NUREG/IA-0163

International

Agreement Report

A Study of the

Dispersed Flow Interfacial

Heat Transfer Model of

RELAP5/MOD2.5 and

RELAP5/MOD3

Prepared by

M. Andreani*

G. Th. Analytis

S. N. Aksan

Thermal-Hydraulics Laboratory *ETHZ, Nuclear Engineering Laboratory

Paul Scherrer Institute (PSI) Swiss Federal Institute of Technology (ETH)

CH-5232 Villigen PSI, Switzerland 8092 Zurich, Switzerland

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

April 1999

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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owned rights.
NUREG/IA-0163

International

Agreement Report

A Study of the

Dispersed Flow Interfacial

Heat Transfer Model of

RELAP5/MOD2.5 and

RELAP5/MOD3

Prepared by

M. Andreani*

G. Th. Analytis

S. N. Aksan

Thermal-Hydraulics Laboratory *ETHZ, Nuclear Engineering Laboratory Paul Scherrer Institute (PSI) Swiss Federal Institute of Technology (ETH)

CH-5232 Vflligen PSI, Switzerland 8092 Zurich, Switzerland

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

April 1999

Prepared as part of

The Agreement on Research Participation and Technical Exchange under the International Code Application and Maintenance ftogram (CAMP)

Published by

U.S. Nuclear Regulatory Commission

A Study of the Dispersed Flow Interfacial Heat Transfer Model of

RELAPS/MOD2.5 and RELAP5/MOD3.

M. Andreani*, G. Th. Analytis, S. N. Aksan

Thermal-Hydralics Laboratory,

Paul Scherrer Institute (PSI),

CH-5232 Vifligen PSI, Switzerland

Abstract

In this work, the model of interfacial heat transfer for the dispersed flow regime used in the RELAP5

computer code is investigated. The validity of the 1-D approach used for calculating the heat exchange

between the droplets and the vapour in the dispersed flow region above a quench front was shown to be

questionable for conditions of low quality at the quench front and low mass flux. Under such conditions, the

interfacial heat transfer calculated assuming a uniform distribution of droplets over the cross-sectional area of

the channel is necessarily overpredicted, and the vapour superheat is strongly underpre~dicted. The purpose of

the present paper is to show that the limitations of the 1-D approach, obtained from the steady-stare anlyses

of slow reflooding experiments, has some impact on the performance of the l-D transient computer codes

like RELAP5/MOD2.5 and RELAPS/MOD3. As an example, the transient analysis of a low flooding rate

experiment in a tube was performed. An early completion of the quench process and a fast desuperheating of

the vapour at the tube exit was obtained by both codes. The too high quench front velocity (four times higher

than in the experiment) could not, however, be put univocally in relation to the underprediction of the vapour

temperature, and the consequent increase of the precursory cooling, as many coupled thermal and hydraulic

transient effects prevailed. Quasi steady-state anlyses of two runs, where the boundary conditions for the

the post-dryout region could be better controlled for a predetermined position of the quench front, were

thus performed. These analyses show that the vapour superheat at the tube exit is strongly underpredicted,

confirming the limitations of the I-D model for interfacial heat transfer in the dispersed flow region.

* ETHZ. Nuclear Engineering Laboratory Swiss Federal Insitute of Technology (ETH),

8092 Zurich, Switzerland
Contents

List of Figures 3

List of Tables 5

I Introduction 6

2 Test cases used for the assessment 7

3 Transient analysis of test case 3051 10

4 Quasi-steady state analysis of Run 3051 12

4.1 Calculation with RELAPS/MOD2.5 .................................... 15

4.2 Calculation with RELAP5/MOD3 ..................................... 17

5 Quasi-steady state analysis of Run 3053 by RELAP5/MOD2.5 18

6 Conclusions 19

Acknowledgements 20

References 21

FIGURES 23

APPENDICES 58

I Input decks for transient calculations 58

A. RELAPS/Mod2.5 input deck 59

B. RELAPS/MOD3 input deck 68

II Sample input deck for quasi-steady state analysis 77

2
List of Figures

1 System flow diagram of the UC-B Test facility (Seban et al., 1978) .............. 25

2 Experimental wall temperatures at various elevations for Run 3051 .............. .. 26

3 Nodalizations used for the calculations of Run 3051 with RELAP5/ MOD2.5 and RELAPS/MOD3 .................................................. 27

4 Comparison of the wall temperatures at 0.66, 1.22, 1.83 and 2.44 m calculated by RELAP5/MOD2.5 and RELAP5/MOD3 with the experimental ones for Test 3051 ........ 28

5 Comparison of the wall temperatures at 3.05 and 3.35 m calculated by RELAP5/MOD2.5

and RELAP5/MOD3 with the experimental ones for Test 3051 ................ 29

6 Calculated vapour temperature at 2.44 m and comparison of the calculated vapour temperature

at the tube exit with the measured value ................................. 30

7 Calculated vapour and liquid carry-out .................................. 31

8 Calculated mass flow rate at the tube exit. ............................... 31

9 Comparison of the calculated accumulated mass, collapsed liquid level, vapour and liquid

entrainment with the experimental data. ................................. 32

10 Comparison of the calculated fluid expelled out of the test section by integrating directly

the mass flow with that using control variables and the primitive variables (velocities, void

fractions and densities): a) RELAPSIMOD2.5 and b) RELAP5IMOD3 ............ .. 33

11 Prerequisites for the "quasi steady-state' analysis of the post-CHF zone above a slowly advancing quench front (under these conditions the interfacial heat transfer model can be assessed 34

12 Nodalization used for the steady-state analysis of Run 3051 by RELAP5/MOD2.5...... 35

13 Initial axial wall temperature distributions used for the steady-state analysis of test 3051 at

the time of the mid-height quench. ................................... 36

14 Average axial wall heat flux distributions during the time span 10-15 s for the three initial

axial wall temperature distributions shown in Fig. 13 ......................... 36

15 Calculated wall temperatures time-history at various elevations, when the experimental axial

wall temperature distribution is imposed. ................................ .37

16 Equilibrium quality at the QF calculated using the three different initial wall temperature

distributions (Fig. 13) ............................................ 37

17 Vapour and liquid mass flow raes at the tube exit. .......................... 38

18 Average values of the total mass flow rates and exit quality over the time span 10-15 sec

The quantities to be averaged are calculated starting from 10 sec .................. 39

19 Comparison of the heat transfer rate to vapour with the total power input to the mixture

measured in the experiment. ....................................... 40

20 Comparison of the calculated vapour temperature at the tube exit with the experimental value

at the time of the mid-height quench ................................... 41

21 Time-average value of the vapour temperature at the tube exit. ................ .. 41

22 Phase velocity difference at z=2.4 m and z=3.0 m ........................... 42

23 Nodalization used for the steady-state analysis of run 3051 by RELAPS/MOD3 ....... . 43

24 Wall temperatures at various elevations calculated by the frozen version of RELAP5/MOD3. 44

25 Equilibrium quality at the QF calculated by three versions of RELAP5/MOD3 ........ . 45

26 Vapour temperature at the tube exit calculated by three versions of RELAP5/MOD3 (see

legend in Fig. 25) ............................................... 46

27 Vapour carry-out calculated by three versions of RELAPSIMOD3 (see legend in Fig. 25). 47

28 Liquid cany-out calculated by three versions of RELAP5/MOD3 (see legend in Fig. 25). 47

3
29 Void fractions (shown on two different scales) at various elevations calculated by the frozen

version of RELAP5IMOD3 ......................................... 48

30 Void fractions (shown on two different scales) at various elevations calculated by RELAP5/MOD3+PSI updates (version bb) .................................. 49

31 Void fractions (shown on two different scales) at various elevations calculated by RELAP5/MOD3+PSI updates+modified CD and di, = 1.25 mm (version UPX)........ .. 50

32 Power input to the mixture in the region above zQF for the three versions of RELAP5/MOD3

(see legend in Fig. 25) ............................................ 51

33 Heat transfer rate to the vapour alone for the three versions of RELAP5/MOD3 (legend in

Fig. 25) ...................... .............................. 51

34 Nodalization used for the steady-state analysis of run 3053 by RELAP5/MOD2.5 ...... . 52

35 Initial axial wall temperature distributions used for the steady-state analysis of test 3053 at

the time of the mid-height quench. ................................... 53

36 Average axial wall heat flux distributions during the time span 10-14 sec. for the two initial

axial wall temperature distributions shown in Fig. 35 ......................... 53

37 Wall temperatures at various elevations for test 3053 calculated by RELAP5/MOD2.5 ... 54

38 Equilibrium quality at the QF for the two different initial axial wall temperature profiles... 54

39 Heat transfer rate from wall to vapour in the post-dryout region for the two initial wall

temperature distributions ........................................... 55

40 Liquid carry-out for the two initial axial wall temperature distributions ............ .. 56

41 Total mass flow at the tube exit for the two initial axial wall temperature distributions .... 56

42 Vapour carry-out for the two initial axial wall temperature distributions ............ .. 57

43 Vapour temperature at the tube exit calculated by RELAP5/MOD2.5 for the two initial wall

temperature distributions. ......................................... 57

4
List of Tables

1 Experimental conditions for test cases UCB-3051 and UCB-3053 .............. 9

2 Calculated variables used for the steady-state analyses ......................... 9

5
1 Introduction

An extensive work of assessment, development and applications of the RELAP5 and TRAC/B reactor safety

computer programs has been underway for some time at the Thermal-Hydraulics Laboratory of the Paul

Scherrer Institute (PSI). One of the reactor safety issues to which the research has been specially addressed

is the reflooding phase of the LOCA: indeed, the accurate calculation of the rod temperature during this

period is of importance for the estimation of the safety margins required for the reactor operation.

The main results of this research [1],[2],[3],[4] and the contributions to the code development in the areas of

wall heat transfer [5],[6], interfacial drag [2],[6], and numerical techniques [6],[7] have been fully documented

and examples of references have been provided here. For several reflooding experiments, it has been observed

that both the trend and the values of the calculated results (wall and vapour temperatures, entrainment,

collapsed liquid level, etc.) obtained by the 'frozen' version of the codes were very different to the measured

experimental data. The discrepancies between the calculations and the data were noticeably large, especially

for low flooding rate conditions. Low flooding rate experiments are characterized by very high void fractions

all over the region above the QF for long time periods; in this zone droplets are dispersed in a stream

of superheated vapour. The incorrect calculation of either the vapour superheat or the wall heat transfer

coefficient resulted in the poor predictions of the 'frozen' versions of the codes.

The implementation of more consistent packages of correlations for wall heat transfer and interfacial shear

(for both the wet and dry wall regions), the accurate selection of several empirical constants entering in the

various models on the base of sensitivity studies, and the modification of interpolation schemes and numerical

techniques lead to improved versions of the programs, which successfully simulate most experiments with

bundles [2],[3],[51,[6],[7].

Modifications to the closure laws (largest droplet size, interfacial heat transfer) related to the dispersed flow

regime above the Quench Front (QF) have also played a role in the achievement of good results, and more

research is underway in this area. These studies are important contributions to the development of reliable

calculation tools for reactor safety assessment, as the test facilities considered for the assessment are closely

reproducing the conditions expected in the nuclear cores.

However, since a few quantities (average temperatures, entrainment rate, quench front velocity, etc.) are

measured and can be used for the assessment, the adequacy of the closure laws used for the" individual

submodels can seldom be demonstrated. In principle, a fortuitous compensation of errors and the lumping of

different effects in one single 'best fit' parameter, could mask the deficiencies in the models implemented in

the codes.. For instance, the lack of the representation of the spacer grids and an incorrect model for interfacial

heat transfer and interfacial shear could be 'compensated' by a proper, although unphysical, choice of the

empirical criterion for defining the droplet diameter.

In the present report, a different approach to the analysis of the reflooding process is presented, which is

more suitable for evaluating the capabilities of one of the submodels used in the codes, namely that for

interfacial heat transfer in the dispersed flow region. The attention to this particular model is motivated by

the results of a recent investigation [8],(9] on post-CHF heat transfer with dispersed flow during reflooding.

It has been shown that, under conditions of low mass flux and low quality at the QF, the interfacial heat

transfer must be overpredicted by the l-D models and some 3-D (such as those implemented in RELAP and

TRACI series of codes) and this results in large underpredictions of the vapour superheat. It was argued

that the neglect of the cross-sectional distribution of the droplets can be the reason for these discrepancies.

The clustering of the droplets around the centre with the consequent reduction in the interfacial heat transfer

has been proposed [8],(9] as a possible mechanism for explaining the high superheating rates (close to those

admissible for 'frozen flow') at some distance from the QF observed in several tube experiments [10]: this

conclusion was obtained using the commonly accepted closure laws for interfacial drag and droplet diameter.

The present analysis using RELAP5 is, therefore, aiming at a verification of such a general statement by

a specific application of the I-D approach. In this respect, the most interesting variable for judging the

'The TRAC code is often quoted as 3-D code, as it can provide information on the 3-D distributions in die corer the large

sizes of the radial mesh which can practically be used. however, do not allow to calculate the radial profiles inside dte individual

jub-dcannels.

6
capability of the model is the vapour temperature at the exit of the test section, and this will be the main

object of the discussions of the results.

For this purpose, it is convenient to assess the codes against data obtained in single-tube experiments; the

simpler tube geometry allows to better test the interfacial heat transfer model, as the 'spurious' effects of the

grids (which cannot be quantified and cannot be taken into consideration by the code, but are present in the

experiments with bundles) on the droplet size and other variables, will not bias the comparison between the

calculated vapour temperature and the experimental data. The exercise presented in this work is the transient

analysis of the low flooding rate experiment 3051 at 2 bar in the single tube test section of the University

of California-Berkeley (UCB) by means of frozen versions of RELAPS/MOD2.5 (Mod2, cycle 36.06)(1 1]

and RELAP5/MOD3 (7vj)[12]; this experiment was chosen because the separate steady-state analyses [9]

showed severe difficulties of the I-D approach.

However, while the transient analysis of the core is the only important application of the reactor safety

codes, much deeper insight in the influence of the interfacial heat transfer on the mixture evolution is made

possible by considering 'snapshots' at different times during the transient. The values of several variables

are considered constant over a short time period, and the analysis is referred to such a 'frozen' picture.

The rationale for the consequent 'quasi steady-state analysis' is shortly outlined here, and discussed more in

depth in §3. For the 'steady-state' calculation, the conditions above the QF must be imposed (approximately)

constant over a short time, and the wall heat flux over the tube length above the QF must be given as input:

if good predictions for both liquid and vapour mass flow rates are obtained, the resulting vapour temperature

at the tube exit (which depends on the total heat exchanged with the droplets) will give a direct estimate

of the adequacy of the interfacial heat transfer closure law used in the code. Indeed, the uniform heat flux

which characterizes the single-tube experiments leads to slowly changing axial wall temperature distributions

and allows a sufficiently good estimate of the local heat flux at all elevations. This, in turn, allows a rather

accurate evaluation of the total power input to the mixture above the QF (on which the vapour superheat

strongly depends) at any selected time, so that one of the unknowns which complicates the interpretation of

the results of the transient analyses, namely the instantaneous 'actual' power input, can be eliminated.

Therefore, in the present work the quasi-steady state analyses of two UCB experiments (Run 3051 at 2 bar

and Run 3053 at 3 bar) for the time of the mid-height quench are also carried-out.

2 Test cases used for the assessment

Well documented experimental results for the reflooding of tubes were obtained at the University of California

at Berkeley [13],[14]. The diagram of the experimental test facility is shown in Fig. 1. The test section used

for the second series of experiments at 2 and 3 bar was a 14.25 mm inside diameter (0.83 mm thickness)

tube, 3.67 m in length. It was quenched by bottom injection of water at flooding rates from 25 to 180 mn/s:

the water subcooling, about 80 K, was the same for all the experiments. Constant electrical heat input was

used, and the cooling of the tube was initiated after an equilibrium condition had been reached- initial wall

temperatures ranged between 316 OC and 760 OC. The tube was not insulated, and large heat losses have

been allowed. The transient was initiated from a condition of constant wall temperatures, the power input

being balanced by the heat losses. The initial axial wall temperature profile was practically flat.

The temporal variation of the tube wall temperature at different elevations and of the fluid temperature at the

tube exit were obtained: the vapour temperature was measured by a sheathed microthermocouple (0.254 mm)

located at about 5 cm above the tube exit. The flows of liquid and vapour at the tube exit were separated by

means of a carefully designed separator, and the entrained liquid was collected in a tank. The vapour mass

flow was also measured. Moreover, the water collected inside the test section (from which the collapsed

liquid level can be calculated) was obtained by measuring the pressure difference between inlet and outlet

of the test section, and neglecting friction and acceleration pressure drops.

During many tests, the wall temperatures at some distance from the QF set to nearly constant values for

long times, being the heat input balanced by the heat losses and the heat transfer to the fluid: typical wall

temperature time-histories are shown for Run 3051 (experimental conditions specified below) in Fig. 2.

7
In the present investigation, the conditions at the earliest time such a steady condition (the quality at the

QF being positive) for the wall temperatures at the higher elevations is attained are chosen for the 'quasi

steady-state' analyses.

In Table 1 the experimental conditions of the two runs used for assessing the codes are summarized.

For the quasi steady-state analyses, the additional information on the equilibrium quality at the QF Zqp and

the axial distribution of the net heat input to the fluid qj", at the selected time are also required.

The equilibrium quality immediately above the quench front zQF is calculated (15] by adding to the quality

immediately below the Quench Front z- (calculated by an energy balance over the wet region) the quality

change across the Quench Front due to the intense vaporization caused by the release of the heat stored in

the tube wall:

X+ X + C.A.(T+ - T-)Uq

H15GA

with.

Put (• AH.,&

HfsGA H1 ,I

where C. is the volumetric heat capacity of the wall, A. the cross sectional area of the wall UQF the

quench front velocity, H1, the vaporization enthalpy, G the total mass flux, A the cross sectional area of the

channel, T+ and T- the wall temperatures above and below the quench front, AH,,b the liquid subcooling

at the inlet of the test section, L the total length of the tube test section, zQF the distance of the QF from

the tube inlet, and Pst the input power

The net heat flux %I, at each elevation is obtained by subtracting from the inside-wall 'imposed' heat flux

q' (Po, divided by the internal surface area), the heat losses q,,j (referred to the inside-wall surface area)

and the sensible heat stored in the wal:

qj" = q. - q,,. - C, d-- (2)

where C" is wall heat capacity per unit inside wall area and T.1 the value of the wall temperature (in K)

obtained by polynomial interpolation of the experimental temperatures. The outside heat flux (losses) is:

=hlC(T. - Tab) (3)

where hL is the global heat transfer coefficient (HTC) from the wall to the surroundings, assumed to be at

temperature T.,b=294 K The experimental values of hL could be correlated by the formula:

hL = 11.949 e[2'2"2 3

-" (".-T....)" (Wlm2 K) (4)

The heat losses per unit area (heat flux to the ambient) referred to the inside-wall heat transfer area is obtained

by:

D4

where D. and D are the external and internal diameters, respectively.

The total heat input to the fluid in the dry region above the Quench Front P&, is calculated by integrating

q' over the wall heat exchange area.

The values of the several variables required for the quasi steady state analyses are given in Table 2 together

with the transient times to which they refer.

8
Run p G Pot~

No. bar kg/mh s kWs K

30511 2.125.1 14.326 812.

3053 3. 125. 1 4.326 812. I

Table 1: Experimental conditions for test cases UCB-3051 and UCB-3053

Run t zQr Uqr T- T+ z- zQF Pd,.r

No. s m mm/s K K W ) kWim 2

3051 150. 1.90 7.3 401. 593. 0.10 .12 923. 2.00 16.91

2.20 15.45

2.40 13.92

2.60 12.38

2.80 10.97

3.00 9.69

3.20 8.83

3.40 8.19

3.58 8.19

3053 150. 2.05 7.5 414. 593. 0.12 .14 874. 2.20 15.19

2.40 14.20

2.60 13.60

2.80 12.98

3.00 12.10

3.20 10.97

3.40 9.90

3.58 9.89

Table 2. Calculated variables used for the steady-state analyses.

9
3 Transient analysis of test case 3051

The basic nodalization scheme for the analysis of Run 3051 is shown in Fig. 3. It consists of one pipe

(component 120) with 19 cells for the test section, one time-dependent junction (110) for the bottom injection,

one junction (125) representing the tube exit, and two time-dependent volumes (105 and 130). The heated

wall of the test section is represented by the same number of heat slabs (19), and uniform linear power

generation is imposed. The heat losses (calculated from a temperature-dependent heat transfer coefficient

hL) to the ambient depend on the external wall temperature, according to correlations (3) and (4). The

values of the heat transfer coefficient hL has been calculated for several temperatures, and a table of hL

vs temperature has been used for the input deck. The ambient temperature has been imposed constant and

equal to 294 K. At the beginning of the transient (t=O s) the test section is filled with superheated steam in

thermal equilibrium with the wall of the tube (T,=T,,,-812 K). The transient is initiated by starting (t=0 s)

the bottom injection of water.

The transient was calculated with the 'frozen' versions of both RELAP5/MOD2.5 and RELAP5/MOD3. For

both calculations the correlation for interfacial drag in tubes was chosen. The reflooding option was chosen

for the calculation with both versions of RELAP5, so that moving fine heat structure meshes were used in

the wall heat transfer calculation. The reflooding heat transfer package was then utilized in the calculation

with RELAP5/MOD2.5, whereas this option does not affect the correlations used by RELAP5/MOD3, as a

unique heat transfer package is used for all conditions. The input decks for the two calculations are given

in App. IThe calculation by RELAP5IMOD2.5 could not be completed using the nodalization with 19 vols. shown

in Fig. 3 (left), as a condition of error (which could not be explained) occurred, at about 160 s, in the

subroutine QFMOVE, which calculates the advancement of the quench front(at this time, the quench front

elevation was already above 2-44 m). Reduction of maximum and minimum time steps, and changes in the

number of radial meshes and maximum number of fine meshes only provoked a negligible time shift of the

error condition. Eventually, the solution was to halve the size of the hydrodXnamic nodes (from 0.2 to 0.1

m) and, correspondingly, of the heat slabs. The resulting nodalization with 36 vols. is also shown in Fig. 3.

The comparisons of the wall temperatures calculated by both codes with the measurements are shown in

Figs. 4 and 5. Even though some minor differences can be observed in the results of the two calculations,

the most striking outcome of the analysis is the very large discrepancy between calculated and measured

wall temperatures, as in (4]. The test section quenches completely in about 550 s, while both codes predict

complete quench within 160 s. While RELAP5/MOD3 predicts correctly a gradual bottom quench progression, RELAPS/MOD2.5 calculates a second quench front, propagating from the top of the test section, as it

can be recognized (Fig. 5) from the early quench at 3.35 m (highest measurement station). The results are

somewhat surprising, if one recalls that most of the analyses with bundles (e.g., NEPTEJN with cosine-shaped

heat flux profiles, but also with several spacer grids) indicated the tendency of RELAP5/MOD2.5 to overpredict the quench times and the wall peak temperatures, especially at the highest elevations [4]. These results

are also in conflict with recent assessment work using data from the Lehigh bundle reflooding experiments,

where a uniformly heated bundle was used: for these cases RELAP5/MOD2.5 was generally predicting peak

temperatures and quench times with acceptable agreement in comparison to experimental data, and the only

underpredictions were found for low elevations (4]. This last assessment work can, however, be barely

associated to the previous ones, because of the short length of the test section; in fact, the analysis was

focussed on lower elevations than in the NEPTUN case.

RELAPS/MOD3, on the contrary, behaves in accordance with the previous experience [4]: it overpredicts

the quench front velocity (or rather the 'cooling front' velocity, as no evident temperature 'knee' can be

observed). For the present tube experiment the effect is amplified, as the QF velocity is four times higher

than in the experiment.

Moreover, it can be observed (Fig. 6) that early in the transient (30 s for the MOD2.5 and 100 s for the

MOD3), the calculated vapour temperatures in the uppermost node drop to low values (about 420-450 K)

close to the saturation temperature (about 400 K), whereas the experimental value sets for a long time to a

value of around 600 K.

10
The reason for such a large disagreement between the predicted and the measured quench front velocities

is not easy to single out, as several effects can contribute to the inaccuracy of the predictions. On the one

hand, too large wall heat fluxes might have driven the wall to rewetting conditions, with consequent drop

in the vapour temperature to the saturation value. Moreover, large mass flow oscillations affect both code

versions (in Figs. 7 and 8 the vapour and liquid carry-out, as well as the total mass flow rate at the tube

exit are shown), and most probably contribute to the excessively fast propagation of the quench front. On

the other hand, the large temperature difference between the wall and the vapour might have resulted in an

additional increase of the already over-predicted wall heat flux [6] and, consequently in an overprediction of

the precursory cooling.

From the comparison of the calculated mass accumulated inside the tube, the collapsed liquid level, the

vapour flowing out of the test section and the total liquid carry-out (Fig. 9), with their respective experimental

counterparts, one can conclude that both code versions predict the 'average' hydraulic behaviour of the system

fairly well.

It should be noted here that the calculation of the total liquid carry-out by RELAP5,MOD2.5 is actually

better than what appears in Fig. 9, as the control variable (cntrlvar) which integrates the liquid expelled over

the time is affected by an error, and underestimates the total liquid carry-out. A short explanation can be

given as follows. The mass flow rate of the phase k at a junction j (in the specific case the junction 125) is

obtained from:

k = pkjcrtjt?&jA

and the total mass flow can be calculated by summing up the gas and the liquid mass flows. The total mass

expelled out of the test section k., is then:

= int(cntdhvar) = + -,s) dt (5)

On the other hand, the total carry-out can be calculated by integrating directly the total mass flow at the tube

exit -Mdi, which is also calculated by the code:

= hit(mflowjl125soo ) j)= d (6)

The comparison of the results of the two expressions is given in Fig. 10: it is observed that a 10% difference

exists in the RELAP5/MODZ5 calculation, while a much smaller discrepancy appears in the RELAPSIMOD3

calculation. As the value of ,,, calculated by Eq. (6) satisfies the mass balance with a high accuracy, it can

be concluded that the mass flows calculated by Eq. (5) underestimate the 'actual' mass flows. Indeed, under

rather unstable liquid mass flow conditions, the spikes in the mass flow rate (due to the spikes in the liquid

mass flow rate) calculated by Eq. (5) are much lower than those present in the mass flow time-history of the

variable mflowj1250000 (see Figs. 7 and 8). This can be explained by the different numerical techniques

used for obtaining the time and space averages of the various quantities at the junctions in these two different

versions of the code.

Even though the comparison between calculations and experiment regards only the exit conditions, it can

be assumed that both wall and interfacial drag are correctly predicted over the transient' and that the axial

liquid mass distribution has also been captured with a good approximation.

Using the hypothesis that the low vapour temperatures drive the wall heat flux increase and the wall temperatures drop (and not vice versa), one could then argue that the failure of the RELAPSIMOD3 in predicting

vapour and wall temperature, is an overprediction of the interfacial heat transfer, which results in an anticipated desuperheating of the vapour up to large distances from the quench front. An indication in favour of

this explanation can be obtained from observing, for example, that at the time of the quench at 2.58 m (t-93

s, interpolated value) the vapour temperature at the tube exit (1.2 m downstream) is about 550 K, whereas

'This does not nec=sarily imply that interfacia and wall drag models are correct.

11
in the experiment at the time this happens (t=298 s) the vapour temperature was 50 K higher. However,

the comparison is not fully meaningful as one cannot compare the actual quality at the quench front (highly

oscillating in the calculation because of the high heat transfer rates across it) and the total heat input to the

vapour in the dry region. In fact, one also observes that, because of the high cooling rates (due also to

the large wall-to-liquid heat transfer) at large distances from the QF, the calculated and experimental wall

temperatures and temperature decrease rates in the dry region are quite different at the time of the 2.58 m

quench. This implies different values of the local heat fluxes and, then, of the total heat input to the vapour.

Therefore, because of the complex interactions between the different heat transfer processes, the transient

effects and the history-dependent thermal-hydraulic conditions, the boundary conditions which influence the

vapour superheating rate are different in the calctilation and in the experiment, so that the model for interfacial

heat transfer in the dispersed flow regime cannot be judged conclusively.

In the case of the MOD2.5 calculation, the occurrence of top quench confuses the picture even more, as

the sudden decrease of the vapour temperature close to the exit is not due to the excessive interfacial heat

transfer, but to the fact that the vapour cannot receive heat from the wet wall any longer and a large quantity

of saturated vapour is generated from the direct wall-to-liquid heat exchange.

From the analysis of this tube experiment, it can be deduced that:

1. The analysis of experiments with tubes can yield rather different results than those obtained for bundles,

even for very similar conditions. The results for tubes, even though they are not truly representative

of the conditions in a nuclear reactor, provide a useful test for the adequacy of the correlations used

in the computer codes, as most of the models implemented are primarily derived for circular channels

(the application to bundles implies the use of some 'equivalent' diameter). This can be recognized

from the excellent prediction of the hydraulics.

2. The calculated quench front progression is much faster than in the experiment for both versions of the

codes. The discrepancy is due to a too large wall heat flux [6] in the post-CHF regime, to which the

underprediction of the vapour temperature also contributes.

3. Run 3051 provides a very severe test for the codes. The difficulty of the 1-D approach in predicting

interfacial heat transfer demonstrated for more specific conditions (9], cannot be definitely confirmed

here, but the large discrepancy in the calculated vapour temperature seems to be consistent with those

phenomenological considerations. In particular the MOD3 calculations seem to be affected by the

overprediction of interfacial heat transfer, but both calculations are probably dominated by deficiencies

in the wall heat transfer models.

4. The transient analysis above does not allow to draw any conclusion about the correctness of the

individual models, as they (and their numerical implementation) interact with each other.

With respect to the main goal of the present work, namely the assessment of the interfacial heat transfer

model for Dispersed Flow, not much can be concluded from the transient analysis; hence, the analysis of

experiment 3053 (for which results similar to those obtained for test 3051 are expected) was not carried-out.

4 Quasi-steady state analysis of Run 3051

The basic idea behind the quasi steady-state analysis has already been introduced in § 1; here the mathematical

formulation of this approach, as well as the conditions which must be fulfilled in order to arrive to the

evaluation of the capabilities of the interfacial heat transfer model using only the experimental vapour

temperature at the tube exit, are discussed.

When the quenching process progresses slowly and at constant rate, as usual when the quality at the QF is

positive, the thermal-hydraulic conditions at any point in the channel downstream from the QF (wall and

fluid temperatures, mass flows, qualities, etc.) vary slowly as well. The quasi steady-state condition is thus

realized if the time necessary to the QF to move a significant distance (e.g., the distance between the centres

12
of the volumes above and below the 'nominal' position of the QF) is larger than the residence time of the

droplet in the post-CHF region. If this prerequisite is satisfied, any slow variation of the QF position (and

quality) will lead to slow variations of the hydrodynamic variables downstream. Under this condition, the

evolution of the mixture can be described by the usual conservation equations, dropping the time-derivatives.

The steady-state form of the vapour continuity and energy equation can be written for a straight tube (like

in the RELAPS code [11], but neglecting dissipation and using the enthalpy instead of the internal energy)

of cross-sectional area A as:

d" 7z(< ýG(7)=< s r ><p, > >u )

< a,>< p,> U"s-H 6 q' = =Q.,+ <Qi,,> - < r > (H- H,..)()

'dz 'dz

where U,, Hb, G#, Q.# and < Q,, > are the cross-sectionally averaged axial vapour velocity, the bulk vapour

enthalpy, the vapour mass flux, the phasic wall heat transfer rate per unit volume and the cross-sectionally

averaged interfacial heat transfer rate per unit volume, respectively.

By integrating between the QF elevation zq. and the tube exit elevation zT Eq. (8) multiplied by the tube

flow area A, one obtains:

IAG, " = j Q,0,Adz + < <Qi, > Adz - r(H6 - Hq,,)Adz

and, integrating by parts the first term in the Left Hand Side of the equation:

,T dG = 7 Q.~,Adz +

AG#'H& - AG#,,qHs.QF - ficpAt d

< Qj, > Adz +

- r(Hb - H,o.)Adz

Using the continuity equation (7) and replacing GA by the vapour mass flow rate M,, it follows:

&IqHbT-M,.QFHb.QF = [?Q,.Adz + < Qi, > Adz + TrH 1 Adz

IAQP XQPAg

f j (H, - H,,,)Adz

Observing that:

JrAdz-M=-&ITQF

one can write (assuming saturated vapour at the QF):

&#THT q+FH&QF+ P y+J <Q,,> Adz + (.•--,.qF)H',. (9)

where P,.dr, is the total heat transfer rate to the vapour alone in the dry-wall region. If all the droplet

population is represented by an average diameter d, the averge interfacial heat transfer rate per unit volume

is:

13
6h•

< Qs >=-h- •< (1 - a) (T -T)> (10)

Since in a 1-D model the droplet concentration is assumed uniform over the cross section and the distribution

coefficient is taken equal to unity, the relation:

< ( ,)( T - T.) ><( >< (T - T.) >=-< (1 -as > T - T.) (11)

leads to:

< Qf >= 6h- < (1 - as) >< (Tb- T,) (12)

This is the relation that under low mass flux and low quality conditions, will lead to the overprediction of

the interfacial heat transfer, and the underprediction of the vapour temperature (9]. To show that RELAP5

is affected by large inaccuracies in the prediction of the vapour temperatures because of the large interfacial

heat transfers, and then that Eq. (12) does not hold, is the goal of the following analysis.

From Eq. (9) one easily realizes that if the gas mass flow rate at the QF and the tube exit calculated by

the computer code are (approximately) equal to the experimental ones, and the total wall-to-vapour heat

transfer is the same as in the experiment, any deviation of the exit enthalpy (temperature) can be attributed

to the incorrect prediction of the interfacial heat transfer. This estimation can be done irrespective of the

wail-to-liquid heat transfer and the consequent vapour production at the wall.

Tlerefore, for the quasi steady-state analysis four conditions (Fig. 11) must be fulfilled in the calculation:

1. the time needed by the the QF to advance a substantial length about the 'reference' QF position is

much longer than the droplet residence time t,. (- (z - zQF)/UI). The average values of the variables

during this time (which we will call 'averaging period' t.,) can be compared with the values obtained

by a steady-state analysis.

2. same (approximately) zqr (k,,QF) as in the experiment during the averaging period t,,.

3. total heat input to the vapour equal to the experimental one.

4. same k as in the experiment.

The first two conditions can be imposed by restricting the analysis to a section of the tube including the

region downstream of the 'nominal' position of the QF and a short length upstream of it. The fluid injected

at the bottom of this 'reduced' length section is a mixture with a quality close to that inferred from the

experiment. The QF progresses from the inlet of the section up to the nominal location of the QF (and the

quality at the QF changes accordingly), for which the tube exit vapour temperature must be compared with

the experimental value. From this time on, and until the time the first node above zq- quenches, the values

of the various variables of interest are considered.

With respect to the third condition, certain assumptions and approximations are needed, since only the total

heat flux to the mixture is known at the few measurement stations. As the experimental wall temperatures

reach nearly steady values at some time during the run, and drop suddenly only when the QF is very

close to the measurement location, the interpolation of the heat fluxes for any elevation of the tube can be

calculated with good accuracy, and the total heat input to the mixture in the dry region of the tube can then

be estimated with reasonable confidence: with this procedure the values in Table 2 were obtained. Still the

partition between the two phases is unknown, so that two approaches can be used in the calculation:

Impose the total heat flux equal to the experimental one, and let the code partition the heat between

the phases. If this option is chosen, the wall-to-liquid heat transfer must be small at least at some

distance from the QF, otherwise an underprediction of the vapour temperature can be due to the low

heat input to the vapour rather than to the overprediction of the interfacial heat transfer.

14
e If the code calculates a too large wall-to-liquid heat transfer, increase the total heat flux as much as

necessary to transfer to the vapour an amount of heat equal (approximately) to the total experimental

one. This second option assumes no wall-to-liquid heat transfer in the experiment. Therefore, this

procedure uses the upper limit for the heat input to the vapour, and, if the vapour temperature is still

underpredicted, the deficiency in the approach for calculating the interfacial heat transfer will become

more evident.

For the fourth condition, namely calculated vapour mass flow close to the experimental value, one can only

hope that the computer program predicts liquid and vapourcarry-out flow rates close to the experimental ones

at the time of the mid-height quench. This expectation is conforted by the results of the transient analysis

above, where both calculated liquid and vapour carry-out were in excellent agreement with the measured

ones. For the purpose of testing the interfacial heat transfer model, lower vapour mass flows than in the

experiment can also be accepted, as the calculated vapour superheat is enhanced by a reduced vapour mass

flow. Similarly, the liquid carry-out must be approximately equal to or lower than the experimental value. A

calculated liquid mass flow much higher than in the experiment would imply a large concentration of liquid

at high elevation: the consequent vapour desuperheating effect could not be related to the l-D approach for

the interfacial heat transfer.

In order to avoid misunderstandings, it is appropriate to repeat here that, as RELAP5 is a transient code,

our 'steady-state' analysis consists in running -the program over the short time period needed by the QF to

advance from the inlet of the tube to the node immediately above the elevation ZqF, chosen for the analysis.

In summary: The analysis will be restricted to a shorter tube length. The quench front moves from the

inlet to a certain predetermined elevation above zqF. From the time the QF is close to zQF, the quality

immediately above must be close to ZQF, and the power input to the vapour in the dry region must be

equal to P,,,; this last condition can be approximated by imposing that the total heat transfer rate to the

mixture is equal to the (experimental) P4f... The mass flows at the tube exit will be observed; if they are in

close agreement with the experimental data (obtained by differentiating the experimental water and vapour

carry-out), the vapour temperature at the tube exit will be compared with the experimental data.

4.1 Calculation with RELAPS/MOD2.5

In Fig. 12 the nodalization used for the 'steady-state' analysis with the MOD2.5 is shown. It consists of

ten short (0.03 m) cells below zQF, and nine cells representing the above-QF region having the same size

as in the 19 vols. nodalization (Fig. 3). The inlet quality zi, was set by a trial-and-error procedure, by

checking the value of zQ- at the end of the transient, as calculated from an energy balance performed by

means of a control variable; at the end z,, has been set equal to about 6%. The need for this procedure

is due to the oscillations in the calculation of the mass flows immediately above the QF, which render the

direct estimation of the quality impossible.

The heat flux at the wall was initially imposed according to Table 2, the initial vapour temperature in the dry

region was given a linear profile (the experimental temperature at t=150 s being used for the highest node),

and the experimental wall temperature was imposed to the heat structures. Unfortunately, the calculation

stopped after a few seconds, and the reason was easily found: the vapour temperature upstream of the quench

front was exceeding 3000 K. This happened because a very large amount of heat was transferred directly to

the vapour, in spite of the fact that a bubbly flow regime was predicted. It was realized that RELAP5, for

imposed heat flux conditions, calculates the heat fluxes to the two phases by splitting the total heat input

into two equal parts, irrespective of the flow regime.

It was then necessary to change our approach of imposing the heat flux, and other boundary conditions were

used: •

" Impose the power generation.

" Impose the experimental heat losses at t=150 s.

15
e Impose the axial temperature profile. As the heat flux is then calculated by the reflooding heat transfer

package, the total heat input to vapour is not controlled, but varies during the short period of time

of the 'quasi steady-state' analysis, and depends mainly on the initial wall temperature. Therefore,

three different axial wall temperaure profiles are imposed, and the resulting heat input to the vapour

is calculated. The three axial wall temperature profiles, shown in Fig. 13, are:

1. exp: experimental wall temperature distribution.

2. low: temperatures decreasing with the distance from the QF. These values have been found (by trial

and error) to yield (approximately) the desired axial heat flux distribution (Fig. 14).

3. high: temperatures higher than those experimentally observed.

The resulting wall heat fluxes averaged over the last 5 s of the calculations are shown in Fig. 14. The input

deck for the calculation using the experimental temperature distribution is given in App. II.

The results of the three computations are shown in Figs 15-21. In Fig. 15 the wall temperatures at several

elevations are shown for the low initial wall temperature distribution. The temperature time-histories in the

other two cases were similar. Two remarks can be done: 1) the wall temperatures at some distance from

the QF (z 2t 2.2 m) are practically constant during the calculation time, and 2) in the time period between

10 and 15 s the QF lies somewhere between 1.855 m and 2.0 m. The time needed for the QF to advance

from 1.855 m to 2.0 m was found to be larger than the droplet residence time t, (= 1-2 s), so that the first

requirement for the 'steady-state' analysis is satisfied.

It is first observed that the equilibrium quality at the QF (Fig. 16) varies during the transient time, but is

close to the experimental value at least for times larger than 10 s.

The vapour mass flow (Fig. 17) at the tube exit is also in close agreement with the experimental value. On

the contrary, the liquid mass flow rate (Fig. 17) exhibits large oscillations: their, amplitude, however, do

not exceed 30% of the measured average flow rate. The fact that the average values (between 10 and 15 s)

of the total mass flow rate and exit quality (Fig. 18) are quite close to the experimental values, suggest that

the hydraulic behaviour is in general well predicted.

The otal power input to the vapour (Fig. 19) is, in all cases, at least as high as the total heat input to the

mixture, so that also the thermal boundary condition on the vapour superheat evolution is satisfied in both

cases.

Under the conditions mentioned above, the average values of the several quantities of interest for our analysis

during the time period from 10 to 15 s are thus comparable with those obtained by a steady-state analysis

for QF 'frozen' at 1.89 in; then, it does make sense to compare the calculated vapour temperature with the

measured one. In Fig. 20, it can be observed that after 10 s the vapour temperature (while being unstable)

is always underpredicted by at least 50 K, even when the power input to the vapour is twice as much as the

experimental total heat transfer rate to the mixture (Fig. 19). It is easily realized that the spikes are due to

instantaneous increases of the void fraction due to sudden drops in the liquid mass flow rate (liquid carry-out).

In Fig. 21 the average value of the vapour temperature is shown, and the large underprediction (150 K)

observed for the initial wall temperature distribution which yields value of Pd,. close to the experimental

ones is in fairly good agreement with that expected on the base of the quasi steady-state separate model [9].

From the RELAP5/MOD2.5 calculations, we can thus conclude that the interfacial heat transfer must be

overpredicted. Let us discuss the possible sources contributing to the overprediction:

1. The liquid fraction might have been overestimated. This, however, can be almost excluded by the

fairly good prediction of the liquid carry-out, which suggests a correct estimate of the axial liquid

distribution.

2. The interfacial heat transfer coefficient (h,) is possibly overpredicted because of a large average value

of the phase velocity difference. Indeed, one can suspect that large oscillations in the liquid velocity

(leading even to negative velocities) can still be compatible with a correct average value of the liquid

16
mass flow, but yield (because of the non-linear dependence of the heat transfer coefficient on the phase

velocity difference (U, - Ua)) large values of h.. This can be also excluded by observing that the

velocity difference at two elevations (Fig. 22) are quite stable.

3. The droplet diameter might be underpredicted. The separate steady-state analyses [8] show, however,

that for a wide range of droplet diameters the results are very similar to those obtained using the

RELAPS/MOD2.5 criterion.

4. The only possibility that is left is, thus, that the interfacial heat transfer rate is based on the incorrect

assumption of a uniform droplet concentration over the cross-sectional area.

Other calculations (not shown) have been carried-out using the normal (blowdown) heat transfer package,

and slowing down the propagation of the quench front by increasing the thickness of the heat slabs in the

region below the QF. The vapour temperature at the tube exit was still underpredicted by at least 100 K.

Furthermore, a version of the code was generated, in which the wall-to-liquid heat transfer was forced to go

to zero within 30 cm from QF. The axial wall temperature profile was imposed in such a way to preserve the

total experimental power input to the dry region (Pft), so that the heat transfer rate to the vapour (P,.,)

was practically equal to the total experimental one. Even under these conditions, the calculated vapour

temperature was still much lower than in the experiment.

4.2 Calculation with RELAPS/MOD3

For the analysis with RELAPS/MOD3, the nodalization had to be slightly modified. This was due to the reason

that the first calculation (not shown by a figure) showed that the quench front was progressing much faster

than in the corresponding RELAP5/MOD2.5 calculation. Under these circumstances, the time necessary for

the QF to propagate from the node below zQF to the node above it (time t.. over which the thermal-hydraulic

conditions of the post-CHF region can be considered approximately insensitive to the actual location of the

quench front around the imposed zqF) became comparable with the droplet transit time: the averaging of

the variables over the time period t.. was consequently becoming of dubious meaning. Therefore, larger

nodes (50 mm instead of 30 mm) were used for the section below ZQF and the node including ZQF (Fig.

23). Moreover, the time, over which the average values of the variables discussed above are calculated, is

made dependent on the transient, i.e., the averaging period is between the time the wall temperature in the

cell below z

2 r (cell 9) drops below the quench temperature (580 K) and the time this happens to the first

node in the 'dry' zone (cell 11). When the wall temperature in node 11 drops below 580 K the transient is

thus terminated.

The first calculation, performed by using initial wall temperatures equal to the experimental ones, also showed

a very high heat transfer rate in the dry zone, nearly double as much than in the experiment, and a much

larger vapour mass flow, resulting from the large evaporation rates caused by the high wall-to-liquid heat

transfer.

Therefore, the low axial wall temperature distribution (Fig. 13) was used for the next calculations. The

results are shown in Figs. 24-29. The wall temperatures are shown in Fig. 24: at large distance from the

QF they are slowly growing, and the QF lies between 1.825 and 2.0 m for the time between 8 and 12 s. The

equilibrium quality at the QF is shown in Fig. 25. It can be remarked that it is highly oscillatory (often,

the case found in the RELAP5/MOD3 calculations (4]), due to the large spikes in the heat transfer from the

wall to the fluid in the wet region. The lower value of the equilibrium quality, which is representative of

the 'steady state' heat transfer rate, is, however, close to the expected value.

The vapour temperature at the tube exit (Fig. 26) is predicted quite well, but one can easily realize from the

good prediction of the vapour carry-out (Fig. 27) and the negligible liquid carry-out (Fig. 28) that this is due

to the deficiency of liquid in the region far away from the QF. Fig. 29 shows the void fractions at various

elevations. The water is accumulated in the region immediately downstream from the QF, so that no heat

sink is available to the steam further up. It was originally believed that one of the reasons for the insufficient

liquid carry-over could be the larger value of the critical Weber number (Wec,=12) used in the MOD3 with

17
respect to that used in the MOD2.5 (Wee,.=3), which implies the calculation of quite large droplets in the

mist-flow region and these drops cannot be lifted by the drag force (see below).

It was then decided to perform the calculation with a modified version of the code developed at PSI (version

bb) which, among other features, uses the same We,. as in MOD2.5. This modified version has been

discussed by Analytis (6], and the description of its features will not be repeated here: it suffices to say that

better numerical stability is achieved and several unphysical effects (like the void fraction passing through a

minimum immediately below the quench front) are eliminated by the utilization of more physical correlations.

The results are shown in Figs. 24-28 and 30. It can be observed that the liquid carry-out (Fig. 28) is a

little higher than before, but still much too low: however, the resulting slight modification in the axial void

fraction profile (Fig. 30) reduces the average outlet vapour temperature (Fig. 26) by about 30 K.

In order to understand the reasons why the utilization of the same criterion for the droplet size yields different

results in RELAP5/MOD2.5 and in RELAP5/MOD3, the overall calculation scheme of the interfacial drag

in RELAP5/MOD3 was reviewed. It was realized that, for reflooding calculations, the droplet diameter was

bound by a lower limit of 2.5 mm. This value appears arbitrary, and for sensitivity analysis purposes it was

changed to 1.25 umm in a new version of the code, which was called UPX. Moreover, it has been found out

that the drag coefficient for droplets is defined as:

CD = min{CD,., 0.45} (13)

where the drag coefficient for the viscous regime CD., is given by:

Cv,, = 24 + .Re (14)

It is clear from Eqs. (13)-(14) that for large Re (large droplets) the drag coefficient attains very low values.

However, the drag coefficient for a solid sphere is asymptotically approaching the value 0.45 for Re> 1000

(Newton regime), being higher for lower values of Re (16]. The value of CD for fluid particles is even

higher, as they get distorted (16]. Eq. (13) is then basically wrong and even if one wants to consider the

droplets as solid spheres, it must be replaced by the usual criterion:

CD = max{CD., 0.45} (15)

Eq. (15) was then implemented in the version UPX.

The results of the calculation by version UPX are shown in Figs. 25-28 and 31. It can be remarked that

the liquid carry-out (Fig. 28) is now larger than in the previous calculations, but it is still too low. The

axial void fraction distribution (Fig. 31) shows a larger presence of liquid at the higher elevations, and this

determines lower vapour temperature (Fig. 26).

It can be observed that the calculated heat transfer rate to the mixture (Fig. 32) was always higher than in

the experiment, but the heat transfer rate to the vapour was about 30% lower (Fig. 33). However, radiation

from wall-to-droplets can account for a significant portion of the wall heat transfer [9], so that the actual

heat transfer rate to vapour P,&, is likely to be predicted realistically.

Further attempts to modify the code in order to get the correct liquid mass flow were not performed. As

a trend, it has been already established that, for decreasing values of the droplet diameter, more and more

liquid can be carried over, and increasing underprediction of the vapour temperature results.

These results are in agreement with those obtained by RELAPSIMOD2.5 and by the steady-state analyses.

5 Quasi-steady state analysis of Run 3053 by RELAP5/MOD2.5

Experiment 3053 was practically the same as Run 3051, the only difference being the higher pressure (3 bar

instead of 2). The analysis has been performed only by the RELAP5/MOD2.5, because of the difficulties

18
in the prediction of the liquid canyout by the MOD3, which render the interpretation of the results more

difficult.

The nodalization used for the analysis of run 3053 is shown in Fig. 34. Only, two axial wall temperature

distributions (Fig. 35) have been tested. The resulting average axial wall heat flux profiles during the time

period between 10 and 14 s are given in Fig. 36.

The transient time for the calculations was 14 s, since the node above the 'nominal' value of zqF quenches at

that time. The time-histories of the wall temperature at various elevations for the 'experimental' initial axial

wall temperature distribution are shown in Fig. 37. The time span over which the results are representative

of a steady quench front at 2.05 m (QF between 2.055 m, vol. 10, and 2.2 m, vol. 12) is between 8 and 14

S.

The results of the calculations are shown in Figs. 38-41. It is noted that the quality at the QF is correctly

obtained (Fig. 38) and the heat transfer to the vapour alone is much larger than the experimental tot&

power input (Fig. 39). The calculated liquid carry-out (Fig. 40) and the calculated total mass flow at

the exit (Fig. 41) decrease from larger values than in the experiment to lower values during the averaging

period. The calculated vapour carry-out (Fig. 42) is, however, larger than in the experiment, due to an

excessive evaporation in the tube. Therefore, one condition necessary for the fully meaningful comparison

(see §3) between calculated and experimental vapour temperature is not satisfied. We note, however, two

circumstances: 1) the experimental total mass flow was probably larger than that measured in the test, as a

certain error in the mass balance was found [14], and 2) the heat input per unit volume of the flowing vapour

is at least as large as in the experiment, since P,,,&, is substantially larger than the total experimental heat

flux (Fig. 39).

It is thus possible to state that the large underprediction of the vapour temperature at the tube exit (Fig. 43)

is again due to the overprediction of the interfacial heat transfer, and not to an overprediction of the vapour

mass flow.

6 Conclusions

The main goal of the present work was the assessment of the interfacial heat transfer model for the Dispersed

Flow Film Boiling region above a quench front implemented in RELAP5/MOD2.5 and RELAP5/MOD3. The

interest for this specific aspect of the DFFB model was raised by a separate work [9] which suggested severe

limitations of the I-D models in calculating the interfacial heat transfer rate under low mass flux conditions

for low qualities at the quench front.

To this aim, a low-flooding rate reflooding experiment at 2 bar has been analyzed by RELAPS/MOD2.5

and by RELAPS/MOD3. The transient analysis has shown for both codes a large disagreement between

the calculated and experimental quench front propagation rates, which can be attributed to an inadequate

modelling of the heat transfer phenomena in the vicinity of the QF and confirms other similar analyses, e.g.

[4]. In fact the good prediction of the carry-out suggests that the hydrodynamic behaviour is fairly well

simulated. The vapour temperature at the tube exit dropped to the saturation value much before than in

the experiment, but it could not be conclusively assessed whether a too high interfacial heat transfer was

responsible for such bad predictions. This was due to the fact that the total instantaneous power input to the

mixture in the post-CHF region at the time the quality at the QF becomes positive, was much larger than

in the experiment, as it depends on the axial transient cooling rate distribution and then on the actual wall

temperature distribution above the QF, which results from the preceeding transient.

In order to control the power input to the mixture, a quasi-steady state analysis has been performed. A short

transient was run, during which the quench front is allowed to advance a short distance around the position

of interest, imposing (approximately) the quality at the quench front. The power input to the mixture can be

controlled, to a certain extent, by imposing the initial wall temperature distribution above the quench front.

The RELAPS/MOD2.5 calculated correctly the total carry-out and overpredicted the total heat transfer to the

vapour. In spite of this, the vapour temperature was substantially underpredicted. It could be so concluded

that interfacial heat transfer was overpredicted.

19
RELAP5/MOD3 showed a large underprediction of the liquid carry-out, and this led to a good prediction of

the vapour temperature. However, when the calculation was repeated using two other versions of the code

where the closure laws affecting the droplet dynamics were modified in a way to allow a larger carry-out

(though still lower than the experimental one), the vapour temperature was again underpredicted.

The steady-state analysis of a second experiment at 3 bar was also performed by RELAP5/MOD2.5, and

results similar to those obtained for the test at 2 bar were obtained.

It can be concluded that, as expected from steady-state analysis, computer codes like RELAP5, which calculate the interfacial heat transfer rate between vapour and droplets assuming a homogeneous distribution of

the droplets over the cross section, overpredict the desuperheating effect of the liquid, and hence underpredict

the vapour temperature at large distances from the quench front. This underprediction of the vapour temperature, does not always lead to the underprediction of wall temperatures, as other deficiencies in the wall

heat transfer modelling and/or hydrodynamics often mask the difficulty of the codes to estimate the correct

interfacial heat transfer. Nevertheless, we hope that this work will motivate some efforts to develop some

empirical approach to calculate the reduction in the interfacial heat transfer rate under low mass fliu/low

quality conditions, since a detailed 3-D model which can calculate these effects (like that developed by one

of the authors [9]) cannot be easily incorporated in a system code.

Acknowledgements

The interest of Prof. Yadigaroglu for the practical implications of the main results of Andreani's dissertation

and his encouragement to carry-out the present investigation were appreciated. The help of M. Sencar in

setting up the nodalization and running the codes is gratefully acknowledged.

20
References

[I] S.N. Aksan, G.Th. Analytis and D. Lulbbesmeyer, 'Switzerland's Code Assessment in support of the

International Code Assessment and Application Program (ICAP)', Proceedings of the 161 Water Reactor

Safety Information Meeting (Oct. 1988), NUREG/CP-0097, Vol. 4, pp. 245-273, March 1989.

makebox[2Opt][J][2]M. Richner, G.Th. Analytis and S.N. Aksan, 'Assessment of RELAP5IMOD2, cycle

36.02, using NEPTUN Reflooding Experimental Data', NUREGAIA-0054, August 1992 (PSI report No. 104,

Oct. 1991)

[3] S.N. Aksan, F. Stierli and G.Th. Analytis, 'Boil-off Experiments with the EIR-NEPTUN Facility:

Analysis and Code Assessment Overview Report', NUREG/IA-0040, March 1992 (PSI-Report, No. EIR629, Sept. 1987).

[4] M. Sencar and S.N. Aksan. 'Evaluation of Reflooding Models in RELAP5/MOD2.5, RELAPSMOD3/v7j

Codes by using Lehigh University and PSI-NEPTUN Bmidle Reflooding Experimental Data', Paper presented

at the 2 - Code Assessment and Maintenance Program (CAMP) Meeting, Brussels, Belgium, May 1993.

[5] G.Th. Analytis, 'A Comparative Study of the Post-CHF Wall Heat Transfer Packages of the RELAP5 codes and Preliminary Assessment of Model Changes in RELAPS/MOD3/v7j' Paper presented at the

Ill Code Assessmntr and Maintenance Program (CAMP) Meeting, V'lligen, PSI, Switzerland, June 1992.

(6] G.Th. Analytis, 'A Summary of Model Changes and Options in RELAPS/MOD3' Paper presented

at the 211 Code Assessment and Maintenance Program (CAMP) Meeting, Brussels, Belgium, May 1993.

[71 G.Th. Analytis and P. Coddington, 'The Effect of Donor Cell Differencing of some Terms of the

Momentum Equations of TRAC-BFI on the Code Predictions', Proceedings of the 5

Y International Topical

Meeting on Reactor Thermal-Hydraulics (NURETH-S), Vol. III, pp. 745-750, Sept. 21-24, 1992, Salt Lake

City, Utah, USA.

[8] M. Andreani and G. Yadigaroglu, 'Effect of the Cross-Sectional Droplet Distribution in Dispersed Flow

Film Boiling at Low Mass Flux', Proceedings of the 5• International Topical Meeting on Reactor ThermalHydraulics (NURETH-5), Vol. III. pp. 823-831, Sept. 21-24, 1992, Salt Lake City, Utah, USA.

[9] M. Andreani, 'Studies of Dispersed Flow Film Boiling with 3-D Lagrangian Hydrodynamics and a

2-D Eulerian Vapour Field', Doctoral Dissertation Nr. 9794, Swiss Federal Institute of Technology, ErHZ,

Zurich (1992).

[10] D.G. Evans, S.W. Webb and L.C. Chen, 'Measurement of Axially Varying Non-Equilibrium in PostCritical-Heat-Flux Boiling in a Vertical Tube', NUREG/CR-3363, Lehigh-Univ., June 1983.

[11) Ransom, V.H. et al. 'RELAP5/MOD2 Code Manual', NUREGICR-4312, Rev. 1, March 1987.

21
[12] Carlson, K.E. et al. 'RELAP5/MOD3 Code Manual', Vol. 1, 2, 3, and 4, Draft, NUREGICR-5535, June

1990.

[13]Seban, R., Grief, R, Yadigaroglu, G., Elias, E., Yu, KL, Abdollahian, D. and Peake, W. 'TC-B Reflood Program: Experimental Data Report', EPRI NP-743, 1978.

[14] Seban, R.A. 'Reflooding of a Vertical Tube at 1,2 and 3 Atmospheres', Report EPRI NP-3191, 1983.

[15]Yu, K.P. 'An experimental investigation of the Reflooding of a bare tubular test section', Dissertation, Univ. of California at Berkeley, October 1978.

(16] Ishii, M. and Mishima, K. 'Two-Fluid Models and Hydrodynamic Constitutive Relations', NucL Eng.

and Design, 82 (1984), 107-126.

22
FIGURES

23

PRESS

CONTF

STES

TAN

mc

----T Ty COLLECTOR

t e.

V-5 

• 1•T

E :ymMETER

UMP TIN-L

DRAIN

Figure 1: System flow diagram of the UC.B Test facility (Seban et at, 1978).

.Pt

25
RUN #3051 (1•0-1-2 )

-&Ad%

IA

03

0

1000

300-

700-

600i' _V ý -- L

-500

-400

,300(.

i

I 4OU . - 1 1 H i ! i i i i ! i 1 1- -MR

I

* I0'f1 '__ _l __ I"o

200 , - , 100

10 .0- 0 0.61 m

1.Z21 -

1.83 a

2.44 a

'I-I-'.-'- X~ 3.35u

u 1 *

so too 200 250 300 350

I

50

[0

rime ( seconds )

400 40SO 50, S

0

SO

Figure .2: Experimental wall tempeanires at various elevations for Run 3051.

26
: (m~~

32.5

13.

-3.1

-3.3

-Lt.

-0.7

UC

~0.3

- A

-0

t2

tz

i1

t?

9

16

14

I

t

7

6o

.5

I.

11..0

34

31

21

27

25

34

23

21

Is

14

12

10

7

I

.S

4

3

2

1

a. position of.

thermocouple

(mPoizr texnp.)

I position of

the:=couPIe

I

I . hteat slatr

Qpipe

.11y nction

time dependent

volume

RELAP5/MOD3 RELAP5/MOD2.5.

Figure 3: Nodalizations used for the calculations of Run 3051 with RELAPS/MOD2.5 (right) and RELAP5•MOD3 (left).

27
oct.-

0x

0

'C0.-'

30.

416 . me. C. .0 M $Go 60. TIMEC (2)

11Mg Cs)ý

E

N

U.

S.

.I

. I

TIME (=)

.1. i I I i I *C•. I .. | .4 i 4

U. "

7$

2 " 1,

3:• AO I

I- - a

3M I

4 I M -. I =. I

*" .O ; ' 'I

3 , 2• O,4a. 3 iO. 14.S

I* 4 * iCC . 4emie TIME ()TIME (9)

Figure 4: Comparison of the wall temperatures at 0.66, 1.22, 1.83 and 2.44 m calculated by RELAPSAMOD2.5

and RELAPS/MOD3 with the experimental ones for Test 3051.

28
01

Oin

Sao,

3

I.

-S~

it |" I

S S S

I

I

I

I

I

I

x

i

C

01

S

N

a

2

a

a

a. E

a

3 .... •_ LL

4. -64. 2ke -* 34 |. - |

- - i s

Figure 5: Comparison of the wall temperzt'ures at 3.05 and 3.35 m calculated by R.ELAPS/MOD2.5 and

RELAP5/MOD3 with the experimental ones for Test 3051.

29
2

C'

N

H

a

aIa

a

a.

2

a

1~

a

a. a

,oa.7

.37

30L.

0.CL oo o. to. Sc

TIME()

I-3I00

x.

0

Ia.2

o. 0

TIME (s)

Figure 6: Calculated vapour temperature at 2.44 m and comparison of the calculated vapour temperature at

the tube exit with the measured value.

30
U

0

a

a

.3a.

d

a

a

U

0

a.

0

- GO-OSACO" - MASVOZ

Figure 7: Calculated vapour and liquid cany-out

S

E

-.-.. ac.uii/iaaos

TIMC (a)

Figure 8: Calculated mass flow rate at the tube exit.

31
L r

3

a

*1

U

a

3

3

U

a

a

S

3

U

.5

.3

a.

.5

U

U

a.

.2

a

'1

IA".

F

0. *00.

Time . 0. 00 riMe (a)

-2CuAS/WL -RLA.PSA'OGLS O•-03

.X

3

C6

1.

1~.

3

a

.5

U

.3

.3

U

.5

2

a.

a

a

LAD.

TAO.

1a .

V

'Cf

-E

'V

I

'C.0

TIME )

4M. I. IC.. 200. .46 . SQg.

TtMC (s)

Figure 9: Comparison of the calculated accumulated mass, collapsed liquid level, vapour and liquid entrainment with the experimental data.

32
a)

0€.

x

SO 1SO

C.

x

"0

0.00

0. 1oo. 200. 300. 400. 500. 600.

Figur omparson 10: o the calultedfluid explled ou of Iathe estsecio by)nertngdrcl

1.S

an dnstis).a REL.P.M•2. 03 and RLPtO3 b) ,<• 

x

0.000 Q:. 30 •Q. S0 0

TIME Cs)

Fi'ure 10: Comparison of the calculated fluid expelled out of the test section by integrating directly

the mass flow with that using control variables and the primitive variables (velocities, void fractions

and densities): a) RELAP5/MOD2-3 and b,) RELAPSIMOD3.

33
Steady-state calculation:

axial temperature

profiles in the post-CKF

region 'shift' with the

position of the QF.

Conditionx.

Vapour

temnperature

=en urement

T wall

droplet

residence

timue

tr .!

t my>>t

Xeq.QIP' X eq.QF.crp

at the end

- - -

4

|

$

O

|

0

e

Q

Axial advancement of the

I quench front during the

reraging' period t. ttr= Z. )-t(z • = Z,)

tin t the post-(W region

is considered representative

of that eisting above a

frozen'quench front at z=zv

water

filled

section

(not represented)

I

zo= z2W at beginning the of

the calculation

Figure 11: Prerequisites for the 'quasi steady-state' analysis of the post-CHF zone above a slowly advancing

quench front (under these conditions the interfacial heat transfer model can be assessed

34
ii

z (M)

-3.5

•3.3

- 3.1

post-CHF

W.7 region

2.3

2.1

Ten small

cells. 3c

1.7 each.

1.5

L3

LI

water

filled -u section (not

.7 repre- -9.7 sented)

0.3

0.1

0

125

R20

Z2

Z1

zo

10

L qz (tay)

This is anupper bound.

The actual elevation of the

quench front at the end

of the calculation wa different

in the three rins, and lower

than that shown.

o position of

thermocouple

(vapour temp.)

S position of

I

R270

ther•mocou'Sle (wall -temp.)

heat slab

pipe

11_0 junction

LJ • (•)ti.e depe~nden•t volumtae

Figure 12: Nodalization used for the steady-state analysis of Run 3051 by RELAP5/MOD2.5.

35
Sao

Sao1.75 2.00 2-=5 L34 275 S.O0 3.2 3-W

Elevation (m)

igure 13: Initial axial wall temperature distributions used for the steady-state analysis of test 3051 at the

time of the mid-height quench.

E

N.-

0

CL

-.km Cr-rmW

Iwof TV

-. hi 8 WElevation (M)

Figure 14: Average axial wall heat flux distributions during the time span 10-15 s for the three initial axial

wall temperavr distributions shown in Fig. 13.

36
- a - La-w-

- a-1.=5

a a & A m

4

I..

0

6

0.

6

0

3

TIME (s)

Figure 15: Calculated wall temperatures time-history at various elevations, when the experimental axial wall

temperature distribution is imposed.

---- --V.------)

btio.. 13).i -(Fig. ,,,)

a.

0-

E 0.(s

Fiue1. :Eufbu ult tteQ acltduigtetredfeetiia attmca•dsi

buos 0 g.1)

37
a

N

0

x

S

A

a

0

a

a

0

2

1~

0

0.

a

a

N

0

.1

a

0

A

0

0

U

U

a

S

0

0*

L - - - -- -- - - - I--- 

10.0

0.0 _________._____ C4 . 40 90 8.* 10.0 12.0 14.2

TMC (s)

- ,JStUo02.JoaT.• ) - -. ,O

- 8PSift002J(@ipT) - XADS/Wi002.OwMT-)

Figure 17: Vapour and liquid mass flow rates at the tub exit.

38
R " . I I I , - - I - I I I I S

0

6

6

.0

a

C

*1

*1

0

E

6

0

a

I.

6

0

x

6

6

.0.

a

0

a.

S

0

C

1~6

0

7.0-

4.0-

S.0-

4.0-

3.0-

1.0-

---------------

a-, - - - - - -

-IA

0

I

. 3.0 4.0 .0 . 10.0 t.0 14.0

0.75-

0.45-

0.30-

0.1S0.00

oa

, I 3 ~

. ---- - -- - - - - - -

I I I I 4. I I

0 2.0 d. 4.0 I 8(0 TMr. Cs) I0.0 1 1

Figure 18: Average values of the total mass flow rates and exit quality over the time span 10-15 sec. The

quantities to be averaged are calculated starting from 10 sec.

39
c

0

0

0c

0

CL

TIME (s)

- RELAP5/L4002.5(hiTw)

-RELAP5/h4002.5(*xpTw)

-a mm @ measured (vap..Ii.q.)

RELAP5/UJO2Z.S(tOwTw)

Figure 19: Comparison of the heat transfer rate to vapour with the total power input to the mixtmr measured

in the experiment.

40
SSao.=

*1.--------------------------------------------

330. * .0 ..1 1 0 A.0 4.0 &a I C.0 t 2.0 12 A.

TIME

- - - - To"

Figur 20: Comparison of the caculae vapour teprtr e at tihe tube exit with the expeimntal value a

tetime of the mid-beigh quenhx

0.

a

.0

I

d.

E

a

I0

C.

a

a

I

4

I I p j I

600. n - - - - - fl - fl - -

550.

500.

'50.

- - . mau.

- WCJPS/IIOOLICISWTW)

- Taut

- ECJPS/MOOZJOITW)

- SCAPSj'W00L5(..PTW)

T i 

. - - 1-," I1wr I I

44 I I Z.0 4.4 &aM . 10.s 12. Id

TIME (s)

Figure 21: Time-average value of the vapour temperature at the tube exit.

41
E

v O.

.. .

4.G.

0 U0

3.%

0-0

" 1 .,

a

0

-1.0

0.0 P 2.: 0 vli ..0 6.0 10.0 12.0 14.0

o 5.

N

N

4.

0"

3..

U.o

• 0.

0.0 2,.0 4.o 6.0 8L.0 10.0 iz• i..0

TIrME" Cs)

Figure 22: Phase velocity difference at z=2.4 m and z=3.0 m.

42
\

z (M) 12 t3Ms

3S5

-- 3.1

post-CHF

2.7 region

-2.5

.2.1

Z2

- LS

Ten small

cells. 5 cm

-L7 each.

1.5

i3

1.1.

- 0.9

AZ, (t a)

position of

thermocouple

(vapour temp.)

position of

thermocouple

(WLn temp.)

za

water :fi.1ed

- 7 section

(not

repre- 0.5 sented)

C.3

- 0.1

-"0

110.

I heat slab

pipe

junction

time dependent

volume

Figure 23: Nodalization used for the steady-state analysis of ran 3051 by RELAP5/MOD3.

43
mu I

1 1 1 1 1 1 1 -- F I I I I I . . . I - I I I I I I

I 4 a

700.-

0

1..

0

3..

0

0.E

0

a

3:

650.-

600.--

550.

500.

450.

400.

0.

* ~ ~92-@~~~*

I 0

-1 - .

-- 1.*~ - -.

-. x. - - - - - g

- - - - - - I I

m - - ~ ---- 1mm - -- - ii

x

:1* -1

0 2.5 5.0 7.5 10.0

TIME (s)

i ...............

12.5 15.0 17.5 20.0

.Tquenc1h (exp.)

Z =3.0m

-- *=z1.5mm

z2.4 m

Figure 24: Wall temperatures at various elevations calculated by the frozen version of RELAP5/MOD3.

44
.I.III,..,I.,.,i,,s.i,... I

0.90-

L&.

0

4-00

0

E

.03

0*

Vm

0.75-

0.60-

1**

.1

I i .; ;,; i~

I ? ; ;;; '. ,

I !i'-; i !iII i? ; ;

! II ! I!; "; I

0.3-

0.15:

0.0

g

J.

I I

I .2. ~

I

0. 0 2.5 5.0 7.5 10.0 20.0

TIME (s)

-. ,,-.- RELAPS/Uo03(UPX) -

-.....- RELAPS/MOO3(froz.) -.

12.5 15.0 17.5

i £ moazurod

*-RELAPS/M003(bb)

Figure 25: Equilibrium quality at the QF calculated by the frozen RELAP5/MOD3 (- .. -. -), by a version

(bb) including the PSI updates (- - .- : Analytis, 1992) and an 'ad hoc' modification (version UPX)

of the RELAP5/MOD3(bb) using a reduced value for the minimum droplet diameter and the correct drag

coefficient for droplets ( e -.

45
f . I . . . . . . I . . . - I. . - i I . . | . I . . . ..-

-- - - -- --- .-------------

&..

550. 4!

,/

S

E

00.

I I A

0.0 5.0 7.5 30.0 12.3 15.0 17.5 20.0

a o I Ja a

Sm.

a

SS0.

. .0 2. S . .3 10.0 .2 13.0 17.3 20.*

-*5o 00./ • - - • m u au STI M

Fig=r 26: Vapou temperatur at the tube exit c-a]mlt by three versions of RELAP5/MVOD3 (see legend

in Fig. 25).

46
.0

0

.C

*2

*1

C

0

0.

C

* '~

4.

3.

2.

II

i.0

a a

I

I

I

I

I.

I.

i

I a

i ;

* a

* a

* a

ma *

aS *

I a

~q

q

p - - *

I -

tmeoiufnd

RCLAPI•I003(bb)

*CLAPS/MOZ3(UPX)

XCLAP5/WCO0fre.)

I !

- 1 i ~l I

a a a a

0.0 2•, S.O 7Y. 10.0

TIME Cs)

I I I 12.5 19.0 17.S 20.0

Figr Z7: Vapour carry-out calculated by three versions of RELAPSMOD3 (see legend in Fig. 25).

- -. inos~i

- . .. RZL..P5/SO03(b~)

- . - ac2.&P5/Ia0o3~Iwr)

- acLAP5/Ia0o3(fu'sL~

.0

x

0

3:

U

Vs

0

C.

4.O

3.,

2.- a '

t.;

.3

I.D.. • : -., ;1 ,"

;g . , '.

II. . as• ".

91-

1.05 11.0I .- . .. 1

0.0 . .0 7.S 1. 1 ¶..S 1S.0 17., TIME (Z) 32.0

Figure 28: Liquid carry-out calculated by dhre versions of RELAPS/MOD3 (see legend in Fig. 25).

47
C

0

0

2.4m

s- JAL" .4

urn2M

TIME (s)

Figure 29: Void fractions (shown on two different scales) at various elevations calculated by the frozen

version of RELAP5/MOD3.

48
.10 -

C

0

C.

C

C

C. I.

C

|

0

0

,.-- jLT---4

V.4'

.50-

0.0

- I-.-

2.3 5.0 7. 1.0 1ot

TIME (s)

I

S 17I ¶0 *17-S -- I 20.0

a I 2L4 M

aU a l7$ ma

33.U

a 2Ai

Figure 30: Void fractions (shown on two different scales) at various elevations calculated by RELAP5/MOD3+PSI updates (version bb).

49
I I &

C0

0

*0

100.0..

50.0..

S8.0~!

'Si

55.0-

a "-* • . * ".* • .*1 •

•S !; "" " . l*8 * * I -

8 !. **f

8. I

0. 2.5 5.0 7.5 10.0 12.3 15.0 17.5 2C.0

a

1"

0

1.10..

I~0..i- 88 s 

*. ;. S

1.7• ..-

I

O.SO.

0.W I

. 5.0 7-3 10.0 12.

TIME (s)

. I . . .

15q.0 8 17..5 I 20.0I

-. a u 2,.,6 -2

a 1.S73 1m

-- 2Am

Figure 31: Void fractions (shown on two different scales) at various elevations calculated by RELAP5IMOD3+PSI updates+modified C' and d4i,, = 1.25 mm (version UPX).

50
h ,

";; '. it

1 0,II I~ I* •. a

C " as S I i

(see le.g, . ,2 1500.I !

. ... 'I 'l ! •

L.

-- I I l..

. ... ., . l.. •.'- . .,- .~~. .. ,-,s,."'

LI ".I* 1Cll 0 12. iS0 75*0.

U0.*P5 i. I.03IZ *u -. m 

(seef~* , Pi. 25) legend"n 

AUDO

-.-.- 30.hPS/U003(lP~ - -. m.suw.d(..p.*q.)

CA

c

0

I.

a

L

S

I i I I I

4.

I'

ii

1I 0.- 1 i

t !! ; ;

15I! ;

TO00.- .. i-..---

-- -- ";t 1 " '.-

•..•, ,.;. ..•

0. • ...

0t.0 2.5 LO 7.s sosi :2.5. ls.c

TIME[ Cs)

F.

C

17.= 20.

Figure 33: Heat transfer rate to the vapour alone for the three versions of RELAPS/MOD3 (legend in Fig.

25).

51
z(m) 125

3.68

- 3.5

- 3.3

-' 3.1

1l270

region

2.?

2.5

7.3

Z.1 Eleven smad

cells. 3 cm

1u- each.

z2

zi

Z-F=20

1.7

1.5

1.3

water

filled

1.1 region

(not

repreo.9 sented)

0.7

0.5

zo

z QF (t av)

a position of

thermocouple

(vapour temp.)

I position of

thermocouple

C(wl temp.) 110

I

110•

heat slab

pipe

junction

time dependent

volume

0.1

0

Figure 34: Nodalizadon used for the steady-state analysis of run 3053 by RELAP5/MOD2.5.

52
7 00.- C

E

goo

x

0

1.75. 2.00 2.2= 2.50 2.7s 3.00 3.= 3.50

Elevation (m)

- --. Tq,.nch(.xp.)

Figure 35: Initial axial wall temperatzre distributions used for the steady-state analysis of test 3053 at

the time of the mid-height quench.

Ia

'I.

ex.p. Tw-

-- Sw TvI

ElevatIon (m)

Figure 36: Avenge axial wall heat flux distributions daring the time span 10-14 sec. for the two initial axial

wall temperamu distributions shown in Fig. 35.

53
S

I..

3

0

S

0.

E

S

3

TtME (s)

- .Tquencli (omp.)

- - 3.0 fm

- z m L2 m

z - 2.5S en

-z - &A4 mo

-z u2.4 ro

z 2.=3 mn

Figure 37: Wall tepera-nres at various clvartions for test 3053 calculated by RELAPS/MOD2.5.

- RrLAPS/1i002.SVTw.

- - - - - - - - - -. - - - - -

0

U.

TIM

Fo

U..

0.* S0.0 2.0 4. 6.0 8.0 10.0 12.0 1,4.0

Figure 38: Equilibriuma quality at ihe QF for the two different initial axial wall temnperature profiles.

54
0

0

CL

1c

0

CL

TIME (s)

-, - , mecsured (vcp:+Iiq.)

RELAP5/MO02.5(expTw)-.-- RELA.P5/M002.5(QowTw)

Figure 39: Heat transfer rate from wall to vapour in the post-dryout region for the two initial wall temperature

distributions.

55
6.0.

XS.O

.,0

3.-

3

-o -- - -

'a

- 0.0 -

"1 12 tA 0

0.0 2..0 4.0 4O S.C 10O.0 2,.0 A

TIMT Cs)

F- gue4/:otamuspT----- x.eLPS/Initial2.,wTw) Rae 

igure 40:. Liquid cazi-ouz for the two inta a~xial wall tempezaire dimibutons.

" 10 i " " I " " I " " ' " " " I | " " i " " i "

7.O

* ..

0 -

€;_ :o- - - RV.,PS/U02.5(u,-T,,)

a

1.-

Thn4E (s)

Figure 41: Total mass flow at the tube exit for the two initial axial wall temperatur distrbutions.

56
0

20

.0

8. -- - - - - - - - - - - -- - - - - - -

E

1--- -- .----- O

0. .

I0.

~P

.... TTME

"-,,-RP ()S//O( -)LS(MT,). 

Figur 42 Vao- u fo ax -my -h-w wa- -nta - t- -d--st-rib- p rtr ution-- s.- -

TIME (s)

Figure 43: Vapour temperaur at the tube exit calculated

temperature dtibutions.

by IRLAP/MOD2S5 for the two intial wall

57
APPENDIX 1

Input decks for transient calculations

58
Appendix IA

RELAP5/Mod 2.5 input deck

59
= UcB351

UNIV. OF CALIFORNIA BERKELEY (UCB) TEST

REFLOOD TEST IN SINGLE TUBE GEOMETRY

* UCB-3051: P=2 BAR, G=25.1 KG/M**2/S

* input : power and heat losses

* RELAP5/MOD3 calculation *SmffU.U*US.US*SSU.UmmU@mmSm...U.U.*.Wm mmmme. .f.....m..mmm..m..UPCU*Ubm

00000100 new transnt

00000101 run

* cpu-lim

00000105 1.0 2.0 3580.

* TIME STEP CONTROL CARDS * REQUIRED

* END TIME MIN DT MAX DT OPTN MNR MJR RST

00000201 10. 1.0-9 0.001 3 100 4000 4000

00000202 1000. 1.0-19 0.01 3 100 4000 4000

* 1MINOR EDIT REQUESTS

00000301 cputime 0

00000302 voidq 120010000

00000303 voidg 120020000

00000304 teodf 120010000

00000305 tempg 120010000

00000306 tempf 120020000

00000307 te=Vf 120020000

00000308 p 120010000

00000309 p 120020000

00000310 voidg 120120000

00000311 voidg 120130000

00000312 voidg 120140000

00000313 voidg 120150000

00000314 voidg 120170000

00000315 voidg 120190000

00000319 tempg 120190000

00000340 mflowj 125000000

00000342 cntrlvar 103

00000343 cntrlvar 112

00000344 cntrlvar 113

00000345 cntrlvar 114

00000346 cntrlvar 115

00000349 cntrlvar 122

00000350 cntrlvar 118

00000351 cntrlvar 119

00000363 cntrlvar 140

00000364 cntrlvar 141

0000036S cn*rlvar 142

00000366 cntrlvar 143

00000370 cntrlvar 150

00000371 cntrlvar 151

00000372 cntrlvar 152

00000373 cntrlvar 161

00000374 cntrlvar 162

00000376 cntrlvar 169

60
00000380

00000381

00000382

00000383

00000386

00000388

00000390

00000391

00000392

00000393

00000394

00000395

t

floreg

floreg

floreg

floreg

floreg

floreg

httemrp

httemp

httemp

httemp

httemp

httemp

120040000

120070000

120100000

120130000

120160000

120180000

120000401

120000701

120001001

120001301

120001601

120001801

*00000399 ztrvt 1200000

*

*

*

TRIP COM0NENTS

*

500 time 0

502 time 0

qt null

gt null

0. 600. 1 * end

0. 0. 1 * start injection

and reflood

600

*

500

*

*

*

HYDRODYNAMIC COMPONENTS

* [105] feedwater tank - inlet thermodynamic conditions

1050000 "feedtan tmdpvol

1050101 0.00015935 0. 1000. 0. 0. 0. 0. 0. 0

1050200 103 * i.c.: p,T; subcooled water

1050201 0. 2.5e5 311.15

* ---- 1 -------- 1 -------- 1 -------- ------- -------- ----

* (110] inlet junction - flow boundary conditions

*----...1--------1-........-1------1---........1--------1--

1100000 "feedjun" tmdpjun

1100101 105000000 120000000 0.

1100200 1 502

1100201 0. 0.004007 0.000 0. *mcell=G*Acell

* ---- -- ---------------- 1 -------- ---- ---1---- ---- 1----

* (120] tube

1200000 "testsec" pipe

1200001 19

---- 1-

---- 1-

----1-

---- 1-

---- 1-

---- 1-

1200101

1200201

1200301

1200302

1200303

1200401

1200501

1200601

1200701

1200702

1200703

1200801

1200901

*1201002

1201002

1201101

1201201

0.0001593!

0.

0.1

0.2

0.16

0."

0.

90.

0.1

0.2

0.16

1.-4

0.

100

0

00000

3

5 19*A'

18* A 1 * ci 18 *

19

19 * Ci

19 *hi

19 *v 1I* ci 18 *

19 *

0.01425

0.

11

19

18

2.00+5

Wcell

junct

eli length dz

ell volume=dz*Avcell

or. angle

ert. angle

ell dz

19

18

* friction data

* loss coefficients

* bundle correlation on

* tube correlation on

812. 0. 0. 0. 19 * initial

* conditions

61
1201300 1

1201301 0. 0.0 0. 18 * mass flows

• [125] outlet junction

•- -- 1.---------1--- ----. ....--- 1-----1--------1--------1-

1250000 "outljuno sngljun

1250101 120010000 130000000 0. 0.0 0.0 00000

1250201 1 0. 0.00 0.0

* -- -- 1 -------- 1 -------- 1 -------- 1 -------- 1 -------- 1 ------ I-

[ (1301 outlet tank

•.... 1--------1-........-1--------1-----1--------1--------1-

1300000 "outltan" tmdpvol

1300101 .01 0. 1000. 0. 0. 0. 0. 0. 0

1300200 003

1300201 0. 2.0+5 812.

$*$**$*$*$*$*$*$*$*$*$*$*$*$*$*****s*********$*$*$*$*$***$**$****

* HEAT STRUCTURES

•* (1200] test section(tube)

11200000 19 10 2 0 0.007125 502 0 32

11200100 0 1

11200101 9 0.00795 * outer tube radius

11200201 1 9

11200301 1. 9

11200400 0

11200401 812. 10 * temperatures of the nodes

11200501 120010000 0 1 1 0.1 1 * inner

11200502 120020000 10000 1 1 0.2 18 * surface

11200503 120190000 0 1 1 0.16 19 * 11200601 -120 0 4110 1 0.1 1 * outer

11200602 -120 0 4110 1 0.2 18 * surface

11200603 -120 0 4110 1 0.16 19 * 11200701 150 0.0273224 0. 0. 01 * power distribution

11200702 150 0.0546448 0. 0. 18 * 11200703 150 0.0437158 0. 0. 19 * * 11200801 0 0. 0. 0. 19 *2

11200801 0. 11. 11. 0. 0. 0. 0. 1. 19 * M3 *--- -1--------1-........-1--------1-----1--------1-. ....-i--

• CONTROL VARIABLES

* cv 103: collapsed liquid level in the test section (m) *----...1--------.1-----1-------1-----1-------1--------1-

20510300 collev sum 1. 0. 1

20510301 0.

20510302 0.1 voidf 120010000

20510303 0.2 voidf 120020000

20510304 0.2 voidf 120030000

20510305 0.2 voidf 120040000

20510306 0.2 voidf 120050000

20510307 0.2 voidf 120060000

20510308 0.2 voidf 120070000

20510309 0.2 voidf 120080000

20510310 0.2 voidf 120090000

20510311 0.2 voidf 120100000

20510312 0.2 voidf 120110000

20510313 0.2 voidf 120120000

20510314 0.2 voidf 120130000

62
20510315

20510316

20510317

20510318

20510319

20510320

0.2

0.2

0.2

0.2

0.2

0.16

voidf

voidf

voidf

voidf

voidf

voidf

120140000

120150000

120160000

120170000

120180000

120190000

*-- -12: -vapou 1-- ----- 1f ..... tt 1-- s1-- -- 1-- -- 1-

* cv 112: -vapour mass flow from the test section (kg/s)

*- - - -... 1-- - - -1-i-........1-........-1-------- 1-.... ....-1-........-

20511200

20511201

20511202

20511203

entgf 1

rhogj

velgj

voidgi

mult

125000000

125000000

125000000

0.00015935 0. 0

*-- - 13: liquid-m ----.f1- 1-- -- - --- 1- -- -1- - . . 1-

* cv 113: liquid mass flow from the test section (kg/s)

*----1-------1 --i ------1--------1-........-1-----1--------1-

20511300

20511301

20511302

20511303

entff 1

rhofj

velfj

voidfj

mult

125000000

125000000

125000000

0.00015935 0. 0

*-- - 14: total- 1-- f1--- ---- out o1f t1-- .. - -1- --- 1

* cv 114: total mass flow out of the test section (kg/s)

*- - --.. .------- 1- --.... .... 1-.... ....-1-.... ....-1-.... ....-1-.... ....- 1-

20511400

20511401

20511402

20511403

entflow sum

0.

1. cntl

1. cntz

I. 1. 0

lvar 112

ivar 113

*-- - 15: qualit at 1-- ------ 1 e.. -------- 1--- ....1---- .1--

* cv 115: quality at the tube exit

*----...1-----1--------1-........-1-........-1--------1--------1

20511500

20511501

*

exitqual div

cntrlvar 114

1. 1.

cntrlvar 112

0

* ------------- 1 ---- ---- 1 -------- 1 -------- 1 -------- 1-

* cv 118: integral of mass flow at tube exit (kg) (R5 var.)

*----.1------1-------.- .---- 1--------1--------1....

---- 1-

---- 1-

20511800 intwgext

20511801 mflowj *

integral 1.

125000000

0. 0

... 1 19: inte --of 1---- --flo1w 1at tube1 .... (1-- . . . - * cv 119: integral of maSS flow at tube exit (kg) (cntrlvar)

*----.1-----1---------1-........-1--------1-........-1-........-1-

20511900

20511901

intwgext integral 1.

cntrlvar 114

0. 0

.... 1-----....-1-........-1-. * ---- --------- 1 1 -------- 1 -------- 1 ---- * cv 122: mass of water accumulated (kg)

*----1---------1--------1--------1-....

*

---- 1 -------- 1---- ---- 1-

20512200

20512201

20512202

20512203

20512204

20512205

20512206

20512207

20512208

20512209

20512210

20512211

accwater sum

0.

0.1 rho

0.2 rho

0.2 rho

0.2 rho

0.2 rho

0.2 rho

0.2 rho

0.2 rho

0.2 rho

0.2 rho

0.00015935

120010000

120020000

120030000

120040000

120050000

120060000

120070000

120080000

120090000

120100000

0. 0

63
20512212 0.2 rho 120110000

20512213 0.2 rho 120120000

20512214 0.2 rho 120130000

20512215 0.2 rho 120140000

20512216 0.2 rho 120150000

20512217 0.2 rho 120160000

20512218 0.2 rho 120170000

20512219 0.2 rho 120180000

20512220 0.16 rho 120190000

* cv 140: integral of vapour mass flow at tube exit (kg)

*----...1--------1--------1--------1--------1--------1-------.--1

20514000 intwgext integral 1.

20514001 cntrlvar 112

0. 0

* ---- 1 -------- 1---- ---- 1 -------- 1 -------- 1 ---------1--- ---- 1-

* cv 141: integral of liquid mass flow at tube exit (kg)

*----...1--------1--------1--------1-........-1-........-1-------.--1

20514100

20514101

intwfext

cntrlvar

integral 1.

113

0. 0

* ---- --------- 1 1 ----.-------- ---- 1 -------- 1 -------- 1 -------- 1-

* cv 142: sum of total mass accumulated and-expelled (kg) *----1-------1 ----------------- 1 ---------1---------1---------.1-

*

20514200

20514201

20514202

20514203

20514204

masstot

0.

1.

1.

1.

sum 1. 0. 0

cntrlvar 140

cntrlvar 141

cntrlvar 122

* ---- 1 -------- 1 -------- 1 -------- 1 -------. ---- ---- 1----

cv 143: integral of liquid mass flow at tube inlet (kg)

*----...1-........-1--------1--------1-........-1-........-1....

--1-

---- 1-l

20514300 intwfin integral 1.

20514301 mflowj 110000000

0. 0

* ------- 1 -- -- -------- 1- I-------- 1 -------- ----

* cv 150: powerý input to wet portion of test section

*----...1--------1------1--------1-........-1--

*

-- 1-----1-

....1-----1-

20515000 qinwet

20515001 0.

1. 0.

20515002

20515003

20515004

20515005

20515006

20515007

20515008

20515009

20515010

20515011

1. 1.

i.

1.

1.

1.

1.

1.

1.

1.

q

q

q

q

q

q

q

q

q

q

120010000

120020000

120030000

120040000

120050000

120060000

120070000

120080000

120090000

120100000

*t------------1 ---- e v 151: power

---- 1 -------- 1 -------- 1 ---- ---- ------- -------- 1-

input to dry portion of test section

---- 1 -------- 1 -------- 1 -------- --- ------- -----

20515100

20515101

20515102

20515103

20515104

20515105

20515106

qindry

0.1.

1.

i.

1.

1.

sum 1. 0. 2 2000.

q

q

q

q

q

120110000

120120000

120130000

120140000

120150000

64
20515107

20515108

20515109

20515110

1. 1.

1.

1.

q

q

q

120160000

120170000

120180000

120190000

* cv 152: total ---- 1 -------- 1 ---------------- 1 ----

power input to fluid in test section

---- 1--------1 ---- ---- 1 -------- 1 ----

---- 2.---------1-

---- 1---------1-

20515200

20515201

20515202

20515203

qindry

0. 1.

i.

sum I. 0. 1

cntrlvar 150

cntrlvar 151

* ---.---- 1 -------- 1 -------- 1 ----.---- 1----

* cv 161: power input above 2.4 m

*- - - -... 1-.... ....-1-.... ....- 1-.... ....- 1-..--

---- 1---------1---------1-

---- 1---------1---------1-

20516100

20516101

20516102

20516103

20516104

20516105

20516106

20516107

qindry24 sum

0.

1. q

1. 1.

1. q

i.I. qZ

1. 0. 1 2 2000.

120140000

120150000

120160000

120170000

120180000

120190000

* ---- 1---- ------------ 1 -------- 1 -------- ---- ---- 1 ---- ---- 1-

* cv 162: power input to vapour above 2.4 m

*----...1-........-1-----1--------1-........-1--------1-------.--1

20516200 qvap24

20516201 0.

20516202 1.

20516203 1.

20516204 1.

20516205 1.

20516206 1.

20516207 1.

1. 0. 1 2 2000.

qwg

qwg

qwg

qwg

qwg

qwg

120140000

120150000

120160000

120170000

120180000

120190000

* ---- 1 -------- 1 -------- 1 -------- 1 -------- 1 ---- ---- 1 --- ---- 1-

* cv 169: heat transfer to vapour only in the dry region ----... 1--------1-........-1--------1--------1-........-1-. ....----1

20516900

20516901

20516902

20516903

20516904

20516905

20516906

20516907

20516908

20516909

20516910

qdryvap sum

0. 1. qlwg

1. qlwg

1. qw•

1. qwg

1. qlwg

1. qwg

1. qwg

1. qwg

1. 0. 1

120110000

120120000

120130000

120140000

120150000

120160000

120170000

120180000

120190000

*s**s*s*s*s*s***s****s*s**$**$*$*s*$*s*s*s*$*s*s*s*s***$**s**s***

* GENERAL TABLES *s*$*$*S*$*$*S*$*$*S*S*$*$*S*S*S*$*$*S*S*$*S*S-*$*$**S*s*$*S$*S*$

* tab 110: heat losses

20211000 htc-temp

20211001 300. 12.13582

20211002 350. 13.81509

20211003 400. 15.72673

20211004 425. 16.77957

20211005 450. 17.90288

65
20211006 475. 19.10140

20211007 500. 20.38016

20211008 520. 21.46453

20211009 540. 22.60660

20211010 560. 23.80943

20211011 580. 25.07626

20211012 600. 26.41051

20211013 620. 27.81574

20211014 630. 28.54614

20211015 640. 29.29573

20211016 650. 30.06500

20211017 660. 30.85448

20211018 670. 31.66468

20211019 680. 32.49616

20211020 690. 33.34947

20211021 700. 34.22519

20211022 710. 35.12390

20211023 720. 36.04622

20211024 730. 36.99275

20211025 740. 37.96413

20211026 750. 38.96103

20211027 760. 39.98410

20211028 770. 41.03404

20211029 780. 42.11154

20211030 790. 43.21734

20211031 800. 44.35218

20211032 810. 45.51682

20211033 820. 46.71204

20211034 830. 47.93864

20211035 840. 49.19745

20211036 850. 50.48932

20211037 860. 51.81511

20211038 870. 53.17572

20211039 880. 54.57205

20211040 890. 56.00505

20211041 900. 57.47568

20211042 910. 58.98492

20211043 920. 60.53380

20211044 930. 62.12335

20211045 940. 63 .75463

20211046 950. 65.42876

20211047 960. 67.14684

20211048 970. 68.91004

20211049 980. 70.71953

20211050 990. 72.57655

20211051 1000. 74.48233

20211052 1010. 76.43815

20211053 1020. 78.44532

20211054 1030. 80.50520

20211055 1040. 82.61918

20211056 1050. 84.78866

20211057 1060. 87.01512

20211058 1070. 89.30003

20211059 1080. 91.64495

20211060 1090. 94.05144

20211061 1100. 96.52112

*tab 120 ambient temperature

20212000 temp

20212001 0. 294.26

20212002 1000. 294.26

66
* ....---- 1 ---- - -1 ---I ---- 1 -... 1 ------ ..

* tab 150 test section power

20215000 power

20215001 0. 4326.

20215002 1000. 4326.

$$$$$$$$$$ HEAT STRUCTURE THERMAL PROPERTY DATA

* (s-steel and inconnel 600 from Bethsy, copper from TM-IN-816)

*----...1-........-1-........-1---------1---------1-........-1--

20100100 tbl/fctn 1 1 * INCONEL-600

* INCONEL-600 - thermal conductivity (W/m-K) [tpO02]

20100101 0. 14. 294.0 14.8 * Ist artiff.

20100102 373.0 15.7

20100103 473.0 17.3

20100104 573.0 18.9

20100105 673.0 20.6

20100106 773.0 22.2

20100107 873.0 23.8

20100108 973.0 25.7

20100109 1073.0 27.4

20100110 1173.0 29.3

20100111 1273.0 31.0

20100112 1373.0 32.7

20100113 1473.0 34.4

20100114 1573.0 36.1

* INCONEL-600 - volumetric heat capacity (J/m3-K)

20100151 0. 3.7+6 294.0 3.740+6 * 1st artiff.

20100152 373.0 3.900+6

20100153 473.0 4.060+6

20100154 573.0 4.180+6

20100155 673.0 4.330+6

20100156 773.0 4.510+6

20100157 873.0 4.650+6

20100158 973.0 4.830+6

20100159 1073.0 5.020+6

20100160 1173.0 5.200+6

20100161 1273.0 5.340+6

20100162 1373.0 5.630+6

20100163 1477.0 5.810+6 * *** END ***

67
Appendix lB

RELAP5/Mod 3 input deck

68
= UCB351

* UNIV. OF CALIFORNIA BERKELEY (UCB) TEST

* REFLOOD TEST IN SINGLE TUBE GEOMETRY

* UCB-3051: P=2 BAR, G=25.1 KG/M**2/S

* input : power and heat losses

* RELAP5SMOD2.5 calculation

* for RELAP5/MOD3 only the card nr. 11200801 has to be changed *..neumem*'ummumumeummmamummmm....m..m..m..mm.....m..fhmmmumemmm.*swmm U

00000100 new taunsnt

00000101 run

* cpu-lim

00000105 1.0 2.0 3560.

* TIME STEP CONTROL CARDS * REQUIRED

* END TIME MIN DT MAX DT OPTN MNR MJR RST

00000201 10. 1.0-7 0.001 3 100 4000 4000

00000202 1000. 1.0-19 0.05 3 40 2000 2000

* MINOR EDIT REQUESTS

00000301 cputime 0

00000302 voidg 120070000

00000303 voidg 120130000

00000304 voidg 120190000

00000305 voidg 120250000

00000306 voidg 120310000

00000307 voidg 120340000

00000309 voidg 120160000

00000310 rhog 120160000

00000311 voidf 120160000

00000312 rhof 120160000

00000313 rho 120160000

00000319 hemog 120360000

00000340 mflowj 125000000

00000342 cntrlvar 103

00000343 cnxtrlvar 112

00000344 =ntrlvar 113

00000345 =ntrlvar 114

00000346 =ntrlvar 115

00000347 cntrlvar 120 00000348 121

00000349 cntrlvar 122

00000350 ntrlvar 118

00000351 c=trlvar 119

00000363 cntrlvar 140

00000364 cntrlvar 141

00000365 cntrlvar 142

00000366 cntrlvar 143

00000380 floreg 120070000

00000381 florel 120130000

00000382 floreg 120190000

00000383 florei 120250000

69
00000384 floreg 120310000

00000385 floreg 120340000

00000390 h*temp 120000701

00000391 httemp 120001301

00000392 httemp 120001901

00000393 httemp 120002501

00000394 httemp 120003101

00000395 httemp 120003401

*00000399 ztrwt 1200000

* TRIP COMPONENTS *

500 time 0 gt null 0. 600. 1 * end

502 time 0 gt null 0. 0. 1 * start injection * and reflood

600 500

* HYDRODYNAMIC COMPONENTS

* [1051 feedwater tank - inlet thermodynamic conditions *---------------1 ---------1---------1-......-1---1------1-........-1-

1050000 "feedtan tmdpvol

1050101 0.00015935 0. 1000. 0. 0. 0. 0. 0. 0

1050200 103 * i.e.: P,T; subcooled water

1050201 0. 2.5e5 311.15

* --------- 1 --------1---------1---------1-------- 1 -------- 1 --------- 1-

* (1101 inlet junction - flow boundary conditions

1100000 *feedjun" tmdpjun

1100101 105000000 120000000 0.

1100200 1 502

1100201 0. 0.004007 0.000 0. *mcellýG*Acell *------------------1 --i------1-........-1-........-1---------1---........1-

* (1201 tube

*-------------1---------------1-........-1--i -....- 1-........-1-........-1-

1200000 "testsec* pipe

1200001 36

1200101 0.00015935 36 * Avcell

1200201 0. 35 * Ajunct

1200301 0.1 1 * cell length dz

1200302 0.1 35 * 1200303 0.16 36

1200401 0. 36 * cell volume=dz*Avcell

1200501 0. 36 * hor. angle

1200601 90. 36 * vert. angle

1200701 0.1 1 * cell dz

1200702 0.1 35 *

1200703 0.16 36 * 1200801 l.-4 0.01425 36 * friction data

1200901 0. 0. 35 * loss coefficients

*1201002 100 11 * bundle correlation on

1201002 0 36 * tube correlation on

1201101 00000 35

1201201 3 2.00+5 812. 0. 0. 0. 36 * initial

* * conditions

70
1201300 1

1201301 0. 0.0 0. 35 * mass flows

* [125] outlet junction

1250000 Ooutljun sngljun

1250101 120010000 130000000 0. 0.0 0.0 00000

1250201 1 0. 0.00 0.0

* ---- 1 -------- 1 ---- ------------ 1 -------- 1 -------- 1 -------- i-

[ (130] outlet tank

1300000 Ooutltan* tmdpvol

1300101 .01 0. 1000. 0. 0. 0. 0. 0. 0

1300200 003

1300201 0. 2.0+5 812.

• HEAT STRUCTURES

• [1200] test section(tube)

11200000 36 10 2 0 0.007125 502 0 16

11200100 0 1

11200101 9 0.00795 * outer tube radius

11200201 1 9

11200301 1. 9

11200400 0

11200401

11200501

11200502

11200503

11200601

11200602

11200603

11200701

11200702

11200703

11200801

*11200801

812.

120010000

120020000

120190000

-120

-120

-120

150

150

150

0

0.

10 * temperatures of

0 1

10000 1

0 1

0 4110

0 4110

0 4110

0.0273224 0.

0.0273224 0.

0.0437158 0.

0. 0.

11. 11.

the nodes

1 0.1 1 * inner

1 0.1 35 * surface

1 0.16 36 *

1 0.1 1 * outer

1 0.1 35 * surface

1 0.16 36 *

0. 01 * power distribution

0. 35*

0. 36 *

36 M2

0. 0. 0. 0. 1. 18 * M3

* CONTRODL VARIABLES

* cv 101: collapsed liquid level in the test section (lower part)

*----...1--------1-........-1-........-1-----1--------1-........-1-

20510100

20510101

20510102

20510103

20510104

20510105

20510106

20510107

20510108

20510109 20510110

20510111

20510112

20510113

20510114

collevl

0.

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

sum

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

1. 0. 1

120010000

120020000

120030000

120040000

120050000

120060000

120070000

120080000

120090000

120100000

120110000

120120000

120130000

71
20510115

20510116

20510117

20510118

20510119

0.1

0.1

0.1

0.1

0.1

voidf

voidf

voidf

voidf

voidf

120140000

120150000

120160000

120170000

120180000

*-- 10 1- ------- 1--------- le 1----d ... 1 te . t-.. --- se n (--.1--- 

* cv 102: collapsed liqluid level in the test section (upper part)

*----...1--------1-----1--i --- 1---------1-........-1-------. --

20510200

20510201

20510202

20510203

20510204

20510205

20510206

20510207

20510208

20510209

20510210

20510211

20510212

20510213

20510214

20510215

20510216

20510217

20510218

20510219

collevu sum

0.

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.1 voidf

0.16 voidf

1. 0. 1

120190000

120200000

120210000

120220000

120230000

120240000

120250000

120260000

120270000

120280000

120290000

120300000

120310000

120320000

120330000

120340000

120350000

120360000

* cv 103: collapsed liquid level in the test section (m)

*----...1-----1--------1-........-1--------1-----1-----1

20510300 collev

20510301 0.

20510302 1.

20510303 1.

$sum 1. 0. 0

cn=trlvar 101

cntrlvar 102

* ---- 1--- - -- -- 1 -------- -------1 -------- 1----

* cv 112: vapour mass flow from the test section (kg/s)

*----...1-----1--------1-........-1-........-1 -

--1-----1-

--1---------1

20511200

20511201

20511202

20511203

entgf 1

rhogi

velgj

voidgi

mult

125000000

125000000

125000000

0.00015935 0. 1

* ---- -- .------- 1---- 1 -------- 1 -------- 1 ---- * cv 113: liquid mass flow from the test section (kg/s)

*----...1-----1--------1-........-1--------1--

1 20511300

20511301

20511302

20511303

entff 1

rhofi

velfi

voidfj

mult

125000000

125000000

125000000

0.00015935 0.

* cv 114: total -- -1- s ------ 1of ---- t1est ----- 1-........1----- -- 1-

mass flow out of the test section (kg/s) ----1-----....-1-........-1--------1-----1- ....----

20511400 entflow sum

20511401 0.

20511402 1. cnt3

20511403 1. cnmt

I. 1. 0

ivar 112

lvar 113

72
* ... 1 -... . .... a th-e t e 1 -1 - * cv 115: quality at the tube exit

*----..--------1-........-1-........-1-----1--------1-------.--

20511500 exitqual div

20511501 cntrlvar 114

1. 1.

cntrlvar 112

0

* cv 118: integral of mass flow at tube exit (kg) (R5 var.)

*----...1--------1-----1--------1-........-1-----1---------

20511800 intwgext

20511801 mflowj

integral 1.

125000000

0. 0

cv 119: integral of mass flow at tube exit (kg) (cntrlvar)

*----...1-........-1-........-1--------1-........-1--------1 -

----1-

----1-

20511900 intwgext

20511901 cntrlvar

integral 1.

114

0. 0

* ---- --------- 1 -------- 1 ---- ---- 1 --------1----

* cv 120: mass of water accumulated (lower part)

*----...1-........-1--------1-----1--------1..-.

----1---------1-

----1---------1-

20512000

20512001

20512002

20512003

20512004

20512005

20512006

20512007

20512008

20512009

20512010

20512011

20512012

20512013

20512014

20512015

20512016

20512017

20512018

20512019

accwatl

0.

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

0.1

sum

rho

rho

rho

rho

rho

rho

rho

rho

rho

rho

rho

rho

rho

rho

rho

rho

rho

rho

0.00015935

120010000

120020000

120030000

120040000

120050000

120060000

120070000

120080000

120090000

120100000

120110000

120120000

120130000

120140000

120150000

120160000

120170000

120180000

0. .. 0

*--- mass 1: o 1f ------ 1---------1--------1---------1(upper..t-1 * cv 121: mass of water accumulated (upper part)

*----.1--------1-........-1-........-1--------1-........-1-------.--

20512100

20512101

20512102

20512103

20512104

20512105

20512106

20512107

20512108

20512109

20512110

20512111

20512112

20512113

20512114

20512115

20512116

20512117

20512118

20512119

*

accwatu sum

0.

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.1 rho

0.16 rho

0.00015935

120190000

120200000

120210000

120220000

120230000

120240000

120250000

120260000

120270000

120280000

120290000

120300000

120310000

120320000

120330000

120340000

120350000

120360000

0. 0

73
* -..-..- 1-.....11....1-.....1--11.. -

* cv 122: mass of water accumulated in the test section (kg)

20512200 accwater sum 1. 0. 0

20512201 0.

20512202 1. cntrlvar 120

20512203 1. cntrlvar 121

* cv 140: integral of vapour mass flow at tube exit (kg)

*----1-------1 --------- 1-........-1--------1-----1 ... 1-

20514000 intwgext integral 1. 0. 0

20514001 cntrlvar 112 *

* cv 141: integral of liquid mass flow at tube exit (kg)

*----...1-........-1--------1-........-1--------1-........-1--------1-

*

20514100 intwfext integral 1. 0. 0

20514101 cntrlvar 113 *

* cv 142: sum of total mass accumulated and expelled (kg)

20514200 intwtext sum 1. 0. 0

20514201 0.

20514202 1. cntrlvar 140

20514203 1. cntnlvar 141

20514204 1. cntrlvar 122

* cv 143: integral of liquid mass flow at tube inlet (kg)

20514300 intwfin integral 1. 0. 0

20514301 mflowj 110000000 *

* GENERAL TABLES *$~*$*S*s*s*s*s*S*S*s*s*s*$*S*$*$*S*s*s*$S*s*$*S*$*$*$*s*S*S*$*$*

*----...1-........-1 '1--------1--------1------1-

* tab 110: heat losses

20211000 htc-tenp

20211001 300. 12.13582

20211002 350. 13.81509

20211003 400. 15.72673

20211004 425. 16.77957

20211005 450. 17.90288

20211006 475. 19.10140

20211007 500. 20.38016

20211008 520. 21.46453

20211009 540. 22.60660

20211010 560. 23.80943

20211011 580. 25.07626

20211012 600. 26.41051

20211013 620. 27.81574

20211014 630. 28.54614

20211015 640. 29.29573

20211016 650. 30.06500

20211017 660. 30.85448

20211018 670. 31.66468

20211019 680. 32.49616

20211020 690. 33.34947

20211021 700. 34.22519

74
20211022 710. 35.12390

20211023 720. 36.04622

20211024 730. 36.99275

20211025 740. 37.96413

20211026 750. 38.96103

20211027 760. 39.98410

20211028 770. 41.03404

20211029 780. 42.11154

20211030 790. 43.21734

20211031 800. 44.35218

20211032 810. 45.51682

20211033 820. 46.71204

20211034 830. 47.93864

20211035 840. 49.19745

20211036 850. 50.48932

20211037 860. 51.81511

20211038 870. 53.17572

20211039 880. 54.57205

20211040 890. 56.00505

20211041 900. 57.47568

20211042 910. 58.98492

20211043 920. 60.53380

20211044 930. 62.12335

20211045 940. 63.75463

20211046 950. 65.42876

20211047 960. 67.14684

20211048 970. 68.91004

20211049 980. 70.71953

20211050 990. 72.57655

20211051 1000. 74.48233

20211052 1010. 76.43815

20211053 1020. 78.44532

20211054 1030. 80.50520

20211055 1040. 82.61918

20211056 1050. 84.78866

20211057 1060. 87.01512

20211058 1070. 89.30003

20211059 1080. 91.64495

20211060 1090. 94.05144

20211061 1100. 96.52112

* tab 120 ambient temperature

20212000 temp

20212001 0. 294.26

20212002 1000. 294.26

* tab 150 test section power

20215000 power

20215001 0. 4326.

20215002 1000. 4326.

75
* $$$$$$$$$$ HEAT STRUCTURE THERMAL PROPERTY DATA

* (s-steel and inconnel 600 from Bethsy, copper from TM-IN-816)

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

* ---- - I 1- - 1--- . .-.- 1 .---- I... 1 ...---- 1 ----

20100100 tbl/fctn 1 1 * INCONEL-600

* .-... 1-...1-..1-.. ... 1-... ..--..

----1----

---- 1----

* ...--- 1 -- -1 ---- - ....-.... 1 ....... - 1 ---.---- 1 ----

* INCONEL-600 - thermal conductivity (W/m-K) [tpO02]

20100101 0. 14. 294.0 14.8 * Ist - artiff.

20100102 373.0 15.7

20100103 473.0 17.3

20100104 573.0 18.9

20100105 673.0 20.6

20100106 773.0 22.2

20100107 873.0 23.8

20100108 973.0 25.7

20100109 1073.0 27.4

20100110 1173.0 29.3

20100111 1273.0 31.0

20100112 1373.0 32.7

20100113 1473.0 34.4

20100114 1573.0 36.1

* INCONEL-600 - volumetric heat capacity (J/m3-K)

20100151 0. 3.7+6 294.0 3.740+6 * 1st - artiff.

20100152 373.0 3.900+6

20100153 473.0 4.060+6

20100154 573.0 4.180+6

20100155 673.0 4.330+6

20100156 773.0 4.510+6

20100157 873.0 4.650+6

20100158 973.0 4.830+6

20100159 1073.0 5.020+6

20100160 1173.0 5.200+6

20100161 1273.0 5.340+6

20100162 1373.0 5.630+6

20100163 1477.0 5.810+6

.*** MND ***

76
,Appendix I -

. - - o .

Sample input deck for quasi-steady state analysis

• o ,I

,. ._ . - : - .:: 

* (p

-''-

C."," . , . . r.. • f. ! r -

,,77
UCB351

* UNIV. OF CALIFORNIA BERKELEY (UCB) TEST

* REFLOOD TEST IN SINGLE TUBE GEOMETRY

* UCB-3051: P=2 BAR, G=25.1 KG/M'*2/S, XQ-=0.09

* GOAL: CALCULATION OF THE AXIAL VAPOUR

* TEMPERATURE DISTRIBUTION WHEN THE QUENCH FRONT

* IS AT HALF-LENGTH ELEVATION (Z=1.89 H)

* REPRESENTATION OF THE SECTION ABOVE QUENCH FRONT ONLY (Z>1.8 M)

* REFLOODING OPTION USED

* input: Power and heat losses

* RELAPS/MOD2.5 calculation

* for RELAPS/MOD3 only the card nr. 11200801 has to be changed

*66050m..66066S*...@USSm.m.u...66m6U660mmSUCUmUm.mam*Ummm*NBU*mUSUm605m

00000100 new transnt

00000101 run

* cpu-lim

00000105 5.0 10.0 880.

00000110 nitrogen

* TIME STEP CONTROL CARDS * REQUIRED

* END TIME MIN DT MAX DT OPTN MNR MJR RST

00000201 15 1.0-9 0.01 3 10 400 400

* MINOR EDIT REQUESTS

00000301 cputime 0

00000308 voidg 120100000

00000309 voidg 120110000

00000310 voidg 120120000

00000311 voidg 120130000

00000312 voidg 120140000

00000313 voidg 120150000

00000314 voidg 120170000

00000315 voidg 120190000

00000319 temvi 120190000

00000320 120110000 00000321 qwg 

00000322 q 120120000

00000323 qwg 120120000

00000324 qwg 120130000

00000325 qwq 120140000

00000326 qwg 120150000

00000327 qwg 120160000

00000328 qwg 120170000

00000329 qwg 120180000

00000330 qwg 120190000

00000331 mflowq 120110000

00000332 mflowj 120120000

00000333 mflowj 120130000

00000334 mfl*wj 120140000

00000335 mflowj 120150000

00000336 mflowj 120160000

00000337 mflowj 120170000

78
00000338 mflowj

00000340 mflowj

00000342 cntrl

00000343 cntrlv

00000344 cntrlv

00000345 cntrlv

00000346 cntrlv

00000348 cntrlv

00000349 cntrlv

00000351 cntrlv

00000352 catrlv

00000353 cntrlv

00000356 cntrlv

00000357 cntrlv

00000358 cntrlv

00000359 cntrlv

00000360 cntrlw

00000361 cntrlw

00000362 cntrlv

00000363 cntrlv

00000364 cntrlv•

00000365 cntxrlv

00000370 cn=trlv.

00000371 cntrlv

00000372 cntrlv

00000373 cntrlv

00000374 cntrl

00000375 cntrlv

00000376 cntrlx

00000377 cntrlv

00000378 =trlvi

00000380 floreg

00000381 floreg

00000382 floreg

00000383 floreg

00000384 floreg

00000385 floreg

00000386 floreg

00000387 floreg

00000388 floreg

00000389 floreg

00000391 httemp

00000392 httemp

00000393 httemp

00000394 httemp

00000395 httemp

*00000397 ztxwt

ar

ar

ar

ar

ar

ar

ar

ar

ar

ar

ar

"at

ar

'at

ar

arIr

!Lr ar

ar

Rr

I•r

ar

ar

ar

ar

ar

ar

AX

120180000

125000000

101

102

103

104

105

106

110

1il

115

120

125

128

129

133

134

138

139

143

144

149

150

151

152

166

167

168

169

173

182

120100000

120110000

120120000

120130000

120140000

120150000

120160000

120170000

120180000

120190000

120001001

120001101

120001301

120001601

120001801

1200000

*

*

*

TRIP COMPONENTS

500 time 0 qt null'

501 quale 120180000 le null

0. 400. 1 *end

0. 20.00001 1 * full

0. 0. 1 * start injection

and reflood

0. 10. 1 * start averaging

502 time 0 gt null

503 gt null

600

time 0

500

79
* HYDRODYNAMIC COMPONENTS

* C105] feedwater tank - inlet thermodynamic conditions

*---- 1----- ---- 1 ---- ------------ 1 -------- 1 -------- 1 -------- 1-

1050000 "feedtan* tmdpvol

1050101 0.00015935 0. 1000. 0. 0. 0. 0. 0. 0

1050200 002 * i.c.: pX; saturated

1050201 0. 2.15+5 .008

* ---- 1 -------- 1 ---- ------------ 1 ---------1 -------- 1 ------ I-

* [110) inlet junction - flow boundary conditions

1100000 "feedjun" tmdpjun

1100101 105000000 120000000 0.

1100200 1 502 time 0

1100201 -1. 0.0 0.0 0.

1100202 0. 0.0 0.0 0.

1100203 0.0001 0.003707 0.00030 0.

1100204 100.0 0.003707 0.00030 0. *mcell=G*Acell

* ------------- 1 ---- ---- 1---- --- 1 -------- 1 -------- 1 -------- 1-

* (120] tube

1200000 Otestsec" pipe

1200001 19

1200101 0.00015935 19 * Avcell

1200201 0. 18 * Ajunct

1200301 0.03 10 * cell length dz

1200302 0.2 18 * 1200303 0.16 19

1200401 0. 19 * cell volume=dz*Avcell

1200501 0. 19 * hor. angle

1200601 90. 19 * vert. angle

1200701 0.03 10 * cell dz

1200702 0.2 18 *

1200703 0.16 19 * 1200801 1.-4 0.01425 19 * friction data

1200901 0. 0. 18 * loss coefficients

*1201002 100 11 * bundle correlation on

1201002 0 19 * tube correlation on

1201101 00000 18

1201201 2 2.00+5 1. 0. 0. 0. 1 * initial

1201202 2 2.00+5 1. 0. 0. 0. 2 * conditions

1201203 2 2.00+5 1. 0. 0. 0. 3

1201204 2 2.00+5 1. 0. 0. 0. 4

1201205 2 2.00+5 1. 0. 0. 0. 5

1201206 2 2.00+5 1. 0. 0. 0. 6

1201207 2 2.00+5 1. 0. 0. 0. 7

1201208 2 2.00+5 1. 0. 0. 0. 8

1201209 2 2.00+5 1. 0. 0. 0. 9

1201210 2 2.00+5 1. 0. 0. 0. 10

1201211 3 2.0+5 420. 0. 0. 0. 11

1201212 3 2.0+5 430. 0. 0. 0. 12

1201213 3 2.0+5 440. 0. 0. 0. 13

1201214 3 2.0+5 450. 0. 0. 0. 14

1201215 3 2.0+5 460. 0. 0. 0. 15

1201216 3 2.0+5 470. 0. 0. 0. 16

1201217 3 2.0+5 480. 0. 0. 0. 17

1201218 3 2.0+5 490. 0. 0. 0. 18

1201219 3 2.0+5 500. 0. 0. 0. 19

1201300 1

1201301 0. 0.00030 0. 1 * mass flows

1201302 0. 0.00035 0. 2

1201303 0. 0.00037 0. 3

80
1201304 0. 0.00039 0. 4

1201305 0. 0.00040 0. 5

1201306 0. 0.00041 0. 6

1201307 0. 0.00042 0. 7

1201308 0. 0.00043 0. 8

1201309 0. 0.00045 0. 9

1201310 0. 0.00047 0. 10

1201311 0. 0.00052 0. 11

1201312 0. 0.00058 0. 12

1201313 0. 0.00063 0. 13

1201314 0. 0.00066 0. 14

1201315 0. 0.00069 0. 15

1201316 0.. 0.00072 0. 16

1201317 0. 0.00075 0. 17

1201318 0. 0.00078 0. 18

* ---- 1 ---- ---- 1 -------- 1 -------- 1---------------- 1 -------- 1-

* [1251 outlet junction

*----...1-----1------1--------1-........-1-........-1-----1-

1250000 aoutljun" sngljun

1250101 120010000 130000000 0. 0.0 0.0 00000

1250201 1 0. 0.00078 0.0

* ---- 1 ---- ---- 1 ---- ---- ---- -------------------- 1 -------- 1-

* [130] outlet tank

*----...1-----1-i-........1-........-1-----1--------1-........-1-

1300000 moutltan. tmdpvol

1300101 .01 0. 1000. 0. 0. 0. 0. 0. 0

1300200 003

1300201 0. 2.0+5 480.

* NEAXT STRUCTURES

*$**$****S*$****$***$S*$*SS*SS*$*

** ---.---- 1 ----.---- 1 -------- 1 ---- (1200] test section(tube) - above

**--- 1--------1--------1-

---- 1 ----. ---- I.-1--- ----... -- 1 .--- *

quench front

----I -------- 1 -------- 1 ------- -*

11200000

11200100

11200101

11200201

11200301

11200400

11200401

+

11200402

+

11200403

1 0

11200404

+

11200405

+

11200406

+

11200407

+

11200408

11200409

+

11200410

+

11200411

4..

11200412

4.

19

0

9

1

1.

-1

400.

400.

450.

450.

555.

555.

575.

575.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

645.

645.

10 2 0

1

0.00795 * outer tube radius

9

9

400.

400.

450.

450.

555.

555.

575.

575.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

645.

645.

400.

400.

450.

450.

555.

555.

575.

575.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

645.

645.

400.

450.

555.

575.

595.

400.

450.

555.

575.

595.

400.

450.

555.

575.

595.

400.

450.

555.

575.

595.

595. 595. 595. 595.

*

*

*

*

*

*

*

*

*

*

*

*

*

*

*

*

*

*

*

*

*1

*1

*1

0.007125 502 0 64

initial

temperatures of

the nodes

(above QF)

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595.

595. 595. 595. 595.

645. 645. 645. 645.

81
11200413

+

11200414

124

11200415

124

11200416

11200417

+

11200418

11200419

+

11200501

11200502

11200503

11200601

11200602

11200603

11200604

11200605

11200606

11200607

11200608 11200609

11200610

11200701 11200702

11200703

11200801

*11200801

670.

670.

690.

690.

710.

710.

720.

720.

726.

726.

733.

733.

733.

733.

670

670

690

690

710

710

720

720

726

726

733

733

733

733

120010000

120110000

120190000

0

0

0

0

0

0

0

0

0

0

150

150

150

0

0.

. 670. 670. 670.

I. 670.

. 690. 690. 690.

. 690.

I. 710. 710. 710.

I. 710.

I. 720. 720. 720.

. 720.

5. 726. 726. 726.

. 726.

1. 733. 733. 733.

1. 733.

. 733. 733. 733.

. 733.

10000 1

10000 1

0 1

0 0

0 2121

0 2122

0 2123

0 2124

0 2125

Q 2126

0 2127

0 2128

0 2129

0.00819672

0.0546448

0.0437158

0. 0.

11. 11.

670. 670.

690. 690.

710. 710.

720. 720.

726. 726.

733. 733.

733. 733.

**

**

**

**

**

**

**

**

**

**

**

**

**

**

1 0.03 10 * inner surf

1 0.2 18 **

1 0.16 19 **

1 0.03 10 * outer surf

1 0.2 11*

1 0.2 12 *

1 0.2 13 *

1 0.2 14*

1 0.2 15*

1 0.2 16 *

1 0.2 17*

1 0.2 18 *

1 0.16 19 *

0. 10 * power distribution

0. 18 *

0. 19 *

0. 19 * M2

0. 0. 0. 0. 1. 19 * M3

0.

0.

0.

* CONTROL VARIABLES

* cv 101: collapsed liquid level in the test section

*----...1-........-1-----1----- 1---------1---------.1-. ....----

20510100

20510101

20510102

20510103

20510104

20510105

20510106

20510107

20510108

20510109

20510110

20510111

20510112

20510113

20510114

20510115

20510116

20510117

20510118

20510119

20510120

*

collev

0.

0.03

0.03

0.03

0.03

0.03

0.03

0.03

0.03

0.03

0.03

0.2

0.2

0.2

0.2

0.2

0.2

0.2

0.2

0.16

sum

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

voidf

1.

120010000

120020000

120030000

120040000

120050000

120060000

120070000

120080000

120090000

120100000

120110000

120120000

120130000

120140000

120150000

120160000

120170000

120180000

120190000

0. 1

82
* cv 102: vapour mass flow from the test section

*----...1-........-1------1-........-1------1-----1--------1-

20510200 entgfl mult 0.00015935 0. 1

20510201 rhogj 125000000

20510202 velgi 125000000

20510203 voidgj 125000000

20510204 cntrlvar 153

*----...1-........-1-----1---........1--------1-........-1-........-1-

* cv 103: liquid mass flow from the test section

20510300 entffl mult 0.00015935 0. 1

20510301 rhofj 125000000

20510302 velfj 125000000

20510303 voidfj 125000000

20510304 cntrlvar 153

* cv 104: total mass flow out of the test section

20510400 entflow sum 1. 1.-I 1 1 1.-i0

20510401 0.

20510402 1. cntrlvar 102

20510403 1. cn=trlvar 103

* cv 105: quality at the tube exit

*- - - 1-------- 1-........-1-........-1-.... ....-1-........-1-.... ....- 1-

20510500 exitqual div 1. 0. 1

20510501 cntrlvar 104 cntrlvar 102

* cv 106: phase velocity difference in volume 13

20510600 diffvl3 sum 1. 0. 1

20510601 0.

20510602 1. velg 120130000

20510603 -1. velf 120130000

* cv 107:

20510700 abscvl06 stdfnctn 1. 0. 1

20510701 abs cntrlvar 106

* cv 108:

*----.1--------1-........-1-........-1-........-1--------1--------1-

20510800 cvlO7slO mult 1. 0. 1

20510801 cntrlvar 153 cntrlvar 107

* cv209:

*----...1--------1-........-1-----1--------1-........-1--------1-

20510900 cvl07int integral 1. 0. 1

20510901 cntrlvar 108

83
* ---- 1--. 1 -- - 1 ---- 1 ...

* cv 110: average phase velocity difference (l0-20s) in

20511000 avdvl3 div 1. 0. 1

20511001 cntrlvar 154 cntrlvar 109

* cv 111: phase velocity difference in volume 16

20511100 diffvl6 sum 1. 0. 1

20511101 0.

20511102 1. velg 120160000

20511103 -1. velf 120160000

* cv 112:

20511200 abscvlll stdfnctn 1. 0. 1

20511201 abs cntrlvar 111

* cv 113:

20511300 cvll2slO mult 1. 0. 1

20511301 cntrlvar 153 cntrlvar 112

* cv 114:

20511400 cvll3int integral 1. 0. 1

20511401 cntrlvar 113

*--- 1----1 -------1-------I. .. 1-........-1....

* cv 115: average phase velocity difference (10-20s) in

20511500 avdvl6 div 1. 0. 1

20511501 cntrlvar 154 cntrlvar 114

* cv 118: vapour mass flow at junction 10

20511800 entgjlO mult 0.00015935 0.

20511801 rhogi 120100000

20511802 velgi 120100000

20511803 voidgi 120100000 t

* cv 119: liquid mass flow at junction 10

".... .1--------1-----1---------1-........-1-...

20511900 entfjlO mult 0.00015935 0.

20511901 rhofj 120100000

20511902 velfj 120100000

20511903 voidfj 120100000

* ..... 1--------1---- ---1-------1------i-

* cv 120: total mass flow at junction 10

20512000 mflowjlO sum 1. 0. 1

20512001 0.

cell 13 ....- 1-----1

---- 1----

---- 1----

---- 1-

---- 1-

----1-----....-1-

....- 1-------.--1

cell 16

Q----1----

1

1

----1-

----1-

----1-

----1-

----1-

----1-

----1-

----1-

----1-

----1-

----1-

----1-

84
20512002 1. cntrlvar 118

20512003 1. cntrlvar 119

* cv 121: quality at junction 10

20512100 qualJilO div 1. 0. 1

20512101 cntrlvar 120 cntrlvar 118

* cv 122: vapour mass flow at junction 11

*----1---------1--i --- 1--------1-........-1--------....-1-.....1-

20512200 entgj11 mult 0.00015935 0. 1

20512201 rhogi 120110000

20512202 velgi 120110000

20512203 voidgj 120110000

20512204 cntrlvar 153

cv 123: liquid mass flow at junction 11

20512300 entfjii mult 0.00015935 0. 1

20512301 rhofj 120110000

20512302 velfj 120110000

20512303 voidfj 120110000

20512304 cntrlvar 153

cv 124: total mass flow at junction 11 ----... 1-........-1--------1-........-1-........-1--------1--------1-

20512400 mflowjll su= 1. 1.-I0 1 1 1.-10

20512401 0.

20512402 1. cntrlvar 122

20512403 1. cntrlvar 123

* cv 125: quality at junction 11

*----.1--------1-........-1-........-1--------1-........-1-----.---i-1

20512500 qualjll div 1. 0. 1

20512501 catrlvar 124 cntrlvar 122

* cv 126: vapour mass flow at junction 16

20512600 entgj16 mult 0.00015935 0. 1

20512601 rhogj 120160000

20512602 velgj 120160000

20512603 voidgj 120160000

20512604 cntrlvar 153

* cv 127: liquid mass flow at junction 16

20512700 entfJ16 mult 0.00015935 0. 1

20512701 rhofj 120160000

20512702 velfj 120160000

20512703 voidfj 120160000

20512704 cntrlvar 153

85
* ----. -.-- 1 . .... --- I .... ....- 1 ----. .... 1 ----. .... 1 .... * cv 128: total mass flow at junction 16

*----1-------1 1---1------1-........-1----. ----- ----1-----....1-

1 1. -10 20512800

20512801

20512802

20512803

mflowjl6 sum

0.

1. cnt3

1. cnt2

1. 1

:lvar 126

rlvar 127

* ---.---- 1 ----.---- 1 ----.---- 1 -------- 1 -------- 1----

* cv 129: quality at junction 16

*----...1-........-1-........-1-----1--------1-....

*

----1-----....-1-

-...- 1-----....-1-

20512900 qualj16 div

20512901 cntrlvar 128

1. 0.

cntrlvar 126

1

* --------- 1 ----.----. 1---- ---1 ----.---- 1 -------- 1 ---- * cv 130: integral of vapour mass flow at junction 11

*- - - -... 1-.... ....-1-.... ....- 1-.... ....- 1-.... ....- 1..-. .... 1-----....-1-

20513000

20513001

intwgj 11

cntrlvar

integral 1.

122

0. 0

* ---.---- 1 ----.---- 1 -------- 1 -------- 1 -------- 1----

c cv 131: integral of liquid mass flow at junction 11

*- - - -... 1-.... ....-1-.... ....- 1-.... ....- 1-.... ....- 1 --..

---- 1-----....-1-

.... 1-------..-I20513100

20513101

intwfj 11

cntrlvar

integral 1.

123

0. 0

* ---.---- 1 -------- 1 -------- 1 -------- 1 -------- 1 ---- * cv 132: integral of total mass flow at junction 11 - - - -... 1-.... ....-1-.... ....- 1-.... ....- 1-.... ....-1-....

20513200 intwtjll sum 1 I.-i0

20513201 0.

20513202 1. cnt2

20513203 1. cnt2

1. 1.-i0 1

rlvar 130

rIvar 131

* ---- 1----- -1- -------- 1---- ---- 1 ---- * cv 133: average total mass flow at junction *- - - -... 1-.... ....-1-.... ....- 1-.... ....- 1-..--

---- 1----

11

---- 1----

-...-1-------..-i-

.... 1-------..-1-

20513300 avwtjll div

20513301 cntrlvar 154

* cv 134: average quality

20513400 avxajll div

20513401 cntrlvar 132

1. 0.

cntrlvar 132

0

---- 1 ---- ---- 1---- -------- at junction 11

---- 1 -------- 1 -------- 1---- .... 1-----....-1-

1. 0.

cntrlvar 130

0

* ----- 1 ----.---- 1 ------ 1- ---- -------- ----1----

cv 135: integral of vapour mass flow at junction 16 .... ...1-------- 1-.... ....- 1-.... ....-1-.... ....-1-- *

...- 1-------..-1-

....-1-------..--

20513500

20513501

intwgj 16

cntrlvar

integral 1.

126

0. 0

* ---- 1 -------- 1 -------- 1 -------- I -------- 1----

cv 136: integral of liquid mass flow at junction 16 ---- 1---------1--------- 1-........-1---------1--- *

....-1-----....-1-

.... 1-----....-1-

20513600 intwfjl6

20513601 cntrlvar

integral 1.

127

0. 0

86
* ----. -.-- 1 ---- --- 1 ---- --- 1 ----. .... 1 ----. .... 1 - - * cv 137: integral of total mass flow at junction 16

*.. . . .1i .. .1- . . . 1

1 1 .- i0 20513700

20513701

20513702

20513703

intwtj16 sum

0.

1. 1

1.

cntrlvar 135

cntrlvar 136

* ---- 1 ---- ---- 1-------- 1 -------- 1 -------- I----

* cv 138: average total mass flow at junction 16

*--- 1------1------1--------1-------i1----

....- 1-------.--1

..-.1-----1-

20513800 avwtjl6 div

20513801 cntrlvar 154

* cv 139: average quality

20513900 avxajl6 div

20513901 cntrlvar 137

1. 0.

cntrlvar 137

0

at junction 16 ----1-----....-1--------1-----1------i

1. 0.

cntrlvar 135

0

* --.---- 1 -------- I---- ------------ ---- ---- ---- * cv 140: integral of vapour mass flow at tube exit

*----------1-----1----- 1--------1-........-1--

---1---------

-1---- ..---- 1

20514000 intwgext

20514001 cntrlvar

integral 1.

102

0. 0

* ----.---- 1 ---------1------- 1 -------- 1 ----.---- 1----

* cv 141: integral of liquid mass flow at tube exit

*------------1------1--------1--------1-........-1--..-

.... 1--- ...- I--1

....- 1--------1-

20514100

20514101

intwfext

cntrlvar

integral 1.

103

0. 0

* --------- 1--- 1 .-- - ----1 -- -------- 1 1 --------- 1 ----

* cv 142: integral of total mass flow at tube exit

*-------------1-........-1-........-1--------1-........-1---..

---- 1----

---- i--

---- i-

---- 1-

20514200

20514201

20514202

20514203

*

intwtext sum

0.

1. cnt2

1. cnt3

1. 1.-10 1 1 1.-I0

•lvar 140

?lvar 141

* --------- 1 .------ - ---- 1 1 ----- -- 1 --- 1- -------- 1 ---- cv 143: average total mass flow at tube exit

*-------------1-........-1-........-1--------1-........-1....

*

....1-------..-1-

....-1------ --1

20514300 avwtext div

20514301 cntrlvar 154 *

1. 0.

cntrlvar 142

0

*-------------.-- ----- -1--- ---- 1 -------- 1 -------- 1----

* cv 144: average quality at tube exit

*-------------1--------1-........-1--------1-........-1...- ....-1-----....-1-

20514400

20514401

avxaext div

cntrlvar 142

1. 0.

cntrlvar 140

0

*.--------- I---------I ----.---- 1- ---- -------- ---- 1 ---- * cv 149: equilibrium quality at the quench front

*-------------1--------1-........-1--------1-........-1...-

*

20514900 0

20514901

20514902

xeqqf SUM

0.075

1.13355258-4

1. 1. 0.

cntrlvar 150

87
* .... .... 1 - - ---- 1 .... .... 1 ---- - I - -- i-

* cv 150: power input to wet portion of test section

*----1-----1----1--------1-----1---........1-........-1-. ....----

20515000 qinwet

20515001 0.

20515002 1.

20515003 1.

20515004 1.

20515005 1.

20515006 1.

20515007 1.

20515008 1.

20515009 1.

20515010 1.

20515011 1.

1. 0. 1

q

q

q

q

q

q

q

q

cz

q

120010000

120020000

120030000

120040000

120050000

120060000

120070000

120080000

120090000

120100000

cv 151: power input to dry portion of test section *----1-------1 --------- 1--------1-........-1-........-1-........--

20515100

20515101

20515102

20515103

20515104

20515105

20515106

20515107

20515108

20515109

20515110

*

qindry

0.

1.1.

1.

1.

1.

1.

1.

1.

1.

sum I. 0. 1

q

q

q

q

q

q

q

q

q

120110000

120120000

120130000

120140000

120150000

120160000

120170000

120180000

120190000

* cv 152: total ----1 -------- 1----

heat input to fluid

---- 1 -------- ---- in test section .... 1-----....-1-........-1-----1

20515200

20515201

20515202

20515203

qintot

0.1.

1.

sum 1. 0. 1

cntrlvar 150

cntrlvar 151

* ---- 1 ---- ------------ 1----

* cv 153: trip unit

*----1-----1----1------1--

....1-----...1-....2.---------1----- --1

.... 1-----....-1-........-1------ --1

20515300 trip

20515301 503

tripuni t 1. 0. 0

* ---- 1 -------- I -------- 1----

cv 154: delay

20515400 delay sum 1.

20515401 -10.

20515402 1. time 0

.... 1-----....-1-........-1--------1

----1 --------- 1 ----

0.0000001 1 ....- 1-------. --

....1-------..-1-

....1-------..-I-

* ---- 1 -------- 1 ------- 1---- ---- 1----

cv 160: heat added to wet part (after 10 s)

*----1--------.....1-----1--------1-

--1--

--1--

20516000 htwetlO mult

20516001 cntrlvar 150

I. 0.

cntrlvar 153

0

*---- 1 -------- 1 -------- 1 --------. ---- * cv 161: heat added to dry part (after 10 s)

*----.1--------1-........-1--------1-

--1--

--1--

....- 1 .... ....-1 -

....1-------..-1-

88
*

20516100

20516101

htdrylO mult

cntrlvar 151

1. 0.

cntrlvar 153

0

* cv 162: heat added to fluid (after 10 s)

*----...1--------1-----1--------1-........-1-........-1-. ....- i--1

20516200 httotlO mult

20516201 cntrIvar 152

1. 0.

cntrlvar 153

0

* ---- -- ------- 1 -------- 1----

* cv 163: total heat added to fluid

20516300 htinwet integral 1.

20516301 cntrlvar 160

* cv 164: total heat added to fluid

*----...1-........-1---------1-Iin wet region (after 10 s) ...-- .. 1------..-----1------ ... --1

0. 1

---- ---- --- 1---- ---- 1 ----

in dry region (after 10 s)

---- 1 ----.---- 1 -------- 1 ----

---- 1-

---- i20516400

20516401

htindry integral 1.

cntrlvar 161

0. 1

* ---- 1 ---- ---- 1 ------------ * cv 165: total heat added to fluid

*----...1-........-1--------1.... (after 10 s) .... 1--------1-........-1-........-1-

20516500 htintot integral 1. 0. 1

20516501 cntrlvar 162

* cv 166: average heat transfer rate in wet region (10-20 s)

*- - - -... 1-.... ....-1-.... ....-1-.... ....- 1-------- 1-........-1-...-

---- 1-

---- 1-

20516600

20516601

*

avhtwet div

cntrlvar 154

1. 0.

cntrlvar 163

0

* --.---- 1 ------------ ---- 1 -------- ---- --- 1---- ---- 1----

* cv 167: average heat transfer rate in dry region (10-20 s)

*----1--------1--------1-........-1--------1-----1-

---- 1-

---- 1-

20516700 avhtdry div

20516701 cntrlvar 154

1. 0.

cutrlvar 164

0

* cv 168: average heat transfer rate to fluid (10-20 s)

*----...1-........-1--------1-........-1--------1-........-1-------.--1

20516800 avhttot

20516801 cntrlvar

div

154

1. 0.

catrlvar 165

0

* ---- 1- -------- 1 -------- 1 -------- 1 -------- 1 -------- 1----

cv 169: heat transfer to vapour only in the dry region ---- 1--------1-........-1-........-1--------1-........-1--...

---- 1-

---- 1-

20516900

20516901

20516902

20516903

20516904

20516905

20516906

20516907

20516908

20516909

20516910

qdryvap 0.

1.

1.

1.

1.

1.

1.

1.

1.

1.

sum

qW9

qw9

qW9

qwg

qwg

qwg

1. 0. 1

120110000

120120000

120130000

120140000

120150000

120160000

120170000

120180000

120190000

89
* ---- --- .. . ....- 1 --- --- -- -I i .. -- .- .- .--I- -- - - 1 ----. .... 1 ----. .... 1-

* cv 171: heat added to vapour in dry region (after 10 s)

*----...1--------1-----1-i-........1-------...---.......1-........-1-

20517100 htdrylO mult 1. 0. 0

20517101 cntrlvar 169 cntrlvar 153

* cv 172: total heat added to vapour in dry region (after 10 s)

*----...1--------1-----1-i-........1--------1-........-1-........-1-

20517200 htgdry integral 1. 0. 1

20517201 cntrlvar 171 1

* cv 173: average heat transfer rate to vapour in dry region (10-20 s)

*----1-------1 ---------1--------1-........-1-........-1-. ....----

20517300 avqwgdry div 1. 0.

20517301 cntrlvar 154 cntrlvar 172

* cv 180: vapour temperature after 10 s

20518000 tempg10 mult 1. 0.

20518001 tempg 120190000 cntrlvar 153

*----...1-----1-----1---........1..-.

cv 181:

20518100 tempint integral 1. 0.

20518101 cntrlvar 180

* cv 182: average temperature (10-20 s)

20518200 avtemp div 1. 0.

0

.... 1------

---- 1----

0

------ 1------

0

---- 1i-

---- i--

---- 1-i

---- 1----

---- 1-

---- 1-

---- 1-

---- 1-

---- 1-

---- 1-

....-1-----....-1....

....-1---------..1....

0

20518201 cntrlvar 154 cntrlvar 181

* GENERAL TABLES

* tab 110-129 heat losses and heat fluxes

20211000 htrnrate

20211001 0. 26402.

20211002 1000. 26402. *

20211100 htrnrate

20211101 0. 16930.

20211102 1000. 16930.

20211200 htrnrate

20211201 0. 15470.

20211202 1000. 15470.

20211300 htrnrate

20211301 0. 13930.

20211302 1000. 13930.

20211400 htrnrate

20211401 0. 12395.

20211402 1000. 12395.

90
20211500 htrnrate

20211501 0. 10980.

20211502 1000. 10980. *

20211600 htrnrate

20211601 0. 9700.

20211602 1000. 9700.

20211700 htrnrate

20211701 0. 8843.

20211702 1000. 8843.

20211800 htrnrate

20211801 0. 8195.

20211802 1000. 8195.

20211900 htrnrate

20211901 0. 8195.

20211902 1000. 8195.

20212100 htrnrate

20212101 0. 8494.

20212102 1000. 8494.

20212200 htrnrate

20212201 0. 9802.

20212202 1000. 9802.

20212300 htrnrate

20212301 0. 11174.

20212302 1000. 11174.

20212400 htrnrate

20212401 0. 12553.

20212402 1000. 12553.

20212500 htrnrate

20212501 0. 13823.

20212502 1000. 13823.

20212600 htznrate

20212601 0. 14973.

20212602 1000. 14973.

20212700 htxnrate

20212701 0. 15737.

20212702 1000. 15737.

20212800 htrnrate

20212801 0. 16317.

20212802 1000. 16317.

20212900 htrnrate

20212901 0. 16317.

20212902 1000. 16317.

* tab 150 test section power

20215000 power

20215001 0. 4326.

20215002 1000. 4326.

91
* $$$$S$$$$$ HEAT STRUCTURE THERMAL PROPERTY DATA

* (s-steel and inconnel 600 from Bethsy, copper from TM-IN-816)

20100100 tbl/fctn 1 1 * INCONEL-600

* INCONEL-600 - thermal conductivity (W/m-K) [tp0021

20100101 0. 14. 294.0 14.8 * ist artiff.

20100102 373.0 15.7

20100103 473.0 17.3

20100104 573.0 18.9

20100105 673.0 20.6

20100106 773.0 22.2

20100107 873.0 23.8

20100108 973.0 25.7

20100109 1073.0 27.4

20100110 1173.0 29.3

20100111 1273.0 31.0

20100112 1373.0 32.7

20100113 1473.0 34.4

20100114 1573.0 36.1

* INCONEL-600 - volumetric heat capacity (J/m3-K)

20100151 0. 3.7+6 294.0 3.740+6 1st artiff.

20100152 373.0 3.900+6

20100153 473.0 4.060+6

20100154 573.0 4.180+6

20100155 673.0 4.330+6

20100156 773.0 4.510+6

20100157 873.0 4.650+6

20100158 973.0 4.830+6

20100159 1073.0 5.020+6

20100160 1173.0 5.200+6

20100161 1273.0 5.340+6

20100162 1373.0 5.630+6

20100163 1477.0 5.810+6 .***EN**

92
NRC FORM s3s US. NUCLEAR REGULATORY COMMISSION 1. REPORT MBER

(24.a) (Asslgred by eRC, Add VoL. Sup.. Rev. NRCM 110Z, Wid Addendu Wxe Wanyf.)

3.= BIBUOGRAPHIC DATA SHEET

2-. TTLE AMn SUBTTLE. NUREGCAA-0163

A Study of the Dispersed Flow Interfacial Heat Transfer Model of

RELAP5/MOD2.5 and RELAP5/MOD3 3. DATE REPORT PUBLISHED

MONTH EAR

April 1999

4. FIN OR GRANT NUMBER

D___7

5. AUTHOR(S) 6. TYPE OF REPORT

M. Andreanr

G. Th. Analytis

S. N. Aksan 7. PERIOD COVERED Onve Daes)

8. PERFORMING ORGANIZATION - NAME AND ADDRESS 1VNRC, •ovideDi•% Ote w fRg, U'S AW&-RegRAWY ya issw and Raft addnag Ioonbcb

mv-de -ame and meftg ,dd*ssJ

Thermal-Hydraulics Laboratory *ETHZ. Nuclear Engineering Laboratory

Paul Scherrer Institute (PSI) Swiss Federal Institute of Technology

CH-5232 Villigen PSI, Switzerland 8092 Zurich Switerland

g. SPONSORING ORGANIZATION - NAME AND ADDRESS XAC p Mom ~ams *bow' AWcvnracbw pwd eNRCav, Mkx 016oW fnx U.& AkcbeerAigzdetuy Cawawnmsi

*nd nftSg dd-sj

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

10. SUPPLEMENTARY NOTES

G. Rhee. NRC Project Manaaer

11. ABSTRACT g200 ards or/ss)

For conditions of low quality at the quench front and low mass flux, the interfacial heat transfer calculated assuming a uniform

distribution of droplets over the cross-sectional area of the channel is necessarily overpredicted, and the vapour superheat is

strongly underpredicted. The purpose of the present paper is to show that lmitations of the 1-0 approach, obtained from the

steady-state analyses of slow reflooding experiments, has some impact on the performance of the 1-D transient computer codes

Ike RELAP5JMOD2.5 and RELAP51MOD3. As an example, the transient analysis of a low flooding rate experiment In a tube was

performed. An early completion of the quench process and a fast desuperheating of the vapour at the tube exdt was obtained by

both codes. The too high quench front velocity (four times higher than In the experiment) could not, however, be put univocally in

relation to the underpredicton of the vapour temperature, and the consequent increase of the precursory cooing, as many coupled

thermal and hydrauic transient effects prevailed. Quasi steady-state analyses of two runs, where the boundary conditions for the

post-dryout region could be better controlled for a predetermined position of the quench front, were thus performed. These

analyses show that the vapour superheat at the tube exdt is strongly underpredicted, confirming the Imitflions of the 1-0 model for

Interfacial heat transfer in the dispersed flow region..

12. KIEY WORDSJDESCRIPTORS &.W -•3ud a.,esef w~assiztse13hws , •cai me ¶.. AVALABL.Y S'rATEMENT

Assessment of RELAPS, Reflood Heat Transfer unrimited

Dispersed Flow, Post-CHF Heat Transfer 1s Cur ac1JssnCA1=oN

Interfacial Heat Transfer Model MT3 P824

RELAP&MOD2.5 unclassified

RELAP5/MOD3 OM Rp"-

unclassified

15. NUMBER OF PAGES

16. PRICE

NRC FORM =S C24M Thig UM Vin ejecOonv=11y prodijmd by Oft Fedual ForM I=
printed1~

Federal Recycling Program

NUREGIIA-0163 A STUDY OF THE DISPERSED FLOW INTERFACIAL HEAT TRANSFER MODEL OF

RELAPS/MOD2.5 AND RELAPS/MOD3

APRIL 1999

UNITED STATES

NUCLEAR REGULATORY COMMISSION

WASHINGTON, DC 20555-0001

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POSTAGE AND FEES PAID

USNRC

PERMIT NO. G-67

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PENALTY FOR PRIVATE USE, $300