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App 15C, Analysis Methods for Steam Line Breaks, to CESSAR Sys 80+ Std Design
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APPENDIX 15C ANALYSIS METHODS FOR STEAM LINE BREAKS l

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O 89040g ppR g h h pko Amendment No. 7 W March 31, 1982

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g EFFECTIVE PAGE LISTINC; CHAPTER 15 APPENDIX 15C Table of Contents Page Amendment j 9 ii 7 Text Page Amendment 15C-1 9 150-2 9 15C-3 9 15C-4 9 15C-5 9 15C-6 9 15C-7 9 15C-8 9 15C-9 9 l 15C-10 9 l 15C-11 9 Tables Table No. Amendment 15C-1 7 15C-2 7 I

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TABLE OF CONTENTS CHAPTER 15 V APPENDIX 15C Section Subject Page No.

15C.1 INTRODUCTION 15C-1 15C.2 MATHEMATICAL MODELS 15C-1 15C.2.1 Primary and Secondary Thermohydraulic 15C-1 Model 15C-2 9 15C.2.2 Nuclear Model 15C.2.3 DNBR Evaluation Methodology 15C-2 15C.3 INPUT PARAMETERS AND INITIAL CONDITIONS 15C-3 15C.3.1 General 15C-3 15C.3.2 Parameters and Conditions for Maximizing 15C-4 Pre-trip Degradation in Fuel Performance 15C.3.3 Parameters ar.d Conditions for Maximizing 15C-6 Post-trip Der.radation in Fuel Performance O

i Background 15C-6 15C.3.3.1 I

15C.3.3.2 Plant Initial Conditions 15C-7 9 '

15C.3.3.3 Analysis Assumptions 15C-8 15C.3.3.4 Single Failures 15C-9 15C.3.4 Parameters and Conditions for Maximizing 15C-10 Secondary System Contribution to Radiological Releases REFERENCES 15C-11 9 l

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APPENDIX 15C l O ANALYSIf METHODS FOR LARGE STEAM LINE BREAKS 15C.1 INTRODUCTION

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This appendix provides a description of methods used in the analysis of nuclear steam supply system (NSSS) response to the steam line break (SLB) events presented in Section 15.1.5. Computer codes and supporting calculational methods used in the analysis are discussed in Section 15C.2. Analysis assumptions which were used to maximize the potential for degradation in fuel cladding performance and to maximize radiological releases are discussed in Section 15C.3.

15C.2 MATHEMATICAL MODELS 15C.2.1 PRIMARY AND SECONDARY SYSTEM THERMAL-HYDRAULIC M00Ei The NSSS response to the steam line break was simulated using the CESEC computer program version described in Reference 4. Major model changes relative to versions of CESEC used for earlier FSAR analyses include: (1)a more detailed reactor coolant system (RCS) thermal-hydraulic model to include the effect of temperature tilt in the reactor core during asymmetric transients and an explicit representation of the reactor vessel upper head region, (2) a reactor coolant pump model which, in combination with an RCS loop momentum model, explicitly calculates the time dependent reactor coolant mass flow rate, (3) a safety injection tank model, (4) an RCS metal heat transfer model, and (5) a three-dimensional reactivity feedback model.

The explicit representation of the reactor vessel upper head region, which 2 nominally receives only about one percent of the total reactor vessel flow, produces a more accurate RCS pressure calculation for those transients which result in steam formation in the reactor vessel, In this region the RCS metal heat transfer model accounts for heat transfer between the upper head region fluid and metal (including the vessel wall and cladding, the upper guide structure and the control element assembly (CEA) guide tube shrouds).

The effect of decreasing reactor vessel downcomer fluid temperature on ex-core neutron detector response during steam line break is explicitly modeled in CESEC through the use of a decalibration factor. Ex-core detector decalibration, which is caused by increased neutron attenuation, delays the occurrence of the high core power reactor trip signal during steam line breaks.

Other reactor trip functions which are creditable for SLBs and which are directly modeled in CESEC include low steam generator pressure and low pressurizer pressure (Section 17 of Reference 4). The reactor trip signal.s which are generated by the core protection calculators (CPCs) are not directly modeled in CESEC. For the SLB analyses which assume a concurrent loss of 9 offsite power, the subsequent reactor coolant pump coastdown would result in a CPC-generated reactor trip not later than 0.6 seconds after event initiation.

Therefore, for these events reactor trip in CESEC is initfated at 0.6 seconds.

For those other SLB transients for which a CPC trip is credited, a CESEC case without trip is used to generate the inputs the CPCs receive (Section O 15C-1 Amendment No. 9 February 27. 1984

7.2.2.2.4). These inputs are then fed to a CPC FORTRAN simulation code. (The functional description of this simulation code is identical to that of the CPCs as given in References 5 and 6. This code implements the functions described 9 in References 5 and 6 and is used in verification of the CPC software, as '

reported in References 7 and 8). The reactor trip time determined by the CPC l FORTRAN simulation code is then used in CESEC for the SLB analysis.

15C.2.2 NUCLEAR MODEL Core power as a function of time is calculated in CESEC using a six delayed group point kinetics model. Moderator, Doppler, boron, and scram rod reactivity contributor are explicitly modeled. The moderator and Doppler reactivity functions are based on two-dimensional PDQ-X (described in subsection 4.3.3.1.1) calculations. Moderator and Doppler reactivities are parameterized as functions of average moderator density and effective fuel temperature, respectively, for use by CESEC, Values used for scram rod worth (with one stuck CEA) as a function of scram rod insertion and for reciprocal  !

boron worth were also calculated using PDQ-X. Reactivity coefficients corresponding to end-of-cycle operation were used for the steam line breaks appearing in Section 15.1.5 to maximize post-trip reactivity insertion.

Calculational uncertainty in the PDQ-X calculations was accounted for through the use of conservative multipliers on the CESEC reactivity functions. j The CESEC three-dimensional reactivity feedback option gives the capability of including three-dimensional reactivity feedback effects associated with core inlet plane temperature distribution, stuck CEA, and changes in core power distribution. The 3-D reactivity contribution is based on HERMITE (described in subsection 4.3.3.1.1) calculations, and is parameterized in CESEC as a function of core inlet plane temperature tilt (difference between hot and cold edge temperatures), core flow, and core fission power. This option was not used for the steam line breaks presented in Section 15.1.5.

15C.2.3 DNBR EVALUATION METHODOLOGY For steam line breaks initiated from full power conditions, pre-trip DNBR in the hot channel was calculated using the TORC computer code discussed in Subsection 15.0.3.1.6, and the CE-1 critical heat flux correlation described in CENPD-162. The initial axial core power distribution was determined by i selecting the most adverse

  • axial shape index (ASI) allowed by the core operating limit supervisory system (COLSS) for steady-state conditions. The allowed axial shapes for steady-state full power conditions were calculated  ;

using the QUIX computer program, which is discussed in Subsection 4.3.3.1.1. l The initial power distribution was used from beginning of the transient until I reactor trip. Average core heat flux, reactor coolant flow rate, RCS pressure, and core inlet temperature from CESEC are provided as input to TORC, The planar radial peaking factor provided as input to TORC corresponds to the most adverse

  • radial peak at the ASI for which the power operating limit (POL) is ,

not exceeded.

l The determination of DNBR for post-trip steam line break conditions requires i methods which differ from those described above. This is due to the fact that the verified range of the CE-1 correlation dose not cover low pressures and low flow ratr,. Therefore the .Macbeth DNBR correlation (References 1 and 2) has been selected to represent margin to DNB during periods of return-to-power.

  • See Section 15C.3.2.

15C-2 Amendment No. 9 February 27, 1984

1 Macbeth correlates critical heat flux to mass flux, inlet subcooling, pressure, heated diameter, and channel length. Application of a channel heat balance allows. the correlation to be converted to a " local conditions" form.

O Using this local conditions form of the correlation, critical heat flux as a function of height in the hot channel (which is located near the stuck CEA location) is calculated, where the effect of non-uniform axial heating is incorporated using the method applied by Lee (Reference 3).

Open core calculations indicate that local quality in the hot channel during l steam line break post-trip return-to-power conditions seldom exceeds a few percent, regardless of fission power rate or core average mass flux. This occurs due to the assembly cross-flow effects. The presence of low density liquid or of voids at the top of the hot channel causes post-trip power 1 generation to occur near the bottom of the core. For return-to-power DNBR I calculations an integrated racial peaking factor of 15 and an axial peaking j factor of 3 are used to bound all possible power distributions unless explicit j 3-D HERMITE calculations are performed. Enthalpy as a function of height is >

comauted by performing a closed channel heat balance. Hot channel inlet entialpy is set equal to the average enthalpy predicted by CESEC for the fluid at the core inlet for that half of the core on the side associated with the )

affected steam generator. Maximum enthalpy is limited to that corresponding 4 to 20% quality at the system pressure, to account for the cross-flow effect.

15.C.3 INPUT PARAMETERS AND INITIAL CONDITIONS l

I 15.C.3.1 GENERAL l

The consequences of steam line breaks are evaluated with respect to criteria j on: I

a. over-pressure l b. fuel performance, and .

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c. radiological releases Steam line breaks are initially depressurization events. During the portion i of the transient after steam generator dryout and before operator action, some '

l repressurization can occur due to safety injection pump flow, decay heat addition, and heat transfer from the hotter walls and structure of the RCS.

However, the emergency feedwater system and the primary and secondary system i safety valves are designed to relieve in excess of the energy available from i these sources while maintaining primary and secondary pressures at, or below, l design pressure. Therefore, input parameters and initial conditions were not chosen to maximize over-pressure for the analyses of SLB initiated transients.

Degradation in fuel performance can occur during SLB initiated events either 1 during the portion of the transient prior to and during reactor trip (henceforth referred to as the pre-trip port;on) or during the post-trip return-to-criticality, or approach-to-criticality, portion of the transient (henceforth referred to as the post-trip portion). Input parameters and initial conditions which maximize the potential for pre-trip degradation in fuel performance are discussed in Section 15C.3.2. Input parameters and p initial conditions which maximize the potential for post-trip degradation in l U 15C-3 Amendment No. 9 February 27, 1984

1 l

l fuel performance are discussed in Section 15C.3.3. The departure from nucleate boiling ratio (DNBR) provides a measure of fuel performance.

Therefore the discussions of potential for degradation in fuel performance will be in terms of those parameters which can decrease local DNBR (i.e.,

degrade fuel performance) for the conditions present in a PWR during SLBs:

a. increase in local heat flux,
b. decrease in coolant flow,
c. decrease in coolant pressure, and
d. increase in coolant temperature.

It there is a potential for degradation in fuel performance such that more than {

a very small fraction (on the order of 0.1% for System 80) of the fuel pins in the core must be assumed to fail, then offsite doses are sufficiently dominated '

by the contribution from primary system activity that assumptions which maximize the potential for degradation in fuel performance also maximize the  !

radiological releases. If there is not a potential for degradation in fuel performance such that more than a very small fraction of the fuel pins in the core must be assumed to ' ail, then offsite doses are sufficiently dominated by the contribution from secondary system activity that assumptions which affect the contribution of the secondary system activity to the offsite dose must be considered. The input parameters and initial conditions which maximize the contribution of the secondary system activity to the offsite dose are discussed in Section 15C.3.4.

15C.3.2 PARAMETERS AND CONDITIONS FOR MAXIMIZING PRE-TRIP DEGRADATION IN FUEL PERFORMANCE Due to the protective action of the core protection calculators (CPCs, Section l9 7.2.1.1.1.4) the pre-trip minimum transient DNBR will be nearly the same for a wide spectrum of steam line break sizes, initial conditions, and analysis assumptions. The CPCs calculate DNBR and generate a trip signal before the DNBR falls below the specified acceptable fuel design limit (SAFDL). The inputs to the CPCs are given in Section 7.2.2.2.4. Neutron flux power and pressurizer I pressure contribute the greatest to the decrease in calculated DNBR and resultant reactor trip during SLB initiated events. The neutron flux power is the dominant of these two contributors, with the contribution of pressurizer l pressure decreasing with decreasing break size. i The CPCs will also generate a trip if RCS hot leg saturation is approached. f l

The CPC inputs used in this calculation are pressurizer pressure and hot leg g i temperature. For many SLB initiated events this trip function will act before a low DNBR trip is generated. However, this trip function was not credited in j the SLB analyses. '

Figure 15C-1 presents the transient minimum DNBR as a function of total SLB flow area. (The total flow area is the steam line break area plus the effective turbine flow area.) Although the minimum transient DNBR is lower than 1.19 for the largest flow areas, the CPC low DNBR trip ensures a minimum transient DNBR such that no more than 0,7% of the fuel pins will be calculated to experience DNB during any outside containment SLB. In the SLB .

transient presented for pre-trip degradation in fuel performance in Section l 15C-4 Amendment No. 9 ,

February 27, 1984  !

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15.l.5 (Case 5), the CPC trip is taken at a' time which illustrates an approach to this limit on fraction of fuel which will be calculated to experience DNB.

The initial conditions chosen for RCS pressure and temperature, core flow, and power are such as (a) to make the initial state near a power operating limit 3 for the values of ASI and radial peaking factors used and (b) to achieve a transient minimum DNBR less than 1.195, thus requiring the protective action of the CPCs. The value of ASI and radial peaking factor, F are chosen to maximize the fraction of fuel pins calculated to experie0c,e DNB for a given transient minimum DNBR. The most negative ASI (-0.3) and the lowest Fo (1.4) were found to yield the largest fraction of fuel pins calculated to experience DNB for a given minimum DNBR. This combination of parameters, together with analysis assumptions which result in the most rapid rute of decrease in DNBR after event initiation, yields the maximum potential for pre-trip degradation in fuel performance. The most negative moderator and the least negative Doppler reactivity coefficients result in the most rapid power increase and thus the most rapid decrease in DNBR prior to reactor trip. The consequent j conservative analysis assumptions which affect pre-trip fuel performance and which were used for the System 80 SLB analysis include:

a) End of equilibrium burnup cycle core conditions to yield the most negative mode stor coefficient.

b) A moderator reactivity versus coolant temperature function determined ,

using the most negative, including uncertainties, technit '

specification mcderator temperature coefficient of -3.5 x 10_4/F at nominal full power conditions, T ave = 594 F (Section 15.0.3.3.2).

c) A 15 percent decrease in the slope of the Doppler reactivity versus fuel temperature function (Section 15.0.3.3.1). g d) Saturated blowdown with no moisture carryover from the steam generators to yield the maximum cooldown rate.

e) Minimum initial steam generator water level to yield the maximum I initial rate of steam generator depressurization and consequently the maximum initial cooldown rate.

Other analysis assumptions which were used for the System 80 SLB analysis, but which have little or no impact on the pre-trip fuel performance include:

f) The CEA of maximum worth stuck in the fully withdrawn position after reactor trip.

g) Maximum pressurizer water level and assumptions concerning heat transfer areas and method of moderator reactivity calculation the same as items (h), (1), and (j) of Section 15C.3.3.3.

For cases initiated from a power operating limit and where loss of offsite power is assumed to occur concurrent with the SLB, there will be a CPC trip on i projected DNBR withi., the first 0.6 second of the initiation of the event.

Thus the CPC trip occurs much earlier in the transient for cases with concurrent loss of offsite power than for cases with offsite power available.

The loss of flow (LOF) due to RCP coastdown causes a more rapid rate of reduction in DNBR for the SLB cases with concurrent loss of offsite power.

15C-5 Amendment No. 9 Februar.y 27, 1984

However the power operating limit is determined such that the CPC trip will prevent the transient minimum DNBR due to a LOF from being less than 1.19.

The only significant additional effect of the SLB, over that of the LOF, up to the time of minimum transient DNBR will be a reduction in RCS pressure.

Therefore for SLBs with concurrent loss of offsite power the transient minimum DNBR will be only incrementally lower than 1.195. Analyses done to determine  ;

the transient ninimum DNBR for SLBs with concurrent loss of offsite power have  !

shown that this minimum DNBR is not less than 1.13. Thus the minimum DNBR for SLBs with concurrent 'oss of offsite power is less adverse than the minimum DNBR for the SLB events with offsite power available discussed in Case 5 of Section 15.1.5.  !

15C.3.3 PARAMETERS AND CONDITIONS FOR MAXIMIZING POST TRIP DEGRADATION IN FUEL PERFORMANCE 15C.3.3.1 Background Degradation in fuel performance during the post-trip portion of SLB initiated transients cam only occur if there is a return-to-power (R-t-P). Therefore i

the primary consideration for maximizing post-trip degradation in fuel performance is to select those parameters and conditions which will maximize R-t-P. The magnitude of R-t-P is primarily determined by the value of the naximum post-trip reactivity, the timing of this reactivity, and the duration ,

of the reactivity peak. (Other parameters which can affect the R-t-P, such as I delayed neutron fraction, have a minor effect within the range of values of the parameters. These other parameters are therefore thosen to be appropriate to the core burnup which yields the maximum transient post-trip total

, reactivi ty. ) The timing of the maximum post-trip reactivity has an important effect on the post-trip R-t-P: the same reactivity will produce less R-t-P I later in a transient since (a) fission power will have decreased to a lower j value nrior to R-t-P, requiring more multiplication to reach a given power level, and (b) the delayed neutron background will be lower, requiring more l reactivity to produce a given, positive rate of change of power. The duration of the reactivity peak is important in that this parameter determines how long the post-trip power will continue to rise (if a R-t-P occurs) before being turned around by decreasing reactivity.

For transients which result in R-t-P, degradation in post-trip fuel cladding performance (measured by the DNBR) is impacted strongly by core flow at the time of R-t-P. Core flow at the time of R-t-P is primarily a function of the analysis assumption on time of reactor coolant pump coastdown. Initial conditions and possible single failures have little or no effect on this core flow. For the range of pressure and temperature involved, the direct effect 4 of pressure and temperature upon post-trip DNBR is small compared with the l impact of these parameters upon fuel performance thrcugh their effect on the magnitude of the R-t-P via the reactivity feedbacks.

l Initial conditions which impact the R-t-P are discussed in Section 15C.3.3.2.

l The effect of analysis assumptions on the R-t-P and the core flow at time of l R-t-P are presented in Section 15C.3.3.3. A discussion of the effect of possible single failures on R-t-P is presented in Section 15C.3.3.4.

O 15C-6 Amendment No. 9 February 27, 190

15C.3.3.2 Plant Initial Conditions The impact of initial conditions on the potential for post-trip degradation in fuel performance is through their effect on R-t-P via the magnitude, timing, anu duration of the oost-trip total reactivity peak. These effects act through their contributions to the moderator reactivity, the Doppler reactivity, and the safety injection boron reactivity.

The ranges of the parameters given in Table 15.0-5 (with the restriction on core inlet coolant temperature given in footnote 2 of that table) were l considered in establishing the most adverse initial plant state for R-t-P.

l (The radial peaking factors given in Table 15.0-5 are not used for post-trip i analysis. See the discussion in Section 15C.2.3.) For System 80 this most i adverse state has been found to be the maximum core power, most positive ASI, i minimum core flowrate, maximum pressurizer water level, maximum core inlet coolant temperature, maximum reactor coolant system pressure, and maximum water level in the affected steam generator with the water level in the unaffected steam generator at the maximum value which can exist initially and still result in emergency feedwater actuation at the time of main steam isolation valve closure (i.e., the transient time of minimum level).

Maximizing the core power and core inlet temperature and minimizing the core flow impact the R-t-P adversely via their effect of maximizing RCS average temperature and core outlet temperature. Maximizing RCS average temperature maximizes the rate of cooldown since it maximizes steam generator pressure.

Maximizing RCS (core) average temperature also causes the cooldown to occur over a more adverse portion of the moderator reactivity function, i.e. the portion having tb greatest rate of change of reactivity with temperature.

, Maximizing core outlet temperature maximizes the energy stored in the water

( and metal of the upper head region of the reactor vessel and also maximizes the saturation pressure of the water in this region. As the RCS pressure falls below the saturation pressure of the liquid in the upper head region, i

the stored energy provides the energy necessary to vaporize this liquid, I

resulting in a low rate of decrease of RCS pressure below the saturation pressure of the liquid in the upper head. This in turn minimizes the safety injection boron reactivity at the time of R-t-P, since the safety injection actuation signal is delayed and the safety injection pump flow is impeded by the higher transient pressures.

Use of the most positive A51 maximizes the delay in insertion of CEA l reactivity following trip. This has little effect on the R-t-P. Maximizing pressurizer water level and pressure maximizes the energy stored in the pressurizer. This maximizes transient RCS pressures, delaying and impeding safety injection flow.

Maximizing steam generator water level in the affected steam generator maximizes the amount of cooldown, thus maximizing the moderator reactivity.

Maximizing the water level in the unaffected steam generator maximizes the amount of steam blowdown from that steam generator before MSIS, since a hiaber  !

initial steam generator water level results in a lower rate of decrease of steam generator pressure causing a lower rate of decrease in steam blowdown flow rate. Thus increasing the initial water level in the unaffected steam generator increases the cooldown due to steam blowdown from this steam generator, also. However, if the initial water level in the unaffected steam 15C-7 Amendment No. 9 February 27, 1984

i generator is above sonta minimum level, the level in this steam generator will not fall below the EFAS low level setpoint during the transient. It has been found that, for System 80, the cooldown provided by emergency feedwater is more than the additional cooldown that would result from initializing the water level at the maximum possible level in the unaffected steam generator.

Therefore cooldown by the intact steam generator is maximized by using the >

maximum initial water level which will result in EFAS at the time of MSIV cloure (the point at which level stops decreasing).

j 15C.3.3.3 Analysis Assumptions The impact of analysis assumptions on the potential for post-trip degradation in fuel performance is through their effect on core flow at the time of R-t-P and their effect on R-t-P via the magnitude and timing of the maximum post-trip reactivi ty.

The analysis assumption which affected core flow at time of R-t-P was the time of reactor coolant pump (RCP) trip. Early RCP trip yields low flow at time of R-t-P and therefore low minimum DNBR. The time of RCP trip (initiation of four-pump coastdown) also affects the magnitude of the R-t-P, primarily via the timing of the maximum post-trip total reactivity -- but also via the magnitude of this reactivity. Higher flows tend to produce a larger R-t-P.

However, the magnitude of the core flow, itself, at the time of R-t-P has more effect upon the minimum DNBR than does the less direct effect via the magnitude of the R-t-P. Therefore, loss of offsite power concurrent with the steam line break yields the greatest potential for degradation in post-trip fuel performance, since the RCPs begin to coastdown at the beginning of the transient. Table 15C-1 shows, for System 80, the effect af time of RCP trip concurrent with the break and at time of SIAS as well as for cases with no RCP trip.

A nunber of analysis assumptions affect the magnitude of the R-t-P.

Conservative analysis assumptions which affect the R-t-P and which were used in the System 80 SLB analysis include:

a) The CEA of maximum worth stuck in the fully withdrawn position after reactor trip.

b) End of equilibrium burnup cycle core conditions to yield the most negative moderator coefficient.

c) Saturated steam blowdown with no moisture carryover from the steam .

generators to yield the maximum energy removal.  !

d) A 10 percent increase for the slopw of the moderator reactivity versus coolant temperature function to assure that the calculation ,

of the reactivity increase due to cooldown of the moderator is I conservative. I e) A 15 percent increase in the slope of the Doppier reactivity versus fuel temperature function to assure that the calculation of the reactivity increase due to cooldown of the fuel is conservative.

f) A 10 percent decrease in the slope of the boron reactivity versus l boron concentration to assure that the calculation of SI reactivity '

is conservative. I 15C-8 Amendment No. 9 February 27, 1934 1

g. The steam line breaks were initiated by a postulated double-ended rupture of one steam line upstream of the MSIV. This break location results in an initial blowdown area' for each steam generator (until the MSIVs close) equivalent to two flow restrictor areas (since there are two steam lines per steam generator). As the MSIVs close, Q steam blowdown from the unaffected steam generator terminates and the effective blowdown area for affected steam generator is reduced to one flow restrictor area (i.e., blowdown through the other steam line for the affected steam generator is terminated by the closed MSIVs). A smaller break delays the time of maximum post-trip reactivity and therefore decreases.the magnitude of the R-t-P generated. '
h. Heat transfer areas in the reactor vessel upper head region were increased by 10% to assure that the heat added by the walls and I structure in this region was conservatively large, causing the transient pressures to be higher and, as a result, the safety injection reactivity to be less.
1) Heat transfer areas in the RCS, other than those in the reactor i vessel upper head region, were decreased by 10% to assure that the heat added by the walls and structure in these regions was conservatively small, causing the RCS cooldown to be increased.-

j) Moderator reactivity was determined as a function of the lowst cold l leg temperature to account conservatively for the effect of uneven temperature distribution on the moderator reactivity. Asymmetric heat removal causes unequal cold leg temperature; at the reactor vessel inlets for the two steam generator loops. Unequal reactor vessel inlet temperatures in combination with incomplete mixing of

( coolant in the reactor vessel downcomer and lower plenum results in a temperature distribution at the core inlet plane. The effect os this temperature distribution is included by basing moderator reactivity on core cold edge moderator density.

15C.3.3.4 Single Failures l

Of the single failures possible for System 80 (Table 15.0-6) only the failure of one MSIV to close and failure of one HPSI pump can affect the potential for l post-trip R-t-P and consequent possible degradation in fuel performance. (The failure of the most reactive CEA to insert and a loss of offsite power are assumed, additionally, for SLB analyses). Whether the additional cooldown provided by the failure of an MSIV on the unaffected steam generator or the decreased safety injection boron reactivity resulting from the failure of a HPSI pump is more adverse for a transient depends upon a number of factors.

In general the failure of a HPSI pump will be more adverse unless transient characteristics (e.g., RCS pressure, tim of R-t-P) are such that little or no safety injection boron reaches the core befcre R-t-P, even when both HPSI pumps are assuned to be operative.

O 15C-9 Amendment No. 9 February 27, 1984

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Table 15C-2 shows the maximum post-trip reactivities, core average powers, and minimum DNBRs with an assumed MSIV failure and with an assumed HPSI pump f ailure for double-ended guillotine SLBs for System 80. Cases are presented for SLBs initiated at full power and at zero power, with and without loss of offsite power. For all cases except the SLB initiated at full power with offsite power present, the HPSI failure produces the most adverse transient results.

15C.3.4 PARAMETERS AND CONDITIONS FOR MAXIMIZING SECONDARY SYSTEM CONTRIBUTION TJ RADIOLOGICAL RELEASES The contribution of the secondary system to radiological releases is maximized by (a) the maximum bitial steam generator inventory in the affected steam generator, (b) a ioss of condenser availability, and (c) the maximum amount of post-accident heat to be removed.

(a) Assuming that the initial steam g merator waite level is at the highest perm:n n,la operating lesal maximizes the potential for radiological release due to the 'ischarge to atmosphere of the contents of the affected stem generator. Further, cases initiated from zero power operating conditions will have the maximum initial steam generator water inventory for a given water level.

(b) A loss of condenser availability (due to loss of offsite power, e.g.) requires that the plant be cooled down by use of the atmospheric dump valves. This causes addition;I radiological releases due to the discharge to atmosphere of water from the unaffected steam generator.

(c) Maximizing the amount of post-accident heat to be removed maximizes the amount of liquid from the unaffected steam generator that must be vaporized and released to the atmosphere (in the absence of condenser availability of (b) above) to achieve cold shutdown. The amount of post-accident heat to be removed is maximized by assuming the maximum initial plant temperature and by assuming that the decay heat to be removed is that appropriate to full power operation, even for cases initiated at zero power: It is assumed that the event occurred at zero power, but that within the previous hsif hour the j plant had been at equilibrium full power conditions. j 1

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15C-10 Amendment No. 9 February 27, 1984

REFERENCES Macbeth, R, V,, "An Appraisal of Forced Convention Burn-out Data,"

O 1.

Proc. Instn. Mech. Engrs, Vol, 180, Pt3c, pp 37-50, 1965-66.

Macbeth, R. V., " Burn-out Analysis - Part 5:

2. Examination of Published World Data for Rod Bundles," A. E. E. W. Report R358,1964.
3. Lee, D, H,, "An Experimental Investigation of Forced Convection Burn-out in High Pressure Water-Part IV, Large Diameter Tubes at About 1600 psia," A. E. E. W. Report R479, 1966,
4. LD-82-001 (dated 1/6/82), "CESEC Digital Simulation of a Combustion Engineering Nuclear Steam Supply System," Enclosure 1-P to letter from A. E. Scherer to D. G. Eisenhut, December, 1981,
5. CEN-147(S)-P, " Functional Design Specification for a Core Protection Calculator, Feb. 81, (Submitted on Dockets 50-361,50-362).
6. Enclosure 1-P to LD-82-039, "CPC/CEAC Software Modifications for System 80," March 1982. 9
7. CEN-217(V)-P, "CPC/CEAC System Phase I Software Verification Test Report," Jan, 83.
8. CEN-219,(V)-P, "CPC/CEAC System Phase II Software Verification Test Report," Jan. 83.

1

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15C-11 Amendment No. 9 February 27, 1984

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c TABLE 15C-1

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( EFFECT OF TIME OF REACTOR COOLANT PUMP TRIP ON MAXIMUM POST-TRIP REACTIVITY, i

CORE AVERAGE POWER, AND DNBR FOR DOUBLE-ENDED GUILLOTINE MAIN STEAM i LINE BREAKS WITH A STUCK CEA AND A SINGLE FAILURE.

Time of Post- Reactor Initial Power Le 'l Reactor- Coolant Trip: Pump Trip FULL ZERO

, 1 l

Maximum

  • core f average power 0 5.5 0.007 I

(% of 3800 MW) l at SIAS 4.1 0.018 no trip 5.1 0.017 Maximum O reac tivity 0 +0.09 -0.64

(% ap) at SIAS -0.28 -0.52 no trip -0.18 -0.19 l

l l

Minimum ONBR 0 2.7 >10 i

at SIAS >10 >10 no trip >10 >10 '

  • or value at time of maximum reactivity, if no return-to-power occurs.

C\

U Amendment No. 7 March 31, 1982

m --

TABLE 15C-2 EFFECT OF SINGLE FAILURE OF MSIV OR ONE HPSI PUMP ON MAXIMUM POST-TRIP REACTIVITY, CORE AVERAGE POWER, AND DNBR FOR DOUBLE-ENDED GUILLOTINE MAIN STEAM LINE BREAKS WITH A STUCK CEA.

INITIAL MAXIMUM POST-TRIP:

POWER OFF-SITE SINGLE LEVEL POWER FAILURE CORE AVERAGE REACTIVITY POWER

(%Ap) (% OF 3800 MWt)

ONE HPSI )

PUMP +0.087 5.5 2.7 )

' l LOSS l OF MSIV -0.005 4.5 >10 l FULL ONE HPSI -0.64 2.4 >10 AVAIL- PUMP ABLE MSIV -0.18 5.1 >10 ONE HPSI -0.64 0.007 >10 0

LOSS PUMP 0F MSIV -1.3 0.007 >10 ZERO ONE HPSI -0.19 0.017 >1G AVAIL- PUMP ABLE MSIV -0.39 0.007 >10 TMaximum value or value at time of of maximum reactivity, if no return-to-power occurs.

j l

Amendment No. 7 March 31, 1982 l

l

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