NUREG/IA-0113

International

Agreement Report

Preliminary Assessment of PWR

Steam Generator Modelling

in RELAP5/MOD3

Prepared by

R. J. Preece, J. M. Putney

National Power

Technology and Environmental Centre

Kelvin Avenue

Leatherhead, Surrey

KT22 7SE, United Kingdom

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

July 1993

Prepared as part of

The Agreement on Research Participation and Technical Exchange

under the International Thermal-Hydraulic Code Assessment

and Application Program (ICAP)

Published by

U.S. Nuclear Regulatory Commission
NOTICE

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

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

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and

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NUREG/JA-0113

International

Agreement Report

Preliminary Assessment of PWR

Steam Generator Modelling

in RELAP5/MOD3

Prepared by

R. J. Preece, J. M. Putney

National Power

Technology and Environmental Centre

Kelvin Avenue

Leatherhead, Surrey

KT22 7SE, United Kingdom

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

July 1993

Prepared as part of

The Agreement on Research Participation and Technical Exchange

under the International Thermal-Hydraulic Code Assessment

and Application Program (ICAP)

Published by

U.S. Nuclear Regulatory Commission
NOTICE

This report is based on work performed under the sponsorship of the

United Kingdom Atomic Energy Authority. The information in this

report has been provided to the USNRC under the terms of the

International Code Assessment and Application Program (ICAP)

between the United States and the United Kingdom (Administrative

Agreement - WH 36047 between the United States Nuclear Regulatory

Commission and the United Kingdom Atomic Energy Authority Relating

to Collaboration in the Field of Modelling of Loss of Coolant

Accidents, February 1985). The United Kingdom has consented to the

publication of this report as a USNRC document in order to allow

the widest possible circulation among the reactor safety community.

Neither the United States Government nor the United Kingdom or any

agency thereof, or any of their employees, makes any warranty,

expressed or implied, or assumes any legal liability of

responsibility for any third party's use, or the results of such

use, or any information, apparatus, product or process disclosed

in this report, or represents that its use by such third party

would not infringe privately owned rights.
PRELIMINARY ASSESSMENT OF PWR

STEAM GENERATOR MODELLING IN

RELAP5/MOD3

R J Preece and J M Putney

Addressee: Mr P Lightfoot, NE PPG

SUMMARY

A preliminary assessment of Steam Generator (SG) modelling in the PWR

thermal-hydraulic code RELAP5/MOD3 is presented. The study is based on

calculations against a series of steady-state commissioning tests carried out on the

Wolf Creek PWR over a range of load conditions. Data from the tests are used to

assess the modelling of primary to secondary side heat transfer and, in particular, to

examine the effect of reverting to the standard form of the Chen heat transfer

correlation in place of the modified form applied in RELAP5/MOD2. Comparisons

between the two versions of the code are also used to show how the new interphase

drag model in RELAP5/MOD3 affects the calculation of SG liquid inventory and the

void fraction profile in the riser.

Conclusions

1. Like RELAP51MOD2, RELAP5/MOD3 under-predicts SG heat transfer for all load

conditions examined (36% - 99%). If the code is initialised with the correct

primary side conditions, this is reflected by an under-prediction of the secondary

side pressure. The deficiency can be attributed to the inappropriate use of Chen

correlation, which was developed using data from flows in tubes and annuli, to

calculate the boiling heat transfer coefficient on the secondary side of the U-tube

bundle. The errors in SG heat transfer are worse in MOD3 because the

correlation is applied in its standard form, whereas MOD2 incorporates a

modification which enhances heat transfer as the void fraction reduces.

2. For both versions of the code, the error in secondary side pressure tends to

reduce as the reactor power reduces. The error in the MOD3 prediction is

typically around 0.6 bar higher than the MOD2 error, and is about 4 bar at full load

conditions.

3. A number of studies performed elsewhere have shown that RELAP5/MOD2

under-predicts SG secondary side liquid inventory if the downcomer level and

recirculation ratio are calculated correctly. The deficiency has been attributed to

an over-prediction of void fraction in the SG bundle, caused by an over-prediction

of the interphase drag force. The Wolf Creek calculations performed here indicate

that the new interphase drag models in RELAP5/MOD3 result in a 25% increase

in SG inventory at full load conditions. The new models also lead to changes in

the void fraction profiles predicted for the SG riser which can be'explained in

terms of the variations in SG geometry. Unfortunately, no Wolf Creek data are

available to check the MOD3 results in these areas.

iii
Task no T5401

Project no PGD 2600

Approved: 

1A L e6

Contents, classification

and distribution ................................

Date ....... ...........

iv
CONTENTS

I INTRODUCTION ..............................

2 BACKGROUND TO RELAP5/MOD3 ................

2.1 Interphase Drag Model .......................

2.2 Boiling Heat Transfer .........................

3 WOLF CREEK COMMISSIONING TESTS ............

3.1 Description of Tests ..........................

3.2 Tests Analysed .............................

4 RELAP5 MODEL ..............................

5 CODE PERFORMANCE .........................

5.1 Steam Pressure ............................. 5.2 Liquid Inventory .............................

5.3 Void Fraction ...............................

6 CONCLUSIONS ...............................

7 REFERENCES ................................

8 NOMENCLATURE ............................

1 

. . . . . . . . . °. . o.. . . . .

.. ... ....5 

................. 11

V

I INTRODUCTION

A previous Report (Putney and Preece, 1991) described an assessment of steam

generator (SG) modelling in the PWR thermal-hydraulics code RELAP5/MOD2

(Ransom. Wagner, Trapp, Johnsen, Miller. Kiser and Riemke, 1987) - which is

currently being used by Nuclear Electric for the analysis of small break LOCAs and

intact primary circuit fault transients in the Sizewell 'B' PWR. The assessment was

based on a review of code assessment calculations performed in the UK and

elsewhere, detailed calculations against a series of commissioning tests carried out

on the Wolf Creek PWR and analytical investigations of the phenomena involved in

normal and abnormal SG operation. A number of modelling deficiencies were

identified and their implications for PWR safety analysis were discussed. The most

significant deficiencies were associated with the modelling of primary to secondary

side heat transfer, secondary side liquid inventory and level behaviour. These

deficiencies were traced to errors in the calculation of boiling heat transfer and

interphase drag on the secondary side of the U-tube bundle.

In January 1990, the RELAP5/MOD3 code was released as a potential successor to

RELAP5/MOD2. Amongst other things, the new version of the code includes model

improvements in the area of interphase drag and uses a revised correlation to

represent boiling heat transfer. Changes in the predictions of PWR SG behaviour may

therefore be expected.

The present Report describes a preliminary assessment of SG modelling in

RELAP5/MOD3, based on calculations against the Wolf Creek commissioning tests

considered in the MOD2 assessment study.

The Wolf Creek PWR is one of two fully operational PWRs based on the Westinghouse

SNUPPS design, which is the basis for the Sizewell 'B' PWR. The test data used for

the present study were obtained during commissioning trials on the station involving

measurements on a specially instrumented SG operating under steady-state

conditions at various load conditions (36 - 99% full power). The calculations were

performed using RELAP5/MOD3 version 5m5 running on an IBM RS6000 work station.

2 BACKGROUND TO RELAP5/MOD3

RELAP5/MOD3 contains a number of modelling changes from the MOD2 version of

the code. Those affecting SG modelling concern the interphase drag and boiling heat

transfer models.

2.1 Interphase Drag Model

The modelling of interphase drag in RELAP5/MOD2 has been the subject of much

criticism in the past and significant changes were introduced in the MOD3 version of

the code.

The deficiencies in the MOD2 models that affect SG modelling may be summarised

as follows

1. For low flow conditions in rod/tube bundles, there is considerable experimental

and theoretical evidence to show that RELAP5/MOD2 over-predicts the

interphase drag force and thus void fraction Putney (1988; 1989b). The deficiency

arises because the code assumes that the flow regime that exists in a bundle is

identical to that in a tube having the same hydraulic diameter. As a consequence,

a slug flow model is used for the interphase drag calculation until a transition to

annular flow is predicted to occur. Observations made by Bestion (1985) at low

void fraction however, indicate a quite different flow regime for a bundle, in which

1
the vapour and liquid phases have a tendency to occupy separate flow paths.

Also, at higher void fractions, Venkateswararao, Semiat and Dukler (1982) found

that under certain circumstances, large Taylor bubbles can occupy several

subchannels. As a result, and confirmed by measurements, the interfacial drag

force in a bundle is much less than that associated with slug flow. This

fundamental deficiency in the code is exacerbated by the use of a slug flow model

which neglects the effects of profile slip.

2. The code's slug flow model is also inappropriate for the case of low flow in a

large diameter pipe (e.g. the unheated part of the SG riser), because the Taylor

bubbles are assumed to have a diameter equal to the pipe diameter. In reality

however, bubbles of this size cannot be sustained due to interfacial stability and

they disintegrate into cap bubbles (Kataoka and Ishii. 1987). Since the specific

area of a Taylor bubble is inversely proportional to its diameter, the assumption

of slug flow in this situation can lead to an under-prediction of interphase drag

when the proportion of vapour in the Taylor bubble (relative to the small bubbles)

becomes significant - and the consequences of neglecting profile slip are

outweighed. This effect is clearly demonstrated by Putney (1988; 1989b).

The review of RELAP5/MOD2 assessment calculations performed by Putney and

Preece (1991) found clear evidence to show that RELAP5/MOD2 under-predicts SG

liquid inventory under steady-state conditions, when the recirculation ratio and

downcomer level are calculated correctly. The deficiency was attributed to an

over-prediction of void fraction in the bundle region, caused by the over-prediction

of the interphase drag force. It was noted however, that the over-prediction of void

fraction in the bundle could be compensated to some extent by an under-prediction

in the unheated riser (which consists of a circular channel having a diameter of

around 3.6 m). This effect is unlikely to be present at full power conditions as annular

flow is generally predicted to exist in the unheated region (and thus deficiencies in

the slug flow model will have no effect). but could become important at lower power

conditions.

RELAP5/MOD3 includes improved interphase drag models for low flow conditions in

both rod bundles and large diameter pipes (Putney, 1989b). Both models are based

on drift flux correlations derived from experimental data appropriate to the conditions

concerned - the former on the EPRI drift flux correlation (Chexal and Lellouche,

1986), the latter on the Kataoka-lshii correlation (Kataoka and Ishii, 1987). The code

may therefore be expected to provide a better prediction of SG liquid inventory than

RELAP5/MOD3.

Other changes related to interphase drag modelling introduced in RELAP5/MOD3

include a move to a junction based drag calculation (rather than volume averaged)

(Riemke, 1989), and modifications to the annular flow transition criteria (Putney,

1989a). These changes however, are unlikely to have a significant effect on the

calculation of SG inventory.

The present study compares RELAP5/MOD3 calculations of steady-state SG

behaviour against data obtained from a series of commissioning tests carried out on

the Wolf Creek PWR. Unfortunately, these tests did not include measurements of SG

liquid inventory. However, the code's predictions of this parameter have been

compared with those obtained from RELAP5/MOD2 to show the changes resulting

from the new interphase drag models.

2.2 Boiling Heat Transfer

For saturated and subcooled boiling, RELAP5/MOD2 applies a modified form of the

Chen correlation (Chen. 1966) to calculate the wall to fluid heat transfer coefficient.

The principal modification introduced has the effect of enhancing heat transfer as the

void fraction reduces. This modification has been removed in RELAP5/MOD3.

2
For saturated boiling (i.e. saturated nucleate boiling and forced convection

vaporization), the standard form of the Chen correlation is

q"w = Hmic (Tw - Ts) + Hmac (Tw- Tf) (1)

where

0.79 0.45 049 0.24 A .75

kP C -p0.00122 f f A T sat - -sat 1

Hmic = 0.5 0.29 0.24 0.24 S (2)

a Pt hot Pg

H"mac Hmac L2 Prr. Re°"6 F (3) De

Ref-=/1 R (1 - x) G De (4)

F = 1.0 X- < 0.1

- 2.35(X- 2 + 0.213)0736 X-* > 0.1 (5)

0o.9 )0.1

x • Martinelli flow parameter (6)

S = [1 + 0.12(ReTP) '14]-l Re-rp < 32.5

= [1 + 0.52 (ReTP)°78]-1 32.5 _ ReTP < 70.0 (7)

- 0.1 ReTP > 70.0

RTP - x) G De F1.25 -4

and the rest of the notation is defined in the Nomenclature Section. Note that for

saturated boiling, Tf = T, (essentially), and thus equation (2) may be written

q"w = (Hmic + Hmac)(w - Ts) (8)

The term Hc in the Chen correlation represents the contribution to heat transfer due

to nucleate boiling, or microconvection, and is derived from the Forster-Zuber

correlation for pool boiling (Forster and Zuber, 1955). The factor S in equation (3)

accounts for a suppression of nucleate boiling under forced convection conditions

and increasing flow rate, due to a reduction in the thermal boundary layer thickness

and, consequently, a degradation of the conditions for growth of vapour bubbles. The

term H,, in the Chen correlation represents the contribution from single phase

convection, or macroconvection, and is based on the Dittus-Boelter (Dittus and

Boelter, 1930) equation with a modifying factor F to account for an enhancement of

heat transfer due to an increase in the mixture velocity.

For subcooled boiling, the Chen correlation is generally applied with the following

modifications

G De

F- 1, ReTP = / D (9)

Also, in equation (1), T, is no longer equal to T,.

The principal modification to the Chen correlation introduced in RELAP5/MOD2 is to

replace the suppression factor S by the modified suppression factor, S*:

S = 1 c<0.3 -- 1 - 2(1 - S) (a - 0. 3) 0. 3 < a :!g 0. 8 (10)

= S a>0.8

This has the effect of increasing heat transfer for void fractions below 0.8.

3
The code assessment exercise performed by Putney and Preece (1991) clearly

established that RELAP5/MOD2 under-predicts SG heat transfer under steady-state

normal operating and start-up conditions. If the code is initialised with the correct

primary side conditions, this will be reflected by an under-prediction of the secondary

side pressure. The effect is seen in both plant and rig calculations, although it tends

to reduce as both reactor power and scale reduce. The deficiency was attributed to

the inappropriate use of Chen correlation to calculate the boiling heat transfer

coefficient on the secondary side of the U-tube bundle. In particular, the correlation

was developed using data from flows in tubes and annuli. The evidence available

however, indicates that boiling heat transfer in a full size SG bundle at full load

conditions, is much greater than that in a tube having an equivalent heated diameter.

Although, as shown above, RELAP5/MOD2 incorporates a modification to this

correlation which enhances heat transfer as the void fraction reduces, the

enhancement is not sufficient for plant calculations. (It also exacerbates an incorrect

trend with bundle elevation.)

In RELAP5/MOD3, the suppression factor in the Chen correlation is applied in its

standard form. The reason for doing this is not clear, apart from a general desire by

the code developers to return all correlations to their literature form if the reasons for

modifications are not known. As a result of this change, the tendency for the code to

under-predict SG heat transfer can be expected to be worse than that associated with

RELAP5/MOD2. This effect is investigated in Section 5.1 using the Wolf Creek data.

3 WOLF CREEK COMMISSIONING TESTS

3.1 Description of Tests

In 1978 Westinghouse and the SNUPPS utilities agreed to sponsor a joint test

programme to provide detailed measurements of the thermal- hydraulic performance

characteristics of one of the first Model 'F' SGs to go into commercial service. The

Wolf Creek generating station was chosen to be the site of this test programme.

Wolf Creek is one of two fully operational PWRs based on the Westinghouse

Standardised Nuclear Unit Power Plant System (SNUPPS) design, which is the basis

for the Sizewell 'B' PWR. The plant consists of 4 loops and has a rated thermal power

of 3425 MW. The SGs are Westinghouse U-type Model 'F'. The Nuclear Steam Supply

System (NSSS) is essentially the same as Sizewell 'B', although there are some

differences in the Balance Of Plant (BOP).

During the commissioning tests measurements were concentrated on a single steam

generator which had been constructed with thirteen additional ports to take special

instrumentation for internal secondary side measurements. The data recorded

(Curlee and Preece, 1987) for each test condition represent plant measurements

averaged over 10 minutes when the plant was operating under nominal steady state

conditions. This covers loads from 23% to full load, together with a number of loads

below 10%. Unfortunately, due to feed flow instability problems, it was not possible

to achieve steady state conditions in the load range up to 23%.

Details of the measurements made in the tests and the parameters derived from them

are given by Putney and Preece (1991). Parameters relevant to the present study

include

4
Primary side reactor power

pressurizer pressure

hot and cold leg temperatures

fluid flow rate

Secondary side SG exit pressure

feedwater flow and inlet temperature

steam flow

downcomer water level

recirculation ratio

The SG exit pressure corresponds to the pressure at the inlet to the steam flow

restrictor nozzle in the upper dome of the SG, and provides an important quantity for

assessing SG heat transfer modelling in RELAP5. This parameter was obtained from

measurements of the steam pressure in the steam main outside the containment

boundary, after taking account of the steam line losses. The pressure measurements

were normally made using three equi-spaced pressure transducers. However, for

some of the earlier tests, these were supplemented by measurements from more

accurate direct-reading Heise gauges mounted close by. For the present study, the

test runs containing the Heise gauge values are the preferred ones because of the

relatively high precision obtainable (±0.1bar compared with _+0.45bar for the

pressure transducers). From standard error analysis, Putney and Preece (1991) found

that the uncertainty in the steam exit pressure was ± 0.143 % at 36.1 % reactor

power, increasing to + 0.147 % at full power.

Other parameters of interest which were derived from the plant measurements were

the primary fluid flow and the recirculation ratio. The primary fluid flow was deduced

from measurements of the hot and cold leg temperatures and the reactor power

(NSSS trend block data, process and control instrumentation). However. because the

cold leg temperature was measured downstream of the reactor coolant pump, a

temperature correction was required to account for the heat input from pumping. The

calculation of the circulation ratio, which was based on mass flow rather than

enthalpy flow, took due account of the blowdown flow.

3.2 Tests Analysed

From the datasets recorded by Curlee and Preece (1987), five test cases were

selected for detailed analysis with RELAP5/MOD3. The operating conditions covered

by these tests are summarised in Table 1. The values given in the table also include

the input values required to run RELAP5.

The tests selected for analysis were identical to those used by Putney and Preece

(1991) for their RELAP5/MOD2 assessment, and covered a wide range of load

conditions (36%-99% of full rated power). Only tests in which steam pressures were

measured using the Heise gauge were used. Very low load conditions were avoided

because of the associated control system instability, which placed in doubt the

achievement of steady-state conditions.

4 RELAP5 MODEL

For each of the tests selected, RELAP5/MOD3 was applied to simulate the

thermal-hydraulic performance of a single SG. The input decks used for the analyses

were created from the MOD2 decks used in the previous study by running the

5
automatic conversion program RICO (Ward, 1990). Additional changes were

introduced manually to set the junction hydraulic diameters equal to the hydraulic

diameters of their upstream volumes (as the default setting was not always

appropriate), and to invoke the rod bundle interphase drag model on the secondary

side of the U-tube bundle (including the U-bend). It was also noted that RICO had set

the axial boiling factors on the 'additional' heat structure boundary cards to be 0.0,

whereas the default value suggested in the input data description is 1.0. However,

when the problem was run, RELAP5/MOD3 overrode these values and set them to 0.0.

The nodalisation scheme used for the calculations is shown in Fig. 1 and

encompasses the primary side pump suction leg and hot leg elements, the secondary

side main feed and the SG steam main element. Excluded from the present model

are the main steam inlet valve, steam relief valves and the power operated relief

valve. The auxiliary feed is also omitted as is the primary side recirculation pump.

In the model, the primary side flow and hot leg temperature are represented by a

time dependent junction, TDJ, and a time dependent volume, TDV, connected to the.

steam generator inlet plenum. Pressure within the TDV is set equal to that in the inlet

plenum. The pump suction pressure is represented by a TDV connected to the steam

generator outlet plenum.

On the secondary side, the main feedwater flowrate and temperature are represented

by a TDJ and TDV2 . The main feed flow through the TDJ is varied to achieve the

desired water level in the steam generator. To do this the adjusted flow is set equal

to the specified flow in addition to a term based on the level error. The steady state

steam outlet flow is represented by a TDJ connecting the top of the steam dome to

a dummy steam main TDV, with steam flow set equal to the specified main feed flow.

With this arrangement the dummy steam main pressure has no effect on the steam

generator performance.

The required circulation ratio is achieved by adjusting the downcomer flow. The

downcomer flow is itself controlled by adjusting the friction losses at the junction

between the downcomer and riser by varying the junction flow area.

The boundary conditions for the SG model are therefore specified as follows:

1. Hot leg temperature

2. Primary flow

3. Cold leg pressure

4. Feed water temperature

5. Feed water flow

6. Desired SG narrow range water level

7. Desired downcomer flow

These may be obtained from Table I

The calculations were performed using RELAP5/MOD3 version 5m5 running on an

IBM RS6000 workstation at NPTEC.

The MOD2 deck had been derived from the two-loop Sizewell 'B' whole plant deck described by Harwood (1986). This was felt to be an acceptable approach as the geometric and material

characteristics of the Wolf Creek and Sizewell 'B' SGs are essentially identical for simulation

purposes. The deck was created by extracting the data cards for the broken loop SG and introducing appropriate boundary conditions to represent the rest of the NSSS.

The feedwater pressure is set to a nominal value, which was felt to be adequate, as this pressure only

affects the local fluid properties.

6
As with the earlier RELAP5/MOD2 calculations, the initial calculations with

RELAP5/MOD3 were made using a requested time step of 0.25s. Both the MOD2 and

MOD3 calculations showed a Courant time step of around 0.2s. However, whereas the

MOD2 time step controller obeyed this limit and reduced the calculational time step

to 0.125s, the MOD3 calculation allowed the requested time step to be used. For an

81% load test, the larger time step in the MOD3 calculation resulted in large

oscillations in the predicted values of several SG parameters - e.g. the downcomer

flow varied between 1400 and 2000 kg s-1. These oscillations were subsequently

eliminated completely by imposing a maximum (requested) time step of 0.125s.

Consequently, it was decided to force the MOD3 calculations to use the same time

step as the MOD2 calculations.

The only other difference in the manner in which the RELAP5/MOD2 and

RELAP5/MOD3 calculations were performed concerns the adjustment of flow area at

the junction between the downcomer and riser to achieve the desired circulation flow.

In the MOD2 calculations, this was done automatically using a control system. Due

to numerical difficulties, this approach was not successful in the MOD3 calculations,

and the appropriate flow area had to be determined by trial and error.

5 CODE PERFORMANCE

5.1 Steam Pressure'

The difference between the calculated steam pressure in the SG dome and that

deduced from the Wolf Creek measurements is shown in Fig. 2 as a function of

reactor power. The calculated steam pressure in this case is that evaluated at the

nodal element 612-3 in Fig. 1, which is compared with the plant value corresponding

to the inlet plane to the steam flow restrictor nozzle. Two curves are shown,

corresponding to the RELAP5/MOD3 and MOD2 predictions. In both cases, the SG

pressure is under-predicted, implying an under-prediction of primary to secondary

side heat transfer, with the error in pressure increasing monotonically with load. As

expected the use of the standard Chen heat transfer correlation in MOD3 results in a

greater under-prediction of. pressure than MOD2. At full load conditions, MOD3

under-predicts the pressure by almost 4 bar, with the error in pressure dropping to

about 1 bar at 36% load. This error is consistently about 0.6 bar greater than the

MOD2 error.

5.2 Liquid Inventory

As discussed earlier, the interphase drag model used in RELAP5/MOD2 results in an

under-prediction of SG liquid inventory under steady-state conditions, when the

recirculation ratio and downcomer level are calculated correctly. The effect of the

new interphase drag model in MOD3 can be seen in Fig. 3. which shows the ratio of

the SG liquid masses obtained from the MOD2 and MOD3 calculations. For the 73%,

81% and 99% load cases, the new interphase drag model results in an increase in

inventory of around 25%. For the 51% and 36% load cases however, there is little

change in the predicted inventory.

The trends shown in Fig. 3 can partly be explained by the fact that as the reactor

power reduces, so does the SG vapour generation rate, and thus errors in interphase

drag modelling will have an smaller impact on the predicted SG inventory. The errors

in the MOD2 interphase drag modelling in bundles are also less significant at lower

void fractions.

A more detailed examination of the RELAP5/MOD2 results reveals that, for the 36%

and 51% load cases, slug flow was predicted to occur in the unheated section of the

7
SG riser (nodal element 604-6 in Fig. 1) and the separator inlet volume (nodal

element 606). At the higher load conditions however, annular-mist flow was predicted

in both cases. In the RELAP5 model, the unheated riser section is represented as a

circular channel of length 2.9 m having a hydraulic diameter of 3.6 m; while the

separator inlet is represent as a circular channel of length 1.4 m having a hydraulic

diameter of 0.5 m. As discussed in Section 2.1, the assumption of slug flow in such

volumes could result in an under-prediction of the local void fraction, which for the

inventory calculation, would tend to compensate for the over-prediction of void

fraction in the bundle. Although this effect could also explain the trend shown in

Fig. 3, Section 5.3 shows that the void fraction profiles calculated for the SG riser do

not support this conclusion (the low vapour generation rate would appear to be the

dominant effect). However, as also discussed in Section 5.3, for the lower load

conditions, RELAP5/MOD2 does tend to under-predict void fraction in the upper part

of the heated bundle, and this could be linked to the assumption of slug flow in the

unheated riser and the use of volume averaging to calculate the interphase drag force

at a junction. Moreover, for intermediate load conditions between the 51% and 73%

load cases examined here, a cancellation of void fraction errors could well lead to a

reasonable prediction of SG inventory.

5.3 Void Fraction

The influence of the new interphase drag model in RELAP5/MOD3 is also reflected in

the void fractions predicted by the two versions of the code for SG riser. These are

shown in Fig. 4 to Fig. 8. Not only are the magnitude of the void fraction predictions

different, but the void profiles are also quite different.

For the higher load conditions. 73% and above, the voidage predicted by MOD3

becomes progressively lower than that predicted by MOD2 as the elevation above the

tube sheet increases. Moreover, at the higher elevations. MOD3 predicts a reduction

in voidage with increasing elevation, resulting in a convex-type profile, while MOD2

predicts a continual increase in voidage.

More detailed examination of the results reveals that, for the 73% load case, both

codes predict a transition from bubbly-slug3 to annular-mist flow between the 4th and

5th nodes in the riser', and for the 99% load case, the transition is predicted to take

place between nodes 3 and 4. For the 81% load case, RELAP5/MOD2 predicts a

transition between nodes 3 and 4, while RELAP5IMOD3 predicts a transition between

node 4 and 5. Interestingly, in all cases, RELAP5/MOD3 (but not RELAP5/MOD2)

predicts a transition back to bubbly-slug flow as the mixture enters the unheated

section of the riser and the hydraulic diameter increases from 1.7 cm to 3.6 m - with

a return to annular flow when the mixture enters the separator inlet region (hydraulic

diameter of 0.5 m). Physically, this implies that as the channel width increases, the

vapour confined within annular core redistributes itself into cap bubbles near the

maximum stable size (which is physically reasonable). As a result of this transition,

RELAP5/MOD3 predicts a reduction in the local void fraction, indicating that the

consequent increase in specific interfacial area (i.e interfacial area per unit volume)

causes an overall increase in interphase drag. The slight reduction in void fraction

in passing from nodes 4 to 5 in the MOD3 calculations is probably due to a reduction

in flow area in node 5 arising from the presence of the U-bend, causing an increase

in the mixture flow rate.

In RELAP5/MOD3 solution, the interphase drag model applied in the bubbly-slug regime is based on

the EPRI drift flux model in the bundle region and the Kataoka.lshii model in the unheated riser. In

the RELAPS/MOD2 solution, a slug flow model is applied in the bubbly-slug regime for all regions of

the riser.

Nodes I - 4 represent the straight section of the heated bundle, node 5 represents the U-bend and

node 6 represents the unheated section of the bundle.

8
For the 36% and 51% load cases, both codes predict slug flow tI-oughout the riser,

although RELAP5/MOD3 still predicts a reduction in voidage as the mixture enters the

unheated section of the riser (and the vapour 'bubbles' confined within the channels

expand to form cap bubbles). The void fractions predicted by each code are very

similar over the first three nodes, but then the MOD2 voidage fal~s below the MOD3

value and stays relatively constant. On the surface, the results do not appear to

indicate a significant under-prediction of void fraction by MOD2 in the unheated

section of the riser, due to the inappropriate assumption of slug flow (see Section

5.2). However, it is possible that the volume-averaged calculation of interphase drag

at each junction in MOD2 has caused this effect to be smeared over the upstream

nodes - and this could be the reason why the voidage profile above node 3 is

relatively flat.

The void fraction predictions shown in Fig. 4 to Fig. 8 are are consistent with the

trends shown in Fig. 3.

6 CONCLUSIONS

1. Like RELAP5/MOD2, RELAP5/MOD3 under-predicts SG heat t-ansfer for all load

conditions examined (36% - 99%). If the code is initialised with the correct

primary side conditions, this is reflected by an under-predicticn of the secondary

side pressure. The deficiency can be attributed to the inapprcrriate use of Chen

correlation, which was developed using data from flows in tubes and annuli, to

calculate the boiling heat transfer coefficient on the secondary side of the U-tube

bundle. The errors in SG heat transfer are worse in MOD3 because the

correlation is applied in its standard form, whereas MO02 incorporates a

modification which enhances heat transfer as the void fractior reduces.

2. For both versions of the code, the error in secondary side pressure tends to

reduce as the reactor power reduces. The. error in the MOD3 prediction is

typically around 0.6 bar higher than the MOD2 error, and is abcut 4 bar at full load

conditions.

3. A number of studies performed elsewhere have shown that RELAP5/MOD2

under-predicts SG secondary side liquid inventory if the downcomer level and

recirculation ratio are calculated correctly. The deficiency has been attributed to

an over-prediction of void fraction in the SG bundle, caused by an over-prediction

of the interphase drag force. The Wolf Creek calculations performed here indicate

that the new interphase drag models in RELAP5/MOD3 result in a 25% increase

in SG inventory at full load conditions. The new models also lead to changes in

the void fraction profiles predicted for the SG riser which can be explained in

terms of the variations in SG geometry. Unfortunately, no Wolf Creek data are

available to check the MOD3 results in these areas.

7 REFERENCES

Bestion, D., 1985, Interfacial Friction determination for the 1D-6 Equations Two Fluid

Model used in the CATHARE Code. European Two-Phase Ftow Group Meeting,

Southampton. England

Chen, J.C.. 1966, Correlation for boiling heat transfer to saturated :;uids in convective

flow, Process Design and Development, Vol. 5. pp. 322-329.

Chexal, B. and Lellouche, G.. 1986, A full-range drift-flux correlation for vertical flows

(Revision 1), EPRI NP-3989-SR.

9
Curlee, N.J. and Preece, R.J., 1987, Final Report: Model F steam generator field

instrumentation program. WCAP-1 1276-FNP, PWR/SGWG/P(86)361.

Davis, D., 1984, The Sizewell 'B' total plant dynamic model SIBDYM, NNC Report No.

PWR/R912.

Dittus, F.W. and Boelter, L.M.K., 1930, Heat transfer in automobile radiators of the

tubular type, Publications in Engineering, University of California, Berkeley, Vol. 2,

pp. 443-461.

Forster, H.K. and Zuber, N., 1955, Dynamics of vapour bubble and boiling heat

transfer, A.I.Ch.E. J., Vol. 1, No. 4, pp. 531-535.

Harwood, C., 1986, RELAP5/MOD2 input data for Sizewell 'B', GD/PE-N/1372 (Issue

1).

Kataoka, I. and Ishii, M., 1987, Drift flux model for large diameter pipe and new

correlation for pool void fraction, Int. J. Heat Mass Transfer, Vol 30, No 9, pp

1927-1939

Preece, R.J., 1987. Steady state thermal-hydraulic performance of the Westinghouse

Model 'F' steam generator from plant measurements, TPRD/L/3079/R86.

Putney, J.M., 1988, Proposals for improving interfacial drag modelling for the bubbly

and slug regimes in RELAP5, CERL Report No. RD/L/3306/R88,

PWR/HTWG/P(88)597

Putney, J.M., 1989a, An assessment of the annular flow transition criteria and

interphase friction models in RELAP5/MOD2, CERL Report No. RD/L/3451/R89,

PWR/HTWG/P(88)653

Putney, J.M., 1989b. Development of a new bubbly-slug interfacial friction model for

RELAP5 - Final Report, NP TEC Report No. ESTD/L/0075/R89, PWR/HTWG/P(89)722

Putney, J.M. and Preece, R.J., 1991, Assessment of PWR steam generator modelling

in RELAP5/MOD2, TEC/L/0471/R91, PWR/THSG/T(91)31

Ransom, V.H., Wagner, R.J., Trapp, J.A., Johnsen, G.W., Miller, C.S., Kiser, D.M. and

Riemke, R.A., 1987, RELAP5/MOD2 code manual - Volume 1: Code structure,

systems models, and solution methods, NUREG/CR-4312, EGG-2396, Rev 1.

Riemke, R.A., 1989, Junction based interphase drag and vertical stratification

modifications for RELAP5/MOD3. EGG-EAST-8580

Venkateswararao, P., Semiat, R. and Dukler, A.E., 1982, Flow pattern transition for

gas-liquid flow in a vertical rod bundle, Int. J. Multiphase Flow, Vol. 8, No. 5, pp.

509-524.

Ward, G., 1990, JEMS application: RICO (RELAP input convertor organiser)

introductory notes, RUG/P(90)45.

10
8 NOMENCLATURE

The following lists the general notation used throughout the Report. Other symbols

used are defined when they appear.

All variables relate to a ID (channel averaged) flow. SI units are assumed

throughout.

k Phase index = g(vapour)/f(liquid)

aXk Volume fraction (a -= a,)

Pk Density

hk. Specific enthalpy at saturation conditions

hgf Latent heat of vaporisation (h,. = h9, - h.1)

Tk Temperature

T. Saturation temperature

T,. Wall temperature

AT.., = T.- T.

AP,ý, Difference between saturation pressure at T. and saturation

pressure at T.

q".. Wall heat flux

Pkw Dynamic viscosity

ký Thermal conductivity

C"• Specific heat at constant pressure

Prk Prandtl number (= C,,, p• k,)

,a Surface tension

vk Velocity

jk Superficial velocity (j, = ayv.)

G Mixture mass flux (G p, jq + PiJ,

x Quality

D. Equivalent heated diameter

11
Test Run Number WC046 WC047 WC049 WCO51 WC068

Reactor Power, % 36.1 50.5 73.3 81.4 98.9

Feedwater temperature, K 452.6 467.3 484.8 489.3 496.9

Feedwater flow, kg/s 152.5 221.5 335.3 376.6 467.1

Total downcomer flow, kg/s 1554.9 1642.7 1686.6 1680.5 1684.4

Blowdown flow, kg/s 5.23 5.14 5.03 5.36 3.91

Primary hot leg 576.1 581.5 589.4 591.9 597.0

temperature, K

Primary fluid pressure, bar 154.77 154.77 155.18 155.18 155.32

Primary fluid flow, kg/s 4592.0 4675.3 4666.0 4612.8 4745.0

SG water level, % 50 50 50 50 49

Steam nozzle inlet pressure, 71.39 71.87 70.69 70.10 70.61

bar

Primary pump suction 563.6 564.6 565.4 565.1 566.1

leg temperature, K

Table 1. Wolf Creek SG Data

In the above table the effective values of the feed flow and downcomer flow are the

measured plant values reduced by the value of the blowdown flow.

12
6 -313

Mid deck plate-----

Intermediate

deck plate

Lower deck plate ------

Main feed

Primary fluid - ho

Primary fluid - pump suction leg

Figure 1 Nodalisation of the model 'F' steam generator

't leg

J MP/CPR(8.8.91)0539/1

13
Fig. 2. Errors In Calculated SG Dome Pressure

DIFFERENCE BETWEEN MEASURED AND

CALCULATED STEAM PRESSURE

IN STEAM DOME (DELTA P)

Oci 2 °rm

.......... ..... . . ......

............................. ..............

30 40 50 60 70 80 90 100

REACTOR POWER (Z)

Fig. 3. SG Liquid Mass Ratio, MOD31MOD2

CY

-j. 51

~1.

.300-

.250- ..................... .................... ....................

.Z,• . ........................................... ..................... .................. ..................... .................... ? ...................

. 5 . .................... ..................... .................... ....... ........... ..................... ..................... ...................

.050- .................... .......... ; .................... .................... A ..................... j ...................

1AAJ-4 .4.~ .~ - - S

0.g00+

30

4

40 1 1

50 60

REACTOR

70

POWER

60 90 100

(Z)

14
15
Fig. 7. Void Profiles in SG Riser at 51% Load

4 --

I. .U -.. .......... ............. ............. ! ............. ............. ............... :............. ............. -f ............. ............. f ...........

o.8o ............. ............ ............. .............. ............. ............. ............ ............. .............. ............-.: ........... 0.60....... .................... ...... 4...... ...... 4............ 4........... * i s"2

0.60- .. .2 .

OAO Mtd2 040............. .B ............. ........ ............ .................. e............. 44 ............ 4 ...........

z

0

0

0

/ I-Ber

I

I Urhecte Riser

/

0.00 0

0 2 I I I I I I

.3 4 5 6 7 8

HEIGHT ABOVE TUBESHEET (m)

9 I I

10 11

16
Fig. 8. Void Profiles in SG Riser at 36% Load

z

o OA

0.4

0

3 4 5 6 7 B

HEIGHT ABOVE TUBESHEET (m)

17
- 1,
:NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

(2-89) (Atsioned by NRC. Add Vol.. Supo.. Rev..

NRCM 1102. DHd Addendum Numbers. If Iny.)

3201.3202 BIBLIOGRAPHIC DATA SHEET

ISe# instructions on the reversel ....... NURFEG/IA-Oll3

2. TITLE AND SUBTITLE

Preliminary Assessment of PWR Steam Generator Modelling in RELAP5/MOD3 3. DATE REPORT PUBLISHED

MONTH YEAR

July 1993

4. FIN OR GRANT NUMBER

L2245

5. AUTHORIS) 6. TYPE OF REPORT

R. J. Preece and J. M. Putney Technical

7. PERIOD COVERED Inctcuu,'e Oatr

a.r~irnmr~ nuueLeIU' 4MM MIL, rL• O t IP,. rVW JV~f.L OrMit . .~............tommI. O.. .. ... .. isoftrct,~rO .... i

8. Y•PERFOMING• ORGAJ•NZ ION/, |ul-- N•AME" ANI•U ALiIUlroa) litIINR.oy OivitLllon, Office or RIVIOno Ul.S- IClear R09910 roY Commission. And Mading addres;2 it contractor, proviO name and melting addr#SLI

National Power

Technology and Environmental Centre

Kelvin Avenue, Leatherhead, Surrey

KT22 7SE, United Kingdom

9. SPONSOR ING ORGANIZATION - NAME AND ADDRESS (I1NRC. ty'pe ',•* •or oi•rr ooirNt i~, Officreor Reslro. U.S. Nu¢/r, /tegwronCommis4;on. and mali/lg ,addrezl.i

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

10. SUPPLEMENTARY NOTES

11. ABSTRACT (200 words oerssi

A preliminary assessment of Steam Generator (SG) modelling in the PWR thermal-hydraulic code RELAP5/MOD3

is presented. The study is based on calculations against a series of steady-state commissioning tests carried out on

the Wolf Creek PWR over a range of load conditions. Data from the tests are used to assess the modelling of

primary to secondary side heat transfer and, in particular, to examine the effect of reverting to the standard form of

the Chen heat transfer correlation in place of the modified form applied in RELAP5/MOD2. Comparisons between

the two versions of the code are also used to show how the new interphase drag model in RELAP5/MOD3 affects

the calculation of SG liquid inventory and the void fraction profile in the riser.

12. KEY WORDSIDESCR!PTORS (List words orphrsase rtar willAssi/st researchers In locating rhe rerort.) 13. AVAILABILITY STATEMENT

ICAP Program Unlimited

RELAP5/MOD3 14. SECURITY CLASSIFICATION

Assessment of PWR Irbl Page)

Unclassified

(this Report f

Unclassified

15. NUMBER OF PAGES

16. PRICE

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