NUREG/IA-0025

International

Agreement Report

RELAP5/MOD3 Subcooled Boiling

Model Assessment

Prepared by

A. S. Devkin and A. S. Podosenov

Nuclear Safety Institute

Russian Research Centre

"Kurchatov Institute"

Kurchatov Square, 1

123182, Moscow

Russia

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

May 1998

Prepared as part of

The Agreement on Research Participation and Technical Exchange

under the International Code Application and Maintenance Program (CAMP)

Published by

U.S. Nuclear Regulatory Commission

Ol(1

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I

NUREG/IA-0025 RELAP5/MOD3 SUBCOOLED BOILING MODEL ASSESSMENT MAY 1998
A
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NUREG/IA-0025

International

Agreement Report

RELAP5/MOD3 Subcooled Boiling

Model Assessment

Prepared by

A. S. Devkin and A. S. Podosenov

Nuclear Safety Institute

Russian Research Centre

"Kurchatov Institute"

Kurchatov Square, 1

123182, Moscow

Russia

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

May 1998

Prepared as part of

The Agreement on Research Participation and Technical Exchange

under the International Code Application and Maintenance Program (CAMP)

Published by

U.S. Nuclear Regulatory Commission
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ABSTRACT

This report presents the assessment of the RELAP5/Mod3 (5m5 version)

code subcooled boiling process model, which is based on a variety of

experiments. The accuracy of the model is confirmed for a wide range of

regime parameters for the case of uniform heating along the channel. The

condensation rate is rather underpredicted, which may lead to considerable

errors in void fraction behavior prediction in subcooled boiling regimes for

nonuniformly or unheated channels.

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Contents

Page 1 was intentionally omitted.

1. Introduction .................................................................................................... 2

2. M odel description ........................................................................................... 3

2.1 Bubble Regim e ...................................................................................... 5

2.2 Slug regim e ........................................................................................... 6

2.3 W all heat transfer ................................................................................ 7

2.4 V apor generation rate ............................................................................ 7

3. Analysis of RELAP5/MOD3 subcooled boiling model results ......................... 9

3.1 Tests description .................................................................................. 9

3.2 Comparison of calculational and experimental data ............................ 9

4. Sensitivity analysis ......................................................................................... 18

5. Run statistics ....................................................................................................... 18

6. Conclusions .................................................................................................... 19

References ...................................................................................................................... 20

APPENDICES

1. Input deck for BARTOLOMEY experiment ........................ 1.1

2. Figures ................................................................................................................ 2.1

V
,I
2

1. INTRODUCTION

The flow in the core for some reactors such as RBMK or BWR is in two -

phase conditions in nominal regimes. Nonboiling reactors such conditions may

exist in the hottest channels, as for WWER-1000. For accidental regimes fluid

boiling in the core appears practically for any of transients scenarios.

Calculational analysis of these processes are fulfilling by using the codes having

the capability for accurate description the appearing and behavior of vapor

phase in reactor channels . The description of such processes in the "best

estimate codes" is based on two-fluid models, which gives the possibility to

describe the process of boiling in the core accurate enough for arbitrary

distribution of power along the channels and for any transients. It being known

that the description of saturated boiling process don't give rise to difficulties.

But as for subcooled boiling and vapor condensation processes the situation is

more complicated because it needs to describe such processes as vapor

appearance and it's generation at the heated walls , vapor condensation in the

subcooled water, interface heat and mass transfer and so on.

This work purpose is the evaluation of the accuracy of the models used in

RELAP5/MOD3 (version 5m5) code [1] for subcooled boiling process by

comparison of the calculational and experimental date in wide range of regime

parameters.

I
3

2. MODEL DESCRIPTION

Governing equations system used in RELAP5/MOD3 code are as following.

Mass equations:

FQagpg) + -L"2-(ccgpgVgA) = Fg, (1)

F(ac pf) + x(czfpf VfA) = -Fg, (2)

where CCg = 1 - af - vapor void fraction ; pg, pf - specific densities of vapor

and liquid phases; Vg, VY - phase velocities , A - cross section area; rg-vapor

generating rate which consists of two parts - volume generation (interface) rate

- rig and wall generation rate - rw , such as

Fg = Fig + Fw (3)

Energy equations

a(agpgUg) + "@XgpgUgVg) g - 7-(ccgVgA) + Qwg + Qig+

+Figh* + Fwhg +DISSg (4)

F( c-pfaUf) + ±(cpfUf Vf) =(PcV - AjCcfVfA) + Qwf + Qif -

-Fifh - Fwh * + DISSf , (5)

where Uk=g,f - specific internal energy of k-phase, P - pressure, Qwkk=g,f) -

specific heat flux from wall to "k"-phase; Qik (k=g,f) - interface heat flux

ig k - interface latent heat rw hk - wall latent heat, DISS k - wall

friction dissipation

The momentum 11 equations are not considered here , because the relative

motion of phases are minor for subcooled boiling process . It should be noted

that the phase velocities are determined by using the reliable enough

IIJ
4

correlations, especially for upward flows in channels. Vapor generation rate in

the two-phase volume is determined as:

H ig(T.s-Tg)+H if(T.s-Tf)

F, * , (6)

where "s" - is related for saturated conditions ,Tk- "k" - phase temperature

* 9 *

hg = hg - for condensation and hg = hg - for vaporization,

* *

hf = hfS - for condensation and hf = hf - for vaporization.

Interface heat fluxes are determined as :

Qig= Hig (Ts-Tg)+ ((1- c) / 2) rw ( hgs - hfS), (7)

Qif= Hif (Ts-Tg)+ ((1+ c) / 2) Tw ( hg5 - hfs) (8)

where =l, if rw>0 and =-1 if rw<0.

Wall generation rate is equal

=Qig+Qij+]Fig (h*-h.

,(9)

Such way the net vapor generation rate is determined as I

Hig(Ts-Tg) + Hif(Ts-Tf)

- , - h *(10)

Hig and Hif in this expression are the products of interface area value Ai and

heat transfer coefficient hik from interface to phase "k" Hik = A hik, which

are depending from the two-phase flow regime: bubble, slug, annular and so on.

Here we not consider the whole spectrum of regimes but only specific for

subcooled boiling process : bubbly and slug.

J
5

2.1 BUBBLE REGIME

If the flow is in a bubble regime and the fluid temperature is below the

saturation, the interface heat transfer coefficient is calculated using Unal

formula [20].

hif = (C F hfg* d)/ (2 * (1/pg - 1/Pf)), (11)

where hfg = hg s - hf s, P k=g, f - specific densities of phases,

F = 1, for Vf < 0.61 m/s, (12)

Vf 10O.47

F =0 1" , for Vf > 0.61 m/s , (13)

C = 65 - 5.69 * 10 5(p -10 5), for 105 < P < 106 Pa,

C =0.25 *101 0*P'1418 ,for 106<P<17.7*10 6Pa,

d - bubble diameter, which is calculated using critical Weber number

Wecrit= P f (V-Vf )2 *dmax / a = 10, (14)

where a - surface tension coefficient.

This expression gives the maximal bubble diameter. Mean bubble diameter is

determined as do= dmax/ 2 , and interface area is

Ai = 3 .

6 ag / do (15)

Using this expression and Unal's formula (11 ) one can obtain

Hif = hif A, = 1.8 ag C F hfg pf pg / ( pf - pg ) (16)

This expression used in the code to calculate the heat transfer between interface

surface and subcooled liquid. The field of parameter recommended to use Unal's
6

formula is : pressure P= 0.1 - 17.7 MPa , heat flux q= 0.47 - 10.64 MWt/m2 ,

liquid velocity Vf = 0.9 - 9.15 m/s , liquid subcooling AT= 3 - 86 K , maximal

bubble diameter d= 0.08-1.24 mm.

The heat transfer with vapor phase in subcooled boiling regime does not

introduce significant influence on the process. It should be noted only that large

values of interface heat transfer coefficient on the vapor side ensure the vapor

conditions closed to saturation.

2.2 SLUG REGIME

Interface heat flux (for volume unit) is determined as

hsAsAT hbAbAT

V + V (17)

where index "s" concerned to slugs, and "b"- to bubbles,

h - interface heat transfer coefficient,

A - interface surface area.

Interface surface area for slugs regime is determined from the expression

Ds =0.88D, where D is the hydraulic diameter and As= 4/Ds = 4.5/D. Volume

fraction for slugs is :

Ccgs= ( ag- a bub )/(1 -a bub), (18)

where a bub- void fraction of small bubbles in liquid bridges and near the wall

which is determined as:

a bub= a bs exp [ -8 *(ag - abs)/(asa- abs)], (19)

where abs- void fraction for the bubble-slug transition, asa- void fraction for

slug-annular transition.

Liquid side interface heat flux equals to:

= 1.18942 * 0.5, Pr 0.5 QifS 1192Ref Pf *(kf/D) *Asabub *(Ts -Tf ), (20)
7

where Prf =Cpf pf /kf, Ref =pf D*min[Vf -Vg ;0.8]/Jf ,

[if, kf - coefficients of dynamic viscosity and heat conductivity of liquid.

Interface heat flux for bubbles is determined by the same way as for bubble

regime taking into account that interface area is equal to

Ab =3.6 abub (1- as) /d 0 , (21)

2.3. WALL HEAT TRANSFER.

All heat flux from the wall for subcooled boiling process is consumed to vapor

generation and liquid heating , so Qwg =0 and heat transfer coefficient is

determined by modified Chen correlation.

2.4. VAPOR GENERATION RATE

Vapor generation rate on the wall is calculated as

rw=qwf Aw X/(V (hgs-hf)), (22)

where q wf - heat flux from wall to liquid phase , Aw - heated surface of cell

with volume V, X - vapor generation fraction of the wall heat flux . This fraction

is:

X= ( hf - hb ) / ( (hfs - hb) (1+ s)), (23)

where

c= pf ( hfs - hb) / (Pg hfg), (24)

hb- the critical enthalpy , which is computed using Saha-Zuber formula:

hb= hfs - St Cpf / 0.0065, at Pe > 70*104, (25)

hb= hfs - Nu Cpf / 4.45, at Pe< 70*10 , (26)

I
8

St =Nu /Pe, Nu=qwf De /kf , Pe =G DeCPf /kf,

De - heated equivalent diameter, G - mass flux, Cpf, kf - specific heat capacity

and heat conductivity of fluid.
9

3. ANALYSIS OF THE RELAP5/MOD3 SUBCOOLED BOILING

MODEL

3.1 TESTS DESCRIPTION

All tests chosen for the comparison with calculational results were simple

enough. They were carried out at steady state conditions and for simple

geometry as a rule. The majority of this works were fulfilled with using the

round tube as a test section [2,5-7,18]. One work was chosen to evaluate the code

capability for some exotic case as [19] for very narrow flat test section. The

main parameter in this tests to be compared with calculational data is the void

fraction distribution along the channel or the dependence of void fraction via

equilibrium quality. The accuracy of void measurement for all experiments is

near equal, the method used for them is identical - y -beam absorption method

with some variations. In references 2, 4, 5 and 18 the wide beam was used that

demands careful graduating. A narrow beam was used in reference 19, that

allows to get a void distribution in the cross section of the channel and a more

accurate mean value of void fraction. The absolute value of the void fraction

error is less than 0.04 for all tests used. The other errors of this tests concerned

to accuracy of the measurements of the regime parameters: pressure, inlet

temperature, heat and mass fluxes. For example maximal values of regime

parameters errors have been estimated in [5] as following:

AT =2K,

SG= 0.02,

P= 0.03,

8 q = 0.01

3.2 Comparison of calculational and experimental data

Subcooled boiling process could take place in the reactors of various type in

wide enough range of regime parameters, especially in transient conditions.

Therefore the assessment of the model have to be checked in wide range of

parameters too and by using the large amount of the experimental data.

The results of the investigations of this process could be found in many works.

A careful study of subcooled boiling was made in Russia , for example see

reference [5] , where the experimental data about void fraction distribution

along the round uniformly heated tube inlet diameter of 12 mm were presented.
10

The nodalization scheme for base case calculations for the tests described in

[7] is presented at fig. 51. It consists of one element "pipe" devided in 14

subvolumes 0.1 m length each, two elements "tmdpvol" - to set the conditions at

the inlet - liquid temperature and outlet of the pipe - pressure. Inlet flowrate

was set at the element "tmdpjun" connecting the inlet "tmdpvol" and "pipe".

Upper "tmdpvol" is connected with "pipe " through the element "sngljun". Heat

structure is connected with element "pipe" and it is devided in 14 parts. First 10

parts have the internal heat sources and last 4 parts are without heating.

The calculation have been performed at "transient" mode until all parameters

were not changed in time.

The experimental (from [5] ) and calculational results for high pressures and

mass fluxes are adduced on figures 1-6. The accordance between calculational

and experimental data is good enough, but it must be noted that some

underprediction of the void fraction on fig. 5-6 and some wrong account of heat

flux influence in the model (fig.1 and fig.5). This test series have been fulfilled

with G and P near equal but with different levels of heat flux, the latest being

higher the discrepancies were higher too.

Similar data for P=11 MPa are submitted on figures 7-10. Coordination of

calculational and experimental data is very good even for low flow rates (G =500

kg/m's, fig.7).

The results for P=7MPa and G=960-998kg/m's at various heat fluxes

q=440-1980 KWt/m 2 are presented at fig. 11-15 and the fig. 16 is the summary

schedule for this series of the experiments. Obviously that for these parameters

the RELAP's model is enough truthfully reflects the parameters influence on the

void fraction behavior.

Data presented on fig. 17-21 illustrate the mass flux influence on void

distribution along the channel (that is equivalent to the void fraction dependence

from equilibrium quality in the case of uniform heat flux distribution along the

channel ). Obviously that the sharp discrepancy between calculational and

experimental data are presented at low flow rates G=405kg/m 2s (fig.18).

Fig. 22-24 illustrate the influence of pressure on void distribution with fixed

values of mass flux G = 990 kg/m's [5]. Here should be paid attention on that at

small voids the calculational data are little below than experimental ones and

that the calculational curve has an a break at quality equals to 0.05 , which is

caused by increasing of vapor drift at transition from slug to annular flow.
11

The analogous behavior of the void fraction have been received by using the

experimental results from [4] at P= 6.8 MPa, G=419kg/m 2 s and q = 443

KWt/m 2.

Figures 25 and 26 illustrate the void fraction behavior as Voidg(quality) and

Voidg(z), where z is axial coordinate of the channel. Note here that the break of

the curve was not fixed in the experimental data. At more higher mass fluxes

(G = 962kg/m's) the transition between slug and annular regimes could be

identified at quality equals to 0.0 as in calculations, as in experiments , though it

was expressed very weak (fig.27).

The comparison of calculational and experimental data for very low flow rates

is showed on fig.36- 39. Obviously that the calculational results for all considered

experiments from [2] are some higher than the experimental ones the

discrepancy being more -depending on the inadequate description of relative

motion of phases.

On fig.38 are adduced (dashed line ) the calculational results with using the

homogenous equilibrium model . The area of equilibrium boiling takes the large

area and significant distinction between calculational (homogeneous model) and

experimental data in it is stipulated only by the relative motion of phases. And,

as it is visible from fig. 36-36, calculational values of vapor drift is some less,

than experimental.

The region of subcooled boiling takes very insignificant area in this tests and

the distinction between calculational and experimental data is caused only by the

reliability of interface friction model. Obviously that vapor drift according to

RELAP5/MOD3 model is considerably underestimated and more real values for

void fraction one could get only with using the dependencies for annular-mist

regime. It testifies in our opinion about the necessity of some updating of flow

regimes map, in particular, for transition between slug and annular-mist regimes

at low pressures.

As it is known the velocity difference between vapor and liquid phases

increased at low pressures. Therefore it is rather interesting to evaluate the

reliability of code models at a very low pressures. The comparison of

calculational and experimental data for P= 1 MPa are adduced on fig.40-42.

Obviously that the RELAP5/mod3 technique gives strongly overestimated

results for these parameters. Especially it is visible on fig.42.

The results of calculations and experimental data from [18] for very low

pressures are presented on fig.43-45 . Also the conditions of this experiments

were the following: geometry of the test section - 0.5 m length annular pipe
12

with heated inner rod 7mm diameter and outer diameter 13 mm, pressure near

the atmospheric P = 1.128 bar, mass flux - G= 1416 kg/m 2/s, heat flux -

q=885KWt/m 2 . The results are presented as the void fraction dependence from

the fluid subcooling (fig.43). One can see that there is a very large discrepancy

between the Relap and experimental data for this conditions.

The analysis of the experimental data showed that in such conditions it is

very important to set the boundary conditions for pressure. In spite of the little

length of the test section the outlet - inlet pressure difference is the same order

as the pressure. Therefore the value of pressure given in [18] may be set at the

inlet of the channel or at its outlet. The results of this calculations are shown on

figure 44 .The void distribution along the channel is rather differed from each

other , but for the dependence of the void fraction from equilibrium quality we

have the same curves for both cases (fig. 45). That is why one must be careful

when using the low pressure data and choosing of their presentation method.

It is interesting to evaluate the trustworthy of the model in some exotic

geometry as used in the [19] . It was used there a very narrow flat channel 50

* 2 mm. The method enabled the distribution of the local void fraction in a cross

section to be measured at about 100 locations (along 2 mm) and from these local

values it was possible to determine accurate mean void fraction values. Fig. 46

demonstrates the calculational and experimental results for one of the tests with

conditions : pressure P= 1411.85 bar , mass flux G =750 kg/m 2/s, heat flux q=0.4

MWt/m'. The different curves on this figure demonstrate the dependence of the

calculational results for different number of cells along the channel (total length

is 1.5 m) and the size of the circles around the experimental points corresponds

to the experimental error of determining of mean value of void fraction. One can

see that the accordance between the experimental and calculational data is good

enough for such geometry too, but the Relap's data are some lower then

experimental ones.

Returning to the results of [4] one can use them to evaluate also the ability

of model to take into account the influence of fluid subcooling at the channel

inlet at other fixed parameters and the influence of non - uniformity of heat flux

distribution along the channel also.

The calculational and experimental results for three tests at inlet

temperatures T=221,240 and 255 C and for increased along the channel heat flux

under the low q(z)=0.397+0.801*z, P=4.4 MPa, G =1000 kg/m2/s (average heat

flux equal to q=436 KWt/m2) are presented on fig.28. The RELAP model takes

into account the influence of inlet temperature enough well, however as in
13 a.

previous cases gives sharp change of vapor drift at transition to annular-mist

flow.

The results for test with decreasing along the channel heat flux q=430

KWt/m2 and q=796.5KWt/m2 are presented on figures 29 and 30. One can see

that for the test with more intensive decreasing of heat flux the calculational

data being higher than experimental ones. It is obviously that the reason of this

discrepancy is the nonadequate description of vapor phase condensation process

in the subcooled boiling region with low heat fluxes.

As described in chapter 1 the rate of vapor generation according to

RELAP5/MOD3 model is determined in main by two processes: the rate of

vapor generation on the channel walls and vapor condensation in the subcooled

liquid. Good enough tuning of this two values can give good results, that

however does not mean that' each of this processes is described enough precisely.

The analysis of results , adduced on fig. 29 -30 permits to assume , that the

RELAP5/mod3 model gives the underestimated rates of condensation

Therefore it is rather interesting to evaluate the reliability of model at absence

of vapor generation on the wall.

On fig.31-32 are presented the calculational and experimental results [7],

received at research of void fraction behavior in the pipe by general length 1.5

m at step change of heat flux q=1200 kWt/m2 on length from 0 up to 1.0 m and

q= 0.0 at last 0.5m. Tests were performed with different values of subcooling at

the channel outlet, so for test presented on fig.32 the conditions of flow at the

outlet of the channel were such that the condensation of vapor does not occur.

One can see from these figures that the discrepancy between experimental

and calculational data is especially for nonheated part of the channel, and in all

considered experiments the calculational rate of void fraction decreasing is less

than experimental one.

It testifies that the calculational rate of condensation which is determined by

interface heat flux on liquid side is essentially underestimated (the interface heat

flux from vapor side in considered mode is rather small as the temperature of

vapor phase is close to saturation).

During the calculation execution the increasing of values interface heat

transfer coefficient Hif in a number of cases (especially for large subcooling) was

not caused any changes of the calculational results. The conducted analysis of

the algorithm has shown that it contains some limitations on interface heat flux

and wall vapor generation values, which introduced into the model to make

stable the numerical scheme.
14

One of such limitations is the "umbrella" one, which decreased the heat

transfer coefficient values when void fraction is near the zero or one.

Hif = min [ Hif, 17539 * max ( 4.724, 472.4 ag (1- ag))] *

* max [0, min (1 , (ag - 1.0 * 10 -10)/ (0.1 - 1.0*1010))] (27)

This limitation realizes only by using the semi-explicit numerical scheme. The

nearly-implicit scheme does not consist such limitation, therefore the results of

calculations with using this two schemes differed from each other as it showed

on fig. 31. Besides that the interface heat flux is limited by the condition:

A1=[r w-Hif *(Ts-Tf)/hfg]*At, (28)

A2 = 0.5 ag P g (1- x), (29)

-Al> A2, (30)

that is the amount of appearing (disappearing) vapor in the volume at one time

step must be less than a half of amount of vapor in this volume. The presence of

such hard limitation leads to considerable lowering of interface heat transfer

coefficient (for high subcoolings and large values of Hif ) and dependence of the

results from time step. The degree of Hif decreasing depends on ratio between

the amount of disappearing vapor and its amount in volume and it is equal to

Hif new = Hif old A2/A1 (31)

It must be noted that value of Hif old is not calculated from (16-20) but is also

the corrected value and it is computed from the time relaxation procedure

Hif old = Hif m+1 (Hif m/Hif m+l)Y (32)

y = exp (- 10 *At)*( 1+ 0.25 (T, - Tf )), (33)

where Hif m+1 is calculated from (16) (20) and Hif m is the Hif from previous time

step. Therefore large values of interface heat transfer coefficients are reduced in
15

some orders and the results of calculations become independent of type of

correlations used.

For regimes with low subcooling or for condensation case in the unheated

part of the channel the limitations described do not play the essential role and

the interface heat transfer coefficient is calculated with using Unal's formula

which in our opinion doesn't, describe truthfully physics of condensation process

in the unheated channels, because this correlation was originally obtained for the

conditions rather differed from the under consideration ones.

Field of applicability of the Unal's formula apparently must be restricted by

the conditions, when the vapor bubbles are attached to the channel wall, i.e.

from the location of their appearance until the departure location. The

appearance location may be determined as following

Tw=T. or Tf= T.- q/ho,-

and the location of vapor bubbles departure as it was shown in [17] is the same

point as the point of intensive growth of vapor void fraction. This point can be

determined by the Zuber- Saha formula [13]. This region is large enough and as

it was estimated in [ 17] as

I Xal = 31 XbI I

where xa - relative enthalpy of fluid (quality) at the point of bubbles appearance

and xb - relative fluid enthalpy at the point of intensive growth of vapor fraction

and calculated by Zuber-Saha formula. More reasonable in our opinion is using

of correlation from [8] obtained for subcooled boiling and condensation processes

and having an experimental confirmation [12].

St = 0.228 * Re 0.3, * Pr-o, *( P / P/(1 - ag)) 0.25 , (33)

where

St=Nu/Pe, Nu=qwf*De/kf , Pe= G De Cpf/kf

This formula was used in our calculations only for nonheated part of the pipe

when the subcooling is riot too high, that is for those conditions when the

restrictions (31)-(32) do not deform the calculational results. The latest ones were

considerably better for bubble mode. Data from [6] are presented on fig.33 for

the following parameters : pressure P=7.0 MPa , mass flux G=2960 kg/s*m2
16

heat flux q=1200kWt/m2 , inlet flow temperature T=526.5 K . The geometry of

the channel : internal diameter of a channel d=12.1mm, heated length 1=1.235m,

unheated length 1=1.235m. There are three curves on the figure: with using

semi-explicit scheme, nearly implicit, and with using (16) and (33) for nonheated

part of the channel (nearly-implicit scheme).

It is obviously, that the presence of "umbrella" restriction reduces the rate of

condensation , other limitations on Hif value do not deform the results of

calculations for unheated zone. We note also that in this made of flow on all

length of the channel was identified on accepted the code model as a bubble one,

top border of existence of which i.e. the transition from bubble to slug regime is

calculated in the code according to:

Cbs = Cbs" , if G < 2000 kg/m 2/s,

abs = abs" [(0.5 - cas" )/1000] (G-2000), if 2000 < G < 3000 kg/m 2/s,

abs = 0.5, if G > 3000 kg/m 2/s,

where abs* = max[ 0.25 min (1.0, (0.045 D')8), 0.001],

D" = D [ g (pf - pg )/a]0.

5 , D - hydraulic diameter, a - surface tension

coefficient. The most reasonable results , well agreed with experimental data

were obtained by use of the formula (33).

For low mass fluxes (G < 2000) and regime parameters and geometry under

consideration the (34) gives the acbs=0.001 that contradicts in our opinion to

experimental data. So for example according to a map used in the code

RETRAN [9] ,for a given mass fluxes this value is abs=0.2-0.4 in code TRAC

Cabs=0. 3 -0.5 that will be agreed with experimental data [11], [15-16] and

theoretical prediction [14].

Share of bubbles in slug regime is defined as

•bub = exp [-8* ( ag - CUbs ) / (aC - Xbs )] * abs (35)

As follows from Cbub in slug mode aims to zero very quickly and at abs- 0 .

0 0 1

this value are actually away, at that the interface heat flux from bubbles has

the main contribution on total heat flux from interface to liquid (due to

considerably greater area of interface surface).
17

The validation of influence of change of abs on the condensation rate was

conducted. The void fraction profiles along the channel are presented on fig.34

for the parameters P=7.0 MPa, G = 730 kg/s m2 , q= 618 kWt/m 2, Tinlet= 492 K

and geometry as for fig. 33. The calculations were carried out with using the

"nearly-implicit" numerical scheme and with abs-0.2 , respectively increased

values of Hif for non-heated part.

Curve 1 on this figure shows the results of original model , curve 2 - the

results with abs-0. 2 . More better accordance for latest case testifies about the

presence of bubble flow regime in the nonheated part of the channel and that

the corrected value of ab,3-0. 2 is more reasonable. As was marked earlier the

correlation (33) in the regime with presence of bubbles may give some better

results. Curve 3 shows the behavior of void fraction by use (33) instead of (16) in

the unheated part of the channel and at abs-0.2. In both latest cases the

interface heat transfer coefficient are higher and the calculated data are more

close to experimental ones.

The results of experiments of [7] has been obtained with test section of 1m

long unheated and 0.4m. heated parts and inner diameter 12.03mm. The

experimental and calculational results of test with following parameters: P=6.95

MPa, G=980 kg/s mi2, q=824 kWt/m2, Tinlet=504 K are presented on fig.35. The

marks on this figure are the same as on previous one. One can see that the

change of the transition from bubbles to slug regimes causes the more good

accordance between the -experimental and calculational data and the rate of

condensation is increased by using (33).

It seems logical to make the system of closer equations for vapor generation

and condensation terms for both parts of the channel including the heated zone.

But such attempts were unsuccessful , because it is impossible to receive the

reasonable values of the parameters because of excess of 507% limit on

condensation and vaporization rates. The attempts of soften this conditions

resulted in catastrophic growth of parameters oscillations.

I
18

4. SENSITIVITY ANALYSIS

The conducted calculations showed that the heaviest inlet parameters

influence on the results were the regimes with vapor condensation in the

unheated zone of channel. Therefore the sensitivity analysis results are

presented for such conditions in main.

The data of void distribution for the experiments from [7] are presented at

fig. 47, where the results for different number of subvolumes are presented.

Obviously that for number of subvolumes greater then N=14 , (that corresponds

to the subvolumes length A z = 0.1 m) the calculational results do not practically

change. Therefore all the calculations had been conducted with approximately

such sizes of subvolumes.

The results of calculations, showed the influence of inlet temperature error

are presented at fig. 48. As a rule this value was approximately 1 K. Obviously

that the calculational results depend strongly on this value.

The influence of heat flux error on void distribution along the channel is

presented in fig. 49. For the experiments from [7] the maximum value of relative

error of heat flux measurements makes 5 q = 3%. Obviously that this error gives

the maximal contribution on the void fraction distribution along the channel.

Other errors - mass flow and pressure ones have much smaller influence in

the ag behavior at given parameters. However, as was indicated earlier, the

influence of pressure error increases for low pressures. Moreover as was found

out, the influence of time step on the results increases at low pressures also.

We note that practically all made calculations which were conducted by use

the RELAP 5 /mod 3 version 5m5 were repeated by use of RELAP5/mod 3

version 7j and RELAP5 / mod3.1 codes. The calculational results of this two

codes, as has appeared, coincide by the RELAP5 /mod 3 version 5m5 code

results, but it was found out that at low pressures 9 for [18] tests at the same

time step the latest two codes may give nonphysical void distribution (fig. 50)

along the channel which one can remove by reducing of time step value.

i
19

5. RUN STATISTICS

All the calculations were carried out with computer IBM PC-386 and only

small part of them with IBM RISK 6000 computer to evaluate the possible

difference between the results. The most part of calculations have been

performed with time step At =0.05s, which guarantees the absence of parameters

oscillation and dependense of calculational results from time step. Further

decreasing of time step have no influence on the results. Therefore the plot of

time steps as a function of real transient time is simple constant function At

=0.05s and it is not provided in this report.

Typical grind time for IBM PC-386 computer was

(CPU time)*10' /((Number of volumes) (Number of time steps)) =

= 35*103/(14*150) = 8.3

For IBM RISK 6000 grind time was 1.08.

6. CONCLUSIONS

Trustworthiness of RELAP5/mod3 version 5m5 models for subcooled boiling

process was verified using a lot of experimental data. It was found out that the

code models give good enough results for uniformly heated channels except of

very low pressure case. For more complicated lows of power distribution along

the channel the rates of vapor generation and condensation are not compatible

and the discrepancies between calculational and experimental data become too

large. The main causes of this are the limitations of heat transfer rate terms

implemented into the code due to imperfections of code numerical scheme.
20

REFERENCES

1. Carlson K.E. et al. RELAP5/Mod3 Code Manual, vol. 1, - NUREG/cR-5535,

1990., Idaho.

2. Rouhani S.Z., Axelsson E. Calculation of Void Volume Fraction in Subcooled

and Quality Boiling Regions, - International Journal of Heat and Mass Transfer,

V.13, pp. 383-393.

3. Marchaterre J.F. Natural and Forced Circulation Boiling Studies, - ANL-5735,

May, 1960.

4. Bartolomey G.G., Sabotinov L. S. Void fraction at subcooled boiling at

different heat distribution lows along the channel. Nuclear Energy, 1976, N3,

Sofia, (in Bulgarian).

5. Bartolomey G.G. et al. Experimental investigation of void fraction at

subcooled boiling mode in tubes. Teploenergetika, 1982, N3, pp. 20-22. (in

Russian).

6. Labuntsov D.A. et al. Void fraction investigation of nonequilibrium two-phase

flows. ENIN transactions, 'Heat transfer and hydrodynamics, v. 35. pp. 88-98. (in

Russian).

7. Bartolomey G.G. Void fraction of diabatic flows in tubes at different heat

distribution lows. Heat and mass transfer - VI, International Conference, Minsk,

1980, v.5, pp. 38-43.

8. Avdeev A.A., Pehterev V.P. Subcooled boiling of liquid at forced convection

conditions, High Temperature Thermal Engineering, 1986, v.24, N5, pp.

912-920.(in Russian).

9. RETRAN -02 - A Program for Transient Thermal-Hydraulic Analysis of

Complex Fluid Flow Systems, EPRI NP-1850-CCMA, V.1, 1984.

10. TRAC-PF1 - An Advanced Best Estimate Computer Program for Pressurized

Water Reactor Analysis. NUREG/CR - 3567, LA-9944-MS, Los Alamos, 1984.

11. Bennet A.W. Hewitt G.F. , - Flow Visualization Studies of Boiling Water at

High Pressure, AERE-R 4874, 1965.

12. Avdeev A.A Reynolds analogy application for subcooled boiling process at

forced convection conditions. High Temperature Thermal Engineering,

1986,v.24,N1, pp. 111-119. (in Russian).

13. Saha P., Zuber N., Point of Net Vapor Generation and Vapor Void Fraction

in Subcooled Boiling, - Proc. 5-th Intern. Heat Transfer Conf., Tokyo, 1974, v.4,

pp. 175-179.

I
21

14. Mishima K., Ishii '..\., Flow Regime Transition Criteria for Upward

Two-phase Flow in Vertical Tubes. International Journal of Heat and Mass

Transfer, V.27, N5, pp.723-737, 1984.

15. Hosler E.R., Flow Patterns in High Pressure Two-phase Flow with Heat

Addition., Chem. Eng. Symp. Ser., V.64, N82, p.54, 1968.

16. Doroschuc V.E., Borevsky L.Ja., Levitan L.L., Holographic Investigation of

Steam-water Flows in Heated and Unheated Channels. Heat Transfer. 1982.

Proc. VII Intern. Heat Transfer Conference, Munich, 1982, V.5, pp. 277-281.

17. Bartolomey G.G. Mihailov V.N. Intensive vapor production beginning

enthalpy. -Teploenergetica, 1987, N2, pp. 17-20. (in Russian).

18. Evangelisti R., Lupoli P. The void fraction in an annular channel at

atmospheric pressure. - International Journal of Heat and Mass Transfer, 1969,

v.12, 3NT_6, pp. 699-711.

19. Martin R. Measurement of the Local Void Fraction at High Pressure in a

Heated Channel, - Nuclear Science and Engineering, 1972, v.48, X9fo2, pp.125-138.

20. Unal H.S. Maximum Bubble Diameter, Maximum Bubble-Growth Time and

Bubble-Growth Rate During the Subcooled Boiling of Water up to 17 Mn/m 2 ,

- International Journal of Heat and Mass Transfer, v.19,1976, pp. 643-649.

)
Appendix 1

Input Deck for BARTOLOMEY Experiment
APPENDIX 1

INPUT DECK FOR BARTOLOMEY EXPERIMENT

= pipe BARIPLOMEYTESTWTrH CONDENSATION

0000100 new transnt

end dtmin dtmax

0000201 30.0 1.0-7 0.025 00011 5 100000 100000

0000301 p 004010000

0000302 p 004140000

000O303 quals 004010000

000O304 quals 004140000

0000305 veltj 004010000

0000307 velg 004140000

0000308 velf 004140000

0000309 voidg 004140000

0000310 quale 004140000

0000312 ug 004140000

0000313 uf004140000

0000314 tempf 004140000

0000315 tempg 004140000

a component DD2

a

0020000 inlet tmdpvol a

a vol area vollength vol vol ahor aver dcev rough dhy

0020101 1.13606-4 0.075 0 0 90.0 0.075 1.0-4 0 0

a ebt

0020200 O3

S time pres temp

0020201 0.0 6.9600+6 504.0

component 003

0030000 inlet tmdpjun

a from to area

0030101 002000000 004000000 1.13606-4

0030200 1

a time flowf flowg

0030201 0.0 0.1113 0.0 0.0

.. ¶

I.

S

1-i NUREG/IA-0025

'7
* component 004

• work pipe

0040000 tube pipe

0040001 14 *nvol

• vol area vol no

0040101 1.13606-4 14

vol length vol no

0040301 0.1000 14 *

e aver vol no

0040601 90.0 • 14

* rough dhy vol no

0040801 1.0-4 0.0 14

• floss dloss jun no

0040901 0.0 0.0 13 S

• pvbfe vol no

0041001 00000 14

* fvcahs jun no

0041101 001000 13

* ebt press temp

0041201 103 6.9600+6 504.0 0 0 0 14

• flowf flowg win junno

0041301 0.1113 0.00 0.0 13

- component 005

0050000 outlet sngIjun

- from to area floss rloss fvctas

0050101 004010000 006000000 1.13606-4 0 0 001000

• flowf flowg win

0050201 1 0.1113 0.00 0.0

* component 006 S

0060000 outlet tmdpvol

vol area vol length vol vol ahor aver elev rough dhy

0060101 1.13606-4 0.2 0 0 90.0 0.2 1.0-4 0 0

* eN1

NUREG/IA-0025. 1-2

. .. . ... .
0060200 102 *

S time pres x

0060201 0.0 6.9600+6 0.000

*

* component 008

* work pipe

nh np geom st-st left

10080000 14 11 2 1 6.015-3 .......... .... -':.

S

10080100 0 1 S

10080101 10 7.015-3

10080201 1 10 S

10080301 1.0 10 S

10080400 0

10080401 500.0 11 *

10080501 004010000 010000 I 1 0.1 14

10080601 0 0 2701 1 0.1 14 S

10080701 025 0.1 0 0 10

10080702 025 0.0 0 0 14 S

left chf Whf lhb gslf gslr glcf gcr bf no

10080801 1.2-2 20. 20. 0. 0. 0. 0. 1. 14

10080901 1.60-2 20. 20. 0. 0. 0. 0. 1. 14

20270100 htnmrate *

20270101 0.0 0.0 without heat losses

20270102 l.Oc6 0.0

20100100 s-steel

20202500 power

20202501 0.0 3.112594+4

20202502 2.0 3.112594+4

20202503 11.0 3.112594+4

20202504 1.0+6 3.112594+4

end of input deck

1-3 NUREG/IA-0025

i
Appendix 2

Figures
Voidg

0.50

0.40

0.30

0.20

0.10

0.00

BARTOLOMEY DATA

D = 0.012

L = 1.50

m

m

MPa

K /

p

G

Q

T

14.7

2014.

1720.

545.0

Quality Fig. 1

2-1 2-1 NUREGJIA-0025
Voidg

0.30 -

0.20 -

0.10 -

0.00-

-0.

BARTOLOMEY DATA

D = 0.012

L = 1.500

P = 14.75

G = 2123.

Q = 1130.

T = 583.0

m

m

MPa

kg/s*noj

k

K

-0.05

Quality Fig.2

NUREG/IA-002522 2-2
Voidg

0.40

0.30

0.20

0.10

0.00

BARTOLOMEY DATA

D = 0.012 m

L = 1.500 m

p

G

Q

T

14.74

1847.

770.0

598.0

MPa

K

Quality Fig.3

2-3 2-3 NUREG/IA-0025
Voidg

0.20

0.15

0.10

0.05

0.00

BARTOLOMEY DATA

D = 0.012 m

L = 1.500 m

P = 14.79

G = 1878.

Q = 420.0

T = 603.0

MPa2

K

Quality Fig.4

NUREG/IA-002524 2-4
Voidg BARTOLOMEY DATA

0.20

0.15

0.10

0.05

0.00

D = 0.012 m

L = 0.750 m

P = 14.99

G = 2012.

Q = 2210.

T = 563.0

MPa

K 1

Quality Fig.5

2-5 NUREG/IA-0025
Voidg

0.50

0.40

0.30

0.20

0.10

BARTOLOMEY DATA

* experi ment *

- Relap

X*

* . , ... , r ,- ,n-1rtfl

D = 0.012 m

L = 1.500 m

P = 14.68 MPa

G = 1000. kg4 S3n2

Q = 1130.

T = 533.0 K

0.00

-0.20 -0.15 --0.10 -0.05 0.00 0.05 0.10

Quality Fig.6

NUREG/IA-0025 2-6
Voidg

0.60 -.

0.50

0.40

0.30

0.20

0.10

0.00

-0.

BARTOLOMEY DATA

D = 0.012 m

L = 1.000 m

P = 11.02

G = 503.0

Q = 990.0

T = 494.0

MPa

K

Fig.7 Quality

2-7 2-7 NUREG/LA-0025
Voidg

0.50

0.40

0.30

0.20

0.10

BARTOLOMEY DATA

* experiment

- Relap

* *t -

D = 0.012 m

L = 1.500 m

P = 10.81

G = 966.0 Q = 1130.

T = 502.0

MPa

km*In~s W/m

K

0.00

).20 -0.10 0.00 0.10

Quality Fig.8

NUREG/IA-0025 2-8
Voidg

0.50 ,

0.40

0.30

0.20

0.10

0.00 -

-0.20

BARTOLOMEY DATA

D = 0.012 m

L = 1.160 m

P = 10.81

G = 1554.

Q = 1160.

T = 563.

MPa

kf 4rn*rn*s

K /*

Quality Fig.9

2-9 NUREG/IA-0025
Voidg

0.55

0.35

0.15

-0.05

BARTOLOMEY DATA

D = 0.012 m

L = 1.500 m

P = 10.84

G = 1959.

Q = 1130.

T = 563.0

MPa

KmK

Quality Fig. 10

NUREG/IA-0025 2-10
Voidg

0.40

0.30

0.20

0.10

-0.00

-0.10

BARTOLOMEY DATA

D = 0.012 m

L = 1.500 m

P = 6.810

G= 998.0

Q =440.0

T = 521.0

MPa2

K /

Quality Fig.1 1

2-11 2-11 NUREG/IA-0025
Voidg

0.50

0.30

0.10

BARTOLOMEY DATA

* experiment

- Relap

F*

*/

|i i | li l~ l l | l l | ;; ~ i J | l ~ ~ l~ J il*l t

D = 0.012 m

L =1.500 m

P

G

Q

T

6.890

965.0

780.0

493.0

MPa

k%.M*rn*s

-0.10 --I

0.20 -0.15 -0.10 -0.05 0.00 0.05 0.10

Quality Fig. 12

NUREG/IA-0025 21-12
Voidg

0.70

0.50

0.30

0.10

-0.10

BARTOLOMEY DATA

D = 0.012 m

L = 1.500 m

P = 6.840

G = 961.0

Q = 1130.

T = 466.0

MPa

K /

Quality Fig. 13

2-13 NUREG/IA-0025
Voidg

0.60

0.40

0.20

0.00

BARTOLOMEY DATA

D = 0.012 m

L 1.400 m

P = 6.740 MPa

G= 988.0 Kg/s*m2

Q = 1700. kW/m 2

T = 416.0 K

0

Quality Fig. 14

NUREG/IA-0025 2-14
Voidg

0.50

0.40

0.30

0.20

0.10

0.00

BARTOLOMEY DATA

D = 0.012 m

L = 1.000 m

p

G

Q

T

= .7.010 = 996.0 = 1980.

= 434.0

MPa 2

kg/s*i2 kW/m

K

Quality Fig. 15

2-15 2 NUREG/IA-0025
BARTOLOMEY DATA

0.60

Voidg

0.50

0.40

0.30

0.20

0.10

Fig.16 Quality

NUREG/IA-0025 2-16
Voidg

0.30

0.20

0.10

-0.00

-0.10

BARTOLOMEY DATA

D = 0.012 m

L = 1.500 m

P = 6.810 MPa

G = 2037. kg/s*rr,• Q 1130. kW/m'

T = 504.0 K

Quality Fig. 17

2-17 NUREG/IA-0025
Voidg

0.80

BARTOLOMEY DATA

D = 0.012 m

L = 1.240 m

0.60

P = 6.890

G = 405.0

Q = 790.0

T.= 421.0

MPa

kW/m

K 0.40

0.20

0.00

Quality Fig. 18

NUREGIIA-0025 212-18
Voidg

0.25

0.20

0.15

0.10

0.05

0.00

BARTOLOMEY DATA

D = 0.012 m

L = 1.160 m

p

G

Q

T

6.890

1467.

770.0

519.0

MPa2

kgv4 s*nm2

K

Quality Fig. 19

2-19 NUREG/IA-0025
Voidg

0.25

0.20

0.15

0.10

0.05

-0.00

-0.05

BARTOLOMEY DATA

D = 0.012 m

L = 1.500 m

P = 6.790

G = 20240

Q = 780.0

T = 520.0

MPa

K

Fig.20 Quality

NUREG/IA-0025 2-20
BARTOLOMEY DATA

Fig.21 Quality

2-21 NUREG/IA-0025
Voidg

0.75 -

0.55

0.35

0.15 -

-0.05-0.:

BARTOLOMEY DATA

D = 0.012 m

L = 1.500 m

P = 4.410

G = 994.0

Q = 900.0

T = 463.0

MPa

K

Quality Fig.22

NUREG/IA-0025 2-22
Voidg

0.70

0.50

0.30

0.10

-0.10 i,

-0.20

BARTOLOMEY DATA

D = 0.012 m

L = 1.000 m

P = 3.010

G = 990.0

Q = 980.0

T = 445.0

MPa 2

k~r/s*nr6

K

Quality Fig.23

2-23 2-23 NUREG/IA-0025
BARTOLOMEY DATA

0.60

Voidg

0.50

0.40

0.30

0.20

0.10

0.00

Fig.24 Quality

NUREG/IA-0025 2-24 22
Voidg

0.80

0.60

0.40

0.20

0.00

SABOTINOV DATA

D = 0.0117

L = 1.500

t

P = 6.806

G = 419.0

Q = 433.0

T = 539.0

m

m

MPa

kg/m*m*s

kW/m*m

K

0.(

L (m)

Fig.25

2-25 NUREG/IA-0025
Voidg

0.80

0.60

0.40

0.20

0.00

SABOTINOV DATA

D = 0.0117

L = 1.500

P = 6.806

G = 419.0

Q = 433.0

T = 539.0

m

m

MPa

kg/m*m*s

kW/m*m

K

Quality Fig.26

NUREGJIA-0025 2-26
Voidg

1.00

0.80

0.60

0.40

0.20

SABOTINOV DATA

* expnriment

- Relap ___ _ "

*

*/

l *ll l I I I I III I I II I II

D = 0.0117

L = 1.500

P = 6.786

G = 962.0

Q = 1688.

T = 461.0

m

m

MPa

kg/m*m*s

kW/m*m

K

0.00

-0.3 -0.2 -0.1 0.0 0.1 0.2 0.3

Quality Fig.27

2-27 NUREGI/A-0025
Voidg

0.75

0.55

0.35

0.15

-0.05 -,,

-0.10

SABOTINOV DATA

D = 0.0116m

L = 1.500 m

P=4.4

G = 1000.

Q = 436.0

MPa

Kq/mrm*s

Quality Fig.28

NUREG/IA-0025 2-28
Voidg

0.70

0.50

0.30

0.10

-0.10

SABOTINOV DATA

D = 0.0117m

L = 1.500 m

p

G

Q

T

6.86

1000.0

430.0

461.0

MPa

kg/m*m*s

kW/m*m

K

Quality .Fig.29

2-29 NUREG/IA-0025
Voidg

0.50

0.40

0.30

0.20

0.10

SABOTINOV DATA

eperi R ment Sel OF

-rT--rT " . .... .. ... .... ... ...

D = 0.0117m

L = 1.500 m

P

G

Q

T

6.86

1000.0

796.5

461.0

MPa

kg/m*m*s

kW/m*m

K

0.00.-- 0.20 -0.15 -0.10 -0.05 0.00 0.05

Fig.30 Quality

NUREGIIA-0025 232-30
Voidg

0.15

0.10

0.05

0.00

BARTOLOMEY DATA

00000 ei periment in Lplicit

ey- plicit

0

00

0N

I0

'a o •,•,

'a. Il %

00 ,/

P

G

T

Q

6.89

1000.0

1200.0

452.0

MPa

kg/m*m*s

kW/m*m

K

min

rn

D = 0.01203

L = 1.0

L= 0.4

0.0 0.4 0.8

Fig.31.

1.2 1.6

L (m)

2-31 NUREG/IA-0025
Voidg

0.60

0.50

0.40

0.30

0.20

0.10

0.00

BARTOLOMEY TEST

D

L

L

P

G

Q

T

0.01203

1.0

0.4

6.89

1000.0

800.0

534.0

m

m

m

MPa

kg/m*m*s

kW/m*m

K

(mn)

NUREG/IA-0025 2-32
LABUNTSOV DATA

0.20

Voidg

0.15

0.10

0.05

0.00

XPCI

G G.-

pappil~ Ip~lpll PpTpJ(T

Relap

GOOOD experiment

D

L

L

P

G

Q

T

0.0121

1.235

1.235

7.00 2960.

1200.

526.5

m

m

m

MPa

kW/m

K

0.5 1.0 1.5 2.0

Fig.33

2.5

Z,m

2-33 NUREG/IA-0025
LABUNTSOV DATA

0.40

Voidg

0.30-

P=7.0 IPa

G =730 kg/ 2 Q=618 cW/m/

T=492 c *' 

0.20

D = 0.( 121 m

L1 =0.35 *

L2 = 1 235 m /"

0.10 !

0.00 1 ii-i-i

Fig.34 Z,m

NUREG/IA-0025 232-34
BARTOLOMEY TEST

0.40

I

Fig.35 • Z,m

2-35 NUREG/IA-0025
Voidg

0.80

0.60

0.40

0.20

ROUHANI DATA

* experi nent " Relap

lI ii | i |I I I iIII I I I I I I I I I I I | I |

D = 0.010

L =0.400

m

m

bar

kg/m*m*s

kW/m*m

K

p

G

Q

T

50.0

129.44

918.77

366.5

0.00 i

0.1 -0.0 0.1 0.2

Quality

Fig.36

NUREG/IA-0025 232-36
Voidg

0.80

0.60

0.40

0.20

ROUHANI DATA

/**

* experiment

- Re op

0.10 -0.00 0.10 0.20 0.3(

D = 0.010

L = 0.440

P = 39.228

G = 127.22

Q = 607.086

T = 441.75

m

m

bar

kg/m*m*s

kW/m*m

0.00 )

0 Quality

Fig.37

2-37 2-37 NUREG/IA-0025
Voidg

0.80

0.60

0.40

0.20

ROUHANI DATA

* experiment * *

Relp *

/ *

***

m|rrrrr rr-miIII I II I A I III Ir ITr

0.10 -005 0.00 0.05 0.10 0.15 0.2w

D = 0.010

L =.0.440

P = 29.225

G = 127.22

Q = 721.06

T = 359.00

m

m

bar

kg/m*m*s

kW/m*m

K

0.00

0

Quality Fig.38

NUREG/IA-0025 2-38
.Voidg

0.80

0.60

0.40

0.20

ROUHANI DATA

• experiment

- Rela/

- nrTr 4 rmm. mmmm

D

L

P

G

Q

T

= 0.010

= 0.440

= 9.611 - 125.838 - 588.48

= 336.25

m

m

bar

kg/m*m*s

kW/m*m

K

0 00 A

0.10 -0.05 0.00 0.05 0.10

Quality Fig.39

2-39 NUREG/IA-0025
Voidg

0.80

0.60

0.40

0.20

0.00

MARCHATERRE DATA

D = 0.010

L = 1.300

m

m

bar

kg/m*m*s /m*m

K

P

G

Q

T

11.180

500.0

132.582

448.12

)6

Quality Fig.40

NUREGIIA-0025 2-40 24
Voidg

0.80

0.60

0.40

0.20

MARCHATERRE DATA

* experiment , -Relc p

/*

II-

*t

D = 0.010

L = 1.300

m

m

bar

kg/m*m*s

kV/m*m

K

P

G

Q

T

= 11.180

= 850.0

= 160.494

=. 451.57

0.00

0.02 -- 0.01 0.00 0.01 0.02 0.03 0.04

Quality Fig.41

2--41 NUREG/IA-0025
Voidg

0.60

0.50

0.40

0.30

.0.20

0.10

MARCHATERRE DATA

*

* * experimer

-Relap

SA4*

I I I I I | I I llI I i i I

D = 0.010

L = 1.300

P = 11.180

G = 1516.67

Q = 160.494

T = 454.32

m

m

bar

kg/m*m*s

kW/m*m

K

0.00

-0.015 -0.005 0.005 0.015 0.025

Quality

Fig.42

NUREG/IA-0025 2-42
EVANGELISTI DATA

0.80

-o

0.60 -

0.40

0.20

0.00 . -30.0

*

***** experiment * Relop *

*

**

**

* **

**

.4...................................

P = 1.128 bar, G =-1416 kg/mrm/

Q = 9.726 kg/m/r

,,,,, .... I,. . ,.. . .. .... I....... . . I

-25.0 -20.0 -15.0 -10.0 -5.0- 0.0 5.0

Outlet subcooling, K

Fig.43

2-43 2-43 NUREGJIA-0025
0.80

.. ... Evangelisti data with

different pressure

boundary conditions

0.60

0)

2 0.40

0

0.20

0.00

Fig.44

NUREG/IA-0025 242-44
0.80

0.60

0.40

0

0.20

0.00

-0.040

P outlet = P exp * P inlet = P exp

-6.630 -0.0 ' -' -'b' 0.0. ...-0.010" 0.000 I

0.010

Quality Fig.45

2-45 2-45 NUREG/IA-0025
0.80

0.60

-Q 0.40

0

0.20

0.00 ,, -0.2

Fig.46

NUREG/IA-0025 2-46
BARTOLOMEY DATA

0.40

Voidg 14 vol- RELAPVoidg 28 vol- RELAP

*** *experiment P=6.89 MPa - - 7 vol -RELAP G=1000.0 k g/m*m*s

0.30 Q=800 kW/m*m

,** T=503.0 K

0.20

0.10 **_

D=O.O1 203m

Ll1=1.0m

,** ____L2=O.4m

0.00 A. AAA a A A AAAA I 1IA

Fig.47 L,m

2-47 2-47 NUREG/IA-0025
0.80

0.60 versions

0.40 * 'RELAP5/MOD3 ver.5m5

0.40 RELAP5/MOD3 ver. 7j. > • RELAP5/MOD3.1

.0.20

0.00 6

-0.050 -0.040 -0.030 -0.020 -0.010 0.000 oolOquality

I-1 g.,-3U

NUREG/IA-0025 2-48
BARTOLOMEY DATA

0.40

Voidg

0.30

0.20

0.10

0.00

L,m

2-49 2-49 NUREG/IA-0025
0.40

Voidg

0.30

0.20

0.10

0.00

BARTOLOMEY DATA

**experiment

OMT--lKST--1K

4 1.• ,, , a . . . .., , ,, , ,

A

0.0 0.4 0.8 1.2 1.6 Fig.48 L,m

NUREG/IA-0025 2-50
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

(2-89) (Assigned by NRC, Add Vol., Supp., Rev.,

NRCM 1102, and Addendum Numbers, if any.) 3201. 3202 BIBLIOGRAPHIC DATA SHEET

(See instructions on the reverse) NUREG/IA-0025

2. TITLE AND SUBTITLE

RELAP5/MOD3 Subcooled Boiling Model Assessment

3. DATE REPORT PUBLISHED

MONTH YEAR

May 1998

4. FIN OR GRANT NUMBER

5. AUTHOR(S) 6. TYPE OF REPORT

A. S. Devkin and A. S.,Podosenov Technical

7. PERIOD COVERED (Inclusive Dales)

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

provide name and mailing address.)

Nuclear Safety Institute, Russian Research Center, "Kurchatov Institute"

Kurchatov Square 1,

123182, Moscow,

Russia

9. SPONSORING ORGANIZATION - NAME AND ADDRESS (If NRC, type 'Same as above'; if contractor, provide NRC Division, Office or Region, U.S. Nuclear Regulatory Commission,

and mailing address.)

Division of Systems Technology

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, D.C. 20005-0001

10. SUPPLEMENTARY NOUEb5

S. Smith, NRC Project Manager

11. ABSTRACT (200 words or less)

This report presents the assessment of the RELAP5/Mod3 (5m5 version) code subcooled boiling process model, which is

based on a variety of experiments. The accuracy of the model is confirmed for a wide range of regime parameters for the case

of uniform heating along the channel. The condensation rate is rather underpredicted, which may lead to considerable errors

in void fraction behavior prediction in subcooled boiling regimes for nonuniformly or unheated channels.

12. KEY WORDS/DESCRIPTORS (List words or phrases that will assist researchers in locating the report.) 13. AVAILABILITY STATEMENT

RELAP5/Mod3 (5m5 version) code unlimited

subcooled boiling regimes 14. SECURITY CLASSIFICATION

void fraction behavior prediction (This Page)

unclassified

(This Report)

unclassified

15. NUMBER OF PAGES

16. PRICE

NRC FORM 335 (2-89)
Federal Recycling Program