown phase of the LOCA.

SATAN-VI also calculates the accumulator water mass and_ internal pressure and the break mass and energy flow rates that are assumed to be vented to tr -

containment during blowdown. At the end of the blowdown phase, the mass and energy release rates during blowdown are transferred to the C0C0 code, detailed in Reference 11, for use in determination tf the containment pressure response during the first phase of the LOCA. Additional SATAN-VI output data from the end-of-blowdown, including the core _ inlet flow rate and enthalpy, the core pressure, and the core power decay transient, are input to the LOCBART code. At the end of ie blowdown, information from SATAN-VI on the state of- the system is trar.sferred.to-the WREFLOOD code which calculates the time to BOC recovery, RCS conditions at B0C and mass and energy release from the break during the reflood phase of the LOCA. Since the mass flow rate to the containment depends upon the core flooding rate and the local core pressure, which is a function of the containment backpressure, the WREFLOOD and C0C0 O codes are interactively linked. The BOC conditions calculated by WREFLOOD and the containment pressure transient calculated by C0C0 are used as input to the BASH code. Data from both SATAN-VI code and the WREFLOOD code out to B-48

APPENDIX B BOC are input to-the LOCBART code which calculates core average conditions at BOC for'use by_the BASH _ code. The BASH code provides a realistic thermal-hydraulic simulation of- the- , reactor core' and RC" during the reflood phase of a large break LOCA. Instantaneous values of the accumulator conditions and safety injection flow at the time of completion of lower plenum refill are provided to BASH by WREFLOOD. Figure 15.4-1B illustrates how BASH has been substituted for WREFLOOD in calculating transient values of core inlet flow, enthalpy, and pressure 'for the detailed fuel rod model, LOCBART. A detailed description of the-BASH code is available in Reference 6. The BASH code provides a sophisticated treatment of steam / water flow phenomena in the reactor coolant system during core reflood. A dynamic interaction between core , thermal-hydraulics and-system behavior is expected, and experiments have shown this behavior. The BART code has been coupled with a loop model to form the BASH code. The loop' model determines the loop flows and pressure drops in response to the calculated core exit flow determined by BART. The updated core inlet. flow calculated by the loop model is used by BART to calculate a new entrainment rate to be fed into the loop code. This process of transferring data between BART, the loop code and back to BART forms the calculational process for analyzing the reflood transient. This coupling df the BART code with a loop code produces a dynamic flooding transient, which reflects the close coupling between core thermal-hydraulics and loop

g. behavior.

The cladding heat-up transient is calculated by LOCBART which is a combination of the LOCTA code with BART. A more detailed description of the LOCBART code can be found in References 3 and 8. During reflood, the LOCBART code provides a significant improvement in the prediction of fuel __ rod behavior. In LOCBART the empirical FLECHT correlation has been replaced by the BART' code. BART employs rigorous mechanistic models to generate heat transfer coefficients appropriate to th.e actual flow and heat transfer regimes experienced by the fuel rods. Modeling ' features.necessary to account for the reactor barrel-baffle region and the reactor fuel assembly thimbles were included in this analysis as presented in Reference 10. 15.4.1.4.3 Input Parameters and Initial Conditions important input parameters and initial conditions used in the analysis are listed in Tables 15.4-2 and 15.4-3. Cases analyzed are given in Table 15.4-1. O The initial steady state fuel pellet temperature and fuel rod internal B-49

APPENDIX B pressure used in the LOCA analysis was generated with the PAD 3.4 Fuel Rod Design Code (Ref_erence 12) which has been approved by the Nuclear Regulatory Commission. A full spectrum break analysis was done in order to justify plant operation at 2652 MWt at initial RCS pressurizer pressure of 2310 psia and initial hot leg temperature of 611.3*F from which the limiting break size was determined, in addition, the analyses conservatively modeled a downflow barrel-baffle configuration. However, the limiting break size was also analyzed with an upflow barrel-baffle configuration, applicable to Unit 1, which demonstrated that the downflow barrel-baffle configuration remained limiting. The upflow barrel-baffle design would not cause the limiting break size to shift since there is only an incremental difference in the core flow during l blowdown as a result of the different barrel-baffle configurations. Since a larger discharge coefficient has substantially better cooling during the blowdown calculation, an incremental difference in the core flow due to the barrel-baffle configuration will not significantly alter the clad temperature response during blowdown. In addition, the barrel-baffle configuration affects the calculation of the reflood portion of the transient through

vessel liquid level changes. The flow of steam through the loop and out the rupture will only be affected by the amount of steam generated in the

 }O-_ reflooding and quenching process.

Also, previous Farley specific analyses, which performed a full spectrum analysis using ti;e upflow barrel-baffle design, demonstrated that the limiting break size was the same as the downflow barrel-baffle configuration. Therefore, a complete spectrum assuming the upflow barrel-baffle configuration was not performed in this analysis sinte the barrel-baffle configuration -will not change the limiting break size. All cases conservatively _ assumed 20* -team generator tube plugging in all three steam generators and an 8% aegradation of both the RHR and high head safety injection pumps with a 10 gpm high head safety injection flow imbalance. Table 15.4-1 describes the cases analyzed. The bases used to select the numerical values that are input parameters to the analysis have been conservatively determined from extensive sensitivity studies (Westinghouse 1974 (Reference 13); Salvatori 1974 (Reference 14); Johnson, Massie, and Thompson 1975 (Reference 15)). In addition, the requirements of Appendix K to 10 CFR 50 (Reference 1) regarding specific rdel features were met by selecting models which provide a significant O overall conservatism in the analysis. The assumptions which were made pertain to the conditions of the reactor and associated safety system B-50

APPENDIX B equipment at the time that t'ae LOCA occurs, and include such items as the core peaking fact es, the containment pressure, and the performance of the ECCS. Decay heat generated throughout the transient is also conservatively calculated as per the requirements of Appendix K to 10 CFR 50 (Reference 1). Another input parameter that affects LOCA analysis results is the assumed axial power shape at the beginning of the accident. Power shbr sensitivity studies performed with Westinghouse ECCS evaluation models have always demonstrated the chopped cosine shape with the peak at the core midplane to be limiting. Westinghouse has performed " spot check" analyses using the BASH reflood evaluation model for power shapes skewed to the top of the core. Results of these analyses have demonstrated the chopped cosine peaked at the core midplane remains the limiting power shape (Reference 16). 15.4.1.4.4 Additional Break Cases Westinghouse ECCS analyses currently assume minimum safeguards for the safety injectlon flow, which minimizes the amount of flow to the RCS by assuming maximum injection line resistances. However, for some Westinghouse plants, including Farley Nuclear Plant Units 1 and 2, the current nature of the ' Appendix K ECCS evaluation model is such that it may be more limiting to assume the = maximum possible ECCS flow delivery. In that case, maximum p safeguards, which assume operation of both trains of RHR (low head) and high P head safety injection pumps, minimum injection line resistances and enhanced ECCS pump performance result in the highest amount of flow delivered to the RCS. Therefore, the worst break for the Farley Nuclear Plant Unit I and 2 was reanalyzed, assuming maximum safeguards. Examination of the LOCA analysis results in Table 15.4-6 demonstrates that mimimum safeguards assumptions result in the limiting peak clad temperature for Farley Nuclear Plant Units 1 and 2. The worst break for the Farley Nuclear Plant Units 1 and 2 was also analyzed for this first transition core assuming 17X17 LOPAR fuel input parameters (the current hot channel enthalpy rise factor of 1.55 and a total core peaking factor of 2.32 were retained). In addition, a conservative assumption of 4000 MWD /MTU minimum barnup for the 17X17 LOPAR fuel and a maximum of 72 fresh 17X17 VANTAGE 5 fuel assemblies was used in the analysis. Since 17X17 VANTAGE 5 and 17X17 LOPAR fuel have different overall hydraulic resistances and grid effects, both the hydraulic transient and the clad heatup transient we-e modeled. .The 17X17 VANTAGE 5 fuel case for the downflow barrel-baffle configuration remained limiting. The use of existing 17X17 LOPAR fuel in subsequent cycles will have obtained sufficient burnup to be bounded by the first transition core analysis, i i B-51 L 1 L - ,_m... , _ . - - , . , ,

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

APPENDIX B A complete spectrum assuming 17X17 LOPAR fuel was not performed based.on Westinghouse experience with three loop plants analyzed with the 1981 Evaluation Model plus BASH. All-three loop plants, including the 17X17 l/]L VANTAGE 5 analysis for Farley, have been shown _ to have a 0.4 limiting break discharge coefficient (C D) independent of the fuel type analyzed (Reference 17). Therefore, only the limiting break size was analyzed for 17X17 LOPAR fuel. 15.4.1.4.5 Transition Core Effects When assessing the effect of transition cores on the large break LOCA alalysis, it must be determined whether the transition core can have a greater calculated peak cladding temperature (PCT) than either a complete core of the 17x17 LOPAR assembly design or a complete cure of the Westinghouse 17x17 VANTAGE 5 design. For a given peaking factor, the only mechanism available to cause a-transition core to have a greater calculated PCT than a full core of either fuel is the possibility of flow redistribution due to fuel assembly hydraulic resistance mismatch. Hydraulic resistance mismatch will exist only for a tran;ition core and is the only unique difference between a complete core of either fuel type and the transition i core. !- T Westinghouse transition core designs, including specific 17X17 LOPAR to 17x17 ( VANTAGE 5 transition core cases, were analyzed. The increase in hydraulic resistance for the VANTAGE 5 assembly was shown to produce a reduction in reflood steam flow rate for the VANTAGE 5 fuel at mixing va.ie grid elevations for transition core configurations. The various fuel assembly specific transition core analyses performed resulted in peak cladding temperature increases of up to 50*F for core axial-elevations that bound the location L of the PCi. Therefore, the maximum PCT penalty possible for VANTAGE 5 fuel L residing in a transition core is 50*F (Reference 18). Once a full core L of VANTAGE 5 fuel is achieved the large break LOCA analysis will apply without the transition core penalty. Since the increased hydraulic resistance for the VANTAGE 5 assemblies produces a reduction in flow rate for the VANTAGE 5 fuel in a mixed fuel situation, the transition core penalty is only applied to VANTAGE 5 analyses and not to the LOPAR fuel analyses.

              -15.4.1.5    Results Based on the results of the LOCA sensitivity studies (Westinghouse 1974 (Reference 9); Salvatori 1974 (Reference 14); Johnson, Massie, and Thompson O           1975 (Reference 15)), the limiting large break was found to be the double-ended cold leg guillotine (DECLG). Therefore, only the DECLG break is B-52

APPENDIX B considered in the large break ECCS performance analysis. Calculations were performed for a range of Moody break discharge coefficients. The results of f these calculations are summarized in Tables 15.4-5 and 15.4-6. The mass and energy release data during reflood for cases A and B are shown in Table 15.4-7. Spilling accumulator flows are provided in Table 15.4-8 and the minimum and maximum pumped safety injection flows are plotted in Figures 15.4-2A and B, respectively. Figures 15.4-3A through 15,4-178 present the results of the cases analyzed for the large_ break LOCA. The alpha designation in the figure number corresponds to the cases as described in Table 15.4-1. ) Figures 15.4-3A-E The system pressure shown is the calculated core pressure, j Figures 15.4-4A-E The flow rate from t e break is plotted as the sum of both ends of the guillotine break. Figures 15.4-5A-E The power transient is plotted as a fraction of the ' nominal pcwer.

 't     Figures 15.4-6A-E         The core flow rate is -shown during the blowdown phase of the transient.

Figures 15.4-7A-E The accumulator flow rate during blowdown is plotted as the sum of that injected into the intact cold legs. Figures 15.4-8A-E- The core and downcomer collapsed liquid water levels are plotted during the reflood phase of the transient. Figures 15.4-9A-E The core inlet flow velocity is shown as it is calculated during the reflood phase. Figures 15.4-10A-E The break energy released to containment as calculated by SATAN is shown. Figures 15.4-llA-E The fluid quality as calculated by LOCBART is shown. Figures 15.4-12A-E The fluid velocity is plotted at the hot spot (the v node which produced the peak clad temperature) on the L hot rod. B-53 1

APPENDlX B Figures 15.4-13A-E The heat transfer coefficient is plotted at the hot spot on the hot rod. g

      ]
 "'                                                                         The fluid temperature at the hot spot on the hot rod Figures 15.4-14A-E is plotted.

Figures 15.4-15A-E The clad temperature at the hot spot is shown for the hot rod. Figures 15.4-16A-E The containment backpressure transient used in the analysis. Figures 15.4-17A-B The containment condensing wall heat transfer coefficient. The limiting peak clad temperature calculated for the Unit 1 17X17 VANTAGE 5 large break is 1982.l*F, which is less than the acceptance criteria limit of 2200*F. This value includes 3*F for containment mini-purge auto isolation, 8'F for increased temperature uncertainty, and 6*F for combined Safe Shutdown Earthquake (SSE) and LOCA events. Addition of the 50*F transition penalty yields a transition core peak clad temperature of ' 2032.l*F. This transition core penalty can be accommodated by the ( 2200*F acceptance criterion of 10 CFR 50.46. The maximum local M] metal-water reaction is 4.96 percent, which is well below the embrittlement limit of 17 percent as required by 10 CFR 50.46. The limiting peak clad temperature calculated for the Unit 217X17 VANTAGE 5 large break is 2090.3*F, which is less than the acceptance criteria limit of 2200*F. This value includes 3*F for containment mini-purge auto isolation, 8'F for increased temperature uncertainty, and 6*F for combined SSE and LOCA events. Addition of the 50*F transition penalty yields a transition core peak clad temperature of 2146.3*F. This O transition core penalty can be accommodated by the 2200*F acceptance l criterion of 10 CFR 50.46. The maximum local metal-water reaction is 6.59 percent, which is well below the embrittlement limit of 17 percent as required by 10 CFR 50.46. The total core metal-water reaction is less than 1.0 percent for all breaks analyzed, corresponding to less than 1.0 percent hydrogen generation, as compared with the 1 percent criterion of 10 CFR 50.46. The clad temperature transient is terminated at a time when the core geometry is still amenable to cooling. As a result, the core temperature will continue to drop and the e ability to remove deca; heat generated in the fuel for an extended period of (. time will be provided. B-54 )

APPENDIX B 15.4.1.6 References

    .                                                                                                l

1. " Acceptance Criteria for Emergency Core Cooling System for light Water l

                  . Cooled Nuclear Power Reactors," 10 CFR 50.46 and Appendix K of 10 CFR 50,

[gde_ral Reaister 1974, Volume 39, Number 3.

2. U. S. Nuclear Regulatory Commission 197L, " Reactor Safety Study - An l Assessment of Accident Risks in U. S. Commercial Nuclear Power Plants" l WASH-1400, NUREG-75/014. ;

3. Bordelon, F. M.; Massie, H. W.; and Zordan, T. A. " Westinghouse ECCS Evaluation Model - Summary," WC, -8339, July 1974.

4. Bordelon, F. M. et al., " SATAN-VI Program: Comprehensive Space-Time Dependent Analysis of Loss-of-Coolant," WCAP-8302 (Proprietary) and WCAP-8306 (Non-Proprietary), June 1974.

5. Kelly, P.. D. et al. , " Calculation model for core Reflooding After a Lass-of-Coolant Accident (WREFLOOD Code)," WCAP-8170 (Proprietary) and WCAP-8171 (Non-Proprietary), June 1974. .

6. Young, M. Y. et al,. "The 1981 Version of the Westinghouse ECCS Evaluation p Model Using the BASH Code," WCAP-10266-P-A Rev. 2 (Proprietary), March p .

1987.

7. Rahe, E. P. (Westinghouse), letter to J. R. Mille- (USNRC), Letter No.

NS-EPRS-2679, November 1982.

8. Rahe, E. P., " Westinghouse ECCS Evaluation Model, 1981 Version,"

WCAP-9220-P-A (Proprietary Version), WCAP-9221-P-A (Non-Proprietary version), Revision 1, February 1982.

9. Bordelon, F. M., et al., " Westinghouse ECCS Evaluation Model -

Supplementary Information," WCAP-8471 (Proprietary) and WCAP-8472 (Non-proprietary), April 1975. l l

10. Special Report NS-NRC-85-3025(NP), "BART-WREFLOOD Input Revision."

i

11. Bordelon, F. M., and Murphy, E. T., " Containment Pressure Analysis Code (C0CO)," WCAP-8327 (Proprietary) and WCAP-8326 (Non-Proprietary), June 1974.

12. Weiner, R. A., et al., " Improved Fuel Performance Models for Westinghouse 1 ( -

Fuel Rod Design and Safety Evaluations", WCAP-10851-P-A, August 1988. B-55

 !-                                                                                                      )

APPENDIX B l

13. " Westinghouse ECCS - Evaluation Model Sensitivity Studies," WCAP-8341 I (Proprietary) and WCAP-8342 (Non-proprietary), July 1974.

O 14. Salvatori, R., " Westinghouse ECCS - Plant Sensitivity Studies," WCAP-8340 (Proprietary) and WCAP-8356 (Non-proprietary), July 1974,

15. Johnson, W. J.; Massie, H. W.; and Thompson, C. M. " Westinghouse ECCS Four Loop Plant (17x17) Sensitivity Studies," WCAP-8565-P-A (Proprietary) I and WCAP-8566-A (Non-Proprietary), July 1975.

16. Besspiata, J.J., et al., ~The 1981 Version of the Westinghouse ECCS Evaluation Model Using the BASH Code, Power Shape Sensitivity Studies",

WCAP-10266-P-A Revision 2 Addendum 1 (Proprietary), December 15, 1987. I

17. Knochel, J. A. (Westinghouse), Letter to J. D. Woodard (Alabama Power Company), Letter No. ALA-91-618. " Response to NRC Questions Regarding Downflow Barrel Baffle", April 1,1991,

18. Davidson, S.L. and Kramer, W.R.; (Ed.) " Reference Core Report VANTAGE 5 fuel Assembly", WCAP-10444-P-A (Proprietary), September 1985. ,

H LO B-56 l l l

APPENDlX B JOSEPH M. FARLEY NUCLEAR PLANT UNITS 1 AND 2 O TABLE 15.4-1 LARGE BREAK LOCA - CASES ANALYZED CASE A - CD =0.4, 2652 MWt Core Power, Fc =2.45, FAH 1.65, i P-BAR-HA=1.4426, Minimum Si anc upflow (Unit 1) barrel-b3ffle configuration. CASE B - CD =0.4, 2652 MWt Core Power. T,=2.45, c FAH 1.65, P-BAR-HA 1.42, Minimum 51 and cownflow (Unit 2) barrel-baffle configuration. CASE C - CD =0.6, 2652 MWt Core Power, F. c 2.45, FAH 1.65, P-BAR-HA 1.42, Minimum Si and cownflow (Unit 2) barrel-baffle configuration. CASE D - CD =0.8, 2652 MWt Core Power, Fq =2.45, F6H 1.65, , P-BAR-HA 1.42, Minimum Si and downflow (Unit 2) barrel-baffle configuration. CASE E_- CD =0.4, 2652 MWt Core Power, F c =2.45, FAH 1.65, P-BAR-HA=1.42. Maximum 51 and cownflow (Unit 2) barrel-baffle configuration, t O B-57

1 APPENDIX B JOSEPH M. FARLEY NUCLEAR PLANT UNITS 1 AND 2 O TABLE 15.4-2 INPUT PARAMETERS USED IN THE LARGE BREAK LOCA ECCS ANALYSIS Barrel-Baffle Configuration < Upflow Downflow (Unit 1) (Unit 2) License Core Power * (MWt) 2652 2652 Peak Li.1 ear Power * (KW/ft) 12.749 12.749 Total Peaking Factor, Fg 2.45 2.45 Axial Peaking Factor, F7 1.485 1.485 Hot Channel Enthalpy Rise Factor, FAH 1.65 1.65 Hot Assembly Average Power, PHA 1.4426 1.42 Power Shape Chupped Cosine Fuel Assembly Array 17 X 17 VANTAGE 5 Intact Accumulator Water Volun: (ft /3 accumulator) 2121 (Min) 2121 (Min) Broken Accumulator Water Volurr 'ft /3 accumulator) 1086 (Nom) 1086 (Noa) 3 Intact Accumulator Tank Volume sit / accumulator) 3001 3001 3 Broken Accumulator Tank Volume (ft / accumulator) 1511 1511' Accumulator Gas Pressure, Minimum (psia) 600 600 Safety injection Pumped Flow (All pumps See Figures 15.4-1

   '(      degraded 8%, HHSI flow imbalance = 10 gpm)                 and 15.4-2A and B Containment Parameters                                            See Table 15.4-4 Initial Loop flow (GPM)                                        86000          86000 Vessel inlet Temperature ('F)                                  540.52         540.52 Ves;el Outlet Temperature ('F)                                 611.28         611.28 Average Reactor Coolant Pressure (psia)                        2310.0         2310.0 Steam Pressure (psia)                                          747.99         747.99 Steam Generator Tube Plugging Level (%)                        20             20 Minimum Refueling Water Storage Tank Temperature ('F)          35.0           35.0 Fuel Backfill Pressure (psig)                                  275            275 Low Pressurizer Pressure Setpoint (psia)

Reactor Trip 1840 1840 Safety injection Signal 1715 1715 Safety injectio? Delay Time (sec) 27 27 i Safety Injection Spilling Containnent Pressure (psig) Blowdown / Refill 20.0 20.0 Reflood 11.5 11.5 Blowdown Containment Pressure (psia) 35.0 35.0 Two percent is added to this power to account for calorimetric error. B-58 l -_ - . , - - .--

1 APPENDlX B JOSEPH M. FARLEY NUCLEAR PLANT UNITS 1 AND 2 1 TABLE 15.4-3 LARGE BREAK LOCA ECCS ANALYSIS SYSTEMS MODELING i Pressurizer Low Pressure Reactor Trip (psia) 1840.0 Pressurizer Low Pressure Safety Injection (psia) 1715.0(a) Containment High Pressure for Safety injection (psia) 22.0 Safety injection Delay (includes signal processing, EDGs start-up, sequencer and pumps to full speed) (sec) 27.0 feedwater Isolation Delay after Reactor Trip (sec) 0.0(b) Steamline Isolation Delay after Reactor Trip (sec) 0.0(b) Number of HHS1 Pumps Operating (Min / Max safeguards) 1/3(C) Number of LHS! Pumps Operating (Min / Max Safeguards) 1/2( ) ; Steam Generator Tube Plugging 20%(d) (a) This setpoint causes actuation of the safety injection at the times shown in Table 15.4-5, for all five cases. (b) Conservative modeling for large break LOCA. . (c) Minimum safeguards assumes one HHS! pump and one RHR pump operating. Maximum safeguards assumes three HHS! pumps and two RHR pumps operating. (d) Uniform 20% steam generator tube plugging assumes 20% steam generator tubes plugged in each steam generator and corresponds to the peak plugging level in any steam generator. This configuration bounds all combinations of non-uniform plugging for LOCA as long as no one steam generator plugging level exceeds 20%. However, it should be noted that the licensed steam generator plugging level is limited to 15% average /20% peak in any one steam generator. B-59

APPENDlX B JASEF" M. FARLEY NUCLEAR PLANT UNITS 1 AND 2 O TABLE 15.4-4 LARGE BREAK LOCA CONTAINMENT DATA

l Net Free Volume 2.300,000 ft 3 Initial Conditions Pressure 14.7 psia Temperature 90.0'F RWST Temperature 35.0*F Service Water Temperature 40.0*F Temperature Outside Containment 20.0*F l Initial Spray Temperature 35.0'F 1 Sorav System Runout Flow for a Spray Pump 2775 gpm Number of Spray Pumps Operating 2 Post-Accident Spray System Initiation Delay 48 see Maximum Spray System Flow 5500 gpm Containment Fan Coolers , Post-Accident Initiation Fan Coolers 27.4 sec Number of Fan Coolers Operating 4 O B-60

     ,,.-v, .v.-s.-.,                      ,   s,    , , , _ , , , __ , , , , , , , , ,   , .

APPENDIX B JOSEPH M. FARLEY NUCLEAR PLANT UNITS 1 AND 2 TABLE 15.4-4 (continued) LARGE BREAK LOCA CONTAINMENT DATA Structural Heat Sinks Wall TAir Area Height T init Thickness ! 3 f*F1 fft 1 f ft1 I'F1 finches 1 < 1 40 75000 150 90 0.25 Carbon Steel / 45 Concrete 2 40 4700 20 90 0.6 Carbon Steel / 45 Concrete 3 90 69800 10 90 9 Concrete 4 90 77000 10 90 0.004 Zinc / 07 Steel 5 90 80500 10 90 0.135 Steel , 6 90 47600 10 90 0.36 Steel i 7 90 23600 10 90 0.7 Steel 8- 90 11800 10 90 2.3 Steel 9 50 13275 10 90 108 Concrete 10 90 7900 10 90 0.25 Stainless Steel / 18 Concrete 11 90 12500 10 90 0.105 Stainless Steel

1. - Containment Wall and Dome

2. - Containment Penetrations, Plates and Liner Stiffeners

3. - Unlined Concrete

4. - Galvanized Carbon Steel

5. - Thin Painted Carbon Steel (< 0.5 Inches)

6. - Painted Steel (< l.0 Inches)

7. - Painted Steel (< 2.0 inches) l 8. - Thick Painted Steel

9. - Floor

10. - Refueling Pool Liner

11. - Unpainted Stainless Steel B-61 l

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

O 'O O JOSEPH M. FARLEY NUCLEAR PLANT UNI 15 1 AND 2 i-TABLE 15.4-5 i LARGE BREAK LOCA ANALYSIS TIME SEQUENCE OF EVENTS i Unit 1 Unit 2 i Case A Case B Case C Case D Case E i CD=0.4 C0 -0.4 CD=0.6 00-0.8 CD=0.4 Min SI Min SI Min SI Min SI Max SI ] Upflow DownfIow Downflow Downf1ow Downf1ow 2652 MWt 2652 MWt 2652 MWt 2652 MWt 2652 MWt Start of LOCA with LOOP (sec) 0.00 0.00 0.00 0.00 0.00 Reactor Trip Setpoint Exceeded (sec) 0.536 0.537 0.528 0.523 0.537 Safety I. 4ction Setpoint Exceedeo (sec) 0.95 0.95 0.79 0.69 0.95 Accumulator Injection Begins (sec) 14.2 14.3 10.9 8.92 14.3 End-of-Bypass (sec) 30.29 30.94 24.23 21.03 30.94 End-of-Blowdown (sec) 30.29 30.94 24.23 21.03 30.94 Pump Injection Begins (sec) 27.95 27.95 27.79 27.69 27.95 Bottom of Core Recovery (sec) 42.40 43.17 36.03 33.16 42.33 Accumulator En.pty (sec) 50.74 51.07 45.82 42.80 51.06

L j o o JOSEPil M. FARLEY NUCLEAR PLANT UNITS 1 Arm 2 o I- TABLE 15.4-6 LARGE BREAK LOCA RESUL13 FUEL CLADDING DATA Unit 1 Unit 2 ] Case A Case B Case C Case D Case E CD=0.4 C0=0.4 CD=0.6 CD=0.8 CD=0.4 i Min SI Min SI Min SI Min SI Max SI ! Upflow Downflow Downflow Downflow Downflow i 2652 MWt 2652 MWt 2652 MWt 2652 MWt 2652 MWt 1 Peak Clad Temperature (OF) 1982.l* 2090.3 1711.0 1629.3 1929.4 ! Peak Clad Temperature - j Location (ft) 7.25 7.25 6.25 6.25 6.25 , Peak Clad Temperature 106.I Iime (sec) 81.5 54.i 52.7 58.9 Local Zr/ll 20 Reaction Maximum (%) 4.96 6.59 1.64 1.07 4.00 Local Zr/H 2O Reaction Location (ft) 6.00 6.00 5.50 7.25 6.00 Total Zr/H 2O Reacttan (%) <l.0 <l.0 <l.0 <l.0 <l.0 i Hot Rod Burst Time (sec) 36.6 38.6 48.6 NA 38.6 l Hot Rod Burst Location (ft) 6.00 6.00 5.50 NA 6.00 Includes the fo110 win 9 effects: Temperature Uncertainty = 8.0 F Containment Mini-Purge Isolation = 3.0 0f Combined SSE and LOCA Events - 6.0"F - Note: A 50 F 0 transition core penalty must be added to tt.e Vantage 5 results until all LOPAR fuel is removed. i

JOSEPH M. FARLEY NUCLEAR PLANT UNITS 1 AND 2 TABLE 15.4-7A Os CASE A - UNIT 1 LARGE BREAK LOCA C D0.4 MINIMUM SAFEGUARDS LOCA REFLOOD MASS AND ENERGY RELEASE RATES Time Mass flow Rate Energy Flow Rate 11gtl (1bm/see) _ (BTV/sec) 42.40 0.0 0.0 43.06 4.4 5741.0 43.60 4.4 5675.0 50.38 319.0 88561.2 65.11 270.0 145688.3 84.31 292.7 144825.1 105.81 301.9 140396.7 129.11 308.8 135373.3 154.11 315.0 130041.8 B-64

 - . +   - . . , . . - ,-                    ,.-.m%w,  .-.-       .
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APPENDIX B I JOSEPH M. FARLEY NUCLEAR PLANT UNITS 1 AND 2 O TABLE 15.4-7B CASE B - UNIT 2 LARGE BREAK LOCA C D-0.4 MINIMUM SAFEGUARDS LOCA REFLOOD MASS AND ENERGY RELEASE RATES Time Mass Flow Rate Energy flow Rate (sec) (Ibm /sec) (BTV/sec) 43.17 0.0 0.0 l 43.82 4.4 5741.5 44.37 4.3 5619.9 l 53.44 38.9 50328.5 i 72.46 49.7 '3536.8 94.61 61.6 821.8 ' 117.41 309.5 139498.0 141.31 333.1 138892.3 4 167.74 353.3 136520.8 s e O l t l l-l 2 s l B-65

APPENDIX B JOSEPH M. FARLEY NUCLEAR PLANT UNITS 1 .: . . 2 TABLE 15.4-8A CASE A - UNIT 1 LARGE BREAK LOCA C =0.4 D MIN! MUM SAFEGUARDS LOCA ACCUMULATOR SPILL RATES I Time Mass flow Rate Energy flow Rate 11111 (lbm/sec) (BTV/sec) 1.0 4261.7 254082.0 2.0 3877.5 231177.6 3.0 3582.1 213564.2 4.0 3346.5 199518.4 5.0 3151.9 187914.3 6.0 2987.3 178105.4 7.0 2844.9 169614.4 8.0 2719.4 162133.9 9.0 2607.7 155469.6 10.0 2507.4 149488.8 11.0 2416.8 144089.5 12.0 2334.5 139181.3 13.0 2259.2 134695.4 (O 14.0 15.0 2190.1 2126.2 130573.7

                                                      -126763.7 16.0                   2066.9                123228.3       '

17.0 2011.6 119932.4 18.0 1959.9 116852.0 19.0 1911.5 113964.4 20.0 1865.9 111243.1 21.0 1823.5 108714.8 22.0 1784.0 106364.4 23.0 1747.1 104159.4 24.0 1712.4 102090.4 25.0 1679.9 100155.8 26.0 1649.2 98323.0 27.0 1620.1 96590.4 O B-66

APPENDIX B JOSEPH M. FARLEY NUCLEAR PLANT UNITS 1 AND 2 TABLE 15.4-8B CASE B - UNIT 2 LARGE BREAK LOCA C D 0.4 MINIMUM SAFEGUARDS LOCA ACCUMULATOR SPILL RATES Time Mass flow Rate Energy Flow Rate 11gg1 (1bm/sec) (BTV/sec) 1.0 4261.8 254088.2 2.0 3878.1 231214.5 3.0 3583.6 213652.6 4.0 3348.3 199620.6 5.0 3154.0 188040.7 6.0 2989.8 178254.2 7.0 2847.5 169765.8 8.0 2722.0 162288.2 9.0 2610.3 155627.2 10.0 2510.0 149645.7 11.0 2419.4 144244.4 12.0 2337.1 139335.7 13.0 2261.8 134849.0 p 14.0 2192.6 130723.0 15.0 2128.7 126911.4 16.0 2069.2 123368.6 17.0 2013.9 120067.3 18.0 1962.1 116979.7 19.0 1913.5 114081.6 20.0 1867.6 111348.4 21.0 1824.8 108793.6 22.0 1785.0 106424.7 23.0 1747.8 104207.2 24.0 1712.9 102123.9 25.0 1680.3 100179.8 26.0 1649.4 98339.2 27.0 1620.2 96597.2 I B-67

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m 400 -r i G l-d w E d 100 I h a o 1000 1500 2000 PRES $UME (pstA)

1 HHS! and 1 RHR Pump Operating i

I SAFETY INJECTION FLOW l' A N'H E PN AATE VERSUS ACS PRESSURE Q Alabama Power g wucuan puur MINInun SAFEGUARDS-FIGURE 15.4-2A

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CORE PRES 5URE UNIT 2 DECLG (CD = 0.4) Alabama Power Q wucuun ewn UNIT 1 AND UNtT 2 FIGURE 15.4-38

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TW ITCl , l I 1 CORE PRESSURE JOSLPH M. FARLEY LMIT 2 DECLG (CD = 0.6) Alabama Power Q NUCLEAR PLANT UNR 1 AND UNM 2 FIGURE 15.4-3C 1

.           . - - -             . . - . - .        . . . - . . . . _       .      .. .-.                                   _      . .        _ = - - - . . . - . ~ . . ~ . - . - . . _ . . . - . - - -                ..
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i t CORE PRESSURE JOSEPH E FARLEY UNIT 2 DECLG (CD=0.4) Alabam'a Power NUCLEAR PLAm MAXIMUM SAFEGUARDS FIGURE 15,4-3E s** y vt,<<-r c m. w,-,.,a..-,9 r Wt v et e 5m W ,ti-* w r= t +w +- aw e-w WW++ F e e-vS - e w-n av vv n w--wm-m+em. w & - Sv %- ~ e - es we w -W h w+ xw~e+~w+t-wwe- rm-+-+weMt-- * - ter +srv+7--+-M-+d

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FIGURE 15.4-4A

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f 1. I-I f JospH W. 7Amt.gy BREAK FLoldRATE A UNIT 2 DECLS (CD = 0.4) Alabama Power g Nuc m w UNIT 1 AND UNFT 2 FIGURE 15.4-48 l

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ML20079D610

Contents

Text

SUMMARY

1.0 INTRODUCTION

SUMMARY

7.0 REFERENCES

1.0 INTRODUCTION

SUMMARY

SUMMARY

7.0 REFERENCES

Subject:

SUMMARY

SUMMARY

3.7 REFERENCES

4.7 REFERENCES

SUMMARY

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ML20079D610

Text

SUMMARY

1.0 INTRODUCTION

SUMMARY

7.0 REFERENCES

1.0 INTRODUCTION

SUMMARY

SUMMARY

7.0 REFERENCES

Subject:

SUMMARY

SUMMARY

3.7 REFERENCES

4.7 REFERENCES

SUMMARY

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