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NUREG/IA-0038STUDSVIK/NP-88/101

InternationalAgreement Report

Assessment of TRAC-PF1/MOD1Against an Inadvertent FeedwaterLine Isolation Transient in theRinghals 4 Power PlantPrepared byA. Sjoberg

Swedish Nuclear Power InspectorateS-61182 NykopingSweden

Office of Nuclear Regulatory ResearchU.S. Nuclear Regulatory CommissionWashington, DC 20555

March 1992

Prepared as part ofThe Agreement on Research Participation and Technical Exchangeunder the International Thermal-Hydraulic Code Assessmentand Application Program (ICAP)

Published byU.S. Nuclear Regulatory Commission

NOTICE

This report was prepared under an international cooperativeagreement for the exchange of technical information: Neitherthe United States Government nor any agency thereof, or any oftheir employees, makes any warranty, expressed or implied, orassumes any legal liability or responsibility for any third party'suse, or the results of such use, of any information, apparatus pro-duct or process disclosed in this report, or represents that its useby such third party Would not infringe privately owned rights.

Available from

Superintendent of DocumentsU.S. Government Printing Office

P.O. Box 37082Washington, D.C. 20013-7082

and

National Technical Information ServiceSpringfield, VA 22161

NUREG/IA-0038STUDSVIK/NP-88/101

InternationalAgreement Report

Assessment of TRAC-PF1/MOD1Against an Inadvertent FeedwaterLine Isolation Transient in theRinghals 4 Power PlantPrepared byA. Sjoberg



March 1992

Prepared as part ofThe Agreement on Research Participation and Technical Exchangeunder the International Thermal-Hydraulic Code Assessmentand Application Program (ICAP)

Published byU.S. Nuclear Regulatory Commission

NOTICE

This report documents work performed under the sponsorship of the SKI/STUDSVIKof Sweden. The information in this report has been provided to the USNRC

under the terms of an information exchange agreement between the United Statesand Sweden (Technical Exchange and Cooperation.Arrangement Between the United

States Nuclear Regulatory Commission and the Swedish Nuclear PowerInspectorate and Studsvik Enerigiteknik AB of Sweden in the field of reactor

safety research and development, February 1985). Sweden has consented to thepublication of this report as a USNRC document in order that it may receive

the widest possible circulation among the reactor safety community. Neither

the United States Government nor Sweden or any agency thereof, or any of theiremployees, makes any warranty, expressed or implied, or assumes any legal

liability of responsibility for any third party's use, or the results of such

use, or any information, apparatus, product or process disclosed in this

report, or represents that its use by such third party would not infringe

privately owned rights.

STUDSVIK NUCLEAR STUDSVIK/NP-88/101

1988-11-10

Anders Sj6berg Swedish State Power BoardProject BMP - 860409 GS

Swedish Nuclear Power InspectorateProject 13.3 - 735/88 - 88114

ICAPASSESSMENT OF TRAC-PF1/MODI AGAINST AN INADVER-TENT FEEDWATER LINE ISOLATION TRANSIENT IN THERINGHALS 4 POWER PLANT

Abstract

An inadvertent feedwater line isolation transientin a three loop Westinghouse PWR has been simu-lated with the frozen version of the TRAC-PFl/MOD1computer code. The results reveal the capacityof the code to quantitatively predict the differ-ent pertinent phenomena. For accurate predictionsof the system response it was found that a carefulnodalization of the steam generator downcomerwas essential for accurate pressure distributionand associated level prediction. Also the coremoderator temperature reactivity coefficient wasrecommended to be decreased somewhat (be lessnegative) whereas the fuel gap conductance shouldbe rather low in order to increase the initialstored energy of the fuel rods. It was also foundduring the course of the calculations that somerestrictions had to be imposed on the allowablemaximum timestep size in order not to violatethe convergence of the solution procedure.

ISBN 90-7010-120-5

Copyright Studsvik AB, 1989 Approved by.1

iii

STUDSVIK NUCLEAR STUDSVIK!NP-88/101

198B-11-10

List of contents

Executive summary

Page

1

2

33.13.23.3

4

55.15.25.2.15.2.25.3

6

Introduction

Plant and transient description

.Code and model descriptionPrimary system nodalizationSecondary system nodalizationControl system and trip logicmodeling

Steady state calculation

Data comparisonBoundary conditionsResults from the simulationBasecase simulationSecond stage simulationsCode performance

Conclusions

1

3

5

88

10

11

13

151517172124

25

27

28

References

Figures

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Executive summary

A TRAC-PFI/MODI simulation has been conducted toassess the capability of the code to predict afeedwater line isolation.

The measured data was obtained from an inadvertentfeedwater line isolation at full power operationin the Ringhals 4 power plant. Ringhals 4 is aWestinghouse PWR with three loops and two turbinesof Stal-Laval design. The nominal thermal poweris 2 775 MW and 915 MW electrical. It is equippedwith three Westinghouse steam generators model D3with a feedwater preheater section located atthe cold leg side of the U-tube bundle and adivision is made of the feedwater flow betweenthis lower feedwater inlet and the top inlet atthe upper part of the downcomer.

During the pretransient stationary phase thetotal feedwater was apportioned so that about10 % of the flow was delivered to the top inletand the rest to the preheater. The circulationratio at this condition was about 2.43.

The transient was initiated by a failure in anelectronic logical circuit causing the feedwaterline isolation valves to close in all three loops.Following the closure of the valves the steamflow through the feedwater preheater train ceasedand a corresponding increase of the flow throughthe turbine was obtained. This was automaticallycompensated for by the throttling of the turbinevalves. The impulse chamber pressure of the tur-bines was as a consequence decreased by about10 per cent. This was felt by the control logicof the turbines as a corresponding load rejectionresulting in deblocking of 25 per cent steamdumping capacity.

Because of the loss of main feedwater flow theaverage temperature of the primary coolant wasincreased while the reference temperature wasdecreased due to the reduced impulse chamberpressure. This deviation resulted in a dumpdemand signal and about 14 s after the feedwaterisolation steam dumping from the turbines wasinitiated.

The continued steam flow resulted in depletionof steam generator liquid inventory and reactorscram was obtained on low downcomer level signal.Isolation of the turbines was activated as wellas initiation of auxiliary feedwater supply. Thelevel was as a consequence slowly increased andfinally resumed normal value.

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In the TRAC-simulation only a single loop rep-resentation was used and the core was modeledby the TRAC neutron point kinetics specifiedwith middle-of-cycle conditions. The completemodel comprised 37 components made up by 144 nodeswith the boundary condition components excluded.

The boundary conditions were either taken di-rectly from the recordings of the plant computeror were inferred from these. The following condi-tions were used:

Main feedwater flow and temperaturevs time

Steam flow vs time, trip controlled

Scram reactivity vs time, trip controlled

Auxiliary feedwater flow and temperaturevs time, trip controlled

Decay heat

The pressurizer control system was modeled indetail and also the trip logic for the scram,steam flow and auxiliary feedwater flow.

The result of the simulation revealed a satis-factory agreement with measured data. However,for accurate predictions of some basic and im-portant parameters the obtained result indicatedareas of model improvements. Because of sensitiv-ity of system response on steam generator down-comer conditions a thorough nodalization of thispart was essential. Also modifications in themoderator temperature reactivity coefficient aswell as the gap conductance of the fuel improvedthe results. Finally some changes in the orig-inally assumed long term steam flow boundaryconditions were found to be beneficial for theoutcome of the simulation.

From the run statistics it was found that a 310 stransient without any restarts used 1 068 time-steps with a maximum allowable timestep size of0.5 s. This required 4 784 CPU-s on a CDC Cyber180-835 computer.

It was observed that when letting the code freelydetermine the timestep size an abnormal termina-tion was obtained after about 160 s transienttime because of convergence failure of the numer-ical solution procedure. This was remedied bydecreasing the maximum timestep size to 0.5 s.

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1 Introduction

This study was conducted under the auspices ofthe Swedish State Power Board and the SwedishNuclear Power Inspectorate. Analytical assistanceduring preparation of the input deck and earlyphases of the steady state analysis, was obtainedfrom Mr F Pelayo, Consejo de Seguridad Nuclear,Spain and was much appreciated.

The Interhational Thermal-Hydraulic Code Assess-ment and Application Program (ICAP) is beingconducted by several countries and coordinatedby the USNRC. The goal of ICAP is to make quanti-tative statements regarding the accuracy of thecurrent state-of-the-art thermal-hydraulic com-puter programs developed under the auspices ofthe USNRC.

Sweden's contribution to ICAP relates both toTRAC-PWR [Ref 13 and RELAP5 [Ref 23. The assess-ment calculations of TRAC have earlier been car-ried out as a joint effort between the SwedishState Power Board (SSPB) and Studsvik AB whereasthe RELAP5 calculations have been conducted byStudsvik for the Swedish Nuclear Power Inspecto-rate (SKI).

Recently a Swedish group was formed for coordina-tion of Swedish efforts within ICAP. This grouphas representatives from SSPB, SKI and Studsvikand has emphasized the importance of using planttransients for assessment purposes. Accordinglythe Swedish future efforts will basically concen-trate on analyzing plant transients with theTRAC-PF1 code. The current assessment matrix isshown in table 1.

In this report the results of an assessment ofTRAC-PFI/MODl against a main feedwater line iso-lation transient are presented. The ability ofTRAC to simulate this transient is assessed bycomparison to measured data from an inadvertentisolation occurrence at 100 per cent power inthe Ringhals 4 power plant.

This report is organized as follows: Section 2describes briefly the Ringhals 4 power plant andthe transient which originated from the feedwaterline isolation. In section 3 the TRAC model usedto simulate the transient is described and sec-tion 4 is a review of the procedure used toobtain the specified steady state. Section 5presents the results from the simulation as wellas performance of the TRAC-code. Also some run

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statistics are given. Conclusions are presentedin section 6.

Table 1

ICAP Assessment Matrix - Sweden.

Code Facility Type Description

TRAC-PF1 Ringbals 4 Integral, Full loadfull scale rejection

TRAC-PFI Ringhals 2 Integral, Inadvertent steamfull scale line isolation

valve closure inone loop

TRAC-PF1 Ringhals 4 Integral, Symmetric loss offull scale feedwater

TRAC-PF1 SPEC Intergral, Symmetric losssmall scale of feedwater

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2 Plant and transient description

Ringhals 4 is a 3-loop, 2 turbine PWR of Westing-house-Stal Laval design. The power is nominally2 775 MW thermal and 915 MW electrical. It isequipped with three Westinghouse steam generatorsmodel D3 with a feedwater preheater section lo-cated at the cold leg side of the U-tube bundle.The feedwater is divided between the top feed-water inlet, which enters into the upper part ofthe downcomer, and the preheater section at thelower end of the riser. Normally only a smallerpart of the total feedwater flow is delivered tothe top inlet and the rest to the preheater.

In the preheater section the flow is apportioneddue-to the flow restrictions in support platesetc. According to specifications at nominal loadand no top feedwater 54.5 % of the flow is fedto the upper part of the riser (U-tube boilersection) while the remaining flow is fed downwardand enters the lower end of the riser on thehotleg side where it mixes with the downcomerflow. The circulation ratio at this condition isspecified to be 2.43.

Prior to the transient the plant was operatingat full power with about equally loaded turbines.Because of malfunction of an electronic logicalcircuit a spurious feedwater isolation signalwas created resulting in closure of the feedwaterline isolation valves to all three steam gener-ators. Thus the feedwater flow ceased but stillthe reactor and the turbines were operating atfull load.

As the high and low pressure feedwater preheatersare in series with the feedwater trains the steamdrainage from the high pressure turbines throughthe preheaters decreased when the feedwater flowceased. Consequently the steam flow through theturbines increased as well as the power. Thiswas automatically compensated for by throttlingthe turbine valves (the control oil pressure wasdecreased). The impulsechamber pressure of theturbines was then decreased by about 10 per cent.This was felt by the control logic of the tur-bines as a corresponding load rejection resultingin deblocking of 25 per cent steam dump capacity.

When the feedwater flow ceased the steam gener-ators were less efficient as heat sinks for theprimary side and as a consequence the averagetemperature of the primary coolant was increased.

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The primary side reference temperature was de-creased because of the lower impulsechamber pres-sure of the turbines. Thus a deviation betweenthe primary side average temperature and thereference temperature was obtained resulting ina dump demand signal and about 14 seconds afterthe feedwater isolation, steam dumping from bothturbines was initiated.

By continued steam flow (about 100 per cent)from the steam generators the downcomer levelsdecreased and about 37 s after the feedwaterisolation reactor scram was obtained due to ex-treme low level in steam generator 2. This re-sulted also in turbine trip as well as turbineisolation. The other two steam generators ex-perienced about the same level decrease.Auxiliary feedwater flow was initiated immedi-ately after scram and two motordriven pumps(steam generators 1 and 2) and one turbine drivenpump (steam generator 3) were started. Duringthe next 40 seconds the steam flow slowly de-creased to about zero flow as a result of de-creased dump demand.

Indicated narrow range downcomer level was reach-ing zero almost immediately after scram and aminimum level of about 20 per cent on the widerange level indication was obtained approximately78 seconds after feedwater isolation. The levelwas from then on slowly increased due to continuedauxiliary feedwater flow.

The faulty equipments were replaced and about14 hours after scram the reactor was criticaland after another 10 hours full power was resumed.

Throughout the transient important plant signalswere monitored and stored on the plant computer.From the plant signal follower, which recordsthe time sequence of trips and control signals,and pertinent time plots a sequence of eventswas obtained. This is shown in table 2, leftcolumn, where all the recorded time points havebeen shifted so that the time 10 seconds corre-sponds to the time when the feedwater isolationsignal occurred.

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Table 2

Sequence of Major Events

Recorded Calculated Eventtime (s) time (s)

0.0 0.0 Pretransient stationary

operation

10.0 10.0 Feedwater line isolation

12.1 12.1 Throttling of turbine valves

13.3 - Deblocking of 25 per cent steamdumping

23.5 23.5 Initiation of steam dumping

47.4 47.9 Reactorscram due to low-lowlevel in SG2

47.5 47.9 Trip of both turbines

47.5 49.3 Initiation of auxiliary feed-water flow

88.1 77.1 Minimum wide range SG level

92.0 94.0 Closing of steam dump valves

138.6 80.0 Extreme low primary T-average(< 289.4 0 C)

143.3 - Extreme low primary T-average,condition P12: Blocking ofsteam dump, however, manualbypass allowed to provide max25 per cent dump if necessary

426.4 Extreme low pressurizer level(< 15 per cent): Let-downisolation

1 367 Manual initiation of steam lineisolation to prevent additionalcooldown of RCS

1 800 The plant resumed normal zero-load condition

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3 Code and model description

The simulation of the transient was made withversion 14.0 of the TRAC-PF1/MOD1 computer code[Ref 1) with an additional update to provideproper functioning of the restart capability ofthe core component. The program was run on a CDCCyber 180-835 computer under the NOS 2.6 operat-ing system with no SCM and LCM partition of thememory. Instead the central processor primarymemory was used together with an extended memorycapability. TRAC was also locally modified toallow writing of signal variables and controlblock output on a separate file for later plot-ting with a separate program. Thus the EXCON andTRAP auxiliary programs were not used for pro-ducing the graphics to this simulation.

In the simulation only a single loop representa-tion was used as shown in figure 1. Differencesbetween the three loops were considered to pro-duce effects of secondary order during the trans-ient. It should be noted that trip margins areusually dependent on conditions in individualloops. For instance, the'reactor scram on lowsteam generator level was initiated by the con-ditions in steam generator 2, though closelyfollowed by the other two. Also the turbinedriven auxiliary feedwater pump provided abouttwice as much flow to steam generator 3 as themotor driven pumps to steam generators 1 and 2.Thus the symmetry between the loops was not per-fect and this will have some influence on thecomparison between calculated and measured re-sults.

The basic TRAC-model of Ringhals 4 nuclear steamsupply system for this analysis is depicted infigures 1 and 2. Only minor modifications havebeen introduced compared to the model used inthe loss of load transient [Ref 3). A new pres-surizer model has been implemented though. Thenodalization comprised 17 components with 81 nodeson the primary side and 20 components with63 nodes on the secondary side making a total of37 components with 144 nodes if the boundarycondition components are excluded.

3.1 Primary system nodalization

The reactor core, denoted by component 5, wasdivided into seven vertical nodes; five nodesrepresenting the active core and one unheatedinlet and outlet node respectively. The axialpower distribution was preserved throughout the

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transient. Default point kinetics together withreactivity feedback with middle-of-cycle charac-teristics were used to simulate the neutronicresponse of the fuel during the transient. Thedecay heat was calculated according to ANSI 5.1.

The upper plenum (component 6) and hot leg inletwas divided into three nodes. The hot leg andsurge line, denoted 710, was represented by atee-component with five nodes in each branch.

The primary side of the steam generator was mo-deled by 18 nodes; one each for the inlet andoutlet plena and sixteen nodes for U-tube bundleand the thermal interaction with the secondaryside.

The cold leg leading from the steam generator tothe vessel inlet was represented by 10 nodes;five on each side of the recirculation pump. Thevessel inlet section was modeled by three nodes,and the downcomer and lower plenum were representedby two nodes and one node, respectively.

The recirculation pump was set up to also provideheating power to the coolant. As the energy dissi-pation terms in the conservation equations areomitted no conversions of mechanical energy (wallfriction and pump pressure) into coolant internalenergy are accounted for when passing throughthe loop. A crude compensation for this was intro-duced by adding a specified power to the pumpcomponent wall structure.

The pressurizer was modeled according to recom-mendations given in the TRAC User's Guide [Ref 4).The bottom of the pressurizer was modeled byusing a pipe component divided into four cellsto assure proper draining and accurate pressureloss computation (component 400). The length ofthis component was specified to equal the lengthof the electrical heaters and the heater powerwas assumed to be deposited directly in the fluid.The main body of the pressurizer was modeled asa tee component number 410. Six cells were consid-ered reasonable to simulate the pressure transi-ents and level behavior. The side tube at thevery top of this component was used to modelconnections to the pressure relief and safetyvalves. The top hemisphere of the pressurizerwas represented by a "prizer" component number 420.One feature of this component was to serve as apressure boundary.condition during the steadystate calculations.

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The spray flow was simulated by attaching a fillcomponent to the upper end of the "prizer" com-ponent. The corresponding junction flow area wasspecified such that the liquid velocity was 4 m/sat a spray flow rate of 19.4 kg/s. This wouldactivate the enhanced interfacial condensationmodel in the "prizer" component and thus allowedfor adequate condensation of vapor when a reason-able spray flow was maintained.

The pressurizer walls were simulated by heatstructures with four radial nodes. The heat lossesto the environment were chosen so that they bal-anced the steady state heater power when a speci-fied spray flow was maintained. The losses werethen about 178 kW.

All the pressurizer valves were sized, as sug-gested in ref 4, to their rated capacities underchoked flow conditions. The pressurizer pressurecontrol was modeled in detail and was earliertested in connection with a Ringhals 2 analysis[Ref 5]. Although the level control was modeledit was bypassed for this specific transient.

3.2 Secondary system nodalization

An outline of the steam generator base model isdepicted in figure 2. Feedwater was supplied tothe steam generator at two locations. Ten percent of the feedwater was supplied as top feedinto the downcomer and ninety per cent into thepreheater section near the outlet of the coldleg side of the U-tube. All the auxiliary feed-water was supplied to the top feed connectionthrough the dummy tee-component 745 as shown infigure 1.

The hot leg boiler and the explicitly modeledpreheater section were both divided into fivenodes with the node boundaries located at thesame elevations. The U-tube boiler was representedby three nodes. An ideal separator which allowedonly steam to escape upwards was assumed at thetop of the riser. A connection was made betweenthe separator node and the upper separator drainflow path. The downcomer was represented by to-tally ten nodes.

The steam generator level measurements representedby differential pressures, were explicitly modeledin order to estimate dynamic contributions fromflow and mass content. Both the narrow range andthe wide range level indications were simulated.

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The steam line was divided into a number of tee-components and the secondary pressure (steamline pressure) was measured in the tee-compo-nent 752, figure 1. Also safety and relief valveswere connected to the steam line. According tothe report of the isolation occurrence none ofthese valves were activated during the transient.However, in the model the activation logic wasmodeled. The part of the steam line denoted by753 was represented by a valve component withtwo nodes simulating the main steam isolationvalve.

The steam flow was measured by means of a dif-ferential pressure between the steam dome pres-sure tap and a tap in the relief valve header.In order to avoid disturbances from the flowcontraction between the steam dome and steamline (junction 750) the noding in the very firstpart of the steam line was made somewhat moredense than elsewhere. This is according to recom-mendations given in ref 4.

The steam flow was divided into two streams -one for each turbine - in the steam line header(754) which was represented by three nodes. Theline for each turbine was further split into twoflow paths - one containing the turbine valveand the other containing the dump valve.

During the transient there was no stem positionmeasurement of the dump and turbine valves. Thusit would have been difficult to accurately evaluatethe involved control logic and operation of thesevalves and the corresponding influence on theflow. For simplicity the actual operation ofthese valves was not simulated in the currentanalysis. In the analysis the dump valves remainedfully closed throughout the transient whereasthe turbine valves were fully opened. The flowresponse of the actual valve operations was simu-lated by providing the measured steam flow asboundary conditions downstream the turbine valves.

3.3 Control system and trip logic modeling

The TRAC control system and trip capability wasused to initiate the transient and to introduceseveral actions from auxiliary- and safetysystems. In this way it was possible to make thecalculation dependent on the initial event only,that is the isolation of the feedwater line.

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The following systems were simulated:

Pressurizer pressure control

Trip logic based on level indication ofsteam generator downcomer. The occurrenceof narrow range low level subsequentlyresulted in:

Reactor scram

Turbine trip

Initiation of auxiliary feed-water

The pressurizer pressure and level control isschematically shown in figure 3. This system wastaken from an earlier made Ringhals 2 analysis[Ref 5]. It was found during that calculationthat unphysical oscillations in the output ofthe PI-controller occurred regularly. By replac-ing the PI-controller by the equivalent set ofcontrol blocks this problem was eliminated.

It was also apparent that due to the explicitnessof the control block numerics, the efficiency ofthe control system depends upon timestep size,particularly if the rate of change of the variablebeing controlled is large.

Both the steam generator narrow range and widerange level were simulated by means of calcula-tions of differences between cell pressures with-out any compensation for vapor columns. The narrowrange level was calculated from the pressuredifference between cells 731 (4) and 738 (1),figure 2. The wide range level was similarlycalculated from pressures in cells 731 (1) and738 (1). Scaling factors were obtained first byequating the narrow range signal with the corre-sponding collapsed level during steady state.Next, also during steady state, the wide rangelevel should be equal to the narrow range level.These scaling factors remained constant through-out the simulation.

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4 Steady state calculation

Prior to the transient simulation the TRAC modelwas adjusted to replicate the plant stationarypre-transient conditions. This had earlier beendone for the current power level (100 %) in con-nection with analysis of the load rejection tran-sient [Ref 3). Only minor changes had to beintroduced for this analysis.

A heat balance calculation of the plant duringthe stationary phase provided valuable informa-tion of recirculation pump power and primarycoolant massflow which were not known from meas-urements. A simple pump speed controller wasintroduced in the model in order to attain aspecified mass flow target during steady state.It was set up in such a way that when switchingTRAC to transient mode the so obtained pump speedremained constant.

The full plant model was taken from the loadrejection analysis and all the initial conditionwere manually updated to correspond to the perti-nent situation. Only minor changes had to beintroduced. The steady state calculation wasthen performed in two steps. After specifyingmeasured feedwater flow (= steam flow) and tempe-rature on secondary side and pressurizer pressureand core power as well as coolant massflow targeton primary side the system was allowed to stabi-lize. Thus during this first step a system steadystate with the new primary massflow was obtained.

During the second step the transient mode wasactivated and a null transient was run. The pres-surizer pressure boundary condition was thusdeactivated and the core kinetics implemented.The so obtained steady state is shown in table 3.

The model steady state condition was saved on adumpfile for later retrieval at the transientsimulation.

The total steady state analysis was run for180 seconds with a maximum time step of about4 seconds (obtained rather late in the calcula-tion). This required 985 CP-seconds and relevantstatistics for the steady state were:

CPU-time/problem time = 5.47

CPU-time/cell and timestep = 0.034 s

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14

Table 3

Results of steady state analysis.

Primary side

Parameter Measured/ Calculatedspecified

Core power (MW) 2 775 2 780

Coolant flow (kg/s) 13 790 13 792

RCP speed (rad/s) - 150.4

T hot leg (K) 594.5 594.6

T cold leg (K) 558.6 558.3

Prz pressure (MPa) 15.52 15.54

Prz level (%) 46.5 46.2

Secondary side

Parameter Measured/ Calculatedspecified

SG dome pressure (MPa) 6.22

Steam line pressure(MPa), 6.19 6.11

SG level (%) 61.9 61.7

CR-ratio (-)- 2.45

Feedwater flow (kg/s) 1 528.9 1 528.9

Feedwater temperature(K) 495.9 495.9

Steam flow (kg/s) 1 530.0 1 529.0

During the steady state 10 per cent ofthe total feedwater flow was suppliedto the top nozzle and the rest to thebottom nozzle.

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5 Data comparison

The simulation was made using a single loop repre-sentation. The measured thermal-hydraulic datawere obtained mainly for loop 2 with some fewdata taken in loop 1 and 3. Thus an averagingand extrapolating procedure had to be applied inorder to provide data for an average single loop.The data processed in this way were

Steam flow

Main feedwater flow

- Narrow range steam generator level

- Wide range steam generator level

- Auxiliary feedwater flow

During the steady state as well as the transientsimulation it was thus assumed that completesymmetry prevailed between the loops.

5.1 Boundary conditions

The main heat source during the transient wasthe core power and decay heat. The core reactivityparameters were specified to correspond to middleof cycle condition. Otherwise default kineticparameters were used. The decay heat was simulatedaccording to the ANSI-curve assuming equilibriumconditions. The rod insertion following the reac-tor trip signal (low steam generator level) wasspecified according to Ringhals 4 FSAR chapter 15.A time delay of 2.0 seconds was used from thetime point the trip was true to the time pointthe control rod reactivity insertion started. Ithas to be realized that this yields conservativereactivity insertion and is intended for safetyanalysis.

The (conservative) control rod reactivity inser-tion curve used in the analysis is shown infigure 4. The maximum reactivity worth of thecontrol rods was -0.08385.

The speed of the reactor recirculation pumps wasmaintained constant at the steady state valuethroughout the transient.

The pressurizer control was fully modeled usingthe rated values for the proportional and back-upheaters. The spray flow was taken from loop sealand was allowed to vary from its trickle flow toits maximum rated value. This flow was extracted

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from loop seal by means of FILL-component 715(figure 1) and the same flow was entered as sprayflow through FILL-component 500. For simplicitya constant spray water temperature was assumedcorresponding to the steady state cold leg tem-perature. This temperature was only slightlychanged during the transient thus justifyingthis assumption.

The feedwater isolation was introduced by a tripcontrolled table providing the main feedwaterflow as boundary conditions to top and bottominlet nozzles, FILL-components 741 and 744 infigure 1. A careful review of pertinent timeplotsand signals revealed a feedwater time functionafter trip as a linear ramp from normal flow tozero flow during 2.5 seconds.

The auxiliary feedwater flow was taken form meas-urements. This flow was measured for each steamgenerator and thus the total flow was obtainedby summing the three separate flows. The measuredflow from the turbine driven pump to steam gener-ator 3 was not correctly recorded as a saturationwas obtained at 30 kg/s. According to specifica-tion this pump should deliver 48 kg/s and thislatter flow was used when specifying the auxiliaryfeedwater flow boundary condition. The flow wasentered as a trip controlled table to FILL-compo-nent 743, figure 1. As the feedwater isolationoccurred in December with fairly cold weather,and the auxiliary feedwater was taken from anoutdoor located storage tank the temperature ofthis water was assumed to be rather low. A reas-onable temperature according to the SSPB wasabout 298 K which was used throughout the simula-tion. The auxiliary feedwater flow as functionof time after the steam generator low level tripis shown in figure 5.

The operation of the turbine stop valves andsteam dump valves was not explicitly modeled.Instead this operation was imposed on the TRAC-model by applying the measured steam flow asboundary condition. When doing this it has to berealized that the steam flow time history mustbe separated in two parts; one describing theflow prior to reactor scram signal that causedturbine isolation and the other describing theflow after scram. The measured steam flow fromloop 2 taken as a pressure difference betweensteam dome and steam line is shown in figure 6.

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The low steam generator level trip occurred at37.4 seconds after the feedwater isolation. Inorder to let the TRAC-model fully control thesteamflow behavior it was assumed that the meas-ured steamflow immediately before the low leveltrip would have remained even if the trip occurredat a later timepoint. Thus the first part of thesteamflow boundary condition was assumed to bedescribed by the curve in figure 7 and this curvewas applicable until the low level trip occurred.

Once this trip has occurred the steam flow willfollow the curve in figure 8.

5.2 Results from the simulation

The transient was simulated for 300 s including10 s of pretransient steady state condition. At10 s the feedwater isolation started with topand bottom feed flow being ramped from fullyflow to zero flow during 2.5 s. From that timeand on the calculated transient developed as thetime sequence of events listed in table 2 aboveindicates. The results are shown in figures 9 to28. The plotted calculated variables wereusually filtered by means of a first order lagfunction with 0.5 s time constant in order tosome degree simulate the plant signalprocessing. In some cases other time constantswere used as indicated in the figures.

The simulation was carried out in basically twoseparate stages. First a basecase was run withnodalization and boundary conditions as describedabove. Second some modifications of the nodaliza-tion were introduced and sensitivity studieswere carried out for a few important parameters.The results from the basecase are discussedfirst.

5.2.1 Basecase simulation

The steam flow (specified as boundary condition)continued after the feedwater isolation andexperienced a reduction due to the throttling ofthe turbine valves and a later increase when thesteam dumping started. In figure 9 the steamflow at the outlet of the steam generator isshown and compared to measured values. The directflow indicates the total flow as represented bya TRAC signal variable. This flow was in fullyagreement with the imposed flow boundary condi-tions downstream of the turbine valves. The flow

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1988-11-10

as taken as differential pressure between thesteam generator dome and the steam line revealeda discrepancy from the direct flow when the flowwas reduced and the pressure increased, figure 10.The basic reason for this behaviour was the omis-sion of pressure dependence in the flow algorithm.Also some minor influences originated from thenodalization of the very first part of the steamline and the steam line pressure drop distribution.Modificatjons were introduced for the calculationsin the following stage.

The steam generator narrow range level is shownin figure 11. A satisfactory agreement betweenthe calculated level from a differential pressureand the measured signal was obtained until thelow level trip setpoint (33 %) was reached at45.9 s. At that time point an oscillation in thecalculated level signal was encountered that hadno correspondence in the collapsed level nor inthe measurements. This is more easily seen infigure 12 showing the same phenomenon in anothertime scale.

No changes in the boundary conditions were encoun-tered at the time when the oscillation started.The auxiliary feedwater flow trip was not latchedmeaning that when the level signal recovered thetrip condition became false again as shown infigure 13 and consequently no flow had time tobe introduced until the trip was true the secondtime at about 50 s. The sharp decrease of steamflow did not occur until the reactor trip wastrue. This trip was latched and delayed 2 s fromthe time point the low level was obtained. Thusthe reactor power was maintained during the timefor the oscillation. A careful review of thelower and upper pressure tap pressure behavior,figure 14 marked area, revealed a high sensitivityof level signal on small dissimilar variationsin lower and upper tap pressures. Especially thelower tap pressure was strongly influenced bythe conditions in the downcomer. For instance atthe time for level oscillation a vapor contentof downcomer cell 4 had just become apparent,the void fraction being rather low though,figure 15. The void fraction time derivative waschanging somewhat due to variations in vaporgeneration, figure 16, which will have some in-fluence on the cell pressure as the volume ofcell 4 was rather small compared to surroundingcells. A more careful nodalization of the down-comer would have alleviated the problem and wasintroduced for. the simulations in the secondstage.

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On the primary side prior to reactor trip thetemperature increased, figure 17, because of theless efficient heat removal on the secondaryside when the feedwater flow ceased and thethrottling of the turbine valves was activated.The difference between the measured high averagetemperature and the calculated values could stemfrom the fact that the measurement representedthe highest value from three loops whereas thecalculated temperature represented average valuesfrom the three loops. However, the pressurizerlevel, which could be interpreted as a measureof the average temperature of the primary side,revealed a somewhat higher measured temperaturethan calculated, figure 18. It is also clearthat the calculated core power during this partof the transient before reactor trip was somewhatlower than the measured power, figure 19, indicat-ing that the specified moderator temperaturereactivity coefficient in the model was somewhattoo high (too negative).

Once the low steam generator level signal wasobtained at about 50 s, figure 11, the isolationof the turbines was initiated and the steam flowaccording to figure 8 was imposed as boundarycondition in the model. Also auxiliary feedwater,figure 5, was fed to the top feedwater nozzleand reactor trip occurred, figure 19.

The steam flow decrease after trip as representedby the differential pressure across the steamgenerator outlet nozzle, figure 9, revealed anexcellent agreement with the measurement untilthe actual steamflow was approaching zero flow.Although the imposed steamflow apparently waszero a non-zero value was obtained from thedp-calculation. This fictitious flow correspondsto the elevation pressure difference between thepressure tap locations and was not compensatedfor in the model.

The measurement indicated a positive constantsmall steamflow prevailing for the rest of thetransient. This was an extrapolated value fromthe only available steamflow measurement (steamline 2). In the operational report of this occur-rence there was no explanation whether it was atrue flow or a signal error from the measurementdevice. In view of the time history plots ofother variables as outline below, it can bereasonable to assume that a small steamflowactually was existent during the latter part ofthe transient.

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The calculated steam line pressure after reactortrip was somewhat low compared to measurement,figure 10. The difference was about 0.2 MPa.This low pressure was caused by a somewhat lowprimary temperature (about 2 K), figure 17. Thelow calculated primary temperature was alsorevealed from the pressurizer level response,figure 18. Thus it seemed apparent that the cal-culated power generated on the primary side afterreactor trip was too low.

The calculated core power was following veryclosely to the measured curve, figure 19. Itshould be emphasized when interpreting figure 19that once the reactor was scrammed the outputfrom the PRM-detectors was not highly accurate,especially when realizing that the basic heatsource then was from y-decay. The simulationwould have been improved if the stored energy inthe fuel was increased. This could have beenobtained if a lower gap conductance was used.This will also delay the energy equalizationprocess between the fuel and coolant. A value of10 kW/m 2K was used in the basecase calculation.For the later simulations two new values for thegap conductance were suggested; 7.5 and 5.0 kW/m'K.

Quite rapidly after the reactor and turbine tripthe liquid level in the steam generator downcomerdecreased below the narrow range lower pressuretap, figure 11. Both the measured signal and thecalculated signal from a differential pressureretained a positive non-zero level. This wascaused by the elevation head between the tapsand was not corrected for in the model, nor inthe plant.

The wide range level is shown in figure 20. Itseemed that the water content of the steam gener-ator was somewhat low in the simulation whichbecame apparent when the internal circulationceased at about 80 s. From that time and on theincrease in level had its origin in the auxiliaryfeedwater flow. The slope of the measured andcalculated curves was during this time about thesame. This indicates that the auxiliary feedwaterflow in the model was about what was used inreality and that the assumption of using designflow when the measurement was uncertain wasadequate.

From the long term behavior of the primary aver-age temperature, figure 17, the pressurizer level,figure 18, and the pressurizer pressure, figure 21it could be concluded that the calculated coolingof the primary side was less than the measurements

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indicated. As the auxiliary feedwater flow seemedto be adequately modeled it is reasonable toassume that a small steam flow actually was pre-vailing during the major part of the transient.A fact is that the turbine driven auxiliary feed-water pump was operating by steam from one ortwo of the loops. However, the steam flow throughthis device was not known. A crude estimationrevealed that a steam flow of about 10-15 kg/swas required to decrease the calculated primarytemperature with a rate similar to the measure-ment. A sensitivity study was suggested with thesteam flows 10 and 15 kg/s.

5.2.2 Second stage simulations

In view of the results from the basecase simu-lation the original model was somewhat modifiedbefore the second set of calculations was carriedout.

First the steam line downstream of the steamgenerator was somewhat more dense nodalized inorder to see any possible nodalization effect onthe pressure drop across the dome outlet. Alsothe pressure drop distribution of the entiresteam line was slightly modified to correspondmore closely to measured values obtained at anearlier made plant performance test.

Second the nodalization of the steam generatordowncomer was changed to avoid the oscillationin liquid level when taken as function of differ-ential pressure. The new nodalization is shownin figure 22. The number of cells in the down-comer was increased from 8 in the basecase modelto 17 in the new model.

Four new calculations were carried out. The fuelgap conductance was reduced from the basecase10 kW/m 2 K to 7.5 kW/m 2 K and 5.0 kW/m 2 K. Thiswould successively increase the stored energy inthe fuel (increase the fuel temperature) priorto the transient. During the transient theincreased initial fuel energy would increase theprimary side coolant temperature and increasethe secondary side pressure. Before the transientsimulation a new core steady state had to beobtained for each new gap conductance. Thesesteady state conditions were saved on files forlater retrieval at the transient runs.

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For each new gap conductance two different longterm steam flow values were used: 10 kg/s and15 kg/s. This steam flow could partly be consid-ered to pass through the turbine driven auxiliaryfeedwater pump. An estimate revealed that theneeded steam flow to provide the specified feed-water flow under assumption of reasonable losseswas comparable or below these values.

Also the .moderator temperature reactivity coef-ficient was decreased (i.e. was made less nega-tive) in order to increase the core power duringthe very first part of the transient prior toreactor trip when the primary coolant temperaturewas increasing. A scrutiny of the fuel conditionsat MOL as given in the nuclear design reportrevealed that the coefficient could be changedby about 3 pcm.

The results of the new simulations with the abovementioned modifications are presented in figures 23through 28 where some important variables areplotted. From figure 23 and earlier figure 10 itwas clear that the more dense nodalization ofthe upstream part of the steam line did not haveany significant influence on the pressure dropacross the steam dome outlet. Thus the nodaliza-tion was conceived to be converged with respectto dome exit pressure drop calculation and defic-iencies in steam flow obtained as function ofthis pressure drop had to be attributed to theflow algorithm. This will be discussed later.

The influence from initial store energy in thefuel and the long term steam flow on the transientis seen in figures 23 through 25. When the fuelgap conductance was decreased more energy wasinitially stored in the fuel. This energy wastransferred to the primary coolant during thetransient. Consequently a higher coolant averagetemperature (or pressurizer level) and pressurewere attained when the gap conductance was reducedas is found in figures 23 and 24. This resultedalso in a higher steam line pressure which isrevealed by figure 25. It was realized that forthe used values of gap conductance the lowestvalue (5.0 kW/m 2 K) resulted in the best comparisonwith measurements.

During the later phase of the transient the steamflow had an easily perceived influence on thecooling of the reactor coolant system. This effectis clearly seen in figures 23 through 25 starting

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at about 100 s. The steam flow resulted in asimilar calculated behaviour of the pressurizerlevel and primary and secondary pressures ascorresponding measurements. It seems that a steamflow of 15 kg/s somewhat overpredicted the coolingwhereas 10 kg/s resulted in a more adequate systemresponse.

With the more dense nodalization of the steamgenerator downcomer the pressure distributionexperienced a more smooth behaviour. The earlierfound oscillation in the downcomer narrow rangelevel when specified as function of a differ-ential pressure, figures 11 and 12, was not foundwith the new nodalization, figure 26. This figureshows the result for just one simulation. However,the other simulations revealed the same behaviourfree from oscillations. Thus a careful nodaliza-tion of the downcomer upper part was essentialfor proper pressure response when the liquidlevel decreased.

The suggested change in the moderator temperaturereactivity coefficient did not result in anynoticeable improvement of the core power whenthe core coolant temperature was increased,figure 27. If compared to the basecase result,figure 19, the new calculation resulted in aneven more pronounced power decrease before reactortrip. The reason seemed to be that in the newcase the coolant temperature experienced a higherincrease than in the basecase. Although the reac-tivity coefficient had been changed the amountof change was obviously not enough to also compen-sate for the higher coolant temperature. Similarbehaviour was also obtained for the other simula-tions.

Finally an investigation was made of the import-ance of including pressure compensation to thesteam flow when calculated as function of domeoutlet pressure drop. In the basecase no suchcompensation was made and a deficiency wasobtained when the flow decreased simultaneouslywith a pressure increase, figure 9. Two compen-sations were introduced. First a compensationfor the elevation head was made and second thecompensation also included the absolute pressureinfluence through the density. The result isshown in figure 28 revealing the importance ofpressure compensation and at low flow rates alsothe compensation for elevation head.

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5.3 Code performance

In a first attempt the 310 s basecase transientwas executed with no limitation by input of thetimestep. Thus TRAC was allowed to use as big atimestep as the solution method permitted. Inthat calculation the execution was terminatedafter about 160 s transient time because of con-verging problems in the steam generator downcomercomponent. The timestep size at this instancewas about 1.32 s. The very same case was sub-sequently rerun but now with a maximum allowabletimestep size of 0.5 s. The execution was nowsuccessfully completed. Thus it seems that thetimestep size algorithm is not efficient enoughto control the numerical solution to a convergentstate under all conceivable circumstances.

The 310 s transient was executed without anyrestarts. These 310 s required 1 068 timestepsand needed 4 784 CPU-s on the CDC Cyber 170-835.

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STUDSVIK PUCLEAR STUDSVIK/NP-88/101 25

1988-11-10

6 Conclusions

An assessment of TRAC-PFI/MOD1 version 14.0 againstan inadvertent feedwater line isolation in theRinghals 4 PWR power plant was conducted. Exten-sive use of results from Ringhals 4 data aquisi-tion system was made to derive the initial condi-tions and also to specify the necessary boundaryconditions.

The results from the TRAC simulation were comparedto measured data. From this comparison it wasclear that the used TRAC-version was capable ofperforming a satisfactory simulation.

However, the results also indicated some areasof model improvements which were investigated inlater calculations. The steam flow taken asproportional to the square root of a pressuredrop revealed for the basecase a discrepancywhen compared to measurement. It was found by anodalization analysis that a change to a moredense nodalization of the upstream part of thesteam line had only a minor influence on thepressure drop. The basic reason for the discrep-ancy was found to be the omission of pressurecompensation in the flow algorithm. When thiscompensation was introduced a favorable compari-son with measured steam flow was obtained.

During the course of steam generator leveldecrease an oscillation was initiated in thenarrow range level signal of the model. The levelsignal was expressed as a differential pressurebetween specified taps in the downcomer. Theoscillation had no correspondence in the measuredsignal nor in the collapsed level calculation ofthe model. It seemed that this behavior was causedby the vapor flow into the downcomer cells incombination with the condensation of the vaporin the cells as long as subcooled conditionsprevailed. A denser nodalization especially ofthe downcomer upper part helped to alleviate theproblem.

It was also found from the early core powerresponse of the model that the specified moder-ator temperature coefficient was somewhat toohigh. Despite the too low calculated core coolanttemperature increase the core power decrease wasexcessively predicted. A lower (less negative)reactivity coefficient would have resulted in ahigher core power that in turn would have raisedthe primary temperature level during this part

STURAPP NPS/AH


1988-11-10

of the transient. It was found in later calcula-tions that the suggested change in reactivitycoefficient was not enough to result in anynoticeable improvement in the core power response.

The primary temperature in the basecase modelwas too low compared to measurements. An increaseof the initial stored energy of the fuel wouldhave raised the coolant temperature. An increaseof the stored energy was obtained by decreasingthe gap conductance of the fuel. In the basecasea value of 10.0 kW/m 2 K was used for the gapconductance. A sensitivity analysis was carriedout with successively lower values. A value of5.0 kW/m 2 K was resulting in a reasonable responseof the reactor system when compared to measurement.

The long term development of the primary tempera-ture revealed in the basecase a less efficientcooling of the primary side in the TRAC-modelthan in the plant. A small steam flow of about10-15 kg/s was estimated to provide this addi-tional cooling. From the measurements there wassome supportive indications that a steam flowactually was prevailing during major parts ofthe transient. Clearly the turbine driven auxili-ary feedwater pump was operating during thispart of the transient; however, the steam flowthrough this device was not known. Sensitivityanalysis was performed to establish the influenceof the long term flow. It was found that a flowrate of 10 kg/s was needed to adequately predictthe long term cooling.

External restriction on the maximum allowabletimestep size had to be imposed for convergenceof the solution procedure. It was found that0.5 s was adequate for the simulation to proceedthroughout the transient. For the 310 s transientunder this restriction 4 748 CPU-s was needed ona CDC Cyber 170-835 computer.

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1988-11-10

References

1 TRAC-PF1/MOD1. An Advanced Best-EstimateComputer Program for Pressurized WaterReactor Thermal-Hydraulic Analysis.NUREG/CR-3858. July 1986.

2 RANSOM, V et alRELAP5/MOD2 Code Manual.NUREG/CR-4312. August 1985.

3 SJOBERG, A, ALMBERGER, J, SANDERVAG, 0ICAP. Assessment of TRAC-PF1/MODIAgainst a Loss of Grid Transient inRinghals 4 Power Plant.Studsvik Nuclear, Sweden 1987.STUDVIK/NP-87/10.

4 BOYACK, B E et alTRAC User's Guide.NUREG/CR-4442. November 1985.

5 PELAYO, F, SJOBERG, AICAP. Assessment of TRAC-PF1/MOD1Against an Inadvertent Steam Line Isola-tion Valve Closure in the Ringhals 2Power Plant.Studsvik Nuclear, Sweden 1988.STUDSVIK/NP-88/14.

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[/) Ii-

U)

763 76 7,,6O1 7•63 764 •

761 sv 762 lRv761 762

731-740 () 741

js 71 751 751 754 72 753 75

474" 745

rr tow 1 R~I17 270(it0 1 74 7 270(0$ osP 05 71

vIS, -- 04L, o

70701 710 172 171 271 Z7Z p U)6 172 171 271 272 t

co14 174 73 273 274 O

I U

710 717 174 173 273 274

710

0 Z

00

Figure 1

TRAC-PF1/MOD1 model of Ringhals 4.


1988-11-10

f.

P--Zsp

Figure 2

Steam generator nodalization.

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1988-11-10

30

Figure 3

Pressurizer pressure and level control logic.

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1988-11-10

1.0

0.9

0.8

0.7

I,-

0.6

1

w, 0.5

U

0

z0.3

0.2

0.1

00 1. 2.0 2.2 3.0

TIME FROM ROD RELEASE - SECONDS

Figure 4

Control rod reactivity insertion curve.

STURAPP NPS/AH

STUDSVIK NUCLEAR STUDVIK 1JCL1~RSTIJDSVIK/NP-88/101 32

1988-11-10

(:iý

060,

40.

20

T00 L50 200

Time ofter low level sionol (s)

Figure 5

Auxiliary feedwater flow vs time.

STURAPP NPS/AH

STUDSVIK NUCLEAR

600"

500-

400"

.300-0

C,

STUDSVIK/NP-88/101

1988-11-10

33

200'

100"

100 150 200Time after feedwoter isolotion (s)

250

Figure 6

Measured steamflow.

STURAPP NPS/AH

STUDSVIK NUCLEAR STUDSVIK/NP-88/1Ol 34

1988-11-10

600-

500-\

400-

0 .300-

200-

100-

00 0o 160 150 260

Time after feedwater isolation (s)25o 300

Figure 7

Steamflow prior to low level signal.

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1988-11-10

600.

500"

400.

00

300.

2.00-

100.

30050 100 150 200Time after low level signol (s)

Figure 8

Steamflow after low level signal.

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1988-11-10

ICAP. RINGHALS 4, FEEDWATER LINE ISOLATION.

00c,

0'

Qn

LEGENDo = DIRECTo= FROM DP+ = MEASURED

Nel

-is

Li

0

In

150.0Time (s)

200.0 250.0 300.0

Figure 9

Steamflow at steam generator outlet nozzle.

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1988-11-10


LEGENDo = STEAM LINEo = STEAM DOMEA = STEAM LINE, MEASURED

0,4

L/)

LiJ

R~0n

R

0

0

50.0 100.0 150.0

Time (s)200.0 250.0

Figure 10

Secondary side pressure.

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1988-11-10


00*.

LEGENDo= COLLAPSEDo= FROM DP&-'= MEASURED

z

0

z-j

Li-j c

0i

Li~

Li

100.0 150.0

Time (s)200.0 250.0 300.0

Figure 11

Steam generator narrow range level.

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1988-11-10


00A

LEGEND= COLLAPSED=FROM DP= MEASURED

0

wiCDz

0

z

-LJ

LIi

Lii

00o

33%

00.

20.0 60.0 80.0 100.0

Time (s)

Figure 12

Steam generator narrow range level, expanded time scale.

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. STUDSVIK NUCLEAR STUDSVIK/NP-88/101 40

1988-11-10


0

F-

0w

IL--

0/

0

0.0 20.0 40.0 60.0 80.0

Time (s)

Figure 13

Auxiliary feedwater flow control trip signal.

100.0

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1988-11-10


LEGENDo = LOWER P-TAPo = UPPER P-TAP

0

wj

80.0

Time (s)

Figure 14

Steam generator downcomer pressures.

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1988-11-10


0

zz0

0

06

0.0 20.0 80.0 100.0

Time (s)

Figure 15

Steam generator downcomer cell void fraction.

STURAPP NPS/AH


1988-11-10


00

0

Lf)

z R

0

z

a-

IU

R67.

40.0 80.0

Time (s)

Figure 16

Steam generator downcomer cell vapor generation term.

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STUDSVIK NUCLEAR STUDVIKNUCLAI ~STUDSVIK/NP-88/101 44

1988-11-10


LEGENDo = UNFILTEREDo = FILTERED 1.0 S

= MEASURED

R

sn

<I

0~<

>~

low

06•

9C

100.0 150.0 200.0 250.0 300.0

Time (s)

Figure 17

Primary side average temperature.

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1988-11-10

ICAP. RINGHALS 4, FEED WATER LINE ISOLATION.

LEGENDo = CALCULATEDo = MEASUREDR

0o

00o

LI

LI

LI)v

LI.-

100.0 150.0

Time (s)200.0 250.0 300.0

Figure 18

Pressurizer level signal.

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1988-11-10


06.

C00O

00

Li,

00C0..0- toC:C0C-)

.4La

0•0-

0

0-

&

LEGEND= CALCULATED= MEASURED

p-..............-.----------.---..--.------. ~.------1.---------'. ~.

0.0 50.0 100.0 150.0

Time (s)200.0 250.0 300.0

Figure 19

Reactor core power.

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1988-11-10


LEGEND0 COLLAPSEDo = FROM DPa = MEASURED

o•0

z

Li

-3

Li

Li

(A1

06

00,

50.0 100.0 150.0

Time (s)200.0 250.0 300.0

Figure 20

Steam generator wide range level.

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1988-11-10


LEGENDo = CALCULATEDo= MEASURED

0r~Z

U,

0:-

wNJ

w,

Lii

0,4.

150.0

Time (s)200.0 250.0 300.0

Figure 21

Pressurizer pressure.

ST1JRAPP NPS/AH


1988-11-10

1;10

Figure 22

Modified steam generator nodalization.

STURAPP NPS/AH


1988-11-10

50

ICAP. RINGHALS A. HGAP=5000, STEAM FLOW 10 KG/S

LEGENDA= ALCULATEDME= EASURED

ICP. RINGHALS 4, HGAP=5000, STEAM fLOW = 15 KG/S.

LEGEND0 = CALCULATED

o 0 =MEASUREDo6.

no.0 200,0

rime (S)C00A0

rime (S)

ICAP. RINGHALS 4, HGAP=700. STEAM rLoW = -0 KG/S #CAP. R1NGHALS . NGAP=7500, STrAu rLOW = 15 KG/S.

LEGEND LEGENDc = CALCULATED 0= CALCULATEDa 0= MEASURED 0 MEASURED

0

150.0 200.0rime (s)

Figure 23

Pressurizer level signal.

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1988-11-10

51

ICAP. RiNGHALS A. HGAP=5000. STEAL' rLoW= 10 v.G/S

LEGENDoCALCULATEDo=MEASURED

tCAP. RINCHALS A. H-GAP=5000. STEAM nLow =1 KG/S.

LEGENDo = CALCULATEDo = MEASURED

C

C. -

N

E

In

~

C

C.

10.0 300.0 310.0 2000 210.0 300.0ITime (s)

ICAP. RINGHALS 4. HGtP=7500. STEAM rLOW = 10 KG/S ICAP. RtNGHALS 4A HGAP-7500, STEAM rLow = ts KG/S.

LEGEND LEGENDo CALCULATED v = CALCULATEDCALCULATED

o = MEASUREDME WASURED •

0.~

Li

IAIn

31,

C

Figure 24

Pressurizer pressure.

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1988-11-10

ICAP. RINOHALS 4. NGAP=5000. STEAM FLOW = 10 KC/S

CA

&AIIn

LEGENDo = STEAU UNEo = STEAM DOME& = STEAM UNE. MEASURED *1

ICAP. RINGHALS 4. HGAP=5000. STEAM rLOW = I1 KG/S.

LEGENDo = STEAM LINEo = STEAM. DOME

= STEAM LINE. MEASURED

Joý

D.

0.0 50.0 I00.0 tse ,Time (s)

200.0 20.0 a 500.0 .0 S0.0 T00.0 e500Time (S)

20o0.0 25'o.0 so3o

ICAP. RINGHALS 4. HGAP=7500. STEAM rLOW = 10 KG/S ICAP. RINGHt.LS 4. HGAP=SO0. STEAM rLO\% = 15 KC/S

:i1

LEGENDa = STEAM LINEo : STEAM DOME

£STEAM LINE. MEASURED:I

Cl

a-

0~~

LEGENDo = STEAM LINEo = STEAM DOME* = STEAM LINE. MEASURED

0.

rn-JOa. a

InIn.4

Oj

ICai© : ....

0 0 00 o 0 1,50.0Time (s)

2ý00. 2;0o0 3000 0.0 50.0 ,0o00 15o 0Time (s)

200.0 250.0 300.0

Figure 25

Secondary side pressure.

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1988-11-10

ICAP. RINGHALS 4, HGAP=5000, STEAM FLOW = 10 KG/S.

LEGENDo = COLLAPSEDo = FROM DP

= MEASURED

z

0

0:

Li

V)

50.0 200.0 250.0 300.0Time (s)

Figure 26

Steam generator narrow range level.

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1988-11-10

ICAP. RINGHALS 4, HGAP=5000, STEAM FLOW = 10 KG/S.

04

C;

LEGENDo = CALCULATEDo = MEASURED

00-

0R

0!--

0

IJ

0

0-

0;

0-_C1

0.0 50.0 100.0 150.0

Time (s)200.0 250.0 300.0

Figure 27

Reactor core power.

STURAPP NPS/AH

STUDSVIK NUCLEAR STUDSVIK/N~P-88/101 555

1988-11-10

ICAP. RINGHALS 4, HGAP=5000.

LEGENDoE= DIRECTo = FROM DPa = FROM DP, E-HEAD COMP.+ = FROM DP, E-HEAD AND P COMP.x = MEASURED

000IA

0~

LLI

LI)

0

0*0

Time (s)

Figure 28

Steamflow at steam generator outlet nozzle.

STTJRAPP NPS/AH

NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER(2.89) (ANWd by NRC. Add Vol., Supp., Rw.,NRCM 1102. d Addendum Numbems. If any.)3201.3202 BIBLIOGRAPHIC DATA SHEET

(See instructions on the reverse) NUREG/IA-00382. TITLE AND SUBTITLE STUDSVIK/NP-88/101

Assessment of TRAC-PF1/MOD1 Against an Inadvertent FeedwaterLine Isolation Transient in the Ringhals 4 Power Plant 3. DATE REPORT PUBLISHED

MONTH YEAR

March 19924. FIN OR GRANT NUMBER

A46825. AUTHOR(S) 6. TYPE OF REPORT

A. Sjoberg7. PERIOD COVERED flnclusive Dated

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (If NRC, provide Division, Officeor Region. U.S. Nuclear Regulatory Commission, and mailing address; If contractor, providenamte and mailing address)


9. SPONSORING ORGANIZATION - NAME AND ADDRESS iIf NRC, type "Same as above", it contractor, provide NRC Division, Office or Region, U.1 Nuclear Regulatory Commison,and mailing address.


10. SUPPLEMENTARY NOTES

11. ABSTRACT (200 words or ess)

An inadvertent feedwater line isolation transient in a three loop Westinghouse PWRhas been simulated with the frozen version of the TRAC-PF1/MOD1 computer code. Theresults revealed the capacity of the code to quantitatively predict the differentpertinent phenomena.

For accurate predictions of the system response, particular care was required in thenodalization of the steam generator downcomer and restrictions had to be imposed onthe allowable maximum time step.

12. KEY WORDS/DESCR!PTORS (LIst words or phreaes that will alst restaChers in locating the report.) 13. AVAILABILITY STATEMENT

ICAP Program Unlimited

Ringhals 4 Power Plant 14. SECURITY CLASSIFICATION

TRAC-PF1/MOD1 Computer Code This Pawe)

Unclassified(This Report)

Unclassified15. NUMBER OF PAGES

16. PRICE

NRC FORM 335 (2-89)

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UNITED STATESNUCLEAR REGULATORY COMMISSION

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