NUREG/IA-0124

ICSP-JC-SPR-R

International

Agreement Report

Assessment of RELAP5/MOD2

Against a Pressurizer Spray

Valve Inadverted Fully Opening

Transient and Recovery by

Natural Circulation in

Jose Cabrera Nuclear Station

Prepared by

R. Arroyo, L. Rebollo

Union Electrica Fenosa

c/Capitan Haya 53

28020 Madrid

Spain

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

June 1993

Prepared as part of

The Agreement on Research Participation and Technical Exchange

under the International Thermal-Hydraulic Code Assessment

and Application Program (ICAP)

Published by

U.S. Nuclear Regulatory Commission
NOTICE

This report was prepared under an international cooperative

agreement for the exchange of technical information. Neither

the United States Government nor any agency thereof, or any of

their employees, makes any warranty, expressed or implied, or

assumes any legal liability or responsibility for any third party's

use, or the results of such use, of 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.

Available from

Superintendent of Documents

U.S. Government Printing Office

P.O. Box 37082

Washington,"D.C. 20013-7082

and

National Technical Information Service

Springfield, VA 22161
NUREG/IA-0124

ICSP-JC-SPR-R

International

Agreement Report

Assessment of RELAP5/MOD2

Against a Pressurizer Spray

Valve Inadverted Fully Opening

Transient and Recovery by

Natural Circulation in

Jose Cabrera Nuclear Station

Prepared by

R. Arroyo, L Rebollo

Union Electrica Fenosa

c/Capitan Haya 53

28020 Madrid

Spain

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

June 1993

Prepared as part of.-

The Agreement on Research Participation and Technical ExOtange

under the Internatbonal Thermal-Hydraulic Code Assessment

and AppLication Progra (iCAP)

Published by

U.S. Nuclear Regulatory Commission
NOTICE

This report documents work performed under the sponsorship of the Consejo De

Seguridad Nuclear of Spain. The information in this report has been provided

to the USNRC under the terms of an information exchange agreement between the

United States and Spain (Technical Exchange and Cooperation-Agreement Between

the United States Nuclear Regulatory Commission and the Consejo De Seguridad

Nuclear of Spain in the field of reactor safety research and development,

November 1985). Spain has consented to the publication 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

Spain or any agency thereof, or any of their employees, 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.
ABSTRACT

This document presents the comparison between the

simulation results and the plant measurements of a real event

that took place in JOSE CABRERA nuclear power plant in August

3 0 th, 1984. The event was originated by the total, continuous

and inadverted opening of the pressurizer spray valve

PCV-400A.

JOSE CABRERA power plant is a single loop Westinghouse

PWR belonging to UNION ELECTRICA FENOSA, S.A. (UNION FENOSA),

an Spanish utility which participates in the International

Code Assessment and Applications Program (ICAP) as a member of

UNIDAD ELECTRICA, S.A. (UNESA). This is the second of its

two contributions to the Program: the first one was an

application case and this is an assessment one.

The simulation has been performed using the RELAPS/MOD2

cycle 36.04 code, running on a CDC CYBER 180/830 computer

under NOS 2.5 operating system.

The main phenomena have been calculated correctly and

some conclusions about the 3D characteristics of the

condensation due to the spray and its simulation with a ID

tool have been got.
EXECUTIVE SUMMARY

This work shows the results of the analysis with

RELAP5/MOD2 cycle 36.04 of a real event that took place in

JOSE CABRERA nuclear power plant in August 3 0 th, 1984. The

-event had its origen in a total, continuous and inadverted

opening of the pressurizer spray valve PCV-400A. This is the

second contribution of UNION ELECTRICA FENOSA, S.A. (UNION

FENOSA) to ICAP.

JOSE CABRERA nuclear plant is a single loop Westinghouse

PWR. A general purpose nodalization of JOSE CABRERA for RELAP5

has been used. This nodalization is being widely used in

thermal-hydraulic applications and has given good results.

The inadverted opening of the spray valve caused a

reactor coolant system depressurization, producing a reactor

trip. The reactor coolant pump was manually stopped half an

hour later in order to stop the spray flow. Afterwards, the

cooling of the primary system was due to natural circulation.

The safety injection system (9.7 MPa shut-off pressure) did

not introduce water into the primary circuit because the RCS

pressure was stabilized at 10.0 MPa.

The calculation purpose was not to study the control

systems response, because they worked mainly in their manual

mode, but to analyse the main thermal-hydraulic phenomena that

appeared during the transient: steam condensation in the

pressurizer due to the continuous spray flow and corresponding

-ii-
RCS depressurization, primary " cooling from the

generator, reactor coolant pump trip effect,

circulation, etc.

steam

natural

This objective was reached, the main phenomena were

reproduced in the simulation and the discrepancies were

justified. Main results of pressurizer separated effects

experiments (MIT and NEPTUNUS) have been confirmed. In

general, RELAP5/MOD2 code has performed properly and the host

computer, CDC CYBER 180/830 has run well with the code with a

good CPU time to real time ratio.

-iii-
LIST OF TABLES

TABLE 4.1

TABLE 5.1

TABLE 5.2

TABLE 5.3

TABLE 6.1

TABLE 6.2

TABLE 7.1

RELAP5/MOD2 model description (nodalization).

Stabilization system actuation logic.

Steady state results at nominal conditions.

Steady state results at 96%.

Sequence of events.

Variables identification in the transient figures.

Run statistics.

-iv-
FIG. 2.1

FIG. 3.1

FIG. 3.2

FIG. 3.3

FIG. 3.4

FIG. 3.5

FIG. 3.6

FIG. 3.7

FIG. 3.8

FIG. 3.9

FIG. 4.1

FIG. 4.2

LIo OF FIGURES

JOSE CABRERA NUCLEAR POWER PLANT REPRESENTATION.

PRESSURIZER PRESSURE (KG/CM2 REL, NARROW RANGE).

PRESSURIZER PRESSURE (KG/CM2 REL, WIDE RANGE) & RCS

DELTA TEMPERATURE (C).

RCS PROGRAMMED AVERAGE TEMPERATURE (C) & RCS AVERAGE

TEMPERATURE (C).

COLD LEG TEMPERATURE (C).

PRESSURIZER PROGRAMMED LEVEL (% OF THE SPAN) &

PRESSURIZER LEVEL (% OF THE SPAN).

RCS CHARGING FLOW (LPM).

STEAM GENERATOR PRESSURE (KG/CM2 REL) & STEAM

GENERATOR DOWNCOMER LEVEL (CM).

FEEDWATER MASS FLOW RATE (TN/H) & STEAM MASS FLOW

RATE (TN/H).

TURBINE LOAD (MWe).

JOSE CABRERA PLANT NODALIZATION.

RCP COASTDOWN. PERCENTAGE OF THE NOMINAL FLOW.

SHORT-TERM.
o_1 or ZGZun (CONT.)

FIG. 4.3 RCP COASTDOWN. PERCENTAGE OF THE NOMINAL FLOW.

LONG-TERM.

FIG. 4.4 AD-HOC REDUCED NODALIZATION.

FIG. 6.1 PRESSURIZER PRESSURE (KG/CM2 REL).

FIG. 6.2 RCS AVERAGE TEMPERATURE (C).

FIG. 6.3 COLD LEG TEMPERATURE (C).

FIG. 6.4 RCS DELTA TEMPERATURE (C).

FIG. 6.5 PRESSURIZER LEVEL (% OF SPAN).

FIG. 6.6 STEAM GENERATOR PRESSURE (KG/CM2 REL).

FIG. 6.7 STEAM GENERATOR DOWNCOMER LEVEL (CM).

FIG. 6.8 FEEDWATER MASS FLOW RATE (KG/S).

FIG. 6.9 STEAM MASS FLOW RATE (KG/S).

FIG. 6.10 CORE POWER (W).

FIG. 6.11 SPRAY MASS FLOW RATE (KG/S).

FIG. 6.12 PRESSURIZER HEATERS POWER (W).

-vi-
LIST aF FIGURES (CONT.)

FIG. 6.13 RCS & STEAM GENERATOR TEMPERATURES (K).

FIG. 6.14 LIQUID & SATURATION CORE EXIT TEMPERATURE (K).

FIG. 6.15 AVERAGE CHANNEL CLAD TEMPERATURE (K).

FIG. 6.16 RCS MASS FLOW RATE (KG/S).

FIG. 6.17 PRESSURIZER TEMPERATURES - NODE 1 (K).

FIG. 6.18 PRESSURIZER TEMPERATURES - NODE 2 (K).

FIG. 6.19 PRESSURIZER TEMPERATURES - NODE 3 (K).

FIG. 6.20 PRESSURIZER TEMPERATURES - NODE 4 (K).

FIG. 6.21 PRESSURIZER TEMPERATURES - NODE 5 (K).

FIG. 6.22 PRESSURIZER TEMPERATURES - NODE 6 (K).

FIG. 6.23 PRESSURIZER TEMPERATURES - NODE 7 (K).

FIG. 6.24 PRESSURIZER TEMPERATURES - NODE 8 (K).

FIG. 6.25 PRESSURIZER VOID FRACTION - NODES 5 & 6.

FIG. 7.1 CPU TIME (S).

-vii-

CONTENTS

ABSTRACT ................................................. 1

EXECUTIVE SUMMARY ........................................ i

LIST OF TABLES ......................... e.................. iv

LIST OF FIGURES .............. oo...... o..... ................ v

1.- INTRODUCTION ... o................. o ....... o ... o.... o...... 1

2.- NUCLEAR STATION DESCRIPTION .......................... 3

3.- EVENT DESCRIPTION ................... ............ o 6

4.- CODE INPUT MODEL DESCRIPTION ......................... 10

4.1 Primary system nodalization ..................... o11

4.2 Secondary system nodalization .................... 14

4.3 Trips and control variables ..................... 19

4.4 Simplifications ............................. o .... 21

5.- STEADY STATE CALCULATION ............................. 22

6.- TRANSIENT RESULTS AND COMPARISON WITH PLANT DATA ..... 24

7.- RUN STATISTICS ............ ...... o ...... 38

8.- CONCLUSIONS .......... ...................... ....... 39

9.- REFERENCES .......... o ... o ...... ............ o .... .. . 43

ADDENDUM A: ASSESSMENT OF RELAP5/MOD2 AGAINST PRESSURIZER

SEPARATED EFFECTS EXPERIMENTS.

ix

FOREWORD

This report has been prepared by Uni6n Electrica Fenosa

in the framework of the ICAP-UNESA Project.

The report represents one of the application calculations

submitted in fulfilment of the bilateral agreement for cooperation in thermalhydraulic activities between the Consejo

de Seguridad Nuclear of Spain (CSN) and the United States

Nuclear Regulatory Commission (USNRC) in the form of Spanish

contribution to the International Code Assessment and Applications Program (ICAP) of the USNRC whose main purpose is

the validation of the TRAC and RELAP system codes.

The Consejo de Seguridad Nuclear has promoted a coordinated Spanish Nuclear Industry effort (ICAP-SPAIN) aiming to

satisfy the requirements of this agreement and to improve

the quality of the technical support groups at the Spanish

Utilities, Spanish Research Establishments, Regulatory Staff

and Engineering Companies, for safety purposes.

This ICAP-SPAIN national program includes agreements

between CSN and each of the following organizations:

- Unidad Elictrica (UNESA)

- Uni6n Iberoamericana de Tecnologla El~ctrica (UITESA)

- Empresa Nacional del Uranio (ENUSA)

- TECNATOM

- EMPRESARIOS AGRUPADOS

- LOFT-ESPARA

The program is executed by 12 working groups and a generic code review group and is coordinated by the "Comit4 de

Coordinaci'n".. This committee has approved the distribution

of this document for ICAP purposes.

xi

D-1/90-MPNV

I.- INTRODUCTION.

The "International Thermal-Hydraulic Code Assessment and

Applications Program (ICAP)" is being developed by some

organizations from several countries coordinated by the U.S.

Nuclear Regulatory Commision' (USNRC). Its purpose is to

obtain a good vision of the USNRC thermal-hydraulic codes

validity over the widest range of possible applications. The

program generic interest derives from the extended diffusion

and use of these computer codes.

The Spanish contributions to ICAP consist of different

calculations, both assessment and application ones, with the

RELAP5/MOD2, TRAC-PFI/MODI and TRAC-BF1 codes. The assessment

cases compare simulations with real data, basically from

commercial operating plants. The application cases study

accidental situations, which prove the main thermal-hydraulic

models of these codes in an exhaustive manner.

UNION ELECTRICA FENOSA, S.A. (UNION FENOSA) participates

in ICAP as a member of UNIDAD ELECTRICA, S.A. (UNESA)

contributing with two calculations: an application case

(ref. 1) and an assessment case.

This report shows the assessment case, in which a real

event in JOSE CABRERA nuclear power plant (NPP), a single-loop

Westinghouse. PWR belonging to UNION FENOSA, is studied. This

event was caused by the total, continuous and inadverted

locked opening of the pressurizer spray valve, producing the

reactor trip. The reactor coolant pump (RCP) was manually
stopped afterwards and this caused the cooling of the primary

system by natural circulation without safety injection into

the primary circuit.

RELAP5/MOD2-cycle 36.04 running on a CDC CYBER 180/830

computer under the NOS 2.5 operating system was used to

analyse the selected plant transient.

The purpose of the calculation was not that of studying

the interphase between the control systems, most of which were

working in manual mode, and the thermohydraulic response of

the plant, but that of reproducing the main event phenomena.

The analysis has been based on the main thermal-hydraulic

phenomena that the event presented: pressurizer steam

condensation and corresponding reactor coolant system (RCS)

depressurization, RCS cooling from the steam generator (SG),

RCP trip effect, natural circulation, etc.

JOSE CABRERA NPP, where the event tooks place, is

described in section 2, and the event itself in section 3.

The model of the plant, used as RELAP5/MOD2 code input deck,

is explained in section 4. The steady state results are

commented in section 5. Section 6 shows the analysis results

and their comparison with available plant data. Section 7

shows some interesting data on computer performance, and,

finally, the main conclusions obtained are summarized in

section 8.

-2-
2.- NUCLEA STThIO DESCRIPTION.

The analysis was done for JOSE CABRERA Nuclear Power

Plant (fig. 2.1), a Westinghouse PWR commercial plant, sited

in Zorita de los Canes (Guadalajara), belonging to UNION

FENOSA, a Spanish utility (ref. 2). The plant had its first

criticallity in 1968 and was the first nuclear station

connected to the Spanish electrical grid.

The Reactor Coolant System (RCS) has only one loop that

includes the cold leg, reactor pressure vessel, hot leg,

pressurizer, steam generator tubes, cross-over leg, and

reactor coolant pump. The Chemical and Volume Control System

(CVCS) as well as the Residual Heat Removal System (RHRS) are

also connected to the reactor coolant loop.

Nominal reactor power of 510 Mwt is generated by the

reactor core that has a configuration of 69 fuel assemblies

(14 x 14) with 2.40 m of active length. Reload fuel has an

average enrichment of 3.60% in U-235. The plant nominal

output electrical power is 160 Mwe with a frequency of 50 Hz.

The Emergency Core Cooling System (ECCS) connects

directly to the downcomer of the reactor vessel and includes

one accumulator, two intermediate pressure safety injection

pumps taking borated water from the reload water storage tank,

and two recirculation pumps and a jet pump taking water from

the containment sump and feeding the injection pumps in the

recirculation phase of a loss of coolant accident (LOCA).

There is one single steam generator that links thermally

-3-
the primary and secondary systems. It includes the downcomer

annulus, riser boiling chamber, separators and dryers.

In the secondary side the typical balance of plant (BOP)

components are included (two 50% main feedwater pumps, steam

line, safety (4) and steam-dump (3) valves, main steam

isolation valve, turbine trip valves (2), main steam control

valves (4), high (1) and low (1) pressure turbines, condenser,

heaters (4), etc).

Main feedwater comes directly into the upper part of the

downcomer without passing through any preheater section inside

the steam generator, being previously heated through the

heaters installed between the condenser and the steam

generator. The circulation ratio in the secondary side of the

steam generator is 1.96 at full power.

The SG auxiliary-emergency feedwater system includes one

turbine driven and two motor operated pumps. Both subsystems

take cold demineralized water from a tank and start their

operation automatically. The turbine driven subsystem (AFWS)

injects into the upper part of the downcomer, injection

requiring the turbine operator intervention, by opening an

isolation valve from the Control Room. The motor operated

subsystem (EFWS) injects directly into the lower part of the

SG tubes, no operator action being necessary to allow the

subsystem injection. For this subsystem there is also the

possibility of an optional injection into the upper part of

the downcomer once the operator lines up properly the system.

The plant operates normally in automatic mode under the

-4-
influence of the reactor control system that maintains the

programmed coolant average temperature in the primary system

by acting on the control bank B position.

The RCS pressure is controlled by the pressurizer

pressure control system which acts on the PORV's, spray valve

and heaters. The pressurizer level control system follows a

level program as a function of the average temperature in the

RCS by acting on the speed of one of the two charging pumps.

The steam generator level is controlled in operation by

acting on the feedwater control valve based on the level error

for the specified level setpoint, the feedwater flow and the

steam flow.

Finally, there is a system to control the primary average

temperature after the turbine trip by opening the steam-dump

valves to the atmosphere (2) and to the condenser (1). This

system can also work under pressure control by following a

manually selected steam generator pressure setpoint.

The safety of the plant is guaranteed by the Reactor

Protection System and the Emergency Safety Features.

-5-
3.- EVET DESCRIPTION.

The event that was selected to be analysed and simulated

with the RELAP5/MOD2 code is a reactor trip in JOSE CABRERA

NPP, which happened at 13:40 h, August 30th, 1984. It was

caused by the total, continuous and inadverted opening of the

pressurizer spray valve PCV-400A that remained locked at its

fully open position for more than half an hour. Until the

beginning of the transient, the plant had been operating under

automatic control with all the main parameters within normal

ranges, at 96% thermal power.

The valve opening fault was not due to a control failure

by the pressurizer pressure control system, but to a failure

in its driving mechanism. At that time the only indication of

the valve position in the Control Room, in percentage of

demanding opening, came from the control system. Therefore,

the indication the reactor operator could see during the

depressurization event was 0% opening, preventing the correct

diagnosis of the event. As a result of the evaluation by the

NPP technical support center and in order to solve this

problem in the future, a valve PCV-400A stem actual position

indication was installed in the Control Room. This solution

was considered acceptable by the Consejo de Seguridad Nuclear

(Spanish regulatory commission).

When the primary pressure decrease following the spray

valve opening was observed, the turbine operator tried to

control it by means of a manual turbine load reduction. The

pressurizer relief and spray valves indications were also

checked showing 0% opening demanded by the pressurizer

-6-
pressure control system.

The pressurizer level control system was set into its

manual control mode, and the second charging pump was manually

started to try to repressurize the primary circuit by partial

refilling of the pressurizer.

As the RCS pressure went on decreasing, the low pressure

reactor trip was produced at the programmed setpoint, so

inducing a simultaneous turbine trip. After the turbine trip,

the steam release through the steam-dump valves, which were

working in their temperature control mode trying to reach the

RCS no-load average temperature (275.0 "C), caused the primary

pressure to go on decreasing very quickly. Therefore, the

steam-dump control system was set into its pressure control

mode by the reactor operator, and was manually controlled from

that moment on.

The safety injection signal was automatically activated.

From then on, the operation team continuously monitored the

intermediate pressure safety injection pumps performance

parameters, specially the driving motors currents and the ECCS

lines flow. The ECCS flow indication was always 0.0 m3 /h, and

the current of the motors was maintained at the normal value

that is measured during the pumps maintenance monthly tests

with the pumps working in the recirculation mode from/to the

borated water tank. This fact demonstrated that cold water

was never introduced into the vessel. This was also

consistent with the fact that the minimum pressurizer pressure

measured during the event was, according to the wide range

recorder, 100.0 Kg/cm2 (rel.), for a safety injection nominal

-7-
shut-off pressure of 99.0 Kg/cm2 (rel.).

When an auxiliar operator could go into the containment

building, once its normal conditions were checked, he

confirmed that the valve PCV-400A was fully open. A little

bit later, and because it was impossible to close the valve

either locally or from the Control Room, the RCP was tripped

to finish the depressurization, by stopping the spray flow.

By this operation, around half an hour after the beginning of

the transient, the ECCS injection of cold borated water into

the vessel downcomer was just prevented avoiding the potential

risk of pressurized thermal shock to the reactor vessel. Some

time later, the auxiliar operator could manually close this

valve locally.

Throughout the incident, the operation team supervised

the core behaviour by reading the core exit thermocouples

temperatures, in order to control the subcooling margin that

was always higher than 20.0 "C.

The minimum measured levels were 25% for the pressurizer

and -80.0 cm for the steam generator.

About 40 minutes after the RCS depressurization was

initiated, the event was already solved and the pressure was

being recovered with natural circulation in the RCS, the fan

coolers of the vessel head functioning, without any

pressurizer level symptom of a steam bubble formation in the

vessel head.

The lack of a process computer at the plant prevented

-8-
knowing exactly the sequence of the event and the precise

chronology of the operators intervention. Both had to be

deduced from the several records and graphics available in the

trip report, within the limits associated with their scales,

speeds and precisions. In spite of these difficulties, this

transient was selected as an assessment case because of its

duration and the amplitude of the variation of the main

thermal-hydraulic plant variables.

Figures 3.1 to 3.9 show the available plant records, on

which this validation study is based. More information about

this incident can be found in references 3 and 4.

-9-
4.- CODE INPUT MODZL DZECRIpTION.

For this assessment analysis, the RELAP5/MOD2/36.04 code

(ref. 5) running on a CYBER 180/830 computer under NOS 2.5

operating system has been used. NEW, RESTART and STRIP modes

of operation were used for the steady state, transient and

plotter applications respectively.

As the DISSPLA plotter package was not available, the

reading-writing POSTRIP (post-STRIP) program has been

developed. This program reads from the "strip file" and

writes a file adapted to the input of GRAPHS (a general

purpose plotter program).

The RELAP5 model of JOSE CABRERA NPP nuclear steam supply

system (NSSS) depicted in figure 4.1 is currently being used

in the transient and safety analysis of the plant. It is a

general purpose model developed specifically for JOSE CABRERA

NPP in order to have a tool to allow the utility to do its own

in-house safety analysis.

The nodalization comprised 124 control volumes or nodes,

15 of which are time dependent volumes, 133 junctions and 63

heat structures.

A transient and accident analysis methodology adapted to

the use of the code, including engineering procedures and

simulation rules, has also been developed.

-10-
4.1.- Primary system nodalization.

The reactor core was divided into eight vertical nodes; a

six nodes pipe (209) representing the active core and two

unheated inlet (211) and outlet (207) nodes respectively. The

upper plenum (206) collects coolant coming from the core, from

the core bypass (210) and from the head of the vessel.

The vessel has a lower (four nodes annulus 208) and an

upper (204, 202) downcomer, a lower (212) and an upper (201)

dome, and an upper plenum (203, 205). Three bypass ways for

the coolant have been considered: core bypass (from 210 to

206), vessel head bypass (from 205 to 206), and cold leg - hot

leg bypass (from 208 to 100). By applying appropriate loss

coefficients, the specified flow distribution between core

flow and each bypass flow was met.

The hot leg was divided into two nodes (100 and 105), the

junction of wich corresponds to the surge line (three nodes

pipe 300) connection.

The pressurizer was modelled by two pipe components; the

two nodes upper one (312) and the six nodes lower one (310),

the connection of wich corresponds to the spray junction (from

354 to 310). This nodalization was chosen in order to allow

the model to introduce coolant spray from the pump discharge

(150) through- the-spray line (three nodes pipe 350, single

volume 354) directly into the steam volume under the influence

of the modulation of the spray control valve (352).

Heat structures to simulate both the pressurizer heaters

-11-
and the heat losses have been modelled. Also the continuous

spray mass flow rate of has been considered. There is an

equilibrium among this continuous spray (0.08 Kg/s),- the

pressurizer heat losses (15 Kw) and the power generated by the

proportional heaters power during the steady state (40 Kw).

The pressurizer relief lines (322, 326), valves (324,

328) and collector (330), as well as the safety lines (314,

318) and valves (316, 320) have been simulated. The four

nodes pipe common safety-relief collector (332) carries the

steam discharges into the pressurizer relief tank that was

simulated as a couple of volumes; the bottom one corresponds

to the water part (334) and the top one corresponds to the

steam-nitrogen part (336). The rupture disc was simulated by

a valve (338) having the disc real section and an opening

set-point equal to its rupture pressure. This valve allows

the discharge of steam directly to the containment atmosphere

simulated as a boundary condition (time dependent volume 340).

The primary side of the steam generator was modeled with

an inlet plenum (110), the portion of the tubes in the "up"

direction inside the tubeplate (115), the eight nodes pipe

(120) representing the U-tubes, the portion of tubes in the

"down" direction inside the tubeplate (125), and the outlet

plenum (130).

The loop-seal between the steam generator outlet and the

pump suction was simulated with three volumes corresponding to

the "down" part (140), the "horizontal" part (142) and the

"up" part (144).

-12-
The reactor coolant pump (150) was represented using the

specific homologous curves obtained from Westinghouse.

Two-phase factors from LOFT facility were used to simulate the

degraded behavior under abnormal conditions of void fraction

as an application of the conclusions of reference 6. The RCP

speed coastdown after a trip has been fitted by comparison

with plant reliable measurements of a specific test performed

at hot zero power conditions. The adjusted parameter has been

the internal friction torque and specially the TF0 parameter

of this torque in the RELAP5/MOD2 input deck. As a

consequence of that, a good agreement has been got between

measured and calculated mass flow rates (fig. 4.2 and 4.3).

The cold leg leading from the pump discharge to the

vessel inlet was represented by two nodes (160, 165). The

injection of the charging system was simulated by a time

dependent volume (164) and a time dependent junction (163).

By applying the appropriate friction and form pressure

loss coefficients, the thermal-hydraulic design reference loop

pressure distribution, total pressure drop and flow were

achieved.

The emergency core cooling system was simulated by a

couple of subsystems. The passive subsystem includes the

accumulator (600), the discharge line (605), the isolation

valve (610), the three nodes pipe discharge line (620), and

the check valve (630). By tuning appropriated coefficients

the referenced Westinghouse accumulator discharge mass flow

rate under LBLOCA conditions was met. The active subsystem

corresponds to the safety injection pumps, modeled as a time

-13-
dependent junction (655) taking borated water from the reload

water storage tank (time dependent volume 650) as a boundary

condition. The injection flow has been defined as a function

of the primary system back-pressure with conservative

assumptions for the line pressure losses.

Heat structures for the accumulator, vessel, reactor core

(average and hot channels), hot leg, surge line, pressurizer,

steam generator plena, U-tubes, loop seal, pump, spray line

and cold leg, have been simulated. In the case of the SG

plena, three different heat structures have been considered:

one connecting each plenum with the containment atmosphere,

one connecting both plena, and one connecting each plenum with

the riser, simulating the tubeplate thermal structure.

The point kinetic model, including best estimate fuel

temperature, coolant temperature and coolant density feed-back

reactivity effects, has been selected for the active heat

structures of the reactor core. Realistic data have also been

used in the estimation of the decay heat which is obtained

from the ANS-79 standard. The reactor control and protection

systems based on the functional diagrams corresponding to the

real gains and delays measured at the power station have been

simulated too.

4.2.- Secondary system nodalization.

Feedwater was simulated as a time dependent junction

(445), connected to the upper part of the downcomer, taking

warm water from the time dependent volume (444) that

-14-
represents the outlet of the 4 th heater. Feedwater

temperature was simulated based on at power and post-trip

operational data.

Turbine driven auxiliary feedwater pump was represented

as a time dependent junction (449), connected to the upper

part of the downcomer, taking cold water from a constant

temperature time dependent volume (448).

Emergency feedwater motor pumps were simulated as a time

dependent junction (457), connected to the lower part of the

riser, taking cold water from a constant temperature time

dependent volume (456).

The steam generator downcomer was simulated by a five

nodes annulus (450). The single junction (455) connects the

downcomer bottom to the riser inlet. The riser was

represented by a five nodes pipe (400) with the same

elevations as their counterparts in the downcomer. The first

four are thermally connected to the primary system through the

U-tube heat structure.

A non-ideal but nearly-real first separator (410) was

simulated at the top of the riser with special detail in the

carry-over and carry-under flow characteristics. To do that,

a geometrical analysis of the real dimensions of the ciclonic

pathways in the separators has been done so as to obtain the

values of the VOVER (carry-over) and VUNDER (carry-under)

parameters for the RELAP5 separator model. A connection (from

410 to 450) representing the separator draining paths has been

provided. The separator bypass (440), connecting the

-15-
downcomer and the steam dome, has been simulated.

By applying the appropriate friction and form loss

coefficients in the natural circulation loop of the steam

generator (400, 410, 450), with the highest resistance located

in the downcomer/riser junction (455), the specified

circulation ratio of 1.96 has been met. Also, by adjusting

the secondary side liquid inventory, the measured downcomer

level has been obtained.

The steam node (420) corresponds to the volume between

the ciclonic separator and the steam dryer. The dryer was

simulated as a nearly-ideal second separator (424) which

allowed nearly-only steam to escape upwards. The drain flow

path (426) represents the real pipes that connect the steam

dryer to the top of the downcomer.

The steam volume at the top of the steam generator dome

has at its bottom a plate of orifices that behaves as a

separator and so it has been simulated as an ideal third

separator (430) allowing only steam to escape upwards. The

drain flow path (428) represents the real pipes that connect

the plate of orifices to the top of the downcomer.

When defining the scope of the plant model, this special

emphasis in the simulation of the three separator stages,

including the real definition of the draining ways, was

considered to be important in the analysis of

depressurizations of the secondary system due to steam line

breaks. These draining pipes behave in such an event as a

riser bypass leakage pathways for the inventory of the steam

-16-
generator that leaves the downcomer without any cooling effect

on the primary coolant through the riser/U-tubes thermal

connection.

The steam line was divided in several parts, a four nodes

pipe (500), two single nodes (502, 504), the main steam

isolation valve (506), a single node ( 508), a three nodes

pipe (510), the turbine trip valve (512), a single node (513),

the main steam control valve (514) and the time dependent

volume (516) that represents the turbine.

The turbine was simulated as a boundary condition

selecting its constant back-pressure high enough just to avoid

critical flow in the main steam line valves.

The real characteristics and actuation logic of each

valve have been modeled. Also, by using the appropriate

friction and form loss coefficients, the reference secondary

pressure distribution was met.

The model includes four safety valves (540, 544, 548 and

552) as well as their relief lines (542, 546, 550,and 554) to

the environment atmosphere simulated as constant time

dependent volumes (560, 561, 562 and 563).

The steam consumption of the turbine driven pump has been

simulated by time dependent junctions discharging to the

environment atmosphere (time dependent volumes 460 and 462)

for the turbining (459) and turbining-pumping (461) modes of

operation.

-17-
The automatic steam-dump system modulates the opening of

the relief valve (532) to the condenser (time dependent volume

538) through the relief line (530, 534) and the opening of the

relief valves (522 and 526) to the environment atmosphere

(time dependent volumes 564 and 565) through their relief

lines (524, 528), looking for the RCS no-load programmed

average temperature. There is a common relief line (520) to

the atmosphere and a general common relief line (518) from the

main steam line. A valve (536) isolates the relief line to

the condenser in case of loss of offsite A.C. power,

protecting the condenser that would be unavailable under this

circunstance.

Heat structures for the steam generator vessel and

internals have also been simulated.

A sensitivity calculation tuning the hydraulic diameter

of the steam generator U-tubes/riser heat structure was done

to fit the pressure in the secondary side at nominal power.

The explanation for this correction can be found in the

substantial amount of crossflow created by the U-tubes support

plates and the "U" curve itself inside the riser. The

crossflow enhances the heat transfer considerably and is not

taken into account in the standard heat transfer correlations.

The tuned hydraulic diameter corresponds to a value

similar to the gap between tubes. This value was only used in

the definition of the U-tubes/riser heat structure,

maintaining the real geometric value of the hydraulic diameter

for the definition of the riser volumes.

-18-
A summary description of the model including concept,

node number and type is given in Table 4.1.

4.3.*- Trips and control variables.

The software of the reactor protection system has been

modelled with a set of trips that forces the reactor scram

when necessary.

A wide range of control variables has been defined to

model the different control systems that work in the plant:

a) RCS average temperature control system:

Before the reactor trip, this system controls the rods

bank insertion and withdrawal taking into account:

- the difference between the primary average

temperature and the programmed one as a function

of the turbine load determined by the pressure in

the impulse chamber.

- the difference between the primary pressure and

the reference one.

After the reactor trip, this system controls the steam

dump from the steam generator to the atmosphere and to

the condenser by following a primary temperature

programme, as well as the stem position of the

feedwater valve by following a RCS temperature

-19-
hysteresis cicle.

b) Pressurizer pressure control system:

This system controls the primary pressure by acting on

the s.pfay, PORV's and heaters, looking for maintaining

the reference pressure.

j

c) Pressurizer level control system:

This system controls the pressurizer level by acting

on the speed of the CVCS charging pumps, looking for

maintaining the programmed level as a function of the

RCS average temperature.

d) Steam generator level control system:

This system controls the level of the steam generator

by acting on the stem position of the feedwater

control valve, looking for maintaining the constant

reference SG level.

For an easier comparison with plant measurements in

assessment cases, all the important Control Room

instrumentation has been simulated with their actual units and

delays (2.0 sec. for liquid temperature, 20.0 sec. for steam

temperature, 1.0 sec. for liquid and steam mass flow rate, and

0.5 sec. for pressure).

-20-
4.4.- simplifications.

In order to save CPU time in the specific transient

analyses of this report, some simplifications have been done

in the general purpose model to generate a reduced ad-hoc one

(fig. 4.4), some control volumes and junctions being deleted

(for instance, and due to the fact that this is a

depressurization transient, the pressurizer relief lines,

PORV's, safety lines and valves as well as the relief tank

have been eliminated), the final nodalization having 73

control volumes, all of which remain as previously defined

excepting the spray valve that has been removed and

substituted by a constant mass flow rate TMDPJUN.

-21-
5.- STEADY STATE CALCULATION.

The first step was to get a steady state condition

representing the normal operation of the plant at full power.

The aim was to get the desired stable condition with the

minimum CPU time consumption. The simulated control system,

that reproduces the real characteristics of the system in the

plant, behaved slowly with high code running time and was not

considered the more appropriate for getting the steady state.

To that aim, a stabilization system has been developed

according to the logic of actuation represented in Table 5.1.

In this way, the pressurizer pressure and level, primary

coolant flow, primary average temperature, steam generator

pressure and steam generator downcomer level have been fitted

to the values measured in the plant.

This stabilization system was based on controllers that,

by using properly selected gains, fitted the system parameters

in such a way that the error signal (defined as the difference

between the desired and the calculated value of each

controlled variable) was reduced to zero as fast and stably as

possible.

Reactor power and feedwater temperature were maintained

constant as boundary conditions representing the nominal

values corresponding to the normal operating condition. For

this first calculation the kinetic model was not used.

After 346.5 sec. reactor time steady state calculation a

fully stable condition for the controlled model has been got.

-22-
The final state has fitted the desired full power measured

operating conditions. The RELAPS/MOD2 code stopped the steady

state calculation once the stability condition was accepted by

its internal checking procedure.

Then, a second step eliminating the stabilization system

and introducing the reactor kinetic model and the real reactor

control and protection systems was carried out. The stable

condition was reached inmediately. Table 5.2 gives a summary

of the comparison between the main variables measured and

calculated in the steady state simulation at full power.

Since the plant conditions at the beginning of the event

were not exactly those corresponding to 100% power, but to

96%, a new steady state corresponding to these conditions was

obtained with the same methodology as explained above.

Because this steady state calculation started from the steady

state at 100% power, the new desired conditions at 96% power

were found very quickly. These new steady state results,

which were used as the start point for the transient, are

shown in table 5.3 compared to the measured values.

-23-
6.- TRANSIENT RESULTS AND COMPARISON WITH PLANT DATA.

The purpose of this analysis was to compare the response

of the RELAP5/MOD2 code with the selected plant transient,

focusing the attention on the following aspects of the RCS:

- depressurization due to the continuous spray,

- response to the cooldown from the steam generator,

- response to the RCP trip,

- natural circulation evolution.

As is regular practice in JOSE CABRERA NPP, the operation

team worked most of the control systems. So these systems

didn't perform according to their automatic logic.

The control systems behaviour during the transient up to

the 2150.0 sec. studied in this simulation may be summarized

as follows:

- RCS average temperature control system:

It worked in automatic mode for the first 435.0 sec. in

the transient (up to the reactor trip), reducing power

to match the manual turbine load decrease.

- Pressurizer pressure control system:

It worked in automatic mode during the whole transient.

From 87.0 sec. on, all the pressurizer heaters worked

at full power (300 Kw).

-24-
- Pressurizer level control system:

Before the reactor trip (435.0 sec.), the reactor

operator started the second charging pump manually and,

from then on, the whole system was manually operated to

maintain the desired pressurizer level. Taking into

account the operator intervention, the charging flow

has been simulated as a TMDPJUN following the plant

register data (fig. 3.6).

- Steam-dump control system:

Just after the turbine trip, the reactor operator

switched the system from RCS average temperature

control mode to SG pressure control mode, setting the

adequate pressure setpoint to obtain the desired RCS

average temperature. Taking into account that, because

of the high capacity of the steam-dump valves, the SG

pressure response reached in a few seconds the demanded

one the measured SG pressure has been used as the

reference for the simulation of the actuation of the

control system.

- Steam generator level control system:

Once the turbine was tripped, it was manually

controlled by the turbine operator to recover the steam

generator downcomer level and to maintain it within an

acceptable range.

In the calculation with RELAP5/MOD2 the automatic and

-25-
manual modes of the control systems behaviour have been

simulated, simplifying their modelation as much as possible,

especially after the operators took the control of the

systems. Anyway, as mentioned above, the main aim of the

calculation was not to validate the control systems

performance, as other plant transients would be more suitable

for this purpose.

Therefore, the response of the pressurizer level and the

steam generator pressure have been fitted with special care.

Great attention has not been given to the steam generator

level evolution, whose impact on other variables, specially in

the primary circuit, is not very important as has been proved

by sensitivity calculations.

The spray mass flow rate has been set as a TMDPJUN in

order to obtain with the RELAP5/MOD2 models related to this

phenomenon and specially with the condensation one, the same

depressurization rate that was measured at the plant until the

reactor trip. The assumption of a constant mass flow rate

TMDPJUN model may be justified taking into account the

following considerations:

- There is a constant geometry in the RCS between the

discharge of the main coolant pump and the spray nozzle

in the pressurizer.

- The failure of the spray valve forced a fully open

position with constant area from the begining of the

transient on.

-26-
- The density of the fluid in the cold leg was

practically constant because of the small variation in

the temperature and the reduced impact of the

depressurization on it.

- A previous calculation of the same transient with

TRAC-PFI/MODI (ref. 7) in which the spray valve was

simulated with a VALVE component showed an almost

constant calculated mass flow rate.

It has been shown that, with this hypothesis, the more

suitable value of the equivalent mass flow rate at nominal

conditions is 2.7 Kg/s, in contrast with the design mass flow

rate of 4.8 Kg/s (reference documentation). It is important

to point out that the real spray flow can not be measured in

the plant and so it is not well defined but only referenced

from the Westinghouse design data.

The discrepancy between both flows might be due to two

different reasons:

- The three-dimensional (3D) characteristics of the

condensation phenomenon caused by the spray (ref. 8):

* Geometry of the nozzle and distance to the

pressurizer level that may be divided into three

spray zones: the continuous liquid. zone, the

break-up zone and the droplet zone.

* Mean drop diameter, velocity, spray angle and

break-up length as a function of the spray mass

-27-
flow rate.

* Form of the liquid sheet - characterized by the

spray angle and the break-up length - that is

responsible for the final form and efficiency of

the spray.

* Break-up length. of the liquid sheet into a swarm

of fine droplets as a function of Weber and Jacob

numbers. Due to the mechanism of aerodynamic

instability, the produced drops are of very

different sizes and fly at different velocities.

It may be summarized that in order to describe the

spray, measurements of the shape and length of the

liquid sheet and of the size and velocity distributions

of the droplets are necessary.

- The limitation of the condensation correlations of the

RELAP5/MOD2 code when analysing the condensation

process only in one dimension (MD).

A similar 1D code (TRAC-PFl/MOD1) yielded a calculated

spray mass flow rate of 2.9 Kg/s (ref. 7), in fact very close

to the RELAPS/MOD2 value of 2.7 Kg/s, in such a way that it

may be concluded that it is not a problem of models and

correlations but of physical effects and/or uncertainty in the

real mass flow rate.

The time sequence of the transient's more significant

events is shown in table 6.1.

-28-
At time t=0.0 sec. the spray valve was completely open,

and a RCS depressurization began due to the pressurizer steam

cooling and condensation (fig. 6.1). This depressurization

activated the pressure control system in such a way that, at

time 87.0 sec., all the pressurizer heaters were working at

full power and maintained this situation until the end of the

simulation.

As the depressurization continued, the shift supervisor

asked the turbine operator to decrease the turbine load in

order to create a reactor/turbine mismatch to force an

increase of the RCS temperature during the first moments in

order to reduce the pressure decrease rate. He therefore

throttled the turbine control valves,- forcing a pressure

increase in the steam generator.

The RCS average temperature control system started

working when this temperature was higher than the system dead

band, decreasing the nuclear power by means of control bank B

insertion (fig. 6.2, 6.3 and 6.4). The throttling action

produced an average temperature increase of between 0.5 and

1.0 "C in the primary circuit and a later temperature decrease

by following automatically the program of average temperature

in the RCS as a function of the turbine load. In this way the

turbine load decreased and, therefore, the nuclear power

decreased from 96%, the starting point, to 85% of- the rated

value.

The mentioned RCS average temperature increase produced a

primary coolant swelling and the corresponding pressurizer

-29-
level increase (fig. 6.5). The level control system, still

performing in automatic mode, decreased the flow of the

controlled CVCS charging pump. Afterwards, when the RCS

average temperature decreased and the pressurizer level

decreased consequently, the same pump automatically increased

its flow. From this moment to the reactor trip, which occurs

very soon, the second charging pump was manually started and

the reactor operator tooks control of the pressurizer level.

The manually started charging pump, injected its maximum flow

of 240.0 1/min into the system.

The primary system depressurization continued until the

low pressurizer pressure reactor trip set point of 125.5

* 2 Kg/cm (rel.) was reached (fig. 6.1). As a result of the

reactor trip, the turbine tripped and the steam-dump valves

tried to get the no-load temperature in the RCS. Some seconds

later, when the primary average temperature was lower than

280.0 OC, the main feedwater control valve began to close, in

order to obtain a mass flow rate of 5% of the nominal one in

70.0 sec. time.

The hot and cold leg temperature decreased and,

therefore, the RCS average temperature reduction (fig. 6.2,

6.3 and 6.4), produced a pressurizer level decrease (fig.

6.5), that was partially compensated by the charging pumps

both working in their manual mode. In normal conditions, the

level after the reactor trip would have automatically reached

a value of 18% of the span. In the transient, under the

manual influence of the reactor operator, a minimum level of

25% of the span was reached. It must be noted that the

primary system pressure decreased sharper in the simulation

-30-
than in the plant. That was because the pressurizer emptying

and refilling caused a sharper depressurization and

pressurization in the analytic results with RELAP5/MOD2 than

observed, as was also demonstrated with the NEPTUNUS Y-05

pressurizer experiment and its subsequent study with this code

(addendum A).

It has also been observed, as in other transients, a

calculated steam generator level drop larger than measured in

the plant after the turbine trip (fig. 6.7). This was due to

the dynamic effects on the instrumentation, to the way the

level in the RELAP5 model was defined (sum of all the

downcomer cells liquid fraction multiplied by their heights),

to the nodalization of the decreasing area annulus node

(450-01) which was between the upper part and the lower part

of the downcomer and finally to the uncertainty in the

circulation ratio and feedwater flow after the reactor trip.

The feedwater flow evolution corresponding to the closing

of the control valve from its nominal position to the 5% one

has been considered to be linear, although it would be logical

to assume that in the first phase of the closing transient the

flow would decrease very slightly with the result of a value

higher than simulated. So, in the analysis, the integrated

mass introduced has been supposed to be lower than it would

actually be at the plant (fig. 6.8). The feedwater

temperature may also have some effect; the best estimate

temperature evolution coming from other turbine trip analysis

has been adopted.

Less than one minute after the turbine trip, the

-31-
steam-dump control system was manually set into its SG

pressure control mode. The reactor operator tuned the

pressure setpoint to meet the desired RCS temperature

evolution, which is equivalent to open the steam-dump valves

when needed.

In the first phase of the steam-dump manual control the

steam generator pressure is allowed to *increase (fig. 6.6),

with the subsequent RCS temperature and pressurizer level

increase in the primary circuit. This fact, in combination

with the additional pressurizer level increase as a result of

the charging pumps manually working, forces a RCS

repressurization which, for the same reason mentioned

regarding the emptying after the reactor trip, is sharper in

the simulation than observed in the plant (fig. 6.1).

After the partial refilling, the pressurizer level

changes as a function of the primary average temperature, with

small deviations due to manual and not very severe charging

flow oscillations around the letdown flow value (fig. 6.5).

There is a phase in the simulation, between 800.0 sec.

and 1150.0 sec., with a pressurizer pressure sharp decrease

and later increase while the primary temperature and

pressurizer level are increasing (fig. 6.1, 6.2 and 6.5).

This tendency, apparently against the physic laws, appears in

the plant records too, but in a smoother manner. It has been

thought that this abnormal tendency could be due to local

effects in the pressurizer, which had been refilled with

subcooled water coming from the hot leg. This fact, together

with the pressurizer water stratification and temperature

-32-
axial gradient that forces the homogenization of the

stratified water zones, the interphase heat transfer and the

pressurizer walls temperature effects, justifies this pressure

response.

The calculated depressurization that appears between

1150.0 and 1350.0 sec. is sharper than measured in the plant.

This discrepancy has been considered to be due to a code

malfunction as a consequence of the calculation of very cold

water introduced in the node 6 of the pressurizer when the

node 5 is just filled of water and the pressurizer level

increasing continues. The very cold water temperature

calculated at node 6, colder than the water at node 5 and

colder than the previous calculated spray temperature at node

6 after its heating in the steam condensation proccess,

results very near to the pressurizer inlet spray temperature.

This non-physical-meaning code malfunction could correspond to

the same error observed in the RELA5/MOD2 simulation of the

MIT pressurizer test (addendum A) when the pressure evolution

was abnormally calculated at the time the level was going from

one node to the next one.

After this phase, from about 1350.0 sec., all the

variables evolution meets the steam generator pressure

evolution very well and with its effect on the primary

temperature (fig. 6.6).

Between 1450.0 sec. and 1550.0 sec. a significant

steam-dump manual depressurization in the steam generator is

forced by the reactor operator (fig. 6.6), producing an RCS

average temperature decrease (fig. 6.2). Approximately at

-33-
that time, the system's lowest.pressure is reached, about

100.0 Kg/cm 2(rel.) according to the plant records. Anyway,

this pressure was greater than the one at which the safety

injection begins injecting water into the system because a

value of 0.0 m3/h safety injection flow was observed. In the

simulation, the lowest pressure is about 2.0 Kg/cm2 lower than

measured and would have forced the ECCS water injection into

the reactor pressure vessel if the corresponding TMDPJUN would

not have been inhibited previously in the input deck.

At 1550.0 sec. the shift supervisor asked the reactor

operator to trip the RCP once the problem has been clearly

identified, in order to stop the spray flow. The main pump

takes about 80.0 sec. to stop and, meanwhile, the spray flow

decreases until it eventually stoped. Once the

,depressurization cause is finished and also due to the pump

trip, the RCS begins increasing its pressure (fig. 6.1).

The hot leg temperature increases due to the RCP trip,

but the cold leg temperature is not affected in the short

term. This produces a primary average temperature and primary

delta temperature increase (fig. 6.2 and 6.4).

At 1650.0 sec. the steam-dump is manually opened by the

reactor operator in order to control the primary heating. The

sharp cooling it produces can be seen more quickly and is

sharper in the cold leg than in the hot leg, due to the low

primary flow in natural circulation, and so, although the

average temperature decreases, the delta temperature

increases. This delta temperature increase, in combination

with the one described above (the first one due to the hot leg

-34-
temperature increasing, the second one due to the cold leg

temperature decreasing), produces a very significant peak of

the delta temperature, about 30.0 "C at the plant and 20.0 *C

in the simulation (fig. 6.2, 6.3 and 6.4).

The difference between these two values has been

concluded to be due in some way to different causes:

- A sharper or slightly shifted in time steam generator

pressure evolution as a consequence of the turbine

operator intervention.

- The records system reliability for such abrupt changes.

- The actual lack of hot leg axial temperature

homogeneity, specially under low flow conditions.

- The temperature instrumentation locations in the upper

part of the hot and cold legs for delta temperature

calculation in the plant.

The last two reasons create a 3D problem beyond the ID

design bases of the RELAP5 code. This effect of axial thermal

gradient in the hot leg combined with the instrumentation

location has been previously observed in the plant, even under

RCS nominal flow with the RCP running, when there is a

pressurizer level decrease because of a malfunction of the

CVCS and saturated water goes out of the pressurizer, is not

mixed and flows through the upper part of the hot leg long

enough to activate the reactor trip on "variable low

pressurizer pressure" that is equivalent in this nuclear

-35-
station to the OTAT reactor trip of a standard Westinghouse

3-loop PWR.

After that, stable natural circulation is intended. Of

course, it is affected by the reactor operator manual action

on steam generator pressure through the manual control of the

steam-dump valves. The natural circulation primary delta

temperature is established at about 12.0 "C at plant and 13.0

"C in the simulation (fig. 6.4). That is because the RELAP5

primary model has been fitted to the reference pressure drop

that corresponds to the thermal design flow (75000. gpm),

8.5% lower than the actual flow. This actual flow is about

82000. gpm, but it is difficult to know it exactly. That is

just a direct consequence of the overestimation in the

pressure drop coefficients of the RCS based on the

Westinghouse design references. 'The RCS nominal flow

deviation justifies the small discrepancy observed in the

simulation in the stabilized delta temperature under natural

circulation: the RCS flow is lower and the delta temperature

is higher than in the plant measurements.

The RELAP5/MOD2 simulation confirms the lack of a steam

bubble formation in the vessel head due to the RCP trip under

no-load RCS temperature and low pressure conditions.

The simulation finishes at 2150.0 sec., with the RCS

pressure being recovered. This pressure is, at the end of the

calculation, 106.0 Kg/cm2 and. it is increasing (fig. 6.1).

Natural circulation is established, although it is affected by

the manual depressurization of the steam generator performed

by the reactor operator.

-36-
A good agreement between the code results and the data

recorded in plant can be observed in all the transient phases,

specially considering that all the interesting phenomena are

clearly shown and the most significant discrepancies have been

explained according to the RELAP5 models or the way some

boundary conditions have been assumed. The higher primary

temperature discrepancies are about 2.0-3.0 °C, and the

2 greater RCS pressure discrepancies are about 5.0-6.0 Kg/cm

The worst agreement corresponds to the steam generator level,

but even that variable shows about 30.0-35.0 cm. as the

greatest discrepancy between plant and simulation.data, which

is considered acceptable.

Finally it is important to point out that the code

assessment conclusion depends on the quality of the records.

The script type thickness with the nominal recording speed of

1 inch/hour, which really is a little bit different from one

record to another, produce an uncertainty of about 70.0 s. The

uncertainty about the value of the variables depends on the

scale .of each one and is higher the higher the scale. Finally

the chronology of events of the NPP trip report is just

approximate. All these reasons make very difficult to

describe the sequence of events with a greater accuracy.

Figures 6.1 to 6.9 show, in plant units and with the

corresponding instrumentation delays, the comparison between

the real data and those obtained in simulation. The remaining

figures, from 6.10 to 6.25, have been attached to make the

transient more easily understandable. Table 6.2 shows the

variables identification for the transient figures.

-37-
7.- RUN STATISTICS.

During the first 1650.0 sec. of the transient, up to

100.0 sec. after the RCP trip, the maximum allowed time step

was 0.5 sec. From this time to the end of the simulation,

500.0 sec. more, the maximum allowed time step was relaxed to

1.0 sec., as the Courant limit was less restrictive during

this phase because of the low velocity of the fluid in the

primary circuit.

Figure 7.1 shows the CPU time comsumption as a function

of real time. Twice in the transient the integral CPU time

goes to 0.0 sec. and corresponds to the time when a restart

with some changes in the input file is made.

The run statistics summary is shown in table 7.1. The

average CPU time (35172. sec.) to real time (2150.0 sec.)

ratio was 16.36 to 1.

-38-
8.- CONCLUSIONS.

RELAAP5/MOD2-cycle 36.04 has been used for an assessment

case, comparing the analysis results with plant measurements

of a real incident that took place in August 1984. The

reference nuclear plant, JOSE CABRERA, has been nodalized and

this nodalization is being used in a wide range of

thermal-hydraulic applications. Both code and the model have

worked friendly and properly.

The comparison with plant measurements has been difficult

due to the lack of a process computer, the quality of the

registers and, consequentely, the data uncertainties.

Nevertheless, the selected transient has a big interest both

generic (because of the thermalhydraulic phenomena that took

place) and specific (behaviour of the single-loop nuclear

power plant).

It has been possible to reproduce the main events in the

transient and reach a good agreement between the simulation

and the plant data. The steam generator level has been the

variable in which a lower fitting effort has been done. It

was proved that the highest discrepancy of 20.0-30.0 cm. has

not a big impact on the other primary and secondary variables.

Concerning RCS temperatures and pressure the highest

discrepancies have been 2.0-3.0 OC and 5.0-6.0 Kg/cm2

respectively.

The spray mass flow rate has been tuned to fit a

depressurization rate similar to the one in the plant before

the reactor trip, The parameter was kept constant for the

-39-
rest of the pump running time. The spray calculated value is

lower than the reference design one. The same applies to

TRAC-PFl/MODl, for which the fitted value is close to the one

of RELAPS. The discrepancy is due to the one-dimensional

characteristic of RELAP5 studying a tri-dimensional phenomenon

and mainly to the fact that the reference is a design value

that can not be measured in the plant. However, with a well

tuned spray equivalent mass flow rate, it is not neccessary a

3D pressurizer model to simulate its effect on the RCS

pressure.

Some experimental results (NEPTUNUS Y-05 and MIT

pressurizer tests) have been simulated to check the code.

From their results with no uncertainty about the spray flow it

may be concluded that the code underpredict the spray

condensation efficiency. Based on this evaluation and the

results of the plant simulation it might be concluded that the

real spray flow in the plant is lower than the fitted one, so

much lower than the design reference value. However, to

support this conclusion, the previously mentioned 3D phenomena

should be considered.

As in the simulation of the experiments, the plant RCS

depressurization and pressurization due to pressurizer partial

emptying and refilling were sharper in the analysis than in

the plant measurements and also an abnormal pressure responsiG

appears when the pressurizer level changes from one node to

the next one.

It is important to- remark the big effect of some

pressurizer local phenomena (such as cold water stratification

-40-
with axial temperature gradient, heat transfer in the

water/steam and spray/steam interphases', heat thansfer from/to

the walls, heat slabs properties, heat losses to the

containment atmosphere) on the.RCS, pressure. A detailed

pressurizer nodalization is required to properly simulate

these effects.

The response of the steam generator level immediately

after the turbine trip is sharper in the simulation than in

the plant. This might be due to the way of modelling the

collapsed liquid level in the steam generator downcomer.. The

necessity of modelling the downcomer delta pressure

instrumentation or, at least, of giving a good table of level

as a function of the mass inventory, specially in the zones

with complex geometry, has been detected.

In general, the primary parameters have followed very

well the cooldown induced from the steam generator. The very

high peak of delta temperature in the RCS after the main

coolant pump trip has been obtained and also the stable delta

temperature in natural circulation has been properly

calculated without a vessel head steam bubble formation. This

means that the friction and form pressure losses coefficients,

fitted with the pump running, are good enough to reproduce the

pressure losses in natural circulation with a Reynolds number

much lower.

The final calculation has been carried out after several

sensitivity Analyses. In spite of the duration of the

transient, these analyses have been possible because of the

good CPU time to real time ratio, specially after the RCP trip

-41-
when the CPU time consumption decreased a lot because the

maximum allowed time step could be higher due to the lower

restriction of the Courant limit. Also, the CDC CYBER 180/830

computer has performed successfully with RELAP5/MOD2.

-42-
9 - REFERENCES.

1.- L., Rebollo, UNION FENOSA,."RELAPS/MOD2 analysis 2 A

postulated 'cold "~g SBLOCA' simultaneous to _ 'total

black-out, event in JOSE CABRERA nuclear station as an

aDDlication .2 'lessons learned' from OECD-LOFT LP-SB-3

experiment. Development of a mitiQation procedure.", ICAP

report, October, 1987.

2.- Josd Cabrera Nuclear Power Plant Final Safety Analysis

Report (1968).

3.- J. Fornieles, UNION FENOSA, "Central Nuclear Jose Cabrera:

Informe de disparo. de la central por disparo del reactor

(motivado pM bala presi6n, causada por la apertura local

de ja PCV-400A)", ID-95, Septiembre, 1984.

4.- R. de la Vega, CSN, Acta de Inspecci6n AIN/123/JC/84 del

Consejo de Seguridad Nuclear, Septiembre, 1984.

5.- V.H.Ramsom & R.J. Wagner. "RELAPS/MOD2 Code Manual."

NUREG/CR-4312, EGG-2396, INEL, USA, December, 1985.

6.- J.B. Adams, D.L. Batt & V.T. Berta. "Xnfluence of LOFT

Pn Simulations 2A Thermal-Hydraulic Aspects 2f Commercial

En Safety", NUCLEAR SAFETY, April-June, 1986.

7.- R. Martinez, TECNATOM, "Assessment of RAC-PFI/MODI

against an inadverted Dressurizer ssray total oveninq jn

the JOSE CABRERA power plant", ICAP report, December,

1990.

-43-
8.- A. Ch~vez, F. Mayinger, Technische Universitdt Minchen,

"B_•av characterization and direct measurement of drop

size and velocitv distributions using pulsed laser

holocraDhy", European Two-Phase Flow Group Meeting,

Brussels, 30-31 May and 1 June 1988.

-44-
TABLE 4.1_RELXS/MOD2 model description (nodalization).

CONCEPT

PRIMARY COOLAMT SYSTEM

NODE NODE

NUMBER TYPE

Cold leg - pump discharge ...................... Charging flow (CVCS) ................. CVCS tank ..................... o. .......... ..

Cold leg - vessel inlet ... ......... . .....

Vessel downcomer - lower part .............. Vessel bottom . . ........ o............ o......o

Core lower plenum ....... .0.V.0....

Core inlet junction .o. .... . ....

Reactor core ....... _.o. ... .... ........... Core outlet junction .............. . ........ Core upper plenum .... .... .. .o.o............

Core bypass ... . ................... o .... . Reactor outlet ........................... Vessel downcomer junction ............. Vessel downcomer - ECCS injection Vessel downcomer junction ....... .... Vessel downcomer - upper part ................. Vessel top .......... Vessel upper-upper plenum .................. Vessel upper plenum .... ooo..o... ... . ... o......

Hot leg - vessel outlet .... . .... .o.o. ... . .... Hot leg - steam generator inlet ...... eo.......

S.G. tube inside the tube-plate (hot)

S.G. tube connected to the riser .. S.G. tube inside the tube-plate (cold)

S.G. outlet plenum .. ... . ..... . .... ... .o.o

Cross over leg - S.G. outlet ............ .. Cross over leg - intermediate o....

Cross over leg - pump inlet ....................

Primary coolant pump ...... Pressurizer surge line....

Pressurizer inlet junction Pressurizer vessel (lower part) ............. Pressurizer junction .. ... ..................... Pressurizer vessel (upper part) ......... Pressurizer safety valve 1 inlet

Pressurizer safety valve 1 ................. Pressurizer safety valve 2 inlet .............. Pressurizer safety valve 2

Pressurizer relief valve 1 inlet'. .... oooo[oo.

Pressurizer relief valve 1 .................. Pressurizer relief valve 2 inlet .............. Pressurizer relief valve 2 .... . ...... Pressurizer relief collector .... .......... Pressurizer relief-safety collector ...........

160

163

164

165

208

212

211

243

209

244

207

210

206

245

204

246

202

201

203

205

100

105

110

115

120

125

130

140

142

144

150

300

301

310

311

312

314

316

318

320

322

324

326

328

330

332

branch

tm.dp. jun. tm.dp.vol.

branch

annulus

branch

single vol.

single jun.

pipe

single jun.

single vol.

pipe

branch

single jun.

annulus

single jun. annulus

branch

branch

branch

branch

branch

branch

branch

pipe

branch

branch

branch

branch

single vol.

pump

pipe

single jun.

pipe

single jun.

pipe

branch

valve

branch

valve

branch

valve

branch

valve

branch

pipe
TABLE 4.1 RELPSZ/MOD2 model description (nodalization). (Cont.)

CONCEPT NODE NODE

NUMBER TYPE

Pressurizer relief tank (water volume) ......... Pressurizer relief tank (steam volume) ......... Pressurizer relief tank rupture disc ........... Containment atmosphere ......................... Cold leg - spray line connection...............

Pressurizer spray line (up) .................... Pressurizer spray valve ........................ Pressurizer spray line (down) .................. Accumulator .................................... discharge line .....................

Accumulator isolation valve .................... Accumulator discharge line................

Accumulator discharge line check valve...

Reload water storage tank ....................

Emergency core cooling system (pumps) ..........

SECONDARY SYSTEM

S.G. downcomer ....... 0 ........ ............. ... .. S.G. downcomer - riser connection ............. S.G. riser .............. ................ ..... S.G. separator ...... .... . ... . ........... .. ... S.G. separator bypass . ........................

S.G. separator outlet steam volume ............. S.G. first dryer .. ... o....o.........o.....o........

S.G. first dryer draining pipe .. S.G. second dryer (orifice plate) .. S.G. second dryer draining pipe ............

S.G. feedwater tank ........ ....... . .... ......

S.G. feedwater flow .... o.. .. .................

AFWS turbine-driven pump FW2 suction tank ......

AFWS turbine-driven pump FW2 junction .......... EFWS motor pump suction tank ........... EFWS motor pump junction ................... Steam line (S.G. outlet) ...................... Steam line ........ .. ......... ............ Steam line ...................... o...............

Steam line isolation valve

Steam line ...... Steam line o............ ................. o........

Turbine trip valve ......................... Steam line ............ ....... .... ...... ... . Main steam control valve Turbine ... ..................................... o

S.G. safety valve 1 .........................

S.G. safety valve 1 discharge line

Environmental atmosphere ...................

334

336

338

340

349

350

352

354

600

605

610

620

630

650

655

450

455

400

410

440

420

424

426

430

428

444

445

448

449

456

457

500

502

504

506

508

510

512

513

514

516

540

542

560

branch

branch

valve

tm.dep.vol

single jun.

pipe

valve

branch

accumulator

singl vol.

valve

pipe

valve

tm.dep.vol.

tm.dep.jun.

annulus

single jun.

pipe

separator

branch

branch

separator

branch

separator

branch

tm.dep.vol

tm.dep.jun.

tm.dep.vol

tm.dep.jun.

tm.dep.vol.

tm.dep.jun.

pipe

branch

branch

valve

branch

pipe

valve

singl vol.

valve

tm.dep.vol.

valve

branch

tm.dep.vol.
TABLE 4.1 RELAPS/MOD2 model description (nodalization). (Cont.)

CONCEPT NODE NODE

NUMBER TYPE

S.G. safety valve 2 ..... ................. S.G. safety valve2 discharge line ............

Environmental atmosphere ... S.G. safety valve 3 ......................... S.G. safety valve 3 discharge line

Environmental atmosphere . S.G. safety valve 4 o .... ..................

S.G. safety valve 4 discharge line ... .......... Environmental atmosphere .................... .

AFWS FW2 steam consumption (injecting)

Environmental atmosphere ............ .

AFWS FW2 steam consumption (not injecting) Environmental atmosphere

Steam relief line ... ..... o...o.....

Steam relief line to the atmosphere ... S.G. relief valve 1 (to the atmosphere) .......

S.G. relief valve 1 discharge line ......

Environmental atmosphere S.G. relief valve 2 (to the atmosphere) S.G. relief valve 2 discharge line

Environmental atmosphere

S.G. relief line to the condenser .... S.G. relief valve 3 (to the condenser)

S.G. relief valve 3 discharge line

Condenser isolation valve ... ...... o .o.... o......... o..........

544

546

561

548

550

562

552

554

563

459

460

461

462

518

520

522

524

564

526

528

565

530

532

534

536

538

valve

branch

tm.dep.vol.

valve

branch

ti.dep.vol.

valve

branch

tm.dep.vol.

tm.dep.jun.

tm. dep.vol. tm.dep.jun.

tm.dep.vol.

branch

single vol.

valve

branch

tm.dep.vol.

valve

branch

tm.dep.vol.

branch

valve

single vol.

valve

tm.dep.vol.
TABLE 5.1 Stabilization system actuation logic

* VARIABLE TO BE FITTED * CONTROL SYSTEM ACTION * * * ."

* Primary pressure * Time dependent .volume connected to the *

* * pressurizer-steam dome *

* Primary coolant average temperature * Main steam control valve modulation *

************************************************************************************** *****************

* * •

* Primary system mass flow rate * Reactor coolant pump speed modulation *

* Pressurizer level * Make-up andlet-down (CVCS) modulation *

Steam generator downcomer level * Main feedwater modulation

* Primary - secondary temperature difference. * *

* * Parametric analysis of the hydraulic diameter *

* Steam generator pressure * *
TABLE 5.2 Steady state results at nominal conditions.

* Variable Units k JOSE CABRERA * RELAP5/MOD2*

* (measured) *(calculated)*' * ** * *

* Reactor power (Mw) 510.00 * 510.00 *

SRCS average teldperature (AK 566.60 * 566.60 , ** e

* Pressurizer level M() 64.00 * 64.00 - ft* * 82

* Pressurizer pressure WPM 13.82 * 13.82 k

RCS mass flow rate * (Kg/s) * 3605.00 * 3605.00 *

* Reactor coolant pump speed * (rpm) * 990.00 * 995.00 f

* Steam generator pressure W (MPa) M 4.63 * 4.63 *

* Steam generator circulation rate * ( ) * 1.96 * 1.96.*

f Steam generat6r collapsed liquid level * ( cm ) * 0.0 * 0.0 *

t ft ft f *

* Steam flow rate f (Kg/s) * 266.40 * 265.90 f

* Feedwater temperature * ( 0K) * 477.00 * 477.00 *

ft **** ******** ************ ft *************** ft ft *** ** *
TABLE 5.3 Steady state results at 96.0%

* Variable * Units * JOSE CABRERA * RELAP5/MOD2*

, * * (measured) *(calculated)*

**************************************************** ********* ************* **************************

, * • * *

* Reactor power * (Mw) * 490.00 * 490.00 *

* RCS average temperature * (0K) * 565.80 * 565.70 *

* Pressurizer level * (M) * 62.00 * 61.90 *

* Pressurizer pressure * (MPa) * 13.82 * 13.82 *

* RCS mass flow rate * (Kg/s) * 3686.80 * 3686.80 *

* Reactor coolant pump speed * (rpm) * 990.00 * 995.00 "

* Steam generator pressure * (MPa) * 4.68 * 4.68 *

Steam generator circulation rate * ( - ) * 1.96 * 1.96 *

* Steam generator collapsed liquid level * (cm) * 0.0 * 0.0 *

* Steam flow rate * (Kg/s) * 255.40 * 256.00 *

* Feedwater temperature * ( 0K) * 477.00 * 477.00 *

*********************************************** * * * *
TABLE 6.1. BeauenDO 21 EveZnt

EVENT TIME (s)

calculated

Spray valve PCV-400A opening

Pressurizer heaters at full power

Turbine load decrease begins

Reactor trip

Turbine trip

Safety injection signal

Manual reactor coolant pump trip

End of simulation

0.0

87.0

135.0

435.0

435.0

462.0

1550.0

2150.0

measured

0.0

120.0

135.0

440.0

440.0

470.0

1550.0
TABLE 6.2 Variables identification JA the transient fitures

CNTRLVAR

CNTRLVAR

CNTRLVAR

CNTRLVAR

CNTRLVAR

CNTRLVAR

CNTRLVAR

MFLOWJ

MFLOWJ

CNTRLVAR

MFLOWJ

CNTRLVAR

TEMPF

TEMPF

TEMPF

TEMPG

TEMPF

SATTEMP

HTTEMP

HTTEMP

HTTEMP

HTTEMP

HTTEMP

HTTEMP

MFLOWJ

TEMPF

TEMPG

SATTEMP

HTTEMP

TEMPF

TEMPG

SATTEMP

HTTEMP

TEMPF

TEMPG

SATTEMP

HTTEMP

TEMPF

TEMPG

SATTEMP

HTTEMP

TEMPF

TEMPG

SATTEMP

HTTEMP

TEMPF

TEMPG

SATTEMP

HTTEMP

TEMPF

TEMPG

SATTEMP

HTTEMP

TEMPF

TEMPG

SATTEMP

HTTEMP

VOIDG

VOIDG

CPUTIME

066

056

054

058

073

076

083

445000000

430010000 111

425000000

100

350030000

165010000

100010000

430010000

207010000

207010000

209100110

209100210

209100310

209100410

209100510

209100610

105010000

310010000

310010000

310010000

310100105

310020000

310020000

310020000

310200105

310030000

310030000

310030000

310200205

310040000

310040000

310040000

310200305

310050000

310050000

310050000

310200405

310060000

310060000

310060000

310200505

312010000

312010000

312010000

310200605

312020000

312020000

312020000

312100105

310050000

310060000

0

- PRZ PRESSURE (KG/CM2 REL) - RCS AVERAGE TEMPERATURE (C)

- COLD LEG TEMPERATURE (C)

- RCS DELTA TEMPERATURE (C)

- PRZ LEVEL (% OF SPAN) - SG PRESSURE (KG/CM2 REL)

- SG DOWNCOMER LEVEL (CM)

- FEEDWATER MASS FLOW RATE (KG/S)

- STEAM MASS FLOW RATE (KG/S)

- CORE POWER (W)

- SPRAY MASS FLOW RATE (KG/S)

- PRZ HEATERS POWER (W)

- SPRAY TEMPERATURE (K)

- COLD LEG TEMPERATURE (K)

- HOT LEG TEMPERATURE (K)

- SG DOME TEMPERATURE (K)

- CORE EXIT TEMPERATURE (K)

- CORE EXIT SATURATION TEMPERATURE (K) - AVERAGE CHANNEL CLAD TEMP. NODE 1 (K) - AVERAGE CHANNEL CLAD TEMP. NODE 2 (K) - AVERAGE CHANNEL CLAD TEMP. NODE 3 (K) - AVERAGE CHANNEL CLAD TEMP. NODE 4 (K) - AVERAGE CHANNEL CLAD TEMP. NODE 5 (K) - AVERAGE CHANNEL CLAD TEMP. NODE 6 (K) - RCS MASS FLOW RATE (KG/S)

- PRZ LIQUID TEMPERATURE NODE 1 (K) - PRZ STEAM TEMPERATURE NODE 1 (K)

- PRZ SATURATION TEMPERATURE NODE 1 (K) - PRZ WALL TEMPERATURE NODE 1 (K)

- PRZ LIQUID TEMPERATURE NODE 2 (K) - PRZ STEAM TEMPERATURE NODE 2 (K)

- PRZ SATURATION TEMPERATURE NODE 2 (K) - PRZ WALL TEMPERATURE NODE 2 (K)

- PRZ LIQUID TEMPERATURE NODE 3 (K) - PRZ STEAM TEMPERATURE NODE 3 (K)

- PRZ SATURATION TEMPERATURE NODE 3 (K) - PRZ WALL TEMPERATURE NODE 3 (K)

- PRZ LIQUID TEMPERATURE NODE 4 (K) - PRZ STEAM TEMPERATURE NODE 4 (K) - PRZ SATURATION TEMPERATURE NODE 4 (K) - PRZ WALL TEMPERATURE NODE 4 (K)

- PRZ LIQUID TEMPERATURE NODE 5 (K) - PRZ STEAM TEMPERATURE NODE 5 (K)

- PRZ SATURATION TEMPERATURE NODE 5 (K) - PRZ WALL TEMPERATURE NODE 5 (K)

- PRZ LIQUID TEMPERATURE NODE 6 (K) - PRZ STEAM TEMPERATURE NODE 6 (K) - PRZ SATURATION TEMPERATURE NODE 6 (K) - PRZ WALL TEMPERATURE NODE 6 (K)

- PRZ LIQUID TEMPERATURE NODE 7 (K) - PRZ STEAM TEMPERATURE NODE 7 (K)

- PRZ SATURATION TEMPERATURE NODE 7 (K) - PRZ WALL TEMPERATURE NODE 7 (K)

- PRZ LIQUID TEMPERATURE NODE 8 (K) - PRZ STEAM TEMPERATURE NODE 8 (K)

- PRZ SATURATION TEMPERATURE NODE 8 (K) - PRZ WALL TEMPERATURE NODE 8 (K)

- PRZ VOID FRACTION NODE 5

- PRZ VOID FRACTION NODE 6

- CPU TIME (S)
TABLE 7.1 - nt

Real time RT = 2150.0 seconds

CPU time CPU - 35172.0 seconds

Total number of volumes C - 73

Total number of time steps DT - 25961

(CPU x 1006)/(c x DT) - 18.5
U.- w ..

•

06

F'.

I I.

~ it.

~ ji; f~

1* Ii , ..* **'

IlI I

RFW I1

"r3

(inm

C-)

:I>mzC

C-) mm

"V

I-z

-4mmmz

-4. z

:2

Eccs
*b IL'III H! Th{-81 ! .

FIG. 3.1. PRESSURIZER PRESSURE (KG/CM2 REL, NARROW RANGE)

10 30 40.,

161

FIG 3.2. PRESSURIZER PRESSURE (KG/CM2 REL, WIDE RANGE)

RCS DELTA TEMPERATURE (C)
FIG. 3.3. RCS PROGRAMMED AVERAGE TEMPERATURE (C)

RCS AVERAGE TEMPERATURE (C)

15 65 115 165 215 26 315

A? - 7.

FIG 3.4. COLD LEG TEMPERATURE (C)
20i 40 * 60 .80 Tact

• 8

FIG. 3.5. PRESSURIZER PROGRAMMED LEVEL (1 OF SPAN)

PRESSURIZER LEVEL (% OF SPAN)

010 it 20

N*174566-012-0T011 _ _ _;

FIG. 3.6. RCS CHARGING FLOW (LPM)
i.

at

...".... " ........ 111 60 90 ;60- 30 3 60 J

• -Z.0,3 -10 

.I I . " : : . •- , . ..~

FIG. 3.7. STEAM GENERATOR PRESSURE (KG/CM2 REL)

STEAM GENERATOR DOWNCOMER LEVEL (CM)

I 

I 400 600 800 1000 1200

43 4a v /7

FIG. 3.8. FEEDWATER MASS FLOW RATE (TN/H)

STEAM MASS FLOW RATE (TN/H)
FIG. 3.9. TURBINE LOAD (MWe)
FIG. 4.1 JOSE CABRERA PLANT NODALIZATION
RELAP5fl¶0020 -I. Il~a1oA/cNJco

MRELAPS/M0020 + EOIDA/CNJCO

FIG. 4.2. RCP COASTDOWN. PERCENTAGE OF THE NOMINAL FLOW.

SHORT-TERM.

FIG. 4.3. RCP'COASTDOWN. PERCENTAGE OF THE NOMINAL FLOW.

LONG-TERM.
351

0

UFIG. 4.4. AD-HOC REDUCED NODALIZATION
E1 CNTRLVAR 66 A PLANT DATA

ID

NZ 0

N

(n

Lco

A

A

A

A

A A

A

I I I I I I I I I I I I I I I I I I i I I I I I

0 t0 20 30 40 50 60 70 e0 90 300 310 120 130 140 150 160 170 180 190 200 210 220 230

TIME (S) *IO'

UNION-FENOSA ICAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 6.1 )
ID CNTRLVAR 56 A PLANT DATA

Go

0N

n"

0)

N

03

m

Cq

:0) Z) c

I--

< •: to <'-a

ILU cc_,O

ILU

LUI

(.20)

<NW

a:

N

< t

N

N

(u)

co

a:O

Nq

A

A

A

A

A A A

A A

A A

A

0 10 20 30 40 50 60 70 8o 90 100 110 120 130 140 150 160 170 180 190 200 210 220 230

TIME (S) *IO'

UNION-FENOSA ICAP INADVERTED PRZ SPRAY TnTAI. nPFNTNIN ( Ff T - . 9
[P CNTRLVAR 54 A PLANT DATA

,N

0

0)

N

N

0)

N

N

UJ

toz

<N

CLit ZN

I-J

N~c

(2dA

0

0 E

0)

to

N

0)

to

N

N

to

N

to

to

N

100 110 120 130 140 150 160 170 180 190 200 210 220 230

TIME (S) w10'

UNJON-FENOSA ICAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG R
M CNTRLVAR 58 A PLANT DATA

w

(.4

(J c 0-

w 0

u

A

A A

A A

A

A

A

A

A

A

a

0 0 20 30 40 o0 60 70 80 90 100 110 120 130 140 150 IGO 170

TIME (S) w10'

180 190 200 210 220 230

UNION-FENOSA ]CAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 6,4)
11] CNTRLVAR 73 A PLANT DATA

to

N

zto

In

(I)

LL t

LL n

0

WV to

m

N.

0

N.

A

s" A

V,

I I I I I I I I I I I I I I I

a 10 20 ..30 40 50 60 70 80 so 100 110 120 1330 140 350 160 170 180 190 200 210 220 230

TIME (S) *1IO'

UNION-FENOSA ICAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 6.5)
El CNTRLVAR 76 A PLANT DATA

to

w

in

U '

LDJ

IZ

'n

in

-D

(n in

- I - I I I I I I I I I I I I I

0 10. 20 .30 4o 50 6O 70 80 90 100 110 120 130 14o 150 160 170 190 190 200 230 220 230

TIME (S) 410'

UNION-FENOSA ]CAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 6.6)
ID CNTRLVAR 83 A PLANT DATA

CD

-3:

0

Li (n

A

A *

A

A A

A

A

0 io 20 so 4o 50 6o 70 go 90 1OO 110 120 130 14O 150 160 170 190 190 200 210 220 230

TIME (S) *10'

UNION-FENOSA ICAP INAOVERTEO PRZ SPRAY TOTAL OPENING (FTGC 9.7
mD MFLOWJ 445000000 A PLANT DATA

0

0

03

co

('4

0

w

CD

co

LU

LUL-3

D--

CD1

LU

kL 0

0

A A

A '-El--rn 0

T I I I I I | I I I I I I I I I

0 30 20 30 40 90 60 70 80 90 100 110 120 130 140 150 160 170 18o 190 200 210 220 230

TIME (S) *lO'

iNIfON-FFNOqA IrAP INAnVFRTFD PR7 SPRAY TfTAI. nPFNTNf- Fir- 6.R i
ID MFLOWJ 430010000 A PLANT DATA

0)

0

CI

N

0

toJ

N

-0

*0

0

Lii

U-)

0~

11,

.... ...~- .......a AIj, I

0 10 20

~I ~ ~ I ' I I I I "• I I *-T 4' ~ I I I I I I l- I *• I -- I

30 40 50 60 70 s0 90 100 I10 120 130 140 150 160 170 190 190 200 210 220 230

TJME (S) *lO'

UNION-FENOSA ]CAP INADVERTED PRZ SPRAY TOTAL OPENING (Fig fq)
[1 CNTRLVAR Ill.

V I I 1 . . . 220 230 a0 1 20 30 40 50 60 70 80 90 100 110 120 1.30 140 150 160 170 180 190 200 210

TIME (S) *10'

UNION-FENOSA ICAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG bIU
FD MFLOWJ 425000000

0

N' M-O Eg 9 E3 E3 E9 E9 m m m

V

N

w

U)-

CLO

I:

N

0

C0

I - - I I~T I~I - I I I~1 U I I j La L~j ~ La

0

I I I I I I I I I I I I I I I I " , I 'q I ý I ,'

10 20 30 40 50 60 70 80 90 100 130 320 330 340 150 160 170 180 190 200 210 220 230

TIME (S) W10'

UNION-FENOSA ICAP INADVERTED PRZ SPRAY TOTAL OPENING (FJG6.11)
mD CNTRLVAR 10lO0-

0oo 110 120 130 140 150 160 170 260 190 200 210 220 230

TIME (S) w10'

UNION-FENOSA ]CAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 6.12)
P] TEMPF 165010000

A TEMPG 430010000

+ TEMPF 100010000

co

00

in

.0

in

tn

(tn

cc

oOin

CL)

r w

tn

(0

in

w

In

0 10 20 30 40 50 60 70 go 90 100 110 120 330 140 )50 160 170 190 190 200 210 220 230

TIME (S) Y4O'

GNION-FENOSA JCAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 6,13)
21 TEMPF 207010000 + SATTEIP 207010000

!2

sa

in

L0 i

0

MWin

in

0,

oin

C-)

r-4

ID n

in

In

i'n

on t-4

In

0 10 20 30 40 50 60 70 80 90 100 110 120 130 140 150 160 170 190 190 200 210 220 230

TIME (S) *10'

UNION-FENOSA ]CAP INADVERTEO PRZ SPRAY TOTAL OPENING (FIG 6.1L)
2D HTTEMP 209100110

A HTTEMP 209100310

X HTTEMP 209100510

* HTTEMP 209100210

* HTTEMP 209100410

7 HTTEMP 209100610

tO 20 30 40 50 so 70 90 90 100 310 J20 130 140 150 160 170 180 190 200 210 220 230

TIME (S) ,I0l

UNJON-FENOSA ICAP ]NADVERTED PRZ SPRAY TOTAl. OPFNINr; (Flr. 6.1-'
F3- tIFLOWJ 105010000

n

03

wn

03Wn in

%0

Ul)

0,

0

C3

C3o

w

Ld N

a)

<0

X23

Ippppp9 F9

U)oC

L) -4-

0)

C3 .4- I I I I I I I I I I I I I

n 20 N0 30 4-0 so 60 70 80 s0 100 110 120 130 40 150 GO0 170 180 190 200 210 220 230

TIME (3) w10'

UNION-FEN OSA ICAP JNADVERTEIJ PRZ SPRAY TOTAL OFi.N]N@ (Fl(-,6.1-9)
WD TEMPF 310010000

A SATTEMP 310010000

+ TEMPG 310010000

* HTTEMP 310100105

co

0

0

03

to

Lo

N .-- t m

0

z~

Ln

C m

LU

Lii

InN 0

tD

wa

tntn

N

toin

co

an

!o :0 20 30 40 50 6o 70 80, 90 100 110 120 1.1o 140 150 160 170 190 190 200 210 220 230.

TIME (S) W10'

UNION-FENOSA JCAP INADVERTED PRZ SPRAY TOTAL OPENING (FIC 6.17)
ED TEMPF 310020000

A SATTEMP 310020000

* TEMPG 310020000

* HTTEMP 310200105

to

to

00

to

Lii

to

U) %n

Lin

co

w

co

to

in

in

to

0 8o 90 100 110 120 130 140 190 16O 170 180 1SO 200 210 220 230

TIME (S) w10'

UNION-FENOSA ICAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG6.18)
mD TEMPF 310030000

A SATTEMP 310030000

* TEMPG 310030000

* HTTEMP 310200205

100 110 J20 130 140 150 IGO 170 180 190 200 210 220 230

TIME (S) *lO'

UNION-FENOSA ]CAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 6.lq)
E2 TEMPF 310040000

A SATTEMP 310040000

+ TEMPG 310040000

* HTTEMP 310200305

03

to

to

0

to

Lo

LU 0

Z: o

a)

LU

M

LU

LU c

N

CL

IN

C,

In

0

co

in

0 10 20 30 40 50 60 70 80 . 90 100 110 120 130 140 150 160 170 180 190 200 210 220 230

TIME (S) w10' I

UNJON-FENOSA ICAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 6.20)
•1D TEMPF 310050000

A SATTEMP 310050000

+ TEMPG 310050000

*÷ HTTEMP 310200405

Lo

nC3

Ni

Lo

0

ZC3

to c 0)

a) Z) in

F--

00

d0~

N

(L

tn

tn

0 10 20 30 40. 50 60 70 81 90 100 110 120 1.30 140 150 160 170 1Or 190 200 210 220 230

TIME (S) x-1O'

UN]ON-FENOSA ]CAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG,6,21)
ff TEMPF 310060000

A SATTEMP 310060000

X TEMPF 350030000

+ TEMPG 310060000

* HTTEMP 310200505

U,

0

(L

to

C)

Zin

mu,

o. Zen

Lo in

0i

tnI/I

%n vI

,n

0

1 00 20 30 40 50 60 70 . 80 90 100 110 120 ISO 140 150 160 170 180 190 200 210 220 230

TIME (S) w10'

UN]ON-FENOSA ICAP INADVERTED PRZ SPRAY TOTAL OPENING (FIGC6.9)
ID TEMPF 312010000

A SATTEMP 312010000

+ TEMPG 312010000

11 HTTEMP 310200605

0

co

a

to

to

0

Lo

to

li0

0

Lo

i 0, C:) U' to

DU')t

LIcr n

LLI

ccn

F- 0

-a

01)

N'

03

a)

.n

100 110 120 130 140 15o 160 170 190 .90 200 210 220 230

TIME (S) w10'

UNION-FENOSA ]CAP INADVERTED PRZ SPRAY TOTAL OPENING (FiG6.23
ID TEMPF 312020000

A SATTEMP 312020000

* TEMPG 312020000

* HTTEMP 312100105

r) !0 20 30 40 50 Go 70 so 90 100 110 120 130 140 150 160 170 180 190 200 210 220 230 TIME (S) w1O'

UNION-FENOSA ICAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 6.24)
ID VOIDG 310050000

0

LO06

Ln~

(o

00

Zd

0

ZO_

0 L

<1-

a: 0

LL 0

. N•

C..

0

0

0;

a)

0i

+ VOJOG 3100GO000

.4-

:1; .: I. -

| | | I | I I I "t" "1" I'-- I I I - I I

0 10 20 30 40 50 60 70 8r) 90 100 ig0 120 130 140 150 360 170 380 390 200 210 220 230

TIME (S) .KIO'

UNION.-FENOSA ]CAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 6.25)
ID CPUTIMEO

C4I

Lo

0

0)

0 C

,-I

C0

Wa

Ld c

y-2

C4

C3

10 20 N 30 40 50 60 70 s0 90 100 10E 120 130 140 150 160 170 1 F0 1.90 200 210 220 230

TIMEI (S) *KI01

UNION-FENOSA ]CAP INADVERTED PRZ SPRAY TOTAL OPENING (FIG 7.1)
A~DDENDUM A

ASSESSMENT OF RE~L.PLkIQP AGANST PREBOURIZB SEPARATE

EFFETS XPEIMENTS!

ASSESSMENT OF RELAPS/MOD2 aGAINST PRESSURIZER SEPARATED EFFECTS

EXPERIMENTS

To evaluate the RELAP5/MOD2 response to pressurizer filling

transients and the interphases impact, two pressurizer separated

effects experiments have been selected. These experiments were

conducted at the NEPTUNUS (Holland) and MIT (USA) experimental

facilities.

The local pressurizer performance is emphasized in these

concerns because it presents the wall/steam, liquid-steam and

spray/steam interphases and, consequently, it generates the global

pressure value to be transmitted throughout the system.

Before discussing the mentioned separated effects

experiments, the most significant pressurizer parameters have been

studied.. These parameters are:

a) Geometry:

Height, diameter, spray valve position, wall

thickness and material properties.

It imay be pointed out that the volumes of these

pressurizers are 0.032 m 3 (MIT) and 1.26 m3

(NEPTUNUS) to confront with the 9.0 M 3 in the

pressurizer of JOSE CABRERA NPP.

- A.I -
b) Operating conditions:

Pressure and level in steady state operation

(saturation equilibrium) prior to each experiment.

c) Interphases:

Liquid-steam, liquid-wall, steam-wall and

steam-spray.

d) Boundary conditions:

Heat losses, heaters power, continuous and controlled

spray flow, surge line flow, relief and safety

valves.

After an evaluation based on the definition of the

"importance" functions for the interphases, it may be concluded

that, due to the scale of the experimental pressurizers with

respect to the one of JOSE CABRERA NPP, the "importance" of the

interphases is amplified by a factor of between 1.65 and 3.48 for

the NEPTUNUS pressurizer, and between 90.0 and 398.0 for the MIT

pressurizer.

Simulation of these pressurizer tests is considered

appropriate since, if with such an amplification the code provides

an acceptable response, the scalability to JOSE CABRERA NPP will be

guaranteed and the uncertainty minimized.

- A.2 -
I.- MIT-A EXPERIMENT SIMULATION.

The pressurizer is a small one of 0.032 m3 (fig A.1) set

up in the laboratories of the Massachusetts Institute of

Technology (MIT), USA, operating much below the reference

reactor nominal values. The experiment started from a thermal

saturated equilibrium condition at a pressure of 6.1+5 Pa and

a fixed level of 0.35 m.

The experiment consisted of a partial filling with cold

water (43 *C) up to a final level of 0.86 m. with spray and

heaters out of operation. Pressure increased due to

compression of the steam bubble that became overheated, thus

producing heat transfer to the subcooled liquid interphase,

wall condensation and heat losses to the environment. coolant

stratification was also observed during the test.

A system model for the RELAP5/MOD2 code has been

generated based on the available drawings, processing the

geometry as well as metalic structures, environmental losses,

and surge line flow.

The following conclusions of the RELAP5/MOD2 post-test

simulation may be emphasized:

- - Pressure simulation is overestimated in case the wall

effect and environmental losses are not simulated. A

code malfunction with sudden pressure decrease whenever

the level goes just from one node to the next one has

been detected (fig. A.2).

- A.3 -
- Steam temperature simulation is very deficient in case

the instrumentation time delay constant is not

simulated (fig. A.3).

- The response is very sensitive to the conductivity and

specific heat of the steel wall.

- The evolution of the liquid, steam and wall

temperatures are correct, both at the lower and upper

part of the pressurizer vessel (fig. A.4 y A.5).

- Final pressure simulation has been satisfactorily

calculated (fig. A.6).

- It becomes necessary to provide the model with a

sufficient number of nodes in the liquid phase, to

allow for a correct simulation of the coolant thermal

stratification and axial gradient.

After the appropriate simulation of the MIT-A experiment,

it may be concluded that RELAP5/MOD2 correctly predicts the

liquid thermal stratification, the steam compression due to

the filling of the pressurizer, the condensation in the wall

and in the steam-liquid interphase, and the environmental

losses, thus generating satisfactory evolutions -of level,

temperatures and pressure in pressurizer slow filling

transients.

- A.4 -
FIG. A.l: MIT pressurizer test.

model.

Facility diagram and RELAP5

*0/P tllI 114.3 cO,

class

E)

0.0 Cm -

To N1onconde sable

'Gas Tank

.Line Stea U

.To Drain 0

FIG. A.2: MIT pressurizer test. Effect of wall simulation on the

RELAP5 calculated system pressure.

TIME (S

smn P~irssmuIrp it *A
FIG. A.3: MIT. Effect of the steam temperature instrumentation

lag on the RELAP5 calculated steam temperature.

NIT In•S$MIZElt TEST 'A'

FIG. A.4: MIT. RELAP5 simulation of the wall and liquid

temperature in the lower part of the pressurizer.

O T-4*ELAPI 0 4 IV 4'IEIcP3 A (99r M000 *1r1 HIP S10010

X N11IlEM 310010*7. 6

,ITI 3090000

a

I

I

U

I

WI,

~I.

4

U

a-a.

U

'-I

~l.

N

I

TIE tSi

m1!l PRE.SSuRIlR ISI 'A'
FIG. A.5: MIT. RELAP5 simulation of the wall and liquid

temperature in the upper part of the pressurizer.

0-1.3d6CXPI 0

A LAG-1I9 3WS 1

X 141ICMP 31INOO

+ 'V.3diCEXP3

+ 0411(EW SIOOSI

Z SATICP SOSCOMo

0

I

"I t 'I ' a so 'a N 'U a 'a 'a W ;0 t S 'a 'a 

ITIMF (S)

q41'P'SSU4k!ICR TET! 'A'

FIG. A.6: MIT. RELAP5 calculated system pressure.

U 4 P 0 PI|XPI a * P 3090000

U

I

I

Sit 04 n a a W a a - w a v U " ""% t a a U TIME (S)

MIT1 PRESSURtIZER9 7LS 'A'

2.- NEPTUNUS Y-05 EXPERIMENT SIMULATION.

The pressurizer has 1.26 m 3 of volume (fig. A.7), is

located at the Delft University (Holland) and operates at

conditions similar to those in an actual power reactor. The

Y-05 experiment started from a thermal saturated equilibrium

condition at a pressure of 126.0+5 Pa, an electrical heating

power of 17. Kw and a fixed level of 1.12 m.

The experiment consisted of successive filling and

emptying of the pressurizer as a result of cyclic oscillations

(fig. A.8) of the lower surge line make up flow rate at

constant temperature and of the upper spray flow rate at

variable temperature with the heating power kept constant.

Level oscillations with regard to the reference level

were observed, as well as pressure oscillations as a function

of the combined effect of the following trends:

- pressure increase due to steam volume decrease

associated to rising level.

- pressure increase due to heaters effect.

- pressure decrease due to steam volume increase

associated to dropping level.

- pressure decrease due to steam condensation in the wall

and in the water and spray interphases.

- A.5 -
A system model for the "RELAP5/MOD2 code has been

generated based on the available drawings, processing the

geometry as well as heaters, metalic structures, environmental

losses, and surge line and spray flow.

The following conclusions of the RELAP5/MOD2 post-test

simulation may be emphasized:

- Pressure and steam temperature increase rates are

predicted conservatively in excess during filling, with

amplification of pressure peaks (fig. A.9).

- Pressure decrease rate due to emptying is predicted in

excess, with satisfactory fitting of pressure valley

points (fig. A.9).

- A filling impact larger than the spray impact is

conservatively predicted, in opposition to what was

experimentally observed (fig. A.9).

- Presence of overheated steam in filling is correctly

predicted as well as saturated steam in emptying.

Experimentally observed level oscillations are

predicted as well (fig. A.10).

- The difference between calculated and measured pressure

peaks increases along the various transient cycles

(fig. A.9) due to the increasing delay between surge

line and spray flows (fig A.8).

- The heat transfer in the interphase with the subcooled

- A.6 -
spray is underestimated.* Predicted depressurization

rate due to spray condensation is lower than measured.

- It is convenient to fit the nodalization so that the

pressurizer average level is in a node center.

- Performance of RELAPS/MOD2 matches TRAC-PFl/MODl as

illustrated in NUREG/CR-3919.

After the NEPTUNUS Y-05 experiment simulation it may be

concluded that RELAP5/MOD2 satisfactorily predicts, with

conservative trend, the liquid thermal stratification, steam

compression due to pressurizer filling, wall, steam-liquid and

steam-spray interphase condensation, and environmental losses,

thus generating a satisfactory response to level, temperature

and pressure evolutions in pressurizer fast transients.

-A.7 -

FIG. A.7: NEPTUNUS Y-05 test. Facility diagram and RELAP5 model.

wal

.06

II

FIG. A.8: NEPTUNUS Y-05 test. RELAP5 simulation of the surge

line and spray mass flow rates.

0 L'OVJ I5OOnOOO + MFLOWJ 25000000

.NEPTUNUS PRESSLAIZEA IESY Y.05
FIG. A.9: NEPTUNUS Y-05 test. RELAP5 calculated system pressure.

a rouiof-.0%) 12050=o 4P.IpriP 0

* U- a a n to so em m In so so on In log e "an m tosoM

71;k( (S;

14CPILWA PKSSULMIZER IEST Y-OC%

FIG. A.10: NEPTUNUS

level.

Y-05 test. RELAP5 calculated pressurizer

U1 s.I qVAR I

2

-j

Wit

W2

N(P7UWuS PRESSURIZER 7CS7 T-05
3.- ASSESSMENT CONCLUSIONS OF RELAPS/MOD2 WITH REGARD TO

PRESSURIZER SEPARTED EFFECTS TRANSIENTS.

Both studies of pressurizer separated effects, with

regard to the performance of RELAP5/MOD2 and its applications

to pressurizer pressure slow and fast transients in commercial

reactors, lead to the following conclusions:

- It is necessary a correct simulation of the thermal

properties of the wall material as well as the

environmental thermal losses.

- Simulation of the instrumentation time delay constant

is fundamental in order to compare steam temperature

calculated and measured values.

- Time steps of 0.05 seg. are adequate.

- Prediction of both liquid/wall and steam/wall heat

.transfer is correct.

- Prediction of axial thermal gradient in the wall as

well as in the water is correct.

- Condensation in the steam/spray interphase is

underestimated.

- Filling pressurization and emptying depressurization

rates are overestimated as a result of an

underestimation in processing condensation and

vaporization rates in the liquid-steam interphase.

- A.8 -
- RELAP5/MOD2 performance is iimilar to the observed in

TRAC-PFl/MODl simulation.

- Because

function

NEPTUNUS

in these

the real

the pressurizer interphases "importance"

is lower in JOSE CABRERA NPP than' in MIT and

experimental facilities, the observed impact

simulations will be reduced when applied to

plant case.

code

well

both

real

This set of conclusions positively qualifies RELAP5/MOD2

to be applied for pressurizer filling pressurization as

as for emptying and spray condensation depressurization

for slow and fast transients simulation in the case of a

nuclear power station.

- A.9 -
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION I. REPORT NUMBER (249) (Aalgned by NRC. Add Vol.. Bump. Rev..

NRCM 1102. Wv Addendum Numborw. If ISY.) 3201.3202 BIBLIOGRAPHIC DATA SHEET

(s-g i,,r.,etions, on, rho ,iv. verREsa)IA-01

2. TITLE AND SUBTITLE NUREG/IA-0124

ICSP-JC-SPR-R

Assessment of RELAP5/MOD2 Against a Pressurizer Spray Valve

Inadverted Fully Opening Transient and Recovery by Natural Circulation 3. DATE REPORT PUBLISHED

in Jose Cabrera Nuclear Station JuNTH e 1 99ýJ^

4. FIN OR GRANT NUMBER

_,L2245

5. AUTHORISI 6. TYPE OF REPORT

R. Arroyo, L Rebollo Technical

7. PERIOD COVERED lneduhe De*teut

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (ii NRC.rovide Division. Ofice oe ReiWon. U.S. Nu kReRlatory Co==. end meakvm esO ll contructo, io.ovide

Union Electrica Fenosa

c/Capitan Haya 53

28020 Madrid

Spain

9. SPO NSO R ING ORGA N IZATI ON -- NAM E A ND ADD0R ESS fit NA C. t.p* ý so &bovw" 0 co,,tractor. prvide NR•C Diviio,,. Office of Region. U.S. Nk•,ker Regtatwy Comhmkn.ot .,d mallfdd'I/

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

10. SUPPLEMENTARY NOTES

1I. ABSTRACT 1200 -F or 1m)

This document presents the comparison between the simulation results and the plant measurements of a real event

that took place in Jose Cibrera nuclear power plant in August 30, 1984. The event was originated by the local,

continuous and inadverted opening of the pressurizer spray valve PCV-400A.

Jose Cabrera power plant is a single loop. Westinghouse PWR belonging to UNION ELECITRICA FENOSA, S.A.

(UNION FENOSA), a Spanish utility which participates in the International Code Assessment and Applications

Program (ICAP) as a member of UNIDAD ELECTRICA, S.A. (UNESA). This is the second of its two

contributions to the Program: The first one was an application case and this is an assessment one. The simulation

has been performed using the RELAP5/MOD2 cycle 36.04 code, rinning on a CDC CYBER 180/830 computer

under NOS 2.5. operating system.

The main phenomena have been calculated correctly and some conclusions about the 3D characteristics of the

condensation due to the spray and its simulation with a ID tool have been got.

12. KEY WORDS/OESCR:PTORS (List word orprows that will tist fe~hars In 0ocadne th" Wol. J 13. AVAILABILIT'Y STATEMENT

ICAP, Spray, Jose RELAP5 Cabrera ,Unlimited 14. SECURITY CLASSIFICATION

Spra, REAP5(TMi PAPe)

Unclassified

I'Thi ReponI

Unclassified

15. NUMBER OF PAGES

16. PRICE

NRC FORM 335 12-191


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