NUREG/IA-00 18

Inteionational

Agreement Report

RELAP5/MOD2 Assessment,

OECD-LOFT Small Break

Experiment LP-SB-03

Prepared by

S. Guntay

Paul Scherrer Institute

CH-5303 Wurenlingen, Switzerland

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

April 1990

Prepared as part of

The Agreement on Research Participation and Technical Exchange

under the International Thermai-Hydraulic Code Assessment

and Application Program (ICAP)

Published by

U.S. Nuclear Regulatory Commission
M

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.

I

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-001 8

Internallonal

Agreement Report

RELAP5/MOD2 Assessment,

OECD-LOFT Small Break

Experiment LP-SB-03

Prepared by

S. Guntay

Paul Scherrer Institute

CH-5303 Wurenlingen, Switzerland

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

April 1990

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 is based on work performed under the sponsorship of the

Swiss Federal Office of Energy. The information in this report has

been provided to the USNRC under the terms of the International

Code Assessment and Application Program (ICAP) between the United

States and Switzerland (Research Participation and Technical

Exchange between the United States Nuclear Regulatory Commission

and the Swiss Federal Office of Energy in the field of reactor

safety research and development, May 1985). Switzerland has

consented to the publication of this report as a USNRC document in

order to allow the widest possible circulation among the reactor

safety community. Neither the United States Government nor

Switzerland 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.
RELAP5/MOD2 Assessment. OECD-LOFT Small Break Experiment

LP-SB-3

S. G~intay

ABSTRACT

An analysis of the experimental results and oost-test calculations

using RELAP5/MOD2 carried out for OECD-LOFT small break experiment

LP-SB-3 are presented. Experiment LP-SB-3 was conducted on

March 5, 1984 in the.Loss-of -Fluid Test (LOFT) facility located at the

Idaho National Engineering Laboratory (INEL). The experiment simulated a

small cold leg break, with concurrent loss of high pressure injection

system, and cooldown and recovery by feed and bleed of the steam

generator secondary side and accumulator injection, respectively.

The analysis was under taken as a part of a program at EIR aimed at

developing experience in using the latest generation of best estimate Loss

of Coolant Accident (LOCA) analysis computer codes, and to improve

understanding of Small Break LOCA transients and as well as a part of a

program aimed at assessing the RELAP5/MOD2 code. The latest available

version (Cycle 33 to 36.1) of the code was used. The particular test

selected for the analysis included several phenomena potentially relevant

to any PWR plant operating in Switzerland.

This report documents a short post-test analysis of the experiment

emphasizing the results of additional analysis prerformed during the course

of this task. RELAPS/MOD2 input model and results of the post-test

calculation are documented. Included in the report is the results of a

sensitivity analysis which show the predicted thermal-hydraulic response

to a different input model.
SUMMARY

This report documents the post-test calculation of OECD-LOFT

Experiment LP-SB-3 using the RELAP5/MOD2 computer code.

Experiment LP-SB-3 simulated a hypothetical loss-of-coolant accident

(LOCA) which resulted from a 4.67 cm diameter single-ended break in the

cold leg of a large commercial pressurized water reactor (PWR) with a

concurrent loss of high pressure emergency core coolant injection

capability. The experiment was initiated from conditions which were

scaled to be representative of the design operating conditions of a large

PWR; reactor power of 50 MWe, core temperature rise of 20 K, and system

pressure of 14.95 MPa. The experiment provided data on (a) core heat

transfer characteristics when core uncovery occurs during relatively slow

boil-off conditions, (b) effectiveness of steam generator feed and bleed to

depressurize and cool a highly voided system, (c) effectiveness of the

accumulator injection in establishing core cooling in a highly voided

system.

Results of Experiment LP-SB-3 showed that decay heat removal from

the 4.66 cm diameter equivalent small break LOCA was dominated by the

break flow and heat transfer to the secondary side. It was believed that

during the boil-off phase, reflux condensation took place for most of the

period. The possibility of condensed liquid draining back through the hot

leg and into the core was supported by the asymmetrical fuel cladding

temperature developments. The feed and bleed of the steam generator

secondary side was effective depressurizing the system. The accumulator

injection established core cooling.

The RELAP5/MOD2 computer code was shown to be valuable in

understanding the physical phenomena in the experiment. Although

differences in detail were observed between the celculational results and

experimental data, the code generally performed well, predicting all the.

key events in the correct sequence and with reasonable timing.

ii
TABLE OF CONTENTS

A bstract .................................................................................................................. i

Sum m ary .................................................................................................................. ii

1. Introduction ..................................................................................................... .1

2. Description and Discussion of Experiment LP-SB-3 3

2.1 Description of Experiment LP-SB-3 ......................................... 3

2.2 Discussion of the Experimental Results .................................. 4

2.2.1 Mass Depletion Phase ...................................................... 4

2.2.2 Core Boil-Off Phase ......................................................... 6

2.2.3 Discussion of Primary-Secondary Relationship

During Experiment LP-SB-3 ......................................... .7

2.2.3. I Reflux Condensation And Heat Transfer

to the Secondary Side ......................................

2.2.3.2 Liquid Drainback and its Effects on Fuel

Rod Temperatures ............................................. 9

2.2.4 Core Cooldown and Recovery Phase ............................ 9

3. RELAP5/MOD2 Computer Code Simulation of Experiment

LP-SB-3 ........................................................................................................... 1 0

3.1 RELAP5/MOD2 Description ........................................................ ....... 10

3.2 RELAP5/MOD2 Nodalization and Input Model for

Experiment LP-SB-3 .................................................................... ....... 11

4. Results of the RELAP5/MOD2 Post-Test Calculation .................... 12

4.1 Calculation of the Steady State ............................................. ....... 12

4.2 Transient Calculation ...................................................................... 14

4.2.1 Mass Depletion Phase ...................................................... 15

4.2.1.1 Primary and Secondary System Pressure

Behavior .................................... ..... 15

4.2.1.2 Break Flow .......................................................... 16

4.2.1 .3 Density ................................................................. 16

4.2.1.4 Pump Behavior ................................................... 17

4.2.1.5 System Mass Inventory ................................... 17

4.2.2 Core Boil-Off Phase ........................................................ 17

4.2.2. 1 Primary-Secondary System Pressure

Benavior ........................................................ ...... 17

4.2.2.2 Density ................................................................. 18
4.2.2.3 Break Flow ...................................................... . .... . 19

4.2.2.4 Core Dryout and Thermal Response ............. 19

4.2.3 Core Cooling and Recovery Phase .......................... ...... 22

4.2.3.1 Primary and Secondary System Pressure

Behavior ........................................................... ....... 22

4.2.3.2 'Core Thermal Behavior ............... ...... 23

4.2.3.3 Accumulator and LPIS Injection .................. 24

4.2.3.4 Accumulator and LPIS Flow ........................... 24

5. Sensitivity Calculation ............................................................................ 24

5. 1 Pressures ............................................................................................... 25

5.2 Core Uncovery ..................................................................................... 25

6. Run Statistics ............................................................................................... 26

7. Conclusions and Recommendations ....................................................... 27

8. References ..................................................................................................... 28

T ables .................................................................................................... . ........ 30

Figures ...................................................................................................... . ....... 32

Appendix A ............................................................................................. . ........ 84
TABLES

I. Chronology of the events

2. Initial conditions

FIGURES

1. Primary and secondary pressures (0-150 s)

2. Primary and secondary pressures (0-7000 s)

3. Steam generator pressure and.MSCV cycling set points

4. Fluid density in the break piping

5. Break flow

6. Fluid density in the intact loop hot leg

7. Fluid density in the intact loop cold leg

6. Fluid velocity in the intact loop hot leg

9. Pressure rise across the primary coolant pumps

10. Fluid density in the intact loop seal

11. Fuel cladding temperatures in the center bundle

12. Cladding temperatures at different radial positions

13. Primary secondary pressure differential

14. Temperature difference between the saturation temperature of the

primary system and the temperature of secondary system

15. Steam superheat in the steam generator inlet plenum

16. Response of reactor vessel liquid level probes, core dryout and quench

data from thermocouples

17. RELAP5/MOD2 nodalization for Experiment LP-SB-3

18.

19.

20.

21.

22.

23.

24.

25.

26.

27.

28.

Comparison of

Comparison of

Comparison of

Comparison of

Comparison of

Comparison of

Comparison of

Comparison of

Comparison of

Comparison of

Comparison of

systems

pressures in the intact loop hot leg

pressures in the steam generator secondary side

break flows

densities in the intact loop hot leg

densities in the intact loop cold leg

densities in the beak piping

pressure differential across pump

system mass inventories

pressures in the intact loop hot leg

pressures in the steam generator secondary side

calculated pressures in the primary and secondary

V
29. Calculated temperature difference between the primary and secondary

systems

30. Vapor generation rate in the steam generator upright

31. Calculated liquid flow rate in the steam generator inlet plenum

32. Calculated liquid and vapor flow rates in the intact loop hot leg

33. Comparison of densities in the intact loop hot leg

34. Comparison of.densities in the intact loop cold leg

35. Comparison of densities in the loop seal

36. Calculated collapsed liquid level in the reactor vessel

37. Comparison of break flow rates

38. Comparison of dryouts measured and calculated

39. Comparison of fuel cladding tempgratures at 12.7 cm elevation

40. Comparison of fuel cladding temperatures at 38 cm elevation

41. Comparison of fuel cladding temperatures at 69 cm elevation

42. Comparison of fuel cladding temperatures at 102 cm elevation

43. Comparison of fuel cladding temperatures at 114 cm elevation

44. Comparison of fuel cladding temperatures at 157 cm elevation

45. Comparison of pressures in the intact loop hot leg

46. Comparison of pressures in the secondary side

47. Comparison of ECC injection flow rates

48. Comparison of integrated ECC injection flows

49. RELAP5/MOD2 nodalization for the sensitivity analysis

50. Comparison of calculated fuel cladding temperatures at level 5

5 1. Comparison of calculated fuel cladding temperatures at level 3

52. CPU time versus transient time and required CPU seconds per real

second

A-I. LOFT configuration for cold leg intact loop small break LP-SB-3

A-2. Experiment break piping instrument configuration

A-3. LOFT core configuration and instrumentation

vi
1. INTRODUCTION

The Loss-of-Fluid Test (LOFT) experimental facility is a 50 MW test

reactor at the Idaho National Engineering Laboratory (INEL) used for the

study of Loss of Coolant Accidents, and other reactor transients. The LOFT

reactor system, in particular the primary coolant system and reactor core,

is a fully operational (scaled) representation of a commercial pressurized

water reactor (PWR). It has a nuclear powered core and a single active

(the normally intact loop), containing a U tube Steam Generator (SG) and

two reactor coolant pumps in parallel branches in the cold leg. A second

loop (the normally broken loop), containing passive hydraulic resistances

to simulate the steam generator and pumps, is also connected, primarily

for use in large break LOCA simulations.

LOFT experiment LP-SB-31 simulated a 4.66 cm (1.84 in) break in the

cold leg of a commercial PWR. Additional features simulated were the

failure of high pressure emergency core cooling injection system, and feed

and bleed of the steam generator secondary side. The experiment was

conducted on March 5, 1984 for the Organization for Economic

Corporation and Development (OECD) consortium. This report reviews

briefly the post test analysis carried out at INEL together with additional

post test analysis carried out during this study and the post test

calculations performed using the RELAP5/MOD2 2 (Cycles 33 to 36. 1). The

data deck used for the analysis was based on pre-test prediction deck used

by the INEL to produce Best Estimate Prediction Document (EPD) 3 . Several

Improvements and corrections on the pre-test deck were done (a) to

convert the deck to the RELAP5/MOD2 format, (b) to 'eliminate some minor

errors, (c) to tune various operations (such as opening or closure of

valves) according to the experimental data, (d) to apply cross cross flow

junction model.

The analysis was undertaken to assess the RELAP5/MOD2 computer

code, in displaying a wide range of thermal-hydraulic phenomena. These

includes:

I
a- Primary-secondary relationship Single and two phase forced

convection

Reflux natural circulation

b-Slow boil-off

c-Core heat transfer

d-Single and two phase break flow

e-Pump performance, degradation

f-Cooldown and recovery of the plant by steam generator feed and

bleed and accumulator injection

Section 2 of this document presents a short description of Experiment

LP-SB-3 and discusses briefly the post-test analysis of the experiment

emphasizing the new findings during the conduction of this study. The

RELAP5/MOD2 code is introduced in Section 3 along with the nodalization

used in the base case calculation. Section 4 discusses the results of

post-test calculation and compares the results of the calculation with the

data. Section 5 discusses the sensitivity study performed and compares

the base and sensitivity case results. Section 6 discusses run statistics.

Conclusions and recommendations are discussed in Section 7. A

description of the LOFT facility is presented in Appendix A. Appendix B

contains details of the input data for the base and sensitivity analyses.

2
2. DESCRIPTION AND DISCUSSION OF EXPERIMENT LP-SB-3

OECD-LOFT experiment series LP-SB-3 was designed to investigate the

system thermal-hydraulic response under (a) slow coolant boil-off leading

to core uncovery at high system pressure, (b) steam generator feed and

bleed operation, and (c) plant recovery with accumulator injection.

A short description of Experiment LP-SB-3 and discussion of post-test

analysis of the experiment beyond what was already presented in the

Quick Look Report (QLR)I are presented in the following subsections.

2.1 Description of Experiment LP-SB-3

Experiment LP-SB-3 simulated a small cold leg break loss-of-coolant

transient, with a scaled break size corresponding to a 4.66 cm (1.84 in.)

pipe diameter in a reference commercial PWR. The experiment was

initiated from conditions representative of those in a commercial PWR by

opening a valve in the normally intact loop cold leg break piping. The scram

was initiated on low pressure in the intact loop hot leg. The primary

system depressurized rapidly until fluid saturation conditions were

reached in the hot leg at about 100s, which resulted in a decrease in the

primary system depressurization rate. As a result, a reduction of the break

mass flow rate occurred with void formation in the coolant. Since the

primary system energy loss from the break due to the size of the break

was not enough to remove the decay heat energy, the heat transfer to the

secondary side resulted in a higher pressurization of the secondary side

which caused the main steam valve to cycle four times and to relieve the

energy of the system. The primary coolant pumps continued operation until

the system mass inventory was reduced to approximately 2Q00 Kg. Single

phase forced convection before the system reached saturated conditions

and two phase forced convection until the pump trip were observed.

Following the pump trip, reflux condensation occurred depending on the

primary-secondary relationship. The pumps during the two phase system

coolant condition delivered a relatively homogeneous mixture of liquid and

3
steam to the intact loop cold leg and the piping upstream of the break

nozzle. The primary and secondary system pressures became almost the

same after the primary system reached saturation coolant conditions and

remained coupled during the rest of the transient. A top to bottom slow

core uncovery was observed after the system mass inventory was reduced

to approximately 1750 Kg. Steam generator feed and bleed operation was

initiated when the peak cladding temperature reached 987 K which let to a

rapid cooldown and depressurization of the secondary side followed

closely by the primary side. A partial rewet of the core from below

followed the initiation of the feed and bleed. Primary pressure dropped

below the accumulator set point shortly after this, causing a further

pressure reduction. The accumulator injection caused core rewet from the

bottom upwards and quenched the core quickly to the coolant saturation

temperature. Low pressure emergency core cooling injection system

(LPIS) became operational after the complete core quench at a system

pressure of 1.03 MPa and provided long term cooling. The experiment was

terminated one minute after the LPIS injection.

The chronology of various events for Experiment LP-SB-3 is presented

in Table 1. For details of the experiment description, Reference I should

be referred.

2.2 Discussion of the Experimental Results

Experimental results have already been presented in the QLR of this

experiment. In order to perform the post-test analysis and compare with

the data, additional post test analysis was performed to identify primarysecondary relationship and its implications. A summary of the post-test

analysis emphasizing the new findings will be discussed in the following

subsections.

The experiment is discussed in there sections covering (a) mass

depletion, (b) core boil-off, and (c) core cooldown and recovery phases.

2.2.1 Mass Depletion Phase

This phase refers to the duration of the transient between the break

4
initiation and primary coolant pumps trip at 1600 s.

Figure I and 2 present the short (150 s) and long (7000 s) term primary

and secondary system pressures. The primary system depressurization

under the control of subcooled break flow until about 100 s. The

depressurization rate increased slightly when the pressurizer emptied at

about 70 s. The primary system depressurized to saturation conditions in

the break piping upstream of the break at about 100 s after the break

initiation and the primary-secondarypressure coupling was established.

The core reached saturation at about 200 s. Energy addition to the primary

system from the decay heat and primary pump heat was mainly removed by

the break flow and by heat transfer to the steam generator secondary side.

The heat input to the secondary side was relieved by periodic cycling of

the main steam control valve (MSCV) during the first 1030 s of the

transient. Figure 3 presents the steam generator secondary pressure with

the opening and closure set points. It appears that instrument noise caused

the valve to open or close exceeding or without actually reaching the valve

set point. The valve did not seat 100 % nor did it seat the same closure,

yielded different leakage from the steam generator during different

periods of the experiment. After the last closure of the valve, the steam

generator pressure showed a similar pressure increase to the pressure

rise rates occurred during the earlier closures. After it reached a

pressure some where between the opening and closure set points, the

pressure increase ceased and started dropping. Two different mechanism

might be responsible for this. The first is that the steam valve's fourth

closure position let more steam leakage than the previous closure

positions. And the second is that the heat generation in the primary side

was not enough to cope with the energy loss from the system, which could

also be related to the first mechanism. The primary and secondary systems

showed a continuous pressure drop after the fourth MSCY cycling until the

break was isolated.

After the saturation coolant condition was established in the break line

at about 100 s, the fluid density at the break continued to decrease due to

voidage of the fluid in the break line was increased. Figure 4 and 5 present

the break line density and mass flow rate. Increased voidage in the

primary coolant system decreased the break flow.

5
Fluid densities in the intact loop hot and cold legs are presented in

Figures 6 and 7. As a result of MSCY cycling, increased cooling in the

steam generator tube bundle affected the intact loop cold leg density and

the intact loop hot leg fluid velocity which is shown in Figure 8. The

differential pressure across the pump continuously decreased as the

primary coolant voidage increased. Oscillation in the pump differential

pressure started at around 700 s at a vapor fraction at the primary coolant

pump inlet of 40 %. The oscillations seen in Figure 9 continued until

875 s. Although pump operation caused a homogeneous density in the

intact loop cold leg, onset of fluid stratification as indicated by the

density measurements in the intact loop hot leg (Figure 6) occurred

shortly after the abrupt drop in pump differential pressure at 875 s. The

pump operation continued until 2800 Kg mass in the system remained. The

actual remaining inventory was 800 Kg more than the desired 2000 Kg as.

specified in Experiment Specification Document (ESD) 4 .

2.2.2 Core Boil-Off Phase

This section discusses the core boil-off phase which refers to the

duration of the transient between the primary coolant pump trip at 1600 s

and the steam generator feed and bleed operation.

Cessation of the pump operation caused also fluid stratification in the

intact loop cold leg as seen from the cold leg density measurements as

illustrated in Figure 7. This stratification caused the break plane to

become uncovered. Thereafter, the break flow was high quality steam.

Consequently, the primary system depressurization rate was increased.

The other affect of the pump trip was on the pump seal density and mass

inventory. Since the fluid in the suction and as well as in the discharge

piping collapsed with the pump caostdown, the liquid collected in the loop

seal caused immediate density increase in the loop seal as presented in

Figure 10. The loop seal remained covered afterwards and prevented

natural loop circulation. The secondary system pressure followed the

primary response, and both pressures decreased at 0.0 kPa/s rate.

6
A slow boil-off process was taking place in the reactor vessel governed by

(a) decay heat energy addition to the fluid, (b) break mess flow, and (c)

pnmary-secondary heat transfer rate. The upper portion of the core

showed dryout at about 3800 s, when the primary system mass inventory

fell to approximately 1750 Kg. The dryout progressed towards bottom of

the core. Based on the dryout times indicated by the cladding

thermocouples, the rate of core uncovery rate was about 1.8 mmls when

the dryout level was within the upper 1/3 of the core, and was reduced to

1. 1 mm/s when the dryout level was between 2/3 and 1/3 of the core

height. The boil-off rate decreased as the core become uncovered because

less energy was transferred into the fluid from the lower portion of the

core below the peak-power region which is 0.68 m above the core inlet.

The cladding temperature measurements below 0.13 m elevation indicated

that the core was not completely uncovered. Figure II presents the fuel

cladding temperatures in the center bundle. Figures 12 presents the fuel

cladding radial temperature response at the 1. 14 m elevation. Differences

in the heat up rates at (a) different axial elevations and (b) at different

radial locations were observed. The QLR suggested that a small amount of

entrained water carried up into the core slowed the heat up rate at higher

elevations. The difference in the heat up rate in the radial direction was

explained with the difference in the radial power factors. Further analysis

of the primary-secondary relationship indicated that the indicated

differences were more likely as a result of liquid drainDack from the hot

leg, generated due to reflux condensation in the steam generator tubes and

its re-evaporation in the core. Detailed discussion of the primary and the

secondary relationship is given in the following subsections.

2.2.3 Discussion of Primar -Secondary Relationship During

Experiment LP-SB-3

Post-test examination of the measured pressures and temperatures in the

primary and secondaru sides and the conditions existing in the inlet and

outlet plena providea that reflux condensation was occurring when the

primary side pressure and/or temperature was higher than the secondary

pressure and/or temperature during the boil-off phase. This conclusion

was further supported by the RELAP5/MOD2 post-test calculations and

will be presented in Section 3. The possibility of drainback of the

7
condensed liquid through the hot leg into the core and interpretation of

various data providing evidence to this conclusion are discussed in the

following two subsections.

2.2.3.1 Reflux Condensation and Heat Transfer to the Secondary Side

Pressure differential between the primary and secondary systems is

presented in Figure 13. As the core uncovery progressed, due to lesser

energy transfer to the fluid, and at the same time due to continuous energy

loss from the break, the primary side pressure dropped below the

secondary side pressure at about 4300 s. The primary side pressure

remained below the secondary level until about 4900 s, slightly after the

break isolation at 4750 s. During this time period the secondary

temperature was higher, as illustrated in Figure 14, than the saturation

temperature of the hot leg. Outside of this period it was believed that

condensation was taking place in the steam aenerator tubes below the

secondary liquid level, especially close to the tube sheet and forcing the

primary pressure to follow that of the secondary pressure. The

condensation heat transfer in the steam generator tubes was the

additional heat removal mechanism in the primary side besides the heat

removal via the break flow. After the break was isolated, it became the

only heat removal mechanism in the primary system other than the heat

loss to the environment. The sharp change in the slope of the pressure

differential suggests that the condensation heat transfer was

reestablished just after the break isolation, presumably increased in

magnitude to compensate the loss of the primary side energy outflow via

the break flow. Therefore, the condensation heat transfer became an

essential factor in the overall heat balance existing in the primary

system.

The heat transfer from from the primary to secondary side was estimated

to be 50 to 150 kW in the QLR with the lower bound attained which the

pressure differential was negative. On an approximate basis, the actual

heat transfer can be excepted between zero to 100 kW with a 50 kW offset

to eliminate the heat from the metal work ( which was not taken into

consideration in the QLR) and the instrument (which indirectly measures

8
the steam generator level) uncertainties. A heat transfer of probable 50

kW is equivalent to a condensation rate of 30 g/s 5 in the primary tubes.

2.2.3.2 iquid Drainback and its Effect on Fuel Rod Temperature Some of

the condensate that formed in the upside might have probably drained back

into the vessel. The indirect evidences of this drainback outside of the

4300-4900 s period were (a) the thermocouples in the steam generator

plena showed that saturation conditions persisted in those regions during

the entire period except from 4200 s to 5000 s, when superheat was

measured (the steam superheat shown in Figure 15 was due primarily to

hot wall effects), (b) the upper end box thermocouples which showed

typically 200 K lower temperature in fuel module 4 (which is the nearest

to the intact loop hot leg) than in equivalent ones in modules 2 and 6, (c)

cladding temperatures in modules 4 were asymmetrically lower than in

modules 2 and 6; by 60 K at 1.14 m, by 50 K at 0.76 m, and approximately

equal at 0.28 m above the core inlet, and (d) the fuel heat up rate, as seen

in Figure 12, considerably decreased after 4900 s with the

reestablishment of the reflux condensation, and (e) corner module liquid

level probes, as illustrated in Figure 16, showed that the probe at 1.77 m

elevation above the core inlet was wet outside of the negative pressure

differential period (4300 to 4900 s) although the dryout level was well in

progress in the core.

2.2.4 Core Cooldown and Recovery Phase

This phase refers to the duration of the transient starting with the

initiation of the feed and bleed operation at 5415 s and with experiment

termination at 6845 s.

The feed and bleed operation was initiated at 5415 s when the peak

cladding temperature in the core reached 967 K. The feed and bleed

operation was affective as anticipated to depressurize the system at an

increased rate and to terminate the cladding heatup and even caused the

lower section (below 0.36 m elevation) of the core quenched as seen in

Figure 11. The primary system pressure dropped to the accumulator

injection set point in about 140 s. The fuel cladding cooling induced by the

9
steam generator feed ari bleed operntion was ennfnnced by thre Initiation of

the accumulator. The core was completely quenched and refilled in 242 s

after the accumulator injection started.

3. RELAP5/MOD2 COMPUTER CODE SIMULATION OF EXPERIMENT LP-SB-3

The RELAP5/MOD2 computer code was used for the post-test

calculation of Experiment LP-SO-3. RELAP5/MOD2 is an advanced, best

estimate computer program developed at the Idaho National Engineering

Laboratory (INEL) for the analysis of Loss-of-Coolant Accident (LOCA) and

other PWR transients. The specific application of the code to the

experiment LP-SO-3 post-test simulation is discussed in this section.

3.1 RELAP5/MOD2 Description

RELAP5/MOD2 employs a finite-difference fluid cell representation of

the primary and secondary coolant systems. The six-equation

hydrodynamic formulation employs separate equations to describe the

conservation of mass, momentum and energy for liquid and steam within

each fluid cell. The description of the hydrodynamics is essentially one

dimensional within each fluid cell. The inclusion of a simplified treatment

of the conservation of momentum in the direction perpendicular to the

main stream flow, where cross flow occurs between parallel volumes and

in branches, brings a special treatment of two-dimensional effects.

Description of the hydrodynamics of choked flow, stratified flow, and

abrupt area changes are carried out with special process models. Special

models are included for simulation of particular components, such as

pumps and accumulators. Flow-regime-dependent constitutive equations

and heat transfer packages are incorporated to complement the

hydrodynamic description. Conduction of heat within metalwork or fuel

rods is calculated with one-dimensional (two-dimensional in fuel cladding

for reflooding simulation) finite difference formulation. A powerful

control and trip logic capability is built into the code.

10
3.2 RELAP5/MOD2 Nodalization and Input Model for Experiment LP-SB-3

The nodalization used for the base case calculation was based on the

RELAP5/MOD I input deck that was used for the planning and prediction of

the experiment at the INEL. The deck was first converted to the

RELAP5/MOD2 format and fallowing changes were applied:

a- to utilize cross flow junction model at the pressurizer connection to

the hot leg to obtain an improved pressurizer drainage,

b- to correct minor errors in the main steam valve flow area, and in

some of the control variables mainly calculating heat loss, mass

etc,

c- to tune various valve opening and closure timing.

Figure 17 presents the nodalization diagram used for the base case

calculation. Some more updates were done during the steady state

calculations and early phase of the transient in order to:

a-match the by pass leak flows between the downcomer and upper

plenum and through the reflood assist piping valve with the

experience,

b-tune drainage timing of the pressurizer by increasing flow loss

coefficients of the surge line,

c-tune MSCV leak to match the first MSCV cycling time,

a-reorientate the by-pass flow junction connecting the downcomer

upper annulus to upper plenum upper volume to avoid subcooled state

being calculated in the upper two volumes of the upper plenum. The

final connection was between the downcomer upper annulus and the

upper plenum nozzle area.

'The final version of the input data listing is supplied in Appendix B.

The input model consists a total of 32 fluid cells for the vessel and

100 cells for the remainder of the primary, secondary and ECC systems.

Characteristics of the model are as follows

a- a split downcomer upper annulus was used with the cross flow

connections,

b- filler gap was separately modeled,

c- core was modeled by a average channel approach and represented by

II
six equal length heat slabs,

d- reactor vessel internals and loop piping, pressurizer and steam

generator walls were simulated with heat slabs. Heat losses from

the piping to the environment was modeled,

e- steam leak through the steam valve was specially modeled,

e- one velocity break flow model was used at the break. Discharge

coefficients for the subcooled and saturated flows were 0.93 and

0.8 1, respectively.

4. RESULTS OF RELAP5/MOD2 POST-TEST CALCULATION

This section presents the thermal-hydraulic results of Experiment

LP-SB-3 post-test calculation. Prior to performing the post-test

calculation, a steady state calculation was performed to obtain the initial.

conditions measured during the experiment. Following the steady state

calculation, transient calculation was started with the trip set points

taken from the experiment. The following subsections discuss the steady

state and transient calculations.

4.1 Calculation of the Steady State

Using a steady state controller package the simulated LOFT system was

brought to the required initial conditions. At first, steady state option of

RELAP5/1MOD2 was used to automaticallu terminate the calculation -when

the required steady state was achieved. However, this option terminated

the calculation at a time when various key parameters were too far off

from their steady state values. Therefore, the steady state calculation

was performed with the transient option. The calculation was continued

until the observed variations of the calculated values of these parameters

from their desired values were acceptable. The key parameters controlled

using the control variables were the primary system pressure, pressurizer

level, cold leg temperature, primary system mass flow rate and steam

generator secondary level. The behaviors of the secondary side feed and

steam flows, pump speed and head, pressurizer heater power, pressurizer

spray valve and steam generator main steam valve positions, and primary

side charge or let down flows were the other parameters checked for the

12
steady state.

The system pressure was controlled by the pressurizer spray which

injected cold leg fluid to the pressurizer to reduce the pressure if the

pressure was calculated to be greater than the measured value. The second

controller on the system pressure was the pressurizer heaters. These

heaters were charged if the pressure was lower than the set point. These

heaters, although in realty ware located close to the bottom of the

pressurizer, were placed at the mixture level in the RELAP5 model to

increase the boiling. The pressurizer level was controlled by two

controllers. One controller which charged fluid at the cold leg temperature

to the cold leg if the pressurizer level was lower than the set point. The

second controller dumped the system fluid to a time dependent volume if

the pressurizer liquid level was higher than the set point. The final values

of the primary pressure and pressurizer level were calculated to be almost

the same as their measured values. The final valve positions controlling

the pressurizer spray, primary system charge or let down flows were zero.

The final pressurizer heater power was zero. The pressurizer surge line

flow was negligible at the end of the steady state calculation.

The primary loop flow was adjusted by using a proportional/integral

controller based on loop flow error to control pump speed. The steady

state intact loop flow was calculated to be the same as the experimental

equivalent. The pump speed and head were in agreement with the measured

initial values. The broken loop flow (from the vessel to the cold leg and via

the reflood assist valve to the hot leg and back to the vessel) was small

and based on the leak flow through the reflood assists by-pass valve. The

total by-pass leak flow based on the flowy loss coefficients used in the

INEL deck was calculated to be about 2% of the total loop flow. This value

compared with the generally accepted 7% of the loop flow was considered

to be too low.

The cold leg temperature was controlled by the main steam valve

position with a proportional/integral control system. Based on the steam

flow rate and heat transfer to the secondary side, the code calculated the

secondary system pressure. Another control logic was used to adjust the

feed flow to control the steam generator required level. This controller

13
was also coupled to the main steam flow. The steam generator level, main

steam and feed water flows were calculated to be the same as measured.

Although the steam and feed water flow rates were correctly calculated,

the steam generator secondary side pressure was the only parameter being

calculated offset by 0.35 Mpa from the measured equivalent.

After about 300 s of calculation the steady state was considered

acceptably stable. Two 150 s long steady state calculations as a

continuation to the previous run were performed:

a-to increase the total by-pass leak flows by decreasing the flow loss

coefficients of the two leak junctions,

b- to increase the secondary side pressure.

The system reached steady state condition and become stable during

the first 150 s long transient and the total by-pass leak flow was

calculated to be 4.5% of the total loop flow. Further trials to increase the

by-pass leak to the generally accepted level were not done due to the

computer cost. The final value was accepted to be adequate. Despite of the

trials done to increase the steam generator secondary pressure, no

success was reached. The complex geometry and atypical internal

structure of the steam generator with rather simple nodalization were

possible causes of the problem.

4.2 Transient Calculation

After the steady state was considered stable, various steady state

controllers associated with the pressurizer heaters, pressurizer spray,

primary system mass charger and let down, pump speed, and various valve

position were removed. The trips for various actions were defined based

on the measured data. Table 2 compares the calculated and the measured

steady state values. The transient calculation was started from time zero

and using the last restart record in the steady state plot-restart file. The

complete transient was calculated in three major restarts. Each restart

used the recent code version available at the time of the run. These were

the cycles 33, 35, and 36.

14
The course of the calculated simulation is discussed in the same

phases that were analyzed in Section 2, namely : (a) mass depletion, (b)

core boil-off, and (c) core cooldown and recovery phases. Table I compares

the timing of various events calculated and measured.

4.2.1 Mass Depletion Phase

This phase begun with the opening of the break, defined as time zero.

Reactor scram followed rapidly (measured at 9.2 1 s, calculated at

10.746 s) when the hot leg pressure fell to the scram setpoint of 14.19

MPa. The steam generator main steam control valve was closed and feed

was isolated upon scram. Primary system coolant pumps were left running

until the primary system inventory fell to 2800 Kg from an initial

inventory of 5600 Kg.

4.2.1.1 Primary and Secondary System Pressure Behavior Calculated

and measured intact loop hot leg and steam generator pressures during the

mass depletion phase are shown respectively in-Figure 18 and 19.

Primary pressure dropped rapidly under the control of the subcooled

blowdown. After the pressurizer level dropped below its indicating range

(corresponding to approximately 8 cm of the calculated level at this time),

the calculated and measured system pressures showed an increased

depressurization rate. This trend continued until saturation condition was

reached at 99 s in the break line upstream of the break during the

experiment, but ceased in the calculation about 15 s after the pressurizer

became empty. This reduction in the calculated depresurization rate

caused the simulated system to reach the end of blowdown condition at

about 200 s. The difference between the depressurization rates measured

between 70 and 99 s and calculated between 77 and 200 s was believed to

be associated with (a) poor modeling of the pressurizer surge line and (b)

the water remaining below the surge line entrance after the pressurizer

level showed below its indicating range and effect of flashing of this

water on the system pressure response. It was doubtful that flow loss

coefficients representing the entrance effects and also the bends in the

line and the nodalization of the 7 meter line with three fluid cells were

adequately modeled. However, for a long experiment having major

phenomena occurring a couple of thousand seconds later, it was believed

15
that this time shift did not affect the subsequent thermal-hudraulic

phenomena. The calculated primary system pressure was controlled

thereafter by the secondary side pressure as in the experiment.

After the main steam control valve was closed and the feed water was

isolated, the secondary side pressure reached the cycling pressure of 7.12

MPa (measured at 87.5 s, and calculated at 76.5 s). The first cycle was

followed by three more cycles, the last one being measured at 1030 s and

calculated at 1175 s. The difference between the measured and calculated

cycle times were associated with the different opening and closure set

points observed during the measurement. After the last cycling, the

calculated secondary pressure continued to increase further, although the

measured pressurize increase ceased shortly after the valve was

(supposedly ) closed. This difference was believed to be either as a result

of improper valve seat position which resulted in more leak then

calculated or more energy transfer calculated to the secondary side.

Calculated primary and secondary pressures were overall in agreement

with the date except with small deviations due to main steam valve

cycling behavior.

4.2.1.2 Break Flow Calculated and measured break flows are shown in

Figure 20. The break flow rose from from zero rapidly to a peak of 6.4 Kg/s

in the calculation and to 6.7 Kg/s in the experiment than fell to 2.5 Kg/s

at the end of the subcooled break flow period and to 1.3 Kg/s just prior to

the pump trip. The break flow calculation except with minor differences

associated with the pressure behavior was in excellent agreement with

the data.

4.2.1.3 Densitg Comparison of the measured and calculated densities

in the intact loop hot and cold legs are shown respectively in Figure 21

and 22. Similar comparison in the break piping is shown in Figure 23.

Overprediction of the density in the hot leg was a result of the difference

between the calculated and measured secondary system pressures after

the fourth main steam valve cycle and its consequent effect on the

primary system pressure. The calculated beak line density was

surprisingly well in agreement with the data.

16
3.6.1.4 Pump behavior Calculated and measured pump pressure

differential are displayed in Figure 24. After the subcooled blowdown

ended the pump pressure differential showed a continuous decrease. The

code calculated the pump degradation indicated in Section 2 slightly

later.The oscillations seen in the data were also observed in the

calculated pump performance but being small in magnitude. The pump

homologous curves (mainly were developed based on the Semiscale pumps6

-nd partially modified using LOFT L3-6 experimental data 7 ) used for this

calculation was responsible for this difference. An excellent agreement

between the calculated pump differential pressure and the data existed

after about 1300 s in the transient.

4.2.1.5 5ystem Mass Inventory Figure 25 presents a comparison of the

primary system mass inventories. An excellent agreement was reached as

a result of a good break mass flow calculation. The 2800 Kg Inventory was

calculated at 1618 s and measured at 1600 s.

4.2.2 Core Boil-Off Phase

This phase starts with the pump trip and continues until the peak

cladding temperature reaches 987 K to Initiate feed and bleed operation.

4.2.2. 1 Primary and Secondanri System Pressure Behavior Comparisons

of calculated and measured pnmary and secondary pressures are presented

in Figures 26 and 27, respectively. The primary side was depressurizing

as the break discharged almost pure steam. The data (Figure 13) showed

that the primary side pressure fell below the secondary level only after

the core uncovery was started. The code calculated similar trend except

between 2300 and 3000 s. when the secondary side pressure was higher

than the primary side as illustrated in Figure 28. This discrepancy was.

related to a trip logic which did not activate the steam valve leak

although the main steam valve was close. The leak was designed to be

activated whenever the steam generator pressure fell below the main

steam valve closure set point. After the last cycle, in absence of this

leek, it took quite a long time until the secondary pressure fell below the

valve closure set point end then the leak was activated.

17
The primary pressure was not calculated to fall below the secondary

pressure after the heat up started. This was due to late core uncovery

calculation, which caused the simulated primary system received slightly

more heat from the core. The core begun to uncover (measured at 3800 s,

calculated at 4008 s) and hence the heat transfer to primary fluid was

reduced. This increased the depressurization rate and the hot leg

saturation temperature fell below the secondary temperature (measured at

4300 s, calculated at 4325 s) as presented in Figure 29 and 14 where the

calculated and measured primary secondary temperature differences are

presented. The sign of the calculated temperature difference reversed, as

in the experiment, shortly after the break isolation at about 4905 s. The

calculated primary secondary pressure difference was much smaller than

the measured difference. As a result of this, the primary secondary

temperature difference was less than the measured difference. Figure 30

presents the condensation rate (negative vapor generation rate) in the

steam generator upright (Volume 115020000). Figure 31 presents the

liquid flow flowing into the steam generator inlet plenum. The code

calculated condensation in the steam generator tubes. The amount of total

condensate was generated at an approximately 40 g/s rate. A reasonable

condensation axial distribution was calculated in the steam generator

tubes, with maximum condensation rate being calculated in the cells next

to the tube sheet and no condensation being calculated in the cells of

which outside pipe surface was uncovered. The condensation ceased in the

steam generator tubes except in the very lower cell at about the initiation

of the core uncoveru. The amount of condensation calculated might be

smaller than what actually occurred during the experiment due to the

calculation of smaller primary-secondary temperature difference.

Figure 32 presents the steam and liquid flow rates in the intact loop hot

leg (Volume 100010000). The code also calculated condensation in the hot

leg due to the heat transfer to the environment. For a very long transient,

the piping inner wall should be at the saturation temperature of the

system couple of thousand seconds in the transient and act as a heat sink.

Since there was no data taken to indicate the transient heat loss from the

piping, it was not possible to verify the calculated condensation rate. The

other important conclusion from Figure 32 is that the liquid and steam

flow rates were the same indicating that the heat and mass transfer and

18
resultant mass balance were calculated properly. The calculated ref lux

condensation in the steam generator upright was restored after the break

isolation. Although the code was not able to calculate the primary

secondary pressure crossover due to different core uncovery timing, the

primary and secondary pressures and continuation of reflux condensation

were satisfactorily calculated

4.2.2.2 Density Figures 33 and 34 present a comparison of the

measured and calculated densities in the intact loop hot and cold legs,

respectively. After the pump trip the liquid in the system fell to the lower

part of the loop, e.i, to the reactor pressure vessel and loop seals,

therefore, a sharp reductions in the densities in the hot and cold leg was

calculated similar to the measured reductions. The code calculated the

liquid build up in the loop seal, as illustrated in Figure 35, and the reactor

vessel, as illustrated in Figure 36 where the collapsed liquid level in the

reactor vessel is shown. The recoveries seen in the measured hot and cold

leg densities immediately following the indicated drops were very well

calculated by the code.

4.2.2.3 Break Flow As noted above, the break flow was of single phase

throughout the boil-off phase, and was very accurately calculated, as

presented in Figure 37.

4.2.2.4 Core Dryout and Thermal Response Figure 38, adopted from the

OLR, presents the measured core dryout information based on the

thermocouples and liquid level probes and also overlaid are the RELAP core

heat slab nodalization and calculated dryout times. RELAP5/MOD2

calculated the dryout initiation in the core approximately 200 s later then

earliest measured dryout based on the thermocouple readings. This figure

shows that the dryout time at 1.397 m (55 in) elevation (which

corresponds to the bottom of the top most heat slab) occurred sometime

between 3830 s (implied by fuel cladding thermocouple reading) and 41032

s (implied by the corner module level probe) (possibly later as will be

explained below). The center bundle thermocouples showed dryout earlier

than the dryout indicated by the center bundle liquid level probes and much

earlier than the dryout indicated by the corner bundle liquid probes.

19
The code (due to the heat transfer logic) waited until the void fraction

in a volume associated with a heat slab reached a critical void fraction

(which is 0.99999 ) after which it could pass to post-dryout heat transfer

mode. Therefore, the calculated dryout times were more relevant to the

dryout indicated by the liquid level probes than the dryout times of the

center bundle cladding thermocouples. The late dryout indicated by the

corner channel liquid probe was believed to be a result of the liquid

drainback from the hot leg before 4300 s or after 4900 s. Since this corner

probe was located under the broken loop hot leg, it could be arguable that

the condensed liquid draining back from the intact loop hot leg was more

effective on the dryout of the fuel moduls 7,4 and 8 which were just

under this leg. Since the code considered homogeneous distribution of

the drained liquid in the nozzle area, which was believed not to be the

case, and homogenous downward liquid flow, the net effect of this drained

liquid postponed the dryout of the core. There were three more uncertain

areas which might be responsiole for this late dryout. These are: (a) not

accurate mass distribution in the primary system, (b) over estimation of

the amount of condensation occurring in the hot leg, and (c) smaller

pressure differential between the hot and cold legs.

Since the code predicted a correct mass inventory in the system as

seen in Figure 25, the distribution of the mass in the system might be

wrong due to artificial liquid hold uo in the hot leg before the dryout begun

and afterwards excessive liquid draining back to the core prevented the

dryout. Because of the center of gravity of the simulated hot or cold legs

are at a different elevation than the center of gravity of the connected

volume in the vessel, liquid hold up might be a problem due to improper

connection simulation. A sensitivity analysis in which the cross flow

connection of the legs to the vessel was planned to investigate the effect

of the cross flow junction model on the dryout initiation.

Since there was no data available to indicate the amount of heat loss

from the hot leg which determines the amount of condensation, no further

analysis was planned.

These junctions might have created a lower back pressure in the upper

plenum than might have consequently caused the liquid to be hanging in

the core longer and resulted in late dryout. Since the flow loss

20
coefficients of the by-pass junctions were decreased during the early

phase of the calculation to increase the by-pass leak flow, another

sensitivity was not planned.

Although the code calculated slightly late initiation of the drjout and

slightly higher core uncover'J rate, initiation end progress of the drijout

was successfully calculated. The calculation indicated that the core was

never completely dryed out, as was in the experiment.

Figures 39 to 44 present a comparison of the thermocouple data with

the calculated temperatures at levels I to 6. Calculated temperatures at

adiabatic boundary condition are also overlaid. It can be seen that the

initial heat up was almost adiabatic, both in the calculated (except at the

sixth elevation) and measured data. As significant cladding to coolant

temperature difference was established more heat transferred to the

steam and the heat up rate, although being higher in the calculation, fell

below the adiabatic rate.

The measured heat up rate at the top of the core was greater than the

adiabatic as the steam temperature was above the cladding temperature

and heat was transferred to the fuel. The calculated cladding temperature

response at the highest core elevation was different than the measured

response due to the followings:

a- the draining liquid was avoiding the heat up of this slab until the

reflux condensation in the steam generator upright ceased at about

4500 s. After this time the heat up rate was approximately the same

as the measurement,

b- the reflux condensation in the steam generator uoright started again

at about 4930 s in the calculation and caused on affective cooling at

this elevation.

Since the incoming liquid was considerably evaporated in the sixth core

volume, the liquid draining back to the lower core elevations did not

affect the other heat slabs. However, in the experiment, the draining liquid

was probably more effective on the fuel rods in the fuel modules 7,4 and 6

after 4900 s. It reduced the cladding and the steam temperatures and

created considerable radial temperature differential between the center

21
of the core and the filler blocks. Since the LOFT core is a small core where

center to filler block distance only 23 times the fuel pitch, this

temperature differential considerably increased the radiation (especially

above 700 K) and convective heat transfers especially from the rods in

the center bundle, which in turn reduced the heatup rate extensively as

seen in Figures 39 to 41 Since three dimensional behavior of the liquid

drainback into the core was not modeled and the code did not have the

radiation heat transfer model, the calculated temperature excursion rates

at the second to the fifth core elevations were much higher than the

measured rate. Therefore, the time at whicn the maximum core

temperature of 816 K was reached (to initiate the closure of the break

valve) was calculated only 163 s later than the measured 4742 s, although

the dryout was calculated approximately 200 s later at the hot

temperature elevation.

4*2.3 Core Cooling and Recovery. Phase

At a peak core cladding temperature of 987 K, the main steam valve

was opened to blow the steam generator and at the same time the feed

water flow was restored. This commenced the feed and bleed phase of the

experiment, the time was 5415 s in the experiment and 5166 s in the

calculation. The feed and bleed operation caused the system to

depressurize to the accumulator injection set point at 5558 s in the

experiment and at 5360 s in the calculation. Accumulator injection

enhanced the cooling of the system and the system pressure was further

reduced to LPIS injection set point at 6785 s in the experiment and at

6222 s in the calculation. The experiment was terminated at 6845 s while

the calculations was stopped at 6560 s.

4.2.3.1 Primary and Secondary System Pressure Behavior Feed and

bleed in the steam generator caused a rapid drop in the secondary side

pressure, followed closely by the highly voided primary. system, showing

that secondary feed and bleed operation was an affective recovery

procedure from a small break LOCA with no HPIS injection. Figures 45 and

46 present a comparison of the measured calculated primary and

secondary system pressures, respectively The rate of depressurization

was very well predicted, except with a slight time shifting caused by the

22
calculated core thermal response during the boil-off phase.

4.2.3.2 Core Thermal Behavior With the feed and bleed operation, the

increased condensation in the steem generator caused an increase in the

steam velocity, and therefore entrainment in the core. This, coupled vfit.",

the level swell caused by the reduction in the pressure resulted in a

partial rewet of the core from bottom up ( Figure 40). The rewet at the

core level below 0.38 m above the core inlet was followed by a re-dryout.

This re-dryout seen was not calculated. The upper section of the core

above 0.38 m above the core inlet started to decrease in temperature. The

code properly calculated the increase in the steam velocity in the core

This was evident from the increase in the temperature of the upper most

heat slab (Figure 44). But this increase ceased and temperature of this

heat slab decreased with the increased cooling in the core. The other core

elevations showed a sharp decrease in temperature until accumulator

injection. The indicated temperature decrease was further enhanced by the

accumulator injection as in the experiment. The difference between the

calculated and measured temperature reduction rates during the period

between the feed and bleed operation and accumulator injection was due

to:

a- entrainment calculation,

b- steam condensation calculation in the steam generator

upright and its drain back to the core,

c- calculated heat transfer coefficient.

The code might have possibly calculated more entrainment which

increased the cooling excessively at elevation higher then 0.38 m. In the

experiment, possibly most of the entrained water evaporated before

reaching the higher elevations for about 200 s after the initiation of the

feed and bleed. This was then followed by a higher entrainment which

caused a very sharp reduction in the temperatures.

The code might have calculated more steam to be condensed in the

steam generator tubes and drained back tothe core.

The code might have over predicted the heat transfer coefficient when
the flow contained less than approximately 20 % liquid void fraction

A split core channel approach might have improved the core behavior,

provided with a detailed upper plenum nodalization, especially in the

nozzle area which should simulate the 3-dimensional licuid drainback

behavior. However, the difficulty in detailed modeling the LOFT upper

plenum which required precise information about the geometry, flow loss

coefficients (especially in the radial direction) did not encourage a

sensitivity study on this issue.

After the accumulator injection started, the core was completely

quenched as in the experiment. The quench progress and temperature, and

rate of temperature drop were well in acreement with the data.

4.2.3.3 Accumulator and LPIS Injection The calculated feed and bleed

operation successfully brought the system pressure, first to the

accumulator injection point, ana shortly 2fter to the LPIS injection point

(Figures 45) as in the experiment. The calculation went smoothly during

both injection periods and no noticeably oscillations were observed.

4.2.3.4 Accumulator and LPIS Flow Figure 47 overlays the calculated

and the measured ECC flows. Figure 48 overlays the integrated ECC flows,

with a time shift applied to the RELAP calculated flow for a better

comparison. As seen from these figures, except very small underprediction

during the first 500 s of the accumulator injection, the calculated

accumulator flow was in agreement with the data. The slight

overprediction of the LPIS flow was observed.

,. SENSITIVITY CALCULATION

A sensitivity study was performed to assess hot and cold leg

connections to the vessel using cross flow junctions. At the time when the

base case calculation was initiated, the original connections of these legs

to the veEsel were kept the same since the standard PELAP5 LOFT deck has

been extensively used using "down connection" of the legs. This connection

scheme created an artificial elevation difference between the center of

the gravities of the volumes representing the legs tnd those of the

24
* volumes to which these legs were connected. This unohysical boundary

condition used by the code might have caused a different hydraulic

behavior during the period when liquid from the hot legs was slowly

draining into the vessel. The cross flow junction model was developed

especially for the flow paths where there are no gravity effects involved.

Since, a late core dryout behavior was calculated, remodeling of the leg

connections using cross flow junctions was tried to see whether any

improvement could be gained.

In order to use the cross flow junction connections, the upper

downcomer annulus, and upper plenum volumes were remodeled. The final

nodalization diagram is presented in Figure 49. The base case calculation

was repeated with the new nodalization from 3200 s to 5000 s.

The following subsections compare various system parameters

calculated using two different types of the lea connections.

5.1 Pressures

The system pressures in the primary and secondary systems were

calculated to be the same as the ones presented earlier. This was expected

and also was a good indication that the model was functioning properly.

5.2 Core Uncovery

Figure 50 and 51 compare the calculated fuel cladding temperatures at

the fifth and third core levels. The dryout was calculated earlier by about

40 s with the cross flow junction connected legs. Since the calculation

was repeated from 3200 s, another approximately 40 s could be gained if

the calculation would have been repeated from the time of the pump trip.

This argument was based on the fact that cross flow junction connection

should have to have effect during the penod when the not legs and the

steam generator tubes contained considerable amount liquid inventory. And

it should have negligible effect during the time when very small amounrt.of

condensed liquid was present in the hot leg. This was verified with the

fact that there was no difference observed between the calculated heat up

rates as seen in Figure 50 and 51.

25
In Section 4.2.2.4, it was indicated that the code predicted the dryout

late by about 200 s. Application of the cross flow junction model at the

hot and the cold legs to vessel connections from the beginning of the

transient would bring this time to approximately to 100 s. This remaining

time difference was believed due to the decreased flow loss coefficients

of the junctions representing the two leak paths which created less

pressure differential between the downcomer and the upper plenum. An

evidence to this conclusion is that as the level progressed into the core,

the pressure differential got smaller and the calculated and the measured

dryout times got closer to each other.

6. RUN STATISTICS

CPU time versus transient time was shown in Figure 52. The mass

depletion phase and early core uncover phases consumed considerable

amount of CPU time. The total CPU time was 7.5 hr on CYBER 176 system

for a total of 6560 s of problem time.. Wherever Courant limit dominated

the time step size (such as Volume 107), an attempt was made to

renodalize that particular volume during the next restart. This problem

mainly occurred when the flow in th6 hot leg started to become stratified.

The code was forced to run faster by assigning relatively larger time step

size as the maximum time step size. This caused the code to violate the

Courant limit for most of the time. The criteria in selection of the

maximum and minimum time step sizes were that number of repeated

time steps was under a reasonable value. During the base case analysis,

10389 time steps were repeated out of a total 114065 time steps. The

grind timea was achieved to be. 1.7931 ms. The run time of the sensitivity

study was practically the same as the base case calculation.

a The grind time is the total CPU time diveded by the total number of

volumes and total number of time steps. A total of 26990 s of CPU

seconds was used for 132 volumes and 1 14065 time steps.

26
7. CONCLUSIONS and RECOMMENDATIONS

This LOFT-OECD LP-SB-3 analyses showed that RELAP5/MOD2 correctly

calculated many major system variables in a small break LOCA, such as

prmary and secondary system pressures, break flow, density and timing

and progression of the dryout. Reflux condensation was calculated in the

steam generator tubes and in the intact loop hot leg.

The major problem encountered was the core heat thermal behavior. The

initial heatup rate was properly calculated except at the highest core

elevations. The deviations started above 600-700 K at the core elevations

between 2 and 5. The heatup rate at the six elevation was more in

agreement with the data after the reflux condensation ceased. Although,

calculated peak cladding temperature was in agreement, the timing was

different. This agreement was due to the termination of the temperature

excursion with feed and bleed initiation based on the peak cladding

temperature. The heat up rate in the experiment was controlled by a strong

cooling which was believed as a result of the draining of the condensed

liquid into the vessel. The condensation was evidently taking place in the

steam generator. Some of the condensed liquid generated in the upright

drained into the vessel. The resultant downword liquid flow was not

uniform across the vessel. Since the LOFT was not instrumented to provide

detailed information on reflux condensation occurring in the steam

generator tubes and possibly in the hot legs, it was not possibleto assess

the code models such as interfacial ,heat'and mass transfer and surface

transfer, controlling the reflux condensation and itM reevaporation on

route to core. The other constraint was that the non homogeneous liquid

distribution in the upper plenum and its asymmetrical fall back into the

core could'not properly be modeled with a one dimensional code, even with

series of cross flow junction connected volumes w•hich simulate upper

plenum and core in detail. The lack of of radiation heat transfer model

increased the overprediction of the heat up rate in addition to the lack of

proper cooling due to the drained liquid. Therefore, it should not be •

expected any better results from the code on the core thermal response.

The code properly calculated the cooling induced by the feed and bleed

operation but the amount of it was overpredicted between the third and

fifth elevations. Although the partial quench seen in the experiment below

0.38 m elevation was reproduced, the code extracted all the stored heat

from the fuel, and could not calculate the re-dryout at this elevation. The
core behavior (quench progression, temperatures) during the accumulator

injection period was very good agreement with the measured data.

The pump degradation was properly modeled. The differences observed

between the calculated and the measured pressure differential across the

pumps were as a result of the pump homologous curves used. These curves

were basically developed using the Semiscale pumps and were partiallu

modified using LOFT Small Break Experiment L3-6 data.

The problem seen in the calculated heatup behavior associated with

improper simulation of the cooling due to the drained liquid and lack of the

radiation heat transfer calculation is believed to be more pronounced in

the small LOFT core. For a large PWR simulation, it is believed that, the

RELAP5/MOD2 code will provide a better simulation of a small break LOCA

occurring under similar conditions, provided that an appropriate

representation of the physical systems should be provided.

In general, the RELAP5/MOD2 code is adequate to predict the transient

thermal-hydraulic behavior of a reactor system undergoing a small break

LOCA. However, the code performance was not only depending on the its

models capabilities, but equally depending on how the physical system was

modeled. The representation of the complicated systems where complex

phenomena are occurrng has always given the problem of in what detailed

a particular component must be modeled. Attempts to represent every

detail may even create problems due to lack of accurate information such

as geometry, flow loss coefficients etc. Experience developed by using the

same code and the same system for different simulations and assessing

with the experimental data is probably the only guidance on this issue.

8. REFERENCES

I. W. B. Grush et al., _uick Look Report on Oecd LOFT Experiment LP-SB-3

OECD LOFT-T-3604, March 1984

2. V. H. Ransom et al., RELAP5/MOD2 Code Manual, EGG-SAAM-6377, April

1984.

3. W. B. Grush et al., Best Estimate Prediction for OECD LOFT Small Break

Experiment LP-SB-3, OECD-LOFT-T-3603. February 1984

4. M. Tanaka, M. Peters, OECD LOFT Experiment Specification Document

28
Cold Leg Small Break Experiment LP-SB-3, OECD-LOFT-T-3602,

December 1983.

5. Private communication with J. Birchley who was UK delegate to the

LOFT-OECD program, May 1985.

6. D. L. Reeder, LOFT. eitm and Test Description, NUREG/CR-0247,

TREE- 1208, July 1978.

7. P. D. Bayless and J. M. Carpenter, Experiment Data Reoort for LOFT

Nuclear Small Break Experiment L3-6 and Severe Core Transient

Experiment LS- I NUREG/CR- 1868,EGG-2075, Januarq 196 1.

29
TABLE I CHRONOLOGY OF EVENTS FOR EXPERIMENT LB-SB-3

Time After

Experiment Initiation

(s)

Measured Data

Event

Small break valve opened

Reactor scrammed on low pressure in intact

hot leg

Main feedwater pump shut off on scram signal

MSCV started to close

Pressurizer level below indicating range

First time MSCV cycled open

Subcooled blowdown ended

Pump differential pressure degradation observed

Last time MSCV cycled open

Primary coolant pumps tripped

Break uncovered

Start of core heatup

Break isolated

Steam generator feed and bleed initiated

Maximum cladding temperature reached

Accumulator injection initiated

Core quenched

Low pressure injection initiated

Experiment terminated

0.0

9.21

9.41

9.5

67.

87.5

98.5

700-875.

1030.

1600.

1612.

3800. ±_50

4742.

5415.

5422.

5558.

5800.

6785.

6845.

Post-Test

Calculation

0.0

10.746

10.746

10.746

94.

75.6

201.

600-1250.

1155.

1618.

1630.

4008.

4905.

5186.

5198.

5360.

5570.

6222.

6560.

30
TABLE 2. 1NITIAL CONDITIONS FOR EXPERIMENT LP-SB-3

Parameter Measured value RELAP-steady state

Primary Coolant System

Power level (MW)

Primary coolant mass flow

rate (kg/s)

Pressurizer pressure (MPa)

Pressurizer liquid level (m)

Cold leg temperature (K)

Hot leg temperature

50.3 ±_1.2

482.6±t2.6

15.26_0. I

1.1 15±0.06

556.6±1

576.6i 1

50.3

482.6

15.24

1.094

556.6

576.12

Secondary Coolant System

Steam generator liquid

levela (m)

Pressure (Mpa)

Mass flow rate (kg/s)

0.21

5.58 ±t0.06

26.67*_0.77

0.21

5.21

27.05

.4

a. The level is defined as 0.0 at 2.95 m above the top of the tube sheet

31
is

-Feed Water

isolated

- -Pressurizer empty

N3•

First MSCV cycled open

Su•.• bcooled blowdown ended

as so 75 1"e Its 1s7 9

T I ae Is

Figure 1. Primary and secondary system pressures (0 - 150 s)
to

a rump aegraaatoIUn - - - -

--. Break uncovered

Pumps tripped CO

.Break isolated

tn

s- - SG feed and bleed initia

aAccumulator injection init

6 lose no" 346 40O8

Time (s)

54" 94""S

Figure 2. Primary and Secondary system pressures (0 - 7000 s)
7. S"

- 7. "a

L.

2:

W 5

in *@2

0 see 1"S is. U..

Time (s)

Figure 3. Steam generator secondary pressure and MSCV cycling set points
to"

-Sa

o4"*

269

DE-PC-SO5B

Pumps tripped

Break isolated

0

104* z,4 3.04 4a So". I... 7To"

0

Time (s)

Figure 4. Fluid density in the break piping
I0."

?. 6s

W0

CLL.

S *."

a. *so

* 0.

-2.6 0

0 8000 3,g00 so"0 O".

Time (s)

Figure 5. Break flow.
110s

- - - Lower beam, U.-P-I,-UU1a -

I[ ~ -Upper beam, DE-PC-002C

-IPumps tripped A - | I i lt ,

"L, 4"

IAccumulator

"J '' - - ~ 'I kill, - 'il in,'ectlonl 1,

- I '11 1."' J

11 

11

•- ~ ~ Jil I ii,

Feed and blee

0 1100 2O 3144 4640o 50 6T0i 7(0s

Time (s)

Figure 6. Fluid density in the intact loop hot leg
s0o"

to - - - Upper beam, DE-PC-OI

coo Pumps tripped

"O 400 I Accumulator

Winjection

' " Feed and bleed,

0 1s alsol 3010 480SD S•@t go"q .1

Time (s)

Figure 7. Fluid density in the intact loop cold leg
Is

" .I ,Pump degradation

• • .. Fourth time F

FIMSCV opened

Tine (s)

Figure 8. Fluid velocity in the intact loop hot leg

IR*
a"*

' 400

Pump degradation

9-

4-)

Ca

QIj

• T 

Time (s)

28~

Figure 9. Pressure rise across the primary coolant pumps
1044

Steam generator feed and bleed

400b

0 if" 0 340 40o o 6oo 7Q4

Time (s)

Figure 10. Fluid density in the intact loop seal
1i66

TE-51107-015

5.- -

- T E - 5 G 0 6 -O 1 1 / " ' -" I

Ew"46 -- 60"

o i ( s )

F g r 11 . Fuel c,0t eb

Time (s)

Figure 11. Fuel cladding teniperaturesin the center bundle
1600

TE-4G24-045 -

TE-2G14-045

700

65 4/64"6"94

Time (s)

Figure 12. Cladding temperatures at different radial positions
- * 900

C_

off

Ln

Tie s

Figure 13. Primary-secondary pressure differential
to

4-3

CL

4-a

tJ

tICi1

3• 35 4 45f 45f o4soW ss~e

Tic (s)

Figure 14. The temperature difference between the saturation temperature of the

primary system and the temperature of the secondary system
1s

E

0 1000 . 2000 3000 4000 S000 6000 7000

Time (s)

Figure 15. Steam superheat in the steam generator inlet plenuni
a, I I l i I I I 1 1 1 1I1IT - 1 1 1 1III

0 Corner fuel module liquid probe 0 0

Ok~ 0

60 • "Center fuel module liquid probe

• 0 Core dryout

Quenchin~, 4A

.0

A A 0

S.-

0

C-3

as.. 400o 4so 4600 go## 6

Time (s)

Figure 16. Response of reactor vessel liquid level probes, core dryout and quench data
125 Component

R2D Component

525 Sam[5J Pipe with 2 volumes

generarIs Fn, o°

5Pump

a gsuie Bra~it fnch with junctionsi 5- Plassurilef -E

5 5 5I5 415 4 2.2 Juncltion of valve

2 - treatstrlCuctue

-7 Reactor 5

vessel 7I3 335 410

S405

3 8 -, 1 m

Figure 17. RELAP5/t1OD2 Nodalization for Experiment LP-SB-3
1~ *oe

52 *

CU

0~

10 ~t.

'U

1-

~ 3.s* 5-

4 * 5@

0 see so". is# e0

Time (s)

Figure 18. Comparison of pressures in the intact loop hot leg
7.60

Lu 'U

di

0 so** "N

Time (s)

Figure 19. Comparison of pressures in the steam generator secondary side
to

a

tn 0

1-~

tS

0d

Z:.

6

4

a

S

-a

0 it"o ages to"

Time (s)

Figure 20. Comparison of break flows
1600

-- V Ii U u u eaIMuuL-ri

- - Lo- Upper heamDE-PC

C .. Pupstrppd iA

'IA

240 Lower beain,OE-PC-002A

* Pumps tripped-'

I I I I I , i I i , i I I

Tseee 1(s

Time (s)

to

Figure 21. Comparison of densities in the intact loop hot leg
to"1

000

w 4

0j

0 6" is" is" so"

Time (s)

Figure 22. Comparison of densities in the intact loop cold leg
1046

e oil

all

I3 0

'V.Tim (s)l

sees go"4

Figure 23. Comparison of densities in the break line
-* .. _____PDE-PC-001

4 J

C

LnL

Lfl

I ~0 IA too#I IS

Time (s)

IS/S3'131.

to"e

Figure 24. Comparison of pressure differentials across pump
I..S

>• _• ... _•Data Break isolated

oo }•D/RELALPIS

z=ee -injection

•- z Pumips

tripped

Accuinu1 at6r

injection

1"4e 2oo 3090. 48H "so So

Time (s)

Figure 25. Comparison of primary system mass inventories
3.00

3.*00

1500 8504 3500 450" mTeI

TIme (s)

Figure 26. Comparison of pressures in the intact loop hot leg
l .06

6.*

L.

4.00

is". 2m0e 44" 55lS

Time (s)

Figure 27. Comparison of pressures in the steam generator secondary side
1. *.

0.O

SCL

4.00

3.00

15"0 as". 3560 4500

Time (s)

Figure 28. Comparison of calculated pressures in the primary and secondary systems
tio. .,

- 1 1O. o S- - - - - - - - - -- -- -

E

a' L--

4~i

40o" 460sees 6eo04o

Time (S)

Figure 29. Calculated temperature difference betwleen tihe primary and secondary systems
1.4

Ul)

E

,k-- -

0b

Volume 11502

' *, 4-0seS -eI so

4@44 4SS* $oeo ;see

Time (s)

Figure 30. Vapor generation rate in the steam generator uprigth
6.2O

tj) --

4-)

10 4C-IS - - - - - - - - - - - -

0

In

ILLL-I I I-_..LI.I1 I L . L LLLI

3040 30SG 40* 4SO0 S60

Time (s)

Figure 31. Calculated liquid flow rate in the steam generator inlet plenum

6S5"
O.a

11 it

4J ~f ~ I/~1~

In

343040666 45606#

TIme (s)

Figure sL.Calculated liquid and vap)or flow rates in the intact loop hot leg
to"

-0 Lower beam, DE.

-- -Upper beam, DEsee

MEI

4.. ii Pumps tripped

o• 400

02s 3%I0 4 rp.

Time (s)

SS."

Figure33.. Comparison of densities in the intact loop hot leg
too.

RELAP5

------- Middle beam, DE-PC-OOIB

Lower beam, DE-PC-OOIA

Upper beam, DE-PC-OOIC

I

EO0

E

ulM

a)

4..

260

,,Pumps tripped

- I 1I f

, , , , , , , , I , , , , , ,I, , , , 1I , , , ,,11 i i i,, , , , , , ,

la

-•t*

2566 3N" 454* SIM

Time (s)

Figure 34. Comparison of densities. in the intact loop cold leg
see

E

cri

2SO. 41e" 550 6644

TIne (s)

Figure 35. Comparison of densities ill the loop seal
7.5

s.. -

Top of the core

.2:

Buttom; of the core

toee ee seeso ,oo"o

Time (s)

Figure 36. Calculated collapsed liquid level in the reactor vessel
a

cnj

2~e s" 25" 4"6e 66"

Time()

Figure 37. Comparison of break flow rates
?s

q--

-c

w

a,

S.

0

Li

2s

0

4640 5560 at"0

Time (s)

Figure 38. Comparfson of dryoutSmeasured and calculated
____TE-5106-005

----- ilELAP5(riode-1)

S700

0.0

E

3S"444 4s" 615548)

Time (s)

Figure 39. Comparison of cladding temperatures at 12.7 cm elevation
.-- _____RELAP5 (node-2)

- e-- Adiabatic heatup

--------- TE-4t113-015

TE- 5H07-015

E - /- .- - - 74 -

354J40" 4600 5S4d 56"J

Tine (s)

0"0

Figure 40. Comparison of fuel cladding temperatures at 38 cm elevation
lose

TE-21115-026 S"~/.- "- -,,\-

Adiabatic heatup

SS.-

3os" 4 "t 4see SS"N g ,o,"

Time (s)

Figure 41. Comparison of fuel cladding tempera-tures at 69 cm elevation
I*"o

TE-51106-43.8 ----"

-. o--- o Adiabatic heatup /

E

j-7-

6"

/- , I TII- (s) Is

6*O*

Figure 42. Comparison of fuel cladding temperature at 102 cm elevation
Ad iaat/ce /t// p"

-/ I,-/ --- Adiabatic heatup

0.o 0 -. I/

E4 77

600

Time (s)

Figure 43 Comparison of fuel cladding temperat:ures at 114 cm elevation
- I*"0. 

"s

see

4-S

Ln

E

600

460t

3S4 40"0 so" 5500W 60"s

Time (s)

Figure 44. Comparison of fuel cladding temperatures at 157 cm elevation
4. 00

N,\•Feed and bleed

Accumulator injection

NI

14 LPIS injection

" 1 - i LPS i

M%

I I ' I I I , .l I II

10 **

ss oo s o 6 oog"o

Time Is)

Figure 45. Ctomparison of pressures in the intact loop hot leg
4.00

CL

tn

a~Ll

S1j

S... it.. 655*

Time (s)

Figure 46. Comparison of pressures in the secondary side
is

I.

C)

4-A

Mu S

0

UInL

-I, I I I I i t

see., 550O0 £0.00

Tine (s)

Figure 47. Comparison of ECC injection flow rates

as" To
so4

M_ Data

o - RELAP5 I-.-

C

0

4-)

CU

*,ma

(T I

Ft Li.

w

ciF SO 80

Ti'Us

Fiue4.Lmaio fitgae C neto lw
5113

Steam

gelnerior

125

522

Cormponent

Component

Pipe with 2 volumes

Pump

DOinci with functions

PFesuflzer

515 unction or valve

bliat stniciute

Reactor

vessegl r

C) C

350

U

224

Figure 49. RELAP5/MOD2 Nodalization for sensitivity analysis
941

S-p

4-1p

COp

CLp

4"0 -2 4446 9 "

Tim (s

Figure 50. Comparison of fuel cladding tempeartures at level 5
------- cross flow juniction connection ror tie legs

m/

- /

4i - co 70/ -

U - - ri

+ee400 44SO 4600 4n9" S"o

,Time (s)

Figure 51. Comparison of fuel cladding tbipeartures at level 3
30444

2zoo#

02

U)

a....

Li

Its" 3"0 5. V so ____-..pw ,-* .

Tie* ($I!)

Figure52 . CPU time versus transient timae and required- CPU seconds per. real second. ..
APPENDIX A

SYSTEM CONFIGURATION

The LOFT facility was designed to simulate the major components and

system responses of a commercial pressurized water reactor (PWR) during

a LOCA. The experimental assembly includes five major subsystems which

have been instrumented such that system variables can be measured and

recorded during a LOCA simulation. The subsystems include the reactor

vessel, the intact loop, the broken loop, the blowdown suppression system

(BST), and the ECC systems. Complete information on the LOFT systems is

provide in Reference A-I and a discussion of the LOFT scaling philosophy

is provided in Reference A-2. details of LOFT measurement capability and

its instrumentation can be found in Reference A-I.

The arrangement of the major LOFT components is shown in Figures

A-I.The intact loop simulated three loops of a commercial four-loop PWR

and contains a steam generator, two primary coolant pumps in parallel, a

pressurizer, a Venturi flowmeter, and connecting piping. A spool piece

was connected to the intact loop cold leg downstream of the pump

discharge. A schematic of the break piping and nozzle configuration is

shown in Figure A-2. The break piping provided the break path during the

blowdown. This line was closed when the maximum temperature in the

core reached 977 K.

The broken loop consists of a hot leg and a cold leg, each of which are

connected to the reactor vessel. The broken loop hot leg also contains a

simulated steam generator and a simulated pump. These simulators have

hydraulic orifice plate assemblies which have similar (passive)

resistances to flow as an active steam generator. Both of the legs have

quick opening and isolation valves. During Experiment LP-SB-3, these

valves remained closed because the break was in the intact loop cold leg.

The two LOFT ECC systems are capable of simulating the emergency

injection of a commercial PWR. They each consists of an accumulator, a

high-pressure injection system, and a low pressure injection system.

84
Du'rint Exreriment LP-SB-3, the high pressure injection system was not

used. Accumulator 'A" was used to inject a volume scaled amount of !iauid

to represent the liquid volume of three accumulators in the reference

three-loop PWR.

The LOFT steam generator, located in the intact loop, is a vertical

U-tube design steam generator. The main steam valve is automatically

regulates the secondary side pressure within given limits.

The reactor contains 1300 unpressurized fuel rods in five square (15 bu

15 assemblies) and four triangular (containing 76 fuel rods) corner fuel

modules, shown in Figure A-3 and described in Reference A- 1. The center

fuel module is highly instrumented. Two of the corner and one of the

square assemblies are not instrumented. The fuel rods have an active

length of 1.67 m and an outside diameter of 10.72 mm.

A- I References

A- i. D. L. Reeder, LOFT System and Test Description, NUREG-CR-02d7.

TREE- 1208, July 1978.

A-2. L.J. Ybarando et al., Examiration of LOFT Scaling, 74-WA-hT-53.

Proceeding of the Winter Meeting of American Society of

Mechanical Engineers, New York, November 1974, CONF-741 104.
Inlact loop Broken loop

" BL. Steam . .. . .. . . :.. Quick opening

• .n .generator I - valve (2) experimental simulator

.. . xpe~lmmlla~l Dovwncorner I

mIneasurement . .-- injection -,. " -. s,,,ion. IStGeam location /-"

" It3e ,eat(: a= / :• F vAai-vr t2 ;

- ,- " ,'-,..•• • ..

Pressurizer.- . Pu i

- ,. .- .nm ..-. a

0) Drag c•. " . ECC cold 1g n ast• n.en, t and turbint -" \ :., i . injection ,sai.;,n.. Pumps Ilocation .

PC.•3 . .Y

e)..erint,-,•,' Reactor

melsurement . " vessel .. Suppress:or,

station . \\ ,. k • . vessel Downco.erBreak nozzle " -- ,.

CV.P.139.57 .. -"Core .

\..£ .. Lower pl"numn:

CV-P139.58 V..

Reactor vessel

INEL-L P S1B,3.1510 IU DZ)I

Fiqure A-1. LOFT configuration for c'ld 1(. in;!..t loop •,l1i break Li'-SB-3
2.64 cm

(1.04 in.) Z

0.4.-4cm R

(0.!7 in.) \

0.88 cm

(0.35 in.),

From PC.1, intact 3.08 cm

(1.21 ir.1

loop cold leg .

1.25-in.dia I

c

Break nozzle insert

Detail 'A'

C'j.

Gamma drnsitometer spool piece

DE.PC-S05A. B 8

DE.PC-SC5ABD. BD"

To BST

CV-P139-58 "C _

• 

C.P139 

.5 7

-TE PC-S09

TE-PC-S0O

PE-PC-SO10

DST POE-PC.SO5 Spool piece

FE.PC.S04

ME-PC-S04

C.

r.~j PE.PC.SO11

dia

Break nozzle

(see Detail A) I.NEL-LP-Sa-3. 12504

Figure A-2. Experiment break piping instrument configuration.

87
00

0O

Broken loop - 4

cold log .. ..

t_;_•_.I .;. ....

II':" II" a1It

.-' -"- -•

leg 'P111 'J t"12"

L... o.ffiI . .

I': . ,- Vt •

'- !' -. ,., , , .. ..•7 'i .7! ... .._!;.;!'. ••L~tt~l. '''IL

Inac loop " ' i

ho .l._e..g --

-'I

4.

-I,

4.

4-

'a-

.1

Broken loop

hot leg 1

! ... LCore posi

L. Fuel

B 

I

Ilion Fuel modulo types

module deslgnation Index

Ii a'!... -'',ii, la:, t

1 j T

a-1" ..... '...'-*4

*li - -:.

a. *ti.'1 .. : 

..L; _,.-Intact loop

.... ....... I ' • * I leg'-- * -

I

iI - r -b.

f]

0

Fuel fr."i

"Fl,•t'.T.oc ",'ptl 196)

h.,,-.Jc(touplC lower tic plate It 7i

[).:nrr . ttlarT;OcOtlA(. er 

Na;ut on flux detector. fixed (4)

Neluutron flux scan (4l

Ca-iduct'vi:y hqta:d level detector 131 .1

-i

ii

ItJEL-A.t6 43?

1:1(ti r I: A- 3. r J.!. ('T •,•' ra n i"•r *f i ti ,,;- ;i ndl Ins;tirnm,'ntt.iti fn
%RC FORM 336- U.t. NUCLIAR REGULATORY COMMISSION I RiPOrT NUMGER (A..,#,Wav riot " d WV v Ao. vWy vi

Q2 NACM 841 1102. 320M. J2, BIBLIOGRAPHIC DATA SHEET NUREG/IA-0018

sit INSTRUCTIONS ON THE RIVIERS$

2 TITLE AND SUITITLE J LEAVE CIANK

RELAP5/MOD2 Assessment, OECD-LOFT Small Break

Experiment LP-SB-3 4 DATE REPORT COMPLETED

MONTM I TEAR

AUT.ORISI , February 1986 AJ OAT(! IEPORT ISSUED

S. Guntay MONC I E ,AR

April 1990

7 PERFORMING ORGANIZATION NAME AND MAI"ING AOODRESS tI.&Z•, Z C* 1 8. P"OJICTITASKIWORK UNIT NUMSER

EIR

CH-5303 Wurenlingen 9 IN OR GRANT NUMOIR

Swi tzerl and A6827

10 SPONSORING ORGANIZATION NAME AND MAILING ADDRESS IlmhIw Z Code)I Ili. TYPE OF REPORT

Office of Nuclear Regulatory Research Technical Report

U.S. Nuclear Regulatory Commission 6.,, COV,., .ft.-

Washington, DC 20555

12 SUPPLEMENTARY NOTES

I1. AGSTRACT (9W' A.* be PO

An analysis of the experimental results and post-test calculations using

RELAP5/MOD2 carried out for OECD-LOFT small break experiment LP-SB,3 are

presented. Experiment LP-SB-3 was conducted on March 5, 1984 in the

Loss-of-Fluid Test (LOFT) facility located at the Idaho National Engineering

Laboratory (INEL). The experiment simulated a small cold leg break, with

concurrent loss of high pressure injection system, and cooldown and recovery

by feed and bleed of the steam generator secondary side and accumulator

injection respectively.

This report documents a short post-test analysis of the experiment emphasizing

the results of additional analysis performed during the course of this task.

RELAP5/MOD2 input model and results of the post-test calculation are

documented. Included in the report is the results of a sensitivity analysis

which show the predicted thermal-hydraulic response to a different input model

IA DOCUMENT ANALYSIS a KSWONIS'OSCRIPTDOS

RELAP5/MOD2, ICAP Program, Small Break LOCA, LOFT

a fO(NTIPI(RsiOPEN| OeO TERMS




UNITED STATES

NUCLEAR REGULATORY COMMISSION

WASHINGTON, D.C. 20555

OFFICIAL BUSINESS

PENALTY FOR PRIVATE USE. $300

SPECAL FOURTH-CLASSRATE

POSAGE& FESPAID

1 USNRCj