ML19338E949

to Exxon Nuclear Co Setpoint Methodology for C-E Reactors.

ML19338E949
Person / Time
Site:	Fort Calhoun
Issue date:	07/18/1980
From:	Krysinski T, Nielsen L, Stout R SIEMENS POWER CORP. (FORMERLY SIEMENS NUCLEAR POWER
To:
Shared Package
ML19338E944	List: ML19338E943 ML19338E949
References
XN-NF-507, XN-NF-507-R01, XN-NF-507-R1, NUDOCS 8010060350
Download: ML19338E949 (72)

Text

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O XN-NF-507 Revision 1 07/18/80 ENC SETPOINT METHODOLOGY FOR CE REACTORS Prepared by: L. A. Nielsen T. '. Krysinski F. T. Adams Approved by:

• M (d ' N R.[B. Stout, Manager Neutronics & Fuel Management 4pffA b6,pst J. N. Morgan, Manager Licensing & Safety Engineering G. A. Sof [eManagd Nuclear fuels Engineering i

I ERON NUCLEAR COMPANY,Inc.

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I 3 i NUCLEAR REGULATORY COMMISSION DISCLAIMER I

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IMPORTANT NOTICE REGARDING CONTEN(S AND USt! OF THIS DOCUMENT

, PLE ASE RE_A_D_C._AR E FULLY J

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j This technical report was derived through research and development I

programs sponsored by Exxon Nuclear Company, Inc. it is being sub 1

1 rnettt*t by Exxon Nuclear to the USNRC as part of a technical contri-Intion to f.x:ilitate saf et y analyses by bretiwes of tlm llSNRC wina h l

j utih/c fxxon Nui. lear f ala n.atail v elo.ul fuel or utire tes hnu al wrvu es l provuhst by E xxori Nut. lear for linht w. ster powet re.n.iors .ind it isrene 1

aruf cof rec t to the lost of Exxon Nuclear's k nowhslip , iii fo r e na tion, j anal telmf. T he inf ormation coniannst herein anay be u'.est by the IJSNHC l en its review of this report, .nul by brensees or appht ants hetore the

! USNRC which are customers of Exxon Nudear in their demonstration I f of comphance with the USN RC's regulations. E l Without derogating f rom the foregoing, in ither Exxon Nuc lear nor 4'

any person actmg nn its la4.alf:

A. Makes any warranty, express or implust, with respect to

} the .u.riar .ii y, c:iungile teriess, air assefiitricss ist tiie nihse

( mat nin a untauissi in llus di(13,14) must be greater than or equal te -

an approved safety limit, which ensures the integrity of the fuel and that the SAFDL's are not violated.

(2) The thermal-hydraulic conditions of the reactor core must be within the limiting range of the empirical correlation.

(3) The bulk coolant temperature at the exit of the core must be less than the saturation temperature.

(4) Potential flow instability limits may not be violated.

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XN-NF-507 I

Revision 1 The W-3 burnout heat flux correlation with correction factors for the presence of a cold wall and nonunifonn axial heat flux was used to esta-l blish the specified acceptable fuel design limit (SAFDL) on the fuel burnout perfonnance (DNBR). The application and interpretation of the W-3 correlation for determination of the reactor setpoints is consistent with the ENC pre-dictive models for DNBR II4) .

The SAFDL on DNBR is protected by limiting the operating values of core power, coolant inlet temperature, and system pressure to the most conser-vative of the following:

e That set of operating values which results in MDNBR = 1.3.

e That set which gives rise to parameter values exceeding the range of the W-3 correlation, e That set which precludes the occurence of bulk coolant boiling at the core exit.

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.

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 instability mode. I The TM/LP is provided to prevent operation from exceeding the above safety limits including allowance for measurement error and uncertaintes. A I

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i XN-NF-507

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Revision 1 typical TM/LP LSSS is shown in Figure 4.5 which graphically defines the limiting values of reactor coolant pressure, core inlet temperature, and

[ reactor power level. The low pressure setpoint of 1,750 psia will always trip th', reactor. The continuous TM/LP LSSS trip function is determined according to the following formula:

P var

= a* O*B+8* TIN + Yg where

( P var

= TM/LP trip (psia)

B = High auctioneered thermal (AT) or nuclear power, in % of rated power PF(B)= A core peaking function that defines the variation of overpower with respect to core power

( T IN

= Core inlet temperature ( F) a = the change in primary pressure needed to maintain a given margin to an allowable DNBR for a given change in core power at a constant inlet temperature and overpower margin I~ 2 s = = the change ir, primary pressure needed TINy TIN 2 to maintain a given margin to an allowable DNBR for a given change in inlet temperature at a constant core

.. power and overpower margin.

y = A pressure bias term used to adjust the calculated 1

P var to account for system uncertainties and measure-ment errors.

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XN-NF-507 I

Revision 1 The TM/LP trip effectively monitors all those NSSS parameters, except mass coolant flow ate and core peaking which affect the thennal-hydraulic safety limits.

The calculated P var is then compared to the mea-sured pressure, and a trip signal is generated when P approaches measured pressure with a specific uncertainty. The APD analysis defines the limiting axial shapes that are to be considered in the TM/LP trip determination.

Therefore, the effect of power distribution on TM/LP trip is implicitly considered, at each thermal-hydraulic safety limit considered, thermal margin limit lines are generated over a wide range of pressure, inlet temperature and power. The variation of these thermal limit Loci with DNB overpower is incorporated in the TM/LP trip by the PF(B) function as indicated above. A procedure to obtain TM/LP LSSS trips using ENC methodology is described in Section 5.2.

4.2.1 Typical Uncertainties I 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 application of the uncertainties in each of these calculations are dec fibed below.

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:

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XN-NF-507 Revision 1 LHGR T

= .T@i

• F *F AT *I where, LHGR T

is the limiting assembly linear heat generation rate with' uncertainties UTCli is the nominal limiting assembly linear heat genera-tion rate .

F is the measurement / calculational uncertainty on F R

(Fg = 1.06)

F ZT is the allowable azimuthal tilt (FAT = 1.03)

F is the core power measurement uncertainty used in

[ TM/LP (F = 1.03)

The treatment of plant system measurement uncertainties and instrument response and delay times are directly accounted for in the verifi-cation of the adequacy of the TM/LP trip function. This includes the following

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typical uncertainties: 2% for core power, 2 F on core inlet temperature, 22 psia on pressure plus a 25 psia operational pressure range. The response times associated with instrument processing were accounted 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 represent the time l interval associated with signal acquisition + processing, and the movement of control rods who M canditioris are encountered. The delay times used j in the analyses . .rectly modeled in the analyses. The trip overshoot

-was modeled as an uncertainty in the overpower trip in the A00 analyses.

This results in the use of an overpower trip 5% higher than the nominal f

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Revision 1 value and is forced to be consistent with existing plant technical specifi-cations.

No explicit accounting of the depressurization transient uncertainty was perfonned in determining th TM/LP trip function. Rather, the TM/LP trip function was detennined from the TM/LP limit lines (SAFDL on DNBR) with sufficient margin to preclude penetration of the SAFDL on DNBR for steady-state operation and anticipated operational occurrences.

4.3 VARIABLE HIGH POWER This trip is designed to protect the core at all power levels against transient initiated core power increases.

Calculations show that CEA insertion causes changes in nuclear pcm king throughout the core. In order to avoid excessive peaking due to CEA insertion and exceeding core power limits, control rod insertion during reactor operation is limited by the Power Dependent Insertion Limit (PDIL).

This limit assures that the nuclear peaking will be within acceptable values at any power level by limiting control rod insertion as a function of reactor power. The variable nuclear overpower trip together with PDIL assures that the peaking limits at specified power levels will not be violated during operation at the power dependent insertion limits and the allowed nuclear overpower value.

The setpoint calculations for axial shape index and TM/LP are made with the bounding values of nuclear peaking at the power levels which will result in a nuclear overpower trip. These calculations are perfonned at maximum allowed power levels consistent with the PDIL and variable overpower trip system. The nuclear overpower trip in conjunction with the TM/LP and I

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XN-NF-50/

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Revision 1 APD trips prevent the specified acceptable fuel design limits (SAFDL) from being exceeded during'a power excursion in any rodded or unrodded core con-figuration.

The adequacy of the existing variable high power setpoints is usually verified by explicitly modeling this trip function in the analyses of anticipated plant transients. This modeling is discussed in detail in Section 6.0.

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! XN-NF-507 I

Revision 1 I

Table 4.1 'ypical i Fuel Augmentation Factors (15) i Non-Collapsed Clad I

Core Core j 11eight lleight Augmentation

(%) (inches) Factor 98.5 134.7 1.057  :

86.6 113.6 1.051 1

77.9 106.5 1.047 l i l i 66.2 90.5 1.041

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4 54.4 74.4 1.035 ,

45.6 62.3 1.030  !

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22.1 30.2 1.017

) 13.2 18.2 1.011 1.5 2.0 1.001 i

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• Augmentation factor for evaluation of fuel centerline melt only.

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XN-NF-507 Revision 1

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5.0 LIMITING CONDITIONS FOR OPERATION (LCO)

I I. The Limiting Safety System Settings, (LSSS) are setpoints input into the

. RPS that cause a reactor trip if the corresponding monitored parameters equal the input values protecting the reactor core from exceeding any Specified Acceptable Fuel Design Limit (SAFDL). SAFDL's must not be violated during

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those anticipated operational occurances which are expected to occur one or

{ more times during the life of the plant. A list of the transients to be investigated are shown in Table 5.1. All the anticipated operational occur-rences except the CEA drop and loss of flow have to be considered for the determination of the LCO's. The LC0 limit resulting from the LOCA analysis

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is monitored by the incore detectors. A backup monitoring envelope is

{ usually in the Technical Specifications in the event that the incore detec-tors ~311.

( During normal operation the peak linear heat rate is monitored through the use of the incore detection system. The peak LHGR is maintained to be below the values which are calculated to result in fuel centerline melt during a CEA drop or a peak clad temperature or 2200U F during the postulated

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loss of coolant accident. In addition, penetration of DNB limits during steady state operation and anticipated transients including the CEA drop is precluded through the implementation of LC0 for DNB monitoring. This is accomplished by determining the axial shape index during operation through the excore detectors and comparing this value against the allowable shape

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index as a function of core power. Analysis of the most limiting transients I

XN-NF-507 I

Revision 1 I

is performed with a variety of axial power shapes. The results of the analysis are reduced to provide allowable core power as a function of axial shape ,

index, (ASI) in a fashion consistent with that presented in Section. 4.1.1. l In the event that the incore detectors are not in operation for an extended period of time, the peak linear heat rate will be monitored through the use of a linear heat rate LCO. This LC0 is determined in a manner simi-lar to the APD limiting safety system setting as previously described except that the allowable power as a function of ASI is determined through analysis of the CEA drop and the postulated loss of coolant accident.

The most limiting transient is of primary interest in determining the axial shape index LC0 limits. An example of generation of transient axial shape index LC0's expected to be most limiting (CEA drop incident) is de-scribed in Section 5.1. In Section 5.2 the axial shape index LC0 that imple-ments the LC0A limits is described.

5.1 CEA DROP AXIAL SHAPE INDEX LIMITS The CEA drop incident is defined as the inadvertent release of a CEA r .using it to drop into the reactor core. The absence of a turbine run-back following a CEA drop at the EOC boron condition will tend to restore the reactor to near full power with an adversely distorted power distribution.

Therefore, it is necessary to maintain the linear heat rate within limits to assure that the SAFDL's are not exceeded during the transient. If the allow-able linear heat rate for the CEA drop incident is limiting, the increase in the three-dimensional power peaking due to the CEA drop must be determined.

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XN-NF-507 .,

Revision 1 The APD centerline melt evelope is then reduced by that amount corresponding to that increase in power to that increase in power peaking calculated to result from a CEA drop. The most adverse nuclear peaking and axial power distributions are used in the analysis 'of the CEA drop.

The LC0's are more restrictive than the LSSS, and the plant opera-tion is administratively maintained within the LC0 limits. Figure 5.1 shows a typical LC0 " barn" for protection of LHR during a CEA drop.

5.2 LOSS OF COOLANT ACCIDENT (LOCA) LC0 LIMITS Ti.e consequence of a loss of coolant accident is evaluated in accordance with the Acceptance Criteria as presented in 10CFR 50.46. These criteria maintain that the rated maximum linear heat generation shall be such that:

1) the calculated peak fuel element clad temperature does not exceed 2200 F.

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2) the amount of fuel element cladding that reacts chemically 4

with water or steam does not exceed 1% of the total zircaloy associated with the active fuel rod length in the reactor.

3) 'the cladding temperature transient is terminated at a time 1

when the core geometry is still amenable to cooling and the j hot fuel rod cladding local oxidation does not exceed 17%. l

4) the system long-term cooling capabilities provided for pre-vious cores remain applicable for EhC fuel.

XN-NF-507 I

Revision 1 I

The allowable linear heat generation rate which satisfies the 10CFR 50.46 is evaluated using F,y and zF pairs identified in the " Fly Speck" I analyses which correlates power peakings and axial shape index. The allow-able Linear Heat Generation Rate, (LHGR) or F , for the core is thus evaluated as a function of axial peaking. The exposure dependence of the F is also evaluated and is superimposed on the F versus axial position curve as required.

The LC0 for linear heat generation rate (LHGR) monitoring, LHGR tent, represents operating limits for allowable core power, (P) as a function of Axial Shape Index (ASI) for those times during the cycle in which the incore detectors are inoperable. The peak LC0 LHGR being monitored is deter-mined by calculating the minimum allowable LHGR resulting from either the increase in power associated with a dropped rod or the LOCA LHGR limits. The methodology used by ENC in determining the LC0 LHGR tent is described below.

The axial power distributions used in the determination of the LC0 LHGR tent were the same axials calculated and used in the LSSS APD calcula-tion. Methods used to determine these axials are discussed in Section 4.1.1.

Typically 1,000 to 2,000 axial power profiles were used in the analysis for the LHGR LCO tent as well as the LSSS APD tent.

Input to the LC0 LHGR tent analysis includes ZF values, all uncer-tainties associated with the LC0's, the Technical Specification values of F r and Fxy, the core average LHGR, and the minimum of either the maximum LHGR I

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XN-NF-507 1 Revision I limit that would protect the core from violating the peak kw/ft for center-f line 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 therefore:

P(ASI) = % cllowable power = 100

• K/Ff

• LHGR - Pm I'

LHGR = Core average linear heat rate.

where K = LHGR Limit F =F xy *F 7(ASI)*h Pm = Power Measurement Overshoot f F xY

= ratio of hot pin to average pin at core elevation Z.

( FZ (ASI') = axial power peaking factor for axial shape index ASI'.

Fh= Uncertainties =F c

• F E *I LHGR *F aug FC = Calculation Uncertainty FE .= Engineering Uncertainty

-FLHGR = Linear Heat Rate Uncertainty F

aug = Augmentation Factors (See Table 4.1)

( ASI = ASI - ASlu ASui = total .04 ASI units (See Section 4.1.3)

The above expression gives ordered pairs of (P, ASI) which results in the LHGR LC0 tent.

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XN-NF-507 I

Revision 1 Table 5.1 Incidents Considered in Transient and Accident Analysis I

Anticipated Operational Occurrences for which the RPS Assures no Violation of SAFDL's:

Control Element Assembly Withdrawal Boron Dilution Startup of an Inactive Reactor Coolant Pump Excess Load Loss of Load of Feedwater Flow I Excess Heat Removal due to Feedwater Malfunction Reactor Coolant System Depressurization Loss of Coolant Flow l Loss of AC Power

\

Anticipated Operational Occurrences which are Dependent on Initial 1 Overpower Margin'for Protection Against Violation of SAFDL's:

Loss of Coolant Flow l Loss of AC Power Full Len7th CEA Drop l Part Le,geh CEA Drop Part length CEA Malpositioning Transients Resulting from Malfunction of One Steam Generator I

1 Requires Low r'... Trip

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- XN-NF-507 i Revision 1

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Used in Transient Analyses M M M M M M M M M M M Ms W W W M M M M

XN-NF-507 Revision 1

-6.0 ENC METHODOLOGY Standard ENC methods and calculational tools are employed to calculate the limiting safety system settings and limiting conditions of operation for the CE reactor protection system. The analysis requires extensive calcula-

' tions in the Neutronics, Thermal Hydraulics and Nuclear Safety areas. The l

plant transients are evaluated through Plant Transient Simulation (PTS) calculations. The following is a general description of ENC methodology employed for determining setpoints for'CE reactors.

6.1 NEUTRONICS The methods and the computer programs used to determine LSSS and LCO-setpoints for CE reactors are the standard ENC methods approved by the

! NRC which are described in References 16,17, and 18. The neutronic computer programs used to do the analysis consist of the cross section generation code XPOSE(I) , the pin-by-pin diffusion theory code PDQ7/ HARMONY (2,3) , and the three-dimensional reactor simulator code XTG b) . A synthesis code CESPT was written specifically to process data caltalated with the PDQ and XTG core models.

i 6.1.1 Cross Sections, XPOSE l

Cross sections for the reactor simulator codes, PDQ7 and XTG, are calculated with the ENC computer code XPOSE. The XPOSE code is a i

l

. modified version of'the industry accepted LEOPARD code. It is used to L

generate fast and thennal neutron spectra and cross sections. Microscopic cross sections are generated by XPOSE for PDQ7/ HARMONY and macroscopic cross sections are determined by XPOSE for XTG. Non-fuel region macroscopic cross

XN-NF-507 I

Revision 1 sections are detennined by XPOSE and reflect the spectrum effects of the I

surrounding fuel.

6.1.2 Simulator PD0 Detailed radial calculations of the core are performe' with computer co :e PDQ7/ HARMONY in two-dimensions. The core is modeled in PDQ7 on a pin-cell basis; i.e., one mesh block per fuel cell. Each pin cell has the a;,propi te nuclide concentration of the burnup history for that pin. The PDQ7 modd is similar to the model described in Reference 16. Output from the PDQ7/ HARMONY calculations include the pin-by-pin radial power distributions.

These values are used as input to XTG for the F xy calculation.

6.1.3 Simulator XTG 3-D XTG The reactor core is modeled and depleted in three-dimensions with the reactor simulator code XTG. As mentioned above, cross sections for XTG are generated with PDQ7 and XPOSE. Each assembly in XTG is represented by four radial nodes and twelve axial nodes. Control rod cross sections and assembly local peaking factors are input into the code by assembly. The comouter code XTG calculates radial and axial power distributions and deter-mines Fxy, the ratio of the hot pin to average pin on a planar basis, i .e.,

with 12 axial nodes, 12 axial F xy values are detennined.

Figure 6.1 shows a typical axial power distribution at about 500 MWD /MT. Other features of XTG used e Integral control rod worth calculations, Figure 6.2 e Control rod bank worths e Core average cross sections by axial node e Core average exposure by axial node, Figure 6.3.

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XN-NF-507

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Revision 1 1-D XTG Sets of axial power distributions and axial shape indices are calculateu with a one-dimensional model of XTG using the axial dependent core average cross sections and exposures-determined in the 3-D-XTG cycle depletion described above. The reactor core is modeled in one-dimension using one radial radial node and 24 axial nodes. The 24 node axial exposure distribution is determined from the 12 node 3-D XTG exposure distribution, (Reference Figure 6.3). The integral rod worths are also calculated with the one-dimensional model. A typical comparison of the integral rod worth measurements to the 3-D and 1-0 XTG integral rod worth calculations are shown in Figure 6.2. The agreement between the measured and calculated integral rod worth is very good.

Using the above inputs, the 1-D XTG procedure to determine ordered pairs of axial power and symmetric offset is:

1. Incite a xenon oscillation at HFP, all control rods out, and with or without thermal hydraulic feedback depending if the calculations are at E0C or B0C, respectively. In the 1-D XTG calculation a xenon oscillation can be incited l by partially inserting control rods for several hours and then "instantly" removing the control rods. Feedbacks are used in all cases except during the initial xenon.

oscillation calculation at the B0C.

2. Select 30 to 40 different xenon distributions from Step 1 and rerun several static 1-D XTG calculations with power levels corresponding to the nuclear overpower trip set-

_ point and control rods at the CEA power dependent insertion limits. Use thermal hydraulic feedback.

XN-NF-507 Ii Revision 1

3. Determine the axial shape index symmetric offset, S, and axial power disibution, F z, f r the various power levels and corresponding PDIL lim;ts in order to deter-mine the percent allowable power. The percent allowable power is determined with the synthesis computer code, CESPT.

6 . '_ . 4 Axial Shape Index Setpoint Code, CESPT The computer code CESPT, is used to determine the ordered pairs of P, % allowable power, and Sp, peripheral axial shape index. Input for CESPT is taken from the previous calculations made with PDQ7, 3-D XTG, and 1-D XTG. CESPT, then, determines the minimum allowable power for each axial power distribution detennined with the 1-D XTG calculation described above. The input to CESPT is conservative. Values of F are input into the code by axial position. The core is divided into axial position. The core is divided into axial zones depending on control rod insertion, (each zone has the identical control rod configuration extending throughout the zone).

Values for F xy are extracted from PDQ7 and 3-D XTG and input into the corres-ponding axial position and zone in the synthesis code. A description of the l

" zone" modeling is shown in Figure 6.4.

The axial power distribution, F , determined with the 1-0 z

XTG model are input by axial height in CESPT. The fuel augmentation factor, F

ug, is also input by axial position. The combination of F,y, and Faug F gives values for the Fg as follows:

7 F

g

= F xy xF 7 xF aug without uncertainties I

XN-NF-507

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Revision 1

.or Ff=Fq xF u

-with uncertainties where F = ratio of hot pin to average pin at core evaluation z F = axial power peaking factor z

F = uncertainties N

F .= Fz xF xy = hot spot in core F = Fu xF aug_

f The percent allowable power is therefore calculated to Le:

P = % allowable power, = 100 x k/F xW aug

[

where W = average core LHGR for 100% power avg l k = kw/ft for centerline melt (a typical value is 21 kw/ft) l l The axial power _ profiles used above are the core average values which give an axial shape index, S, for the overall core. As stated

-in Section 4.1.2,'the core average axial shape index, S, must be modified by an adjustment Factor, F, to account for rod shadowing effects on the exccre

!' detectors.. The rod shadowing effects are directly applied to S, in the above calculation giving -(% power, Sp). These are the sets of ordered pairs used in' the setpoint analysis. Figure 6.5- shows typical fly-speck values of the

O E

XN-NF-507 44-Revision 1 percent allowable power, P, as a function of symmetrical offset, without uncertainties, F .

u Figure 6.6 illustrates an example of the axial shape index limiting safety s: stem setting for centerline melt at 21 kw/ft. The application f uncertainties to the axial shape index points shown in Figcre 6.5 is shown in Figure 6 ; Tiie uncertainties applied to the setpoint an11ysis for axial shape index centerline melt are discussed in Section 4.1.3.

The above neutronic discussions deal with the methodology used to detemine the ordered pairs of symmetric offset vs. % allowable power (S,P). Ir, i:m course of doing the calculations the axial power profile and peaking limits are determined which are applied to the IM/LP calculation for setpoints.

6.2 THERMAL-HYDRAULIC Methods and programs used to determine thennal margin and justify the current ENC DNB correlation for PWR's are described in References 5 and

14. The criteria established to assure that the thermal-hydraulic limits are not exceeded are listed in Section 4.2. The ENC thermal-hydraulic setpoint methodology and related computer codes are described below.

6.2.1 XCOBRA-IIIC Code XCOBRA-IIIC(5) is an ENC modified version of COBRA-IIIC computer code (19) The XCOBRA-IIIC code produces both transient and sMdy-I state calculation capabilities while including the effect of cross f t..

mixing between fuel assemblies and subchannels. The thermal-hydraulic paratters, such as DNBR, local quality and void fraction, are calculated for each node. The degree of core-wide nodalization and the modeling options I

I

)

XN-NF-507 Revision 1 availaule in the code provide calculational flexibility. This code is used for thermal-hydraulic parameter evaluation to generate TM/LP LSSS.

The thermal-hydraulic models used to calculate the TM/LP and LC0 for DNB 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 MDNBR, is x

accomplished in two steps: (1) Core flow distribution calculation to deter-mine the limiting assembly flow rate, and (2) Limiting assembly calculation for evaluation of 1he core thermal margin (MDNBR).

The core flow distribution calculation directly models the thermal and hydraulic performance of each fuel type as appropriate single hydr ulic channels. The thermal performance is evaluated using ENC neucronics 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 indi-cate which fuel type will experience the-least coolant flaw rate and that fuel type is selected for the TM/LP and LC0 calculations.

The limiting assembly calculations model the limiting (highast power) fuel assembly into appropriate subchannels with the assembly flow rate as determied above. The calculation is consistent with the methodology used for the core flow distribution calculations. This calculation determines the limiting subchannel flow rate used in the ensuing MDNBR calculation necessary to establish both TM/LP and LC0 setpoints.

6.2.2 PTS-PWR Code The PTS-PWR code (6) is an ENC digital computer program developed to describe the behavior of pressurized water reactors subjected to

XN-NF-507 g

Revision 1 5 abnormal operating conditions. The model is based on the solution of the basic transient conservation equations for the primary and secondary coolant systems, the transient conduction equation for the fuel rods, of the f. int kinetics _ equation for the core neutronics. The program calculates fluid conditions, such as flow, pressure, quality, heat flux, DNBR, reactor power, and reactivity during the transient. This code is used for thermal-hydraulic safety limits evaluation tc generate LC0's for A00 transients or simply used for checking the behavior of the reactor with existing setpoints during a specified transient.

6.2.3 TM/LP LSSS The TM/LP trip protects the reactor core from exceeding SAFDL when the core pressure, coolant inlet temperature, and power deviate from normal. The procedure used to generate TM/LP trips with the XCOBRA-IIIC models is outlined below.

(1) Select a power. The PAD offset trip defines the limiting axial shapes that are to be considered in the TM/LP trip as a function of power.

(2) Select a reactor pressure.

(3) Select a coolant inlet temperature.

(4) Determine the most limiting axial power profile at the selected power. This is the power distribution which results in the least MDNBR or other appropriate SAFDL value in the XCOBRA-IIIC calculation.

(5) Set up an XCOBRA-IIIC model implementing (2), (3), and (4) as input data.

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XN-NF-507

~ ~

Revision 1 (6) Compare the XCOBRA-IIIC results with the thermal-hydraulic safety limits described in Section 4.2. Return to Step (3) adjusting T IN and repeat the process until the results are equal to the T-H limits.

Steps (1) through (6) establish a single poi-t on a safety

. limit line. Additional data points on that isobar may be obta1ned by repeating these -steps by selecting a new power (and thus a new limiting axial power distribution). Additional safety limit lines (at different pressures) are obtained in the same fashion.

Based upon the results of the above calculations, a set of TM/LP safety limit lines are fit with a function of the form:

P =

var a*PF(B)

• B + 8

• TIN + Y g where the variables have the same definition as given in Section 4.2. The values of a, 6, and y are determined by the procedure given below.

The value of 8 is calculated from two contiguous TM/LP safety limit lines as var 1 -P var 2 g _

'T IN -T IN 1 2 where Subscripts 1 and 2 represent two XCOBRA-IIIC calculations at the same power level but different T IN.

Values of 6 are determined from each pair of contiguous safety 1 init lines, as defined above, and numerically averaged to obtain the final value of 6.

The value of-a is alculated as the product of the slope of a safety line and the appropriate 8 as obtained above, i.e.

a =

• 6

XN-NF-507 I

Revision 1 Several values of a are determined for each safety limit line and numerically averaged to obtain the final value for a.

- The value of y , is obtained by application of the P equation and is chosen to conservatively envelope all safety limit lines. l 1

The family of curves generated in this manner represent a conservative envelope l of the thermal-hydraulic safety limits for all the possible operating pressures at specific power levels and coolant inlet conditions. j 1

The TM/LP trip equation differs from the above expression j for P yy only in the value of the additive constant y:

y = y y+b The value of y (or equivalently b) is selected to protect the SAFDL on DNB during anticipated transients, provide a reasonable administrative operating bond, and to include measurement uncertainties associated with operating conditions such as power, pressure, and temperature.

PF(B) is a linear function which intensifies the variation I of P var with power in the part power operating range. It is incorporated in the TM/LP trip function to describe the effects of the APD LSSS implicit in a l g1 the safety limit lines. PF(B) is constrained:

6B + c  ; 100>B>B g PF(B) = 1.0  ; B>100 6Bg + c  ; BB 1g i Bg is the operating power allowed by the APD LSSS at its lower limiting ASI.

The constraints fix c in terms of 6 The value of 6 is determined as follows:

I

XN-NF-507 Revision 1 (1) Selec; the TM/LP safety limit point (B, Tin}i which mini-mizes the absolute value of the quantity [W/AT) \ computed P

as AT (T100 - Tj )

(a P (100 - Bj ) constant P (2) Using the point (G, Tin)i selected in (1), compute the value of 6 from T +

. JB ,in(B=100) - Tin (B=Bj ) f 100-B; f 2 2 100 -B j

- 100100-B}

where f ie, initially equal to 1.0 and 6 is obtained by writing the equation for P var in difference form on an isobar and rearranging.

(3) Plot the part-power safety limit function determined by the value of 6 computed above, and compare to the TM/LP safety limit points.

(4) Repeat Steps (2) and (3) if necessary, adjusting f between 1.0 and 0.0 to obtain a safety limit function which l conservatively envelopes the safety limit points. The value of v3 may require adjustment to obtain such a function.

The above procedure establishes an appropriate PF(B) which is then compared to RPS limitations on TM/LP for reactor in question. If f

XN-NF-507 I

Revision 1 necessary, the constants 6 and c are adjusted to allow implementation of PF(B) in existing RPS hardware. Such adjustments are made to conservatively envelope the data points.

6.2.4 Verification of the Adequacy of TM/LP - First Reload Cycle The verification of the adequacy of reactor setpoints determined using ENC methodology is accomplished through the analysis of anticipated operational occurences for the plant in question. These analyses are generally performed in a fashion to envelope not only the cycle in question but also future operating cycles. This is accomplished by proper selection of initial reactor core parameters which will affect the plant transient response.

These parameters account for core hydraulics and kinetics. A typical list of transients analyzed is given in Table 5.1.

As previously stated, the ENC methodology for the deter-mination of TM/Lp does not consider any prior knowledge of system transient performance to allow an explicit accounting of dynamic effects such as depres-  !

surization, system time response, trip overshoot, 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 1

against penetrating appropriate SAFDL values during steady-state and anticipated operational occurrences. Thus, any degradation 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 setpoint I methodology includes both steady-state and transient performance evaluations.

I

XN-NF-507 Revision 1 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 system response as well as the core neutronic response anticipated for all cycles. Any appropriate time delays associated with the reactor protection system are explicitly defined in the plant transient analyses and are consistent with the anticipated RPS performance for the plant in question. The uncertainties in nuclear peaking and control system i instrumentation readings are consistent with ENC methodology as previously described. Uncertainties associated with other LSSS trip functions were I

i applied to the nominal value 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.

6.2.5 Verification of the Adequacy of TM/LP - Future Reload Cycles The final TM/LP LSSS function is obtained from writing the expression for P var as:

l P =

l var a* PF(B)*B + 6* Tin + Y where Y Yj + Y u +b The value of b is calculated as y-y)-yu, where uy is an effective trip unit allowance for the most rapid depressurization transient.

The parameter b is defined as the available overshoot equivalent pressure

-( A0EP'), and may be interpreted as the pressure equivalent of thermal margin between the trip point and penetration of SAFDL's.

l

XN-NF-507 Revision 1 The required overshoot equivalent pressure (R0EP) is defined as the pressure equivalent of the maximum observed post-trip thermal margin decay, and represents the least allowable value of A0EP. R0EP is determined via the following steps:

(1) Select an applicable PTSPWR plant transient simulation.

(2) From the results of the selected simulation, calculate AB, the magnitude of the post-trip power increase, deiined as the difference between the peak post-trip power in percent of rated and the power in percent of rated at the time of trip-signal generation.

(3) Calculate AT, the temperature increase over the time interval between trip signal generation and the occurrence of MDNBR.

(4) Calculate an overshoot equivalent pressure (0EP) for the I transient via the TM/LP trip equation:

0EP = a*AB + B*AT (5) Repeat steps (1) through (4) for each applicable PTSPWR transient.

(6) Determine R0EP as the maximum of the 0EP's.

Compare ROEP thus calculated with A0EP. If A0EP is larger than R0EP, the TM/LP LSSS affords sufficient thermal margin to protect SAFDL's. If A0EP is less than R0EP, a new value of y in the above equation for P var must be calculated, and the resulting TM/LP LSSS must be verified.

I

XN-NF-507 Revision 1 6.2.6 Low Flow Trip Setpoints Loss of coolant flow transients are caused by a loss of electrical power to the primary coolant pumps and a corresponding increase in coolant temperature. This increase, combined with the reduced flow, is anti-cipated to reduce thermal margin (MDNBR). The two most severe transients of this type are: (1) the loss of the four primary coolant pumps, and (?) the loss of two primary coolant pumps in opposite coolant legs. Since these transients result in thanges 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 is determined by analyzing the above two loss of coolant flow transients for plant / cycle in question.

6.2.7 LC0 for DNB Monitoring In addition to the loss-of-coolant flow transients, the CEA drop transient results in changes in core power and inlet temperature not covered by the TM/LP function. The CEA drop results in a non-symmetric core 1 power distirubtion generally resulting in no reactor trip. Hence, additional l reactor operating = limits must be determined which consider the calculated ,

1 power increase. associated with the full length CEA drop and which preclude l penetration of the SAFDL on DNBR (required overpower margin).

The CEA drop transient is analyzed using standard ENC pre- l dictive methodology for DNBR(5,14) and a thennal-hydraulic model which is essentially the same used in the determination of TM/LP. The analysis includes s

I.

XN-NF-507 Revision 1 E

E consideration of anticipated axial power profiles for the plant in question, changes in the core power distribution consequent to the CEA drop, and appli-cation of appropriate uncertainties in core and reactor system measurements.

The results of the analysis are used to determine the LC0 for DNB monitoring which represents allowable core power as a function of axial shape index. The LC0 for DNB monitoring thereby provides administrative limits to preclude penetration of the SAFDL on MDNBR during the CEA drop *ransient. The methodology used in dctermining these limits is described below in detail.

(1) Axial power Profile Selection ,

The core average axial power profiles used in the deter- '

mination of the DNB tent were calculated using ENC neutronics methodology (16,17,18) and rep.esent that set of power l profiles anticipated during the cycle in question. The pcwer profiles were sorted according to axial shape index.

Within cach small ASI increment, sufficient axial profiles were selected for analysis to ensure proper determination of the DNB tent.

(2) Definition of Analysis Input l

The establishment of the DNBR analysis input includes I allowance for reactor operating uncertainties, nuclear peaking uncertainties, and changes in peaking during the l CEA drop. In addition, any change in assembly flow as a consequence of changes in peaking is included directly in the MDNBR analysis. The change in assembly flow is deter-mined directly from appropriate core flow distribution l calculations.

i 1

C XN-NF-507

~

Revision 1

{

The averaqe Ll;GR associated with the limiting assembly was calculated according to the following expression:

[,

LHGR*Ff*FU,pAUG

{ LHGR =

F g *Ff*F where, LHGR = core average linear heat generation ra'te

[ -

= peaking limits as defined by the Technical F{ Specifications U

F = peaking measurement uncertainty (FU S1.06)

F AUG

= calculated CEA drop peaking augmentaion Fp = asse ably local peaking factor

( F = calculational factor to account for-changes in nominal hot subchannel area due to allowed manufacturing tolerances ,

U F = core power measurement uncertainty (F = 0.98)

[

(3) Power Interation

[: The limiting hot assembly LHGR is defined as that value which results in a. calculated MDNBR equal to the SAFDL on DNB. The calculated MDNBR includes the engineering heat flux factor as well as a correction factor to account for that fraction of total generated energy which appears as rod surface heat flux.

{

The allowable core power fraction, P, is then defined as the i

ratio of LHGR crit to LHGR ref for each ASI value. The ordered pair (ASI, P) r l-

, XN-NF-507 I

Revision 1 defines a point on the unadjusted DNB tent. The ASI value associated with each such point is then adjusted for neutronics uncertainty and for the excore calibration uncertainty associated with the determination of the ASI value. A typical DNB tent is shown in Figure 6.7.

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4; I

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XN-NF-507 Revision 1 1

4 4

f 1.4

• CECORE Measured x 3-D XTG 1-D XTG 1.2 -

O O 0 y V no 5

. o4 ox Y Y 1.0 o f o

. T 1 L 0 o as t

d' . 0. 8 -

e,

.; +o o.

?$ x

~

! 0.6

. T

" o 0

i 0.4 -

0,2 -

4 t I I l' 1 I I i i i 1 0 1 2 3 4 5 6 7 8 9 -10 11 12~

s

, Top Axial Node Bottom Figure 6.1 Measure'd and Calculated Axial Power d Distributions, 500 MWD /MT

I XN-NF-507 Revision 1 5.0 -

_ g i -

o 1D XTG E X 30 XTG 4.5 - Measured ,

4.0 -

{

1 3.5 _

^

3.0 -

j - .

~

5

.h  ?.5 ,

Ea .

Yn 2.0 - '

s .

l

, 1.5 I

~

3 1.0 -

0.5 _

x .

e f , I f i l i f a i i  ! i l A t 1 i I t i t I i n 0 25 50 /5 0 25 50 75 100 125 0 25 50 75 100 125 Grou; 2 Group 3 Group 4 E

Group 1 5 Rod Group Position - Inches With Drawn Figure 6.2 Integral Rod worth (Regulating CEA Grt ups 1, a 2, 3, 4) vs Rod Group Position g

XN-NF-507 Revision 1 14 -

13 -

I? -

11 -

10 -

t

~.

~

[$ 9

• w 8 -

E

9.

i

/ _

?.'

, 6 -

E t

5 ~

.E 4 -

3 -

2 -

1 ._ l l

l I l l t i l l I I i l

()

0 1 2 3 4 5 6 7 8 9 10 11 12 fop Axial Node Bottom Figure 6.3 Core Average Exposure Distribution, 1-D and 3-D XTG, 9,750 MWD /MTU Core Average Exposure

I

XN-NF-507 i Revision 1 I

I Core Top I

Group 4 Rods Zone 1 (use Group 4 rodded peaking 3

3 factor for Fxy)

I Zone 2 No Rods (use unrodded Peaking Factor For F )

V Core Bottom I

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figure 6,4 CEA lonirig Pattern f or input to CESPT I

XN-NF-507

~01- Revision 1 l 200 d

180' - -

e ico -

. \ s

.y *

\ ,

5' 7., .

ll' 140 l ,

r

• I

~

%' 120 o .

a f

3 100 80 60 1 1 1 1 I I I I l 1

.5~ .4 3 ~.2 .1 0 +.1 +.2 +.3 +.4 +.5  !

i Axial Shape Index, S figure 6.5 Axial Shape Index - LSSS for Centerline Melt llithout Uncertainties b '

l i

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XN-NF-507 Revision 1, l

l i

1 P/> -

I 1 I 160 -

Ae

+

l y

7 4 L

si 143 - *

,p . A u 7, 9 I f

G m

,_ 120 -

r r O ,

L ,

f;' 7 7 h ./

S ,

a 100 -

e s

E 80 -

I 60 -

I i i # 1 e i i '

.5 .4 .3 .2 .1 0 +.1 +.2 + .3 +.5 E i.4 m Axial Sh,ipe Index, S Figure 6.6 Axial Shape Index LSSS for fuel Centerline Melt g With lincertainties g

( XN-NF-507 Revision 1

[. ,,

(

160

+.+ + +

tr -T **

1 +

+ 4 .c

+

4 t t t

(. 150 - ,

t t

,+

+

+ + ++ + +

+4 # + +

4 + x 140 ,. 4

{ c. W*

t

.- +

4.* ., .

j 1 30 --

~.a 3

o

( [j 120 ~

[

110 1

( 100 -- -

[..

90 -.

-80 -40 0 40 S, Percent f Figure 0.7 Iypical DNGR .. Allowable Power vs

. Axial Shape Index, S f

r t

l XN-NF-507 Revision 1

7.0 REFERENCES

1. F. B. Skogen, "XPOSE - The Exxon Nuclear Revised LEOPARD", XN-CC-21, Revision 3 Exxon Nuclear Company, April 1975.

2. W. R. Caldwell, "PDQ7 Reference Manual", WAPD-TM-678, Westinghouse Electric Corporation, January 1965.

3. R. J. Breen, O. C. Marlowe, and C. J. Pfeifer, " Harmony: System for Nuclear Reactor Depletion Computation", WAPD-TM-478, January 1965.

4. R. B. Stout, "XTG - A Two-Group Three-Dimensional Reactor Simulator Utilizing Coarse Mesh Spacing", XN-CC-28, Revision 3, Exxon Nuclear Company, April 1975.

5. XN-75-21 "XCOBRA-IIIC: A Computer Code to Determine the Distri-bution of Coolant During Steady-State and Transient Core Operation",

April 1975.

6. XN-NF-75-5, " Description of the Exxon Nuclear Plant Transient Simulation Model for Pressurized Water Reactors", Revision 1, May 1975.

7. "GAPEX: A Computer Program for Predicting Pellet-to-Cladding Heat Transfer Coefficients", XN-73-25, August 13, 1973.

8. M. F. Lyons, et.al., "U0, Thermal Conductivity from Irradiation with Central Modeling", GEAP-4624, July 1964.

9. H. -Hausner, " Determination of the Melting Point of Uranium Dioxide",

J. Nucl. Mat'1., 15(3) 179-183, 1965.

10. J. A. Christensen, et.al., " Melting Point of Irradiated Uranium Dioxide", Trans. Am. Nucl . Soc. 7(2) 390-391, 1964.

11. J. A. Christensen, et.al. , " Melting Point of Irradiated UO2"'

WCAP-6065.

12. L. M. Petric and N. M. Greene, "XSDRNPM: AMPX Module with One-Dimensional Sn Capability for Spatial Weighting".

13. L. S. Tong, " Boiling Crises and Critical Heat Flux", AEC Critical Review Series, TID-25887, 1972.

14. XN-75-48, " Definition and Justification of Exxon Nuclear Company DNB Correlation for PWR's", October 1975.

i

,l XN-NF-507 Revision 1

15. XN-207 " Power Spike Model for Pressurized Water Reactor Fuel",

March 1979.

16. F. B. Skogen, Exxon Nuclear Neutronics Design Methods for Pressurized Water Reactors", XN-75-27, June 1976.

17. Supplement 1 to Reference 16, September 1976.

18. Supplement 2 to Reference 16, December 1977.

19. D. S. Rowe, " COBRA-IIIC: A Digital Computer Program for Steady-State and Transient Thermal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements", BNWL-1095.

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