ML19257D365

Forwards Supplemental Matl in Support of 790717 Application for Amend to License DPR-40 to Permit Cycle 6 Operation Following Core Reload.Proposed Tech Specs & Response to NRC 791030 Setpoint Methodology Questions Encl

ML19257D365
Person / Time
Site:	Fort Calhoun
Issue date:	01/28/1980
From:	William Jones OMAHA PUBLIC POWER DISTRICT
To:	Reid R Office of Nuclear Reactor Regulation
Shared Package
ML19257D366	List: ML19257D365 ML19257D369
References
NUDOCS 8002040215
Download: ML19257D365 (69)

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. t g Omaha Public Power District 1623 HARNEY a OMAHA. NEBRASKA 68102 a TELEPHONE 536 4000 A7EA CODE 402 January 28, 1980 Director of Iluelear Reactor Regulation ATTII:

Mr. Robert W. Reid, Chief Operating Reactors Branch Ilo. 4 U. S. Iluelear Regulatory Commission Washington, D. C.

20555

Reference:

Docket Ilo. 50-285 Gentlemen:

Omaha Public Power District hereby submits forty (40) copies of supplemental material in support of (1) the Application for Amendment of Operating License (" Stretch Application"), filed July 17, 1979, which seeks to amend Facility Operating License Ilo. DPR h0 to permit Cycle 6 operation following core reload at an increased power level of 1500 MWt, and (2) the Application for Amendment of Operating License (" Reload Application), filed July 17, 1979, whJch seeks to permit Cycle 6 operation following core reload. Forty (h0) copies of the following materials are enclosed:

(1) Revised Startup Physics Testing summary.

(2) Responses to IIRC setpoint methodology questions received October 30, 1979 (3) Proposed Technical Specifications addressing RCS heatup and cooldown pressure / temperature limitations.

(h) Discussion supporting item (3), proposed Technical Speci-fications.

Should you desire additional information on these materials, please advise us.

Sincerely, i,'fi\\

I;u Q..',,

g,t; a

W. C. Jones Division Manager Production Operations WCJ/EJI!/BJII:Jmmh Enclosures cc: LeBoeuf, Lamb, Leiby & MacRae Mr. Peter B. Erickson (IIRC) 8002040 2#

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n +. 32 Effective Full Power fearc (lF' Y) if opera.icn. The resulting hea.up and ecclacwn curvec have been adju sted to acec."2t for the total ch1ft of the Reference Li1 ductility Trarci-tion Temperature (ET g ) of 2230F. The thift of 223CF has bren added to tho 'r: celine, ncn-irr i'liate'i r.n ial curvec to rec 11t in the new corrected curvec to be ured thrcach th*= end of fuel 'ycle L<. n +o .t.,... +4.,_, a

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4 n.. c,,o,, n. rv-1-m. w a ,.1 m,, w + me .a 1.+e_ m .m 0. ~ to exte m the curvec for cperaticr. bo nd Cyclc 6. The Low Temperature O'.erpreocure Fr tecticn (LTOP) cyc.en is decirnea to "revent the.nrir.ar.v a.vaten.creccure from exceed inr the i o precbure-terperature linite (Tecnnical Greci: ' cation Ficure s P-1A and 2-1B) in the event of cn inadvertent tacc or energy addition. LTOF cycte actuation cetpcints, as well ac terperatures fc r dis-ablir ni;h Pressure Cafet;. Injecticn (HPSI) pumpc, will b-deter-t tir ca accu-ing failure of one of the two POEV'c. Calculat.cns will be based upon a POEV diccharge coefficient of.k5, wh ch a et prehensive tectinc procran has chcwn to be ecnservative for cubcoolea liquidc. Inadvertt.t actuation of three (3) HFOI pu pc and th ce (3) char,-ing purpc, coincident with the opening of cne of tho two POEV'c, wculd result in a pea'. primarj cysten precoure of 1190 psia. 1190 psia correcpondo with a minitr. pe: iccible temperature of 320Ci er Ficure 2-13. Tr.uc, n' leant cne ' FBI prp is disabled at 320CF. 868 309 2.....L.

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1 s D-1 SECTIOII D 0:UdiA PUBLIC POER DISTRICT FORT CALHOU:I STATIO:; STARTUP PHYSICS TESTI:iG 1868 311

D-2 1. Introduction The principal tests in the proposed Cycle 6 startup physics test progran are listed below. These tests are sufficient to show that the as-loaded core's parameters are within the bounds of the safety analysis, thus permitting continued safe operation. Acceptance and review criteria, along with the action to be taken if these criteria are not met, are also discussed. The predicted values of the parameters being measured will not be calculated until after the end of Cycle 5 when the actual fuel exposures are known. 2. Startun Tests 2.1 Hot Functional Tests Prior to the approach to the initial criticality of Cycle 6, normal surveillance testing and operating procedures vill'oe completed. During this sequence, applicable surveillance tests are perforned to check CEA position indication and all other interlock and control features of the rod d-ive system. 2.2 Initial Criticality and Low Power Physics Tests Following the dilution to initial criticality, the following reactivity parameters will be measured at less than 10-15 of rated power: 2.2.1 Critical Boron Concentration 2.2.1.1 Hot Zero Power, Partial Insertion of Group 4 2.2.1.2 Hot Zero Power, All Rods Out 2.2.2 Isothe2ral Temperature Coefficient - HZP, Nor.inal ARO 2.2.3 CEA Group Worths 2.2.3.1 Individual Bank Worths of Regulating Groups 4, 3, 2, and 1 - HZP

2. 2. 3. 2 Sequential (overlapping) Worth of Regulating Groups 1, 2, 3, and 1-2.2.h CEA Symmetry Checks - HZP 1868 312

D-3 2.3 Power Accension Testa 2.3.1 Tecta Performed at a nominal 50% of Rated Power Following the acceptable comparison of measured and predicted reactivity parameterc, reactor power will be increased to a nominal 50% of rated power and a power distribution verification performed. This verification will be performed in a non-equilibrium xenon state with measurement of the following parameters. 2.3.1.1 Total Unrodded Planar Radial Peaking T Factor (Fxy ) 2.3.1.2 Total Integrated Radial Peaking Factor (FR) 2.3.1.3 Azimuthal Power Tilt, Incore Detectors 2.3.2 Tests Performed at a Nominal 70% of Rated Power A power distribution verification will be performed at this power level after equilibrum xenon has been established with measurements of the same parameters as in Section 2.3.1. 2.3.3 Tests at nominal 100% of Rated Power Following completion of testing at the lower at-power levels, power will be increased to a nominal 100% rated thermal power at a rate commencurate with fuel performance guidelines. After the ectablishment of equilibrium xenon, the following parameters will be teacured. 2.3.3.1 Icothermal Temperature Coefficient 2.3.3.2 Power Coefficient 2.3.3.3 Critical Boron Concentration, ARO 2.3.3.h Total Unrodded Planar Radial Peaking T Factor (Fxy ), ARO P.3.3.5 Total Integrated Radial Peaking Factor (FR ), ARO 2.3.3.6 Azimuthal Power Tilt, Incore Detectora, ARO 1868 313

D-h 3 Acceptance - Review Criteria Acceptance criteria are applied to the test results, after concervatively adding measurement uncertainty to the measured value, to incure that the core conforms to the physics decign and that plant response to transients is in accordance with the safety analysis. Review criteria are also applied to hichlight any lesser deviation which may indicate that the core was incorrectly loaded or to confirm that the assumptions used in the design analyaes are valid. Acceptance and review criteria for the low power physics parameters measured are listed below. Parameter _ Acceptance Criteria Review Criteria Rod Drop Technical Specifications Previous values Critical Boron j; 90 ppm of predicted f; 50 ppm of predicted Concentration Isothermal Technical Specification + 0.3 x 10~ ap/ F Temperature limits

• Moderator Coefficient Temperaw e Coefficient CEA Group Worths + 15% of predicted f_ 15% of predicted Total Regulating ~10% of predicted to f; 10% of predicted CEA Group ensure adequate shut-Worth down margin CEA Syr. metry Hone The greater of: 1.5d Checks dcviation frca group average or 15% devia-tion from group average The review criteria for the CEA symmetry checks will be 1,1 5? or 15% deviation from the group average, whichever is greater.

The reason that two types of criteria are stated 13 th at for high worth roda, a percent deviation is appropriate as is applied to other rod worth measuremento. For small worth reds, however, an abcolute deviation is required which in the same type of allowance as specified for rnactivity coefficient meacurements. This is not intended to be a go-no-go criterion, but rather an indication of the degree of tilt that might cauce the Acimuthal Power Tilt specification to be exceeded. If the criterion is not met during the t'at, the tent program will be extended to reconfirm the r.ieasured value. If the values still fall outside the stated criteria, the reculta vill be reviewed to determine the potential irpact upon plant operationc. 1868 314

D-S The acceptance criteria for power distribution verifications are the limita cited in the Technical Specifications. The review criterion for the comparison of the predicted and measured full core power distributions of the instrumented accedbliec is a 55 standard deviation. Acceptance and review criteria for the at-power critical boron concentration measurements are the came ac for low power phycica tecting. Acceptance criteria for the nominal 100% power Isothermal Temperature Coefficient (ITC) and Power Coefficient (PC) chall be the Technical Specification limits on the Moderctor Temperature Coefficient (MTC). h. Action and Review Planc The following plan of action la provided if a measured parameter differs from the predicted value by more than the acceptance criteria. h.1 The physica tect program will be extended to reconfirm the measured value. If the total regulating CEA group worth ic less than 10% of the predicted value, shutdown bank worth measurementa vill be made. g7g h.2 The predicted value will be reviewed to ensure that it accurately reflects the particular plant conditionc under which the teacurement wac made and refined if appropriate. 4.3 If, after the above two stepc, the dicagreement persists, the safety analycic will be reviewed to determine whether the measured value of the particular parameter in question, when combineu with al] of the other cafety related para-meterc, increacec the severity or consequencea of accidents or anticipated operational occurrencec. If equivalent cafety for the plant can be demonstrated, the test results will be deemed acceptable. h.h The other phycica related cafety parameterc will be verified to be within acceptable limits by additional teacurements if necescary. h.S If the combination of cafety parametera determined above fall outside of range of cafety parameterc used to cupport the proposed operation of the plant, the plant operating limito will be ad, justed to prevent conditionc which could result in exceeding the cpecified acceptable fuel decign limitc. 1868 315

D-6 The Plant Review Committee and Technical Services vill review the results of the low power physics tests and ensure that the acceptance criteria are met prior to allowing escalation above five percent of rated thermal power. The at-power testing results will be reviewed prior to reaching 100% power. If after review of the data it is determined that a Technical Specification limit has been exceeded, then appropriate action as required by Technical Specifications vill be taken. Results of startup testing vill be submitted to the NRC within 90 days following completion of the tests. This report will summarise the test results and include a comparison of the measured and predicted values of low power physics parameters and a full core power distribution comparison including deviations between the measured and predicted relative power densities of operable instrumented assemblies. If the difference between the measured and predicted values exceed the acceptance and/or review criteria, the report vill discuss the actions that were taken and also justify the adequacy of these actions. i868 516

OMAHA FUBLIC POWER DISTRICT FORT CALHOUII STATIO:I DOC:GT !!O.,^-285 RESPO:ISE TO IIRC SETPOIIIT METHODOLOGY QUESTIO:IS 1868 317 QUESTICN 1 List all operational mancuvera or conditions considered in generating axial pouct chapca. Also lict the control rod configuration and burnup. Describe hJu the XCHon oGaillationG ucre induced. Alco dcccribe chather the plant was base loaded or in a load follou configura-tian. Girc the pcuer history ace:cncd. Thie inferna !cn ma.n bc l'rct>idcd rm t efficient.ly in matri: form. t Justify that only theca maneuvero necd be considered for gancrating ca Tal pouce charca. For Ft. Calhoun Cycle 6, hou many axial power chapes acre generated and hou many ucre uced in the cet point analysia?

RESPONSE

The Fort Calhoun Cycle 6 core was modeled in three-dimensions with the computer code XTG(I) and depleted in a base load (1500 MWt) all rods out confi gura tion. This reference power history and rod configuration for Cycle 6 was chosen based on the power and control rod histories of previous cycles (Cycles 3, 4, and 5) and OPPD's anticipated operating requirements for the Fort Calhoun Nuclear Power Plant throughout Cycle 6. (The plant is currently th finishing up the 5 cycle.) A bar chart showing the actual plant power and control rod configuration by month for Cycles 3, 4, and 5 and the anticipated power and rod histories assumed in the analysis for the determination of Cycle 6 setpoints are shown in Figures 1.1 and 1.2. Three different burnups, from the three-dimensional XTG Cycle 6 core depletions, were chosen as the base cases for the starting point of more than 1868 318 1500 axial power profiles generated and used in the determination of Cycle,6 setpoints. The burnups chosen were O MWD /MT (equilibrium xenon), 6000 MWD /MT, ' and 10,330 i1WD/MT. The burnups correspond approximately to the beginning, middle and end of the Cycle 6 exposure range. At each of the above mentioned exposure points many possible Cycle 6 axial power profiles were calculated by inducing xenon oscillations. The xenon oscillations were incited by inserting the Control Element Assemblies (control rods or CEA's) to the Power Dependent Insertion Limits (PDIL) at 100% power for 8 hours and then ir.stantly removing the CEA's. In order to analyze the possible axial power distributions within a wide range of Axial Shape Index (ASI), severe xenon oscillations were created. The oscillations resulted in axial power distributions covering a range of A;;ial Shape Index (ASI) units for values ranging from -60% to +60% offset. A negative ASI is top peaked in the core. The normal axial shape index for the based loaded Fort Calhoun core covers the range from -3% to +2% ASI units. Figure 1.3 shows the anticipated ARO steady state axial shape index units calculated for Cycle 6. Control rod effects on the axial power profiles were also determined. At selected points throughout the xenon oscillation control element assemblies were inserted to the PDIL limits in 10% power increments. At each point selected the xenon distribution was fixed allowing the axials to change depending on rod movement only. These maneuvers gave the many possible axial power profiles at partial power and CEA insertion. 1868 319

.i Fort Calhoun Measured and Anticipated Core Power by Month 1800 + l -l - - - - -~- - --i- - l i l, .,p., l l I 1600 ~ l i " T-I l. I ~ 1400 bl i l l I i l 1 i l 1200 - -,--l--- l l t I' I l i i .._j .i. -l----- l- ----f. - -F--- -i 1000 I . !. i-. I t. i l - l j 1 800 -l-j- - ;- - - - ' - -i- -p i g e i 3 1 l l 1 f 600 --l-j-- I' 1 .a _. I l l 400 J - -- +- -- - ;- 200 i,-- l i l i. g I i. $$i NL ffh*O'b h?hj $.f$$ h I I ' 1.917,_

1978,

19_79 l 1980 _,

~ Month Figure 1.1 9 ~ 1868 320

_L_ Fort Calhoun 4 P.easured and Anticipated Control Element Assembly Position by Month i 1 i. I i I i i i 180 m e r1 e i .c i i ]gQ I i l i i. c 140 .L - - -. - - -----__{-_----.'- 1 o i i 1 l u - l - - -- 120 j- ]1-7- m U 100 -l4 -i, - -i--- l j '. o e i r ---f-- I-- 80 l 2 i j .i. I p l t i c 60 1 - s- - + - - a E .I I. I .I. o I l-l G 40 t o p .y_ i i 5-I i E l i u 0 .x s4 s.x o. x A, ss $ ?E $3l?$s:f:f[$ S 3EE $h$ W $g$& S f i l l 1977 1978

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I Figure 1.2 i 1868 321 i Fort Calhoun Cycle 6 Axial Shape Index vs. Exposure ARO, 1,500 MWt i +2 ~ +1 3 te 5y 0 E .a , i; m Q a _1 i -2 ~ 0 1 2 3 4 '5 6 7 8 9 10 Exposure GWD/MT Figure 1.3 4 l868 322 QUESTION 2 For the crial power shpaca in Qucation 1, describe the analytical methods used to calculate these shapaa. Provide the cpat*al noding schemes for each computer calculation and the size of the tima otcpa uccd. Dccaribe hou reactivity facdback cffecto acre included. Hou many cncrgy groups are used in thcac calculations? sustify thic n:cher of energy groups.

RESPONSE

Analytical methods used to calculate the axial power shapes adhered to in Question 1 were determined using the Exxon Nuclear Neutronic design methods for PWR's (NRC approved) described in References 2, 3, and 4 and Section 6.1 of Reference 5. Specifically the methods used to calculate the axial power shapes for Cycle 6 included the computer codes XPOSE,(6) PDQ,(7M) and XTG(1) The computer code XPOSE, a modified version of the industry accepted LEOPARD code, was used to generate fast and thermal spectra and cross sections in two energy groups for the reactor simulator codes, XTG and PDQ7. Detailed two-dimensional pin-by-pin radial power distributions for the core were determined with PDQ7. This information was used in conjunction with the XTG results to determine values of F for the core. The reactor simulator code XTG was used to determine the wide range of axial power profiles referenced in Question 1. The core was modeled and depleted in three-dimensions with the XTG code. The 3-D XTG model was used to produce reference axial power profiles and core average cross sections and exposures as a function of core height. These core average cross sections and exposures were then usert in the one,. dimensional XTG core model to determine 1868 323 all possible core average axial power distributions for the Cycle 6 Fort Calhoun core. The one-dimensional XTG model, consisting of 1 radial node and 24 radial nodes, was normalized to the three-dimensional XTG model con-taining 4 radial and 12 axial nodes or 48 nodes per assembly. In order to sustain the xenon osci.llation at the beginning, middle and end of Cycle 6, reactivity feedback effects were varied in the one-dimensional XTG model. At the beginning of the cycle the reactivity feedbacks effects due to Doppler and Moderator density were removed from the calculation. A sustaining and slightly divergent xenon oscillation was not possible with these feedbacks in the calculation. The feedback effects were used in the middle and end of cycle axial power shape calculations. Without the effects of the feedback included in the calculation at the end of cycle, the xenon oscillation and corresponding power distributions became very divergent and unrealistic. One hour time steps were used throughout all the xenon oscillation calculations. 1868 324 QUESTIOll 3 Deccribe in more detail the calculation of F donc to datemina diccucced in Section 4.1.1 Xll-i!F-507. Daccribe the PDQ model and shoo hcu (and chich peaking factora frcm PDQ arc uced in XTG to datomina F Alco, shou the XTG 3D modeling. Into hou many incrementa arc the CEA incertionc divided for the variouc F calculationc? Hou r:any encrjy groupt are uccd in theca calculationc? Juctify using this nw-bar of ene.gy groups. RE5r0NSE 2 The calculation of the ratio of the power of the peak fuel pin to the average fuel pin in the core at height Z is commonly referred to as F and is explicitly calculated in the three-dimensional XTG core simulator code referenced in Questions 1 and 2. To determine F , XTG calculates a three-dimensional power distribution for the core. The typical core power distribution consists of powers in 12 axial planes with four radial nodes per assembly in each plane. XTG calculates an F in e ch of the 12 planes by determining the xy peak nodal power times the local peaking factor (F in each plane, the average g nodal power in each plane, and divides the peak by the average to determine the planar F The core F is defined as the maximum of the planar F xy. xy xy values. Values for the local peaking factors, F re directly input into the L XTG calculation by assembly or by fuel type. The local peaking factors are explicitly calculated in the quarter core PDQ model. The PDQ7 core simulator is modeled to perform detailed two-dimensional radial calculations. The core is modeled in PDQ7 on a pin-cell basis; i.e. one 1868 325

-9_ mesh block per fuel cell. Each pin-cell has the appropriate nuclide concentra-tion of the burnup history for that pin. The Fort Calhoun PDQ7 model is similar to the model described in Reference 2. Output from the PDQ/ HARMONY calculations include the pin-by-pin radial power distributions, F. The local r pin power peaking, F, from PDQ7 is used as. input in XTG for the XTG F and r F calculation. Values of F are derived from PDQ7 results for each assembly x by dividing the average assembly power peaking into the peak pin power in that assembly. Values for F are determined throughout Cycle 6 at full power, with the xy CEA's withdrawn from the core At part power the F values were determined with the CEA's at the transient power dependent insertion limits (PDIL). The power distribution calculations were made in increments of 10; power. The CEA's were inserted to the maximum allowable transient PDIL limits at that power. All the calculations done witn the core simulator model are made in two energy groups. Methods are discussed in Reference 2, 3, and 4. O 1868 326 QUESTI0il 4 Discuco to uhat c tent the projected (utility planncd) pcuer hictory is included in the cet point analycca (e.g., dcplation calculationc.'. Diccuco the effect of a difference betucen pro,jected poucr history and actual power history cn the cet pointa. If thic effcat is concidered, dcmonstrate with ccncitivity ctudica for variouc control rod mancuverc and changcc in baration that poucr his tor;t effcetc arc adequatcl o conndcrcd. i Diccuco the typcc of poucr mancuvers accumed (e.g., basa loaded, 200-50-100, etc.).

RESPONSE

Operating projections by 0 PPD for the Fort Calhoun Nuclear Power Plant in Cycle 6 show operation of the plant at 1500 MWt at a 90% capacity factor for 320 days. This operating projection was used in the analysis determining set points. The answer to Question 1 shows the power and CEA configuration assumed for Cycle 6. The axial power shapes calculated for Cycle 6 and used in the set point analysis bound all possible axial power shapes for Cycle - In addition, sensitivity studies were made to determine axial power so es possible for operation of the core at power levels below 100% as well as at CEA insertion to the loag term PDIL limits. The Technical Specifications preclude operation of the core for extended periods of time with CEA's inserted past the long term PDIL limits. In all cases the calculated axial shapes or the core peaking, F, versus ASI were within the bounds of the setpoint calculations. Control q rod histories assumed in the sensitivity studies included CEA insertions +,o 121 inches and 96 inches for the entire cycle. Power levels included in the sensitivity studies included 50% and 75% power! 186B kn7 QUESTICll 5 Discuco the initiali::ation of conditions for the beginning of a cycle for act point calculatienc.

RESPONSE

Both core simulator models used in the Fort Calhoun set point analysis, XTG and PCQ7, were verified against measured data from Cycles 1 through 5. Both simulator codes were depleted through all previous cycles; i.e. Cycle 1, 2, 3, 4, and 5. In each cycle, comparisons between measured and calculated data were made. An assembly power comparison between measured data (CECORE) and calculated data (3-D XTG, and PDQ7) for Cycle 4 at 4000 MWD /MT is shown in Figure 5.1. Agreement between the three power distributions is good. Initial conditions for the Cycle 6 setpoint work were based on the cycle depletions from the previous cycle depletions of Fort Calhoun, namely Cycles 1 through 5. Such cycle depletions were made with both referenced simulator codes, XTG in three dimensions and PDQ7. For Cycle 6 the burnup history of the fuel was explicitly accounted for. G 1868 328

Fort Calhoun Assembly Power Distribution Comparison of Measured to Calculated Cycle 4, 4,000 MWD /MT, HFP .856 1.050 1.209 .977 1.184 1.184 .874 1.053 1.225 .977 1.190 1.128 .851 1.038 1.233 .980 1.202 1.160 ~ .886 1.042 .857 1.131 1.101 .984 .865 1.043 4'e 5 1.105 1.076 .966 .704 1.043 .865 1.115 1.093 .990 .711 1.264 1.116 1.178 1.081 .703 1.264 1.091 1.176 1.119 1.272 1.091 1.195 1.090 1.241 .845 .946 1.179 .851 .949 1.202 .847 .938 1.123 .589 + CECORE (Measured) 1.104 .620 + 3D XTG (ENC) 1.123 .614 + PDQ7 (ENC) Figure 5.1 1868 329 QUESTIO1 6 (c) in:at critical hcat flux correlation is uced by Exxon in calculating the SAFDL on it!BR? (b) Deceribe hou limitations in the range of the correlation (e.g., the 25% quality limitation of the W-3 correlaticn) arc accomodated in thic SAFDL. (c) Provide the cpecific critaria uced to prcycnt c:cccding flou atability ~ limita and provide the ductification for thace critaria. (d) The fuct molting limit is given on Page 10 of Xll-1lF-507 ao 21 ku/ft. Io thic ntenbcr the result of an E::an calculation? Juctify the use of thic n:c:bcr for tuo different fuel designc (Combuction Enginccring and Exxon). In:at fuct molting tcmperatura cac aco:e:cd for thic calculation?

Response

The W-3 burnout heat flux correlation with correction factors for the presence of a cold wall and nonuniform axial heat flux was used to establish the speci-fied acceptable fuel design limit (SAFDL) on the fuel burnout performance ( D::B R). The application and interpretation of the W-3 correlation for deter-mination of the Fort Calhoun reactor set points is consistent with the Ei!C predictive models for DriBR (Xti-75-48). The SAFDL on Df;BR is protected by limiting the operating values of core power, coolant inlet temperature, and system pressure to the most conservative of the following: that set of operating values which results in MDT:BR = 1.3. o that set which gives rise to parameter values exceeding the range of o the W-3 correlation. Thus, the W-3 correlation is never used with parameters outside its acceptable range, and the limits on its range are implicitly included in the SAFDL. 1 8 6 8 N

- 1!+ - The specific criteria to preclude potential flow instability are: (1) a calculated subchannel quality of;less than or equal or 15%, and (2) a core average exit quality less than or; equal to zero. Adherence to these criteria is ensured by the SAFDL on D!iBR and.the core saturation limits. In addition, other reactor protection system limits:such as the low pressure trip and variable overpower trip preclude reactor operation in a potential flow instability mode. The EllC fuel melting limit of.21 kw/ft was calculated via the GAPEX code (Xi!-73-25) in conformance with the.USf1RC approved Et1C fuel densification model for PWR fuels (USfiRC report dated', February 27,1975). The value of 21 kw/ft as an LHGR limit for the existing: fuel is documented in the Fort Calhoun Technical Specifications. Thus e:the value of 21 kw/f t as a maximum fuel rod LHGR for both fuel types is judged to be acceptable. A fuel melting temperature.of 2790 C was assumed at beginning-of-life,- and this value was decreased at the rate of 32 C per 10,000 fGD/MTV and com-pared against the fuel temperatures calculated at 21 kw/ft throughout fuel lifetime. 18683Il Questien 7 Provide a ecmplete deceription of the determination of the shape annealing factor, SAF, for Fort Calhoun. Is the determination of this factor consistent with Exxon calculational methodo? Responce The shape annealing factor (SAF) is defined as the ratio of the internal axial chape (ASI) to the external axial shape index. It can be exprecced as: SAF = ASI (internal) = ASI (IUCA) ASI (external) ASI (excore) where ASI = L-U, L is the power in the lower half core and U is b* the power in the upper half core. The SAF of the Fort Calhoun core van determined by the induction of an axial occillation transient during initial ctartup testing. The axial oscillations were induced by inserting the control rod from ARO conditicnc to approximately the reactor mid-plane.and then returning the control rod to the ARO position. The oscillationc were charactericed by the axial chape index. During the occillation ASI (IUCA) data frcm the incore IUCA code and ASI (excore) data of the excore detectors were monitored every four hours. The ASI data from IUCA and the excores were plotted and a linear least squares fit analysic was performed. The slope of the fitted curve yielded a SAF velue of 2.86. Thic SAF value has been used in the axial power dictribution ( APD) calculator cince Cycle 1. A subcequent measuremento of the SAF per-formed at mid-Cycle 1 and Cycle 2 confirmed that the SAF value used for the Fort Calhoun Station maintains a degree of conservatica. The SAF is a function of raterial and geometry between the core and the detectors and the detectors thencelves. It is independent of fuel type. The existing SAF and its original determination is consistent with Exxon calculational methods. 1868 332 QUESTIO:: 8 Provide thc E==on definition of rod ahadouing. Deactibc in dataii the calcu-lationc donc to determine the adductment to the core averaac crial chapc index to account for rod chadouing effectc. Gioc the results for Ft. Calhoun Cycle G. Will thcca resulta c1'ange frca cycle to cycle? Ccmbuction Engincering performa rod chad:uing calculations using neutron trangart thcary (CE:,TD 1H, Sccticn .1.1. -l.1). It g cava that E. aran rclica on diffucion thacry. Justify thic difference. Alco, c plain hou uncertaintica in thic calculation are taken into account. Show the calculational modela uced to perform thace calculationc. Diccuco tha effecta of trancients which change CEA pocition on rod chadouing factors. Hou is thia cffect included in the act point calculations?

RESPONSE

Effects of rod shadowing on the response of the Ft. Calhoun excore detectors are explicitly accounted for in the determination of set points. A conservative value of +.02 Axial Shape Index units ( ASI) is removed from the operating margin in the final calculation of all Limiting Safety System Settings and Limiting Conditions of operation requiring excore detector response as input. This value is presented as an uncertainty in Section 4.1.3 of Reference 5. Rod shadowing is the effect of control rods distorting the flux in the peripheral assemblies which are the primary contributor to the signal seen by the excore detectors. The calculations determining set points use the core average axial offset or axial shape index (ASI). Therefore, +.he ASI measured with the excore detectors which are inputs to the set points must be adjusted for rod shadowing effects. This adjustment of the excore detectors response was determined to 've less than 10.02 A5i ' units-far C)cle 6. 1868 333 Calculational methods and models (codes) used to determine the effects of rod shadowing on the excore detectors is the transport theory code XSDRtlPM,(9) PDQ7, and 3D-XTG. XSDRNPM is a one-dimensional code used to determine the attention of signal or flux encountered when traversing the water and steel region exterior to the outer assemblies. PDQ7 is used to determine the atten-tion in signal through adjacent assemblies. The 3D-XTG code was used to calculate the ASI in all assemblies and the core average. By using the above calculations to determine the relative effect each assembly has on the excore detector response, and combining the actual effect with the ASI of each assembly, the excore response of ASI is dctermined. Comparing the calculated excore detector response to the calculated core average will give the adjustment in excore detector response required for Cycle 6. This value was calculated to be less than +0.02 ASI units for Cycle 6. Each cycle will require the determination of excore detectnr response due to rod shadowing. The conservative adjustment factor of +0.02 ASI units will be adjusted accordingly on a cycle by cycle basis. Transients affecting the CEA positions will not increase the rod shadowing adjustment factor above a value of 10.02 ASI units. This value was used for an all rods out core configurations as well as for rods in core configurations. 1868 334 Question 9 (a) For the uncertainty values listed on page 15 of XN-NF-507, justify the value listed. List any experiments used to obtain these values and give a full descriptien of the calculations done to obtain these uncertainties frca experimental results. Where values from previous cycles of operation are used, justify that these are appropriate for Exxon calculational methods. If any of these uncertainties are composed of several components, list all com-ponents and justify all the component values. (b) For the trip overshoot, is this based on a transient analysic? Describe the analysis methods used to derive this value. (c) For erch uncertainty listed, give a statistical statement character-izing the confidence in this value. (d) For each uncertainty, show, in detail, how it is included in the setpoint analysis. Respanne (a) and (c) The justification for the uncertainty values listed are given below: (1) Physics calculation measurement uncertainty Peak LHGR 7% F 6% R These uncertainties are applicable to the Technical Speci-fication limits on F T and F T xy R and are effectively applied to both the limit and the measured value of these parameters. (Since the same uncertainty would be applied to both the limit and the measured value, the comparison of the two cancels out the uncertainty term. Therefore it is not applied to the Technical Specifications. ) The values of F i T and F have xy beenandvillbemeasureddurigCycle6usingtheCECORcode n and the methodology described in CENPD-lh5 The 7% uncertainty on peak LHGR and 6% uncertainty on FR vere accepted by NRC for CE reactors as interin uncertainties at the October 2, 1978, meeting between URC and CE. These uncertainties were further confirmed at the March 6, 1979, CE/NRC meeting. Presently, the District is participating in the CE power distrioution uncertainty progrma to justify lower uncertainties. The revised topical report, CENPD-153, is being prepared for submittal. The compcnents and statistical confidence of the values is explained in drafts of this report previously given to NRC. (2) Aximuthal tilt allowance, 3% This allowaace is the current LCO contained in the Fort Calhoun Technical Specificaticnc. As a LCO, it reprecents a limit on tilt. Operatien data has confirmed the suitability of this limit. 1868 335 ~ (3) Engineering tolerance uncerta1nty, 3% The engineering tolerance uncertainty is included in the analysis because manufacturing tolerances in the specificaticn of pellet density, pellet diameter, and pellet enrichment could produce an additional rod surface heat flux at a local hot spot. The engineering heat flux factor is determined from the following characteristic cubfactora derived from the manufacturing tolerances: Subfactor Pellet Density, TD 94.011.5 1.0160 Pellet Enrichment, v/o 3.510.05 1.01h3 Pellet Diameter, in. 0.370010.0005 1.001h Clad Diancter, in. 0.bh2+0.0035 1.0080 -0.0015 Statistical Total 1.0229 It was accured that the finished fuel pellet in the hot spot will deviate from nominal values by the specified allowance. The pellet density uncertainty to a 20 variation while all other values represent an absolute limit. The resultant individual cubractors, when statistically summed, yield a tolerance uncertainty of 1.0229 Fuel denairication results in the shortening of the heated length which causes an increase in the linear heat generation rate. In-reactor shortening of the active fuel column length vac concervatively evt tated from the following expression: AL = 0.965 - pi L 2 where: AL = decrease in fuel column length L = fuel column length pi = initial mean pellet density Baced on a mean pellet density of 9h percent TD and a fuel column length of 128 inchec, the column shrinkage was evalu-ated to be 1.6 inches which yields an increase in local heat flux of 1.25 percent at conotant core power. Compensating for the fuel column length decreace due to densification is the increace in fuel column length resulting from thermal expan-sion. This length incr2 ace is approximately 1.35 inches, which reducec the local heat flux by 1.06 percent. Thuc, the net increase in local heat flux due to axial fuel column shortening is partially compensated for by the increase in fuel column length due to thermal expancien. The net 0.2 percent increace in heat flux is accour.._. three percent engineering tolcrance uncertainty. 1868 336 (4) Power measurement uncertainty 2% of rated for LCO 5% of rated for LSSS These uncertainties are a function of instrument and USSS design and are independent of fuel type and reload cycles. These values have been used for a]1 For Calhoun cores and are given in CEUPD-199 (5) Trip overshoot Discussed below. (6) Physics uncertainty in predicting CEA +.02 ASI distribution effect on excore detectors This uncertainty is addressed in the answer to Question 8. (7) Physics ancertainty in applying shape 0.01 ASI annealing correction axial shapa index limits The shape annealing factor is extremely difficult to predict analytically because of the rather complex neutron scattering involved, it is determined experimentally as discussed in the responsc to Question 7 The procedure used involved the fitting of a linear relationship between the shape index determined from excore detecter signals and the core average axial shape index. Analysis of the data including consider-ation of data scatter and excore detector level calibration indicate that an uncertainty in the analysis of the shape annealing factor equivalent to 0.01 ASI units conservatively represents the ability of procedures to determine the shape annealing function. Again, this uncertainty is a functior of reactor geometry and is independent of fuel type and reload cycle. The statistical confidence of this uncertainty is discussed in CEUPD-199 (8) Excore detector subchannel calibration using 1,.01 ASI incore detectors This uncertainty represents the limit to which the excore ASI can be calibrated to the incore ASI. As such it represents an absolute limit which cannot be exceeded when the calibration is performed. (9) Trip system processing i.02 ASI This uncertainty is currently being used at Fort Calhoun and is a function of the measurement and process instrumentation. Therefore, it is not sensitive to reload cycle cr fuel type. (b) The trip overshoot uncertainty of 5% as listed on page 15 of XN-NF-507 is treated as an uncertainty of the exieting reactor overpower trip to account for the possible variation in trip poiqt gto gljbration and measurement errors. This results in the use df M25 Schpower trip in the analysis of anticipated plant transients (XII-IIF-79-79) and is consistent with existing Fort Calhoun Technical Specifications (section 1.3). (d) The application of uncertainties is explained in detail in the response to Question 10 for the TM/LP analysis, and in the response to Question 23 for the APD and DIIB barn. i868 338 QUESTION 10 Section 4.2.1 lists typical uncertaintica included in the TM/LP analycec. Explain in detail hou each of these uncertainties arc included in the cet point calculations. Dafinc chat effecto are covered by instr:enant procaccing error. Pleacc stata uhather the depracc:adcation trancicnt uncertainty io the only adjust":cnt to the TM/LP for trancient effecto. See Qucation 17.

Response

In determining the TM/LP trip function, uncertainties to account for the magnitude of nuclear peaking, engineering tolerances, and instrument processing are included in both the XCOBRA-IIIC (DNBR) and PTSPWR2 (transient) calculations. The applica tion of the uncertaintics in each of these calculations are described bel ow. The TM/LP trip function is derived directly from the TM/LP safety limit lines, and uncertainties included in the TM/LP safety limit analysis are implicitly included in the TM/LP trip function. The TM/LP safety limit analysis included appropriate uncertainties in the calculation of the limiting assembly LHGR as: LHGR

• F"
• F
• F"P LHGR

= T M AT

where, LHGR is the limiting assembly linear heat generation rate with T

uncertainties LHGR is the nominal limiting assembly linear heat generation rate U F is the measurement / calculational uncertainty on F g R (F = 1.06) M F is the allowable azimuthal tilt (F = 1.03) AT AT U F is the core power measurement uncertainty used in TM/LP p (1.03) 1868 339 In addition to the above, an engineering uncertainty of 3% was applied to the Di;B limiting rod (F = 1.03) and the nominal geometry of the limiting sub-channel was adjusted to account for engineering tolerances associated with the fuel rod / guide tube dimensions. The verificatior, of the adequacy of the Cycle 6 TM/LP is accomplished by analyses of the anticipated operational occurrences (A00) during reactor operation. The results of those analyses have been documented (Xtt-t!F-79-79) and direct accounting of instrument response was included in t.:u analyses. This included the following uncertainties: 2% for core power, 2 F on core inlet temperature, 22 psia on pressure plus a 25 psia operational pressure Theresponsetime$associatedwithinstrumentprocessingwereaccounted range. for in the A00 analyses by assigning the appropriate total delay times to each of the LSSS trip functions as modeled in the analyses. These delay times repre-sent the time interval associated with signal acquisiticn + processing, and the movement of control rods when trip conditions are encountered. The delay times used in the analyses are listed in Xil-flF-79-79. The trip overshoot (Xti-fiF-507) was modeled as an uncertainty in the overpower trip in the A00 analyses (see response to Question 9). ?!c explicit accounting of the depressurization transient uncertainty was performed in determining the TM/LP trip function. Rather, the TM/LP trip function was determined from the TM/LP limit lines (SAFDL on Di;BR) with sufficient margin tc preclude penetration of the SAFDL on Di:BR for steady-state operation and anticipated operational occurrences (see response to Question 24). 1868 340

-2h-Question 11 Deleted by the NRC. 1868 341 QUESTION 12

cu is flu: tilt included in cet point calculations?

Response

A 3f. azimuthal tilt allowance is included directly in the analysis of the TM/LP, and the APD and DNB barns such that the limiting rod power (peak pellet) analyzed includes the 0.03 value for T consistent with the proposed Cycle 6 q Technical Specifications, in which the limiting pea. king is defined as R (1 + T ) and Ff, = Fxy (1 + T ). F =F R q q 1868 341 QUESTI0tl 13 Will E :an design of C.E. set pointa affect the Variabic Overycucr Trip Set point? If co, deceribe hou E==on designa thic trip function.

Response

The Eric set point methodology does not affect the Variable Overpower Trip Set point. The existing overpower set point was modeled in the analysis in s'upport of Cycle 6 operation for Fort Calhoun (XN-flF-79-79). 1868 342

-2T-Q!ESTIO 14 (a) Deceribe how the cerczt reactivity is included in the cet point calculations. Lict cach application of thic item. Describe how the ceram reactivity ic determined. (b) Deceribe or reference the Exxon modeling of CEA chutdoun reactivity as a function of CEA pocition. RESP 0:lSE The Generic Scram curve and worth are used directly in the analysis of the A00's (Anticipated Operation Occurances) to verify the adequacy of the Cycle 6 set points. Exxon Nuclear has generated a conservative generic trip reactivity curve by using a power distribution which was severly skewed toward the bottom of the core (+50% axial offset). The XTG computer code was used in the generation of the scram curve. Part of the study was made with a 3D, quarter core repre-sentation of ar, operating reactor. The remainder of the anlaysis was made with the ENC 1-D Power Distribution Control (PDC) model. The quarter core 3-D nodel was uscd to predict the negative scram reactivity insertion for nominal conditions (BOC and E0C) expected for steady state reactor operation. The 1-D model was used to study the effect of induced bottom peaked axial offsets up to an offset value of +503. In both models the control rod banks were incre-mentally inserted in each case to determine the scram curves. The EilC generic scram curve computed from the XTG models is based on the minimum scram negative reactivity insertion (N-1 rod worth) and the maximum allowable (by Technical Specifications) scram time. Figure 14.1 shows the Fort Calhoun scram curve for an allowable scram time of 2.5 seconds. The mea-sured plant scram time is much shorter than 2.5 seconds. 1868 343 k s .i i. 1200 0 C .? 80 t i i s. 8 .5 60 2? 3.- 40 u 'c3 20 as 0 .5 1.0 1.5 2.0 2.5 Time (sec) Figure 14.1 Fort Calhoun Scram Reactivity Curve O 1868 344 QUESTI0fl 15 Deceribe in detail and provide appropriato equation to chcu huu the various peaking factorc and hot channel factora are included in the various thermt hydraulica calculationc for the TWLP trip. If this ic done in the came way as other thermal hydraulica calculations and thic information hac aircady been provided to the Staff, the reference is cufficient. Resoonse The XCO3RA-IIIC code in determining the TM/LP trip function allows explicit representation of appropriate peaking factors and uncertainties. The linear heat generation rate associated with the subchannel (Df'BR) calculation is the limiting assembly peaking and is represented by the equation shown in the response to Question 10. The value of LHGR in that expression was selected to result in a minimum assembly flow rate for use in the subchannel calculations. This was accomplished by using the value of F as specified by the plant Technical Speci-R fications divided by the minimum hot rod peaking anticipated throughout tne fuel lifetime. The application of the engineering heat flux factor and geometrical changes associated with tolerances on fuel rod / guide tube dimensions is discussed in the response to Question 10. The XCOBRA-IIIC code uses the following expression to determine the local rod surface heat flux: d (ft,J) UiGRgyg

• F) *F
• P
• K
• 1/nD

= 2

where, d (fl,J) local rod surface heat flux (BTV/sec-f t )

= LHGRgyg average LHGR for the model (bundle or pin) = th F) fi pin or bundle peaking factor = th F g nodal axial peaking factor 2 P 1.0 for steady-state calculations = K conversion factor (kw to BTV/sec) = D fuel pin diameter (ft) = 1868 345 QUESTION 16 Describc the calculation of the saturation liztit cuz". cs. List any computer progr=:s used for this calculation, if required. List all uncertaintics included in the analyses and chou hou they are used.

Response

The saturation limit curves define those sets of reactor operating con-ditions (pressure, core power, core inlet temperature, core flow) which preclude a core average exit temperature equal to or greater than saturation temperature. The reactor average coolant exit conditions are judged not to be a. function of fuel type and, as such, the existing saturation limit curves are applicable for Cyle 6 operation. 1868 346 The specific criteria to preclude potential flow instability are: (1) a calculated subchannel quality of less than or equal or 15%, and (2) a core average exit quality less than or equal to zero. Adherence to these criteria is ensured by the SAFDL on DNBR and the core saturation limits. In addition, other reactor protection system limits such as the low pressure trip and variable overpower trip preclude reactor operation in a potential flow ins' ability mode. The ENC fuel melting limit of 21 kw/ft was calculated via the GAPEX code (XN-73-25) in conformance with the USNRC approved ENC fuel densification model for PWR fuels (USNRC report dated February 27, 1975). The value of 21 kw/ft as an LHGR lin'it for the existing fuel is documented in the Fort Calhoun Technical Specifications. Thus, the value of 21 kw/ft as a raaximum fuel rod LHGR for both fuel types is judged to be acceptable. A fuel melting temperature of 2790 C was assumed at beginning-of-life, and this value was decreased at the rate of 32 C per 10,000 MWD /MTU and com-pared against the fuel temperatures calculated at 21 kw/f t throughout fuel lifetime. 1868 347 QUESTION 17 Describe hou dynamic effects arc accounted for in calculating the T10LP trip. Vnat ccmputer codec are used? Wat transients are analyzed to determine the dyn=nic term? Hou are plant corditions initialized for the calculationc to determine dynamic effcats? List the valuca included for RPS delays to control rod actuation. Vnat core power chapes are accumed? What reactivity feedback mechanicma are included? %at values of reactivity feedback coefficients arc 2d? Justify thace values. Discucs the modeling of the core sencoro which would meacure core or loop conditions to assure adequate consideraticn of delays in system responce.

Response

The ENC methodology for the determination of TM/LP does not consider any a priori knowledge of system transient performance to allow an explicit accounting of dynamic effects such as depressurization, system time response, trip over-shoot, etc. Rather, a normalization of the set of lines representing those conditions corresponding to obtaining a SAFDL or DNB (MDNBR = 1.30) is performed so as to provide adequate protection against penetrating appropriate SAFDL values during steady-state and anticipated operational occurrences. Thus, any degra-dation of thermal margins due to changes in the reactor coolant conditions, time delays in instrument and scram response, power overshoot, must and are explicitly modeled in the transient calculations used to verify the adequacy of TM/LP.

Thus, the ENC set point methodology includes both steady-state and transient performance evaluations.

The plant transient calculations directly determine the time dependent plant responses as they affect the core coolant conditions and the MDNBR is calculated during the transient analysis. The analyses include the thermal hydraulic 1868 348 system response as well as the core neutronic response anticipated for Cycle 6. Any appropriate time delays associated with the reactor protection system are explicitly defined in the plant transient analyses (XN-NF-79-79) and are con-sistent with the anticipated RPS performance for Fort Calhoun. The uncertainties in nucl^ar peaking and control system instrumentation readings are consistent ..th ENC methodology (see response to Question 10). Uncertainties associated with other LSSS trip functioil: were applied to the nominal values in a limiting

fashion, i.e., set to those values which result in increased thermal margin degradation during the analysis of the anticipated operational occurrences.

The trip settings and delay times for these latter trip functions as used in the analysis are presented in XU-NF-79-79. 1868 349 Question 18 Is the core AT-power calculator equation included in Exxon setpoint calculations for OPPD? If it is, provide a description of the procedure used to calculate the coefficients and give their values. If it vill not be changed when Exxon calculates setpoints, explain why this is acceptable in view of the fact that different analytical methods will be used by Exxon. In particular, explain why this is acceptable for Fort Calhoan, which is planning operation at a higher power level.

Response

The core AT-power calculator equation was not included because it is independent of fuel type, power level and all cther parameters associated with the reload. 1868 350 QUESTION 19 Deceribe in datail the methods used by E==cn to incorporato into the cet point calculations those trancient cuenta ui:ich req ire a lov fku trip. Explain hou E::cn treata the equivalent of the Combuction Engineering concept of a Required Overpcuer Nbrgin.

Response

Loss of coolant flow transients are caused by a loss of electrical power to the primary coolant pumps and a corresponding increase in coolant tempera-ture. This increase, combined with the reduced flow, is anticipated to reduce thermal margin (MDNBR). The two most severe transients of this type (1) the loss of the four primary coolant pumps, and (2) the loss of are: two primary coolant pumps in opposite coolant legs. Since these transients result in changes in core power and inlet temperature not covered by the TM/LP function, core protection is provided by the RPS through the low flow trip. The adequacy of the existing low flow trip was determined by analyzing the above two loss of coolant flow transients for Cycle 6. Those analyses are reported in XN-NF-79-79. The required overpower margin is defined as that amount of reactor power increase necessary to prevent exceeding the SAFDL on DNBR during a full length CEA drop. The CEA drop results in a non-symmetric core power distribution, ar.d these changes in core peaking are considered in the analysis of the CEA drop. Reactor protection against SAFDL penetration is provided in determining the allowable core power as a function of axial shape index (LCO for DNB monitoring). The method used by ENC is described in detail in the response to Question 23. 1868 351 QUESTIO 1 20 Discuca in detail'the method used by E :on to incorporata into the cet point calculations those tranciento chich requira no reactor trip. Lict all the evento considered in thic category. Explain the step-by-ctep procedure that is used to calculate the required protection.

Response

In addition to the loss-of-coolant flow transients (see response to Question 19), the CEA drop transient results in changes in core power and inlet temperature not covered by the iM/LP function. In order to provide core protection against penetration of MDt;3R limits (SAFDL on Dt BR), the limiting condition for operation (LCO) for Dt B monitoring is determined and provided as part of the reactor operating limits. The CEA drop transient is analyzed using standard EflC predictive method-ology for DilBR (Xti-75-48) and a thermal hydraulic model which is essentially the same used in the determination of TM/LP (see response to Question 10). Included in the analysis are changes in the core power distribution consequent to the CEA drop. The method used to determine the LCO for DilB monitoring is described in detail in the response to Question 23 and the results presented in the Fort Calhoun Cycle 6 Safety Analysis Report (Xft-t!F-79-77). Conformance to these limits provides core protection against penetration of MD:BR limits for the CEA drop (no reactor trip) transient. 1868 352

~ QUESTIO:l 21 Provide a comparison of ENC cet point calculations uith those from a previous cycle of Ft. Calhoun ac diccucccd on page 3 of XN-NF-507. Explain the cause of any differences batueen the tuo calculaticnc.

RESPONSE

Calculations were performed determining the possible axial power profiles for the Cycle 4 Fort Calhoun core. Methods used were those addressed in the set point document; Reference 5. Sections 4.1.1 and 6.1 of Reference 5 show several parameters calculated for the Cycle 4 core. In the analysis Cycle 4 Technical Specification values were used for the SAFDL's and LCO limits. A comparison of the ENC calculated LC0 Axial Power Distribution " Barn" to that of the Fort Calhoun Cycle 4 Technical Specification, LCO, Axial Power Distri-bution, " Barn" is shown in Figure 5.1 of Reference 5. Small differences show up in the comparison between the Technical Specifi-cation APD LCO curve and that generated with ENC methods. Without 3.dditional information about the calculation made to determine the APD LC0 curve in the Technical Specifications, the cause of the small differences cannot be ex-plained.. Areas of possible differences could exist with the values of the uncertainties used in the two independent calculations, methods, codes, and or models. Specific details of both calculations would be requir i to resolve the small differences. 1868 353 Question 22 Provide the curves of Pfdn and BOPM versus peripheral axial shape index for the CEA insertons used for the Fort Calhoun Cycle 6 setpoint calculations. Re spon s_e_ Curves of Pfdn and BOPM versus peripheral axial shape index for the CEA insertions were not generated using the EIIC setpoint methodo-logy as detailed in the response to Question 2h. s 1868 354 QUESTION 23 Daccribe in detail the method used to calculate the linear hcat generation rate and DilB tento (such as Figurac and in the Fort Calhoun Technical Specificationo).

Response

The LC0 for Di;B monitoring (DNB tent) represents allowable core power as a function of axial shape index and provides administrative limits to preclude penetration of MDNBR limits during the CEA drop transient. The methodology used by ENC in determining the DNB tent is described below in detail. o Axial Power Profiles The core average axial power profiles used in the determination of the DNB tent were calculated using ENC neutronics methodology (XN-75-27) and represent that set of power profiles anticipated during Cycle 6 operation. The power profiles were sorted according to axial shape index. Within each small ASI increment, sufficient axial profiles were selected for analysis to ensure proper determination of the DNB tent. o Definition of Analysis Input The establishment of the DNBR analysis input includes allowance for reactor operating uncertainties, nuclear peaking uncertainties, and changes in peaking during the CEA drop. The core operating conditions used in the analysis were 2053 psia (minimal operating pressure minus 22 psia uncertainty), 547"F core inlet temperature (545 + 2 F uncertainty),1530 MWt total core power (1500 + 25 power uncertainty). The effect of the CEA drop upon hot assembly conditions was determined to result in increased nuclear peaking and decreased assembly flow. A peaking i868 355 augmentation of 121% was determined from the neutronics analysis of the worst CEA drop, and a reduction of 6% in the hot assembly flow was determined from appropriate core flow distribution calculations. The accounting for the above two effects was directly applied in the DriBR analysis. The average LHGR associated with the limiting assembly was calculated according to the following expression: LHGR*Ff*F,

• F AVG LHGR F
• g pU,pU

= G P

where, LHGP, core average linear heat generation rate

= F peaking limits as defined by the Technical Specifications = F peaking measurement uncertainty (1.06) = g F calculated CEA drop peaking augmentation (1.21) AVG F assembly local peaking factor (1.09) = g U F calculational factor to account for changes in nominal = G hot subchannel area due to allowed manufacturing tolerances (.938) U Fp core power measurement uncertainty (.98) = The above expression results in an LHGR reference (LHGR ref.) value of 11.46 at 100% of rated power. O Power Interation The limiting hot assembly LHGR is defined as that value which results in a calculatea f4DriBR = 1.306 (LHGRg). The 1.306 value is the product of the 18$8356

hl-SAFDL on Dt!B, engineering heat flux factor, and a correction factnr to account for that fraction of total generated energy which appears as rod surface heat flux (1.036 = 1.30

• 1.03
• 0.975).

The allowable core power fraction, P, is then defined as the ratio of LHGR to LHGR for each ASI value. The ordered pair (ASI, P) defines crit ef a point on the unadjusted DilB tent. The ASI value associated with each such point is then adjusted for neutronics uncertainty and for the excore cali-bration uncertainty associated with the determination of the ASI value in accordance wi th Xfi-i1F-507. The DlB tent as reported in Xfi-flF-79-77 is the loci of points thus generated. 1868 357

h2-The LCO for linear heat generation rate (LHGR) mon'.toring, LHGR tent, represents operating limis for allowable core power, (P) as a function of Axial Shape Index ( ASI) for those times during the cycle in which the 'ncore detectors are inoperable. The peak LC0 LHGR being monitored is determined by calculating the minimum allowable LHGR resulting from either the increase in power asso-ciated with a dropped rod or the LOCA LHGR limit. The methodology used by ENC in determining the LC0 LHGR tent is described below. AXIALS The axial power distributions used in the determination of the LC0 LHGR tent were the same axials calculated and used in the LSSS APD calculation. Methods used to detennine these axials are discussed in the response to Question 2. Approximately 1500 axial power profiles were used in the analysis for the LHGR LCO tent as well as the LSSS APD tent. IflPUT AtlD CALCULATIOt Input to the LCO LHGR tent analysis includes F va u s, all uncertainties Z associated with the LCO's, the Technical Specification values of F and F,, r the core average LHGR, and the minimum of either the maximum LHGR limit that would protect the core from violating the peak kw/f t for centerline melt during a dropped rod incident or the maximum LOCA LHGR limit. The calculation performed to determine the percent allowable power as a function of ASI is 1868 358 therefore: P(ASI) = % allowable power = 100

• K/F
• LHGR - Pm LHGR = Core average linear heat rate.

where K = LHGR Limit T L F =F

• FZ (ASI)
• FU Q

xy Pm = Power Measurement Overshoot = 2.0 F = ratio of hot pin to average pin at core

• 7 elevation Z.

F (ASI')= axial power peaking factor for axial Z shape index ASI'. Fh= Uncertainties = F

• F c

E LHGR aug F = Calculation = 1.07 C F = Engineering = 1.03 E F = Linear Heat Rate = 1.005 LHGR F = Augmentation Factors (Table A.1 of Reference 11 d"9 if dropped rod limited and 1.0 in LOCA limited) ASI = ASI - ASlu ASlu total.04 ASI units (flo trip system processing, Reference 5, since no reactor trip associated with LCO's.) The above expression gives ordered pairs of (P, ASI) which results in the LHGR LCO tent for Fort Calhoun. 1868 359

Response

Parts 1 through 8 of Section 6.2.3 (XN-NF-507) describe the construction of TM/LP safety limit lines. These lines are isobars on a plot of coolant inlet temperature versus allowable core power (Figure 24.1). Each isobar defines as a function of power the lowest coolant inlet temperature which precludes violation of one of the following limits: o MD"BR < 1.35 o Maximum hot channel quality > 15% o Core average exit temperature > saturation temperature The TM/LP safety limit lines thus define steady-state temperature, pressure, and power conditions which may not be exceeded without violating the SAFDL on DNB. (a) Selection of Worst Axial Power Distribution An XCOBRA-IIIC calculation is performed using each of the axial shapes under consideration. The power, inlet temperature, and pressure are fixed according to Steps (1), (2) and (3). Only the axial profile is varied. The worst axial profile at the selected power is the one which yields the least MDNBR or other appropriate SAFDL value in the XCOBRA-IIIC calculation. Table 24.1 lists the worst axials determined as above are used to construct the TM/LP safety limit lines. (b) Steps (1) through (6) establish a single point on a safety limit line. A second point on that isobar may be obtained by selecting a new power (and thus a new worst axial) and again performing a series of iterative XCOBRA-IIIC calculations to determine the limiting coolant inlet temperature. System pressure is constant in all XCOCRA-IIIC calculations used to generate the points on a single safety limit line. 1868 360

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1. 4 4$' 4%

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-hT-Sufficient points are generated at a fixed pressure, plotted, and connected by a piecewise smooth curve which represents a single safety limit line. (c) The procedures outlined above are repeated at other pressures to obtain the set of TM/LP safety limit lines. The reference to Figure 5.1 is in error and should refer to a set of TM/LP safety limit lines obtained in accordance with the steps outlined above. (d) The TM/LP safety limit lines are rit with a function of the form a PF(B)B + 8 Tin + Y3 P (l) = The fit is determined to conservatively envelope the TM/LP safety limit lines (see Figure 24.2). Equation (1) is not the TM/LP trip equation. The final TM/LP trip (TM/LP LSSS) equation is obtained from (1) by writing a PF(B)B + 8 Tin + Y (2) P = y where +b (3) = Y Y) and b is a positive constant which converts the TM/LP safety limit function to the TM/LP trip function and provides protection against penetration of the appropriate SAFDL during anticipated transients. The values of a, PF(B), and 8 are carried over to equation (2) without change. The procedure used to obtain a, S, y and PF(B) is detailed in the following. The value of 8 is calculated frcm two TM/LP safety limit lines as follows: 1869 002 (1) Select a pair of contiguous ;afety 'linit lines. For demonstration, select the 2100 psia and 2250 psia isobars. Then the value of P -P in Step (10) is given by: vag P -P 2250 - 2100 150 psia = = ur) 2 ur (2) Determine the value of T allcwed by each sa fety limit line at in 100% power (Figure 24.1). For the demonstration case, T 567.3 F = in) T 560.3 F = in 2 and T -T 567.3 - 560.3 7F = = in) in 2 (3) A value of 8 is then computed according to the formula in Step (TC): P -P var) var 2 150 psia B = 21.43 = T. - T. 7F in) in 2 Steps (1) through (3) above are repeated for each pair of contiguous safety lines in Figure 24.1. The value of e is finally fixed equal to the arithmetic average of the values thus ehtained. / 1869 003 The value of a is calculated as fcllowsi (1) Select a pair of contiguous safety limit lines. For demons tra tion, select 2l00 psia and 2250 psia isobars. (2) Calculate the value of s associated with those two safety limit lines, as described above (e = 21.43), (3) Calculate the slope of each safety limit line in the straight line portion: (a) for the 2100 psia isobar, slope = A 560.3 - % 7.4 .2417 AB 112 - 100 4 (b) for the 2250 psia isobar, aT 567.3 - 564. slope = 73 D2 - 10'0-- = .2750 -= (4) Compute a value of.x as ie product of the appropriate 6 (Step 2 above) and the arithmetic meca of the absolute values of found in Step (3) above: 21.43 [.2417 +.2750](1/2) = 5.54 a = (5) Repeat Steps (1) through (4) above for each pair of contiguous sa fety limit lines in Figure 24.1. The final value o f a is fixed equal to the arithmetic avere<je of the values thus obtained. s The value of y) is chosen to yield a function which conservatively envelopes all the sa fety limi t lines, lhis is accomplished as follows : (1 ) Select a sa fety limi t line. For demonstration, choose the 1750 psia i soba r. 1869 004 (2) Determine from Figure 24.1 the allowed inlet temperature at the pressure selected in (1) and 100% power For the demon-stration case, 545 F T. = in (3) Write equation (1) at the conditions 545 F T. = I" 5.54 = a 100 B = 22.48 8 = 1750 psia P = and solve for y) to get: y) P - a B PF(B) - 8 T = in 1750 - 554 - 22.48(545) = -11,055.6 = The function thus developed is a closed form representative of the TM/LP safety limit lines. The TM/LP trip equation differs from that function only in the value of the additive constant. The value of y in equation (2) is selected to protect the SAFDL on DNB during anticipated trar.sients and to allow a reasonable administrative operating band. To obtain the value of y, cquation (2) is rearranged and evaluated at conditions consistent with those requirements, as shown below. P - a PF(B) B - s T = y var in 2075 - 5.54 x 100.5 - 22.48

• 548

= -10,801. =

where, 2075 is the least rated pressure allowed at full power operation 100.5 is the maximum operating power allowed by the LCO for DNB monitoring 548 F is the rated coolant inlet temperature at full source operation plus a 3 F offset to allow administrative control of reactor operating variables.

1869 005

The TM/LP trip function is P 5.54 PF(B)

• B + 22.48
• T

- 10,801. var in The value of b +t. .ation (3) may be computed as 255 psi. About 55 psi. of that are allocated as uncertainties, leaving 200 psi margin between reactor trip and penetration of the SAFDL on DNB as represented by the TM/LP safety limit lines. That 200 psi is equivalent to a 36 percent overpower margin. The curve fitting described above might have been accomplished via least squares analysis. The present method has at least two advantages over that alternative: 1) simplicity and ease of application, and 2) direct derivation from the physical meaning of the parameters. (e) The function PF(B) was selected by trial and error to provide a safety limit function

• which conservatively envelopes the TM/LP safety limit lines.

Variations in axial shape defined by APD were modeled explicitly in the calculations which established the safety linit lines and are therefore implicit in PF(B). No explicit relationship between this function and allowed axial flux shapes is necessary within the context of ENC methodology. (f) A complete discussion of procedures used to verify the adequacy of the TM/LP trip function is cont.ined in XN-NF-79-79 and XN-NF-79-79, Supplement 1, " Fort Calhoun Cycle 6 Rel sad Plant Transient Analysis Report." (g) The response to Question 15 discusses the application of appropriate peaking factors used to construct the TM/LP trip function.

• See equation (1), response to Question 24(d).

1869 006

620 o cI 600 N 2 E. 8 ~ 2A00 5 m ~ 2250 N 0 560 2100 0 1900 0 1750 540 co 1 i i I 70 80 90 100 110 m Ci Core Power, Pct. of Rated CD N FIGURE 24.1 TM/LP Safety Limit Lines for 1500 MW, 4-pump Operation

Pressure = P) e B C E. _2 5 E e y 3 N y 2 S o 1 Loci of Safety Limit Points 2 Equation (1) 3 TM/LP Trip Function (Eq. 2) 1 t 1 1 1 1 I Q Os Core Power, Pct. of Rated e O FIGURE 24.2 Comparison of Safety Limits (1) with (2) A Closed g Form Function which Conservatively Envel] pes them m and (3) The TM/LP Trip Function

-p-QUESTIDN 25 Diccuss in deta 2 how the decign differcnces of differcn: typec of fuel bundles i1 a reload core are included in the cet point calculations. In particular, differeneca in fuct hw:iice manufactured by tuo differcn: fuc t vandcre (e. g., E==on and Ccmbustion Engineering for the Ft. Calhoun Cucle 6 reloci) chould be adi*c.=d. Hou does E==cn model fuel hundica of anothcr manufacturcr? E=a.ples of possibic arcas of possible insufficient infomation are (1) fuel densificaticn characteristics (2) fuel bundic ~lco dictribution characterictics (3) fuct rod internal fill cas prese:ee Resconse The differences in fuel design between the Etic and existing fuels were accounted for in the determination of the A'D, TM/LP and LC0 for Drib monitoring for Cycle 6. Local power augmentation as a result of axial densification must be accounted for in the determination of fuel pellet peaking as used in the deter-mination of the APD barn. The peaking augmentation for the EliC fuel was deter-mined and compared against the values currently stipulated for the existing Ft. Calhoun fuel. The larger peaking augmentation as a function of axial height between the ENC and existing fuel was chosen as the appropriate value in deter-mining the Cycle 6 APD barn. The thermal hydraulic models used to calculate the TM/LP and LC0 for Ot;B monitoring explicitly modeled the hydraulic performance of both fuel types in order to determine the appropriate limiting subchannel flow. The determination of the limiting assembly flow, and subsequent MONBR, is accomplished in two steps: (1) Core flow distribution to determine the limiting assembly flow rate, and (2) Limiting assembly calculation for evaluation of the core thermal margin (MDNBR). 1869 009 The core flow distribution calculation directly models the thermal and hydraulic performance of each fuel type as appropriate single hydraulic channels. The thermal performance is evaluated using ENC neutronics methods to determine the core and assembly peaking distribution while the hydraulic performance is determined using the results of pressure drop testing performed by ENC for both fuel types. The results of the calculations inai ate that both limiting fuel assemblies will experience no less than 55 of core average assembly flow rates with the ENC assembly having slightly lesser flow. Thus, the limiting ENC fuel assembly was selected for TM/LP and LCO calculations. The limiting assembly calculations model the limiting ENC fuel assembly into appropriate subchannels with the assembly flow rate as determined above. The calculation is consistent with the methodology used for the core flow dis-tribution calculations. This calculation results in determining the limiting subchannel flow rate used in the ensuing MDN3R calculation (see response to Question 23). 1869 010 QUESTION 26 Diccuco protcetion providad by sat pcint: tc ccy'=ctric trancicnto (Cuch. as th.cce affecting only one stc = guneratcr). Reference The loss of load to one steam generator was analyzed using the same plant transient simulation model for the analysis of other plant transients (XN-NF-79-79). In this analysis, all proposed Cycle 6 LSSS trip functions were explicitly nodeled. The results of the analysis were documented and transmitted as supplementary information in support of the Cycle 6 -Licensing Application (Letter Report from W. C. Jones to R. Reid dated December 4,1979). The results of the analysis indicate that adequate core protection is provided for Cycle 6 with the proposed set points to preclude penetration of the appropriate SAFDL for Cycle 6. 1869 011 Questien What equaticns are used to calculate the D::BR used in the XCOERA code?

Response

The equation of the W-3 correlation along with the equations for the non-unifor= axial heat flux facter and the unheated boundary factor are given in X.-75 l+5 The precedure for using the W-3 correlation and correction factcrs is also given. 1869 Gi2 REFERENCES 1. R. B. Stout, XTG-A Two-Group Three Dimensional Reactor Simulator Utilizing Coarse Mesn Soacing, XN-CC-28, Revision 5, Exxon Nuclear Company, July 1979. 2. F. B. Skogen, Exxon Nuclear Neutronic Design Methods for Pressurized Wa ter Reactors, XN-75-27, Exxon Nuciear Company, June 1975. 3. F. B. Skogen, Exxgn_ Nuclear Neutronic Design Methods for Pressurized Wa te r Reactors, XN-75-27, Supplement 1, Exxon Nuclear Company, September 1976. 4. F. B. Skogen, Exxon Nuclear Neutronic Design 'tethods for Pressurized Water Reactors, XN-75-27, Supplement 2, Exxon Nuclear Company, DecemDer 1977. 5. L. A. Nielsen, ENC Set Point Methodolcav For CE Reactors, XN-NF-507, Exxon Nuclear C Epany, June 1979. 6. F. B. Skogen, XPOSE - The Exxon Nuclear Revised LEOPARc XN-CC-21, Revision 2, Exxon Nuclear Ccmpany, April 1975. 7. W. R. Caldwell, PD07 Reference Manual, WAPD-TM-678, Westinghouse Electric Corporation, January 1965. 8. R. J. Breen, et al., HARMONY: System for Nuclear Reactor Deoletion Computation, WAPD-TM-478, Westingnouse Electric Corporation, January 1975. 9. L. M. Petric and N. M. Greene, XSDRNPM: AM?X Module With One-Dimensional Sn Capabili ty for Spatial Weighting. 10. "Po.ser Spike Model for Pressurized Water Reactor Fuel", March 1979, XN-207. 1869 013 -}}