ML19347F341
| ML19347F341 | |
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|---|---|
| Site: | 07000192 |
| Issue date: | 03/31/1981 |
| From: | Carew J, Diamond D, Neogy P BROOKHAVEN NATIONAL LABORATORY |
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| References | |
| CON-FIN-A-3374 BNL-NUREG-29336, NUDOCS 8105180490 | |
| Download: ML19347F341 (46) | |
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-29336
- L NUREG
- FORMAL REPORT LIMITED DISTRIBUTION A SPATIAL KINETIC ANALYSIS OF A PWR CORE RESPONSE TO STEAM LINE BREAK P. NEOGY, D. DIAMOND AND J. CAREW DATE PUBLISHED - MARCH 1981 CORE PERFORMANCE GROUP DEPARTMENT OF NUCLEAR ENERGY BROOKHAVEN Nail 0NAL LABORATORY UPTON, NEW YORK 11373 Prepored for the U.S. Nuclear Regulatory Comm ssion CMce of Nuclear Reactor Regulation Ccer-act No CE-ACO2-76CMCC016 l
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BNL-NUREG-39336 i
INFORMAL REPORT LIMITED DISTRIBUTICN i
A SPATIAL KINETIC ANALYSIS OF A PWR CORE RESPONSE TO STEAM LINE BREAK P. Neogy, D. Diamond and J. Carew Core Performance Group Department of Nuclear Energy Brookhaven National Laboratory Upton, New York ~,1973 March 1981 Prepared for U.S. Nuclear Regulatory Commission Washington, D. C.
20555 Under Interagency Agreement DE-AC02-76CH00016 NRC FIN No. A-3374
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NOTICE: This document contains preliminary information and was ptepared primarily for interim use.
Since it may be subject to rovision or correction and coes not represent a final report, it should not be cited as reference without the expressed consent of the t
authors.
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ABSTRACT The loss of secondary coolant due to a break in the steam line of a pres-surized water reactor (PWR) is accompanied by a lowering of the reactor cool-ant temperature and pressure. This in turn leads to a reactor trip, usually on low vessel pressure or high neutron flux.
If a high worth control rod is stuck out of the reactor core, the depressurization induced by the accident, together with the power peaking at the stuck rod location leads to a reduction in the Departure from Nucleate Boiling Ratio (DNBR) margin. The response of a PWR core to the steam line break (SLB) accident has been analyzed under these conditions with MEKIN-B to ensure that sufficient margin to DNB 1: maintained at the stuck rod location throughout the course of the transient. The analy-sis was performed under a typical set of conditions rather than under all in-clusive conditions. However, the sensitivity of the DNBR to further depres-surization in the core was investigated with an auxiliary code that utilized the MEKIN-B results and the W-3 correlation to compute critical heat fluxes.
The results of this study show that adequate DNB margin is maintained at the stuck rod location even when the reactor core is allowed to depressurize to 80% of the ncminal operating pressure.
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TABLE OF CONTENTS Page ABSTRACT............................
iii LIST OF FIGURES.........................
v LIST OF TABLES.........................
v I.
INTRODUCTION..........................
1 II.
THE STEAM LINE BREAK ACCIDENT..................
1
- III, THE C ALCULATIONAL MODEL.....................
2 IV.
THE REACTOR MODEL........................
3 V.
STEADY STATE RESULTS......................
4 VI.
TRANSIENT RESULTS....
4 VII.
DISCUSSION...................
5 APPENDIX A...........................
7 ACKN OWL ED GEMENTS........................
9 REFERENCES.....................,.....
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i LIST CF FIGURES i
Figu re 1
Core inlet pressure forcing function 2
Core inlet temperature forcing function 3
Typical core layout - 3 loop PWR 4
Thermal-hydraulic channel representation 5
Local / average channel power at t=0.3 6
Minimum DNBR vs. core height at t=0.0 7
Control rod insertion vs. time 8
Reactivity components vs. time 9
Core thermal power vs. time 10 Coolant temperature vs. time 11 Coolant density vs. time 12 Peak fuel temperature vs. time 13 Local / average channel power at t=4.0 14 Local / average channel power at t=4.5 15 Local / average channel power at t=5.0 16 Local / average channel power at t=5.5 17 Local / average channel power at t=6.0 18 Average channel local / average axial power 19 Average channel heat flux 20 Central channel heat flux 21 Central channel axial power distribution 22 Minimum DNBR in core vs. time 23 Peak heat flux in central channel vs. time 24 Sensitivity of MCNBR to corc pressure LIST CF TABLES Table 1
Time sequence of events for the SLB accident 2
Steady state conditions and scram characteristics 3
Hydraulics model correlation options used 4
Stuck rod location MDNBR and core-w1de MONBR at different times
-v-
I.
INTRODUCTION This analysis was performed to remove some of the uncertainties asso-ciated with the power distribution at the stuck control rod location in a PWR scram.
If the scram is initiated by a steam line break (SLB) accident, the depressurization induced by the accident, together with the power peaking at the stuck rod location, leads to a reduction in the Departure from Nucleate Boiling (DNB) margin. The main objective of this analysis was to detemine whether sufficient margin remained during the entire course of the transient to prevent DNB at the stuck rod location.
To achieve this, the response of a PWR core to the SLB accident was com-1 puted with MEXIN-B, the BNL version of the tnree-dimensional, coupled 2
nuclear /themal-hydraulic core transient code, MEKIN.
Since MEKIN-B models the reactor core alone, it was necessary to use pressure and temperature forc-ing functions at the inlet of the reactor core (obtained frem a plant tran-sient analysis of the SLB accident) to simulate the effects of the SLS on the primary coolant system. Plant transient analyses usually employ a point ki-netics model for the reactor neutronics.
The present analysis, therefore, can be viewed as the second step in an iteration process.
The steam generator response, obtained in the first step, is utilized to detemine a more realis-tic and more detailed core transient behavior. The use of core inlet forcing functions, and the expensive nature of the MEKIN-B calculations required that the analysis be performed under a typical set of conditions rather than under all inclusive conditions. However, having established depressurization as the key mechanism for reducing the DNBR during this transient, tha sensitivity of the DNBR to further depressurization in the core was investigated with an aux-iliary code that used the W-3 correlation 3 and MEKIN-B results to compute critical heat fluxes.
Section II of the report contains a description of the SLB accident.
The calculational model and the reactor model used in this analysis are described in Sections III and IV. The results of the steady state and transient calcu-lations are given in Sections V and VI. A discussion cf the results obtained appear in Section VIf.
The auxiliary code, CHFW3, used to study the sensitiv-ity of the MDNBR to depressuriz: tion is described in Appendix A.
II. THE STEAM LINE BREAK ACCIDENT The loss of secondary coolant due to a break in a steam line between the steam generator and the turbine causes a cecrease in steam pressure, and thus places a demand on the control system for increased feedwater flow. Becau se the primary and secondary systems are coupled through the steam generators, the increased feedwater flow and the stemt flow through the break and the turbine stop valves lower the reactor ccolant temperature and pressure. The reactor trips usually on icw vessel pressure or high neutron flux depending on the break size. The forcing functions for core inlet pressure and reactor coolant temperature used in this analysis are shown in Figs.1 and 2.
They are characteristic of a double-ended steam line rupture between a once-through steam generatt and the main steam line isolation valve, and were obtained i
f 4 of the SLB accident.
from a Babcock and Wilcox Standard Safety Analysis The time sequence of events leading to the reactor trip obtained from that analysis is given in Table 1.
A description of the steady state operating conditions and reactor trip characteristics is given in Table 2.
III. THE CALCULATICNAL MODEL The calculational model in MEKIN-B comprises a three-dimensional neu-tronic model, a thermal-hydraulic model, and a feedback model 1!nking the two.
In the neutronic model, the two-group neutron fluxes and delayed neutron pre-cursor concentrations are related through the transient, source-free neutron diffusion equations and the delayed neutron precursor equations. The reactor volume is assumed to be composed of homogeneous regions of identical size.
Each reactor region is subdivided into an integer number of three-dimensional mesh volumes. The semi-discrete fonn of the neutron diffusion equations and delayed neutron precursor equations are obtained by integrating the equations over each mesh volume, and by dividing by the volume of the mesh increments.
The semi-discrete equations are solved by the non-symmetric alternating direc-tion explicit method. Although up to six delayed neutron precursgr grcups can be used in MEKIN-B, one effective delayed neutron precursor group 3 was used in the present calculation to allow a finer neutronic mesh to be input. The effective delayed neutron fraction used was 0.00738, and the effective decay constant used was 0.4353 sec-1 A decay heat model is provided in MEKIN-B and was used in the present r
cal cul ation. The reactor is assumed to have been operating at steady state for time to, and the fraction of the steady state power due to decay pro-cesses is given by 0.0603 [(0.1)-0.639 - (0.1+to)-0.639),
The decay power remains constant at its initial value until 0.1 sec after time t*, when the reactor power has decreased to a fraction PALFAP of its initial value. Thereafter, the decay power fraction is computed from the expression
' (t-t*)-0.0639 - (0.1+t )-0.0639 o
(0.1)-0.0639 - (0.1+to)-0.0639 PALFAP=0.04833 and t =107 sec were used in the present calculation to a
simul ate near-equilibrium conditions.
The thennal-hydraulic model in MEKIN-B is a modified version of the sub-6 l
channel analysis code CCBRA-IIIC. The reactor core is divided into a num-i l
ber of interecting channels, each channel represanting a fuel assembly.
The l
axial and radio ? power distributions in the reactor are calculated from
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the neutron flux and are used as input in the thermal-hydraulics segment.
The equations of conservation of mass, energy and mcmentum are solved to give the enthalpy ano flow distribution througtcut the reactor.
The temperature pro-file in the fuel is calculated from the heat generation rate, the heat trans-fer coefficient between the clad and the coolant, and other related para-meters. From this information, the coolant temperature, the coolant density and the metal temperature are calculated and passed back to the neutronics.
The thermal-hydraulic mesh consists of the vertical channels divided into equal axial intervals of lengths. When channel conditions are calculated, a
" smeared model" is assumed, giving, for each axial node, the average coolant conditions and the average metal temperatures for the channel. The hydraulics model correlation options of MEKIN-B used in this calculation are presented in Table 3.
Temperature-dependent correlations for the fuel rod material proper-ties have been introduced into MEKIN-B and were used in the present calcula-tion. The DNBR canputation scheme of COBRA-IIIC has been incorporated in MEKIN-B. The Westinghouse W-3 correlation was utilized in the present calcu-l ation to determine DNBR. A local peaking factor of 1.20 was used to allow for power peaking in the fuel pin and axial peaking within an axial node.
The feedback model in MEKIN-B is determined by the representation of the two-group cross-sections as a function of the average fuel temperature, the coolant temperature, the coolant density, the control state and the equili-brium xenon concentration. The Doppler feedback is introduced thrcugh a term proportional to the square root of the fuel temperature.
The moderator feed-back is represented by a linear function of the coolant temperature and a quadratic function of the coolant density.
The moderator density feedback is allowed to be different for controlled and uncontrolled conditions.
IV.
THE REACTOR MODEL The reactor model used in this analysis is identical to one that has been 7
used earlier in a Rod Ejection Accident calculation with MEKIN-8. A b-ief description of the model will be included here for the sake of completeness.
The model used is a three-dimensional, quadrant symmetric model of a typical three-loop PWR at the beginning of life.
The reactor core layout is presented in Fig. 3.
The location of the control rod banks are marked "R" in the diagram.
The location of the stuck rod, the central channel (H-8), is denoted "S".
There are 15715x15 fuel assembif es with three distinct enrichments ar-ranged in a checkerboard pattern.
Fuel types A and B contain UO2 enriched to 1.85 and 2.55 w/
respectively; while fuel types C and C' are both en-riched to 3.10 w/o.o,Each type B and type C' fuel assembly contains 12 burn-able poison rods.
A quadrant of the reactor core is partitioned into 47 thermal-hydraulic channels (Fig. 4) and 18 axial thermal-hydraulic planes. The axial reflectors are not explicitly included, their effects being represented by :uitable albe-do boundary conditions applied at the upper and lower boundaries sf the active core. Each thermal-hydraulic region is farther partitioned into 9 finer neu-tronic mesh volumes, resulting in horizontal and axial mesh spacings of 7.197 cm and 10.200 cm, respectively. Cross-sections were generated at reference I
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and off-reference conditions, varying the fuel temperature, riodgrator density and control fraction, using the collision probability code, CPM. The va-lidity of the nuclear data used was tested by comparing steady state predic-tions to operating data taken under steady state conditions.
V.
STEADY STATE RESULTS Steady state convergence was achieved using a pointwise flux convergence criterion of 10-5, and convergence criteria for regionwise power and keff of 10-4 and 10-5, respectively. The radial power distribution at steady state is indicated in Fig. 5.
The n:ximum radial power occurs in the central channci, W11ch is the potential stuck rod location. The utal power distribu-tions at steady state in the average channel and in the central channel a*e represented by the curves marked "0.0SEC" on Figs.18 and 19, respectively.
The axial distribution of core-wide minimum DNBR at steady state is given in Fig. 6.
A global :ninimum DNBR of 2.274 occurs at a height of 96.4 inches in channel 2, adjacent to the central chnnel.
VI. TRANSIENT RESULTS A neutronic time step size of.005 see was used for the first 4 sec of the transient. This is a comparatively quiescent period in the transient characterized by slow reactivity insertion through a reduction in the modera-tor temperature, and a slow rise in reactor power. The reactor trip is ini-tiated at 4.2 sec, and the control rods, with the exception of the stuck rod, are completely inserted by 6.0 sec.
A plot of the con rol rod insertion as a function of the time after trip is given in Fig. 7.
The heat production be-yond 6.0 sec is essentially due to decay heat production alone.
A time step size of.0025 see was used in the period from 4.0 sec to 6.5 sec. The tran-sient calculation was terminated at 6.5 sec.
Fig. 8 shows the total re6ctivitv, and its components due to the modera-tor temperature, moderator density, Ooppler and scram, as a function of time.
Beyond 4.2 secs, scram is the dominant reactivity insertion mechanism; while, at earlier times, a small positive reactivity is inserted through the modera-tor temperature and mcderator density components. Note that beyond 5.9 see the total reactivity is shown to be slightly more negative than the scram reactivity alone. This is a consequence of the fact that MEXIN-B uses an ini-tial, approximate adjoint flux weighting scheme rather than an exact adjoint flux weighting scheme in the reactivity calculation.
Figure 9 shows a plot of the core thennal power as a function of time throughout the transient. The acak power is reached at 2274 MW, or at 113.7%
of the rated thermal power. This peak power is consistent with the 112% high flux trip setpoint assumed in the SLB plant transient calculation.
No trip delay time was assumed in the present analysis.
The wiggles in the reactivity curve are produced by the approximate way in which cross-sections are calcula-ted for a particular controlled axial node in MEXIN-8.
The cross-section for such a node is taken to be a weighted mean of the controlled and uncontrolled..
cross-sections, the control fraction being used directly as the weighting factor. Figs.10 and 11 show the core-average coolant temperature and density as functions of time. Fig.12 shows the peak fuel temperature as a function of time. Since a " smeared thermal-hydraulic model" is assumed in this calcu-lation, the peak fuel temperature is the fuel temperature of the average rod in the hot channel, rather than that of the limiting rod in the hot channel.
Figs.13 through 17 show the variations in the radial power distribution with time during scram. As expected, the stuck rod location receives an in-creasing share of the core power as the reactor scrams.
Fig.18 shows the change in the average channel axial power shape during the course of the tran-sient, while Fig.19 shows the corresponding changes in the average channel heat flux distribution. Fig. 20 shows the central channel heat flux distribu-tions at various times during the transient. The corresponding axial power distributions are given in Fig. 21. The axial power profile changes markedly during the scram, indicating the effect of the control rods on the axial power distribution at the stuck rod location. Note, too, the large reduction in the absolute power density in the enannel as the reactor scrams.
The variation of the minimum DNBR in the core with time is shown in Fig.
21.
The smallest value of PUNBR,1.978, occurs at 4.30 sec in channel 2.
The values of the MDNBR at the stuck rod location together with the values of the core-wide MDNBR at different times during the transient are listed in Table 4.
As indicated in the table, the core-wide MDNBR occurs throughout the transient in channel 1, the channel containing the stuck control rod, or the adjacent channel 2.
The variation in MONBR with time is best understood when Fig. 22 is studied in conjunction with Fig.1, which depicts the variation of pressure in the core and Fig. 23, which shows the variation of the peak heat flux in channel 1, with time.
During the first 5 sec, the peak heat flux shows only small changes. The change in MDNBR is detarmined mainly, therefore, by the depressurization in the core. Beyond 5 sec, the peak heat flux decreases rap-idly as the reactor scrams, leading to a rap 1d rise in MDNBR.
VII. DISCUSSICN In this analysis, the detailed response of a PWR core to the SLB accident was determined in order to study the power distribution at the stuck rod loca-tion and its effects on the DNCR margin. The results show that al though there is an increase in the relative radial power at the stuck rod location during the scram, the absolute magnitude of the peak heat flux at the stuck rod loca-tion decreases sharply in the same period, leading to an increase in DNBR.
In this early phase of the transient, therefore, reduction in DNBR is associated with depressurization in the core, and not with the high power peaking factors present at the stuck rod location.
If the highest worth rod is stuck in its fully withdrawn position after the reactor scrams, there is an increased prob-ability that the core will become critical and return to power at a later time. A return to power following SLB is a potential concern mainly because of the high power peaking present at the stuck rod location. This phase of the transient has not been considcred in the present study.
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This study shows tr at a core-wide MDNBR of 1.978 is reached 4.3 sec into the transient, a 13% reduction from the steady state MDNBR of 2.275.
It is clear from Fig.1 and Fig. 22 that the value of MDNBR reached during the tran-sient is sensitively dependent on the time at which scram is initiated.
A de-lay in the initiation of reactor scram beyord 4.2 see would lead to a further reduction in MDNBR owing to the progressive depressurization of the core. The two possible mechanisms initiating reactor scram that have been considered in the plant transient analysis of the SLB accident are high neutron flux trip and low reactor system pressure trip. The latter occurs typically at 80% of ncminal operating system pressure.
Fig.1 shows that a low pressure trip would occur at a point in time beyond 6.5 sec. The time at which the reactor scrams due to high neutron flux (the actual trip mechanism used in the plant transient analysis) is dependent upon the rate of reactivity insertion dus to the cooling of the moderator; i.e., upon the moderator temparature coeffi-cient. The plant transient analysis is usually carried out at end of life conditions with the moderator temperature coefficient having a large negative val ue. This leads to conservatism in the analysis of the latter part of the transients which is characterized by possible recriticality and return to powe r.
However, during the early phase of the transient, a large negative moderator temperature coefficient leads to a large positive reactivity inser-tion rate, and hence an early reactor scraa. A conservative evaluation of the MDNBR during this phase of the transient requires that the reactor core be al-l lowed to depressurize to 80% of the nominal operating pressure befor N high neutron flux trip becomes operational.
In order to investigate this point further, it was necessary to determine l
the sensitivity of the MDNBR to depressurization, holding the other parameters constant. A stand alone program, CHFW3 (see Appendix A), was written to com-pute the critical heat flux and DNBR using the W-3 correlation and ut lizing thermal-hydraulic parameters calculated by MEKIN-B. To ensure conservatism, l
the values of the coolant mass velocity, channel enthalpy, inlet enthalpy and l
heat flux used were those calculated by MEKIN-B for the stuck-rod channel at l
4.2 sec. The reference pressure was varied from the nominal value (2250 psia) to 80% of the nominal value (1800 psia). The results are presented in Fig. 24 as the calculated MDNBR as a function of the relative core pressure.
The value of MDNBR at 1800 psia was found to be 1.795, indicating that adequate l
DNB margin is available even when the core is depressurized to 30% of the L
l nominal pressure.
I It is necessary to emphasize that the results obtained hold for the given set of conditions, namely, full power operation with control rods fully with-drawn prior to tne transient, and core depressurization to 80% of the ncminal pressure during the transient.
The expensive nature of the MEXIN-B calcula-tions makes it impractical to do an exhaustive set of parametric studies to
(
include all possible initial conditions.
It is conceivable that a different l
set of initial conditions, e.g., full or partial power operation with control rods at their insertion limits and the stuck scram rod at an off-center loca-tion, would lead to a transient with more severe consequences, j
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APPENDIX A Program CHEM 3 CHFW3 is a FORTRAN code to calculate the critical heat flux and the DNBR utilizing the W-3 correlation and several thermal-hydraulic parameters as calcu-lated by MEKIN-B fcr the channel of interest. The input parameters used are Oh and D, the equivalent heated and wetted diameters of the channel (in feet);
g pref, the reference core pressure (psia); G, the coolant mass velocity 1b/hrxft ); L, the channel enthalpy (BTU /1b'); Hin, the iniet enthalpy 2
(106 H
2 (BTU /lb); and F, the heat flux (MBTU/hrxft ) at the axf al nodes of interest.
The equivalent uniform critical heat flux is given by q0NB,EU/106 = (2.022
.0004302xpref)+(0.1722-0.0000984xpref) x exp((18.2 - 0.004129xp,f)xX) p x((0.1484-1.596xX+1.729xXlXl)xG+1.037) x (1.157-0.869X)x(0.2664+08357xexp(-37.812D )
h x (0.8258 + 0.000794x(H -Hin))'
f X = (L '"f)/Hfg where h
H
= H -H f
g f H = enthalpy of the vapor at p g
ref Hf = enthalpy of the liquid at pref If (1-O /D ) is positive, an unheated wall correction factor, F, is g h 7
applied to the equivalent uniform '
.ical heat flux:
1=(1.-Rg(13.76-1.372xemp(1.78X)4.732/G.0535 F
.0619(p
/1000.)0.14-11.101xD
))
ref h
where R = 1.0 - O /0 u
g h Finally, to account for non-uniform heat-flux distribution, the equiva-lent uniform critical heat flux is divided by an axial shape factor, Fj:
F = T(J,J )xexp(-Cx(J )/12.)/(F(J )/.0036)/
2 3 x x
x (1.0-exp(-Cx(X(J )-x(J -1))/12.))
x s
Jx where T(J,,J,) =
(F(J)/.0036)x(exp(CxX(J)/12.)-exp(CxX(J-1)/12.))
J=Js C = 1.8x(1.0-X)4.31jg.478 0
J = Axial node of interest x
J = Axial node just above the critical boiling location s
X(J) = Elevation of the M axial node If there is no boiling in the core, F2 = 1.0.
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a In the course of this study, a logic error was uncovered in the subroutine that calculates the critical heat flux in COBRA-IIIC and in MEKIN-B, speci-fically in the calculation of F,
The magnitude of this error is small for 2
the axial nodes of interest. The results presented in Table 4 and Fig. 22 are affected by this error, the values of the MONBR shown being 1". to 2".
too small. The error has since been corrected in MEKIN-B.
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i ACXNOWLEDGEMENTS i
We wish to thank Mr. M. Dunenfeld of the USNRC for participating in many helpful discussions, and for his encouragement throughout the entire course of this work. Thanks are due to Mr. H. Richings of the USNRC for a critical re-view of the manuscript. The expert assistance and advice of A. L. Aronson in all aspects of this work are gratefully acknowledged.
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REFERENCES 1.
A. L Aronson, H. S. Cheng, D. J. Diamond, and M. S. Lu, "MEKIN-B: The BNL Version of the LWR Core Dynamics Code MEKIN," BNL-NUREG-28071, June (1980).
2.
B. A. Zolotar, "MEKIN:
Nuclear Reactor Core Kinetics Code (Revision 1),"
EPRI Interim Report, March (1978).
3.
L. S. Tong, J. Nucl. Energy, 21:
241 (1967); and L. S. Tong et al.,
Chem. Eng. Progr. Symp. Ser.,T2 (64): 35 (1965).
4.
Babcock and Wilcox Company, Standard Safety Evaluation Report, B-SAR-205 (Rev 0), FebnJary (1976).
5.
D. J. Diamond, "Use of One Delayed Neutron Precursor Group," BNL Memoran-dum, June 27,(1979).
6.
D. S. Rowe, "CCBRA-IIIC:
A Digital Computer Program for Steady State and Transient Thennal-Hydraulic Analysis of Rod Bundle Nuclear Fuel Elements," BNWL-1695, March (1973).
7.
H. S. Cheng, A. L. Aronson, J. W. Herczeg, and D. J. Diamond, "The Use of MEKIN-B for Light Water Reactor Transient Calcul ations," BNL-NUREG-28785, November (1980).
8.
W. J. Eich, " Advanced Recycle Methodology Program, Research Project 118-1, Electric Power Research Institute (1976).
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P00R831BINM PWR CCRE RESPONSE TO STERM LINE EREAK CCRE INLET PRESSURE FORC[NG FUNCTICN FIGURE 1 l'.00 0.30 -
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FWR CORE RESPONSE TO STERM LINE BREAK THERMAL-HYCRAULIC CPPNNEL REPRESENTATION FIGURE 4 to 46.00 47.33 9
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5 17.00 18.00 19.C0 20.C0 21.00 22.00 23.C0 4
9.000 10.C0 11.00 12.00 13.00 14.00 15.C0 16.00 3
1.000 2.0C0 3.0C0 4.CCO 5.000 6.000 7.000 8.000 3
4 5
6 7
8 9
to X O! RECT!CN CCCRDINATC i
l
[
l l
l l
f FWR CORE RESPONSE TO STERM LINE SRERK LCCPL/RVERPGE CHANNEL PCWER RI I-C.C FIGLRE 5 10 0.383 0.575 3
1.011
!.C:
0.349 0.527 a
1.101
!. C71 - '7.9?!
0.979 C.525 E
E I
E 7 1.133 1.127 1.065
- 0. 365.,
0.796 0.526 5u E
C5 1.120 1.153 t.129 1.000 0.366 0.37'3 0.527 3'
l 0.343l s
i.ise i.:n i.is2 1.::2,:.Ces
- 0. 3n l
l l
4 1.223 1.2CE I.134 1.153
- 1. 27 1.37!
1.012 0.5 5 3
1.224 1.223 1.133 1.130 1.139 1.101 1.01:
0.953 3
4 5
5 7
9 10 X Ol*C07:CN 00090!NAIC 1
I l
J 4
2::
c.c BW"M C%l Cl3 C2:ll CD C3 PNR CORE RESPON$E TO STERM LINE 3RE K MINIMLM CNSR VS. CCRE FE!GHT AT T-0.0 FIGURE 6 5.0
- 5. 5 -
- 5. 0 -
x
- 4. 5 -
E 2.
E
- 4. 0 -
E O2 C
- 3. 5 -
- 3. 0 -
/
- 2. 5 -
l 2.0 0.1 0.2 0.3 0.4 0.5 0.5 0.7 0.3
- 0. 9 l'. 0 NCRPPLIZEC CCRE PCIGHT l
l t
l l
\\
%+
h FWR CCRE RESFCNSE TO STERM LINE SRERK CCNIRCL RCO INSERI!CN VS. TIME FIGURE 7 14.0
. --e 12.0 -
7 10.0-b
~
- 9. 0 -
g d
E
- 6. 0 -
z E
4.0 -
Ea
- 2. 0 -
0.0 0.0 0.2 0.4 0.6 0.'
1.0 1.2 1.4 1.5 1.9 IIME RIIER IRIP (SECI I
u 42 CJ3 N
CD E
CD C3 PWR CCRE RESPCNSE TO STE?M LINE SRERK Ekm REFCTIVITY CCMPCNENTS VS. TIME FIGURE 8 0.01 c.
1 1
-c. t -
4
%x 5
-c. 02 -
- u9
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m H=y
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- 2. 0 2.s 3.a 2.s 4.0 4.s s.a s.s s.c ss a.c o.s TIME INTO TRENSIENTtSEC:
i I
i i
l i
e 6
PhR CCRE RESPCNSE TC STEPM LINE SREPK CCRE THERPFL PChER VS. TIPE FIGLRE 9 2sco.3
- ccc.c-
-xr
.e N
1sC0.0 -
cc.
J5 EuI IC00. 3 -
i--
Uxou sec.a -
'w 3.3
~
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~
a.c o.s T!PE [NTO TRRhS!E'4T 3EC!
J 42:ll C.D M
C""3 M
Cll3 Cll3 CL PWR CCRE RESPCNSE TO STERM LINE BREAK CCCLRNT TEMPERATURE VS. TIME ISU?E 10 se2.3 sec.c-s7s.c -
L._
e-r 575, _
L, 6=
W=
s74. : -
c S
S 572.3 -
ueC s70.0 -
\\
ru>C see.c -
uzoU ses.3 -
t ss4.a -
j se:.c i.o t.s 2.0 2.s 3.0 3.s 4.a 4.s s.a s.s s.a s.s a.c o.s l
TIME INTO TRPNSICITiSEC1 1
b L
r L
4>>
k' " '
.;4 s.
L sJ e = =%.r-.(
., 1.=.v.
--r
~~
L L t",e
.a r e. F rJ 4 r. A U
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ii i v.
/t CCCLANT CENSITY VS. T!*C FIGURE 11 4s.o f-44.5 -
5u N
@q 44.s-
=
~
44.4 -
=w t
44.2 -
2 5eu 44.0 -
U=z t;
43.9-
>c
?J 5
i3.5 -
u 43.4 i.a 1.5 2.3 2.5 3.c
- .5 4.c 4.s s.:
s,s s.:
s.s c.c c.s
- c..a.: : :., c.cc 1 o.r
- NLG, LL L
L...<...es.
w
N/
FWR CORE RE5FCNSE TO 5 TERM LINE SRERK PE?K FUEL TEMPERATUPE VS, TIME FIGURE 12 24e0.0 2463.3 -
244c.:-
Cg :42c.3 -
x c.
E 24CO. 3 -
=W
$ 23ec.a -
a u 2:sa,3 -
2 s 2:4 0.
u O
23:c.a -
2:co.: -
2:50.a
~
$.0 5'. 5 6'. 0 5.5 0.0
- 6. 5 1.0 l'. 5 2.3 2.5 3.0 3.5 4.0 4.5
~
TIME !NTO TPFh3:E'4T(SEC; s
f >' n o n ',I n einq P00R ORIGINAL.
PWR CORE RESPONSE TO STERM LINE SRERK LCCFL/*vCRPGE CHPNNEL PC'ER AT T-4.0 n
FIGURE 13 10 0.**5 0.575 3
1.013 1.015 0.950 0.527 9
1.1C2 1.072 0.393 0.881 0.576 3
5 4 7 1.109
- 1. 25 1.CE6 0.355 O.796 3.5 5 b
a 55 1.179 1.157 1.129 1.053 0.955 0.881 0.537 Y
5
~
5 1.195 1.192 1.151 1.129 1.066 0.193 0.350 1
1.220 1.203 1.192 1.157 1.!!S 1.072 1.015 0.675 3
!.221 1.220 1.195
- .179 1.139 1.102 1.013 0.885 3
4 5
6 7
3 3
10 x O!ROOTION 00080!Mi0 j
i i
l I
l 1
b i
.o,AR CO.:,
,=, r g.o O N e t.
1 n $ T. r..=.i,.
-.:4-
- c.,mc.a.K ne s
i m.
LCCFL/RVERAGE CPANNEL PC'ER AT T-4.5 a
FIGURE 14 10 0.994 3.581 9
1.019 1.013 0.350 3.5 9 9
1.102 1.;E9 0.194 0.!E5 0.529 u
5 5
7 1.132 1.123 L.0E2 0.355 0.aC0 0.6:9 5
l 55 1.174 1.150 1.125 1.C55 0.368 [ 3.385 0.529 si E
~
5 1.190 1.;91 1.161
.125l1.052 0.334 3.350 l1.!$0 1
4
. 221 1.205 1.131
- .123 1.C69 1.C19 3.581 3
!.225 1.22!
1.130
- . 74 1.132 1.
- 02 1.31?
0.994 3
4 5
6 7
9 9
iC x ClaCOTICr4 00CPO!M T0 l
l t
l l
l l
ft.
P00R ORIGINAL PWR CORE RESPONSE TO STEPM LINE BRERK LCCFL/AVCRAGC CHFNNCL PC'nCR AT I-3.0 FIGLRC 15 to 0.915 0.594 9
t.0:E
!.03 0.947 0.543
,8 1.102 1.C61 0.995 0.!94 0.549 I
~
7
!.!!r 1.113 1.C50 0.370 0.!C9 0.648 5
U$
t.!$2 1.134 1.I19 1.043 0.970 0.!94 0.643 t
5
~
5 1.175 1.!!9 1.151 1.119 1.050 0.995 0.347 4
1.220 1.213 1.199 1.134 1.!!3
- 1. ?.61 1.03 0.694 1
3 1.240 1.2 C 1.175 1.162 1.!!6 1.102 1.0:5 1.316 3
4 5
6 7
8 3
( Ol*C TIOra COCRCI W C
~
- +-
Rto,0Lc-10 S i tni i L I s.-
- c n~.s,
-y P,A CO0-,st r
t nt K K
LCCFL/RVERAGE CHFNNEL PCWER AT T-5.5 FIGURE 15 10 0.330 0.740 9
1.084 1.059 0.323 0.555 s
1.095 1.02:
0.390 0.902 0.582 g
5 5g7
!.C59 1.080 1.0C2 0.377 0.342 0.632 c
5 55 1.130 1.CSI 1.103 1.000 0.977 0.322 0.555 w
f I
~
5 1.103 1.199 1.173 1.103 1.002 0.390 0.323 4
- .222 1.250
- .199 1.0S!
1.080 1.02t 1.059 J.740 i
0.390 l 3
!.317 1.292 1.108 1.120 1.059 1.095 1.084 1
I 3
4 5
5 7
8 3
10 X O!PC00:CN COCPOINPTC l
l l
)
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.r..
P00R ORIGINAL
~
~
RWR CORE RESRONSE TO STERM LINE SRERK LCCFL/AVEFFGE Cl-PNNE FC'ER AT T-9.0 n
FIGLFE 17 10 1.299 0.349 9
1.263 1.1:3 0.771 0.7:7 8
1.041 0.314 0.346 1.043 0.*55 si=
C$7 3.9C5 0.325 0.799 1.020
't.010 0.855 C
_E.
G5 0.395 0.857 1.047 0.3 5 1.020 1.043 0.717 N
l s
~
5 0.371 1.258 1.255 1.047 0.739 0.346 0.773 4
1.1:3 1.492 1.258 0.357 3.325 0.814 1.!!9 0.949 3
1.585 1.533 0.37 0.395 0.*C5 1.041 1.253 1.298 I
I 3
4 5
6 7
3 3
13
- CISCOTION C00R0lrAit u
l
/
s g" 6 ilu 3
PWR CCRE RESPCNSE TO STERM LINE SRERK RVERPGE C.HANNEL LCCPL/RVERPGE AxIPL PChER r!GURE iS
~
2.s LE3 ENC
- - 6.0 SEC o.:;; u ::E_,-
ac 4-
- t..
- - 4. 5 5E:
- 2. 0 -
.5.. u c. r. -
x.
.. c..- c.c-a..
O
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C' t.s -
r:9 6
o
?
/
/
tC-Cd
.1 a.s -
xs N
~
\\
c.c C0 3.t c.2 a.3 a.4 a.s a.s a.-
a.3 a', 3
', o NCRPAL[ZEJ CCRE nE;g-7 k
i I
b r
l 1
5
)v.4' 8
7 PODR ORIGINAL
- i e
Ir,.
s
- r.
i
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Cc a.r. i, p c a r s Lc r T.'V q r.:s i a
M g
pi p w.i.
.N 6-i et.
RVERPGE Ch?hNE'. rE?T rLUX
- lC-URE 13
~
0.:c LE3 ENC
[_=
N~
- - 5.C SEC o.:..
c.C.r.
fN
- 1:8 Hi
- 3. :5 -
s 4. C :..
e-N
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=
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cr.:
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q x
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/
s F.
/
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x W
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- 0. :0 -
L.2
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5;2 8'
L
>C 0.05 -
3.00 0.0 0.1 0.2 0.3 0.4 0.5 0.5 0.7 0.9 0.3
- .0 slL..uPa.
r.m.rc.e u.r. [ tw.
ac.
i s
.s
P'WR CCRE RESPCNSE TO STERM LINE SREAK CENTRPL CHPNNEL MEPT ILUX FIGLRC 20 0.40 g
LEGEND g
- - 5.0 SE:
o= 5.5 SE:
0.35 -
Y a - 5.0 SEC
[
- - 4. 5 SEC E
\\
" - 4 C SEC 5
- 0. :0 -
5 X
- - 0.] 5E:
i y
- 0. 2s -
xo&
,/
n
~
n
- 0. r -
c
^
w 0.15 -
j O
\\
- 0. t 0 -
W
- t E
50.0s-l
+
u
- 0. 0 0.0 0.1 0.2 0' 3 o'. 4 0'. S 0'. 6 0.?
0.6 0.3 1.0 NCFFALIZED CCRC PCIGHT l
t l
l l
l i
1 l
I i
l l
g
.m +
P00R ORIGINil PWP. CORE RESPON?C TO STERM LINE BRERK CENTRPL Ch8NNEL ;iXIAL POWER DISTRIBUTICN FIGURE 21
~I LEGENO o - S.0 SEC o - 5.5 SEC 4 - 5.0 SEC 3,a.
+ - 4. 5 SEC x - 4.0 SEC
'N o - 0.0 SEC
/
- 0. 8 -
N g
e N
e
/
\\
$_ 0
.5-
/
5 8
C.
- 0. 4 -
- 0. 2 -
-c 0.0.
0.2 0.3 0.4 0.5 0.6 0.7 0.9 0.9 1.0 0.0 0.1 NCRMPLIZED CORE HEIGH!
i
.. c5 P00R ORIBlHR
~
PWR CCRE RESPONSE TO STERM LINE SREEK MINIMCM CNSR IN CCRE VS. TIME FIGURE 22 2.7
- 2. 6 -
2.5 i
a-
- 2. 4 -
E C
E
- 2. 3 -
2=
O E
- 2. 2 -
2.1 -
^
t
- 2. 0 -
I.9 I
0.0 0'. 5 l'. 0 1.5 2'. 0 2'. 5 3'. 0 3'. S 1". 0 4'. 5 5'. 0 5'. 5 6'. 0 6.5 I
TIME INTO TRENSICNT ($CCl l
E i
i l
r
)
{
I i
~~
~
POOR ORIGINy PWR CCRE RESPCNSE TO STERM LINE SRERK PERK FEAT FLUX IN CENTRPL ChPNNEL VS. tit'E FIGURE 23 0.40 0.33 -
5 0.38 -
0.37 -
E_
y 0.36 -
d
?y 0.35 -
\\
0.34 -
\\
0.33 1.5 2.0 2.5 3.0 3.5 4.0 1.5 5.0 5.5 6.0 6.5 0.0
- 2. 5 1.0 TIME INTO TRENSIENT (SEC)
=
i i
i i
i i
i i
j PWR CORE RESPONSE TO STEAM i
~
LINE BREAK
}
2*5 SENSITIVITY OF MDNBR TO CORE PRESSURE FIGURE 24
~. _ -
x w
z 2.0 o
3 1.5 1
I I
I I
I i
1.0 0.950 0.900 0.850 0.800 RELATIVE CORE PRESSURE e
.. m
- - -,,,,, ~, - - -
m,
l 1
i i
Table 1.
Time Sequence of Events for the SLB Accident Event Time (sec)
Main steam pipe nJpture 0.0 Law steam generator pressure reached for main steam isolation valve closure 3.2 High flux trip setpoint reached 3.8 Control rods begin to drop 4.2 I
i
Table 2.
Steady State Conditions and Scram Characteristics A.
Steady. State Reactor Conditions Core Thermal Power (MW) 2200 Core Burnup Beginning of Life Average Fuel Temperature 1353.1'F Inlet Coolant Temperature 546.2*F Average Coolant Temperature 581.4*F Average Moderator Density 43.54 lbs/ft3 Control Bank Position All Rods Out Soluble Baron Concentration (ppm) 825 Equilibrium Xenon Included Cecay Heat Included System Pressure (psia) 2250 Inlet Flow Rate (106 lbm/hr) 101.5 B.
Reactor Scram Characteristics Reactor Scram Time 1.80 see Scram Acceleration 226 cm/sec2 Stuck Rod Location Central Channel l
r i
l l
Table 3.
Hydraulics Model Correlation Options Used Mixing Coefficient:
3 = 0.020 x Re0 Single Phase Friction:
F = 0.184 x Re-0.2 + 0,0 Two Phase Friction Multiplier:
Hemogeneous Theory Void Fraction:
No Subcooled Void; Slip. Ratio = 1.0 i
l
~
.a.
Table 4.
Stuck Rod Location MONBR and Core-Wide MONBR at different times.
Numbers in parenthesis indicate the channel number and the axial l ocation.
MONBR MONBR Time (sec)
(at stuck rod location)
(core-wide) 0.0 2.311 (1.96.4")
2.275 (2,96.4")
1.0 2.085 (",
)
2.062 (",
)
2.0 2.097 ( *,
)
2.075 (",
)
3.0 2.086 (",
)
2.065 (",
)
4.0 2.031 (",
)
2.014 (",
)
1 4.1 2.014 (",
)
1.999 (",
)
4.2 2.003 (",
)
1.988 (",
)
4.3 1.992 (",
)
1.978 (",
)
4.4 1.985 (",
)
1.985 (1,96.4")
4.5 1.997 (",
)
1.997 (",
)
4.6 2.005 (",
)
2.005 (",
)
4.7 2.013 (",
)
2.013 (",
)
4.8 2.023 (",
)
2.023 (",
)
4.9 2.036 (",
)
2.036 (",
)
5.0 2.055 (",
)
2.055 (",
)
5.5 2.209 (1,88.4")
?.198 (2,88.4")
6.0 2.447 (",
)
2.447 (1,88.4")
6.5 2.701 (",
)
2.701 (",
)
1 1
DISTRIBUTION LIST U.S. Nuclear Regulatory Connission H. Denton W. Dircks M. Dunenfeld (2)
D. Fieno (7)
W. Johnston R. Minogue T. Murley H. Richings D. Ross L. Rubenstein Public Document Room Bethesda Technical Library Advisory Committee Reactor Safeguards (16)
Brookhaven National Laboratory W. Y. Kato Core Performance Group DNE Associate Chairmen Nuclear Safety Group Leaders External T. Anderson, W S. Bian, B&W -
D. Lanning, MIT R. Mills, CE G. Owsl ey, ENC B. Sehgal, EPRI G. Sherwood, GE J. Taylor, B&W B. Zolotar, EPR!
l
_.