NUREG/IA-0147

International

Agreement Report

Assessment of RELAP5/MOD3.2

for Steam Condensation

Experiments in the Presence of

Noncondensibles in a Vertical Tube

of PCCS

Prepared by

H. S. Park, H. C. No/KAIST

Y. S. Bang, K. W Seul, H. J. Kim/KINS

Korea Advanced Institute of Science & Technology Korea Institute of Nuclear Safety

373-1 Kuseong-Dong, Yusong-Ku P.O. Box 114

Taejon, Korea Yusong, Taejon

Korea 305-600

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

September 1998

Prepared as part of

The Agreement on Research Participation and Technical Exchange under the

International Thermal-Hydraulic Code Assessment and Maintenance Program (CAMP)

Published by

U.S. Nuclear Regulatory Commission
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NUREG/IA-0147

International

Agreement Report

Assessment of RELAP5/MOD3.2

for Steam Condensation

Experiments in the Presence of

Noncondensibles in a Vertical Tube

of PCCS

Prepared by

H. S. Park, H. C. No/KAIST

Y. S. Bang, K. W Seul, H. J. Kim/KINS

Korea Advanced Institute of Science & Technology Korea Institute of Nuclear Safety

373-1 Kuseong-Dong, Yusong-Ku P.O. Box 114

Taejon, Korea Yusong, Taejon

Korea 305-600

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

September 1998

Prepared as part of

The Agreement on Research Participation and Technical Exchange under the

International Thermal-Hydraulic Code Assessment and Maintenance Program (CAMP)

Published by

U.S. Nuclear Regulatory Commission

Assessment of RELAP5/MOD3.2 for Steam Condensation Experiments in

the Presence of Noncondensables in a Vertical Tube of PCCS

Abstract

This report deals with the application of RELAP5/MOD3.2 to condensation experiments

in the presence of noncondensable gases in a vertical tube of Passive Containment Cooling

System. When steam-noncondensable gas mixture was injected into the vertical tube, steam

was condensed on the inner surface of the condensing tube but the noncondensable gas greatly

inhibited the condensation of the steam. As the scattering of previous experimental data was

large, the present experimental apparatus was set up to get a reliable data on the condensation heat transfer coefficient of the steam-noncondensable gas mixture in a vertical tube. The

experimental results show that the condensation heat transfer coefficient increases as the inlet

steam-air mixture flow rate increases, the inlet air mass fraction decreases, and the inlet saturated steam temperature decreases.! There are two wall film condensation models, the default

model and the alternative model, in RELAP5/MOD3.2. After a condensation database was

constructed, two models were assessed directly with the data of the database. The experimental apparatus was also modeled with RELAP5/MOD3.2, and simulations were performed

for several sub-tests to be compared with the experimental results. The simulation results

show that in overall sense the default model of RELAP5/MOD3.2 under-predicts the heat

transfer coefficients, but that the alternative model of RELAP5/MOD3.2 over-predicts them

throughout the condensing tube. Different from the modeling results of RELAP5/MOD3.1,

the change in the number of the nodes for condensing tube has little influence on the simulation results of RELAP5/MOD3.2 both with the default and with the alternative model.

From the sensitivity study of input parameters it is also shown that the effects of the coolant

flow rate, the inlet coolant temperature and the vented mixture temperature are negligible,

but that the effects of the inlet mixture flow rate, the inlet saturated steam temperature and

the inlet air mass fraction are significant. Run statistics show that the grind time of the

default model is always about 23% higher than that of the alternative model.

iii

Contents

Abstract iii

Table of Contents v

List of Tables vii

List of Figures viii

Executive Summary xiii

Acknowledgment xvii

1 Introduction 1

1.1 Background ..................................... 1

1.2 Objectives and Report Organization ....................... 4

2 Condensation Experiments 5

2.1 Objectives of the Present Experiments ...................... 5

2.2 Test Facility and Its Instrumentation ....................... 6

2.3 Test Matrix and Experimental Procedure .................... 10

2.4 Experimental Results and Discussions ...................... 11

2.4.1 Effects of inlet steam-air mixture flow rate .................... 14

2.4.2 Effects of inlet air mass fraction .......................... 17

V
2.4.3 Effects of inlet saturated steam temperature .................. 20

3 Assessment of RELAP5/MOD3.2 22

3.1 Two Wall Film Condensation Models of RELAP5/MOD3.2 ............. 22

3.1.1 Default condensation model .............................. 22

3.1.2 Alternative condensation model ........................... 25

3.2 Direct Assessment of Two Wall Film Condensation Models ............ 26

3.3 RELAP5/MOD3.2 Nodalization ......................... 31

3.4 Simulation Results and Discussion ........................ 33

3.4.1 Base case calculation .................................. 33

3.4.2 Nodalization study ................................... 43

3.4.3 Sensitivity study of input parameters ....................... 43

3.4.4 Run statistics ....................................... 51

4 Conclusions and Recommendations 60

References 63

Appendices 67

A Test facility and its instrumentation 67

B Data reduction method 71

C RELAP5/MOD3.2 input listing 75

vi
List of Tables

2.1 Comparison of experimental operating ranges between Siddique's and Kuhn's 6

2.2 Instrumentation of KAIST condensation experimental apparatus ......... 9

2.3 Test matrix of KAIST condensation experiment ..................... 10

3.1 Heat transfer database for the vertical laminar film condensation ....... .. 26

3.2 Input parameters for base case calculations ........................ 34

3.3 Input parameters for the sensitivity study ......................... 45

3.4 The CPU time and the grind time of the simulation of KAIST condensation

experiment ........ ..................................... 58

3.5 The CPU time and the grind time of the sensitivity study .............. 59

vii
List of Figures

1.1 Passive containment cooling system of CP-1300 ..................... 2

2.1 A schematic diagram of KAIST condensation experimental apparatus . . .. 8

2.2 Condensation phenomena in a vertical tube in the presence of noncondensable

gas ......... .......................................... 12

2.3 Local temperature distributions in experiment E13b .................. 14

2.4 Local heat flux distribution in experiment E13b ..................... 15

2.5 Comparison of experimental heat fluxes with the variation of the inlet steamair mixture flow rate ....................................... 16

2.6 Comparison of experimental heat transfer coefficients with the variation of the

inlet steam-air mixture flow rate ............................... 16

2.7 Comparison of experimental temperature distributions with the variation of

the inlet steam-air mixture flow rate at the inlet mixture temperature of 120*G

with 20% air ........ .................................... 18

2.8 Comparison of experimental heat fluxes with the variation of the inlet air mass

fraction ........ ....................................... 19

2.9 Comparison of experimental heat transfer coefficients with the variation of the

inlet air mass fraction ....... ............................... 19

2.10 Comparison of experimental heat fluxes with the variation of the inlet saturated steam temperature .................................... 21

viii
2.11 Comparison of experimental heat transfer coefficients with the variation of the

inlet saturated steam temperature .............................. 21

3.1 Assessment of default model of RELAPS/MOD3.2 ................... 27

3.2 Assessment of two wall film condensation models of RELAP5/MOD3.2 for

pure steam ........ ..................................... 28

3.3 Multiplication factors in alternative model of RELAP5/MOD3.2 ....... .. 29

3.4 Assessment of alternative model of RELAP5/MOD3.2 with steam-air experimental data .......... ................................... 30

3.5 Assessment of alternative model of RELAP5/MOD3.2 with steam-helium experimental data .......................................... 31

3.6 Nodalization scheme of the RELAP5/MOD3.2 code for PCCS experimental

facility ................................................ 32

3.7 Comparison of steam velocities calculated using two condensation models with

the experimental data of E13b ...... .......................... 34

3.8 Comparison of air mass fractions calculated using two condensation models

with the experimental data of E13b ...... ....................... 35

3.9 Comparison of temperatures calculated using two condensation models with

the experimental data of E13b ...... .......................... 36

3.10 Comparison of coolant side heat transfer coefficients calculated using two condensation models with the experimental data of E13b ................. 37

3.11 Comparison of heat fluxes calculated using two condensation models with the

experimental data of E13b ....... ............................ 38

3.12 Comparison of condensation heat transfer coefficients calculated using two

condensation models with the experimental data of E13b ............... 39

3.13 Comparison of condensation heat transfer coefficients calculated using two

condensation models with the experimental data of El2a .............. 40

ix
3.14 Comparison of condensation heat transfer coefficients calculated using two

condensation models with the experimental data of End ............... 40

3.15 Comparison of condensation heat transfer coefficients calculated using two

condensation models with the experimental data of E12b .............. 41

3.16 Comparison of condensation heat transfer coefficients calculated using two

condensation models with the experimental data of E4d ............... 42

3.17 Comparison of calculated heat transfer coefficients with the variation of the

node number ........ .................................... 44

3.18 Comparison of calculated air mass fractions with the variation of the steam-air

mixture flow rate in default model ................ ............. 46

3.19 Comparison of calculated air mass fractions with the variation of the steam-air

mixture flow rate in alternative model ............................ 46

3.20 Comparison of calculated heat transfer coefficients with the variation of the

steam-air mixture flow rate in default model ....................... 47

3.21 Comparison of calculated heat transfer coefficients with the variation of the

steam-air mixture flow rate in alternative model ..... ................ 47

3.22 Comparison of calculated air mass fractions with the variation of the inlet air

mass fraction in default model ...... .......................... 49

3.23 Comparison of calculated air mass fractions with the variation of the inlet air

mass fraction in alternative model ...... ........................ 49

3.24 Comparison of calculated heat transfer coefficients with the variation of the

inlet air mass fraction in default model ........................... 50

3.25 Comparison of calculated heat transfer coefficients with the variation of the

inlet air mass fraction in alternative model ........................ 50

3.26 Comparison of calculated air mass fractions with the variation of the inlet

saturated temperature in default model .......................... 52

3.27 Comparison of calculated air mass fractions with the variation of the inlet

saturated temperature in alternative model ........................ 52

X
3.28 Comparison of calculated heat transfer coefficients with the variation of the

inlet saturated temperature in default model ....................... 53

3.29 Comparison of calculated heat transfer coefficients with the variation of the

inlet saturated temperature in alternative model ..................... 53

3.30 Comparison of required CPU times of two condensation models ........ .. 54

3.31 Comparison of time step sizes of two condensation models .............. 55

3.32 Comparison of code determined time steps of two condensation models. . . . 56

xi

Executive Summary

The RELAP5/MOD3.2 code is assessed with steam condensation experiments in the

presence of noncondensables in a vertical tube of PCCS.

CARR Passive reactor(CP-1300), a large passive PWR concepts which was developed at

Center for Advanced Reactor Research(CARR), has a concrete containment and the final

safety functions are achieved through passive systems such as accumulator, Core Makeup

Tank(CMT), Secondary Condenser(SC), and Passive Containment Cooling System(PCCS).

Two concepts using internal condensers or an external condensers are proposed as PCCSs

of CP-1300. The main concern of PCCS with the external condenser is the degradation of

condensation heat transfer due to the accumulation of the air inside the condensing tubes.

The best-estimate safety analysis code, RELAP5/MOD3.2, which was developed at Idaho

National Engineering Laboratory(INEL) based on international cooperation, CAMP, organized by U.S. NRC, is used to analyze the transients and the LOCA of a nuclear power plant.

However much uncertainties are known to exist in some heat transfer correlations of RELAP5/MOD3.2, especially in modeling the condensation phenomena with noncondensable

gases in a vertical condensing tube which is applicable to the design of PCCS and SC. The

experimental apparatus were set up to get reliable data on the condensation heat transfer

coefficient of the steam-noncondensable gas mixture in a vertical tube.

The objective of the present work is to assess the analysis capability of the RELAP5/MOD3.2

code on the condensation of the steam-noncondensable gas mixture in a vertical tube. First

of all, two wall film condensation models in RELAP5/MOD3.2, the default model and the

alternative model, were assessed with the constructed database. Two wall film condensation

models of RELAP5/MOD3.2 were also assessed extensively for several experiments simulating

steam condensation in the presence of noncondensables in a vertical tube of PCCS.

The test facility was nodalized so as to be suitable for simulating the important experimental parameters, which is known to be important in condensation experiments. The base

case calculations were executed, the results were compared with the experimental data, two

wall film condensation models of RELAP5/MOD3.2 were compared each other, and the code

xiii
predictability on the condensation phenomena was discussed. Nodalization study was performed to investigate the effect of the divided node number in the test section. Sensitivity

studies were performed to investigate the effects of input parameters on heat transfer characteristics. Run statistics of two simulation results with different condensation models were

also compared.

From the studies, the followings are obtained:

" The experimental results show that the inlet steam-air mixture flow rate, the inlet air

mass fraction, and the inlet saturated steam temperature have significant effects on

condensation heat transfer. The condensation heat transfer coefficient increases as the

inlet steam-air mixture flow rate increases, the inlet air mass fraction decreases, and

the inlet saturated steam temperature decreases.

" Two wall film condensation models in RELAP5/MOD3.2 are directly assessed with the

constructed condensation database. The default model under-predicts the experimental

data for the low heat transfer coefficient range, but it over-predicts the experimental

data for the high heat transfer coefficient range. In case of the alternative model, the

predicted values are always higher than the experimental data.

" As base case calculations several experiments are simulated by RELAPS/MOD3.2 with

both the default model and the alternative model. When two wall film condensation

models are compared each other, the calculated air mass fraction from the default model

is always lower than that of the alternative model, all the calculated temperatures from

the default model decrease more slowly than those from the alternative model, and

the calculated heat flux from the default model is always lower than that from the

alternative model in the upper part of the condensing tube, and they are similar in the

lower part.

" Those results also show that with low inlet mixture flow rate and high inlet air mass

fraction the default model under-predicts the experimental heat transfer coefficient, but

xiv
that the alternative model over-predicts it. With high inlet mixture flow rate and low

inlet air mass fraction, the default model predicts the experimental data well, but the

alternative model still over-predicts it, especially in the middle of the test section.

" From the nodalization study of RELAP5/MOD3.2, the change in the number of the

nodes for condensing tube has little influence on the simulation results of the condensation phenomena with RELAP5/MOD3.2.

" As a sensitivity study several input parameters are changed to show their effects on

heat transfer characteristics. The simulation results show that the effects of the inlet

steam-air mixture flow rate, the inlet air mass fraction, and the inlet saturated steam

temperature are most influential in condensation experiments in the presence of noncondensable gas in a vertical tube. However the effects of the coolant flow rate, the

inlet coolant temperature, and the vented mixture temperature are shown to be less

influential.

" Run statistics show that the grind time of the default model is always higher than that

of the alternative model and it is a little increased for both models when more volumes

are included in RELAP5/MOD3.2 simulation. The required CPU time is the highest

when the inlet mixture flow rate is the highest and both the inlet air mass fraction and

the inlet saturated steam temperature are the lowest.

* As both the default model and the alternative model can not predict well the experimental data, it is needed to develop a new correlation for wall film condensation based

on the experimental data produced over various operating ranges, and it can be applied as a new condensation model to simulate the condensation heat transfer of the

steam-noncondensable gas mixture in a vertical tube.

xv

Acknowledgment

This report was performed under the sponsorship of Korea Atomic Energy Research

Institute(KAERI). Authors appreciate the KAERI for the financial support.

xvii

Chapter 1

Introduction

1.1 Background

CP-1300[1, 2], a next generation reactor which is a large passive pressurized water reactor

concept, was developed conceptually by CARR. It adopts several passive engineered-safetyfeatures such as accumulator, CMT, SC, and PCCS, which are investigated extensively[3, 4].

Two types of PCCS are proposed as the long term cooling methods of the concrete

containment[5]. An internal condenser concept is the one and an external condenser concept

is the other. The internal condenser concept utilizes the natural circulation of the PCCS pool

water through the condenser tubes inside a containment, while the external condenser concept

utilizes the condensation in condensing tubes, which are located in a water pool outside a

containment. Figure 1.1 is the conceptual design of the PCCS with external condenser of

CP-1300. The objectives of PCCS are to simplify the design of nuclear power plant by the

use of passive concepts and to keep the integrity and safety of the containment during the

reactor accidents. PCCS is operated to remove the steam, which is generated and flows into

a quiescent containment atmosphere during the postulated reactor accidents such as LOCA.

The condensing tubes of PCCS ingest and cool the mixture of steam and noncondensables

from the containment. The concept is similar to the PCCS of SBWR[6, 7, 8], but it is adopted

to apply to CP-1300 with concrete containment. Heat transfer of condensing steam in the

1
Steam-Noncondensable Gas Wnet

Steam

Figure 1.1: Passive containment cooling system of CP-1300

2
presence of noncondensables in a vertical tube of PCCS is necessary to be investigated to

design the PCCS for the CP-1300 and for the SBWR as well.

The local heat transfer phenomena in a vertical tube with noncondensable gases has been

studied extensively by Vierow[9, 10], Ogg[11], Siddique[12, 13], Hasanein[14], Araki[15], and

Kuhn[16], and many researchers have developed their own correlations based on their own

experimental data. However, it was noticed that the existing correlations were greatly scattered when compared with the applicable experimental data. New experimental apparatus

is set up to minimize the entrance effect and multi-dimensional effect, more reliable data

is acquired, and the effects of the various parameters on condensation with noncondensable

gas are investigated. The objectives of the present experiments are described in detail in

Section 2.1.

The capability to deal with steam-noncondensable gas mixtures in vertical tubes of PCCS

is also an important technical problem in the design of the PCCS, and the RELAP5/MOD3.2

code is selected to simulate it.

The RELAP5/MOD3.2 code is an advanced best-estimate thermal-hydraulic transient

analysis code developed at INEL under the sponsorship of the U.S.NRC. The code has been

developed primarily to perform LOss of Coolant Accident(LOCA) analysis for pressurized

water reactor systems. However, being a generic code, the code can be used to simulate

various system configurations so that it has been extensively used for various simulation,

validation, experimental data analysis and for plant/system analysis purposes. It has been

exhaustively assessed with the various experimental data and extensively applied in evaluating

the safety of both the commercial nuclear power plants and the next generation of advanced

passive reactor designs such as SBWR, AP600, and CP-1300. It was also applied to nonLOCA, transient thermal hydraulic analysis for steam-water-noncondensable mixtures[17].

However, it is known that some heat transfer correlations in RELAP5/MOD3.2 have much

uncertainties in predicting condensation-related phenomena. In particular, more reliable

model is needed on condensation phenomena with noncondensable gases in a vertical tube,

which is applicable to the design of the PCCS and SC of CP-1300.

3
1.2 Objectives and Report Organization

The main objective of the present work is to assess the capability of the RELAP5/MOD3.2

code on the condensation of the steam-noncondensable gas mixture in a vertical tube representing condensing tubes in PCCS.

For this purpose, firstly a set of experiments were performed to get a reliable data on the

condensation heat transfer coefficient of the steam-noncondensable gas mixture in a vertical

tube. Secondly two wall condensation models of RELAP5/MOD3.2 were assessed for steam

condensation experiments in the presence of noncondensables. The condensation heat transfer database was constructed by collecting the available local data[18], and the calculation

results with two wall film condensation models in RELAP5/MOD3.2 were compared with the

constructed database to investigate their abilities to predict the heat transfer characteristics.

Also to evaluate the code predictability, the test facility was modeled so as to be suitable

for simulating the important experimental parameters in condensation experiments, and the

simulation results were analyzed.

In Chapter 2 the experimental facility, its instrumentation, the test matrix, and the experimental procedure were described briefly and the experimental results were discussed in

detail. In Chapter 3 the RELAP5/MOD3.2 code was described briefly, their two condensation

models themselves were assessed directly with the constructed database, the nodalization of

the experimental apparatus was summarized, and the simulation results from the base case

calculations, the nodalization study, and the sensitivity study were analyzed and discussed.

The runs statistics were also described. In Chapter 4 the conclusions and recommendations

were summarized. Finally the experimental test facility and its instrumentation were described in detail in Appendix A and the data reduction method was described in Appendix

B. Also the RELAP5/MOD3.2 input decks for the present experiments was listed in Appendix

C.

4
Chapter 2

Condensation Experiments

2.1 Objectives of the Present Experiments

The steam condensation phenomena in a vertical tube with noncondensable gas has begun

to be investigated just recently. Experiments have been performed by Vierow[9, 10], Ogg[l 1],

Siddique[12, 13], Hasanein[14], Araki[15], and Kuhn[16]. However the scattering of each

experimental data is large and they give different experimental results one another. It is due

to the different experimental objectives and operating ranges. As an example two typical

experimental works were compared, which are performed by Siddique and by Kuhn. Kuhn's

experiment covered lower inlet noncondensable gas mass fractions and higher steam flow rates

than Siddique's, as shown in Table 2.1.

Differences between two experiments also include steam properties, air mass fraction,

Reynolds number, etc. Siddique[12] performed experiments at relatively lower steam flow

rates and Kut[16] did at relatively higher steam flow rates. As a result, two experiments

provided a quite a different results of the heat transfer coefficient.

The present experiments are performed at the steam flow rates of 8 to 40kg/hr. The first

objective of the present experiment is to get more reliable data within all operating ranges,

and the second one is to understand the effects of the various parameters on the condensation

heat transfer in the presence of noncondensable gas in a vertical tube.

5
Table 2.1: Comparison of experimental operating ranges between Siddique's and

Kuhn's

Parameters Siddique[MIT] Kuhn[UCB]

Inlet mass 10-35 0 - 40 for Air

fraction(%) 2 - 10 0 - 15 for Helium

Steam flow(kg/hr) 8 - 32 30 - 60

Vapor super-heat none 0 - 29.3 for Air

(00) 0 - 12.5 for Helium

Mixture T.("C) 100, 120, 140 super-heated

Mixture pres.(kpa) saturated 114.3 - 517.4 for Air

388.0 - 433.0 for Helium

Inlet mixture 5000 - 22700 13000 - 45600 for Air

Reynolds number 5000 - 11400 13100 - 31400 for Helium

Heat transfer 100- 25000 500 - 8000

coefficients high variation low variation

(W/m 2 .o c) sensitive to Remix insensitive

Steam supply four 7kW heaters UCB steam supply system

Coolant water T. center of the annulus inner and outer surface

air-vortex induced turbulence k - e model

Mass fraction high variation low variation

2.2 Test Facility and Its Instrumentation

To meet the above objectives the present experimental apparatus is newly designed to simulate the condensation phenomena of steam-noncondensable gas mixture in a vertical condensing tube.

The followings are considered to design a peculiar experimental apparatus.

1. Since the steam-air mixture velocity should be fully developed in the annulus to calculate the heat flux accurately, a honey comb is used to give the uniform initial velocity

in the inlet of the coolant jacket.

2. The exit of the coolant jacket is designed to reduce the flow resistance and to minimize

multi-dimensional effect.

6
3. Two suface temperatures in the coolant annulus are measured at every axially local

positions, and the coolant bulk temperatures are calculated using an unique calculation

methodology.

4. The local heat flux is directly measured with a RdF heat flux sensor at a certain location,

and it is compared with the calculated heat flux from the temperature gradient of the

coolant bulk temperatures.

The schematic diagram of the experimental facility is shown in Figure 2.1. The experimental facility consists of a steam tank including a 100kW heater, a steam-noncondensable

gas mixture supply line, a test section with a condensing tube and its surrounding coolant

jacket, a lower plenum, venting and draining systems, and a unit of data acquisition system.

The steam tank functions as a containment. Steam is generated from the steam tank and

is mixed with the supplied noncondensable gas in the steam tank. The noncondensable gas

is supplied into the steam tank continuously to give a continuous noncondensable gas mass

fraction at the inlet of the test section, or to keep a total air inventory in the tank by the

noncondensable gas supply line. The steam-noncondensable gas mixture flow is made to keep

uniform velocity distribution, and its temperature and pressure are measured in the inlet of

the test section.

The test section consists of an inner condensing tube and an outer coolant jacket. The

inner tube of the heat exchanger is a stainless steel pipe of 50.8mm in outer diameter,

1.5mm in thickness, and 2400mm in length. At 12 different axial locations each J-type

thermocouple is welded on the outer surface of the condensing tube to measure the outer

surface temperatures, welded through the condensing tube to measure the mixture bulk

temperatures, and installed on the outer wall of the coolant annulus to measure the adiabatic

wall temperatures. These three kinds of temperature sensor are located irregularly so that the

total heat transfer is kept to be same in the interval between two close temperature measuring

probes. The cooling water flows upward through the annulus and the steam-noncondensable

gas mixture flows downward in a condensing tube. The cooling water is supplied to the lower

7
Droplet

removal

bfixtre flowrate

mixtur

119z=c

146cm

175cm

208-Srn tem

+ Air

Steam

tank

guage,

>IOOkW~

Coolant water

Drain line

Figure 2.1: A schematic diagram of KAIST condensation experimental

apparatus

8
end of the jacket pipe in an once-through mode by a calibrated rotameter and a flow control

valve, and is dumped to the drain after the increase of its temperature.

Table 2.2 summarizes the instrumentation of KAIST condensation experimental apparatus and their uncertainties.

Table 2.2: Instrumentation of KAIST condensation experimental apparatus

Parts Measuring Model Measuring Units Uncerparameters number ranges tainty

Steam Pressure ABB PT-624 0 - 8.0 kgf/cm'7 0.1%

tank Temperature OMEGA KMQSS max. 1000 IC 0.5%

Air Flow rate Dwyer RMC-103 20 - 200 SCFH 2%

source Pressure OMEGA PX-425 0- 100 psig 0.2%

T/S : Pressure OMEGA PX-425 0- 100 psig 0.2%

inlet Temperature OMEGA JMQSS max. 500 *C 0.5%

Flow rate OMEGA FV-510B 0 - 500 lb/hr 1.5%

T/S : Mixture bulk T. OMEGA JMQSS max. 500 0C 0.5%

tube Outer wall T. OMEGA JMQSS max. 500 0C 0.5%

Heat flux RdF 20453-1 0 - 50 Btu/ft2s 5%

T/S : Temperature OMEGA JMQSS max. 500 OC 0.5%

coolant Flow rate Dwyer RMC-142 0.2 - 2.2 gpm 2%

Lower Pressure OMEGA PX-425 0 - 100 psig 0.2%

plenum Temperature OMEGA JMQSS max. 500 0C 0.5%

Drain Pressure Foxboro E11AH 0 - 750 kpa 0.5%

plenum Temperature OMEGA JMQSS max. 500 OC 0.5%

Venting Pressure OMEGA PX-425 0- 100 psig 0.2%

plenum Temperature OMEGA JMQSS max. 500 0C 0.5%

Flow rate Dwyer VFB-53 10 - 100 SCFH 3%

All the signals from detectors are current outputs or voltage outputs including outputs

from thermocouples. 4 -, 2OmA current signals from the pressure transducer and the vortex

flow meter are converted to 1 - 5V voltage signals using an multi-channel current-to-voltage

converter. The temperature signals are compensated to give accurate temperature data. All

signals are applied to the Hewlett Packard 44708H high voltage relay multiplexer to give

a reliable temperature and voltage data. Sampled experimental data is acquired using the

9
HP3852A control unit and the data is transferred to IBM-PC/AT using the RS232C interface.

The test facility and its instrumentation are described in detail in Appendix A.

2.3 Test Matrix and Experimental Procedure

As shown in Table 2.3, total 19 sub-tests were performed varying the following input parameters: the saturated steam temperature at the inlet, Ti, the inlet air mass fraction, AMF, its

inlet total pressure, Pt&, the inlet steam flow rate, SF, and the inlet air flow rate, AF. These

Table 2.3: Test matrix of KAIST condensation experiment

I.D. Ti. AMF Ptot SF AF Exp.

(°C) (kpa) (kg/h) (kg/h) No.

T1A2a 110.4 0.204 168.4 22.0 5.5 E13c

T1A3a 110.7 0.297 185.4 7.6 3.2 E9b

T1A3b 110.5 0.303 185.4 18.2 7.8 El3b

T1A4a 110.8 0.395 207.3 7.7 4.9 E9a

T1A4b 110.5 0.407 208.2 11.7 7.9 El3a

T2Ala 120.5 0.103 216.6 25.7 2.9 Ellf

T2A2a 121.4 0.195 239.0 14.8 3.6 E4a

T2A2b 121.4 0.200 239.9 21.3 5.2 Elld

T2A2c 120.8 0.196 234.4 25.8 6.2 Elle

T2A3a 120.9 0.306 260.8 9.8 4.3 Ellb

T2A3b 120.4 0.297 254.9 15.2 6.4 E7a

T2A3c 120.5 0.296 254.5 16.5 6.8 El1c

T2A4a 120.9 0.408 292.0 9.5 6.4 Ella

T3Ala 129.0 0.102 281.3 32.8 3.7 E4d

T4A1a 137.7 0.105 363.6 40.0 4.6 E8e

T4A2a 143.4 0.215 465.5 32.7 8.8 E12b

T4A3a 140.4 0.301 463.2 18.4 7.8 E12a

T4A3b 139.6 0.307 456.2 19.9 8.7 E8b

T4A4a 140.6 0.363 498.4 18.8 10.5 E8a

experiments span the ranges of conditions expected for the PCCS design. These sub-tests

can be grouped into four cases by inlet saturated steam temperatures: T1 series(110 0C); T2

series(1200C); T3 series(130 0C); and T4 series(1401C). In each group, the effect of inlet air

10
mass fraction and inlet steam-air mixture flow rate were investigated.

The steam-noncondensable gas mixture in the test section is in a saturated state, and the

system pressure is the sum of the saturated steam pressure and the partial pressure of the

noncondensable gas.

The test procedures are as follows:

1. At the beginning, the steam tank is isolated from the rest of the system.

2. The water level of the steam tank is adjusted and recorded.

3. The noncondensable gas is removed from, or supplied to the steam tank to meet a

predetermined initial condition.

4. The heater is turned on and the water is boiled to reach the required initial temperature

and pressure from which the initial air mass fraction can be calculated.

5. The heater is controlled to generate the steam to give a predetermined pressure.

6. The coolant flow rate is adjusted to the determined value.

7. The air is injected into the tank to keep a constant initial condition.

8. After the heater power is readjusted, the accumulated air is vented and the condensate

is drained to stabilize the system pressure.

9. Data are acquired from sensors and transducers after the steady state is achieved.

10. The local heat transfer coefficients and the other reduced data are calculated from the

raw data.

2.4 Experimental Results and Discussions

As steam-noncondensable gas mixture enters into a vertical condensing tube of the PCCS, as

shown in figure 2.2, steam begins to condense at the inlet of the tube. The steam condenses on

11
Steam-air mixture

Stainless

steel Air P.

tube tube .= • ~Temperature : ;i

Steam-air mixture .. •.

boundary lay

Condensate !)i!i

Figure 2.2: Condensation phenomena in a vertical tube in the presence of

noncondensable gas

12
the inner wall of the tube and the condensed film flows as an annular film along the condensing

tube. A gas-vapor boundary layer forms next to the condensate interface, through which the

water vapor must pass by diffusion and convection, and it thickens between the condensate

film layer and the steam-noncondensable gas mixture layer. The partial pressures of gas and

vapor vary through the boundary layer. At a certain axial location, the boundary layer grows

and blocks the tube and there is no longer a central core.

Generally both the local air mass fraction and the local condensate flow rate increases,

but the local steam-air mixture flow rate decreases along the condensing tube.

Temperatures of the coolant, the outer wall of the tube, and the steam-air mixture bulk

are measured. The coolant bulk temperature is calculated by the unique numerical calculation

method using two suface temperatures and the coolant flow rate in the coolant annulus. The

local heat flux is calculated from the profile of the coolant bulk temperature, and the inner

wall temperature of the test section is calculated from the outer wall temperature of the

condensing tube and the calculated heat flux. The local heat flux at 1.75m apart from the

tube inlet is directly measured with the RdF heat flux sensor to cross-check the calculated

heat flux.

The data reduction method is described in detail in Appendix B.

Figure 2.3 shows the local temperature distributions of the steam-air mixture bulk, the

inner surface, the outer surface, the coolant adiabatic, and the coolant bulk along the test

section in experiment E13b with the inlet saturated steam temperature of 140*C and the

inlet air mass fraction of 20%. All temperatures decrease smoothly along the test section,

except for the inversion of two wall temperatures.

The inversion of the outer wall temperature occurs in the middle of the condensing tube.

One anticipated reason for such a temperature inversion is from the mechanical error by

wrong silver soldering, and another reason is the secondary flow of the coolant flow in an

annulus jacket, which is originated from Hasanein[14]. To determine whether the temperature

inversion is caused by the secondary flow or not, the radial temperature distribution is planned

to check. The coolant temperatures are measured with movable thermocouples for several

13
160 Experiment E13B : Temperatures 160. - -- Mixture bulk,

.. 4 -.. Coolant bulk, -.o.- Coolant adiabatic, 140-

A Inner wall. , Outer wall.

120.......... Inner wall(fit), Outer wall(fit) 0 10

(D 100

CIS 80- A.

40-

0120.

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 2.3: Local temperature distributions in experiment E13b

radial points at 5 axial locations, which are near the position of the temperature inversion.

The measured outer wall temperatures are fitted to give a smooth slope along the test section,

and the inner wall temperatures are calculated again.

As shown in Figure 2.4, generally the local heat flux decreases rapidly as the local air

mass fraction increases and the local mixture flow rate decreases along the condensing tube.

The local heat flux at 1.75m apart from the tube inlet is directly measured with the RdF

heat flux sensor, and it gives the heat flux similar to the calculated one.

Some sub-tests are compared to investigate the parametric effects of the condensation

experiments.

2.4.1 Effects of inlet steam-air mixture flow rate

The comparisons are performed at three different inlet steam flow rates of 15, 21, and 26kg/hr

for a condition of air mass fraction of 20% and a saturated steam temperature of 120*C.

Figure 2.5 and 2.6 show comparisons of both experimental heat fluxes and heat transfer

coefficients with different inlet steam-air mixture flow rate, respectively.

14
,uuuuU

60000, Experiment E13B : Heat flux

-o- Calculated heat flux

50000- ' Measured heat flux

.with RdF heat flux sensor

• 40000-

- 30000

20000-

10000-

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 2.4: Local heat flux distribution in experiment E13b

As shown in Figure 2.5, with higher inlet steam-air mixture flow rate, the local heat

flux decreases more smoothly in the inlet, and it always keeps higher values throughout the

condensing tube than that with lower one. It is due to the fact that it keeps an higher local

mixture Reynolds number along the condensing tube with a high inlet mixture flow rate.

In the inlet of the test section, it gives similar value regardless of the different initial

mixture flow rate. As the condensate liquid film just begins to develop and the mixture

boundary layer is not developed, the mixture is almost directly in contact with the outer wall

of the condensing tube in that region. It is considered that heat flux depends only on the air

mass fraction and the saturated steam temperature in that region.

Since more steam is removed from the steam-air mixture flow with higher inlet mixture

flow throughout the condensing tube, the local steam flow rate is decreased, and the local

air mass fraction is increased rapidly along the condensing tube. Thus the heat flux becomes

similar also in the outlet of the test section.

As shown in Figure 2.6, the heat transfer coefficient shows the similar tendencies to the

heat flux. It is shown to be similar regardless of the inlet mixture flow rate both in the inlet

15
Cu

Cu

z

100000-

80000.

60000.

40000-

20000.

I-1--E4a. I5korI

~.E 110.21 kq/hr

E116,26kg/hr

0.0 0.o 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 2.5: Comparison of experimental heat fluxes with the variation of the

inlet steam-air mixture flow rate

4000- E 00 1 ;20*C. 20% air "

4-1- E4&, 15kglhr

S--E--E11d. 21 kg/hr

3000- -A.-El, 2II/r

o 2000

1000-

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 2.6: Comparison of experimental heat transfer coefficients with the

variation of the inlet steam-air mixture flow rate

16
and in the outlet of the condensing tube, but it always keeps higher values with higher inlet

mixture flow rate throughout the condensing tube.

Figure 2.7 shows the axial temperature distributions of the steam-air mixture bulk, the

inner surface, the outer surface, the coolant adiabatic, and the coolant bulk along the test

section for experiments E4a, E11d, and Elle with different initial steam-air mixture flow

rate. All the five temperatures show smoother profiles with a higher steam-air mixture flow

rate than with a lower one.

The air contained inside the test section must be vented to initiate the condensation

of the steam in the presence of air, to maintain a constant pressure, and to prevent the

accumulation of the air throughout the test section. As the steam-air mixture penetrates

more deeply, higher mixture bulk temperature is kept throughout the test section, which is

the main cause of the smoother temperature distribution and the higher heat transfer.

2.4.2 Effects of inlet air mass fraction

The comparisons are performed at 2 different values of the inlet air mass fractions of 20, and

30% for a fixed inlet saturated steam temperature of 1101C and inlet mixture flow rate of

about 20kg/hr. Figure 2.8 and 2.9 show comparisons of both experimental heat fluxes and

heat transfer coefficients with different inlet air mass fraction, respectively.

As shown in Figure 2.8, the local heat flux is much higher in the inlet of the test section

with lower inlet air mass fraction. It decreases more rapidly throughout the condensing tube

than that with higher inlet air mass fraction, and as a result it becomes similar in the outlet

of the condensing tube. The local air mass fraction is always kept low with low inlet air mass

fraction along the condensing tube. With lower air mass fraction, more steam is removed from

the steam-air mixture flow throughout the condensing tube. As the local mixture flow rate

is decreased and the local air mass fraction is increased, the heat transfer by condensation is

reduced and the heat transfer by convection of mixture becomes dominant in the outlet of

the test section.

As shown in Figure 2.9, the heat transfer coefficient shows the similar tendencies to the

17
9

ox

0)

c0

a)

0.

E

a)

I160-

140-

120-

100.

so40.

20.

Exponmnen E4a: 15kg/r

-3- Mixute bulk

---- 1- Inner wall

-A- Outer wall

-V- Coolant bul m

"A'• LL• -- Coolant acn'"ati .. ,__,__

CD

a)

CL

0.

CL

S

a)

160

140

120

100.

80

60

40

20

0

160

140

120.

100.

80.

so -

60.

40.

20.

.

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

experiment El Id: 21kg/hr

-3- Mixure bulk

-- - Inner wall

3- m -A-,-Outer well

--- Coolant bulk

-I +CooLant adiabatic

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Experiment Elw a26kg/hr

-.- Mixure bulk

-,-Inner wall

-A- Outer wall

-V- Coolant bulktI

A"A 1 --- Coolant adiabati

V I

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 2.7: Comparison of experimental temperature distributions with the

variation of the inlet steam-air mixture flow rate at the inlet mixture

temperature of 1201C with 20% air

18
E

XS60000-

40000-

20000

0 .

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 2.8: Comparison of experimental heat fluxes with the variation of the

inlet air mass fraction

5000 -

~.E E 4000 110"C, 20kg/hr

3: --U-113b. 30% air

3 

-4--El3c. 20% air

"G3000-

C)

0

0

2000

000

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 2.9: Comparison of experimental heat transfer coefficients with the

variation of the inlet air mass fraction

19
heat flux. It is shown to be much higher in the inlet, and always keep higher value throughout

the condensing tube with higher inlet air mass fraction, and it becomes similar in the outlet

of the condensing tube regardless of the inlet air mass fraction.

2.4.3 Effects of inlet saturated steam temperature

The comparisons are performed at 2 different values of the inlet saturated steam temperatures

of 110 and 140°C for a fixed air mass fraction of 30% and a steam flow rate of about 18kg/hr.

Figure 2.10 and 2.11 show comparisons of both experimental heat fluxes and heat transfer

coefficients with different inlet saturated steam temperature, respectively.

As shown in Figure 2.10, the local heat flux is always much higher with the increase of the

inlet saturated steam temperature in the inlet of the test section, but the difference between

heat fluxes of different inlet saturated steam temperature is negligible in the outlet of the

test section.

However, as shown in Figure 2.11, the local heat transfer coefficient decreases more rapidly

with higher inlet saturated steam temperature. It always keeps lower value throughout the

condensing tube than that with low saturated steam temperature, except for the inlet of the

test section, where it gives similar values regardless of the different inlet saturated steam

temperature. The higher wall sub-cooling permits higher heat flux due to the higher thermal

driving force, or the temperature difference between the mixture bulk and the inner wall.

However, when the heat flux is divided by the temperature difference to give the heat transfer

coefficient, the local heat transfer coefficient is always lower with high inlet saturated steam

temperature than that with lower one.

20
E

X

100000.

80000.

60000.

40000.

-a-E13b,11 001

-L-E2al 40*C

2000:1-

0.0 0.o 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 2.10: Comparison of experimental heat fluxes with the variation of the

inlet saturated steam temperature

zOVU

Ap

SE

C;

2000.

1500.

1000.

500.

0.

1kqh.3%air.

j-u-M- 13b.1 iO'C

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube Inlet (m)

Figure 2.11: Comparison of experimental heat transfer coefficients with the

variation of the inlet saturated steam temperature

21
Chapter 3

Assessment of RELAP5/MOD3.2

3.1 Two Wall Film Condensation Models of RELAP5/MOD3.2

A standard RELAP5/MOD3.2 was used for the present calculation. Related with the calculation for the condensation experiment with noncondensable gas in a vertical condensing tube,

the RELAP5/MOD3.2 code was known to have more improved models than the previous

version, RELAP5/MOD3.1, in wall condensation model and transport of non-condensable

gas[19].

The condensation of steam-noncondensable gas mixture in a vertical condensing tube

can be simulated with two wall film condensation models in RELAP5/MOD3.2[18]. For an

inclined surface the Nusselt-Shah-Colburn-Hougen correlations are used as a default model,

and the Nusselt correlation with UCB multipliers is used as an alternative model. For the

horizontal case the Nusselt correlation is replaced by the Chato correlation both in the default

and alternative models.

3.1.1 Default condensation model

The default model is to use the maximum of Nusselt[20]'s and Shah[21]'s with the ColburnHougen[22]'s diffusion calculation when noncondensable gases are present. The Nusselt[20]'s

22
expression for the vertical surface uses the film thickness.

hf = k~f (3.1)

where the film thickness, 6, is

6 = [ 3pf ] 1131.~ 11/3 (3.2) Lgp- ApJ L4gpPfAp ,

where r is the mass flux per unit width and Ref the liquid film Reynolds number.

Chato[23] developed a modification to the Nusselt formulation which applies to the laminar condensation on the inside of a horizontal tube. The correlation takes the form

r gpfAphfgbkf 1/4

F DhItf (Tppb - TWI)] (3.3)

where k! is the liquid conductivity, pj the liquid viscosity, p! the liquid density, Ap the

difference in densities between the liquid and the vapor, g the gravitational constant, hf9 b

the steam saturation enthalpy at steam partial pressure minus liquid saturation enthalpy in

the bulk, and Tppb the saturation temperature based on steam partial pressure in the bulk.

The F is the term which corrects for the liquid level in the tube bottom and a value of 0.296

is recommended for free flow from a horizontal tube.

The Shah[21]'s model is used for the modeling of film condensation with turbulent flow

as follows;

h3 = h8 (1 + 3-8 (3.4)

where

Z X Pitdq (3.5)

and X is the static vapor quality, Pred the reduced bulk pressure, P/Pcr.tcal, and h,!f the

23
superficial heat transfer coefficient, which is given by

hsf = hi(1 - X)0

°8 , (3.6)

where hi is the Dittus-Boelter coefficient assuming all fluid is liquid, given by

h, = 0.023(±-)Re?" Pr?", (3.7)

where the Reynolds number is given by Rej = Gtot.,Dh/plf, Gtot the total mass flux, Dh

the hydraulic diameter, 14f the liquid viscosity, and Pr1 the liquid Prandtl number. The Shah

correlation is based on the database of both horizontal and vertical data, and it is activated

when its heat transfer coefficient becomes larger than that of Nusselt correlation or that of

Chato correlation.

The formulation of the Colburn-Hougen model is based on the principle that the amount

of heat transferred by condensing vapor to the liquid-vapor interface by diffusing through the

steam-noncondensable gas mixture boundary layer is equal to the heat transferred through

the condensate. The heat flux due to vapor mass flux is

it,= [1- Phi/P, qV= hmhfgbPvbln [1 --P-P'j, (3.8)

where P is the total pressure, Pb the steam partial pressure in the bulk, PFi the partial

pressure of steam at liquid-gas-vapor interface, PO the saturation vapor density at PFb and

hm the mass transfer coefficient. The mass transfer coefficient, h,,, depends on the flow

condition. When the vapor flow is turbulent, the Gilliand correlation[24] is used, and when

laminar, the Rohsenow-Choi correlation[24] is used. The heat heat flux from the liquid to

the wall is calculated by

tq, = h, (Tt, i - T.), (3.9)

where Tj, is the saturation temperature corresponding to the interface vapor pressure, T,

the wall temperature, and h, the predicted condensation heat transfer coefficient. From the

24
energy balance the partial vapor pressure at the interface and its corresponding temperature

are determined by iteration.

hc Tvi-T) = hmhfgbPt~bll [1 11- - P, P2'bPJ" 1/P 1.(3.10)

3.1.2 Alternative condensation model

The alternative model is the Nusselt model with UCB multipliers, which is revised to include

the effects of the interfacial shear and the presence of the noncondensable gas in a vertical

tube as follows:

hUCB" = fl- f2, (3.11)

where fi and f2 are formulated from the curve fits to the experimental data as follows:

l = 1 + 2.88 x 10 5Rel-, (3.12)

where the steam-noncondensable gas mixture Reynolds number is given by Remix = GgDh/pg,

where G9 is the gas bulk mass flux, Dh the hydraulic diameter, and pA the gas bulk viscosity.

1 - 10Mg for Ma < 0.063,

12 = 1 - 0.938Mg- 13 for 0.063 < M= < 0.6, (3.13)

1 - MaO'. for M. > 0.6,

where M. is the fraction of noncondensable gas in the vapor-gas mixture. In this alternative

model the Nusselt correlation and the UCB fi factor is used instead of the maximum of the

Shah and the Nusselt correlation in the default model, and the UCB f2 factor instead of the

Colburn-Hougen diffusion method. The enhancement factor, fl, accounts for the effects of

the shear of the steam-noncondensable gas mixture on the liquid film, and the degradation

factor, f2, accounts for the effects of the noncondensable gas on the heat transfer coefficient.

The maximum value allowed for fi is 2.0 to prevent over-predicting in the high shear region

such as in the entrance region. For a horizontal tube with laminar flow the Nusselt correlation

25
is also replaced by the Chato correlation.

3.2 Direct Assessment of Two Wall Film Condensation Models

The condensation models in RELAP5/MOD3.2 require various local properties to give condensation heat transfer rates. The condensation heat transfer database is constructed by

collecting the local data in literature or by calculating the local parameters from the correlations published[18]. It is done on the basis of local values to ease handling of the database

and assessment with correlations. The detailed information of the constructed database is

given in Table 3.1.

Table 3.1: Heat transfer database for the vertical laminar film condensation

Author Data [Run Pressure fSteam flow NC g as

I # (MPa) (kg/h) Non air fraction

Vierow[9] 297 36 0.03 -- 0.45 5.9 -24.95 air 0 - 0.14

319 42 0.109 ,- 0.518 28.3 -61.5 none 0

Siddique[121 416 52 0.107 ,- 0.485 7.9 - 31.9 air 0.08- 0.42

159 22 0.114 - 0.466 8.6 -20.5 He 0.02-,, 0.11

291 44 0.101, 0.27 10 ,- 40 He 0.025 -0.2

Hasanein[14] 600 76 0.101, 0.27 10 -40 air 0.05- 0.2

1 1 (He) (0.025)',, (0.15)

Kuhn[16] 627 71 0.114 -- 0.517 29.5 -, 61.0 air 0.01,- 0.4

192 24 0.388 - 0.433 29.5 61.9 He 0.003- 0.15

Since the constructed database can give all local values which are required by the two wall

film condensation models in RELAP5/MOD3.2 code, those models can be assessed directly,

or the heat transfer coefficients calculated directly with two condensation models themselves

are compared with those from the constructed data base. Therefore, it does not cause any

systematic effect which might occur from the internal code calculation. Both the default and

the alternative models used for the laminar film condensation heat transfer inside a vertical

26
condensing tube are assessed.

In the case of the default model, its prediction is compared with the Hasanein[14]'s experimental data to assess the Colburn-Hougen diffusion method[22].

C.,

0

CL 401 0.

,,J

M

Jo 301

201

o 101 I0 1000 2000 3000 4000

HTC(W/m 2

-K) from Experiment

5000 600

Figure 3.1: Assessment of default model of RELAP5/MOD3.2

For the Hasanein's data[14] the RELAP5/MOD3.2 model under-predicts the experimental

data for the low heat transfer coefficient range, but it over-predicts the experimental data

for the high heat transfer coefficient range as shown in Figure 3.1. As the heat transfer

coefficients calculated from the default model show much large scattering, a modification

of the Colburn-Hougen diffusion method is needed for a better prediction of laminar film

condensation with noncondensable gas.

In the case of the alternative model, its predictions are compared with 3 kind of experiments for pure steam, steam-air, and steam-helium condensation. Figure 3.2 shows the

27
C4 ZUUUU

CI

0

S15000-

E

75 10000-

RELAP5/MOD3.2

"E j3 Alternative model

5000- (rms error: 64.8%)

z - Default model

(rms error: 17.2%)

. 0 1- . - " 

0 5000 10000 15000 20000 250(

HTC(W/m2-K) from experiment

Figure 3.2: Assessment of two wall film condensation models of

RELAP5/M 'D3.2 for pure steam

assessment results for the alternative model of RELAP5/MOD3.2 with Kuhn's[16] experimental data for pure steam. Since the effect of the shear stress of steam flow was included in

Kuhn's data to increase the heat transfer coefficient, the prediction of the Nusselt correlation

shows a little smaller than Kuhn's experimental data. The predicted value of the alternative

model of RELAP5/MOD3.2 is always higher than that of the Nusselt correlation by almost 2

times, and it gives higher values compared with the Kuhn's experimental data. The effect of

shear stress by the steam flow is overestimated. For most of the experimental data compared,

the enhancement factors, fl, have the limit values of 2. As Kuhn's experiments are performed

at high mixture Reynolds numbers, its data exceed the operating range of UCB correlation

in the alternative model of RELAP5/MOD3.2.

28
2.0

Vi 1.5 ++:6 +A

o +

ot 1.o LU +

x f2 factor :Effects of noncondensable gas I

ini

A Ill factor •Effects of shear (enhancement)

+- MF. -f1"2 - Multiplication factor

0.0

0.0 0.5 1.0 1.5 2.0

Multiplication factor of Kuhn's experiment

Figure 3.3: Multiplication factors in alternative model of RELAP5/MOD3.2

Comparison of multiplication factors in the alternative model of RELAPS/MOD3.2 with

the Kuhn's air-steam experimental data is shown in Figure 3.3. The enhancement factor, fl,

the degradation factor, f2, and the overall multiplication factor, /,are compared. Similar

to the assessment results for the pure steam condensation, the/ factor converges to the

limit value of 2 for the experiment with noncondensable gases. The degradation factor f2 is

also calculated using equation 3.13 with Kuhn[16]'s raw data. f, and f2 is multiplied to give

the overall multiplication factor, f. For the low heat transfer coefficient range with high air

mass fraction and low shear stress the overall multiplication factors of both Kuhm's[161 and

RELAP5/MOD3.2 correlation give similar results as the overestimation of fj is compensated

by the underestimation f2. However, the multiplication factor f is overestimated by the

alternative model for a range to produce the high heat transfer coefficients.

29
cs 0 Vierow (UCB)

S15000- (rms error: 117.9%) CL

U,

cc

E a

*£ ioo

0[]-d~ oDO 

(3 0 3 P0 3 o 0D

z i •0 0

S5000- 3 0 1

030

0 5000 10000 15000 2000

HTC(W/m2 -K) from Vierow's experiment

Figure 3.4: Assessment of alternative model of RELAP5/MOD3.2 with

steam-air experimental data

The UCB multiplication factors in the alternative model are developed based on the

Vierow's experimental data[9]. As Vierow's experimental data used for the development of

the UCB multiplier in the alternative model show great scattering as shown in Figure 3.4, it

is desirable to develop a better correlation with a larger data pool.

Figure 3.5 shows the assessment results for the alternative model of RELAP5/MOD3.2

with the steam-helium experimental data by Hasanein[14] and Kuhn[16]. As the enhancement

factor, fl, is always overestimated, the model predicts higher values than both Hasanein's

and Kuhn's experimental data.The difference between Kuhn's data and Hasanein's data is

from the difference between their experimental conditions. As Hasanein's experiments are

performed over a range of low mixture Reynolds number compared with the Kuhn's exper30
15000

A

0 1000 A A A

Result of Relap correlation

AA (Nusselt/Chato-UCB model)

A :Kuhn(UCs)

(rms error: 94.9%)

0 : Hasanein (MIT)

0 (rms error: 101.8%)

0

0 5000 10000 15000

HTC(W/m.-K) from Experiment

Figure 3.5: Assessment of alternative model of RELAP5/MOD3.2 with

steam-helium experimental data

imental data, most of his data have low heat transfer coefficients and its correlation gives

better prediction on the data with low mixture Reynolds number.

3.3 RELAP5/MOD3.2 Nodalization

The nodalization scheme of the RELAP5/MOD3.2 code of the present experimental facility

is shown in Figure 3.6.

The present RELAP5/MOD3.2 nodalization used for this simulation contains 41 control

volumes, 6 junctions, a valve and a heat structure.

Time-dependent volumes acting as infinite sources or sinks are used to represent boundary

conditions both for the steam-noncondensable gas mixture flow in a condensing tube and for

31
tdv28O

b120

pipel4O

hx14O

an240

tdv2OO

tdj2IO

sj156

pipe157

Figure 3.6: Nodalization scheme of the RELAP5/MOD3.2 code for PCCS

experimental facility

32
the coolant flow in a coolant jacket. For the simulation of the coolant jacket, two time

dependent volumes 200 and 280 are connected to the annulus 240 with 11 volumes via a

time dependent juction 210 and a single junction 270. Similarly, for the simulation of the

steam-noncondensable gas mixture flow, two time dependent volumes 100 and 180, a pipe

with 13 volumes, a time dependent juction 105 and a single juntion 151 are also used.

A branch 120 is used to simulate an upper plenum and three pipe volumes 150, 160 and

157 are used to simulate a lower plenum, a drain tank and a connecting pipe between the

lower plenum and the drain tank, respectively. The above three pipes are connected using

single junctions 155, 156 and 158. A valve 175 is used to regulate the venting of the mixture

of the residual steam and the noncondensable gas.

A heat structure 140 with 11 volumes is used to represent the heat transferred from the

steam-noncondensable gas mixture to the coolant through the condensing tube.

3.4 Simulation Results and Discussion

3.4.1 Base case calculation

For base case calculations, five sub-tests are simulated by RELAP5/MOD3.2. As listed in

Table 3.2, the following 7 input parameters are varied: the pressure at the inlet of the test

section, Pi,; its temperature, T,; the inlet steam-air mixture flow rate, MF; the inlet air

mass fraction, AMF; the temperature at the outlet of the test section, Tout; the temperature

at the inlet of the coolant, T,,im; the coolant flow rate, CF.

Steady state calculations were performed to determine whether or not it can describe

properly the steam condensation experiments in the presence of noncondensables in a vertical

tube of PCCS. These simulations used both the default model and the alternative model of

the RELAP5/MOD3.2 code to be compared each other.

33
Table 3.2: Input parameters for base case calculations

Exp. Pi. Ti. MF AMF Tot Tc,in CF

No. (kpa) (°C) (kg/h) (%) (1C) (°C) (gpm)

4d 281.3 129.0 32.8 10.2 62.3 16.4 2.4

l1d 239.9 121.4 21.3 20.0 81.1 28.9 2.4

12a 463.2 140.4 18.4 30.1 66.6 28.7 2.4

12b 465.5 143.4 32.7 21.5 81.6 29.3 2.4

13b 185.4 110.5 18.2 30.3 81.7 28.4 2.4

Comparison of two simulation results with experimental data

Two simulation results of each steady state calculation were compared with each other, and

they were also compared with the experimental data. Two simulation results with different condensation models were compared with the experimental data of E13b in Figures 3.7

through 3.12.

The following main characteristics are found:

3.0

2.5-

2.0-

1.5-

o 1.0- 0.

0.5-

111--Deaultmodel1

[ 2 Alternative model

0.0 0.5 1.0 1.5 2.0

Lenath from the tube inlet (m)

2.5

Figure 3.7: Comparison of steam velocities calculated using two condensation

models with the experimental data of E13b

34
* As shown in Figure 3.7, the calculated steam velocity from the default model is always

higher than that from the alternative model. They begin to condense with the same

inlet velocities at the inlet, but the local steam velocity calculated from the default

model decreases more slowly than that from the alternative model. As the amount

of the mass transferred to the liquid film by the steam condensation is large for the

simulation using the alternative model, the steam flow rate is calculated to be smaller

for the alternative model than for the default model.

1.0

0.8 -A "A

0

Z 0.6

E 0.4

-i-- Default model I 0.2. -

1 --- Alternative model

- -- Experimental data

0.0

0.0 0.5 1.0 1.5 2.0 2.5

Lenoth from the tube inlet (m)

Figure 3.8: Comparison of air mass fractions calculated using two condensation

models with the experimental data of E13b

" As shown in Figure 3.8, the calculated air mass fraction from the default model is always

lower than that from the alternative model. The calculated air mass fraction from the

default model increases linearly and is always slightly higher than the experimental

data. However the calculated air mass fraction from the alternative model increases

rapidly in the upper part of the condensing tube and the inclination is decreased in the

lower part, it always keeps higher value than the experimental data except for the inlet.

" Figure 3.9 shows the distributions of the mixture bulk, the saturated steam, the inner

35
120,

100-

(,.)

* A. 00 80

-X-T.W1

60

E 40- 40 - !-rl-Tmk

-0-0w

-&-u

20. -,o T0-

0.0 0.5 1.0 1.5 2.0 2.5

Lenoth from the tube inlet (m)

Figure 3.9: Comparison of temperatures calculated using two condensation

models with the experimental data of El3b

and the outer tube wall and the coolant bulk temperatures along the condensing tube.

All the calculated temperatures from the default model decrease more slowly than those

from the alternative model. So the mixture bulk temperature from the default model

always keeps a higher value than that from the alternative model, and its coolant

bulk temperature always keeps a lower value. Most of the inner and the outer wall

temperatures from the default model keep lower values than those from the alternative

model except for the outlet of the condensing tube. The experimental mixture bulk

temperature is always similar to that from the default model. The experimental inner

and outer wall temperatures keep lower values than those from both the default model

and the alternative model, but they crossed with each other in the upper part of the

condensing tube. The experimental coolant bulk temperature goes between that from

the default model and that from the alternative model along the condensing tube.

* As shown in Figure 3.10, the calculated coolant-side heat transfer coefficient from the

default model is lower than that from the alternative model in the upper region of

36
• | |

p

E -I - Defautt model

1500- -- Alternative model

1000

0

50)

0,

0.0 0.5 1. 1.5 2.0 2.5

Lenath from the tube inlet (m)

Figure 3.10: Comparison of coolant side heat transfer coefficients calculated

using two condensation models with the experimental data of E13b

the condensing tube, but the tendency is opposite in the lower part of the condensing

tube. The change of the simulated results in the coolant side has little effect on the

condensation heat transfer characteristics in a condensing tube. Sometimes the rapid

increase of the simulated one from the alternative model is due to the change of the

heat transfer characteristics from the single phase convective mode to the two phase

boiling mode near the inlet of the condensing tube.

As shown in Figure 3.11, the calculated heat flux from the default model is always

lower than that from the alternative model in the upper region of the condensing tube,

and they are similar in the lower region. The condensation heat flux calculated from

the default model is always much lower than the experimental data throughout the

condensing tube. However, the condensation heat flux calculated from the alternative

model is a little higher than the experimental data in the upper region of the condensing

tube, and they crossed each other in the middle of the condensing tube.

•As shown in Figure 3.12, the calculated condensation heat transfer coefficient from the

37
I

60000, [•\• Default model

.-- Aternative model

5000, ••. -A--Experimental data

.~50000.dt

n 40000-

30000

IL

S20000-

10000-

0~~=

0o.0 0.5 1.0 1.5 2.0 2.5

Lenoth from the tube inlet (m)

Figure 3.11: Comparison of heat fluxes calculated using two condensation

models with the experimental data of E13b

default model is always lower than that from the alternative model throughout the

condensing tube. Similar to the comparison results of the heat flux, the calculated

heat transfer coefficient from the default model is always lower than the experimental

data, while the calculated one from the alternative model is always higher than the

experimental data. Three heat transfer coefficients, or two simulated results and one

experimental data, are greatly different in the inlet of the test section,. but they are

similar in the outlet of the condensing tube, where the amount of steam is greatly

reduced by condensation and the convective heat transfer is dominant.

Several simulation results for experiments with different initial condition

From experimental studies, it is known that the main parameters controlling the condensation

with noncondensables in a vertical tube are the inlet steam-air mixture flow rate, the inlet

air mass fraction, and the inlet saturated steam temperature. Simulations on experiments

are performed to show the parametric effects with both the default model and the alternative

38
E 5 --a- Default model 5000 I -- Attemariv. model

-Experimental data

*C 4000.

3000

0

CO 2000

1000

0

0 , . , . , I

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.12: Comparison of condensation heat transfer coefficients calculated

using two condensation models with the experimental data of £13b

model of RELAP5/MOD3.2, and they are compared with the experimental data.

The following main characteristics are found :

Figure 3.13 and 3.14 show simulations for experiments El2a and Eld, respectively.

Those two experiments have high inlet air mass fractions above 10% and low inlet

steam-air mixture flow rates below 30kg/hr. The simulated heat transfer coefficient

using the alternative model is always higher than that using the default model in the

upper part of the condensing tube, and they are approaching along the condensing tube

to give similar values in the lower part. The heat transfer coefficient calculated-from

the default model is always lower, but that from the alternative model is higher than

the experimental data.

" Figure 3.15 shows simulation for experiments El2b, which was performed for relatively

high inlet steam-air mixture flow rate above 30kg/hr. The calculated heat transfer

coefficient from the alternative model becomes similar to the experimental data both at

the inlet and the outlet of the condensing tube, but it is higher than the experimental

39
-11- Default model

E 4000 I---Afternative model

-Experiment dat

• 3000.

to

02000. A

0)

0

0 , .... A

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.13: Comparison of condensation heat transfer coefficients calculated

using two condensation models with the experimental data of El2a

I.-

0

A 6000.

E

5000.

* 4000.

Q

3000. 0

0• 2000.

1000

0.

0-I- Default model

SAlttemrntve model

-&-Experimental data

0.0 0.5 1.0 1.5 2.0

Length from the tube inlet (m)

2.5

Figure 3.14: Comparison of condensation heat transfer coefficients calculated

using two condensation models with the experimental data of Eld

40
data in the middle of the condensing tube. The heat transfer coefficient calculated

using the default model is lower than the experimental data in the inlet, and they are

approaching along the condensing tube to give similar values in the lower part. This

tendency of the default model is unchanged regardless of the change of the inlet mixture

flow rate.

6000.

E 5000- -4M- Default model

-0- Alternative model

--- Experimental data

4000-

2-.

3000- 0

S2000.

) 1000

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.15: Comparison of condensation heat transfer coefficients calculated

using two condensation models with the experimental data of E12b

e Figure 3.16 shows simulation for experiments E4d, which was performed for relatively

high inlet steam-air mixture flow rate above 30kg/hr and low inlet air mass fraction

below 10%. Similar to the case of high inlet mixture flow rate above 30kg/hr, the

calculated heat transfer coefficient from the alternative model becomes similar to the

experimental data both at the inlet and the outlet of the condensing tube, but it predicts

higher values in the middle of the condensing tube. This tendency of the alternative

model is unchanged regardless of the change of the inlet air mass fraction. The heat

transfer coefficient calculated using the default model simulates well the experimental

data throughout the condensing tube. 30kg/hr and low inlet air mass fraction below

41
-2- Default model

8000- -*-Alternative model

Experimental data

A? 8000- C.

a,

0

o 4000,

ccC2000

Cu

0.0 0.5 1 .0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.16: Comparison of condensation heat transfer coefficients calculated

using two condensation models with the experimental data of E4d

10%

9 When the inlet steam-air mixture flow rate is low below 30kg/hr and the inlet air mass

fraction is high above 10%, both the default model and the alternative model predict

wrong, and they are needed to be improved. With the increase of the mixture flow

rate, the default model predicts wrong except for the outlet of the condensing tube,

while the alternative model predicts correctly both in the inlet and outlet, but it also

over-predicts the heat transfer coefficient in the middle of the condensing tube. For

the condition of relatively low inlet air mass fraction below 10% and high inlet mixture

flow rate above 30kg/hr, the default model predicts well the experimental data. It is

considered to be that the effect of the inlet steam-air mixture flow rate is not properly

considered in the default model, and the effect of the inlet air mass fraction in the

alternative model.

42
3.4.2 Nodalization study

The RELAP5/MOD3.2 code permits the user to vary the nodalization. By changing the

number of nodes in the heat structure, it is possible to investigate whether or not the node

number affects the heat transfer characteristics in the RELAP5/MOD3.2 code. It is known

that the change of the number of the divided nodes has little influence on the simulation results of the condensation phenomena with RELAP5/MOD3.2. The sub-program conden.f is

based on the film thickness to make it a local form instead of the average value used in

RELAP5/MOD3 up to MOD3.1.1.1 version. There was much effort to eliminate the dependence on the node size in condensation heat transfer coefficient for an inclined surface of

RELAP5/MOD3.1[4].

To investigate this effect the test section tube is divided into 4, 8, 12, and 16 nodes

regularly instead of the irregular 11 nodes in the base case calculation of El2b. The simulation

is performed at the condition that the total pressure is 0.46547Mpa and the inlet saturated

steam temperature is 143.4 0C(416.55K). The calculated heat transfer coefficients vary almost

the same along the condensing tube for the different node numbers, as shown in Figure 3.17.

From the analysis of the simulated results it is concluded that the heat transfer characteristics

both in the default and in the alternative model of RELAP5/MOD3.2 are little affected by

the change of the node number in the heat structure.

3.4.3 Sensitivity study of input parameters

As listed in Table 3.3, input parameters are varied to compare the calculation results with

changes of input parameters. The experiment El2b is selected as a base case. The effects of

the coolant flow rate, the inlet coolant temperature, and the vented mixture temperature give

negligible effects on the heat transfer coefficient in the condensing tube except for the minor

variations due to the changes of the condition of the coolant. However, the inlet steam-air

mixture flow rate, the inlet air mass fraction and the inlet saturated steam temperature give

significant changes of the heat transfer coefficient in the condensing tube. In this section,

43
o .A. Alternative model

' E 5000- -S.-Bass : 11 node

E -N-Caa.1 :4 nod*

__ -@-Ca,,2 6Snode

on 4000- -A--Cue3:12 node

--- -- 4--:16 node

C)

3000 S. Default model

O -0-BSao* : 11 node

o -0--Casel : 4 node

0) 2000. --O-Ca*e2: 8 node

-A-Cao3 :12 node C

- -V,-Caae4 :15 node

- 1000

0)

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.17: Comparison of calculated heat transfer coefficients with the

variation of the node number

three important parameters are changed to show their sensitivity.

Simulations are performed for each case with two wall film condensation models, the default model and the alternative model. The difference of heat transfer coefficients is decreased

between two simulation results, as the inlet steam-air mixture flow rate increases, the inlet

air mass fraction decreases, and the inlet saturated steam temperature decreases. The heat

transfer coefficient calculated using the alternative model is always higher than that using

the default model, especially in the inlet of the condensing tube. It is due to the fact that

the correlation of the alternative model is based on the film thickness, which is very thin in

the inlet.

Effects of inlet steam-air mixture flow rate

The inlet steam-air mixture flow rate varies between 10, 20, 40, and 50kg/hr instead of the

33kg/hr of the base case.

Figures 3.18 and 3.19 show the simulation results of local air mass fraction along the

condensing tube with the change of the inlet steam-air mixture flow rate at the condition

44
Table 3.3: Input parameters for the sensitivity study

I.D. Node Pi. Ti. MF AMF Tot Ton CF

Number (kpa) (°C) (kg/h) (%) (°C) (°C) (gpm)

ss0 42 465.5 143.4 32.7 21.5 81.6 29.3 2.4

ndl 35 465.5 143.4 32.7 21.5 81.6 29.3 2.4

nd2 39 465.5 143.4 32.7 21.5 81.6 29.3 2.4

nd3 43 465.5 143.4 32.7 21.5 81.6 29.3 2.4

nd4 47 465.5 143.4 32.7 21.5 81.6 29.3 2.4

sfl 42 465.5 143.4 10.0 21.5 81.6 29.3 2.4

sf2 42 465.5 143.4 20.0 21.5 81.6 29.3 2.4

sf3 42 465.5 143.4 40.0 21.5 81.6 29.3 2.4

sf4 42 465.5 143.4 50.0 21.5 81.6 29.3 2.4

afl 42 403.9 143.4 32.7 2.5 81.6 29.3 2.4

af2 42 410.5 143.4 32.7 5.0 81.6 29.3 2.4

af3 42 425.0 143.4 32.7 10.0 81.6 29.3 2.4

af4 42 562.3 143.4 32.7 40.0 81.6 29.3 2.4

stl 42 118.3 100.0 32.7 21.5 81.6 29.3 2.4

st2 42 197.3 115.0 32.7 21.5 81.6 29.3 2.4

st3 42 315.3 130.0 32.7 21.5 81.6 29.3 2.4

st4 42 555.6 150.0 32.7 21.5 81.6 29.3 2.4

which the other parameters are the same

and the alternative model, respectively.

with the base case, using both the default model

The local air mass fraction of the default model

is linearly increased along the condensing tube keeping a relatively low air mass fraction

compared with that of the alternative model, which varies greatly along the condensing tube

especially for the low inlet steam-air mixture flow rate.

Figures 3.20 and 3.21 show the simulation results of local heat transfer coefficient, using

both the default model and the alternative model, respectively. For the simulation using

the default model, the local heat transfer coefficient is increased throughout the condensing

tube with the increase of the inlet steam-air mixture flow rate, and it is much more increased

especially in the inlet. However, for the simulation using the alternative model, it is independent of the inlet steam-air mixture flow rate in the inlet of the tube, but it is very much

changed in the middle of the condensing tube.

45
].O1 U.

0.8-

0

0.6.

0 0.4.

E

- *-Bss:32.7k~r

~-1n-Casel1Okghr~

I--A.-Cas*234Dkg/hrl

0.2J

u.u.l

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.18: Comparison of calculated air mass fractions with the variation of

the steam-air mixture flow rate in default model

0.8

ca

4C

E

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.19: Comparison of calculated air mass fractions with the variation of

the steam-air mixture flow rate in alternative model

46
NE5000 5000 - --0-Baso: 32.Tkg/hr . -&* ~ gh ,E .-l-case¶ lkoIthnr

-4-Case2 20kgmr

-A-Case: 40kglhr _ 4000- -.-- Cas04 : $Okg/hr

3o 4000

3002000.z

CA

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.20: Comparison of calculated heat transfer coefficients with the

variation of the steam-air mixture flow rate in default model

6000. ,

SG•s*l - : 22 kgthr " E --M--Casel :I tkg/lhr

--- Caso2 :0kg/hr

-4000,CasS 4Ok/hr

S4000.U -y--Case4 50kg/hr

3000.

0

2000.

0.0 0.5 1.0 1.\ 2.0 2.5

Length from the tube inlet (m)

Figure 3.21: Comparison of calculated heat transfer coefficients with the

variation of the steam-air mixture flow rate in alternative model

47
The two simulated heat transfer coefficients are greatly different in the inlet, but they

are similar in the outlet of the condensing tube. As the inlet mixture flow rate increases,

the heat transfer coefficient using the default model increases to be similar with that using

the alternative model. Particularly with the low steam flow rate, the simulated heat transfer

coefficients by the alternative model is more highly evaluated than that by the default model.

Effects of inlet air mass fraction

The inlet air mass fraction varies between 2.5, 5, 10, and 40% instead of the 20% of the base

case.

Figures 3.22 and 3.23 show the simulation results of local air mass fraction along the

condensing tube with the change of the inlet air mass fraction at the condition which the

other parameters are the same with the base case, using both the default model and the

alternative model, respectively. The local air mass fraction of the default model is linearly

increased along the condensing tube keeping a relatively low air mass fraction compared with

that of the alternative model, which varies greatly along the condensing tube, especially for

the low inlet air mass fraction.

Figures 3.24 and 3.25 show the simulation results of local heat transfer coefficient, using

both the default model and the alternative model, respectively. For the simulations using

both the default model and the alternative model, the heat transfer coefficients are highly

decreased in the inlet of the condensing tube with the increase of the air mass fraction, but

they become similar in the outlet. However, the heat transfer coefficient calculated using the

default model is always lower than that using the alternative model except for the lower part

of the condensing tube. The effect of inlet air mass fraction is shown only in the upper part

of the condensing tube, but it is negligible in the lower part.

Effects of inlet saturated steam temperature

The inlet saturated steam temperature varies between 100, 115, 130, and 1501C instead of

the 1431C of the base case.

48
1.0

0.8

0.6

E 0.4

0.2

0.0

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.22: Comparison of calculated air mass fractions with the variation of

the inlet air mass fraction in default model

C

0

Go

Eo

1.0

0.8.

0.6.

0.4.

0.2.

0.0 0.5 1.0 1.5

Length from the tube inlet (m)

2.0 2.5

Figure 3.23: Comparison of calculated air mass fractions with the variation of

the inlet air mass fraction in alternative model

49
10000- -*-Base: -J- •Casel 21%2.5%

I---Cass2 5%

S8000 -A-Cas*3: 10%

'E ~-y-C&se4 :40%

6000

0

.) 4000-

Cn

- 2000 0. CO0)

0 i &

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.24: Comparison of calculated heat transfer coefficients with the

variation of the inlet air mass fraction in default model

12000

10000 -0-Base :21% I 0 0 0 - 0Q - C as el ; 25 %

3 00 --- C.se2 : 6%

-A--Cases : 10%

8000 -- Cue4 :40%

0

6000

0

0

0 4000.

C CU.

-2000 ' *

0.5 1.0 1.5

Length from the tube inlet (m)

Figure 3.25: Comparison of calculated heat transfer coefficients with the

variation of the inlet air mass fraction in alternative model

50
Figures 3.26 and 3.27 show the simulation results of local air mass fraction along the

condensing tube with the change of the inlet saturated steam temperature at the condition

which the other parameters are the same with the base case, using both the default model

and the alternative model, respectively. The local air mass fraction of the default model

is linearly increased along the condensing tube keeping a relatively low air mass fraction

compared with that of the alternative model, which varies greatly along the condensing tube

especially for the high inlet saturated steam temperature.

Figures 3.28 and 3.29 show the simulation results of local heat transfer coefficient, using

both the default model and the alternative model, respectively. For the simulation using the

default model, the heat transfer coefficient is highly decreased with the increase of the inlet

saturated steam temperature in the inlet of the condensing tube, but it is little changed in the

outlet. However, for the simulation using the alternative model, the heat transfer coefficient

is a little decreased with the increase of the inlet saturated steam temperature in the upper

part of the condensing tube, but it is much more decreased in the lower part.

3.4.4 Run statistics

The computer used in the calculation is a SUN SPARC 10 with SunOS 4.1.3-KL operating

system. The random access memory is 32Mbyte, the calculation speed is 86.1MIPS, and

the clock speed of CPU is 36MHz.

The CPU time, the time step size, and the grind time were compared between two calculation results with different wall film condensation models. Figure 3.30 shows the required

CPU times with respect to the real problem time for the base case calculation of El2b. The

required CPU times increase linearly for both the default model and the alternative model,

except for the initial transient situation. The CPU time is greatly changed near 40sec. When

the default model is used, the required CPU time is slightly longer than that of the alternative

model. Figure 3.31 shows the time step sizes with respect to the real problem time for the

base case calculation of El2b. The time steps fluctuate between 0.0125, 0.025, and 0.05sec

for both the default model and the alternative model. When the default condensation model

51
1.0 I

0.8-

0

o 0.6-

E 0.4.

0.2.

-0-B&SO. 143.4*C

-M-Casel : 00C

-*--C&sa2 :115'C

-A-Cas*3 :130'C

yC&S64: IBOC0

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.26: Comparison of calculated air mass fractions with the variation of

the inlet saturated temperature in default model

1.0 

.

.

1.0

0.8

0.

0/) 0o

-0.4-

0.2.

0.0 0.5 1.0 1.5 2.0

Length from the tube inlet (m)

2.5

Figure 3.27: Comparison of calculated air mass fractions with the variation of

the inlet saturated temperature in alternative model

52
7000

C.,

C

E

a,

.2

a,

0

U

a,

a,

a,

a,

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.28: Comparison of calculated heat transfer coefficients with the

variation of the inlet saturated temperature in default model

,U.JU

0

0•

0

q)

x

6000.

5000.

4000.

3000.

2000.

1000.

0

~-0- Base: 143.4"G

-a- C"01 : 100*C

• -01- Cae3 : I: I0"C

-- I--CSS4 : 11C

V .O4 5*

A~<

0.0 0.5 1.0 1.5 2.0 2.5

Length from the tube inlet (m)

Figure 3.29: Comparison of calculated heat transfer coefficients with the

variation of the inlet saturated temperature in alternative model

53
10000.

0

* 1000-

a,

E

0 100.-

10.

10 100 1000

Problem time (second)

Figure 3.30: Comparison of required CPU times of two condensation models

is used, the time step size is smaller, and its fluctuation is more frequent than that of the

alternative model.

The time step determined by the code is also shown in Figure 3.32 with respect to the user

specified maximum time step for the base calculation of E12b. The code determined time

step with the default model is observed to scatter much more than that with the alternative

model. It validates larger fluctuation of the time step and longer required CPU time.

The grind time is expressed as:

CPU x 103

Grind time = CxAT (3.14)

where CPU is the CPU time, C is the total number.of model volumes, and AT is the number

of time steps. It means the CPU time used for calculating a volume during a second.

Table 3.4 shows the required CPU time and the grind time for the base case calculations,

and Table 3.5 shows those for the nodalization study and the sensitivity study.

The required CPU time is the highest when the inlet mixture flow rate is the highest and

54
A.^A

0

"o

2

0

0

CA

W

N

0.

I.-

0.09-

0.08.

0.07.

0.06-

0.05-

0.04.

0.03-

0.02.

0.01.

0.00

alternative model

I

o1'000 2000' 3o000' 4o00

Problem time (second)

5o00

0

0

0

0

0

0

a.

0

0

0

0.10

0.09.

0.08.

0.07-

0.06-

0.05-

0.04-

0.03-

0.02-

0.01.

0.00

defaut model

0 1000 2000' 3000 4o00

Problem time (second)

5o00

Figure 3.31: Comparison of time step sizes of two condensation models

55
0.10

e-0.09-

0.093 alternative model

i 0.08

0.

C 0.07

E

a,

E 0.06

.,,.

4)

£

"E 0.04

a 0.03

", 0 0] 0

.o 0.02

0

00.01000

0.0011133

0.00 0.02 0.04 0.08 0.08 0.10

User specified maximum time step (second)

0.10.

C 0.09. 0 *l default model

S 0.08-

• 0 7.

0.06. E

o 0.04.

C

S0.04.

0.02. 0

0.00

0.00 0.02 0.04 0.06 0.68 0.10

User specified maximum time step (second)

Figure 3.32: Comparison of code determined time steps of two condensation

models

56
both the air mass fraction and the inlet saturated steam temperature is the lowest. The grind

time of the default model is about 23% higher than that of the alternative model. The grind

time is a little increased when more volumes are included in RELAP5/MOD3.2 simulation,

but it is little changed with the same node number when the other parameters are changed.

57
Table 3.4: The CPU time and the grind time of the simulation of KAIST

condensation experiment

I.D. Problem time condensation CPU time Number of Grind time

(second) model (second) time step

4d 1000 alternative 7320.5 61748 2.823

default 8740.3 60302 3.451

11d 1000 alternative 7178.6 59666 2.865

default 8986.4 59942 3.569

12a 1000 alternative 3434.6 28177 2.902

default 4286.2 28326 3.603

12b 1000 alternative 5944.8 46591 3.038

default 7202.1 46671 3.674

5000 alternative 28519.5 233309 2.910

default 35098 233281 3.582

13b 1000 alternative 7558.4 61716 2.916

_ default 9902.7 66476 3.547

58
Table 3.5: The CPU time and the grind time of the sensitivity study

I.D. Node Problem The alternative model The default model

time(s) CPU No. of Grind CPU No. of Grind

time(s) time step time time(s) time step time

ss0 42 250 1511.6 11647 3.090 1790.1 11668 3.653

ndl 35 250 1053.1 11772 2.556 1127 11337 2.840

nd2 39 250 1380.8 12276 2.884 1576.7 12002 3.368

nd3 43 250 1680.6 12389 3.155 1928.8 11779 3.808

nd4 47 250 1862.7 11649 3.402 2289.8 11659 4.179

sfl 42 250 849.4 6854 2.951 728.7 4940 3.512

sf2 42 250 990.9 7995 2.951 1174.8 7903 3.539

sf3 42 250 1870.3 14974 2.974 2275.6 14843 3.650

sf4 42 250 2501.7 19888 2.995 3019.9 19850 3.622

afl 42 250 4463.4 36664 2.899 5415.8 36578 3.525

af2 42 250 3552.8 29639 2.854 4243.1 29052 3.477

af3 42 250 1997.7 16033 2.967 2307.2 15958 3.442

af4 42 250 1193.6 9369 3.033 1376.3 9094 3.603

stl 42 250 2587.1 19866 3.101 2364.7 15628 3.603

st2 42 250 1858.3 15328 2.887 2344.7 15276 3.655

st3 42 250 1863.4 15157 2.927 2302.4 15112 3.628

st4 42 250 1202 9728 2.942 1404.5 9593 3.486

59
Chapter 4

Conclusions and Recommendations

The capability of the RELAP5/MOD3.2 code in modeling the condensation heat transfer in

the presence of noncondensable gas has been investigated. Several experiments are performed

to get a reliable data on that condensation phenomena, to show the parametric effects on

condensation heat transfer, and to elucidate the cause of data scattering among the previous

experiments. With the constructed database direct assessments are carried out for both the

default and alternative condensation models in RELAP5/MOD3.2. After the experimental

apparatus being modeled with RELAP5/MOD3.2, the simulation results are compared between two wall film condensation models, and they are also compared with the experimental

data. Nodalization study is performed to investigate the effect of the divided node number

of the heat structure. Also the sensitivity study on the important experimental parameters

is performed, and the run statistics are checked.

From the studies, the followings are concluded:

1. From the comparison of the experimental results, it is known that there are three

important experimental parameters: the inlet steam-air mixture flow rate; the inlet air

mass fraction; the inlet saturated steam temperature. The experimental results show

that the condensation heat transfer coefficient increases as the inlet steam-air mixture

flow rate increases, the inlet saturated steam temperature decreases, and the inlet air

60
mass fraction decreases.

2. From the direct assessment of two condensation models of RELAP5/MOD3.2, the default condensation model under-predicts the experimental data for the low heat transfer

coefficient range, but it over-predicts the experimental data for the high heat transfer

coefficient range. In case of the alternative condensation model, the predicted values

are always higher than the experimental data.

3. When the two wall film condensation models of RELAP5/MOD3.2 are compared each

other, the calculated air mass fraction from the default model is always lower than

that of the alternative model, all the calculated temperatures from the default model

decrease more slowly than those from the alternative model, and the calculated heat

flux from the default model is always lower than that from the alternative model in the

upper part of the condensing tube, and they are similar in the lower part.

4. Simulation results of base case calculations show that the default model of RELAP5/MOD3.2

predicts well the experimental heat transfer coefficient with high inlet mixture flow rate

and low inlet air mass fraction, but it always under-predicts the experimental data

except for the inlet condition. It is also known that the alternative model of RELAP5/MOD3.2 over-predicts the experimental data throughout the condensing tube,

but it predicts well the experimental data with high inlet mixture flow rate except for

the middle part of the condensing tube. From base case calculations it is considered to

be that the effect of the inlet steam-air mixture flow rate is not properly considered in

the default model, and the effect of the inlet air mass fraction in the alternative model.

5. The change of the divided node number in heat structure has little influence on the

simulation results of the condensation phenomena with RELAP5/MOD3.2 both for the

default and for the alternative model.

6. The simulation results for the sensitivity study show that the effects of the inlet steamair mixture flow rate, the inlet air mass fraction, and the inlet saturated steam tempera61
ture are most influential in condensation experiments in the presence of noncondensable

gas in a vertical tube. However, the effects of the coolant flow rate, the inlet coolant

temperature, and the vented mixture temperature are less influential.

7. Run statistics show that the grind time of the default model is always higher than that

of the alternative model and it is a little increased for both models when more volumes

are included in RELAP5/MOD3.2 simulation. The required CPU time is the highest

when the inlet mixture flow rate is the highest and both the inlet air mass fraction and

the inlet saturated steam temperature are the lowest.

8. As both the default model and the alternative model can not predict well the experimental data, it is needed to develop a new correlation based on the experimental data

produced over various operating ranges, and it can be applied as a new condensation

model to simulate the condensation heat transfer of the steam-noncondensable gas mixture.

62
Bibliography

[1] S. H. Chang et al. "Korea Looks Beyond the Next Generation". Nuclear Engineering

International, pages 12-16, Feb 1997.

[2] H. C. NO, S. I. Lee, H. S. Park, S. J. Kim, and H. T. Kim. "Validation Experiments

and Code Development for Passive Safety Systems of CP-1300". In Proc. of the 2nd

Japan-Korea Seminar on Advanced Reactors, pages 77-85, 1996.

[3] Sang R Lee and Hee Cheon NO. "Assessment of RELAP5/MOD3.1 for Direct-Contact

Condensation in the Core Makeup Tank of the CARR Passive Reactor". Annals of

Nuclear Energy, 24(7):553-562, 1997.

[4] Hyoung The Kim and Hee Cheon NO. "Modification of Condensation Heat Transfer

Model of RELAP5/MOD3.1 for the Simulation of Secondary Condensers". Nuclear

Technology, 119:98-104, 1997.

[5] S.W. Lee, W.P. Baek and S. H. Chang. "Assessment of Passive Containment Cooling Concepts for Advanced Pressurized Water Reactors". Annals of Nuclear Energy,

24(6):467-475, 1997.

[6] B. S. Rao et al. "Simplified Boiling Water Reactor Design". In 1st JSME-ASME Int.

Conf. on Nuclear Eng.(ICONE-1), pages 295-300, 1991.

[7] B. S. Shiralkar et al. "Thermal Hydraulic Aspects of the SBWR Design". Int. Conf. on

Design and Safety of Advanced NPPs, 3:31.1.1-31.1.7, 1992.

63
[81 H. A. Upton et al. "Simplified Boiling Water Reactor Passive Safety Features". In 2st

ASME-JSME Int. Conf. on Nuclear Eng. (ICONE-2), volume 1, pages 705-712, 1993.

[9] K. M. Vierow. "Behavior of Steam-Air Systems Condensing in Concurrent Vertical

Downflow". Ms thesis, U.C.B., 1990.

[10] K. M. Vierow and V. E. Schrock. "Condensation in a Natural Circulation Loop with

Noncondensable Gases: Part I - Heat Transfer". In Proc. of the Int. Conf. on Multiphase

Flows '91, pages 183-186, 1991.

[11] D. G. Ogg. "Vertical Downflow Condensation Heat Transfer in Gas-Steam Mixtures".

Ms thesis, U.C.B., 1991.

[12] M. Siddique, M. W. Golay and M. S. Kazimi. "Local Heat Transfer Coefficients for

Forced-Convection of Steam in a Vertical Tube in the Presence of a Noncondensable

Gas". Nuclear technology, 102:386-402, 1993.

[13] M. Siddique, M. W. Golay and M. S. Kazimi. "Theoretical Modeling of Forced Convection Condensation of Steam in a Vertical Tube in the Presence of a Noncondensable

Gas". Nuclear technology, 106:202-214, 1994.

[14] H. A. Hasanein, M. W. Golay and M. S. Kazimi. "Steam Condensation in the Presence

of Noncondensable Gases under Forced Convection Conditions". M.I.T., MIT-ANP-TR024, 1994.

[15] H. Araki, Y. Kataoka and M. Murase. "Measurement of Condensation Heat Transfer

Coefficient Inside a Vertical Tube in the Presence of Noncondensable Gas". Journal of

nuclear science and technology, 32(6):517-526, 1995.

[161 S. Z. Kuhn. "Investigation of Heat Transfer from Condensing Steam-Gas Mixtures and

Turbulent Films Flowing Downward Inside a Vertical Tube". Phd thesis, M.I.T., 1995.

64
[17] W.S. Yeung, J. Shirkov and F. Seifaee. "RELAP5/MOD3 Modeling of Water Column

Rejoining and a Water Slug Propelled by Noncondensable Gas". Nuclear Technology,

108:387-394, 1994.

[18] Ki Yong CHOI, Hyun Sik PARK, Sang Jae KIM, Hee Cheon NO and Yong Seok BANG.

"Assessment and Improvement of Condensation Models in RELAP5/MOD3.2". submitted to the Nuclear Technology, 1997.

[19] V. H. Ransom et al. "RELAP5/MOD3 Code Manual". NUREG/CR-5535, INEL, June

1995.

[20] W. Nusselt. "Die Oberflachenkondensation des Wasserdampfes". Z. Ver. Deutsch. Ing.,

1916.

[21] M. M. Shah. "A General Correlation for Heat Transfer during Film Condensation Inside

Pipes". Int.J. Heat Mass Transfer, 22:547-556, 1979.

[22] A. P. Colburn and 0. A. Hougen. "Design of Cooler Condensers for Mixtures of Vapors

with Non-condensing Gases". Industrial and Engineering Chemistry, 26(11):1178-1182,

1934.

[23] J. C. Chato. "Laminar Condensation inside Horizontal and Inclined Tubes". American

Society of Heating, Refrigeration and Air Conditioning Engineering Journal, 4:52-60,

1962.

[24] W. M. Rohsenow and Y. H. Choi. "Heat Mass and Momentum Transfer". Prentice Hall,

New Jersey, 1961.

[25] W. M. Kays and E. Y. Leung. "Heat transfer in annular passages-hydrodynamically developed turbulent flow with arbitrarily prescribed heat flux". Int.J.Heat Mass Transfer,

6:537-557, 1963.

[26] W. M. Kays and M. E. Crawford. "Convective Heat and Mass Transfer". McGraw-Hill,

Inc., 1993.

65

Appendix A

Test facility and its instrumentation

The experimental facility consists of a steam tank including a 100kW heater, a steamnoncondensable gas mixture supply line, a test section with a condensing tube and its surrounding coolant jacket, a lower plenum, venting and draining systems, and a unit of data

acquisition system. The schematic diagram of the experimental apparatus is shown in Figure 2.1.

Steam and noncondensable gas supply

The steam tank functions as a containment. It is a cylindrical stainless steel pipe, which is

1540mam in height, 1000mm in inner diameter and 10mm in thickness on the top and bottom

of which 12-mm-thick stainless steel caps are welded. The total height is 2000mm and the

total volume is 1.332m . It has a heater, two level gauges, a manhole, two sight glass and a

relief valve.

The 108kW heater is a highly insulated, sheathed and flanged immersion heater which

is located in the lower part of the steam tank in a horizontal direction. 36 heaters whose

capacities are 3kW each are connected in 6-parallel. It is finely controlled by a 350A SCR

power controller with a high precision voltmeter, an ampere-meter, and a watt-meter, which

can be controlled automatically or manually.

67
A water refill line is attached to the upper demineralized water reservoir and during

refilling the demineralized water is supplied from the reservoir by gravity. The two level

gauges is installed to indicate the initial water level.

Steam is generated from the steam tank and is mixed with the supplied noncondensable

gas in the steam tank. The noncondensable gas is supplied by the noncondensable gas supply

line, and the vacuum can be achieved using a vacuum pump. Steam-noncondensable gas

mixture gives a predetermined pressure in the steam tank to give an initial condition in

thermal equilibrium and in homogeneous state. The noncondensable gas is injected into the

steam tank continuously to give a continuous initial steam-noncondensable gas mass fraction,

or to keep a total air inventory.

Even though the pressure of the steam tank is a little changed due to the injection of

the noncondensable gas and the ejecting of the steam-noncondensable mixture, the steam

tank will be in quasi-steady state because the noncondensable gas is injected deliberately to

preserve the pressure, the temperature and the initial mass fraction of the noncondensable

gas.

Test section

The steam-noncondensable gas mixture is supplied through a 1 inch pipe with a vortex flow

meter. The mixture is made to keep uniform velocity distribution and its temperature and

pressure are measured in the inlet of the test section. The upper plenum is a 2 inch pipe with

tabs to install a thermocouple and a pressure transducer.

The test section consists of an inner condensing tube and an outer coolant jacket. The

inner tube of the heat exchanger is a stainless steel pipe of 50.8mm in outer diameter,

1.65mm in thickness, and 2400mm in length. At 12 different axial locations each J-type

thermocouple is welded on the outer surface of the condensing tube to measure the outer

surface temperatures, welded through the condensing tube to measure the mixture bulk

temperatures, and installed on the outer side of the coolant jacket to measure the coolant

temperatures. These three kind of temperature sensors axe located irregularly so that the

68
distance between each temperature measuring probes are chosen to be the same total heat

transfer in each interval. They are located at 0, 10, 22, 36.5, 53.5, 73, 95, 119.5, 146, 175,

206.5, 240cm from the inlet of the uninsulated test section. The thermocouples are located

densely at the inlet of the steam-noncondensable gas mixture because the heat transfer is

high and most important in that region.

The inner tube is surrounded with a concentric coolant jacket pipe which is 100mm in

inner diameter and 10mm in thickness, and the concentric jacket can be divided with three

parts with the total length of 2400m. There are another two small jackets at the inlet and

outlet of the coolant jacket. They are installed to stabilize the flow inside the coolant jacket.

The lower one is designed to minimize the effects of the developing region at the inlet of the

jacket and the upper one is designed to minimize the multi-dimensional effects at the outlet

of the jacket.

The cooling water flows upward through the annulus and the steam-noncondensable gas

mixture flows downward. The cooling water is supplied to the lower end of the jacket pipe in

an once-through mode by a calibrated rotameter and a flow control valve, and is dumped to

the drain after its temperature is increased. Two rotameters are used to measure the coolant

flow rates. The coolant flow in the cooling jacket is in a turbulent regime because Reynolds

number is about 2800 for the minimum flow rate of 363kg/hr.

Lower plenum

The lower plenum is a small hexahedral tank with the dimension of 200mm and 200mm in

the bottom plate and 150mm in height. It is located at the outlet of the test section to

help measure the temperature with a J-type thermocouple, and the pressure with a pressure

transmitter. The level gauge is also used to check the level of the condensate in the lower

plenum. It is connected to drain lines, a venting line and the outlet of the test section.

69
Venting and drain systems

The steam-noncondensable gas mixture is vented through the venting line and the condensate

liquid from the outlet of the test section is drained down to the drain tank separately. The

drain tank is connected to the lower plenum with the drain line and the pressure balancing

line. The condensate liquid is drained to the drain tank through the drain line, and the

pressure balancing line is used to equalize the pressure between the drain tank and the lower

plenum. The valve on the vent line is regulated to keep the system pressure constant, and to

change the venting rate. The vented mixture flow rate is measured.

Instrumentation and data acquisition system

Table 2.2 summarizes the instrumentation of PCCS experimental apparatus. All the signals

from detectors are current output or voltage output including the output from the thermocouple. 4 ,- 20mA current signals from the pressure transducer and the vortex flow meter

are converted to 1 - 5V voltage signals using an multi-channel current-to-voltage converter.

The temperature signals are compensated to give accurate temperature data. All signals

are applied to the Hewlett Packard 44708H high voltage relay multiplexer to give a reliable

temperature and voltage data. Sampled experimental data is acquired using the HP3852A

control unit and the data is transferred to IBM-PC/AT using the RS232C interface.

70
Appendix B

Data reduction method

Calculation of coolant bulk mean temperature

Kuhn[16] calculated the coolant bulk temperature in an annulus by measuring temperatures

at the inner and the outer walls of the annulus, and using the numerically calculated temperature shape factors. The shape factors, F, are calculated for various inner and outer wall

temperatures and coolant flow rates by solving the turbulent k - c equations.

A simplified method to calculate the mean bulk coolant temperature is developed, which

is a modified form of Kays[25]. It uses the empirical velocity profile and eddy diffusivity

profile without solving the turbulent k - c equations.

If the heat transfer in annular passages is considered for the hydrodynamically developed

turbulent flow with arbitrarily prescribed heat flux, the energy differential equation can be

written as follows:

i r(a+eHi)aTr . = ru(r) 09•,x (B.1)

Or Ox

The radius at which the maximum axial velocity is observed can be expressed by the following

relationships[25].

S= = (r*)0.343 (B.2)

where r* is the annulus radius ratio, ri/ro, and s* is (W - r*)/(1 -'§). T is the ratio of the

71
radius at the maximum axial velocity, s, to the outer radius of the annulus, r,.

The flow area in annular passage is broken into four sections radially at a certain axial location: a sublayer near the inner surface(Region I); a sublayer near the outer surface(Region

II); a fully turbulent region from the inner sublayer to the point of maximum velocity(Region

III); and a fully turbulent region from the outer sublayer to the point of maximum velocity(Region IV). Their velocity profiles and eddy diffusivity profiles can be described respectively, as follows:

Region I :y+<10.8, u+=y+ (B.3)

Region II :y+ > 10.8, u+ = ay +b (B.4)

Region III : y+ > 10.8, u+ = 2.5y+ + 5.5 (B.5)

Region IV :y,+<l10.8, u+ =Y (B.6)

where yo = ro - r, yi = r - ri, u+ = Ulur = /Uo4 /, and y+ = yuTIA.

The velocity profile for the wall coordinate is used at Region I and Region IV, and the

velocity profile of Nikuradse equation[26] for the tube coordinate is used at Region III. At

region II the velocity profile for the tube is used and the unknown constants, a and b, are

calculated using Kay's method. The momentum eddy diffusivity eM/l is also calculated from

Kay's method and the thermal eddy diffusivity is given from the calculated momentum eddy

diffusivity as follows:

PrT = ATC-P = 0.9 (B.7)

AT

Since the velocity, the momentum and the thermal eddy diffusivity profiles are all known,

the energy equation can be solved to give the radial temperature distribution and thus the

local mean coolant temperature.

72
Local heat flux

Local heat flux is typically obtained in terms of the axial gradient of the measured coolant

bulk temperature.

q, = ncp dTLoo (B.8) -7rd dL

The accuracy of the calculated heat flux mainly depends on the coolant bulk temperature

distribution.

Another method available to obtain the local heat flux is to use the Fourier's conduction

law.

q k . (T.,i. - TW,o0 t) (B.9)

twaul

To determine the heat transfer two temperature measuring devices are installed on either

side of the material to be measured. The accuracy of the heat flux mainly depends on the

wall temperature measurement device, but error can be generated during the positioning of

the device. To measure reasonably accurate inner and outer wall temperatures, the original

thickness of the isolation condenser must be increased.

As an effective method to measure the heat transfer through any surface material like the

condensing tube, the heat flux sensor is developed by RdF corporation. The heat flux sensor

measure the heat transfer through a surface by differentiating temperature between opposite

sides of certain rigid materials thereby allowing a direct measurement of the heat transfer

through the material surface with a known thermal resistance.

In these experiments the heat flux is calculated by the temperature gradient of the coolant

bulk temperature and is cross-checked with the directly measured one with the RdF heat flux

sensor at a specific location of the condensing tube.

Condensation heat transfer coefficient

To obtain local heat transfer coefficient experimentally, it is necessary to determine the local

heat flux and the temperature difference between the steam-noncondensable gas mixture bulk

73
and the inner wall.

h (Tmi, -h Kwin) q'F(B.10) (.0

where Tmix is the steam-noncondensable gas mixture bulk temperature and Tw,in is the

inner wall temperature of the condensing tube.

The inner wall temperature of the condensing tube

The inner wall temperature of the condensing tube can be expressed using the measured

outer wall temperature and the calculated heat flux, as follows:

Tin(X) = Tw,ot(x) + q" (x)W ln(Dout/Din)Dhydrautic (B.11) 7rkss30 4

where Din is the inner wall diameter, Dour the outer wall diameter, Dhydrauiic the hydraulic

diameter, and k,,304 the thermal conductivity of the condensing tube.

74
Appendix C

RELPA5/MOD3.2 Input Listing

C-1. Base case calculation with default model

(Experiments : #12b)

=RELAP5/MOD3.2 Simulation

******** ** * ******** ****** ** ******$*

*Condensation Experiment in the presence

*ofnoncondensible gas in a Vertical Tube

*Dec 9. 1997 by HSPARK

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

* miscellaneous control cards

100 new transnt

101 run

10510.20.

110 air

* *** ***0.*** *** *** **** * ** **0  ** ***

* time step cards

* *** **$** * **** *** **** *0* *** *** 0****

* end min.st max ctrl minor major rstplt

201 10. 1.0e-g 0.025 3 100400 400

202 100. 1.Oe-8 0.05 3 500 2000 2000

203 1000. 1.Oe-8 0.1 3 2500 10000 10000

204 5000. 1.Oe-8 0.1 3 5000 20000 20000

*204 5000. l.Oe-8 0.1 3 5000 20000 20000

*201 10. 1.Oe-8 0.00001 3 2000000 10000000 10000000

* original time step

*201 500. l.Oe-8 0.005 3 2000 10000 10000

* trip logic

502 time 0 It null 0. 0. n 0.0

506 time 0 gt null 0. 0. n -1.0

508 quala 150010000 gt null 0. 0.8 n 0

* hydrodynamic data:

* steam-air mixture tube

* component 100 : SIG using time dependent volume

* name type

1000000 tdv 100 tmdpvol

* area length vol x angle elev rough dhydr vflag

1000101 2.0 2.0 0.0 0.0 0.0 0.0 L.e-4 0.0 00000

* noncondensible gas(air)

* cntrl

1000200004

* var pressure temp. eq.quality

1000201 0.0 0.46546675e6 0.4165528e3 1.0 ,

* component 105 : steam flow initiation in kg/s

* name type

1050000 tdj 105 tmdpjun

* from to area

* cccOO0000 cccvvOOOn

1050101 100000000 115010001 0.00049

* cntrl trip

1050200 1 502

* var waterflow steamflow interface vel.

75
1050201 -1.0 0.0 0.009087 0,0

1050202 0 0.0 0.009087 0.0 ,

* component 115 : pipe to upper plenum

* name type

1150000 sv115 snglvol

* area length vol x angle

1150101 0.00049 0.07 0.0 0.0 90.

* dev rough dh vflag

1150102 0.07 l.e-4 0.0 00000

* cntrl pressure quality

*1150200 002 0.46546675e6 1.0

* cntrl pressure temp.

1150200 003 0.46546675e6 0.35474181e3

* component 120: upper plenum

* name type

1200000 b120 branch

* no. jun cntrl

1200001 2 1

* area=lOcm length vol x angle elev rough dh vflag

1200101 0.007854 0.20 0.00. -90. -0.20 L.e-4 0.0 11000

* cntrl pressure quality

* 1200200 002 0.46546675e6 1.0

* cntri pressure temp.

1200200 003 0.46546675e6 0.35474181 e3

* from to area kforw kbackwjflag

* cccnlOI cccvvOOOn

1201101 115010002 120010002 0.000490 225.847 120.438

000000

1202101 120010002 140010001 0.001772 7.608 11.783

000000

* waterflow steamflow x

* cccn20l

1201201 0.0 0.009087 0.0

1202201 0.0 0.009087 0.0

* component 140: condensing tube

* name type

1400000 pipe 140 pipe

* nv: no. vol

1400001 13

* area nv

1400101 0.001772 13

* length vn

* ccc0301 - ccc0399

1400301 0.500 1

1400302 0.100 2

1400303 0.120 3

1400304 0.145 4

1400305 0.170 5

1400306 0.195 6

1400307 0.220 7

1400308 0.245 8

1400309 0.265 9

1400310 0.290 10

1400311 0.315 11

1400312 0.335 12

1400313 0.400 13

* ccc03OI - ccc0399

*1400301 0.500 1

* 1400302 0.200 12

*1400313 0.400 13

* volume nv

*1400401 0.0 13

* angle vn

1400601 -90:0 13

* rough dhydr no. vol

1400801 l.e-4 0.0 13

* vflag no. vol

1401001 00000 13

* jflag no. jun

1401101 000000 12

* cntrl pressure temp. nv

*ccc1201-ccc1299

1800200003

1401201 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 1

1401202 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 2

1401203 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 3

1401204 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 4

1401205 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 5

1401206 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 6

1401207 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 7

1401208 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 8

1401209 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 9

1401210 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 10

1401211 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 11

1401212 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 12

1401213 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 13

* cntrl

76
1401300 1

* water flow steam flow int.vel jn

1401301 0.0 0.0 0.0 12

* component 145 : condensing tube to lower plenum

* name type

1450000 sj145 sngljun

* from to area kforw kbackw jflag

1450101 140130002 150010001 0.001772 11.783 7.608

0001000

* cntrl waterflow steamflow x

1450201 1 0.0 0.0 0.0

* component 150: lower plenum

* name type

1500000 pipe 150 pipe

* nv

1500001 3

* area nv

1500101 0.007854 3

* length vn

1500301 0.10 1

1500302 0.10 2

1500303 0.10 3

* volume nv

1500401 0.0 3

* angle nv

1500601 -90.0 3

* rough dhydr nv

1500801 L.e-4 0.0 3

* vflag nv

1501001 00000 3

* jflag nj

1501101000000 2

* cntrl pressure temp. eq.quality vn

1501201 004 0.46546675e6 0.35474181e3 1.0 0.0 0.0 1

1501202 004 0.46546675e6 0.35474181e3 1.0 0.0 0.0 2

1501203 004 0.46546675e6 0.35474181e3 1.0 0.0 0.0 3

* cntrl

1501300 1

* waterflow steamflow int.vel nj

1501301 0.0 0.0 0.0 2

•******************

* hydrodynamic data: venting line

* component 155 : lower plenum to drain tank

* name type

1550000 sj155 sngljun

* from to areav kforw kbackw jflag

1550101 150030002 160050002 0.00001 923317.24

923317.24 0001000

* cntrl waterflow steamflow x

1550201 10.0 0.0 0.0

* component 156: lower plenum to drain pipe

* name type

1560000 sj156 sngljun

* from to areav kforw kbackw jflag

1560101 150030002 157050002 0.00012 2109.13 4153.80

0001000

* cntrl waterflow steamflow x

1560201 10.0 0.0 0.0 #

* component 157 : drain pipe

* name type

1570000 pipe157 pipe

* nv

1570001 5

* area nv

1570101 0.00012 5

* length vn

1570301 0.20 5

* volume nv

1570401 0.0 5

* angle nv

1570601 90.0 5

* rough dhydr nv

1570801 1.e-4 0.0 5

* vflag nv

157100100000 5

* jflag nj

1571101000000 4

* cntrl pressure temp. eq.quality vn

1571201 004 0.46546675e6 0.35474181e3 1.0 0.0 0.0 1

1571202 004 0.46546675e6 0.35474181e3 0.5 0.0 0.0 2

1571203 004 0.46546675e6 0.35474181e3 0.0 0.0 0.0 3

77
1571204 004 0.46546675e6 0.35474181e3 0.0 0.0 0.0 4

1571205 004 0.46546675e6 0.35474181e3 0.0 0.0 0.0 5

* cntrl

1571300 1

* waterflow stearnflow int.vel nj

1571301 0.0 0.0 0.0 4

* component 158 : drain pipe to drain tank

* name type

1580000 sj158 sngijun

* from to areav kforw kbackw jflag

1580101 157010001 160010001 0.00012 42823936.0

21415240.0 0001000

* cntrl waterflow steamflow x

1580201 1 0.0 0.0 0.0

* component 160: drain tank

* name type

1600000 pipel60 pipe

* nv

1600001 5

* area nv

1600101 0.7854 5

* length vn

1600301 0.20 5

* volume nv

1600401 0.0 5

* angle nv

1600601 90.0 5

* rough dhydr nv

1600801 1.e-4 0.0 5

* vflag nv

160100100000 5

* jflag nj

16011010000004

* cntrl pressure temp. eq.quality vn

1601201 004 0.46546675e6 0.35474181e3 1.0 0.0 0.0 1

1601202 004 0.46546675e6 0.35474181e3 0.5 0.0 0.0 2

1601203 004 0.46546675e6 0.35474181e3 0.0 0.0 0.0 3

1601204 004 0.46546675e6 0.35474181e3 0.0 0.0 0.0 4

1601205 004 0.46546675e6 0.35474181e3 0.0 0.0 0.0 5

* cntri

1601300 1

* waterflow steamflow intvel nj

1601301 0.0 0.0 0.04

* componet 175 : valve for venting

* name type

1750000 vv175 valve

* from to areav kforw kback jflag

1750101 150030002 180000000 0.00012 2109.13 4153.80

000100

* cntrl waterflow steamflow interface velocity

1750201 10.0 0.0 0.0

* valve type

1750300 trpvlv

* trip number(open when true)

1750301 506

* component 180 : venting simulation using tdv

* name type

1800000 tdv 181 tmdpvol

* area length vol x angle elev rough dhydr vflag

1800101 2.0 2.0 0.0 0.0 0.0 0.0 0.0 0.0 00000

* cntrl

1800200003

* var pressure temp.

1800201 0.0 0.46546675e6 0.35474181e3

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

* hydrodynamic data: coolant water annulus

* component 200 : coolant source simulation using tdv

* name type

2000000 tdv200 tmdpvol

* area length vol x angle elev rough dhydr vflag

2000101 2.0 2.0 0.0 0.0 0.0 0.0 0.0 0.0 00000

* cntrl

2000200003

* var p(latm) temp.

2000201 0.0 0.1013e6 0.3024605e3

* component 210 : coolant water flow initiation -> 0.13kg/s

* component 210'

* name type

2100000 tdj210 tmdpjun

78
* from to area

2100101 200000000 240010001 0.00049

* cntrl trip

2100200 1 502

* var waterflow steamnflow x

2100201 -1.0 0.1514 0.0 0.0

2100202 0. 0.1514 0.0 0.0

* component 240: outer tube for coolant water

* name type

2400000 outtube annulus

* nv

2400001 11

* area nv

24001010.005151 11

* length vn

*2400301 0.200 I1

2400301 0.335 1

2400302 0.315 2

2400303 0.290 3

2400304 0.265 4

2400305 0.245 5

2400306 0.220 6

2400307 0.195 7

2400308 0.170 8

2400309 0.145 9

2400310 0.120 10

24003110.100 11

* volume nv

2400401 0.0 11

* angle nv

2400601 90.0 11

* rough hd nv

2400801 1.e-4 0.0 11

* vflag nv

2401001 00010 11

* jflag nj

2401101 000020 10

* cntrl pressure temp. vn

2401201 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 1

2401202 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 2

2401203 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 3

2401204 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 4

2401205 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 5

2401206 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 6

2401207 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 7

2401208 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 8

2401209 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 9

2401210 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 10

2401211 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 11

* cntrl

2401300 1

* waterflow steamflow int.vel nj

2401301 0.1514 0.0 0.0 10

* component 270: outer tube to coolant outlet

* name type

2700000 sj270 sngljun

* from to area kforw kbackw jflag

2700101 240110002 280000000 0.00049 50.0 90.48

0001000

* cntrl waterflow steamflow x

27002011 0.1514 0.0 0.0

* component 280: coolant water dumping

* name type

2800000 tdv28O tmdpvol

* area length vol x angle elev rough dhydr vflag

2800101 2.0 2.0 0.0 0.0 0.0 0.0 0.0 0.0 00000

* cntrl

2800200003

* var p(latm) temp.

2800201 0.0 0.1013e6 0.3024605e3

* heat structure input : pipel40 <-> annulus240

* (condensing tube simulation)

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

* heat structure 140 : heat transfer simulation through pipe

11400000 1142 10.02375

1140010001

11400101 3 0.02540

114002015 3

1140030103

11400401 373.0 4

*11400501 1400200000 101 10.200 1

*11400502 140030000 10000 101 1 0.200 11

79
*114006012401100000 101 10.200 1

*11400602 240100000 -10000 101 10.200 11

11400501 140020000 0 101 10.1001

11400502 1400300000 101 0.1202

11400503 1400400000 101 1 0.145 3

11400504 140050000 0 101 10.1704

11400505 140060000 0 101 10.1955

11400506 140070000 0 101 10.2206

11400507 140080000 0 101 10.245 7

11400508 1400900000 101 0.2658

11400509 140100000 0 101 0.2909

11400510 140110000 0 101 10.31510

11400511 140120000 0 101 0.33511

11400601240110000 0 101 10.1001

11400602 240100000 0 101 10.1202

11400603 2400900000 101 10.1453

11400604 240080000 0 101 10.1704

11400605 240070000 0 1011 0195 5

11400606 2400600000 101 1 0.2206

11400607 240050000 0 101 1 0.245 7

11400608 240040000 0 1011 0.2658

11400609 240030000 0 101 1 0.290 9

11400610 240020000 0 101 I 0.315 10

11400611240010000 0 010.335 01 II

114007010 0.0 0.0 0.0 11

114008010 10.0 10.0 10.01 0.0 0 0 1.0 11

114009010 10.0 10.0 10.0 10.0 0 0 1.0 11

* heat structure thermal property data : SUS (005) 18cr-8ni

20100500 tbl/fctn I 1

* temp(k) thermal conductivity(w/m.k)

20100501 0.273261 le+03 0.1489124e+02

20100502 0.294261 le+03 0.1489124e+02

20100503 0.3109278e+03 0.1505739e+02

20100504 0.3664834e+03 0.1609584e+02

20100505 0.4220389e+03 0.1696813e+02

20100506 0.4775945e+03 0. 1800657e+02

20100507 0.5331500e+03 0.1885809e+02

20100508 0.5887056e+03 0.1956423e+02

20100509 0.6442611 e+03 0.2041575e+02

20100510 0.8109278e+03 0.2297030e+02

20100511 0.9220389e+03 0.2423029e+02

20100512 1.9220389e+03 0.2423029e+02

* temp(k) volumetric heat capacityO/m3.k)

20100551 0.274261 Ie+03 0.3831330e+07

20100552 0.3109278e+03 0.3831330e+07

20100553 0.3664834e+03 0.3985580e+07

20100554 0.4220389e+03 0.4105300e+07

20100555 0.4775945e+03 0.4224090e+07

20100556 0.5331500e+03 0.4308800e+07

20100557 0.5887056e+03 0.4359790e+07

20100558 0.644261 Ie+03 0.4410320e+07

20100559 0.8109278e+03 0.4561910e+07

20100560 0.9220389e+03 0.4625250e+07

20100561 1.9220389e+03 0.4625250e+07

end of file

80
C-2. Base case calculation with alternative model

(Experiments : #12b)

Modification of C-1

1. heat structure input

** ** ***   S.** * * *** •**5*5*5*5* *5****•@

* heat structure input : pipe 140 <-> annulus240

* (condensing tube simulation)

* heat structure 140 : heat transfer simulation through pipe

11400000 114 2 10.02375

1140010001

11400101 3 0.02540

1140020153

11400301 03

11400401 373.0 4

*11400501 140020000 0 153 1 0.200 1

*11400502 140030000 10000 153 1 0.200 11

*

*114006012401100000 153 1 0.200 1

* 11400602 240100000 -10000 153 10.200 11

11400501 1400200000 153 1 0.100 1

11400502 140030000 0 153 10.120 2

11400503 1400400000 153 10.1453

11400504 1400500000 153 10.1704

11400505 140060000 0 153 10.195 5

11400506 140070000 0 153 10.220 6

11400507 140080000 0 153 1 0.245 7

11400508 140090000 0 153 1 0.265 8

11400509 1401000000 153 10.2909

11400510 140110000 0 153 10.315 10

11400511 140120000 0 153 10.335 11

11400601 2401100000 153 10.100 1

114006022401000000 153 10.1202

11400603 240090000 0 153 10.145 3

11400604 240080000 0 153 10.170 4

11400605 240070000 0 153 10.195 5

114006062400600000 153 1 0.2206

11400607 240050000 0 153 1 0.245 7

11400608 240040000 0 153 1 0.265 8

11400609 240030000 0 1531 0.290 9

11400610 240020000 0 153 1 0.315 10

11400611 2400100000 153 10.335 11

114007010 0.0 0.0 0.0 11

114008010 10.0 10.0 10.0 10.00 0 1.0 11

11400901 0 10.0 10.0 10.0 10.0 0 0 1.0 11 *t

81
C-3. Simulation for sensitivity study with default model

(Simulation : #ssO)

Modification of C-1

. time step

-RELAP5/MOD3.2 Simulation

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

*Condensation Experiment in the presence

*of noncondensible gas in a Vertical Tube

*Dec 9. 1997 by HSPARK

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

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

* miscellaneous control cards

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

100 new transnt

101 run

105 10.20.

110 air

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

* time step cards

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

* end min.st max ctrl minor major rstplt

201 10. 1.Oe-8 0.025 3 100 400 400

202 100. 1.Oe-8 0.05 3 500 2000 2000

203 250. 1.Oe-8 0.1 3 2500 10000 10000

*204 5000. 1.Oe-8 0.1 3 5000 20000 20000

*201 10. 1.0e-8 0.00001 3 2000000 10000000 10000000

* original time step

*201 500. 1.Oe-8 0.005 3 2000 10000 10000

* trip logic ****$***$***   ** * *** * ** $*** ***** ***

502 time 0 It null 0. 0. n 0.0

506 time 0 gt null 0. 0. n -1.0

508 quala 150010000 gt null 0. 0.8 n 0

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

* hydrodynamic data: steam-air mixture tube

* component 100 : S/G using time dependent volume

* name type

1000000 tdvl00 tmdpvol

* area length vol x angle elev rough dhydr vflag

1000101 2.0 2.0 0.0 0.0 0.0 0.0 L.e-4 0.0 00000

* noncondensible gas(air)

* cntrl

1000200004

* var pressure temp. eq.quality

1000201 0.0 0.46546675e6 0.4165528e3 1.0

* component 105: steam flow initiation in kg/s

* name type

1050000 tdj 105 tmdpjun

* from to area

* cccOOOO00 cccvvOOOn

1050101 100000000 115010001 0.00049

* cntrl trip

1050200 1 502

* var waterflow steamflow interface vel.

1050201 -1.0 0.0 0.009087 0.0

82
1050202 0 0.0 0.009087 0.0

* component 115 : pipe to upper plenum

* name type

1150000 sv1 15 snglvol

* area length vol x angle

1150101 0.00049 0.07 0.0 0.0 90.

* elev rough dh vflag

1150102 0.07 I.e-4 0.0 00000

* cntrl pressure quality

* 1150200 002 0.46546675e6 1.0

* cntrl pressure temp.

1150200 003 0.46546675e6 0.35474181e3

* component 120: upper plenum

* name type

1200000 bh20 branch

* no. jun cntrl

1200001 2 1

* area=IOcm length vol x angle elev rough dh vflag

1200101 0.007854 0.20 0.00. -90. -0.20 1.e-4 0.0 11000

* cntrl pressure quality

* 1200200 002 0.46546675e6 1.0

* cntrl pressure temp.

1200200 003 0.46546675e6 0.35474181e3

* from to area kforw kbackw jflag

* cccnl0l cccvvOOOn

1201101115010002 120010002 0.000490 225.847 120.438

000000

1202101 120010002 140010001 0.001772 7.608 11.783

000000

* waterflow stearnflow x

" cccn20l

1201201 0.0 0.009087 0.0

1202201 0.0 0.009087 0.0

* component 140: condensing tube

* name type

1400000 pipel40 pipe

* nv: no. vol

140000113

* area nv

1400101 0.001772 13

* length vn

* cccO3OI - ccc0399

1400301 0.500 1

1400302 0.100 2

1400303 0.120 3

1400304 0.145 4

1400305 0.170 5

1400306 0.195 6

1400307 0.220 7

1400308 0.245 8

1400309 0.265 9

1400310 0.290 10

14003110.315 11

1400312 0.335 12

1400313 0.400 13

* ccc03O0 - ccc0399

*1400301 0.500 1

* 1400302 0.200 12

*1400313 0.400 13

* volume nv

*1400401 0.0 13

* angle vn

1400601 -90.0 13

* rough dhydr no. vol

1400801 l.e-4 0.0 13

* vflag no. vol

1401001 00000 13

* jflag no. jun

1401101000000 12

* cntrl pressure temp. nv

*cccl201-ccc1299

1800200003

1401201 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 1

1401202 003 0.4654667506 0.35474181e3 0.0 0.0 0.0 2

1401203 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 3

1401204 003 0.46546675e6 0.3547418163 0.0 0.0 0.0 4

1401205 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 5

1401206 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 6

1401207 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 7

1401208 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 8

1401209 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 9

1401210 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 10

1401211 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 11

1401212 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 12

1401213 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 13

* cntrl

83
1401300 1

* water flow steam flow int.vel jn

14013010.0 0.0 0.0 12

* component 145: condensing tube to lower plenum

* name type

1450000 sj145 sngljun

* from to area kforw kbackw jflag

1450101 140130002 150010001 0.001772 11.783 7.608

0001000

* cntrl waterflow steamflow x

1450201 1 0.0 0.0 0.0

* component 150: lower plenum

* name type

1500000 pipel50 pipe

* nf

1500001 3

* area nv

15001010.007854 3

* length vn

15003010.10 1

1500302 0.10 2

1500303 0.10 3

* volume nv

1500401 0.0 3

* angle nv

1500601 -90.0 3

* rough dhydr nv

1500801 l.e-4 0.0 3

* vflag nv

150100100000 3

* jflag nj

1501101000000 2

* cntrl pressure temp. eq.quality vn

1501201 004 0.46546675e6 0.35474181e3 1.0 0.0 0.0 1

1501202 004 0.46546675e6 0.35474181e3 1.0 0.0 0.0 2

1501203 004 0.46546675e6 0.35474181e3 1.0 0.0 0.0 3

* cntrl

1501300 1

* waterflow steamflow int.vel nj

1501301 0.0 0.0 0.0 2

* hydrodynamic data: venting line

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

* component 155: lower plenum to drain tank

* name type

1550000 sj155 sngljun

* from to areav kforw kbackw jflag

1550101 150030002 160050002 0.00001 923317.24

923317.24 0001000

* cntrl waterflow steamflow x

15502011 0.0 0.0 0.0

* component 156: lower plenum to drain pipe

* name type

1560000 sj156 sngljun

* from to areav kforw kbackw jflag

1560101 150030002 157050002 0.00012 2109.13 4153.80

0001000

* cntrl waterflow steamflow x

15602011 0.0 0.0 0.0

* component 157: drain pipe

* name type

1570000 pipe 157 pipe

* nv

1570001 5

* area nv

1570101 0.00012 5

* length vn

1570301 0.20 5

* volume nv

1570401 0.0 5

* angle nv

1570601 90.0 5

* rough dhydr nv

1570801 1.e-4 0.0 5

* vflag nv

157100100000 5

* jflag nj

15711010000004

* cntrl pressure temp. eq.quality vn

1571201 004 0.46546675e6 0.35474181e3 1.0 0.0 0.0 1

84
1571202 004 0.46546675e6 0.35474181e3 0.5 0.0 0.0 2

1571203 004 0.46546675e6 0.35474181e3 0.0 0.0 0.0 3

1571204 004 0.46546675e6 0.35474181e3 0.0 0.0 0.0 4

1571205 004 0.46546675e6 0.35474181e3 0.0 0.0 0.0 5

* cntrl

1571300 1

* waterflow steamflow int.vel nj

1571301 0.0 0.0 0.0 4

* component 158 : drain pipe to drain tank

* name type

1580000 sj 158 sngljun

* from to areav kforw kbackw jflag

1580101 157010001 160010001 0.00012 42823936.0

21415240.0 0001000

* cntrl waterflow steamflow x

1580201 1 0.0 0.0 0.0

* component 160: drain tank

* name type

1600000 pipel 60 pipe

* nv

1600001 5

* area nv

1600101 0.7854 5

* length vn

1600301 0.20 5

* volume nv

16004010.0 5

* angle nv

160060190.0 5

* rough dhydr nv

1600801 1.e-4 0.0 5

* vflag nv

160100100000 5

* jflag nj

16011010000004

* cntrl pressure temp. eq.quality vn

1601201 004 0A6546675e6 0.35474181e3 1.0 0.0 0.0 1

1601202 004 0A6546675e6 0.35474181e3 0.5 0.0 0.0 2

1601203 004 0A6546675e6 0.35474181e3 0.0 0.0 0.0 3

1601204 004 0.46546675e6 0.35474181 e3 0.0 0.0 0.0 4

1601205 004 0.46546675e6 0.35474181 e3 0.0 0.0 0.0 5

* cntrl

1601300 1

* waterflow steamflow int.vel nj

16013010.00.00.04

* componet 175 : valve for venting

* name type

1750000 vv 175 valve

* from to areav kforw kbackjflag

1750101 150030002 180000000 0.00012 2109.13 4153.80

000100

* cntrl waterflow steamflow interface velocity

1750201 1 0.0 0.0 0.0

* valve type

1750300 trpvlv

* trip number(open when true)

1750301 506

* component 180 : venting simulation using tdv

* name type

1800000 tdvl81 tmdpvol

* area length vol x angle elev rough dhydr vflag

1800101 2.0 2.0 0.0 0.0 0.0 0.0 0.0 0.0 00000

* cntrl

1800200003

* var pressure temp.

1800201 0.0 0.46546675c6 0.35474181e3 *

* hydrodynamic data : coolant water annulus

* component 200 : coolant source simulation using tdv

* name type

2000000 tdv200 tmdpvol

* area length vol x angle elev rough dhydr vflag

2000101 2.0 2.0 0.0 0.0 0.0 0.0 0.0 0.0 00000

* cntrl

2000200003

* var p(latm) temp.

2000201 0.0 0.1013e6 0.3024605e3

85
* component 210: coolant water flow initiation-> 0.13kg/s

* component 210 *

* name type

2100000 tdj2IO tmdpjun

* from to area

2100101 200000000 240010001 0.00049

* cntrl trip

2100200 1 502

* var waterflow steamflow x

2100201 -1.0 0.1514 0.0 0.0

2100202 0. 0.1514 0.0 0.0

* component 240 : outer tube for coolant water

* name type

2400000 out-tube annulus

* nv

2400001 11

* area nv

24001010.005151 11

* length vn

*24003010.200 11

2400301 0.335 1

2400302 0.315 2

2400303 0.290 3

2400304 0.265 4

2400305 0.245 5

2400306 0.220 6

2400307 0.195 7

2400308 0.170 8

2400309 0.145 9

2400310 0.120 10

24003110.100 11

* volume nv

24004010.0 11

* angle nv

240060190.0 11

* rough hd nv

2400801 1.e-4 0.0 11

* vflag nv

240100100010 11

* jflag nj

2401101000020 10

* cntrl pressure temp. vn

2401201 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 1

2401202 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 2

2401203 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 3

2401204 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 4

2401205 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 5

2401206 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 6

2401207 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 7

2401208 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 8

2401209 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 9

2401210 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 10

2401211 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 11

* cntrl

2401300 1

* waterflow steamflow int.vel nj

2401301 0.15140.0 0.0 10

* component 270: outer tube to coolant outlet

* name type

2700000 sj270 sngljun

* from to area kforw kbackw jflag

2700101 240110002 280000000 0.00049 50.0 90.48

0001000

* cutrl waterflow steamflow x

2700201 10.1514 0.0 0.0

* component 280: coolant water dumping

* name type

2800000 tdv280 tmdpvol

* area length vol x angle elev rough dhydr vflag

2800101 2.0 2.0 0.0 0.0 0.0 0.0 0.0 0.0 00000

* cntrl

2800200003

* var p(I atm) temp.

2800201 0.0 0.1013e6 0.3024605e3

* heat structure input : pipel40 <-> annulus240

* (condensing tube simulation)

* heat structure 140 : heat transfer simulation through pipe

11400000 11 4 2 1 0.02375

114001000 1

11400101 3 0.02540

86
1140020153

1140030103

11400401 373.0 4

* 11400501 140020000 0 101 10.200 1

*11400502 140030000 10000 101 10.200 11

,

* 11400601 2401100000 1011 0.200 1

* 11400602 240100000 -10000 101 10.200 11

11400501 1400200000 101 10.1001

11400502 1400300000 101 10.1202

11400503 1400400000 101 10.1453

11400504 1400500000 101 10.1704

11400505 1400600000 101 10.1955

11400506 1400700000 101 10.2206

11400507 140080000 0 101 10.2457

11400508 140090000 0 101 10.2658

11400509 140100000 0 101 0.2909

11400510 1401100000 101 10.31510

11400511 1401200000101 10.33511

11400601240110000 0 101 10.1001

114006022401000000 10110.1202

11400603 240090000 0 101 10.145 3

114006042400800000 10110.1704

11400605 240070000 0 101 10.195 5

114006062400600000 101 10.2206

11400607 240050000 0 101 10.2457

11400608 240040000 0 101 10.2658

114006092400300000 101 10.2909

114006102400200000 101 10.315 10

11400611 2400100000 101 10.33511

114007010 0.0 0.0 0.0 11

11400801 0 10.0 10.0 10.0 10.0 0 0 1.0 11

114009010 10.0 10.0 10.0 10.0 0 0 1.0 11

*

*

* heat structure thermal property data: SUS (005) 18cr-8ni

20100500 tbl/fctn I I

* temp(k) thermal conductivity(w/m.k)

20100501 0.2732611 e+03 0.1489124e+02

20100502 0.2942611 e+03 0.1489124e+02

20100503 0.3 109278e+03 0.1505739e+02

20100504 0.3664834e+03 0. 1609584e+02

20100505 0.4220389e+03 0.1696813e+02

20100506 0.4775945e+03 0.1800657e+02

20100507 0.5331500e+03 0.1885809e+02

20100508 0.5887056e+03 0.1956423e+02

20100509 0.6442611 e+03 0.2041575e+02

20100510 0.8109278e+03 0.2297030e+02

20100511 0.9220389c+03 0.2423029e+02

20100512 1.9220389e+03 0.2423029e+02

* temp(k) volumetric heat capacity(j/m3.k)

20100551 0.274261 le+03 0.3831330e+07

20100552 0.3 109278e+03 0.3831330e+07

20100553 0.3664834e+03 0.3985580e+07

20100554 0.4220389e+03 0.4105300e+07

20100555 0.4775945e+03 0.4224090e+07

20100556 0.5331500e+03 0.4308800e+07

20100557 0.5887056e+03 0.4359790e+07

20100558 0.644261 Ie+03 0.4410320e+07

20100559 0.8109278e+03 0.4561910e+07

20100560 0.9220389e+03 0.4625250e+07

20100561 1.9220389e+03 0.4625250e+07

end of file

87
C4. Simulation for sensitivity study with alternative model

(Simulation : #ssO)

Modification of C-3

1. heat structure input

**   ** * .** ** **********   ** **   ** *

* heat structure input: pipet40 <-> annulus240

* (condensing tube simulation)

* heat structure 140: heat transfer simulation through pipe

11400000 11421 0.02375

1140010001

11400101 3 0.02540

11400201 53

1140030103

11400401 373.0 4

*11400501 1400200000 153 10.200 1

*11400502 140030000 10000 153 1 0.200 11

*114006012401100000 153 10.200 1

*11400602 240100000-10000 153 10.200 11

11400501 140020000 0 153 1 100.100 

11400502 140030000 0 153 10.120 2

11400503 140040000 0 153 10.145 3

11400504 1400500000 153 10.1704

11400505 1400600000 153 10.195 5

11400506 1400700000 153 10.2206

11400507 1400800000 153 10.245 7

11400508 1400900000 153 10.2658

11400509 1401000000 153 10.2909

11400510 1401100000 153 10.315 10

11400511 1401200000 153 10.335 11

114006012401100000 153 1 0.100 1

11400602 240100000 0 153 10.120 2

11400603 2400900000 153 10.145 3

114006042400800000 153 10.1704

11400605 240070000 0 153 10.195 5

11400606 240060000 0 153 10.220 6

11400607 240050000 0 153 1 0.245 7

11400608 240040000 0 153 1 0.265 8

11400609 240030000 0 1531 0.290 9

114006102400200000 153 10.315 10

114006112400100000 153 1 0.335 11 *

114007010 0.0 0.0 0.0 11

11400801 0 10.0 10.0 10.0 10.0 0 ) 1.0 11

11400901 0 10.0 10.0 10.0 10.000 1.0 11

88
C-5. Simulation for different number of nodes with default model

(Simulation : #nd4)

Modification of C-3

1. component 140 and 145

2. component 240 and 270

3. heat structure input

4. simular case: ndl, nd2, nd3

* component 140: condensing tube

* name type

1400000 pipel40 pipe

* nv: no. vol

1400001 18

* area nv

1400101 0.001772 18

* length vn

* ccc0301 - ccc0399

1400301 0.500 1

1400302 0.150 17

1400303 0.400 18

* volume nv

*1400401 0.0 18

* angle vn

1400601 -90.0 18

* rough dhydr no. vol

1400801 l.e-4 0.0 18

* vflag no. vol

1401001 00000 18

* jflag no. jun

1401101000000 17

* cntrl pressure temp. nv

*cccl201-cccl299

1800200003

1401201 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 1

1401202 003 0.46546675e6 0.35474181e3 0.0 0.0 0.0 18

* cntrl

1401300 1

* water flow steam flow int.vel jn

1401301 0.0 0.0 0.0 17

* component 145 : condensing tube to lower plenum

* name type

1450000 sj145 sngljun

* from to area kforw kbackw jflag

1450101 140180002 150010001 0.001772 11.783 7.608

0001000

* cntrl waterflow stearnflow x

1450201 1 0.0 0.0 0.0

* component 240: outer tube for coolant water

* name type

2400000 out-tube annulus

* nv

2400001 16

* area nv

2400101 0.005151 16

* length vn

2400301 0.150 16

* volume nv

2400401 0.0 16

* angle nv

2400601 90.0 16

* rough hd nv

2400801 l.e-4 0.0 16

* vflag nv

2401001 00010 16

* jflag nj

2401101000020 15

* cntrl pressure temp. vn

2401201 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 1

2401202 003 0.1013e6 0.3024605e3 0.0 0.0 0.0 16

89
* cntrl

2401300 1

* waterflow steamflow int.vel nj

24013010.1514 0.0 0.0 15

* component 270 : outer tube to coolant outlet

* name type

2700000 sj270 sngljun

* from to area kforw kbackw jflag

2700101 240160002 280000000 0.00049 50.0 90.48

0001000

* cntrl waterflow steamflow x

2700201 1 0.15140.0 0.0

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

* heat structure input: pipe 140 <-> annulus240

* (condensing tube simulation)

* *** ** *** ** * *****$  **  *$* * **$***** ***

$

* heat structure 140: heat transfer simulation through pipe

11400000 16 4 2 1 0.02375

1140010001

11400101 3 0.02540

1140020153

1140030103

11400401 373.0 4

*11400501 140020000 0 101 10.150 1

* 11400502 140170000 10000 101 1 0.150 16

*11400601 2401600000 1011 0.150 1

*11400602 240010000-10000 101 10.150 16

11400501 1400200000 101 10.1501

11400502 1400300000 101 0.1502

11400503 1400400000 101 0.1503

11400504 140050000 0 101 10.1504

11400505 140060000 0 101 10.1505

11400506 1400700000 101 10.1506

11400507 1400800000 101 10.1507

11400508 1400900000 101 10.1508

11400509 1401000000 101 10.1509

11400510 140110000 0 101 10.15010

11400511 1401200000 101 0.150 11

11400512 1401300000 101 10.150 12

11400513 140140000 0 101 10.15013

11400514 140150000 0 101 10.15014

11400515 1401600000 101 10.15015

11400516 140170000 0 101 10.15016

11400601*2401600000 101 10.1501

11400602 2401500000 101 1 0.150 2

11400603 240140000 0 101 10.1503

114006042401300000 101 10.1504

11400605 2401200000 101 10.1505

114006062401100000 101 10.1506

11400607 240100000 0 101 10.1507

11400608 2400900000 101 10.1508

11400609 240080000 0 101 10.1509

11400610 2400700000 101 1 0.15010

114006112400600000 101 10.15011

11400612 2400500000 101 1 0.15012

11400613 2400400000 101 10.15013

11400614 240030000 0 101 1 0.150 14

11400615 240020000 0 101 1 0.150 15

11400616 2400100000 101 1 0.150 16

114007010 0.0 0.0 0.0 16

11400801 0 10.0 10.0 10.0 10.0 0 0 1.0 16

114009010 10.0 10.0 10.0 10.0 0 0 1.0 16

90
C -6. Simulation for different number of nodes with alternative

model (Simulation : #nd4)

Modification of C-5

1. heat stlructure input

* heat structure input :pipel40 <-> annulus240

* (condensing tube simulation)

*heat structure 140 :heat transfer simulation through pipe

11400000 164 2 10.02375

1140010001

11400101 3 0.02540

1140020153

11400301 03

11400401 373.0 4

* 11400501 140020000 0 153 10.150 1

* 11400502 140170000 10000 153 1 0.150 16

*11400601 240160000 0 153 10.150 1

*11400602 240010000 -.10000 153 1 0.150 16

11400501 140020000 0 153 10.150 1

11400502 140030000 0 153 1 0.150 2

11400503 140040000 0 153 10.150 3

11400504 140050000 0153 10.150 4

11400505 140060000 0153 10.150 5

11400506 140070000 0 153 1 0.150 6

11400507 140080000 0153 10.150 7

11400508 140090000 0153 10.150 8

11400509 140100000 0153 10.150 9

11400510 140110000 0 153 1 0.150 10

11400511 140120000 0153 10.150 11

11400512 140130000 0153 1 0.150 12

11400513 140140000 0 153 1 0.150 13

11400514 140150000 0153 1 0.150 14

11400515 140160000 0153 10.150 15

11400516 140170000 0153 10.150 16

11400601 240160000 0 153 1 0.150 1

11400602 240150000 0153 1 0.150 2

11400603 240140000 0153 10.150 3

11400604 240130000 0153 10.150 4

11400605 240120000 0153 10.150 5

11400606 240110000 0153 10.150 6

11400607 240100000 0153 10.150 7

11400608 240090000 0153 10.15089

11 400609 240080000 0153 10.150 9

11400610 240070000 0153 10.150 10

11400611 240060000 0153 10.150 11

11400612 240050000 0153 1 0.150 12

11400613 240040000 0 153 10.150 13

11400614 240030000 0153 10.150 14

11400615 240020000 0153 10.150 15

1 1400616 240010000 0153 10.150 16

114007010 0.0 0.0 0.0 16

1 1400801 0 10.0 10.0 10.0 10.0 00 1.0 16

114009010 10.0 10.0 10.0 10.0 0 01.0 16

91

NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

(24) (Assgne by NRC. Add Val. SuIp. Rev.

NRCM 110Z and Addendum NumbersI alny.) 32D.320 BIBLIOGRAPHIC DATA SHEET

2. TITLE AND SUBTITLEI NUREGAA-0147

Assessment of RELAP5/MOD3.2 for Steam Condensation Experiments in the Presence of

Noncondenslbles in a Vertical Tube of PCCS 3. DATE REPORT PUBLISHED

MONTHYEAR

September 1998

4. FIN OR GRANT NUMBER

D6227

S. AUTHOR(S) 6. TYPE OF REPORT

H. S. Park, H. C. No/IKAIST

Y. S. Bang, K W. Seul, H. J. Kim/KINS

7. PERIOD COVERED alnctusive Deas)

8 PERFORMNG ORGANIZATION - NAME AND ADORESS (NRCA ,rovide Div&iovx ocre rRfgtx U.S. Nuabew Reg to~rymiw and nmalig adoews, ifoonb,

prmv nme and m=&Ving add••sj

Korea Advanced Institute of Science and Technology Korea Institute of Nuclear Safety

373-1 Kuseong-Dong, Yusong-Ku P.O. Box 114

Taejon, Korea Yusong, Taejon

Korea 305-600

g. SPONSORING ORGANIZATION - NAME AND ADDRESS pWC, M Sam. as abov.' foon-uckc, povide NRC =vio. Orice or Rgion U.S. Ncke.- i•gu.by Conma,- &dW ma' addwssj

Division of Systems Technology

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

10. SUPPReMNENTARY NOTES

G. Rhee, NRC Project Manager

11. ABSTRACT M2W0 war* alrm)

This report deals with the application of RELAPSIMOD3.2 to condensation experiments in the presence of noncondensable gases

In a vertical tube of Passive Containment Cooling System. When steam-noncondensable gas mixture was injected into the vertical

tube, steam was condensed on the Inner surface of the condensing tube but the noncondensable gas greatly inhibited the

condensation of the steam. As the scattering of previous experimental data was large, the present experimental appparatus was

set up to get reliable data on the condensation heat transfer coefficient of the steam-noncondensable gas mixture in a vertical tube.

The experimental results show that the condensation heat transfer coefficient increases as the inlet steam-air mixture flow rate

increases, the inlet air mass fraction decreases, and the Inlet saturated steam temperature decreases.

There are two wall film condensation models, the default model and the alternative model, in RELAP5/MOD3.2. After a

condensation database was constructed, two models were assessed directly with the data of the database. The experimental

apparatus was also modeled with RELAP5/MOD3.2, and simulations were performed for several sub-tests to be compared with

the experimental results. The simulation results show that in an overall sense the default model of RELAPSiMOD3.2

under-predicts the heat transfer coefficients, but that the alternative model of RELAP5/MOD3.2 over-predicts them throughout the

condensing tube.

12. KEY WORDSIDESCRIPTORS (Lt wordds rrases fmt wwiasseisars in kmcatn "impoij I&. AVAZLABLI•Y STATEMENT

steam condensation unlimited

condensation Inside tubes 14. SECURI1YCUASSIFAON

non-condensible gas effect Mis Poe)

RELAP5/MOD3.2 unclassified

assessment MTs R•opO

passive containment system unclassified

15. NUMBER OF PAGES

16. PRICE

NRC FORM 335 (2-W
Federal Recycling Program

NUREGIIA-0147 ASSESSMENT OF RELAPS/MOD3.2 FOR STEAM CONDENSATION EXPERIMENTS

IN THE PRESENCE OF NONCONDENSIBLES IN A VERTICAL TUBE OF PCCS

SEPTEMBER 1998

UNITED STATES

NUCLEAR REGULATORY COMMISSION

WASHINGTON, DC 20555-0001

SPECIAL STANDARD MAIL

POSTAGE AND FEES PAID

USNRC

PERMIT NO. G-67

OFFICIAL BUSINESS

PENALTY FOR PRIVATE USE, $300