NUREG/IA-0106

-TEC/L/0471/R91

International

Agreement Report

Assessment of PWR

Steam Generator Modelling

in RELAP5/MOD2

Prepared by

J. M. Putney, R. J. Preece

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

June 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 was prepared under an international cooperative

agreement for the exchange of technical information. Neither

the United States Government nor any agency thereof, or any of

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

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

use, or the results of such use, of any information, apparatus product or process disclosed in this report, or represents that its use

by such third party Would not infringe privately owned rights.

Available from

Superintendent of Documents

U.S. Government Printing Office

P.O. Box 37082

Washington, D.C. 20013-7082

and

National Technical Information Service

Springfield, VA 22161
NUREG/IA-0106

TEC/L/0471/R91 t ! International

Agreement Report

Assessment of PWR

Steam Generator Modelling

in RELAP5/MOD2

Prepared by

J. M. Putney, R. J. Preece

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

June 1993

Prepared as part of

The Agreement on Research Participation and Technical Exchange

under the International Thermal-HIydraulic 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.
NP Restricted TECILIO1471/1R91

PWR/THSG/T(91)31

ASSESSMENT OF PWR STEAM

GENERATOR MODELLING IN

RELAP51MOD2

J M Putney and R J Preece

Addressee: Mr P Lightfoot, Nuclear Electric, PPG

SUMMARY

An assessment of Steam Generator (SG) modelling in the PWR thermal-hydraulic

code RELAP5/MOD2 is presented. The assessment is 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 are identified and their implications for

PWR safety analysis are discussed - including methods for compensating for the

deficiencies through changes to the input deck. Consideration is also given as to

whether the deficiencies will still be present in the successor code RELAP5/MOD3.

Conclusions

1. 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 is 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. For the Sizewell 'B' SG

operating at full load conditions, the error in secondary side pressure is around

3.5 bar. The deficiency can be attributed to the application of the 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. Although 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. Various methods are available for compensating for the

deficiency, but in most cases the most defensible approach is to artificially

increase the U-tube thermal conductivity. In general, the errors in SG heat

transfer behaviour do not hold serious implications for PWR safety analysis,

although they are likely to be worse in RELAP5IMOD3 as the Chen correlation is

applied in its standard form.

2. At steady-state normal operating conditions, RELAP5/MOD2 under-predicts the

SG secondary side liquid inventory if the recirculation ratio is calculated

correctly. The deficiency can be attributed to an over-prediction of void fraction

in the SG bundle, caused by an over-prediction of the interphase drag force

(which is a known deficiency in the code). The error in inventory that results,

which tends to reduce as reactor power reduces, affects the prediction of the

downcomer level following a reactor trip and the calculated time of SG dryout in

loss of feedwater accidents. The latter holds the more serious implications for

PWR safety analysis and is particularly severe for transients in which most of the

boildown occurs under decay heat levels - as the excess inventory on the plant

iii
NP Restricted TECIL/04711R91

takes a long time to boiloff. As a result, the predicted dryout may be advanced

by around 1000 s, which has a dramatic affect on the timings of all subsequent

events. Although the initial inventory may be artificially increased by increasing

the downcomer flow area or the recirculation ratio (through raising the initial SG

level or reducing the flow losses), the adjustments required to compensate for

the void fraction errors may not be physically sensible. The initial inventory can

also be 'corrected' by specifying a very large hydraulic diameter on the

secondary side of the U-tube bundle (=250 cm). However, this may be difficult to

justify. The calculation of SG inventory is expected to be better in RELAP5/MOD3

due to the new interphase drag models.

3. For fast SG boildown transients typical of those occurring under full power

conditions, RELAP5/MOD2 tends to over-predict the rate at which the downcomer

level falls, in addition to under-predicting the initial inventory. Moreover,

increasing the initial inventory by increasing the downcomer flow area or

recirculation ratio, does not have a significant effect on this rate - it only serves

to delay the time at which SG dryout occurs. These effects arise because, in

addition to determining the initial inventory, the interphase drag models in the

bundle have a direct effect on the mixture and downcomer level trajectories

during the boildown process itself. The effect is less noticeable for slow boildown

transients where the vapour generation rate is relatively low, as the errors in

initial inventory tend to dominate. Theoretical analyses also indicate that in order

to predict the correct level behaviour during boildown, it is necessary to calculate

both the local and profile slip components of the drag force correctly. In contrast,

in order to obtain the correct initial inventory, it is only necessary to calculate the

total drag force correctly. This implies that attempts to correct the initial inventory

by specifying a large secondary side hydraulic diameter, which only affects the

calculation of the local drag force, may not necessarily produce the correct level

behaviour. It also means that this type of problem should provide a good test for

the new interphase drag models in RELAP5/MOD3.

4. During SG boildown and blowdown transients, RELAP5/MOD2 tends to

over-predict wall heat transfer above the mixture level. The evidence available

indicates that the code predicts wet wall conditions for the whole of the bundle

until the SG has virtually dried out, and that this is due to the calculation of

excessive liquid entrainment from the mixture region. It is believed that this

problem is a consequence of the known deficiencies in the bundle interphase

drag model, exacerbated by the absence of a proper vertical stratified flow model

and, possibly, the criteria applied to determine wet/dry wall conditions. Some

improvement may be expected in RELAP5/MOD3 due to the new interphase drag

and vertical stratification models.

5. The deficiencies in the bundle interphase drag model are believed to be

responsible for a tendency of RELAP5/MOD2 to calculate excessive dynamic level

swell in response to steam discharge from the SG. They may also be the cause

of instabilities seen in the calculated separator performance, although the

non-mechanistic nature of the separator model may also be a contributing factor.

In the past. a tendency to over-predict liquid entrainment to the break in steam

line break transients have been attributed to shortcoming in interphase drag

modelling. Recent studies however, have indicated that this is not the case, and

the most likely explanation is the failure to model the accumulation of liquid on

structures in the upper SG. Other deficiencies in SG modelling in the code

include the modelling of counter current flow limitation (CCFL) and liquid hold-up

in the U-tubes and inlet plena. These reflect a general problem in modelling

flooding type phenomena in RELAP5/MOD2 over a range of geometries and may

be linked to the interphase drag models and flow regime maps - although in many

cases the primary source of the deficiency lies outside the area of SG modelling.

Nevertheless, some improvement may be expected in RELAP5/MOD3, primarily

as a result of the new user controlled CCFL model.

iv
NP Restricted TEC/L/O471/R91

6. In practice, the above deficiencies in SG modelling do not diminish the value of

RELAP5/MOD2 for performing safety case calculations for the majority of PWR

LOCAs and intact primary fault sequences of interest. The only real exception

concerns the modelling of transients in which the SG boils dry very slowly, when

the under-prediction of initial inventory will result in the SG being calculated to

dry out too early. A typical example is a total loss of feedwater accident with

normal reactor trip and closure of the MSIV. Due to the inaccuracies in the timing

of major events, it may be difficult to use the results of the calculation for safety

assessment purposes.

Task No.

Project No.

T5401

PGD2600

Approved:

Contents. classification

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

Power Technologies

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

V

NP Restricted TEC/L/0471/R91

CONTENTS

1 INTRODUCTION ................................................ 1

2 PRELIMINARY CONSIDERATIONS .................................. 2

2.1 Conceptual steam generator models ............................... 2

2.2 Steam generator initialisation .................................... 2

2.3 Steam generator heat transfer .................................... 3

2.4 Chen correlation .............................................. 5

2.5 Steam generator inventory ....................................... 8

2.6 Modelling practices and their effect on heat transfer and inventory ........ 8

2.7 Measure of steam generator heat transfer .......................... 10

3 REVIEW OF RELAP5/MOD2 ASSESSMENT CALCULATIONS .............. 12

3.1 Wolf Creek commissioning tests ................................. 12

3.1.1 50% load reduction test ...................................... 13

3.1.2 Plant trip from 100% power .................................... 14

3.1.3 Steady-state tests ........................................... 14

3.2 Sizewell B design basis ........................................ 15

3.3 KNU #1 loss of offsite power transient ............................. 15

3.4 Doel 2 SGTR accident ......................................... 16

3.5 Doel 4 loss of load test ........................................ 17

3.6 Tihange 2 plant trip test ........................................ 18

3.7 ROSA-IV 5% cold leg break test (ISP-26) ........................... 18

3.8 LO FT tests .................................................. 19

3.8.1 Steam generator steady-state heat transfer ........................ 20

3.8.2 Separator performance ....................................... 21

3.8.3 LOCA tests LP-SB-01, LP-SB-02, LP-SB-03, L2-5, L3-5, L3-6 and L3-7 ..... 22

3.8.4 Intact circuit fault tests LP-FW-01, L9-3 and L9-4 .................... 23

3.9 M B-2 tests .................................................. 27

3.9.1 Loss of feedwater tests no. I and 2 .............................. 27

3.9.2 Steam line break test T-2013 ................................... 28

3.10 SPES loss of feedwater test (ISP-22) ............................. 30

3.11 LOBI-M OD2 tests ............................................ 32

3.11.1 Steam generator steady-state heat transfer ....................... 32

.3.11.2 LOCA test BL-02 ........................................... 32

3.11.3 Intact circuit fault tests ST-02 and A2-90 ......................... 33

3.11.4 Steam generator tube rupture test BL-21 ......................... 34

3.11.5 Steam line break test BT-12 ................................... 34

3.12 Steady-state analysis for Once-Through Integral System .............. 35

3.13 Sem iscale tests ............................................. 36

3.13.1 LOCA tests S-LH-1 and S-LH-2 ................................. 36

3.13.2 Steam line break test S-FS-1 .................................. 37

3.14 Studies performed at Texas A&M University ....................... 38

3.14.1 Steady-state heat transfer in a OTSG ............................ 38

3.14.2 Steady-state heat transfer in a U-tube SG ......................... 39

vii
NP Restricted TECIL/04711/R91

4 ANALYSIS OF WOLF CREEK STEADY-STATE COMMISSIONING TESTS ..... 40

4.1 Description of tests ........................................... 40

4.2 Tests analysed ............................................... 42

4.3 RELAP5/MOD2 model ......................................... 43

4.4 Base case calculations ........................................ 44

4.5 Sensitivity calculations ........................................ 44

5 ASSESSMENT OF STEAM GENERATOR MODELLING DEFICIENCIES ........ 45

5.1 Primary to secondary heat transfer under normal operating conditions .... 46

5.1.1 Code deficiencies ........................................... 46

5.1.2 M odels at fault ............................................. 48

5.1.3 Implications for PWR safety analysis ............................. 51

5.1.4 Techniques for compensating for deficiencies ....................... 53

5.1.5 RELAP5/MOD3 considerations .................................. 54

5.2 Secondary side inventory under normal operating conditions ........... 55

5.2.1 Code deficiencies ........................................... 55

5.2.2 M odels at fault ............................................. 56

5.2.3 Implications for PWR safety analysis ............................. 57

5.2.4 Techniques for compensating for deficiencies ...................... 57

5.2.5 RELAP5/MOD3 considerations .................................. 58

5.3 Level trajectories during fast boildown transients .................... 58

5.3.1 Code deficiencies and models at fault ............................ 58

5.3.2 Implications for PWR safety analysis and compensation of deficiencies .. 60

5.3.3 RELAP5/MOD3 considerations .................................. 60

5.4 Wall heat transfer above mixture level ............................. 61

5.5 Dynamic level swell ........................................... 62

5.6 Liquid carryover during steam line break transients .................. 63

5.7 Separator performance ........................................ 64

5.8 CCFL and liquid hold-up on primary side ........................... 65

5.9 Secondary side condensation effects during repressurisation transients ... 66

6 CONCLUSIONS ................................................ 66

7 REFERENCES ................................................. 68

8 NOMENCLATURE .............................................. 73

Appendix A CONCEPTUAL STEAM GENERATOR MODELS ................. 75

A.1 Primary to secondary heat transfer ............................... 75

A.2 Steady-state operation and effect of recirculation flow ................. 76

A.2.1 Conceptual m odel ............................................ 76

A.2.2 Steady-state conditions ....................................... 79

A.2.3 Effect of recirculation flow on RELAP5 separator performance ......... 79

A.2.4 Effect of recirculation flow on SG inventory ........................ 80

A.3 Steam generator boildown ..................................... 82

A.3.1 Modelling assumptions ....................................... 82

A.3.2 Vapour generation rate ........................................ 84

A.3.3 Vapour volumetric flux equation ................................ 85

viii
NP Restricted TEC/L/0471/R191

A.3.4 Void fraction equation ........................................ 85

A.3.5 Level and inventory equations ................................. 85

A.3.6 M odel solution ............................................. 86

A.3.7 Sample calculations .......................................... 87

Appendix B MEASUREMENT UNCERTAINTIES FROM COMBINED ERRORS .... 90

TABLES

Table 1. Summary of PWR plants and facilities considered in review ........ 91

Table 2. Errors in total SG heat transfer coefficient HTC-A ................ 92

Table 3. Errors in total SG heat transfer coefficient HTC-O ................ 95

Table 4. Wolf Creek SG data ....................................... 97

Table 5. SG heat transfer adjustment factors for Wolf Creek analysis ........ 98

FIGURES

Fig. 1. Sensitivity of increasing SG heat transfer ........................ 99

Fig. 2. Sensitivity of decreasing SG heat transfer ....................... 99

Fig. 3. Variation of boiling heat transfer coefficient with heated diameter .... 100

Fig. 4. Variation of boiling heat transfer coefficient with mass flux ......... 100

Fig. 5. Variation of boiling heat transfer coefficient with quality ............ 101

Fig. 6. Locations of special instrumentation in Wolf Creek steam generator .. 102

Fig. 7. Nodalisation of the model 'F' steam generator ................... 103

Fig. 8. Errors in the calculated SG dome pressure, Wolf Creek ............ 104

Fig. 9. Component pressure drop below mid-deck plate level, Wolf Creek ... 104

Fig. 10. Component pressure drop above mid-deck plate level, Wolf Creek ... 105

Fig. 11. Component pressure drop across the mid-deck plate, Wolf Creek .... 105

Fig. 12. RELAP5/MOD2 SG secondary side HTC for Wolf Creek at 99% load ... 106

Fig. 13. RELAP5/MOD2 SG secondary side HTC for Wolf Creek at 36% load . . 106

Fig. 14. Comparison of Chen correlation with data from Semiscale SG ....... 107

Fig. 15. Liquid level trajectories for simple SG boildown transient .......... 107

Fig. 16. Inventory trajectories for simple SG boildown transient ............ 108

Fig. 17. Mixture level trajectories for simple SG boildown transient ......... 108

Fig. 18. Errors in HTC-A vs SG heat flux .............................. 109

Fig. 19. Errors in HTC-O vs SG heat flux .............................. 110

Fig. 20. Errors in HTC-A vs primary to secondary temperature difference ..... 111

ix

NP Restricted TEC/L/O471/R91

I INTRODUCTION

Steam Generators (SGs) play an important role in controlling the behaviour of PWRs

under a wide range of normal and abnormal conditions. In some transients for

example, the SG acts as an excessive heat sink, while in others its heat sink capacity

may be lost completely. The SG also creates the driving force for natural circulation

in the primary loops following a failure of the reactor coolant pumps, and provides the

operators with a key heat removal option for controlled reactor cooldown during

many types of small break LOCAs. In steam generator tube rupture accidents, the

fluid conditions in the SG can have an important effect on the retention of radioactive

species.

An accurate modelling of SG behaviour is therefore an important requirement for

thermal-hydraulic computer codes used for PWR safety analysis. In the past

however, this requirement has often been circumvented by using pessimistic

modelling assumptions as a means of producing appropriate bounding solutions.

With the move towards the use of best-estimate methods for licensing purposes, and

an increased awareness of the role played by the SG in accident mitigation and

recovery, the need to validate codes in the area of SG modelling has become

increasingly evident.

The RELAP5/MOD2 code (Ransom, Wagner, Trapp, Johnsen, Miller, Kiser and

Riemke, 1987) 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. Since

its release as a 'frozen code' in April 1985, numerous assessment calculations have

been carried out in the UK and elsewhere using data from a large number of integral

and separate effects experiments. Although these studies provide a wealth of

information relevant to SG modelling in the code, this information has never been

assessed together with this objective in mind.

The purpose of the present study is to carry out an assessment of SG modelling in

RELAP5/MOD2. Particular areas of interest are SG heat transfer under steady-state

conditions and transient level behaviour. The assessment includes the following

tasks

" A review of RELAP5/MOD2 assessment calculations in which deficiencies in SG

modelling have been identified (even though they may not have been

highlighted).

" Detailed calculations of SG performance for a series of steady-state

commissioning tests carried out on the Wolf Creek PWR at various load

conditions.

" Analytical investigations of the phenomena involved in normal and abnormal SG

operation.

The objectives of the study are to identify code deficiencies in SG modelling together

with the specific models responsible, quantify the errors involved and assess the

implications for PWR safety analysis. Methods for compensating for the code

deficiencies through changes to the input deck are also discussed. Consideration is

also given as to whether the deficiencies identified will still be present in the

successor code RELAP5/MOD3, and recommendations for future code development

and assessment studies are made. These may be addressed in a subsequent study

aimed at assessing SG modelling in RELAP5/MOD3.

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NP Restricted TECILI04711R91

2 PRELIMINARY CONSIDERATIONS

2.1 Conceptual steam generator models

In order to obtain a better understanding of SG behaviour and its sensitivity to key

parameters, it is useful to construct simple analytical models of the main phenomena

involved in normal and abnormal operation. A number of such 'conceptual' models

are developed in Appendix A, and the results from these models are referred to in the

main body of this report. The particular areas addressed are

* Primary to secondary heat transfer.

* Steady-state operation and effect of recirculation flow.

* Steam generator boildown.

Although the models developed are very simple and somewhat idealistic, they are

still believed to be representative. In the context of the present study, they are

particularly useful for interpreting the results of RELAP5 analyses and for identifying

modelling deficiencies responsible for apparent code deficiencies.

2.2 Steam generator initialisation

Part of the present study is concerned with the ability of RELAP5/MOD2 to model SG

performance under steady-state, normal operating (or start-up) conditions. Before

carrying out an assessment, it is instructive to examine how many of the major SG

parameters may be specified directly to be equal to plant values, and the key phenomena which govern the calculation of the others.

To focus the discussion, consider the steady-state solution for a single SG which has

been isolated from the rest of the NSSS, using a time dependent junction (TDJ) and

volume (TDV) connected to the SG inlet plenum to set the primary loop flow rate and

hot leg temperature. A second TDV attached to the SG outlet plenum may be used

to fix the primary side pressure in the pump suction leg, which completely defines the

primary side boundary conditions. The pressure specified in the hot leg TDV only

affects the properties of the fluid entering the SG inlet plenum and may therefore be

controlled to be equal to the pressure calculated in the inlet plenum. The fluid

temperature specified in the pump suction TDV has no effect on the calculation

(provided forward flow is specified).

On the secondary side, the required feedwater temperature may be specified using

a TDV connected to the top of the downcomer. A TDV at the top of the steam dome

may be used to set outlet pressure, and this may be controlled to achieve the desired

pump suction leg temperature, thereby ensuring that the total power delivered to the

SG matches that measured on the plant. The SG can be operated over a range of

values of downcomer water level and thus, in principle', the water level in the

calculation may be controlled to match the plant value. The recirculation flow can also

be fixed at the plant value by adjusting loss coefficients in the recirculation loop.

With the above arrangement, steady-state conditions can be imposed by controlling

the feedwater flow (via a TDJ) to be equal to the calculated steam outlet flow. This

leaves three major parameters to be found by the code and which could therefore

differ from those measured on the plant: secondary side pressure, steam/feedwater

flow and liquid inventory (i.e. the total mass of liquid in the SG). Since the primary

side power and temperature are effectively fixed at plant values by the boundary

conditions, the secondary side pressure calculated by the code will correspond to the

I As discussed later, in practice, this may not always be possible.

-2-
NP Restricted TEC/L/O471/R91

value necessary to deliver this power into the secondary side. Any error in this

pressure may therefore be attributed to errors in primary to secondary side heat

transfer modelling - e.g. if the models under-predict heat transfer overall, the code

will compensate by lowering the secondary side pressure relative to the plant value.

An error in the steam/feedwater flow on the other hand. may be attributed to errors

in the vapour generation rate models. The accuracy of the liquid inventory predicted

by the code will depend on the accuracy of the liquid distribution calculation.

The above specification is often adopted when initialising a SG for whole plant

calculations, as it allows the reactor thermal power (including the pump power) and

primary side temperatures to be matched with the measured values. Clearly,

steady-state solutions could also be obtained with different sets of parameters

specified as boundary conditions. For example, the feedwater and steam flows may

be set equal to plant values and the steam outlet pressure left uncontrolled, leaving

the pump suction leg temperature to be determined by the code. In this case, the

predicted secondary side pressure will correspond to the value for which the total

steam generation rate in the heated riser is equal to the specified steam outlet flow

rate. The accuracy of this prediction will depend on the accuracies of the code's

models for both vapour generation and primary to secondary side heat transfer.

Since the calculated pressure effectively forces the total steam generation rate to be

equal to the plant value, an error in the pump suction leg temperature may then be

attributed solely to errors in the vapour generation rate models.

The important point to note is that, when initialising a SG, there are three major

parameters that cannot be specified directly to match plant values. Two of these

parameters relate to the thermal performance of the SG, and their prediction depends

on the modelling of primary to secondary side heat transfer and vapour generation.

However, since most of the SG fluid will be at saturation, it is likely that the

calculation of vapour generation will be relatively accurate. Thus, in both examples

described above, any errors in the thermal modelling of the SG are likely to be seen

mainly in the predicted secondary side pressure.

The other phenomenon where modelling deficiencies could cause a major parameter

to differ from a plant value is the calculation of the liquid distribution throughout the

SG. In the above examples, in which the downcomer level and recirculation flow were

specified, the pertinent parameter was the liquid inventory. This is the most likely*

situation, as the downcomer level and recirculation flow are generally well known on

the plant, whereas the inventory is not.

The present study includes an assessment of RELAP5/MOD2's modelling of SG

steady-state heat transfer and inventory, for which a number of analysts have

suggested that deficiencies may be present.

2.3 Steam generator heat transfer

In order to examine the calculation of SG heat transfer in RELAP5/MOD2 in more

detail and to acquire a feeling for its sensitivity to key parameters, it is useful to

construct a simple model of the processes taking place. Consider therefore, the heat

transfer across a short length of a single tube. In Appendix A.1, it is shown that by

solving the steady-state 1D heat conduction equation, the total heat transfer rate

between the primary and secondary sides, Q (Watts), may be approximated by:

27r R p 6SL (Tp - Ts) (1)

Q± 1p~ + = I- +

where

Tp primary side temperature

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NP Restricted TECILI0471/R91

Ts secondary side temperature

3L length of tube considered

R, tube radius on primary sidrp

t tube thickness

k tube thermal conductivity

H, secondary side heat transfer coefficient

H, primary side heat transfer coefficient

The principal assumptions behind this result are that the thermal conductivity is

constant across the tube and that the tube thickness is sufficiently small such that

ln(1 + t/Rp)-=t/Rp .

Equation (1) shows how various SG parameters affect the local heat transfer across

a short length of tube. Furthermore, by taking Tp and Ts to represent appropriate

average primary and secondary temperatures, and setting 6L equal to the combined

length of all tubes, it may also be used to 'characterise' the overall primary to

secondary heat transfer. The equation then shows that, for a given average primary

to secondary temperature difference, an increase in SG heat transfer may be

achieved by increasing one or all of the following parameters

* primary or secondary heat transfer coefficient

* tube conductivity

* heated length

Also, although less obvious, for a fixed primary side radius, an increase in heat

transfer may be achieved by decreasing the tube wall thickness.

Equation (1) also enables rough estimates to be obtained for the sensitivity of SG heat

transfer to changes in the individual parameters. For the Sizewell 'B' SG (Harwood,

1986)

Rp 7.724 x 10-3 m

t - 1.016 x 10- 3 m

and for normal operating conditions

k = 19Wmn-1 °K-1

In addition typical 'representative' values for the primary and secondary heat transfer

coefficients are 40 000 and 30 000, respectively. Using these parameter values to

evaluate the terms in the denominator of equation (1) gives

p t --L = 2 .9 x 10 -5 RRp + t IHs

- 2.5 x 10-5

Htp

t= 5.3 x 10-'

k

2For the Sizewell "13' PWR, In(1 + t / RP) = 0. 1235 and t / Rp = 0. 1315.

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NP Restricted TECILIO471/R91

These terms are all of a similar magnitude, which indicates that heat transfer

behaviour at normal operating conditions is not dominated by any one of the three

processes described - i.e. primary side heat transfer, secondary side heat transfer,

or tube conduction. Also, as evident from the nature of equation (1), the heat transfer

rate will not be particularly sensitive to an increase in the primary or secondary side

heat transfer coefficient on its own, or the tube wall conductance (i.e. k/t). Moreover,

successive increases in these parameters will have a diminishing effect on

increasing heat transfer. This is illustrated in Fig. 1, which shows the factor increase

in heat transfer rate for a given temperature difference that results from an increase

in the primary or secondary heat transfer coefficients, thermal conductivity or heated

length, or a decrease in the tube thickness. In particular, for the values assumed, the

maximum increase in heat transfer that could result from an increase in the

secondary side heat transfer coefficient is less than 40 %. Note also that Fig. 1 shows

that the increase in wall conductance that results from a reduction in tube wall

thickness (for a fixed primary side radius), outweighs the corresponding decrease in

heat transfer area.

In the case of a decrease in heat transfer coefficient or wall conductance, the

opposite effect occurs as the corresponding term in equation (1) quickly dominates

the heat transfer behaviour. This is clearly evident from Fig. 2, where it can be seen

that a decrease in the heat transfer rate varies in a linear manner with a decrease in

the primary or secondary heat transfer coefficient, tube conductivity or heated length,

or an increase in tube thickness.

It must be stressed that Fig. I and Fig. 2 are based on typical representative values,

and thus the quantitative results shown should only be taken as a guide to the heat

transfer sensitivity. Nevertheless, many of the qualitative effects demonstrated are

believed to be correct. In particular, both figures clearly show that the heat transfer

is most sensitive to the heated length, which indicates the importance of obtaining

an accurate estimate for this parameter3 . Also, Fig. 1 indicates that, if an increase in

the calculated heat transfer rate is desired, it is more effective to artificially increase

the thermal conductivity than either the primary or secondary side heat transfer

coefficients, although an increase in the heated length would have an even greater

effect. Fig. 2 on the other hand, while initially presented to illustrate the sensitivity

of SG heat transfer to a decrease in heat transfer coefficient or wall conductance, also

shows the situation that could arise if the primary or secondary heat transfer

coefficient degrades.

2.4 Chen correlation

One of the issues under consideration in the present study is a tendency for

RELAP5/MOD2 to under-predict SG heat transfer under steady-state, normal

operating conditions. Several analysts have attributed this deficiency to the use of the

Chen (1966) correlation (albeit modified) to represent boiling heat transfer in a tube

bundle, noting in particular that the correlation was developed using data from

experiments carried out in tubes and annuli. Thus, before carrying out a review of

RELAP5/MOD2 calculations which have exhibited this deficiency, it is useful to state

the Chen correlation and the modifications adopted in the code.

Saturated boiling

The Chen correlation may be applied to both saturated and subcooled boiling but has

different forms in each case. For saturated boiling (i.e. saturated nucleate boiling and

forced convection vaporisation), the standard form of the correlation is

q"w = Hmic (T - Ts) + Hmac (Tw - Tf) (2)

3 In RELAP5, the heated length need not be equal to the hydraulic length and, in fact. different heated

lengths may be specified independently on the primary and secondary sides.

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NP Restricted TEC/L104711R91

where

0.00122 k 0.79 C"0.45 0.49 AT0.24 ^P0.75 Hmi f -p ,-,-sat •'sat "] (3

Hmicp--"P 0.5 0.29 0.24 0.24 S (3) o" Pf h f Pi

H0.ac = [0023 -L' Pro' ROe.` ] F (4)

Mac De

Re (1- x) G De (5) Pt

F = 1.0 X- 1 < 0.1

= 2.35(Xt 2 ± 0.213)0.736 X_ > 0.1 (6)

X'1= ( x )0.9 (Pt)O.S (/ig)0.1 Martinelli flow parameter (7)

S = E1 + 0.12 (ReTp)l" 14]-1 ReTP < 32.5

= [1 + 0.52 (ReTp)°'78I- 1 32.5 _ ReTP < 70.0 (8)

= 0.1 ReTP _ 70.0

ReTP = (1 - x) G De F1.

25 X 10- 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 = (Hric + Hmac) (T, - Ts) (9)

The term Hmi 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.

In RELAP5/MOD2, the following modifications are made to the Chen correlation

(Dimenna, Larson, Johnson, Larson, Miller, Streit, Hanson and Kiser, 1988):

1. The actual value of H,,,, used in equation (2) is taken as the maximum of that

given by equation (4) and a heat transfer coefficient for forced convection to fully

developed laminar flow:

H = 4.36 Dekf (10)

For normal SG operating conditions, equation (4) will always provide the

maximum value.

2. The suppression factor S is replaced by the modified suppression factor, S':

S = 1 c<0.3

= I - 2(1 - S)(a - 0.3) 0.3 < a < 0.8 (11)

= S a >0.8

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

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NP Restricted TECILIO471/R91

3. The value of S for ReTP > 70 is set to 0.0797 rather than 0.1 (see equation (8)),

presumably to make S vary continuously with ReTP.

4. Further modifications may be made if vertical stratified flow is predicted to occur

in a volume, but these need not be considered at the present stage of the

investigation.

The code assumes that saturated boiling occurs if post-CHF conditions have not been

reached and if

T, >_ Ts and Tf > Ts - 5

0 C (12)

and deposits the heat transferred from the wall into the liquid phase (i.e. the wall is

assumed to be wet).

Subcooled boiling

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

modifications

F = 1, ReTP - G De (13)

Also, in equation (2), T, is no longer equal to T,.

In RELAP5/MOD2, the following modifications are introduced

1. F is linearly interpolated between the saturated value at zero subcooling and the

subcooled value (i.e. 1) at 5 °C subcooling, so that H..ic varies smoothly between

subcooled and saturated boiling.

2. A modified suppression factor, S', is calculated from equation (11) and then

multiplied by a 'boiling heat transfer factor' which is linearly interpolated between

0 and I over the range

[ ATonb - 2, ATOnb], or [ ATnvg - 2, ATnvg]

where ATo.b is the wall superheat required for the onset of nucleate boiling, AT,,,

is the wall superheat required for net vapour generation, and the maximum value

calculated is used4. This procedure is designed to ensure that H.,in is introduced

smoothly. However, the formulation of the boiling heat transfer factor assumes

that subcooled boiling will only take place while the void fraction is below 0.3, in

which case the modified suppression factor S' will be unity. The present authors

have observed that this is not always true in RELAP5/MOD2 calculations of

boiling heat transfer in a SG riser. Consequently, a potential discontinuity exists

in the code in the transition of H,j, between subcooled and saturated boiling, due

to the change in the effective liquid flow rate used in the calculation of the

unmodified suppression factor.

3. As for saturated boiling, further modifications may be made if vertical stratified

flow is predicted to occur in a volume, but again these need not be considered

at the present stage of the investigation.

Outside the above interpolation procedures, the code assumes that subcooled boiling

occurs if saturated conditions have not been reached and

T, > T, and Tf < T, - 5°C (14)

For T. < T,, the code assumes convective heat transfer to single phase liquid and,

for turbulent forced convection conditions, applies the Dittus-Boelter equation (i.e.

H,,., with F = 1).

4 Separate equations are used to calculate AT0,b and AT,,.

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NP Restricted TECILI0471/R91

2.5 Steam generator inventory

The previous two sections have highlighted a number of aspects relating to SG heat

transfer modelling in RELAP5/MOD2. As discussed in Section 2.2 however, another

important area where modelling deficiencies can lead to inaccurate predictions of SG

performance is the calculation of liquid distribution on the secondary side. Further,

for the particular case where the SG is initialised with a given downcomer level and

recirculation flow, such deficiencies will be manifested in an incorrect calculation of

the initial liquid inventory.

Results obtained from a number of code assessment studies that have been reported

in the literature do seem to indicate that RELAP5/MOD2 under-predicts SG inventory

under steady-state conditions. These calculations will be reviewed later. It should

be noted however, that the true inventory may not always be known accurately, and

often an inaccuracy in the calculation of initial inventory may only be inferred from

errors in the prediction of a related parameter during the course of a transient

analysis.

The calculation of SG inventory in RELAP5/MOD2 depends, primarily, on both the

code's models for interphase friction and subcooled void, and the geometric

modelling of the SG components (e.g. downcomer flow area). The latter should not

be overlooked, as the appropriate details are not always known accurately and

additional errors may be introduced when setting up a 1D representation.

Nevertheless, a number of previous studies (e.g. Putney, 1988; 1989) have shown that

RELAP5/MOD2 over-predicts interphase friction in the bubbly and slug regimes for

low flow conditions in PWR-type fuel assemblies. Evidently, if this deficiency is also

present for flows in a SG tube bundle at normal operating conditions, it will tend to

lower the inventory prediction.

It should also be recognised that the calculations of SG heat transfer and inventory

in RELAP5/MOD2 are not totally independent. For example, for a fixed SG power or

feedwater/steam flow, an under-prediction of SG heat transfer will result in a lower

secondary side pressure, which will tend to increase the voidage and thus reduce the

inventory. (An incorrect pressure distribution in the heated riser may also have an

effect). Also, since the modified Chen suppression factor (S') used in the code

reduces with increasing void fraction in the range 0.3 < cc < 0.8 (equation (11)), an

over-prediction of interphase friction may result in a lower suppression factor and

thus reduce the secondary side heat transfer coefficient. The actual situation is a little

more complex than this because the code evaluates the Chen correlation using

calculated values of volume fraction and phase velocity (rather than quality), but the

variation of S with void fraction is likely to dominate.

2.6 Modelling practices and their effect on heat transfer and inventory

Before carrying out a review of RELAP5/MOD2 calculations which have revealed

deficiencies in SG modelling, it is important to point out certain practices that are

sometimes adopted in plant modelling that can have a significant impact on code

predictions, and thus obscure the performance of the physical models.

The first point to note is that the tendency for RELAP5/MOD2 to under-predict SG heat

transfer at steady-state conditions has been known for some time, and is often

compensated for by adjusting appropriate input parameters. Thus, when reviewing

calculations with the aim of quantifying the modelling errors and establishing any

effects of scale, it is important to ascertain whether the results under consideration

were obtained using a physically correct geometric description and best-estimate

values of thermal conductivity. On a related issue, the different definitions of the

secondary side heated diameter (D,) used by analysts include hydraulic diameter,

tube OD and inter-tube gap. For the Sizewell 'B' SG, the corresponding values are

2.765, 1.748 and 0.741 cm, respectively. Fig. 3 shows how the RELAP5/MOD2 form

of the Chen correlation (Section 2.4) varies with diameter for a set of

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NP Restricted TECIL/0471/R91

thermal-hydraulic conditions 'typical' of the upper regions of the hot side of a PWR

SG at full power5 . It can be seen that as the heated diameter is reduced from the

hydraulic diameter to the inter-tube gap, the heat transfer coefficient increases by

approximately a factor of 2 (the increase is slightly more for the standard form of the

Chen correlation). Although Section 2.3 showed that the primary to secondary side

heat transfer rate is not particularly sensitive to increases in the secondary side heat

transfer coefficient, a factor of 2 increase could typically result in an increase of

around 10 % or more in the local heat transfer rate (Fig. 1), which is not insignificant.

The diameter dependence of the Chen and modified Chen correlations arises from

the suppression factor in H,ic and the Dittus-Boelter equation in H,,,m. The

Forster-Zuber equation in Hm,, and the F factor in H,,,, are both independent of

diameter (see equations (2) to (8)). Fig. 3 shows that the suppression factor is more

sensitive to diameter over the range of interest. At very small diameters however, the

Dittus-Boelter contribution, which varies as D; 0

°, increases dramatically.

Another area where input parameters for RELAP5/MOD2 calculations may be set

differently from the true plant values concerns the performance of the SG separator.

Several analysts have encountered difficulties in obtaining acceptable steady-state

conditions due to the periodic emptying and filling of this component or the

occurrence of steam carryunder into the downcomer. Such difficulties may well be

a consequence of an under-prediction of the liquid inventory, and some analysts have

overcome the problem by setting the liquid level in the downcomer artificially high,

or by reducing the friction losses in the recirculation loop. As shown in Appendix A.2,

both practices result in an increase in the recirculation flow, which has the effect of

reducing the void fraction of the fluid entering the separator, thereby reducing or

eliminating steam carryunder'. However, as also shown in Appendix A.2, an increase

in the recirculation flow will result in an increase in the liquid inventory (more so if

it is achieved by raising the SG water level), which makes it difficult to quantify

deficiencies in the models governing the inventory calculation.

An increase in the recirculation flow may also affect the calculation of the primary to

secondary heat transfer. Fig. 4 shows how an increase of 40% in recirculation

ratio 7 affects the boiling heat transfer coefficient calculation in RELAP5/MOD2, for the

typical conditions considered in Fig. 3 and a heated diameter equal to the Sizewell

'B' tube nominal OD. Evidently, the increase in mass flux results in an increase in

H,,,, (through the Dittus-Boelter contribution) and a decrease in Hmic (through the

suppression factor). For the particular conditions examined, these two effects largely

cancel and the total heat transfer coefficient is relatively constant.

In reality however, as demonstrated in Appendix A.2, an increase in mass flux will

also tend to reduce the quality and thus void fraction throughout the saturated boiling

region in the SG riser (but leave the beginning of saturated boiling largely

unchanged). Fig. 5 shows how changes in quality affect the boiling heat transfer

coefficient calculation in RELAP5/MOD2 for a fixed mass flux and the same typical

conditions considered in Fig. 4. For qualities between 0.04 and 0.2, the heat transfer

coefficient used in RELAP5 increases significantly with decreasing quality, whereas

outside this range it is largely independent of quality. This variation with quality is

evidently a consequence of the void fraction dependence of the modified suppression

factor, S', and the variation of void fraction with quality. In contrast, the standard form

5 When evaluating the modified suppression factor in Fig. 3 (equation (11)), the RELAP5/MOD2 models

for interphase and wall friction were extracted from the code and used to determine void fraction from

quality and mass flux for steady-state, fully developed flow conditions.

6 In RELAP5IMOD2, steam carryunder from the separator is modelled as a simple function of the total

liquid fraction.

7 The recirculation ratio is defined as the total downcomer flow divided by the feedwater flow, with

some allowance for the blowdown flow.

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NP Restricted TEC/L/0471/R91

of the Chen correlation, which has no direct dependence on void fraction, is relatively

unaffected by changes in quality.

It may therefore be concluded that an increase in the recirculation flow is likely to

lead to an increase in the secondary side heat transfer coefficient used by

RELAP5/MOD2.

Finally, in the initialisation procedure described in Section 2.2 it was conveniently

assumed that the code would be forced to calculate the correct recirculation flow by

adjusting appropriate pressure loss coefficients. This however, may not always be the

case, which further complicates an assessment of modelling deficiencies. Moreover,

the failure to predict the correct recirculation flow using physically correct loss

coefficients, should strictly be attributed to deficiencies in the modelling of two-phase

friction effects (and liquid distribution). However, since the two-phase loss

coefficients are not generally well known, and, apart from the recirculation ratio,

errors in the two-phase friction modelling are unlikely to have a very large influence

on predicted SG performance, this area of SG modelling will not be addressed in the

present study.

2.7 Measure of steam generator heat transfer

As mentioned earlier, one area of particular interest in the present study is the ability

of RELAP5/MOD2 to model SG heat transfer under steady-state, normal operating (or

start-up) conditions. In order to quantify modelling errors in this area and to establish

their dependence on SG conditions and scale, it is desirable to define a parameter which provides a measure of the overall primary to secondary heat transfer. This

parameter should have some physical meaning, but be simple enough to be

evaluated from the information normally reported in plant assessment studies (i.e.

measured and calculated initial values for the main plant variables). An obvious

candidate is the secondary side pressure, as most RELAP5 analyses attempt to match

calculated and measured primary side conditions at steady-state, and thus any

deficiencies in SG heat transfer modelling will manifest themselves in an error in the

predicted secondary side pressure (see Section 2.2). However, the magnitude of this

error may not provide a particularly useful measure of the underlying modelling

errors as the pressure error would be expected to reduce with reactor power (as the

primary to secondary side temperature difference reduces), making it difficult, for

example, to establish whether the modelling errors vary with local heat flux. Also, the

error in secondary side pressure takes no account of inaccuracies in the values of the

other parameters which affect SG heat transfer (e.g. cold leg temperature).

A better measure of SG heat transfer is provided by the total primary to secondary

heat transfer coefficient defined by

HTC- " - " (5 HTC - AT Tp -T (15)

where

WAh W[hf(TI,P) - hf(To,P) (16)

Nt 7r DoE Nt n DoLE

Here &" is the average SG heat flux and T, and Ts are appropriate average primary

and secondary side temperatures, respectively. Other variables in equations (15) and

(16) are defined as

W total mass flow rate on primary side of SG U-tubes (kg s-1)

T, ,T0 inlet and outlet primary side temperatures

P primary side pressure

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NP Restricted TECIL/04711R 91

h, liquid enthalpy (from steam tables)

N, total number of U-tubes

D, OD of U-tubes

E average length of U-tubes

Since most of the fluid on the secondary side of a SG U-tube bundle is close to

saturation conditions, Ts in equation (15) may be approximated by the saturation

temperature. An appropriate value for Tp is less obvious, particularly as most

RELAP5 analyses reported in the literature only quote initial values for T, and T" .

However, for the Wolf Creek calculations described in Section 4. the U-tube

length-weighted average primary side temperature is given by

TPL = co T, + (1 -- () To (17)

where a) lies between 0.324 and 0.332 for reactor powers in the range 36 to 99% load

(i.e. the temperature profile is essentially load independent). There is no guarantee

that this result applies to other plants and facilities, or even that it represents the 'effective' primary temperature that controls the SG heat transfer. Nevertheless, it is

reasonable to assume that the appropriate value of TP lies somewhere between the

arithmetic average of the inlet and outlet primary side temperatures:

TA = 0.5(Ti + TO) (18)

and the outlet temperature itself (i.e. To).

For the purposes of the present study therefore, two definitions of HTC have been

considered; one based on Tp = TA , the other based on Tp = T. . These parameters

will be referred to as HTC-A and HTC-O respectively, and the corresponding primary

to secondary temperature differences in equation (15) will be denoted by DT-A and

DT-O.

The parameter HTC just described combines the important variables which control

SG heat transfer, to provide a gross measure of the overall primary to secondary side

heat transfer effect. In particular, it accounts for the reduction in the primary to

secondary temperature difference with decreasing reactor power, and its calculation

is sensitive to inaccuracies in both the primary and secondary side fluid conditions.

Also, from the simple analysis performed in Section 2.3, the parameter would be

expected to behave as

2ir Rp Nt L (19)

Re 1 I 1 t

Rp +t I HH-H k

where the variables are defined under equation (1) (or above) and represent

appropriate average values.

An indication of the extent of any SG modelling errors in a RELAP5 analysis may

therefore be obtained by evaluating the % error in the value of HTC calculated by the

code; i.e.

100 X HTCcalc- HTCmeas (20) HTCmeas

(so that a negative error is indicative of an under-prediction of SG heat transfer). This

has been done for all RELAP5/MOD2 analyses found in the literature, which provide

sufficient information on the measured and predicted initial conditions. The results

are summarised in Table 2 (HTC-A) and Table 3 (HTC-O), along with some other

details of the calculations. These results will be referred to and discussed in

subsequent sections.

- 11 -
NP Restricted TECIL/O4711R91

One drawback of using the parameter HTC to quantify errors in SG heat transfer

modelling errors is that, because Ah and AT can be small (particularly at low power),

the uncertainty in the measured value can be large. Thus, in order to make it easier

to ascertain whether errors in the calculated values of HTC are significant, Table 2

and Table 3 also include an estimate of this uncertainty. It must be stressed

however, that these estimates are very approximate and should only really be used

as a guide to the likely data unc ertainty. They were obtained using information

available on the measurement errors for the datasets considered (where available)

and some engineering judgement. The actual procedure adopted was as follows.

* Individual uncertainties were determined for W, AT and Ah. For Ah and AT, the

uncertainties were evaluated as the RMS combination of the uncertainties in the

individual components.

* For each of W, Ah and AT alone, the corresponding uncertainty in HTC was

evaluated.

An uncertainty in HTC due to the uncertainties in both W and Ah was calculated

as the RMS combination of the uncertainties due to each parameter alone. An

uncertainty in HTC due to the uncertainties in both W and AT was calculated 'in

the same way.

* The total uncertainty in HTC was calculated as the arithmetic average, of the

above results. This was felt to provide a better indication of the true uncertainty

than the RMS combination of the uncertainties due to W, Ah and AT alone, as in

many situations, some error cancellation may be expected between the

uncertainties in Ah and AT.

3 REVIEW OF RELAP51MOD2 ASSESSMENT CALCULATIONS

This section presents a review of RELAP5/MOD2 analyses carried out in the UK and

elsewhere (mainly under ICAP or CSNI), that have revealed potential code

deficiencies in the area of SG modelling. The review considers calculations against

both full sized PWRs and scaled experimental facilities, and the deficiencies identified

are discussed for each plant/facility in turn in the order of decreasing geometric

scale. A brief description of each plant/facility is also included and a summary of the

important SG characteristics for all cases considered is given in Table 1.

3.1 Wolf Creek commissioning tests

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 four 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).

3.1.1 50% load reduction test

Keevill (1989) presents results from a RELAP5/MOD2 analysis of a Wolf Creek

commissioning test which involved a 50% load reduction from 100% power (test

SU7-0010.2). The calculation performed by Nuclear Electric Technology Division (NE

TD) using a modified version of a four-loop model developed for the Sizewell 'B' PWR.

The averaging process was applied to the minimum (i.e. negative) and maximum (i.e. positive)

uncertainties separately, so that the final values did not account for any cancellation of uncertainties

having opposite signs.

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NP Restricted TECL/LI4711R91

In initialising the SG model, the steam flow at the MSIV was adjusted to obtain the

correct cold leg temperature, and the feedwater flow was varied to match the

calculated steam flow and achieve the correct SG level (see Section 2.2). The initial

steady-state values that resulted were in good agreement with plant values on the

SG primary side (i.e. loop flow rate and hot and cold leg temperatures), but the

secondary side pressure was under-predicted by some 2.4 bar - indicating an

under-prediction of the overall primary to secondary heat transfer mechanism. This

conclusion is supported by the corresponding errors in HTC-A and HTC-O, which as

shown in Table 2 and Table 3, are equal to -9.6 and -24.4% respectively, and lie

somewhat outside the uncertainty in the data. Examination of the steam and feed

flows calculated by the code show only small differences from 'the plant values,

confirming the assertion made in Section 2.2 that the code's vapour generation rate

models are sufficiently accurate for the conditions present'.

Keevill (1989) also found that recirculation ratio calculated by RELAP5 at full load was

about 5, compared with a plant value of 3.5. The author points out that, if necessary,

the calculated recirculation ratio could be reduced by altering the loss coefficients

at the junction between the downcomer and riser. However, he also states that the

code predicts approximately the correct SG inventory and, as shown in Section 2.6,

reducing the recirculation ratio would probably result in a lower inventory prediction.

Section 2.6 also shows that reducing the recirculation ratio is likely to result in a

decrease in the heat transfer coefficient on the secondary side of the SG tubes - implying that the under-prediction of SG pressure may be worse if the recirculation

ratio was calculated correctly.

Despite the errors in SG initial conditions, Keevill (1989) found that the Wolf Creek

load reduction transient was generally well predicted by RELAP5 when the core

power, steam pressure and feed flow were specified as boundary conditions, and the

steam pressure values were normalised to the values calculated by the code for the

initial steady-state - i.e. the secondary side pressure boundary condition was set

artificially low in the transient to compensate for the under-prediction in heat transfer

at full power conditions. The main phenomena of interest were the pressuriser

pressure response during an insurge and in the presence of pressuriser spray, and

both effects were calculated correctly. However, once the system had settled down

after the initial load reduction transient, the cold and hot leg temperatures were

under-predicted by around I °C and 3 °C, respectively. Although these errors are

relatively small and could probably be explained by measurement uncertainties, they

could also suggest that the pressure error needed to compensate for the

under-prediction of SG heat transfer at full power conditions, over compensates for

the deficiencies at 50% load. This conclusion is supported by results from RELAP5

calculations against steady-state commissioning tests at Wolf Creek (see below),

which show that the pressure error tends to decrease with load (Fig. 8). It is also

interesting to note that the relative magnitudes of the errors in the predicted hot and

cold leg temperatures, are broadly consistent with the contributions that these

temperatures make to the SG length weighted average primary temperature (see

Section 2.7).

Some differences from the plant measurements were also evident in the calculated

SG liquid level during the load reduction transient, and these were attributed to

instabilities in the RELAP5 separator model and measurement uncertainties.

3.1.2 Plant trip from 100% power

NE TD (Proctor, 1990a) have recently applied RELAP5/MOD2 and a modified two-loop

model of Sizewell 'B' to analyse a commissioning test at Wolf Creek involving a plant

trip from 100% power (Test SU7-0011). As with the load reduction test, the initial

secondary side pressure was under-predicted by the code, indicating an

9 This feature was evident in all calculations reviewed and thus will not be remarked upon again.

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NP Restricted TEC/L/0471/R191

under-prediction of SG heat transfer, while the other initial conditions were calculated

correctly. For the plant trip test however, the pressure error is smaller, 1.3 bar against

2.4 bar and, as shown in Table 2 and Table 3, the corresponding errors in HTC-A and

HTC-O are also smaller by approximately a factor of 2. Interestingly, the calculated

initial conditions for both tests are virtually identical, and the apparent improvement

in SG heat transfer seen in the plant trip test occurs primarily as a result of a lower

measured SG pressure. The present authors believe that the results for the load

reduction test provide the more reliable indication of the errors in heat transfer

modelling, for the following reasons:

" Proctor (1990a) notes that the quality of the data for the plant trip test is

questionable and an uncertainty of + 1 bar in SG pressure is not unrealistic

(Proctor, 1990b private communication).

" The results for the load reduction test are broadly consistent with those obtained

in the analysis of a steady-state commissioning test at 98.89% power (see below),

where the SG pressure was measured by a Heise gauge having an accuracy of

+ 0.1 bar.

Due to the failure of part of the data acquisition system, plant measurements for the

transient phase of the plant trip test were only available from strip chart recordings

(for 30 s after the trip) and an independent data logger set up to monitor a single SG.

For the RELAP5 analysis of this period, the characteristics of the steam and feed flow

were imposed as secondary side boundary conditions, as well as the SG pressure

recorded by the strip charts. Examination of the results presented by Proctor (1990a)

shows that the strip chart pressure is typically 1.5 bar lower than the pressures

recorded by the data logger over the latter part of the transient. Furthermore, at the

time of trip, the strip chart pressure agrees with the initial steady-state pressure

calculated by the code. It is suggested that imposing the strip chart pressure as a

boundary condition may have had the effect of compensating for the under-prediction

of SG heat transfer. Although Proctor (1990a) claims that this effect was compensated

for by allowing a mismatch between the calculated steam and feed flows, this

mismatch is more likely to be due to uncertainties in other boundary conditions.

Unfortunately, due to data uncertainties and uncertainties over the response times

of the reactor inlet and outlet temperature measurements, it is difficult to draw any

conclusions regarding the heat transfer modelling from the predicted values of the

hot and cold leg temperatures. If anything, the temperatures tend to be

under-predicted, which is not inconsistent with an over-compensation of SG heat

transfer.

Overall, the primary side conditions during the plant trip transient were generally

calculated correctly by RELAP5, although the comparison was rather limited due to

the sparsity of reliable data. For the same reason, it was not possible to draw any

conclusions regarding the modelling of SG levels.

3.1.3 Steady-state tests

As part of the present study, RELAP5 has been applied to analyse a series of

steady-state commissioning tests carried out at Wolf Creek over a variety of load

conditions. These calculations are described in detail in Section 4, but the errors in

HTC are also included in Table 2 and Table 3. Again an under-prediction of SG heat

transfer is evident, with the errors in HTC-A and HTC-O tending to reduce with load.

Although the reduction is not strictly monotonic, this can probably be explained by

the uncertainties in the data.

3.2 Sizewell B design basis

Difficulties in calculating correct SG initial conditions with RELAP5/MOD2 have also

been encountered when modelling Sizewell 'B'. Hirst and Keevill (1989) for example,

report problems in obtaining the design values for both the recirculation ratio and

14-
NP Restricted TECILIO4711R91

inventory, and state that if loss coefficients are adjusted to give the correct

recirculation ratio, the inventory is low by 10%. This is not inconsistent with the

results for the Wolf Creek load reduction test, where the correct inventory was

predicted when the recirculation ratio was over-predicted. It is also consistent with

the simple theoretical analysis presented in Section 2.6 and Appendix A.2.

Hirst and Keevill (1990) on the other hand, claim that it is not possible to achieve the

design combination of SG pressure and cold leg temperature with the standard

Sizewell 'B' deck. In particular, to obtain the correct cold leg temperature, it is

necessary to adopt a lower-than-correct secondary side pressure (by 3.4 bar,

although it is not clear whether this corresponds to the calculation of the correct

recirculation ratio or inventory). The authors attribute this deficiency to the use of the

Chen correlation (albeit modified) to calculate the heat transfer coefficient on the

secondary side of the SG tubes. They also note that the value of this heat transfer

coefficient may be artificially increased by reducing the equivalent heated diameter

on the secondary side, but in order to calculate the correct SG conditions, the

diameter must be set equal to the inter-tube gap width divided by approximately 200

(as opposed to the outer diameter of the tubes in the standard deck). As

demonstrated in Sections 2.3 and 2.4, this reflects the fact that the overall primary to

secondary heat transfer rate is relatively insensitive to increases in the secondary

side heat transfer coefficient, while the Chen correlation only exhibits a strong

dependence on diameter at very small diameters (see Fig. 3).

3.3 KNU #1 loss of offsite power transient

Korea Nuclear Unit 1 (KNU #1) is a two-loop Westinghouse PWR rated at 1724 MWt

and commissioned in 1978, for which BECHTEL was the architect/engineer. The SGs

are U-type Model 51.

Chung, Kim and Lee (1990) present results from a RELAP5IMOD2 analysis of a loss

of offsite power transient that occurred at KNU #1 in June 1981, following the loss of

feedwater in one SG while the plant was operating at 77.5% reactor power. The

analysis was performed for code assessment purposes, and was carried out by the

Korea Advanced Energy Research Institute (KAERI) and Seoul National University.

The code calculations were made using a two loop model with a nodalisation

philosophy based on guide lines given in the code manual.

Before embarking on the transient analysis, the authors applied the code to calculate

steady-state values for full power operation. In order to obtain good agreement with

design values, it was necessary to artificially increase the total heat transfer area of

the SG U-tubes by 9%. This is again indicative of an under-prediction of primary to

secondary heat transfer, but as shown in Section 4, the increase in heat transfer area

required to compensate for this effect is somewhat less than that found necessary for

the Wolf Creek PWR at nominally the same SG conditions (viz. a 28% increase). Also,

the calculated SG inventory for KNU # 1 showed excellent agreement with the plant

value, whereas the Wolf Creek and Sizewell 'B' calculations reviewed above indicate

that the code under-predicts inventory if the SG level and recirculation and feed flows

are calculated correctly (as they are in the KNU R1 analysis).

Examination of the RELAP5 input deck used by Chung et al. (1990) for the KNU

calculation reveals a value for the equivalent heated diameter on the secondary side

of the SG U-tubes which is consistent with the physical hydraulic diameter. The

actual hydraulic diameter however, which is used in the wall and interfacial friction

models, is set to be consistent with the total flow area assuming a circular channel,

which results in a value of the order of 250 cm. Adopting such a large value will have

the effect of artificially reducing the interfacial friction calculated for the slug flow

regime, by reducing the interfacial area per unit volume. Evidently, this will tend to

depress the average void fraction calculated in the bundle region, which could

explain why the inventory was predicted correctly. Furthermore, the lower void

fractions will result in greater values of the modified suppression factor S" in the

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NP Restricted TEC/L/O471/R191

modified Chen heat transfer correlation, leading to larger values of the boiling heat

transfer coefficient on the secondary side of the SG tubes. This may be the reason

why only a modest increase in heat transfer area was needed to obtain the correct

primary and secondary side conditions.

With the above modelling. RELAP5/MOD2 provided accurate predictions of most plant

parameters for the KNU #1 transient. In the base case calculation, the code did not

predict the decrease in the affected SG level that occurred while the feedwater flow

was decreasing in the loss of feed phase of the transient, and spurious level peaks

were calculated at the turbine trip in both SGs. These discrepancies however, were

subsequently eliminated by introducing a 'bypass' flow path between the top of the

downcomer and steam dome. This type of modelling is standard in UK Sizewell 'B'

SG models.

3.4 Doel 2 SGTR accident

Doel 2 is a Westinghouse two-loop 1192 MWt PWR located in East Flanders, Belgium.

TRACTEBEL was the architect/engineer for the plant which was commissioned in

1975. The SGs are U-type Model 44.

In June 1979, a steam generator tube rupture accident occurred at Doel 2 while the

primary system was in the heat up phase after a 24 hour stop for repair work on the

MSIVs. Plant measurements from the transient were later made available as

OECD/CSNI International Standard Problem (ISP) 20. De Feu, Firnhaber, Pochard and

Stubbe (1988) review the results of this exercise, which included the following

calculations made using RELAP5/MOD2.

Organisation Version

TRACTEBEL (Belgium) Cycle 36.05

Technical Research Centre of Finland Cycle 36.04

Joint Research Centre - Ispra IBM Version

University of Pisa Cycle 36.04

ENEA (Italy)

An earlier calculation performed for ICAP by TRACTEBEL using RELAP5/MOD2 Cycle

36.01 is described in detail by Stubbe (1986)

Unfortunately, due to the rather limited quantity and quality of the available data, and

uncertainties in the plant initial and boundary conditions, it is difficult to draw clear

conclusions from the ISP-20 analyses regarding SG modelling in RELAP5(MOD2.

Nevertheless, two clear shortcomings were identified:

1. RELAP5 over-predicted the water level swell that occurred in the intact SG when

cooldown was initiated by opening the SRV, and also when steam was

discharged to the turbine driven auxiliary feedwater pumps. This was attributed

to the an over-prediction of interphase drag in the riser region, combined with

liquid fall back in the separator region - which resulted in excessive mixture level

swell in the riser, leading to a hold-up of water in the downcomer and the region

located at the narrow range measurement taps.

2. After the affected SG generator had been completely isolated (zero steam and

feed flows), the predicted SG pressure remained relatively constant (or

decreased slightly), whereas the measured pressure increased. During this

period, a large temperature stratification existed in the SG. In addition, water

level was high and increasing as the break flow filled the SG. The pressure

increase recorded on the plant resulted from the piston effect caused by the

rising level on the vapour in the steam dome. The failure to predict this trend

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NP Restricted TECILIO471/R91

was traced to the calculation of an excessive vapour condensation rate in the

dome, which was believed to be due to the use of constitutive equations

appropriate to the slug flow regime under stratified flow conditions. Calculations

performed by TRACTEBEL however, showed that an increase in pressure could

be calculated if the steam dome was subdivided into several volumes.

3.5 Doel 4 loss of load test

Doel 4 is a Westinghouse three-loop 3000 MWt PWR located in East Flanders.

TRACTEBEL was the architect/engineer for the plant which was first connected to the

Belgium grid in April 1985. The reactor is equipped with U-type Model E (preheater)

SGs and benefits from a high quality data acquisition system (DAS).

Stubbe, Vanhoenacker, Moeyaert and De Schutter (1988) present results from a

RELAP5/MOD2 analysis of a loss of load test at Doel 4, which was performed to verify

that the plant control systems were capable of bringing the reactor power to house

load (= 5%) without scram. Although an acceptable overall agreement was obtained

between measured and calculated data, the following errors were observed in the

predicted SG level:

" The rapid fall in the SG level" that occurred on the plant following the sudden

increase in SG pressure when the main steam line valves closed, and which was

due to steam collapse in the riser, was under-predicted by RELAP5.

* The level swell that occurred (due to bulk boiling in the SG) when the steam

dump valves opened was over-predicted by the code.

Stubbe et al. (1988) attribute the first level discrepancy to separator modelling

deficiencies and the second to excessive interphase drag in the riser region at low

void fraction, which caused the calculation of excessive water entrainment into the

separator region where the DP level measurement is located. The present authors

believe however, that the first discrepancy could also be caused by an

over-prediction of interphase drag in the riser. It should be noted that the level errors

in the calculation did not persist for very long, as the feedwater control system

reacted quickly to restore the SG level. However, due to the strong thermal coupling

between feedwater flow rate and cold leg temperature associated with pre-heater

SGs, the error in feedwater flow that resulted from the over-prediction of level swell

manifested itself in a small over-prediction of the cold leg temperature (=1 C'). This

strong thermal coupling is a particular feature of preheater type SGs and the same

error in cold leg temperature would not be expected for a U-tube SG.

3.6 Tihange 2 plant trip test

Tihange 2 is a 2785 MWt, 3 loop PWR designed by Framatome and located in the

southern part of Belgium. TRACTEBEL was the architect/engineer for the plant which

started commercial operation in June 1983. The SGs are U-type Series 51M.

As part of a series of commissioning tests, a plant trip test at 100% power was

performed at Tihange 2 in January 1983. The test was initiated by manually opening

the feedwater control valve in one SG for about 5 s, which resulted in a high water

level alarm and a consequent turbine trip and reactor scram. Two minutes later, the

operator transferred the steam dump to manual control to force a cooldown to below

the no load reference temperature.

Results from a RELAP5/MOD2 analysis of the transient performed by TRACTEBEL are

described by Rouel and Stubbe (1989). Overall, the agreement obtained between

calculated and measured parameters was generally good. Although a number of

10 Unless stated otherwise, the term SG level should be interpretted as the downcomer water level, not

the bundle mixture (or collapsed liquid) level.

- 17-
NP Restricted TECIL/0471IR91

anomalies were found, in most cases these did not necessarily point to modelling

deficiencies as the errors could be explained by uncertainties in the plant data -

which despite the presence of a high quality DAS, were relatively large compared

with the data uncertainties associated with separate or integral test facilities. This

was especially true for the secondary side modelling as, as shown by sensitivity

calculations, apparent minor changes in the timing and dynamics of the steam dump

could induce relatively large variations in the SG parameters (e.g. water level).

Nevertheless, two shortcomings in SG modelling were evident:

1. Prior to the manual operation of the steam dump, the calculated SG pressure

reached a final plateau that was very close to the measured pressure. However,

in the period immediately following the turbine and reactor trips, during which the

SG pressure was increasing rapidly, the calculated pressure was significantly

under-predicted. This behaviour was also accompanied by an abnormal rise in

the calculated water level. Investigations performed by Rouel and Stubbe (1989)

indicated that the incorrect pressure prediction was due to excessive

condensation in regions where the feedwater was injected, which had the effect

of driving the vapour to saturation and led to a temporary stagnation in the

pressure evolution. Since the plant data exhibited a similar pressure plateau but

at a higher pressure, the authors deduced that the code's bubbly flow models

cannot maintain strong thermal disequilibrium under fast pressurisation

conditions, and that due to premature condensation the vapour temperature

returns to quasi saturation conditions too quickly.

2. Following the manual opening of the steam dump valves to reduce the SG

pressure and force cooldown, a small temporary increase in the SG level was

observed on the plant narrow range level recorder, but not in the RELAP5

simulation. This is the opposite effect to that seen in the Doel calculations

described above, where an over-prediction of dynamic level swell was evident in

a number of situations. However, at the time when the steam dump was changed

to manual control in the Tihange test, the residual water levels in the SGs were

very low. As noted by Rouel and Stubbe (1989), the failure to predict a level swell

could be explained by an under-prediction of the liquid inventory in the SG,

resulting in the calculated level being below the level gauge tap, such that the

small level swell was not sufficient to enter the measurement range.

3.7 ROSA-IV 5% cold leg break test (ISP-26)

The ROSA-IV Large Scale Test Facility (LSTF) is a 1/48 volumetrically scaled, full

height model of a Westinghouse four-loop (3423 MWt) PWR located at the Japan

Atomic Energy Research Institute (JAERI). The facility is configured symmetrically

with two primary loops, and has an electrically heated core of sufficient power to

simulate typical decay heat starting from a few seconds after reactor trip (maximum

of 14% of scaled reference PWR rated power). Each loop, representing two loops in

the reference plant, has an active SG which consists of 141 full size U-tubes (the

reference single SG having a total of 3382 tubes).

Data from a 5%" cold leg break experiment with normal pump trip on ROSA-IV have

been made available as OECD/CSNI ISP 26. Cooper and Froushan (1990) present

results from RELAP5/MOD2 analysis of this problem, which was performed by NE TD

and HSD using RELAP5/MOD2 and a two-loop model provided by EG&G, Idaho.

Before calculating initial conditions for the full model, a number of steady-state

calculations were performed for a single SG generator that had been isolated from

the NSSS by specifying appropriate boundary conditions. These showed, in particular,

that there was only a restricted range of values of SG inventory for which satisfactory

steady-states could be obtained. Values outside this range either flooded the

11 Based on flow area.

- 18-
NP Restricted TECIL10471IR91

separator, giving water carryover, or gave conditions whereby steam was present in

the liquid return.

The preliminary calculations also demonstrated that it was not possible to obtain a

steady-state solution with the correct combination of SG level, collapsed liquid level

in the riser (or boiling range) and recirculation ratio. For the full plant model, the

authors adopted a compromise solution in which the SG level was matched to the

measured data, but the recirculation ratio was over-predicted and the collapsed liquid

level in the riser was under-predicted, resulting in an under-prediction of the SG

inventory. Although the recirculation ratio could be reduced slightly by increasing the

frictional loss at the junction at the bottom of the downcomer, the authors found that

this led to a reduction in the SG inventory. This type of behaviour is consistent with

steady-state calculations for Wolf Creek and Sizewell 'B' and the theoretical analysis

presented in Section 2.6 and Appendix A.2.

The initial secondary side pressures calculated by RELAP5 for ISP 26 were only

slightly lower than the measured values (0.5 and 0.8 bar for the two loops). Due to

errors in the calculated hot and cold leg temperatures though, the power removed

by each SG was also under-predicted. As a result, and as indicated by the errors in

HTC-A and HTC-O given in Table 2 and Table 3, the under-prediction of SG heat

transfer by the code was more significant than that implied by the pressure error.

Note however, that the data uncertainties for the test are quite large, in particular the

accuracy of the hot and cold leg temperature measurements were ± 3.3 *C.

The transient calculations performed by Cooper and Froushan (1990) gave a

reasonable prediction of the system thermal-hydraulic behaviour during the test. A

few discrepancies were found, but none of these were attributable directly to

deficiencies in SG modelling. However, following loop seal clearance, RELAP5

predicted a period of liquid hold-up in the SG inlet plena which was not observed in

the test. Use of a modified version of RELAP5 which allowed for the existence of

stratified flow in the inclined section of the hot leg improved the prediction, but the

time of draining was still over-predicted - leading to an under-prediction of the core

liquid inventory during the period of core boildown.

3.8 LOFT tests

LOFT is an integral effects test facility designed to simulate the major components

and system response of a commercial four-loop PWR during LOCAs and anticipated

transients. The facility has a nuclear core and two primary loops:

1. An 'intact' (active) loop with an active SG, pressuriser, and two primary coolant

pumps connected in parallel.

2. A 'broken' (test) loop with resistance and elevation changes designed to simulate

the SG and pump resistance, and two quick acting blowdown valve

assemblies"2 .

The scaling philosophy adopted in the design was to reduce the coolant volumes and

flow areas for the components by the ratio of the LOFT core power to that of a

commercial PWR (viz. 50/3400 = 1/68). This was not completely achieved however,

with the result that some of the components were oversized. Also, the vertical scaling

was not preserved with several components considerably shorter than their

counterparts in the reference plant. In particular, the SG riser section in LOFT is

approximately one-third the height of commercial SG riser, with the SG heat transfer

area nominally preserved by having six times as many U-tubes (1845 in total).

12 The SG and pump simulators were often disconnected during intact primary circuit fault experiments,

and the majority of small break tests were conducted with the break in the 'intact' loop. In both cases,

the broken loop pipework acted simply as a fluid storage reservoir.

- 19 -
NP Restricted TECILIO4711R91

Despite the shorter riser, the recirculation ratio is similar in magnitude to that in a

commercial PWR.

3.8.1 Steam generator steady-state heat transfer

Some analysts have observed an under-prediction of SG pressure when applying

RELAP5/MOD2 to calculate initial steady-state conditions for LOFT tests. For example,

NE TD found that the initial pressure for test LP-SB-03 (0.4% cold leg break) was

under-predicted by 1.1 bar (Harwood and Brown, 1990), while for test L3-7 (1" cold leg

break) KAERI obtained an under-prediction of 3.3 bar (Lee, Chung and Kim, 1990).

Other authors however, have encountered no significant difficulties in obtaining

correct initial conditions for LOFT and, as shown in Table 2 and Table 3, the errors

in HTC-A and HTC-O for all calculations reported are fairly well spread. Moreover, in

many cases, the errors lie within the experimental uncertainties, although there is a

clear bias in the direction of an under-prediction of SG heat transfer.

Trends in the HTC errors shown in Table 2 and Table 3 will be discussed later. For

the present purposes, it should be noted that the UK calculations were made with the

equivalent heated diameter on the SG secondary side set to the inter-tube gap,

whereas the KAERI and Studsvik calculations used the hydraulic diameter". Also the

predicted recirculation ratio is not known in most of the calculations considered and

the experimental value was not measured. As far as can be ascertained, all

calculations were performed using an input deck that originated from EG&G, Idaho.

However, for the UK, KAERI and INER-China analyses, the SG heated riser region was

modelled using four fluid nodes, whereas for the Studsvik calculations, it was

represented using three nodes. These calculations also used a slightly longer heated

length (by 6%).

Table 2 and Table 3 include results from two calculations made against LOFT test

L2-5 (200% cold leg break) - one carried out by KAERI (Bang, Lee and Kim, 1990), the

other by INER-China (Kao, 1988). Interestingly, the errors in HTC-A and HTC-O are

much larger for the latter calculation, despite the fact that the SG nodalisation used

in each case is virtually identical. As noted above, the equivalent heated diameter

used for the KAERI analysis was equal to the hydraulic diameter, and the SG heat

transfer area and tube thermal conductivity were also physically correct. Although the

values of these parameters used in INER-China analysis are not known, they are

unlikely to have resulted in poorer SG heat transfer (e.g. the analysts are unlikely to

have used a heated diameter greater than the hydraulic diameter). The results

obtained in the two calculations therefore suggest that there may be factors other

than those already identified, which can have an effect on SG heat transfer. Note

however, that the results from the INER-China analysis were taken from a

presentation given at an ICAP specialists meeting, and a final assessment report has

not yet been issued.

Table 2 and Table 3 include details of two calculations performed by Winfrith

Technology Centre (WTC) (Birchley, 1988) against LOFT test L9-3 (loss-of-feedwater

ATWS). In the second (Revised) calculation, a number of changes were made to the

input deck in an attempt to obtain a better prediction of the SG boildown behaviour -

principally

* The bottom nodes in the SG riser and downcomer were subdivided into two.

* The flow area in the lower part of the SG downcomer was increased in line with

engineering data.

* The flow resistance of the SG was reduced to increase the recirculation ratio.

13 The hydraulic diameter was set to 2.34 cm for the KAERI calculations and 1.92 cm for the Studsvik

calculations. This compares with an inter-tube gap size of 0.64 cm.

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NP Restricted TECIL/O471/R91

For the initial condition calculation, these changes resulted in a slight reduction in the

hot and cold leg temperatures (by 0.40C) - which, as evident from the tables, had the

effect of reducing the errors in HTC-A and HTC-O. It is suspected that this apparent

improvement in SG heat transfer was primarily a consequence of the increased

recirculation ratio.

The tables also include details of two calculations made by Studsvik against test L3-6

(2.5% cold leg break). In the second calculation (Case B), changes were made to the

riser geometry to increase the initial inventory (by 140 kg). As can be seen, this had

the effect of reducing the primary to secondary heat transfer slightly, although the

errors in HTC-A and HTC-O are still within the experimental uncertainty. This again

shows the sensitivity of these parameters to SG modelling.

3.8.2 Separator performance

Several of the UK analyses of LOFT tests have experienced difficulties in achieving

satisfactory separator performance in the initial steady-state or during the early part

of the transient (viz. from those shown in Table 2 and Table 3, tests LP-SB-03,

LP-SB-01, LP-SB-02, LP-FW-01 and L9-4 14). The particular problems encountered

were the periodic emptying and filling of the separator volume and steam carryunder

to the downcomer, which are not inconsistent with the difficulties experienced in the

ROSA-IV ISP-26 analysis discussed in Section 3.7. In the LOFT analyses, the various

solutions adopted to overcome these problems were

1. Set the SG water level artificially high.

2. Reduce the SG flow resistances.

3. Specify a very small value for the carryunder parameter VUNDER15.

It is suggested that the separator problems encountered on LOFT are a consequence

of an over-prediction of interphase drag in the tube bundle region, resulting in the

calculation of high void fraction in the separator volume and a liquid fraction less than

VUNDER. Not only will this lead to an incorrect prediction of carryunder to the

downcomer, but the presence of voidage in the downcomer could also be responsible

for the periodic emptying and filling of the separator observed in some calculations.

(For example, the onset of carryunder results in a surge of two-phase fluid from the

downcomer to the steam dome, which could lead to an increase in vapour generation

as water flashes to steam.) As discussed in Section 2.6 and Appendix A.2, increasing

the SG level or reducing the SG flow resistances has the effect of reducing the vapour

fraction entering the separator volume, thereby reducing the potential for carryunder.

Both actions however, will tend to increase the SG inventory.

Other apparent deficiencies in the RELAP5 separator model have also been observed

in some transient situations:

1. In the RELAP5 analysis of LOFT test LP-SB-03 (0.4% cold leg break), oscillations

in the predicted SG level were present throughout the transient (Harwood and

Brown, 1990). Instabilities in SG level behaviour were also observed in the

analysis of the Wolf Creek load reduction test reviewed in Section 3.1.1, and can

probably be attributed to the non-mechanistic nature of the separator model.

2. Preliminary calculations against test L9-4 (loss of offsite power, ATWS) were

found to be sensitive to the calculational time step (Keevill, 1988). This sensitivity

was eventually found to be caused by numerical oscillations in SG separator

14 The UK LOFT calculations in Table 2 and Table 3 are shown in the order that they were performed,

although the reports were not issued in this order.

Is In the RELAP5 separator model, vapour carryunder to the downcomer cannot occur until the

separator liquid fraction falls below VUNDER.

-21 -
NP Restricted TECIL/04711R91

during the first few seconds of the transient, which were subsequently removed

by changing the flow resistance of the liquid fall back junction. When this was

done the sensitivity to time step was no longer present.

3.8.3 LOCA tests LP-SB-01, LP-SB-02, LP-SB-03, L2-5, L3-5, L3-6 and L3-7

In general, although RELAP5 calculations against LOCA tests on LOFT have revealed

some deficiencies in the code, none of these can be attributed specifically to SG

modelling. This may be partly due to the fact that, for break sizes typically greater

than 5%, SGs only play a small part in LOCA transients as much of the primary

system energy is removed by the break flow, and the fact that errors in the predicted

SG behaviour do not have a very large impact on the primary side parameters

important for safety analysis. Nevertheless, a number of minor discrepancies have

been observed and these may be summarised as follows.

1. In the analysis of test LP-SB-03 (0.4% cold leg break), RELAP5 failed to predict

continued reflux condensation after the primary pressure fell below the

secondary pressure - although the code did go on to predict restoration of the

reflux cooling in the SG bleed and feed phase reasonably well. The analysts

(Harwood and Brown, 1990) suggested that the inclusions of a small node at the

bottom of the riser might have resolved this discrepancy.

2. Throughout the transient phase of test LP-SB-01 (1% hot leg break with normal

pump trip), RELAP5 gave a systematic over-prediction of the secondary side

pressure of the order of 2.5 - 5 bar. Hall and Brown (1990) suggested that this

may have arisen from a tendency to over-estimate SG heat transfer (which is the

opposite effect to that generally observed at steady-state"6 ), or from errors in

modelling heat losses or steam leakage via the main steam control valve. The

latter seems the most likely explanation, and for the subsequent analysis of test

LP-SB-02 (1% hot leg break with delayed pump trip), Hall (1987) 'tuned' the

leakage flow to obtain reasonable predictions of the secondary side pressure.

3. In the UK analysis of LOFT test L3-6 (2.5% cold leg break with delayed pump trip),

RELAP5 under-predicted the SG liquid level after the reactor had tripped and the

secondary side fluid had settled (i.e. when the recirculation flow had ceased and

the vapour generation. rate had fallen off). As noted by Scriven (1990), this

discrepancy could be caused by a number of factors:

" An over-prediction of the voidage in the SG riser at full power conditions (i.e.

when initialising the SG).

" An under-prediction of the initial recirculation ratio.

" The incorrect prediction of liquid hold-up in the separator.

" Errors in representing the variation of cross-sectional area with height in the

input deck.

4. Scriven (1990) also found that the SG secondary side pressure in L3-6 was

progressively under-predicted by RELAP5 from around 1000 s onwards, while the

predicted primary side pressure drifted above the measured value between 1000

and 2000 s. Further examination revealed that this occurred because the

calculated primary and secondary side pressures did not decouple until nearly

2000 s (at which point the secondary pressure 'held-up' above the primary

pressure), whereas in LOFT the pressures decoupled at around 1000 s. In the

RELAP5 calculation, the decoupling of the primary and secondary side pressures

was associated with the predicted draining of the SG U-tubes, which left the tubes

vapour filled and resulted in a large degradation in SG heat transfer. This led

Is As shown in Table 2 and Table 3, the initial steady-state SG heat transfer is predicted very well.

- 22 -
NP Restricted TEC/L104711R91

Scriven (1990) to conclude that the discrepancies in pressure observed were the

consequence of a late prediction of U-tube draining.

5. The SG level and secondary side pressure errors observed in the UK analysis of

L3-6 by Scriven (1990) were also seen in a calculation performed by Eriksson

(1990) at Studsvik Engineering AB, Sweden. In an attempt to improve the level

prediction, the author performed a second calculation in which the boiler

geometry was modified to be more realistic (which had the effect of increasing

the initial inventory), and the closure of the feedwater value was delayed.

Although this resulted in some improvement in the level prediction at the time

of trip, the level error then proceded to increase with time, whereas in both the

original and Scriven's calculations it remained relatively constant.

6. In the KAERI analysis of L2-5 (200% cold leg break), the predicted secondary side

pressure started to deviate from the measured value at around 40 s (which is

after the low head safety injection system had been initiated), and by the end of

the calculation (115 s) was some 3 - 4 bar lower. Bang et al. (1990) note this

discrepancy but offer no explanation for it. However, the behaviour of the

predicted secondary side pressure is similar to that observed in the L3-6

calculation, indicating that it may be explained by a delay in predicting the

draining of the SG U-tubes. The results for L2-5 presented by INER-China (Kao,

1988) do not include the secondary side pressure.

It is worth noting that the incorrect prediction of the secondary side behaviour

observed in L3-6 calculation was not seen in the analysis of the companion test L3-5,

in which the pumps were tripped early. Of the other small break LOCA tests

considered which involved a delayed pump trip (LP-SB-03 - 0.4% cold leg break and

LP-SB-02 - 1% hot leg break), the break size was much smaller and the primary and

secondary pressures did not decouple while the pumps were running.

3.8.4 Intact circuit fault tests LP-FW-01, L9-3 and 19-4

In general, SG behaviour has a more important role in determining the course of

intact circuit fault transients than LOCAs, and thus SG modelling errors can have a

more significant effect on code predictions. To date, three LOFT intact circuit fault

tests have been analysed with RELAP5/MOD2 in the UK: LP-FW-01 and L9-4 by NE

TD and L9-3 by WTC. The SG modelling deficiencies identified in these studies are

discussed below.

LP-FW-01 (complete loss-of-feedwater with recovery by primary side feed-and-bleed)

This transient was initiated by terminating the main feedwater flow to the SG and

disabling the auxiliary feed. The SG inventory then boiled down, eventually emptying

at around 90 s. Since the reactor did not trip until 48 s (on high primary side

pressure), most of this boildown occurred under full power conditions.

In the RELAP5/MOD2 analysis of the transient (Croxford, Harwood and Hall, 1988), the

SG collapsed liquid level was predicted to fall at a somewhat faster rate than the

experimental value. However, since the initial level for the calculation had been set

slightly high, the time at which the SG was predicted to empty was approximately

correct.

The secondary side pressure calculated by RELAP5 during the boildown phase of the

transient indicated a slight delay in the predicted onset of the degradation of primary

to secondary heat transfer"7 . Evidently, this behaviour could be explained both by an

over-prediction of the mixture level in the riser, and an over-prediction of liquid

entrainment above this level. As discussed below, the latter effect was evident in the

17 This delay was believed to be the cause of a small error in the predicted pressuriser level behaviour

that was present up to and just after reactor trip.

- 23 -
NP Restricted TECIL/0471/R91

RELAP5 analysis of LOFT test L9-3. Interestingly, the secondary side pressures for

LP-FW-01 appears to indicate that once SG heat degraded, the rate of degradation

was more rapid in the calculation. This behaviour was also seen in the L9-3 analysis.

No other discrepancies between the calculated and measured results for LP-FW-01

were attributed to deficiencies in SG modelling. However, the main phenomenon of

interest in the analysis was the primary side behaviour, which was relatively

insensitive to SG conditions after the latter had emptied. Nevertheless, the long term

secondary side pressure was still well predicted - although this mainly indicates that

the steam leakage from the main steam control valve (MSCV) (which was closed at

reactor trip) was well modelled.

L9-4 (loss-of-offsite-power ATWS)

LOFT test L9-4 simulated a loss-of-offsite-power accident without reactor trip. The

experiment was initiated from typical commercial PWR operating conditions by

tripping the primary coolant and main feedwater pumps, and by closing the SG MSCV.

Auxiliary feedwater was initiated after a delay of 10 s to simulate the start-up time of

the diesel generators, and the pressuriser PORV and spray were both inoperative

throughout the transient. In the UK RELAP5 analysis of the test (Keevill, 1988), the

reactor power was specified using a table of power vs time (as there was insufficient

reactor physics data to use the code's point kinetics model), and the pump coastdown

was modelled using an input table of pump velocity vs time.

A number of discrepancies between the RELAP5 calculation and the experimental

data occurred in the initial SG boildown phase of the transient. In particular, the

predicted SG level fell more rapidly than the experimental level, with the SG

emptying at 370 s against 600 s in the experiment. This resulted in a more rapid

heat-up of the primary side while the heat sink was lost (even though the calculated

heat-up started slightly later), with 25 pressuriser SRV cycles necessary to control the

pressure compared with the 7 cycles that occurred in the test. The early loss of heat

sink in the calculation also kept the primary coolant temperature high after the

boildown had finished, although the long term primary depressurisation was

predicted reasonably well.

In order to investigate whether the rapid boildown in the RELAP5 calculation was due

to insufficient water in the SG secondary side, Keevill (1988) performed a second

calculation in which the recirculation ratio was increased from 5.1 to 8. This had the

effect of increasing the initial inventory by 8%, but although the level calculation was

improved between 50 and 200 s, significant differences still remained at the beginning

and end of the boildown. It was also found that the revised inventory did not

significantly improve the primary pressure calculation. The author concluded that a

very large recirculation ratio would be needed to increase the initial inventory to the

value needed to match the experimental boildown behaviour, and suggested that the

errors in the initial inventory calculation may have arisen from errors in the void

fraction calculation in the riser region. It is interesting to note however, that although

increasing the initial inventory had the effect of delaying the time at which the SG

emptied by approximately 50 s, it did not have a significant affect of the rate of

boildown.

Keevill (1988) also suggested that errors in the input power and pump characteristics,

and the rather coarse secondary side nodalisation adopted, may have contributed to

the incorrect boildown prediction - noting, in particular, that since reactivity feedback

was not modelled explicitly, the calculation was very sensitive to the primary coolant

flow rate.

Despite the errors in SG level calculation, the initial 40 s of the L9-4 transient

containing the initial heat-up and first phase of SRV cycling, and when DNB is most

likely to occur in the core, was well predicted by RELAP5/MOD2.

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NP Restricted TECILIO4711Rg1

L9-3 (loss-of-feedwater ATWS)

LOFT test L9-3 was characterised by a rapid boildown of the SG, following the

loss-of-feedwater, with the reactor remaining close to its initial power until the SG

became sufficiently uncovered for primary to secondary heat transfer to be

significantly reduced. One important area where the RELAP5/MOD2 analysis of the

transient (Birchley, 1988) differs from that of L9-4, is that the point kinetics model was

used to calculate the reactor power, thereby ensuring that calculated reactivity and

thermal-hydraulic transients remained in step. Also, since the pumps were running

throughout the transient, the primary to secondary heat transfer was essentially

controlled by secondary side conditions, whereas in L9-4 SG heat transfer was

dominated by the low heat transfer coefficient on the primary side of the tubes until

the SG dried out.

For the first 30 s of L9-3, the SG level calculated by RELAP5 was in excellent

agreement with the data. From this point until the MSCV was closed at 67 s however,

the calculated level fell more rapidly than the measured level. In the experiment, the

SG level was just above the bottom of the indicating range at 67 s, whereas in the

calculation the SG was almost empty. The primary side heat-up that occurred as the

SG tubes became progressively uncovered was slightly delayed in the calculation,

but once initiated the primary coolant temperatures increased more rapidly.

Examination of the integrated steam flows and SG levels at the time of MSCV closure

indicated that the initial SG inventory was too low in the RELAP5 calculation. In

addition, a plot of SG heat transfer against level showed that the calculated heat

transfer remained almost constant during boildown until the SG was almost empty,

whereupon a sudden drop in heat transfer occurred. In contrast, a gradual

degradation of SG heat transfer occurred in the experiment as the level fell.

In order to investigate the effects of these discrepancies in more detail, Birchley

(1988) performed an additional calculation in which

1. The initial SG inventory was increased by increasing the downcomer flow area

(in line with engineering data) and the recirculation ratio.

2. A more gradual degradation of SG heat transfer was sought by subdividing the

bottom nodes in the downcomer and riser into two."

Surprisingly, the revised calculation gave results that were qualitatively similar to the

original calculation. The increase in initial inventory did cause the SG to boil dry

slightly later, and a delay in the onset of heat transfer degradation, but the rate of fall

in the SG level was not significantly affected. Also, the finer noding at the bottom of

the SG did not result in a smoother degradation in heat transfer as the tubes

uncovered. On the primary side, the calculated heat-up was delayed slightly due to

the higher initial inventory, but the rapidity of the heat-up was largely unchanged.

The integrated steam flow and SG level in the revised calculation indicated that the

initial SG inventory was probably still too small. Since the changes made to the model

were as large as was thought sensible, Birchley (1988) deduced that the

discrepancies observed were most likely caused by an over-prediction of the void

fraction in the riser region.

A detailed examination of the liquid fraction and heat transfer calculated for each

node in the riser showed that, prior to the closure of the MSCV, the liquid fractions

decreased more or less together, so that instead of a sequential emptying of the

nodes, the liquid was smeared along the riser. As a result, the heat transfer remained

high in all the nodes until they were nearly all empty (largely simultaneously),

whereupon the heat transfer fell sharply in all of them (which explains the rapid, but

18 Some changes to the trip setting were also introduced to match the measured conditions at actuation

more closely.

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NP Restricted TECIL/0471/R91

delayed, primary heat-up). This behaviour is indicative of an over-prediction of liquid

entrainment above the riser mixture level - although errors in the calculation of this

level could also have an effect.

3.9 MB-2 tests

Main Boiler No.2 (MB-2) represents an approximately 0.8% power scaled, full height

model of a Westinghouse Model F SG (the SG used in Sizewell 'B'), which is

geometrically and thermal-hydraulically similar in many important areas. In

particular, the tube bundle region consists of 52 Inconel 600"9 tubes arranged in a 13

x 4 rectangular array, which have the same OD, wall thickness and tube pitch as the

tubes in the full size Model F design. Some differences from the reference plant do

exist in the separator component, although the design adopted was considered to be

more representative than attempting to scale down the Model F assembly in some

way. The primary side of MB-2 acts solely as a heat source and consists of a

pressurised loop (155 bar) under forced circulation and heated by a 10 MW gas heater

- which may be augmented by a further 7 MW if the test specification requires

isothermal conditions to be maintained. At 100% scaled reference power (6.67 MWt),

the boiler produces dry, saturated steam at 69 bar, the same as Model F. The facility

is heavily instrumented and has been used to study steam generator tube rupture,

loss of feedwater and steam line break transients.

3.9.1 Loss of feedwater tests no. I and 2

MB-2 loss of feedwater (LOF) tests no. 1 and 2 were performed to investigate the preand post-reactor trip periods of a complete LOF, and to provide basic data on SG

responses to changes in power and the loss of inventory. Both tests were initiated

from 100% power conditions by shutting off the main feedwater flow. In LOF Test no.

1, the remaining boundary conditions were kept constant, while in Test no. 2 the

power was rapidly reduced to 3%.

Wu and Wang (1988) and Wang and Wu (1988) summarise results from a number of

RELAP5/MOD2 calculations against LOF tests no. 1 and 2, which were carried out at

the Institute of Nuclear Energy Research, Taiwan (INER-Taiwan). Significantly, in

order to achieve the correct steady-state initial conditions in the calculations, the SG

heat transfer area in the RELAP5 model had to be artificially increased by 21.5%. The

authors claim that with this bias, the code did 'reasonably' well in predicting SG

behaviour for the two transients, although the only result shown is the downcomer

water level response during LOF test no. 1. For the first 50 s of the transient, the

calculated level is seen to be in good agreement with the data. However, the level

then falls more rapidly, eventually indicating dryout some 20 s too early (at around

80 s). Overall, the comparison is not dissimilar to that observed for the Winfrith

analysis of LOFT LOF test L9-4 (see Section 3.8.4).

The authors also found that by refining the nodalisation used in the evaporator region,

the bias in heat transfer area necessary to achieve the correct steady-state initial

conditions could be reduced to 12.6%. However, this did not lead to any significant

improvements in the transient predictions for the major SG parameters. Refinements

to the nodalisation in other regions (such as the downcomer) and changes to the flow

junction between the dryer and drain tube, also resulted in only minor improvements

to the calculated results.

Calculations were also performed using a version of RELAP5/MOD2 in which the

Thom correlation was applied in place of the Chen correlation, to calculate the boiling

heat transfer coefficient on the secondary side of the tubes. With this version of the

code and the optimised evaporator nodalisation, the increase in heat transfer area

required to calculate correct initial conditions was only 2.7%. As before however,

19 The Sizewell 'B' SG tubes are manufactured from Inconel 690.

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NP Restricted TEC/L/04711R91

no significant improvements were evident in the predicted transient behaviour,

except that the local heat fluxes and heat transfer coefficients were calculated more

accurately.

A RELAP5/MOD2 analysis of MB-2 LOF test I carried out at Texas A&M University is

briefly described by Kalyanasundaram and Hassan (1988). The only result shown is

the bundle collapsed water level, which is systematically under-predicted during the

boildown process.

3.9.2 Steam line break test T-2013

MB-2 test T-2013 simulated a 100% steam line break transient initiated from hot zero

power conditions, with the initial SG level set for normal operating conditions. The

transient was characterised by the following phases

1. A short period lasting around 5 s in which single phase steam was discharged

through the break leading to fast two-phase level swell.

2. A period lasting from 5 to 33 s in which there was significant liquid carryover

(caused by the level rise flooding the separators and interrupting the separation

process), resulting in a two-phase mixture being discharged from the break.

3. A final period in which single phase vapour was again discharged from the break.

During the experiment, the primary side flow was maintained and the hot leg

temperature was kept approximately constant by means of gas fired heaters. A

RELAP5/MOD2 analysis of the test is described by Rogers (1989b).

No significant difficulties were encountered in setting up initial steady-state

conditions for the RELAP5 analysis, despite the existence of a very weak thermal

coupling between the primary and secondary sides and very low secondary flows..

However, the difference in temperature between the SG inlet and outlet in the

calculation was greater than the experiment, indicating an over-prediction of the

initial primary to secondary heat transfer rate. Although this is the opposite effect to

that generally observed at higher reactor power conditions, due to the very small

temperature differences and thus large uncertainties involved, it is not possible to say

for certain whether the errors were due to deficiencies in heat transfer modelling. For

this reason, the corresponding errors in HTC-A and HTC-O (267 and 433%,

respectively) have not been included in Table 2 and Table 3. The apparent

over-prediction of initial heat transfer was not expected to have any significant

implications for the transient analysis of the test, since after the break had opened the

heat transfer rate increased by nearly two orders of magnitude.

The main differences between the RELAP5 prediction and the test measurements in

each phase of the blowdown behaviour may be summarised as follows.

Phase I

The onset of two-phase flow through the break occurred 3.5 s earlier in the

calculation than the experiment (at 1.5 s against 5 s). As noted by Rogers (1989b),

this was indicative of the prediction of too rapid a level swell - a conclusion supported

by the comparison of the predicted and calculated SG DP levels. The conclusion is

also consistent with the results of plant calculations reviewed in this report, in which

RELAP5 has been found to over-predict level swell when steam is discharged from

the SG.

Phase 2

In the RELAP5 calculation, the rate of depressurisation slowed abruptly at the onset

of two-phase flow through the break. No such change was observed in the

experiment and, as a result, the depressurisation rate was under-predicted during

- 27 -
NP Restricted TECILIO471IR91

this phase. This discrepancy was traced to an under-prediction of the steam flow rate

at the break, resulting in a under-prediction of the energy removed from the system.

As a consequence, the primary to secondary heat transfer rate was under-predicted,

causing the cold leg temperature to be over-predicted by up to 80C (i.e. the primary

cooldown was under-predicted). Some uncertainty was present in the liquid flow rate

at the break but, apart from the beginning of two-phase discharge where the

uncertainty was greatest, the calculated flow rate was in excellent agreement with the

data, with the termination of liquid flow predicted to occur only one second after it

was observed.

Rogers (1989b) attributed the under-prediction of vapour flow rate to the calculation

of too much liquid carryover to the break, and suggested that this may have been

caused by an over-estimation of interphase drag by the code. However, Agar (1990)

has recently repeated the MB-2 test analysis using RELAP5/MOD2.5, which has

improved models for interphase drag in the bubbly and slug flow regimes, including

a special model for bundle geometries. Despite performing various sensitivity

calculations, no significant improvement over the MOD2 results was obtained.

The only explanation suggested by Agar (1990) for the apparent over-prediction of

liquid carryover, is that the steam dryers in the facility, which were not modelled in

the RELAP5 calculations, may have impeded the passage of liquid to the steam line,

and/or deflected liquid away from it. Although the present authors and their

colleagues do not believe that the dryers themselves could have such a significant

effect, it is suggested that liquid carryover to the break in the experiment may have

been reduced by the accumulation of liquid on other structures in the upper part of

the SG, as a result of eddies in the local vapour flow.

The similarity between the RELAP5/MOD2 and RELAP5/MOD2.5 results can probably

be explained by the fact that the liquid carryover in the calculations would have been

primarily controlled by the interphase drag at and above the two-phase mixture level.

The relevant flow regimes are therefore likely to be the annular-mist or post-dryout

regimes, for which the interphase drag models are the same in both versions of the

code. Alternatively, the flow velocities may have been so high that the carryover was

not particularly sensitive to interphase drag.

Phase 3

At the beginning of the third phase in the experiment (termination of water carryover),

water held-up in the top of the downcomer, separator or other structures drained

back into the boiler region and caused a temporary increase in the boiler collapsed

liquid level. This increase was accompanied by a more rapid fall in cold leg

temperature, and the data suggests that some of the water flowed into the top of the

boiler and rewetted the entire bundle. This insurge of water was not observed in the

calculation, although the termination of carryover did result in the collapsed liquid

level in the boiler remaining constant for a while, and caused the cold leg

temperature to stop falling.

Rogers (1989b) notes that failure of RELAP5 to predict drain back into the boiler could

be due to incorrect modelling of counter current flow limitation (CCFL) in vertical

components by RELAP5. However, he goes to suggest that it is more likely to be a

result of not modelling the deck plates in the upper separator region, which have the

capability to provide a reservoir of trapped water. This sort of explanation ties in with

the explanation suggested above for the apparent under-prediction of liquid carryover

in phase 2 of the transient. In fact, it is possible that the drain back seen in the

experiment corresponds to water accumulated in various structure in the upper SG

that was predicted to carryover to the break in the RELAP5 calculation.

The RELAP5 calculation also predicted a spurious return to two-phase discharge

during the last phase of the transient. It was subsequently found that the effect could

be prevented by specifying a large loss coefficient for the separator drain lines, which

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NP Restricted TEC/L/0471/R91

had been modelled as a junction flowing into the top of the downcomer. However, this

caused only a marginal improvement in the calculated results.

Overall agreement

In general terms, RELAP5/MOD2 calculation gave reasonable agreement with the test

data, especially the timing of key events. However, due to the above discrepancies,

the primary side cooldown rate was under-predicted.

3.10 SPES loss of feedwater test (ISP-22)

The Italian SPES integral test facility is a 1/427th power-scaled, full height model of

a Westinghouse 312 type (2775 MWt) PWR, which forms the basis for the proposed

Italian PUN reactor. The facility has an electrically heated core and simulates all

three loops of the reference plant explicitly. Each loop has an active SG which

consists of 13 full sized U-tubes (implying approximately 5550 tubes per SG at full

scale). SPES is well instrumented and is the largest PWR integral test facility

currently in operation which is capable of representing full reactor power.

SPES test SP-FW-02, performed in December 1988, simulated a complete loss of feed

from full power conditions with delayed auxiliary feed initiation to one SG. The

transient, which formed the basis of OECD/CSNI ISP-22, was characterised by the

following stages:

1. SG boildown under nominal full power conditions until closure of the MSIV and

reactor trip on low SG level.

2. SG boildown through PORVs under decay heat conditions, with associated

primary side heat-up, until SGs dry out.

3. Continued primary side heat-up and subsequent boiloff through pressuriser PORV

until core uncovery.

4. Auxiliary feed initiation to one SG on high heater rod temperature, leading to

primary cooldown and core recovery (due to pressuriser draining).

The UK RELAP5/MOD2 submissions for the problem consisted of a pre-test 'double

blind' calculation performed jointly by NE TD and HSD, and a post-test calculation

carried out by PPG. Comparisons of the results from both calculations with the

experimental data are discussed by Willmott (1990). The results illustrate a number

of difficulties in modelling SG level and inventory.

In the pre-test calculation, it was not possible to obtain correct initial values for both

the SG inventory and downcomer level. Since the inventory was deemed to be the

more important parameter, the level was set as high as practically possible (1 m

above the experimental level). This led to an initial inventory that was low by only 6%

(9.3 kg), but also resulted in a recirculation ratio approximately twice that measured

on the SPES facility.

Immediately after the reactor trip however, when the SG secondary side had settled

down and the recirculation ratio and riser voidage had reduced considerably, the

calculated level was substantially lower than the experimental level. As a result,

although the calculation correctly predicted the order and nature of the main

phenomena, they all occurred significantly earlier than in the experiment - an effect

that was exacerbated by an under-estimation of the pump heat losses. The low level

prediction was attributed to a greater loss of inventory being calculated for the

pre-trip period, caused by a delay of almost 20 s in the predicted trip time (on low SG

level). The delay itself was a consequence of an over-prediction of the SG level in the

pre-trip period, which was believed to be due to two factors:

1. Setting the initial SG level in the calculation higher than the experimental level.

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NP Restricted TECILlO471IR91

2. The fact that the calculated level was derived from the liquid fractions in the

downcomer volumes, whereas the experimental value was based on DP

measurements - which, due to frictional pressure losses, were likely to indicate

a lower level than the static value during the high flow conditions at full power.

For the post-test calculation an attempt was made to rectify the second discrepancy

by using the predicted DP values to determine the SG level. However, due to the very

high recirculation ratio that had been used in the calculation to increase the initial

inventory, this resulted in an initial SG level that was below the experimental level

at reactor trip. Thus, since the SG inventory at reactor trip was known to have a

significant impact on the rest of the transient, it was decided to trip the

post-calculation at the experimental trip time (44 s), regardless of the calculated SG

level. A number of improvements to the modelling of heat losses were also

introduced.

As expected, the use of the experimental trip time led to a better representation of the

SG level behaviour in the pre-trip period, although the level was still over-predicted

due to the reasons noted above. However, following the trip, the level again settled

to a lower value than the experimental value, with the level error approximately half

that of the pre-test calculation. The measured and experimental levels then fell slowly

at a comparable rate, but with the code predicting degradation of SG heat transfer to

occur at around 1800 s rather than 2600 s, and SG dryout at 2330 s instead of 3355 s.

These discrepancies clearly affected the timings of subsequent events.

Heat balance analyses performed by Willmott (1990) show that the early prediction

of SG dryout in the post-test calculation could be explained solely by the error in SG

inventory at MSIV closure, which had the following contributions:

* An under-prediction of 9.3 kg in the initial inventory (as noted above).

* An error in the experimental value of the initial inventory, implying that the

RELAP5 prediction was low by an additional 6.6 kg.

• Discrepancies in the modelling of the MFIV closure time and drain lines, which

had the effect of further reducing the RELAP5 inventory by 4.4 kg with respect to

the experimental inventory.

A significant part of the timing discrepancies may therefore be attributed to the

under-prediction of the initial SG inventory, which the authors attribute to an

over-prediction of void fraction in the riser region.

It is important to recognise that there are certain features in the SPES LOF transient

that exacerbate the under-prediction of the initial inventory - viz.

1. As the SG boildown occurs mostly under decay heat levels, the difference in

initial inventory between the calculation and experiment takes a long time to

boiloff, leading to significant differences in the timings of events.

2. Since the reactor trip was on low SG level, raising the SG level to increase the

initial inventory had the effect of delaying the time of trip, which further reduced

the predicted SG inventory at the time of reactor trip.

Other deficiencies in SG modelling noted in the post-test calculation were

• The primary side natural circulation model' following core recovery was

incorrectly predicted due to shortcomings in the modelling of CCFL in the SG

U-tubes.

20 The pumps were tripped during the primary side heat-phase when the subcooling in the pump suction

fell to 3.9TC.

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NP Restricted TECILIO4711R91

When the auxiliary feedwater was initiated (which occurred at different times in

the calculation and experiment), the SG level increased at a more rapid rate in

the calculation than in the experiment. The author attributes this to an

over-prediction of voidage in the riser, causing more water to accumulate in the

downcomer.

Finally, no difficulties were encountered in the post-test calculation in obtaining an

acceptable steady-state prediction for the initial SG heat transfer. As shown in

Table 2 and Table 3, the errors in HTC-A and HTC-O are both within the experimental

uncertainties, although they do indicate a slight over-prediction in heat transfer. As

noted above however, the recirculation ratio in the calculation was high by a factor

of two, and this would have had the effect of enhancing the heat transfer. It is also

worth noting that initial secondary side conditions were slightly unstable in the test,

and oscillations were present in the SG level and pressure. Consequently,

appropriate average values were used for the RELAP5 analysis.

3.11 LOBI-MOD2 tests

The LOBI-MOD2 facility, which is located at the Joint European Research Centre at

Ispra in Italy, is a 1/712th scale, full height model of the German Biblis (= 3760 MWt)

PWR. The primary circuit consists of a reactor vessel with an electrically heated core

and two loops: an 'intact' loop which simulates three loops in the reference plant, and

a 'broken' loop with pressuriser representing a single loop on the plant. Each loop

has a single active SG - the intact loop SG contains 24 inverted U-tubes and the

broken loop SG has eight tubes.

3.11.1 Steam generator steady-state heat transfer

No difficulties in obtaining acceptable SG initial conditions for LOBI tests have been

reported for RELAP5/MOD2 calculations carried out in the UK - although some

analysts have had to overcome problems associated with nonsteady experimental

conditions and data uncertainties. Scriven (1988) notes however, that in order to

achieve the correct steady-state primary to secondary temperature difference, the UK

deck was 'tuned' slightly by reducing the U-tube wall thickness and setting the

equivalent heated diameter equal of the secondary side equal to the inter-tube gap.

As explained earlier, both actions have the effect of enhancing primary to secondary

heat transfer.

The only analyses which included a comparison of predicted and measured initial

conditions were those performed by WTC for BT-12 and A2-90. The former simulated

a steam line break test from typical hot stand-by conditions (=3% power), and was

therefore characterised by very low temperature differences between both the SG

inlet and outlet plena (1°C), and the primary and secondary sides. Consequently, it

does not provide reliable data for assessing SG heat transfer modelling.

In the case of A2-90, which involved a loss of offsite and normal on site electrical

power from full load conditions, the estimated errors in HTC-A and HTC-O are shown

in Table 2 and Table 3. Both parameters indicate an over-prediction of SG heat

transfer in the intact loop and a slight under-prediction in the broken loop. However,

although the uncertainties in the test measurements are not known, the errors

indicated are not thought to be significant.

3.11.2 LOCA test BL-02

LOBI test BL-02 simulated a 3% cold leg break with normal pump trip. As seen in

other LOCA test analyses, the RELAP5/MOD2 calculation of the transient (Scriven,

1987) did not reveal any significant shortcomings in SG modelling, primarily because

the SG itself did not play a very important part in controlling the transient.

Nevertheless, the secondary side pressure response was reasonably well predicted,

although this was only achieved after particular attention had been given to the

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NP Restricted TEClLlO471/R91

modelling of the relief valve, steam leak and auxiliary feed flow actuation. The

decoupling of the primary and secondary side pressures was also well predicted,

although this discrepancy in the equivalent LOFT test was only observed when the

pump trip was delayed (i.e. L3-6. 2.5% cold leg break). There was some evidence

that the code was calculating excessive draining from the SG inlet plena, but further

investigations indicated that this was due to poor modelling of CCFL along the

horizonal part of the hot leg.

3.11.3 Intact circuit fault tests ST-02 and A2-90

ST-02 loss of feed water test

LOBI test ST-02 (subsequently renamed BT-O0) simulated a loss of main feedwater

transient leading via SG dryout to a long term cooldown using bleed and feed. A

RELAP5/MOD2 analysis of the test is described by Scriven (1988).

Unfortunately, due to numerous uncertainties in the boundary conditions and operator

actions, the test did not prove to be particularly suitable from the point of code

assessment. However, two shortcomings in the area of SG modelling were identified:

1. During the period when the SG was boiling down after reactor scram, and before

the auxiliary feedwater was tripped when the downcomer level reached 1 m, the

primary to secondary side temperature difference was over-predicted. Further

investigation into this effect revealed that the SG heat transfer coefficient was

being over-predicted by roughly a factor of 2. Scriven (1988) notes that this could

be partly due to the use of a relatively coarse noding on the secondary side, and

the fact that the code does not recognise the presence of a mixture level in a

node, so that the wall is assumed to be wet until the node is virtually empty. He

also points out however, that even when this effect is allowed for, the RELAP5

calculation still appears to over-predict the heat transfer to the secondary side,

and suggests that this might be due to the fact that the LOBI deck had been tuned

to obtain the correct primary to secondary temperature difference under

steady-state full power conditions. As noted above the modelling adopted had the

effect of enhancing heat transfer, and it is possible that this could have resulted

in an over-compensation of heat transfer for the low power, low flow conditions

prevalent during boildown.

The present authors believe that another factor contributing to the over-prediction

of heat transfer could be the calculation of excessive entrainment above the riser

mixture level, which would have had the effect of keeping the tubes wet until the

the SG was almost empty. As discussed in Section 3.8.4, this behaviour was

observed in the RELAP5 analysis of LOFT test L9-3 and possibly LP-FW-01,

although it is recognised that in these cases the boildown took place under

essentially full power conditions. However, the possibility of an over-prediction

of entrainment in ST-02 is supported by a graph presented by Scriven (1988),

which shows that the primary side temperature continues to fall gradually as the

SG mixture level crosses the top of the second lowest node - since in the

absence of entrainment and any vertical stratification modelling, RELAP5 would

be expected to predict a step rise in temperature at this point.

2. After reactor trip, the SG downcomer levels were lower in the RELAP5 calculation

than the experiment, despite initially being set higher. (Although, in the broken

loop, the initial level in the experiment was rather low and not well known due to

an instrument failure.)

A2-90 LONOP-ATWS

Whittall (1990) presents results from a RELAP5 analysis of LOBI test A2-90, which

represented a loss of offsite and normal onsite electrical power without scram.

Unfortunately, the SG boildown phase of the transient was not analysed in very much

detail, and the results shown are complicated by the fact that the initial downcomer

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NP Restricted TECIL/0471/R 91

levels were incorrectly set. Consequently, the calculation does not provide suitable

information for assessing RELAP5 SG modelling, apart from the modelling of

steady-state heat transfer at full power conditions (see above).

3.11.4 Steam generator tube rupture test BL-21

LOBI test BL-21 simulated a 0.4% .steam generator tube rupture with failure of both

the main and auxiliary feedwater systems and the high head safety injection system.

This led to core uncovery at high pressure, after which the operator opened the

pressuriser PORV to reduce the primary pressure to the accumulator set point and

so restore adequate cooling.

Results from a RELAP5 analysis of the transient are discussed by Willmott (1989). The

following shortcomings in SG modelling were noted.

1. The downcomer level in the intact SG boiled down more rapidly in the RELAP5

calculation than the experiment. (in the broken SG of course, the level soon rose

above the measurement range as the SG filled up.) Willmott (1989) suggests two

possible reasons for this discrepancy. Firstly, as was found in the ST-02 analysis

discussed above, the heat transfer to the secondary side may have been

over-predicted during the boildown process. Secondly, the set pressure of the

broken loop SG PORV during the experiment was apparently 81 bar, rather than

80 bar as was specified and used in the deck. Evidently, both effects could have

resulted in a greater proportion of the decay heat being calculated to be

transferred to the intact SG. Although errors in the riser void fraction prediction

could also have contributed to the level discrepancy, it is not possible to prove

this conclusively.

2. The calculation provided examples of situations where the code failed to model

CCFL correctly in the SG U-tubes (as for the SPES loss feedwater test), and where

excessive liquid hold-up was calculated for the U-tubes. Willmott (1989) attributes

the latter to an over-prediction of interphase drag in the vertical section of the hot

leg preventing liquid from draining back (but gives no evidence to support this

view).

3.11.5 Steam line break test BT-12

LOBI test BT-12 simulated a 100% steam line break transient initiated from hot

shutdown conditions, with the break downstream of the flow restrictor. Isolation of the

intact loop SG occurred after 4 s, and thus the blowdown of this SG had little effect

on the transient. Smethurst (1990) presents results from a RELAP5/MOD2 analysis of

the test in which sensitivity calculations were used to examine the effects of heat

losses on the intact loop SG, break discharge coefficient, water in the steam lines and

the code's nearly implicit numerics scheme.

In contrast to the analysis of MB-2 steam line break test T-2013, which was initiated

from similar power conditions, the initial conditions calculated by RELAP5 for BT-12

indicated an under-prediction of SG heat transfer (the primary side temperature

difference across the SG being calculated as 0.41°C compared with a measured value

of 1VC). Again however, due to the relatively large uncertainty associated with such

small temperature differences, these results are not believed to provide a reliable

indicator of SG heat transfer modelling in RELAP5. Consequently, the corresponding

values of HTC-A and HTC-O (-23 and -30%, respectively) have not been included

Table 2 and Table 3.

It is interesting to note that in order to obtain the correct initial inventory in the broken

loop SG for the RELAP5 calculation, the SG level had to be set artificially high (at 8.68

m instead of 8.23 m). Although this problem has been encountered during the

analysis of other test and plant transients, due to the very low power conditions in

BT-12, it is unlikely to be a consequence of an incorrect prediction of the riser void

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NP Restricted TEC/L/0471/1R91

fraction distribution. The most probable explanation are errors in representing the

SG geometry.

Qualitatively, the general trends of BT-12 transient were well predicted by RELAP5

and, in particular, the timing of key events was fairly accurate. However, the code

always over-predicted the cooldown of the primary side, which was the opposite

effect to that seen with the MB-2 steam line break analysis. For the broken loop, the

error in cold leg temperature was up to 13TC in the broken loop cold leg, although

by the end of the transient the difference had been reduced to 7°C. compared to an

overall experimental cooldown of 24°C.

Possible reasons for the 7°C discrepancy in the overall cooldown are discussed by

Smethurst (1990). The possibility that this could be due to an under-prediction of the

liquid entrained in the break flow is ruled out because, as for the MB-2 analysis, the

available evidence indicates that RELAP5 over-predicted this quantity. A possible

contribution of 1.30C from temporary variations in the auxiliary feedwater temperature

was identified, but this was clearly not enough to explain the discrepancy. The most

plausible explanation was that the metalwork in the broken loop SG contributed more

heat to the fluid than predicted, resulting in liquid being calculated to return to the

boiler and remove primary side heat, rather than being evaporated by metalwork heat

in the upper regions of the SG. The information available indicated that the mass of

SG metal in the model was under-estimated by about 15%, and the author calculated

that the corrected metalwork mass would remove the final 7°C discrepancy.

In the broken loop SG, the secondary side pressure calculated by RELAP5 agreed

reasonably well with the data for the first 100 s of the blowdown transient. After this

point however, the predicted pressure fell more slowly than the measured pressure,

until 160 s when a very rapid fall in pressure occurred as the SG dried out. This

temporary hold-up in pressure was accompanied by an over-prediction in the primary

side cooldown, and was responsible for the peak difference of 13TC between the

predicted and measured cold leg temperature. The reasons for this discrepancy are

not discussed by Smethurst (1990). However, the effects observed are similar to

those seen in some boildown transients, where the code has apparently

over-predicted the liquid entrainment above the mixture level, causing the tubes to

remain wet and provide good heat transfer until the SG has 'virtually emptied.

3.12 Steady-state analysis for Once-Through Integral System

The Once-Through Integral System (OTIS) test facility was designed to simulate

conditions in the reactor coolant system and steam generator of a raised-loop B&W

PWR, during the natural circulation phases of a small break LOCA. It consists of a

one-by-one (one hot leg, one cold leg) electrically heated loop in which key

components of the reference plant are represented at 1/1632 scale, and is capable

of supplying scaled powers ranging from 1 to 5% of full power. The loop has a single

19-tube once-through steam generator (OTSG).

Note that all other plant and facilities examined in the present review have U-tube

SGs, as does Sizewell 'B'. The reason for including results from the OTIS facility is

to allow a comparison to be made of steady-state heat transfer characteristics of

U-tube and once-through SGs, so that the modelling requirements for each design

may be established.

Dimenna, Hall, Kullberg, Larson, Larson and Watkins (1989) present results from a

RELAP5 analysis of a loss of feedwater transient performed on the OTIS, in which the

initial conditions corresponded to plant conditions at approximately 1.5 minutes after

reactor trip and pump coastdown, with auxiliary feedwater activated. In the RELAP5

calculation, the initial SG pressure was under-predicted by 3 bar (from 80 bar), and

the initial core AT was over-predicted by 2TC. As shown in Table 2 and Table 3, the

corresponding errors in HTC-A and HTC-O are -10 and -52.5% respectively, indicating

an under-prediction of SG heat transfer, although the uncertainties in the test

measurements are not known.

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INP Restricted TEC/L/0471/R91

It is pertinent to note that difference between the hot and cold leg temperatures in the

OTIS test was 22°C, whereas the measured SG heat flux was only 0.5 W cm-2 . This

temperature difference is significantly greater than that seen in experiments carried

out in U-tube SGs at similar heat fluxes (viz. the steam line break tests review in this

report). For example, for LOBI test BT-12. the initial SG heat flux was also 0.5

Wcm-2 but the difference between the hot and cold leg temperatures was only 1°C.

Also, in MB-2 test T-2013 and Semiscale test S-FS-1, the initial heat flux was 0.3

W cm-2 and the temperature differences were 0.6 and 0.7°C, respectively. 21 Although

this could indicate that SG heat transfer is better in a U-tube SG than in a OTSG,

implying that different models may be needed in each case, it should be remembered

that the mixture level is above the tube bundle in a U-tube SG but approximately half

way up in a OTSG. Consequently, the actual heat flux to the mixture region is greater

than that quoted in the tables by roughly a factor of 2 22. In addition, the mass flux is

somewhat less in a U-tube SG and the mixture quality is unlikely to rise above 30%

or less, whereas in a OTSG the quality reaches 100% at the mixture level. The

thermal-hydraulic conditions are therefore quite different in the two types of SG,

implying that the apparent difference in heat transfer characteristics may not

necessarily be due to geometric features alone.

3.13 Semiscale tests

The Semiscale Mod-2C facility represents a 1/1706th scale, full height model of a

four-loop (3411 MWt) PWR. It consists of a reactor vessel with an electrically heated

core and two primary loops, identified as the intact loop and the broken loop. The

intact loop is scaled to represent three loops of the reference plant and the broken

loop represents the remaining PWR loop. Each loop has an active U-tube SG in which

the component flow areas, volumes, lengths and pressure drops are sized to simulate

a Westinghouse Model 51 SG, although both SGs incorporate an external downcomer

and have a separator which is not typical of any PWR plant. The broken loop SG has

only two U-tubes and the intact loop SG has six.

3.13.1 LOCA tests S-LH-1 and S-LH-2

Semiscale tests S-LH-1 and S-LH-2 both simulated a 5% cold leg break transient with

a normal pump trip. The only difference between the tests was the presence of a

large core bypass flow in S-LH-2. As a result, the core level depression prior to loop

seal clearance was less severe in S-LH-2 and, unlike S-LH-1, did not lead to core

uncovery. Following loop seal clearance however, both tests experienced a period

of core uncovery due to coolant boiloff.

Results from RELAP5/MOD2 calculation of S-LH-1 and S-LH-2 are described by Hall

and Bull (1989) and Brodie (1989). Like other RELAP5 LOCA test analyses, no

significant deficiencies in SG modelling were identified (as the SG played only a

minor role in the overall primary system energy removal). However, the following

points are worth noting.

1. When calculating the initial steady-state conditions for the tests, the SG level for

the intact loop had to be set artificially low to allow RELAP5 to simulate stable

operation of the SG. As a result, the SG liquid inventory was under-predicted by

some 20%. It is possible that errors in the void fraction modelling in the riser

could have contributed to this discrepancy (as seen in other calculations) -

21 In fact, in all these tests, the temperature differences were too low to allow reliable estimates to be

made of HTC-A and HTC-O, which is not the case for the OTIS experiment.

22 It should also be noted however, the Dimenna et al. (1989) do not quote the tube diameter and heated

length for the OTIS SG, and the dimensions used in this study for estimating the average heat flux were those given by Hassan (1988) for an unspecified 19-tube OTSG model.

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NP Restricted TECIL/04711R91

although, in the broken loop SG, the inventory was predicted to within the

experimental uncertainty.23

2. No difficulties were encountered in obtaining satisfactory steady-state predictions

for the SG pressures and primary side temperatures. As shown in Table 2 and

Table 3, the errors in H-ITC-A and HTC-O are relatively small. Although some

errors are a little larger than the uncertainty estimates (particularly those

associated with the broken loop SG in test S-LH-2), these are not thought to be

significant.

3. In the transient calculations, the SG pressures were systematically

over-predicted at all times after closure of the MSIVs, but the errors appeared to

have only a minor influence on the test predictions. For S-LH-1, some sensitivity

calculations were performed which indicated that tile discrepancies observed

later in the test could be accounted for by assuming a MSIV leakage of less than

0.16%.

4. Errors were present in the RELAP5 predictions of the liquid hold-up in the SG

tubes. However, no clear systematic trend was evident and the core level

depression prior to loop seal clearance was reasonably well predicted for both

tests.

3.13.2 Steam line break test S-FS-1

Semiscale test S-FS-1 simulated a 100% steam line break transient initiated from hot

shutdown conditions, with the break downstream of the flow restrictor. The intact loop

SG blew down through a separate break assembly but was isolated after 20 s, when

a safety injection (SI) signal was generated on low SG pressure. A loss of offsite

power was assumed coincident with the SI signal, causing the primary pumps to trip

and delaying the operation of the high head safety injection and auxiliary feedwater

systems. Results from a RELAP5/MOD2 analysis of the test are presented by Rogers

(1989a). For this analysis, the measured secondary side pressure behaviour in the

intact SG was imposed as a boundary condition.

Some difficulties were encountered in the RELAP5 analysis in calculating initial

steady-state conditions for the SGs. These difficulties were really a consequence of

the hot standby conditions - in particular the almost stagnant liquid conditions present

on the secondary side and the small heat transfer across the tubes2

". The difficulties

were eventually overcome by modifying the control systems used to adjust the

feedwater and steam valves and achieve the desired SG mass. The values of HTC-A

and HTC-O corresponding to the calculated and measured initial conditions indicate

a large over-prediction of SG heat transfer. However, as noted for the other steam line

break tests, due to the large uncertainty in these parameters at very low power

conditions, they do not provide a reliable indicator of heat transfer modelling.

Overall, the RELAP5 predictions for the transient phase of the test gave acceptable

agreement with the experimental results. A number of differences were apparent

however, and these are discussed below.

1. For the short period while single phase steam was being discharged from the

break (0-3 s), the predicted SG depressurisation matched the experimental

behaviour almost exactly. When the break flow became two-phase however, the

22 Note that for both tests, the analysts do not quote the calculated and measured SG levels, only the

inventories. It is possible that the experimental levels were not known and that the calculations

attempted to match the experimental inventories. If this was the case, the low inventory required for

stable operation of the intact SG could be explained by an under-prediction of void fraction in the

riser, and the correct prediction for the broken loop SG would not necessarily indicate correct void

fraction modelling in its riser.

24 It is likely that similar problems were encountered when setting up initial conditions for the

experiment.

-36 -
NP Restricted TECIL/04711R91

predicted pressure and cold leg temperature fell less rapidly. This indicates that

the liquid carryover to the break was greater in the calculation than the

experiment, a conclusion supported by the predicted and measured break flow

densities. A similar effect was observed in the RELAP5 analysis of the MB-2 and

LOBI steam line break tests. In the Semiscale calculation, the carryover

appeared to correct itself by about 15 s, with the subsequent depressurisation

being over-predicted.

2. The SG tubes dried out somewhat earlier in tile RELAP5 calculation than they did

in the experiment (at approximately 40 s as opposed to 70 s) and this, together

with the lower initial depressurisation, resulted in an under-prediction of the

overall primary side cooldown (by around 20TC). The later dryout in the

experiment could be explained by liquid that had accumulated in structures in the

upper SG during the early period of the blowdown, which was not modelled in the

calculation, draining back into the top of the riser. This type of effect was

apparent in the RELAP5 analysis of the MB-2 steam line break test, and it was

also suggested that accumulation of liquid in the experiment could explain the

over-prediction of carryover in the RELAP5 calculation. Although the opposite

behaviour was evident in the analysis of the LOBI steam line break test, with the

predicted dryout occurring late, much of the liquid carried into the upper SG in

the experiment was believed to have been boiled off by the release of metal work

heat.

3. When the SG dried out in the RELAP5 calculation, the heat transfer reduced in a

step-wise manner as each node progressively changed heat transfer mode (even

though the SG was relatively finely noded in the U-tube region). In contrast, the

experimental data indicated a rapid drop in primary to secondary heat transfer

when dryout occurred. This step effect is clearly a consequence of the absence

of any partitioning of heat transfer in a node between the regions above and

below a mixture level.

3.14 Studies performed at Texas A&M University

A number of studies of SG modelling in RELAP5IMOD2 have been carried out by Prof.

Y A Hassan and his co-workers at Texas A&M University. The results of these studies

are summarised below.

3.14.1 Steady-state heat transfer In a OTSG

Hassan and Blanchat (1990) describe a series of steady-state RELAP5 calculations for

a full sized model of a OTSG, that were carried out using modified versions of

RELAP5/MOD2. At 100% load conditions, the standard version of the code was found

to predict an outlet superheat of only 110C, compared to a 22-33°C superheat

suggested by plant test datazl. The authors attributed this discrepancy to the use of

correlations derived from tube data to represent thermal-hydraulic phenomena in the

bundle region. Following a review of the literature, the code was therefore modified

to apply more appropriate criteria for the bubbly to slug and slug to annular flow

regime transitions. These changes however, made little difference to the predicted

temperature profile - although fixing the slug to annular transition in the downcomer

to occur at a void fraction of 0.75 did result in near saturation conditions at the tube

bundle inlet (as would be expected), whereas tile standard version of the code

predicted subcooled conditions.

n5 In a OTSG, the primary fluid enters the SG through a nozzle at the top, flows down through more than

15 000 straight tubes and exits at the bottom head through two outlet nozzles. On the secondary side,

subcooled feedwater is sprayed downwards into an annulus between the SG shell and tube bundle

shroud, where it is heated to saturation by direct contact with steam aspirated from the bundle. It

then enters the bottom of the tube bundle where nucleate boiling begins. After reaching 100% quality

at about the mid-bundle elevation, the steam is superheated in the upper half of the bundle, flows

down through the steam annulus and exits through two steam outlet nozzles.

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NP Restricted TECILIO471 1R91

The authors then turned their attention to the heat transfer correlations applied in the

bundle region. The first modification introduced was to change the constant in the

Dittus-Boelter equation, which is used in the code to model convective heat transfer

to single phase liquid (and vapour), to be consistent with the results of Weisman

(1959), who correlated available heat transfer data for water flowing parallel to tube

bundles. This resulted in an increase in the value of the constant from 0.023 to 0.0304,

but since single phase forced convection to liquid was never predicted in the OTSG

bundle, the modification had no effect on the temperature profile (even though it was

presumably applied in the single phase vapour region). However, introducing the

same modification to Hma in the code's Chen correlation (see Section 2.4) resulted in

an increase in the predicted superheat of 1.10C.

Based partly on an equation for the liquid-only two-phase friction factor developed

by Ishihara, Paleo and Taborek (1980) using data from a variety of tube banks, the

authors then performed a number of investigations in which the factor F in the Chen

correlation was modified in various ways to increase the heat transfer coefficient to

the range determined experimentally by Boucher (1987b) for the Semiscale SG. The

expression ultimately adopted was

F =+ Xtt 12 ]0.75 (21) X• J

which, together with the other modifications, produced an increase in the predicted

superheat for the OTSG model of 80C The combined modifications also resulted in

better predictions of the primary and secondary side temperature profiles for a

19-tube integral economiser OTSG test facility, although the outlet superheat was still

under-predicted.

In a subsequent study, Blanchat and Hassan (1989) examined the application of

different CHF correlations for the secondary side of a OTSG, and recommended the

use of the lookup table developed by Groeneveld, Cheng and Doan (1986) - although

this had no effect on the steady-state predictions. Rais and Hassan (1990) later

showed that with the proposed bundle forms of the boiling heat transfer and CHF

correlations, and if the interphase drag in the bundle was also determined using the

EPRI drift flux correlation, RELAP5/MOD2 correctly predicted the degree of superheat

for a 30-tube OTSG facility at 65, 75 and 100% load conditions. The authors evidently

attribute the improved agreement over the earlier comparisons to the new interphase

drag model. The present authors believe that this is due to a lower void fraction being

calculated in the mixture region, resulting in an increase in the modified suppression

factor used in the Chen correlation.

3.14.2 Steady-state heat transfer in a U-tube SG

Hassan and Salim (1990) describe the development of a boiling heat transfer

correlation for RELAP5/MOD2 for application in a U-tube SG. The work was performed

using a two-loop model of a full size PWR with pre-heated type SGs. During the

steady-state initialisation of the model, it was found that the code under-predicted the

primary to secondary heat transfer. In particular, RELAP5 calculated a cold leg

temperature of 299.45 degreeC, which was significantly higher than the design

temperature of 293.350C.

In an attempt to rectify this effect, the bundle heat transfer and critical heat flux

correlations developed for the OTSG case were implemented in the code. As a result,

the predicted cold leg temperature was reduced to 296.65°C, which was still 3.3°C too

high, indicating that a further improvement in heat transfer was required. A short

investigation was then carried out using different expressions for the factor F in the

Chen correlation, from which the following equation was recommended

5L -- 17

+]0.75 _.L X• <5 5 (22)

Xtt X• J

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NP Restricted TECILIO4711R91

This particular equation produced a cold leg temperature of 294.95°C (which was only

1.6°C too high) and also gave good agreement with experimental data obtained from

the Semiscale SG. In particular, both tne data and modified correlation exhibited a

convex variation of heat transfer coefficient along the U-tubes, whereas the standard

RELAP5/MOD2 correlation shows a slightly concave variation.

Note that the expressions for F developed at Texas A&M for the OTSG and U-tube SG

situations (i.e. equations (21) and (22)) have a very similar form, except that the

equation for a U-tube SG produces a more rapid increase in F with X.I and thus

quality. This may reflect the fact that the quality in the boiling region of a U-tube SG

is not as high as in a OTSG. In particular, X"-1 varies from 0.2 to 50 in the nucleate

boiling regime of a typical OTSG (Hassan and Blanchat, 1990), but for a U-tube SG,

does not normally exceed a value of 5 (Hassan and Salim, 1990).

It is also worth noting that equation (21) (and (22)) predict a much stronger

dependence of F on quality than the standard expression, viz.

F= 2.35(X- 2 + 0.213)"'7 (23)

and this dependence seems to be supported for U-tube SGs by measurements of the

boiling heat transfer coefficient from Semiscale. However, it should not be forgotten

that the calculations performed by Hassan and Salim (1990) using their modified form

of RELAP5/MOD2 also (presumably) made use of the modified suppression factor S'.

As noted in Section 2.4, this tends to enhance the heat transfer coefficient as the void

fraction (and thus quality) reduces, which is the opposite trend to that observed in the

Semiscale data.

In a separate study using the Texas A&M RELAP5/MOD2 U-tube PWR model, Salim

and Hassan (1990) found that the desired heat transfer between the primary and

secondary sides could be achieved by reducing the hydraulic diameter of the SG

U-tubes. The specific value used is not quoted, but it is evident from the values of the

heat transfer coefficients shown that a reduced hydraulic diameter was used on both

the primary and secondary sides. The authors also claim that the use of a smaller

hydraulic diameter did not significantly affect the predicted transient behaviour for a

small break LOCA calculation.

4 ANALYSIS OF WOLF CREEK STEADY-STATE

COMMISSIONING TESTS

This section describes a RELAP5/MOD2 analysis of SG behaviour recorded during a

series of steady state commissioning tests carried out on the Wolf Creek PWR over

a range of load conditions. The results are used to provide information on SG heat

transfer modelling in the code, which is directly relevant to Sizewell 'B' plant

calculations. Sensitivity studies are performed to ascertain the adjustments required

to selected heat transfer parameters in order to achieve agreement with the test data.

A brief description of the Wolf Creek PWR is given at the beginning of Section 3.1 and

in Table 1.

4.1 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, a four loop system with a rated thermal output of 3425

MW, was chosen to be the site of this test programme.

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NP Restricted TECILIO471/R91

During the commissioning tests measurements were concentrated on a single steam

generator which had been constructed with 13 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 10% to 23%.

Internal SG measurements relevant to the present study are:

1. Downcomer flow velocity.

2. Downcomer fluid temperature.

3. Component pressure drop.

The main reason for measuring fluid velocity and temperature within the downcomer

is to enable the circulation ratio to be evaluated. Location of static pressure tapping

points throughout the steam generator are shown in Fig. 6. Pressure drop

measurements of special interest are summarised in the following table.

Differential Pressure measurement

pressure Component location

PT7-PT8 Mid-deck plate Across mid-deck plate

Tube bundle primary From low in the downcomer

PT9-PT8 separator and mid to the steam dome

-deck plate

PT8-Pi Primary and From mid-deck plate

secondary dryer to steam dome exit

A full description of the special instrumentation used and location details can be

found in Curlee and Preece (1987).

In addition to the special instrumentation, measurements were supplemented by

general plant data from permanent process and control instrumentation which form

part of the nuclear steam supply system (NSSS). These data include:

1. Primary circuit - pressuriser pressure

- hot and cold leg temperature

- reactor power

2. Secondary circuit - steam flow, temperature and pressure

- feed flow

- feedwater inlet temperature

- downcomer water level

The steam pressure measurements, which provide an important quantity for

assessing SG heat transfer modelling in RELAP5, were normally made using three

pressure transducers equi-spaced around the steam main at the containment

boundary. However, for some of the earlier tests, these were supplemented with a

Heise direct-reading gauge mounted on the steam main close to the pressure

transducers. The Heise gauge values are the preferred ones because of the

relatively high precision achievable, _0.1bar, compared with + 0.45bar for the

pressure transducer.

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NP Restricted TECILIO471/R91

In addition to this other relevant parameters are derived from plant measurements.

For example the primary fluid flow is deduced from measurements of the cold and

hot leg temperatures and thermal power (NSSS trend block data). However, because

the cold leg temperature is measured downstream of the recirculation pump, some

temperature correction is initially required to allow for the heat input, Qp, from

pumping. Thus, the corrected cold leg temperature, T',, is found from:

OP

T'C TC CP rm'p (24)

where Q, = 4954.25kW, (Davis, 1984) and rh', is an initial guessed value for the

primary fluid flow. Since enthalpy on the cold leg side, He, is found from H, = H,(T',, P,) , then

Q rm P (Hý_-H'c) (25)

The calculation is concluded when 'p mp

The effective values of the downcomer flow and feed flow are the plant values minus

the blowdown flow. Thus the circulation ratio, on a mass flow basis, is given by:

(rhd - rhb) (26)

(rif - rnb)

Before describing the way in which the Wolf Creek plant data are used in

RELAP5/MOD2 it is important to understand the limitations in the plant measurements

and then to quantify these. This is all the more important since in practice the steam

pressure is not measured in the area of interest in the inlet plane of the steam

generator flow restrictor nozzle, but at some considerable distance downstream in

the steam main at the containment boundary. Careful account then has to be taken

of the static and dynamic pressure head loss in order to deduce an accurate value

of the steam pressure in the plane of interest. The problem with this method is that

it can introduce errors in addition to the basic pressure measurement error and this

needs to be addressed. The standard method employed to quantify compound errors

of this type is outlined in Appendix B. The results of this are summarised in the

following table in terms of uncertainties about the mean value for a given steam

generator duty.

Load, % 36.1 98.9

Mean p ear 72.400 70.738

pressure, bar

Uncertainty, % _.0.143 -I-0.147

4.2 Tests analysed

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

detailed analysis using the RELAP5 code. The SG conditions covered by these tests

are summarised in Table 4. The data selected concentrate on representative tests

which encompass a wide range of load conditions (36% to 99% of the full thermal

output of 856 MW from a single steam generator). The selection makes use of the

most accurate measurements of steam pressure, based on the Heise gauge values,

whilst avoiding the low load cases which may suffer from some form of control

system instability.

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NP Restricted TEC/L/O471/R91

4.3 RELAP5/MOD2 model

For each of the tests selected, RELAP5/MOD2 was applied to simulate the

thermal-hydraulic performance of a single SG. The input deck used for these

calculations was 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 nodalisation scheme used for the calculations is shown in Fig. 7 and

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

side main feed and the steam generator 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 TDV26. 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 steam generator 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 4

The calculations were performed using an IBM version of RELAP5/MOD2 cycle 36.04

operating on the NP TEC VM system. Some of the calculations were also repeated

using the PWR-Project version (P01) of the code.

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

affects the local fluid properties.

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NP Restricted TEC/L/0471/R91

The calculated values of the secondary side pressure and pump suction leg

temperature were used to assess SG heat transfer and vapour generation rate

modelling in the code. As discussed in Section 2.2. for the boundary conditions

imposed, an error in the secondary side pressure may be attributed to errors in the

primary to secondary heat transfer and vapour generation rate models. An error in

the pump suction leg temperature on the other hand, may be attributed solely to

errors in the vapour generation rate models.

4.4 Base case calculations

The difference between the calculated value of steam pressure in the steam

generator dome and that deduced from plant measurements at Wolf Creek is

indicated in Fig. 8 over a range of reactor power. The calculated steam pressure in

this case is that evaluated at the nodal element 612-03 (see Fig. 7), which is

compared with the plant value corresponding to the inlet to the steam flow restrictor

nozzle (see Fig. 6). The difference between the measured and calculated values are

shown to increase monotonically with load, leading to a maximum discrepancy of

about 3.5 bar at full load. The initial question raised is whether or not this could be

the result of inadequate modelling of the pressure loss components throughout the

height of the steam generator.

In practice the main pressure drop component measured on plant is that between the

bottom of the downcomer to a point just above the mid-deck plate. Comparison of the

calculated and measured values are shown in Fig. 9. Since agreement is to within

about 4% over a substantial load range, this tends to confirm the adequacy of the

component pressure loss modelling up to this level. Individual, although small

component pressure losses, however, might not be necessarily well modelled as

shown for example in Fig. 10, concerning the pressure loss across the mid-deck

plate. Better modelled is the pressure loss component between the mid-deck plate

and inlet to the steam restrictor nozzle shown in Fig. 11, but again this component is

relatively small. Since, therefore, the overall component pressure drop appears to

be reasonably well modelled, this leads to the conclusion that the bulk of the

discrepancy lies in the calculation of the absolute value of the steam pressure as

indicated in Fig. 8.

To investigate this, a comparison between the measured and calculated temperature

at the SG outlet plenum (pump suction leg) has been carried out over the load range

36% to 99%. The measured temperature used in this case incorporates a correction

to allow for the heat input from pumping as described in Section 4.1. The calculated

temperature corresponds to that in the nodal element 330!01 (see Fig. 7).

The comparison in fact indicates very close agreement with the measured data to

within +0.3 to -0.04 deg.C. This effectively validates the vapour generation rate

model and as such eliminates it as a prime cause of the error in the calculation of the

steam outlet pressure. This error may therefore be attributed to inaccuracies in the

modelling of primary to secondary side heat transfer.

4.5 Sensitivity calculations

As discussed in Section 2, it is possible to compensate for an under-prediction of SG

heat transfer in RELAP5/MOD2 by:

1. increasing the SG tube thermal conductivity,

2. increasing the total heated length of the SG tube, or

3. reducing the heated diameter on the secondary side of the tube bundle, to

increase the local heat transfer coefficient.

These changes can be achieved by modifying the RELAP5/MOD2 input deck.

- 43 -
NP Restricted TECILI0471/1R91

A series of sensitivity calculations, therefore, was carried out to determine the

parameter adjustments required to force the code to calculate the correct secondary

side steam outlet pressure.for the Wolf Creek tests at 36% and 99% load. The results

are summarised in Table 5. Bearing in mind that there is a degree of uncertainty

associated with the plant measurements, the effect of this, expressed in terms of

upper and lower bounds on the adjustment factors, is also indicated.

In all cases the adjustment factor increases with the power level. Of the heat transfer

parameters examined, the heated tube length appears to be the most sensitive, with

an adjustment of only 8 to 28% required over the power range examined. Even

allowing for plant measurement uncertainties the range of adjustment extends only

to 6 to 29%. The heat transfer appears also to be fairly sensitive to the tube thermal

conductivity. In contrast to this the factor applying to the heated-tube diameter has

considerably less impact. In this case the heated diameter (expressed as a fraction

of the inter- tube gap) ranges from 1/8.3 to 1/164. With plant measurement

uncertainties taken into account the factor range extends from 1/5.7 to 1/179. These

results are consistent with the theoretical analysis discussed in Sections 2.4 and 2.6.

A further example of the way in which the heat transfer is affected as a result of a

forced parameter change in order to achieve the correct steam outlet pressure is

illustrated in Fig." 12, for the heated tube diameter. This shows the variation of the

heat transfer coefficient, at 99% load, with the developed SG tube length, in a

comparison between the enhanced value and the base case value. The

enhancement, shown as a ratio in the inset figure, indicates over a three-fold

increase in heat transfer in the upper region of the tube bundle. Most of the gain in

the heat transfer occurs on the cold leg side, although on this side the heat transfer

coefficient is clearly less.

The load dependence of the effect of heated tube diameter on heat transfer is further

illustrated in Fig. 13, for the 36% load case. As expected, a much smaller change is

required as suggested by a maximum increase in the heat transfer of only about 30%,

again in the upper region of the tube bundle.

5 ASSESSMENT OF STEAM GENERATOR MODELLING

DEFICIENCIES

The Sections 3 and 4 have discussed and investigated various deficiencies in SG

modelling that have come to light during RELAP5/MOD2 analyses of plant and rig

data. The principal phenomena involved may be summarised as follows.

1. Primary to secondary heat transfer under steady-state. normal operating and

start-up conditions.

2. Secondary side liquid inventory under steady-state conditions.

3. Downcomer level trajectories under fast boildown conditions (e.g. at full power).

4. Wall heat transfer above the mixture level during boildown and blowdown

transients.

5. Dynamic level swell in response to steam discharge from the SG.

6. Liquid carryover to break during steam line break transients.

7. Separator performance.

8. CCFL and liquid hold-up on primary side.

9. Secondary side condensation effects during repressurisation transients.

- 44 -
NP Restricted TEC/LI04711R91

This section presents an assessment of RELAP5/MOD2 in the above areas. In each

case, the principal code deficiencies are identified together with the specific models

responsible. Where appropriate, the errors involved are quantified and the

underlying parametric trends are noted. The implications for PWR safety analysis are

also assessed and methods for compensating for the deficiencies through changes

to the input deck, which may be used in sensitivity studies, are discussed.

Consideration is also given as to whether the deficiencies identified will still be

present in RELAP5/MOD3 and recommendations for future code development and

assessment studies are made. These may be addressed in a subsequent study

aimed at assessing SG modelling in RELAP5/MOD3.

At this stage however, it is worth noting that although the list of deficiencies may

appear long, in most cases the implications for PWR safety analysis are not serious.

5.1 Primary to secondary heat transfer under normal operating

conditions

5.1.1 Code deficiencies

In the review of code assessment calculations presented in Section 3, numerous

examples were noted in which RELAP5/MOD2 had under-predicted the total primary

to secondary heat transfer under steady-state normal operating conditions. The Wolf

Creek analyses described in Section 4 confirmed this effect and also showed that it

was not restricted to full load conditions.

As discussed in Section 2.7, a convenient way of quantifying modelling errors in this

area is to compare predicted and measured values of the overall primary to

secondary heat transfer coefficient. Two definitions of this parameter were

considered (HTC-A and HTC-O) and, as part of the review, the % errors in the

predicted values of both parameters were evaluated for all RELAP5/MOD2

calculations considered, for which sufficient information had been reported. The

results of this exercise are summarised Table 2 and Table 3.

In order to obtain an overall picture of SG modelling in RELAP5/MOD2 and to

ascertain whether the deficiencies observed depend on SG conditions and scale, the

errors in HTC-A and HTC-O have been plotted in Fig. 18 and Fig. 19 against the

measured value of the average SG heat flux. Additionally, Fig. 20 shows the errors

in HTC-A plotted against the average primary to secondary temperature difference

DT-A. In each case, where the error in HTC lies within the experimental uncertainty,

the uncertainty is shown as a vertical line (although the uncertainties in the LOBI and

OTIS data are not known).

Before considering the results presented it is important to recognise that the

uncertainty in HTC can be relatively large. In addition, as discussed in Section 2.6,

the predicted primary to secondary heat transfer is dependent on the definition used

for the secondary side heated diameter and the calculated value of the recirculation

ratio, both of which are not always known. The review of RELAP5 analyses also found

that other modelling parameters could affect calculated SG heat transfer - e.g.

nodalisation. All these factors tend to obscure the trends exhibited in the figures,

making it difficult to formulate precise conclusions regarding the underlying

modelling deficiencies. Nevertheless, the following statements may be made.

The figures provide substantial evidence that RELAP5/MOD2 under-predicts SG

heat transfer under steady-state conditions and over a wide range of load

conditions. Although many of the error points lie within the experimental

uncertainty, there is a clear bias in the direction of an under-prediction of heat

transfer.

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NP Restricted TECILIO471/R91

* The Wolf Creek results obtained at NP TEC with the secondary side heated

diameter set equal to the tube OD, and for which the recirculation ratio was

calculated correctly, show that the errors in HTC tend to reduce with decreasing

reactor power - which may be interpreted as a reduction with decreasing heat

flux or decreasing primary to secondary temperature difference. The deviation

from a strictly monotonic trend possibly results from data uncertainties.

A reduction in the error in SG heat transfer modelling with reactor power was

also apparent in the Wolf Creek sensitivity calculations described in Section 4.5.

In particular, the parameter adjustments needed to achieve agreement with the

data at 99% load were significantly greater than those required at 36% load.

" Of the two Wolf Creek calculations performed at NE TD (which were also made

with the secondary side heated diameter set equal to the tube OD), the

calculation providing the smaller errors in HTC corresponds to the plant trip test.

As discussed in Section 3.1.2, the data for this test may not be very reliable. The

results for other calculation are similar to the NP TEC results, and the slightly

lower errors in HTC are consistent with a higher prediction for the recirculation

ratio.

For the smallest scale facilities incorporating U-tube SGs - viz. SPES at 1/427th

scale, LOBI at 1/712th and Semiscale at 1/1706th - the tendency for

RELAP5/MOD2 to under-predict SG heat transfer is less noticeable (if anything,

the results indicate a slight tendency to over-predict heat transfer). This implies

that the errors in heat transfer may reduce with scale. Note however, that the

LOBI results were obtained using a secondary side heated diameter equal to the

inter-tube gap (and a reduced wall thickness), and the recirculation ratio for the

SPES calculation was high by a factor of two. (The values of these parameters for

the Semiscale calculations are not known.) Both effects would tend to enhance

the predicted SG heat transfer with respect to the Wolf Creek results, although it

is unlikely that they would compensate for the errors seen completely. This view

is supported by the results from an analysis of Wolf Creek at full load, performed

with the secondary side heated diameter equal to the inter-tube gap instead of

the OD. As evident in the figures, the errors in HTC are still greater than those

associated with the LOBI calculation. It should also be noted that although the

SPES calculation used a high recirculation ratio, the secondary side heated

diameter was set equal to the hydraulic diameter which is greater than the tube

OD. In addition, the SPES, LOBI and Semiscale results correspond to a higher

SG heat flux and primary to secondary temperature difference than the Wolf

Creek results, and thus the tendency to under-predict heat transfer would be

expected to be greater.

The results from the LOFT calculations (1/68th scale) show considerable scatter,

although a tendency to under-predict SG heat transfer is still evident. In general,

the errors corresponding to the UK calculations tend to be smaller than those

associated with the Non-UK calculations, reflecting the fact that most (if not all)

of the UK calculations were made with the secondary side heated diameter set

equal to the inter-tube gap, whereas the others probably used the hydraulic

diameter.27 Overall, the results indicate that the errors in heat transfer modelling

are greater than those corresponding to the smallest facilities, and perhaps

slightly less than those associated with Wolf Creek2t . This trend is more evident

when the errors in HTC are plotted against the primary to secondary temperature

difference.

27 It is apparent in the figures that one UK calculation gave an error in HTC somewhat larger than the

others. This calculation was the first LOFT calculation performed using RELAP5IMOD2 in the UK

(LP-SB-03), and it is possible that the secondary side heated diameter was set equal to the hydraulic

diameter.

28 Note however, that unlike the other facilities, the LOFT SG does not provide a full height

representation of a PWR U-tube SG.

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NP Restricted TEC/LI04711R91

0 The ROSA-IV results (1/48th) appear inconsistent with the results from other

facilities having U-tube SGs. However, the uncertainty in the experimental value

of HTC is very large, and the errors in HTC-A are broadly consistent with those

for Wolf Creek and LOFT when they are plotted against the primary to secondary

temperature difference.

0 The results for the OTIS OTSG facility (1/1632th scale) indicate a much larger

error in SG heat transfer modelling than would be expected for a U-tube SG of

similar scale and reactor power. However, as noted in Sections 3.12 and 3.14, the

thermal-hydraulic conditions in the two types of SG are quite different. Also, the

average SG heat flux considered in the figures is not very representative of the

actual heat flux below the mixture level in a OTSG (and in fact the errors in HTC-A

are slightly more consistent if the errors are plotted against the primary to

secondary temperature difference). Thus, the differences seen may not

necessarily be due to geometric differences between the. two designs.

It is appreciated that due to data and modelling uncertainties, the above type of

analysis can only really provide a guide as to how the errors in SG heat transfer

modelling in RELAP5/MOD2 vary with thermal-hydraulic conditions and geometry.

Nevertheless, there is clear evidence that the tendency to under-predict heat transfer

reduces with reactor power, and some evidence that it also reduces with the scale

of the facility.

The errors in the HTC-A and HTC-O shown Table 2 and Table 3 and Fig. 18 to

Fig. 20 provide some means of quantifying the errors inherent in RELAP5/MOD2's

modelling of SG heat transfer. Other measures appropriate to Sizewell 'B' are

provided by the secondary side pressure errors for the Wolf Creek analyses (Fig. 8),

and the parameter adjustments required to achieve agreement with the data

(Table 5).

5.1.2 Models at fault

Section 2.3 and, in particular, equation (1) show how various SG parameters affect the

calculation of SG heat transfer in RELAP5/MOD2. From this, it is clear that an

accurate representation of primary to secondary heat transfer requires the

appropriate modelling of the following features

1. Primary side heat transfer coefficient.

2. Heat conduction through tube wall.

3. Secondary side heat transfer coefficient.

4. Total heat transfer area.

From the Wolf Creek calculations described in Section 4, and the parameter

adjustments required to obtain agreement with the data, it would appear that the

tendency of the code to under-predict the overall SG heat transfer is unlikely to be

due to an inaccurate calculation of the heat conduction effect or heat transfer area.

In particular

0 As shown in Table 5, the increase in tube thermal conductivity necessary to

obtain the correct primary to secondary heat transfer at full load conditions was

74%. Alternatively, as evident from equation (1), the same effect could have been

achieved by increasing lit by at least this amount. These levels of adjustment are

significantly greater than the data uncertainties associated with these

parameters.

0 The increase in heated length and thus heat transfer area required to calculate

the total heat transfer correctly at full load conditions was 28%, which is again

significantly greater than the uncertainty in this parameter.

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NP Restricted TECILIO471/R91

Table 5 also shows that RELAP5/MOD2 was able to calculate the total SG heat

transfer correctly at full load conditions when the secondary side heated diameter

was set equal to the inter-tube gap divided by 167. As shown in Fig. 12. this resulted

in an increase in the secondary side heat transfer coefficient over the normal value

by up to a factor of more than 3. From the discussion in Section 2.3, it is apparent that

the same effect on SG heat transfer could have been achieved by increasing the

primary side heat transfer coefficient by a similar amount. For this calculation,

RELAP5/MOD2 used the standard Dittus-Boelter correlation, which for the case of

water flowing in straight tube at thermal-hydraulic conditions typical of those present

on the primary side of a SG, is generally quoted with an accuracy of ±25% (Dimenna

et al., 1988). Thus, even allowing for some enhancement in heat transfer due to

entrance and bend effects and the presence of fouling deposits on the inside of the

tubes, the tendency to under-predict SG heat transfer cannot be associated with the

application of the Dittus-Boelter correlation to determine the primary side heat

transfer coefficient. This effect must therefore be attributed to errors in the calculation

of the secondary side heat transfer coefficient.

The heat transfer behaviour of the secondary side of the U-tube bundle is controlled

by subcooled and saturated boiling, with the latter being the dominant mechanism"9 .

As discussed in Section 2.4, for both regimes, RELAP5/MOD2 uses a modified form

of the standard Chen heat transfer correlation, in which the modification provides an

enhancement to heat transfer as the void fraction reduces.

For the case of boiling heat transfer to water flowing in tubes and annuli, Gungor and

Winterton (1987) have compared the standard form of the Chen correlation against a

large databank of experimental measurements. The results of this comparison are

summarised below.

29 Heat transfer to single phase liquid is rarely present on the SG secondary side in RELAP5/MOD2

calculations at full load conditions.

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NP Restricted TECIL/O4711R91

Saturated Boiling

in a Tube

0.7 - 70

Saturated Boiling

in an Annulus

Subcooled Boiling

in a Tube

Pressure

(bar)

1 - 198 1 - 203

Mass Flux

(kg M- 2 s-')

Quality (%)

Subcooling ('C)

Heat Flux

(W cm- 2)

Wall Superheat

(OC)

Heated Diameter3"

(cm)

No. Data Points

Average Error (%)

Mean Absolute

Error (%)

59 - 8179

0- 70

0.5 - 228

0-49

0.3 - 3.2

70- 5700

0.2-66

3- 262

1 - 78

1120-61518

0.5- 174

12- 9153

1.7 -107

0.5- 2.2 0.3- 1.5

1454 546

-21.4

25.8

758

-3.5

22.5

-4.5 to 15.5 "

8.1 to 28.9

" Depending on dataset

Measurements made during the Wolf Creek commissioning tests considered in this

report, show that for operation in the range of 27 to 99% load, the pressure on the

SG secondary side varies between 69 and 72 bar and the mass flux in bundle region

varies from 295 to 460 kg m-1 s- 1 (Curlee and Preece, 1987). In addition, the base case

RELAP5/MOD2 analysis of the 99% load case described in Section 4.4, predicted

minimum subcooling of 4.2°C, a maximum bundle quality of 30%, a wall heat flux in

the range 9 to 40 W cm- 2 , and a wall superheat between 3.6 and 10.30C. Gungor and

Winterton (1987)'s comparisons therefore suggest that for saturated and subcooled

boiling in tubular and annular geometries, the standard form of the Chen correlation

provides a reasonably accurate prediction of heat transfer coefficient for the

thermal-hydraulic conditions typical of those present in a PWR U-tube SG.

It must therefore be concluded that boiling heat transfer in a full size SG tube bundle

at full load conditions, is much greater than that in a tube having an equivalent heated

diameter. As a result, even though the Chen correlation in RELAP5/MOD2 is modified

to enhance heat transfer, a significant under-prediction of the heat transfer coefficient

still results.

The above conclusion is also supported by experimental data on local heat transfer

in tube/rod bundles. In particular

1. Data obtained by Boucher (1987b) for the Semiscale Mod-2C broken loop SG (type

Ill) operating at full power conditions, show that the standard form of the Chen

correlation under-predicts the secondary side heat transfer coefficient by a factor

of typically 2 to 5. As evident from Table 1, this SG consists of two inverted

30 Defined as 4 x flow area divided by heated perimeter

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NP Restricted TEC/L/O471/R91

U-tubes (giving a four tube bundle) and represents a 1/1706th scale full height

model of a Westinghouse Model 51 SG. The experimental heat transfer

coefficients were derived from local temperature measurements for the primary

and secondary side fluid and tube outer wall, obtained from thermocouple triplets

located at various elevations. To minimise the uncertainties associated with

measuring small temperature differences, the results were calibrated against

data obtained under zero heat transfer conditions.

2. Heat transfer measurements made by Kumamaru, Koizumi, and Tasaka (1987) in

a 5 x 5 electrically heated fuel bundle also indicate that the Chen correlation

under-predicts the local boiling heat transfer coefficient by a considerable

amount. The experiments were performed in the JAERI two-phase test facility

(TPTF) and the results presented correspond to a pressure of 30 bar, mass fluxes

around 300 kg m-2 s-1 , heat fluxes in the range 4.9 and 25 W cm 2, and inlet

qualities between 12 and 57%. The authors note that due to the very small

temperature differences involved, the uncertainties in the experimental heat

transfer coefficients were quite high - varying from -36% to + 125% at 3 x 10,

W cm-2 °C-' , to -63% to +oo% at 9 x 1 04 W cm 2 *C-1. Nevertheless, all results

imply an under-prediction of heat transfer and in many cases the errors are

outside the uncertainty in the data.

Although the Semiscale SG is very small scale, the heat transfer date obtained by

Boucher (1987b) is particularly relevant to the present study. As shown in Fig. 14, the

local heat transfer coefficients derived from the data generally increase with bundle

elevation3", which may be interpreted for example as an increase with increasing void

fraction. In contrast, the standard form of the Chen correlation provides a heat

transfer coefficient that reduces slightly with elevation, in addition to under-predicting

the data by a large amount.

As noted earlier, RELAP5/MOD2 applies a modified form of the Chen correlation in

which the heat transfer coefficient is enhanced as the void fraction reduces. This

modification however, achieves the opposite effect to that observed in the Semiscale

data. This has been illustrated in Fig. 14 by showing (from Fig. 12) the secondary

side heat transfer coefficient predicted by RELAP5/MOD2 for the Wolf Creek SG at full

load conditions (base case calculation) - which is representative of what the code

would calculate for the Semiscale SG. Although the RELAP5 modification clearly

enhances the heat transfer coefficient, it exacerbates the incorrect trend inherent in

the standard form of the Chen correlation. Interestingly however, the average heat

transfer coefficient for the bundle in the RELAP5 calculation is not very different from

that implied by the data, which explains why the code did not do too badly when

calculating initial conditions for the Semiscale tests considered in Table 1. By

comparing Fig. 14 with Fig. 12 though, it is apparent that the level of heat transfer

coefficient necessary to achieve agreement with the Wolf Creek data is somewhat

higher than that measured on Semiscale, giving further evidence that the boiling heat

transfer coefficient for a SG bundle depends on scale.

5.1.3 Implications for PWR safety analysis

For steady-state PWR analysis, the deficiencies in RELAP5/MOD2's modelling of SG

heat transfer will result in an under-prediction of either the power dissipated in the

SG (and thus the core power), the primary to secondary temperature difference, or

some combination of the two. It is likely, however, that the code would be initialised

with the correct core power and primary side temperatures, and thus the deficiencies

31 Boucher (1987b) suggests that the reduction in the experimental heat transfer coefficient at an

elevation of 8.86 m is probably due to the increase in flow area and consequent reduction in mass

flux at this elevation. He also argues that the apparently anomalous reduction at 5.64 m is possibly

caused by a flow perturbation or flow regime transition nearby, resulting in some multidimensional

effects which significantly reduce the local flow rate. This effect is not evident in the data for the short

tube hot side.

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NP Restricted TEC/L/0471/R91

would manifest themselves primarily in an under-prediction of the secondary side

pressure. For the Sizewell 'B' SG, the magnitude of this will be similar to that seen

in the Wolf Creek calculations (see Fig. 8). Note that the pressure errors reduce with

load because the modelling errors (i.e. errors in HTC) and the primary to secondary

temperature difference both reduce with reactor power.

Before discussing the implications for transient analysis, it is important to note that

while the primary and secondary sides of the NSSS are well coupled, i.e. wet wall

conditions exist on both sides of the SG tubes, virtually all the power generated in the

core will be removed by the SGs. Thus, any errors in SG heat transfer modelling will

chiefly be seen as an error in the primary to secondary temperature difference, and

should have little effect on the predicted SG heat transfer rate.

As explained above, at the beginning of a transient calculation, the deficiencies in SG

modelling will be reflected by an under-prediction of the secondary side pressure. It

is possible of course, that as the transient evolves this error could be transformed

into an over-prediction of the primary side temperature, and this might result for

example, in the early calculation of two-phase conditions. For a transient involving

a reduction in reactor power though, it is more likely that the initially low SG pressure

will eventually over-compensate for the deficiency in heat transfer modelling.

Evidently, this could manifest itself in an under-prediction of the primary side

temperature and perhaps a late prediction of two-phase conditions. In both cases

however, the error in the primary side temperature is unlikely to be more than a few

degrees C, and thus should not have a significant impact on the predicted primary

side behaviour. This is particularly true for LOCA analyses, as the SGs do not have

a very large effect on the primary system energy removal.

On the secondary side, the main consequence of an incorrect pressure calculation

are errors in the predicted timings for the operation of the automatic safety and

control systems, and postulated operator actions. In a SGTR transient, the time of

break flow isolation could also be affected. Given that the maximum error in the

initial pressure is only around 3.5 bar however, it is difficult to believe that the

modelling deficiencies would have a serious impact on the results of a safety

analysis. A possible exception though, is the predicted behaviour of the SG safety

relief valves, as the PORV set point is only around 8 to 10 bar above the SG normal

operating pressure, so that an under-prediction of 3.5 bar in the initial pressure may

not be insignificant. In the worst case scenario, this error could result in the code

failing to predict the opening of the PORV at all, for a transient in which the valve does

open and is postulated to stick open. Even so, provided that this possibility is

recognised, it should not diminish the value of the code for safety assessment

studies.

It is also worth noting that, following a reactor trip and the closure of the MSIV, the

secondary side pressure is quite sensitive to the small leakage through the valve

(and the heat losses to the environment). This leakage is not generally well known,

and thus the corresponding errors in SG pressure could obscure those arising from

deficiencies in SG heat transfer modelling.

The discussion up to now has been restricted to situations in which the primary and

secondary sides are well coupled. If this coupling is lost, as a result of dryout on

either side of the U-tubes, an error in SG heat transfer modelling could lead to an

error in the predicted SG heat transfer rate and thus the primary (or secondary) side

cooldown. Once dryout occurs however, the heat transfer calculation will no longer

be controlled by boiling heat transfer on the secondary side, where the deficiencies

occur. Thus, the only consequences of inaccuracies in the modelling of this

phenomenon, will be those arising from an error in primary to secondary temperature

difference at the beginning of dryout. Again, these may be expected to be relatively

minor.

It may be concluded, therefore, that the deficiencies in RELAP5/MOD2's modelling

of SG heat transfer at normal operating conditions, which are most severe at full

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NP Restricted TECIL/0471/R91

power and for full scale facilities, and may be attributed to an under-prediction of the

boiling heat transfer coefficient on the secondary side, should not have a serious

impact on the use of the code for PWR safety studies. The deficiencies are more an

inconvenience as they introduce errors into the calculation before the transient

begins, and although the implications of these errors are relatively minor, they must

still be taken into account when interpreting the results. This task is quite feasible

given the large body of information available from assessment calculations against

separate and integral test facilities, although it obviously complicates the arguments

that must be put together when making a safety case.

5.1.4 Techniques for compensating for deficiencies

Several approaches may be used to compensate for the deficiencies in

RELAP5/MOD2's modelling of SG heat transfer without introducing changes to the

code. These may be applied when carrying out sensitivity type calculations as part

of a safety case analysis.

The various techniques available for enhancing SG heat transfer through changes to

a RELAP5/MOD2 input deck are discussed in Section 2. Those recommended for

consideration are as follows.

* Increase the thermal conductivity of the SG U-tubes.

* Increase the total heated length of the tubes.

* Use a very small heated diameter on the SG secondary side.

None of these approaches of course have any physical basis, they are merely

artificial ways of increasing primary to secondary heat transfer. The approximate

changes necessary to obtain the correct SG heat transfer for Sizewell 'B' should be

similar to those required for the Wolf Creek SG, which are detailed in Table 5. Note

that an enhancement in heat transfer could also be achieved by reducing the U-tube

wall thickness, but since the key parameter is tube conductivity divided by thickness,

it seems sensible to restrict the changes to the conductivity component and keep the

U-tube geometry correct (particularly as reducing the tube thickness will also reduce

the secondary side heat transfer area slightly if the internal diameter is kept

constant). An enhancement in heat transfer could also be obtained by using a very

small heated diameter on the primary side of the tubes. However, since the

modelling deficiencies are associated with the calculation of the secondary side heat

transfer coefficient, it seems sensible to restrict any artificial changes made to the

heat transfer coefficient to this side.

The most appropriate way of enhancing SG heat transfer in RELAP5/MOD2 obviously

depends on the transient being analysed, and which ever technique applied must of

course be justified. The principal difficulty will lie in establishing the consequences

that the changes have when the SG conditions differ from those in the initial

steady-state. The following comments may help the reader in making a suitable

choice.

Secondary side heated diameter

One of the disadvantages of setting a very small heated diameter on the

secondary side, is that it is difficult to judge how this affects the Chen correlation

over the full range of thermal-hydraulic condition that might be encountered. This

problem is made worse by the fact that, as shown in Fig. 3, the size of diameter

required forces the correlation into a region where it becomes particularly

sensitive to changes in diameter.

A more important consideration however, is the effect that the small diameter

will have on SG heat transfer if the flow regime on the secondary side of the

bundle changes to one in which a different heat transfer coefficient correlation is

applied. In particular, as discussed in Section 2.3, the reason why SG heat

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NP Restricted TECILI0471IR91

transfer under normal operating conditions is not very sensitive to small changes

in the secondary side heat transfer coefficient, is that the primary and secondary

side heat transfer coefficients and wall conductance all have similar magnitudes.

If the secondary side dries out however, the SG heat transfer will be dominated

by the local heat transfer coefficient. Consequently, an inappropriate value of the

heated diameter could have a dramatic effect on SG heat transfer and, since the

primary and secondary sides would no longer be coupled thermally, this could

have a significant effect on the primary side cooldown.

The present authors would not therefore recommend that this approach be

adopted.

Bundle heated length

As discussed in Section 2.3, SG heat transfer is directly proportional to the total

heated length of the U-tube bundle, and thus an increase in this parameter should

have a similar effect on heat transfer for all thermal-hydraulic conditions

encountered. Evidently, this could result in an over-prediction of SG heat transfer

if the reactor power falls, but the error involved should be less than that

associated with adopting a low secondary side pressure.

Tube thermal conductivity

Similar comments regarding an over-prediction of heat transfer with reducing

reactor power, apply if the initial SG heat transfer is enhanced by increasing the

thermal conductivity of the U-tubes. However. an increase in conductivity has the

advantage that its influence reduces if the heat transfer coefficient on either side

of the SG decreases. In particular, its effect becomes negligible if a significant

degradation in heat transfer occurs. For this reason, the present authors believe

that this technique would be the preferred approach in most situations.

5.1.5 RELAP5/MOD3 considerations

In RELAP5/MOD3, the Chen correlation has been returned to its standard form and

thus the tendency to under-predict SG heat can be expected to be greater than that

associated with RELAP5/MOD2. This has already been observed in early

RELAP5/MOD3 calculations (Coney, 1990).

Despite the comments made in Section 5.1.3 regarding implications for PWR safety

analysis, it is suggested that a study be undertaken with the objective of developing

an improved correlation for RELAP5/MOD3 for boiling heat transfer in bundle

geometries. This work should clearly take account of the experimental data on local

heat transfer obtained for the Semiscale SG, for which a more detailed description is

given by Boucher (1987a). Some of the experiments carried out on the LOBI facility

may also provide data on local heat transfer. There are also a wealth of data relevant

to total SG heat transfer from a whole range of PWR plants and facilities, and these

may be used to characterise the scale effect.

The work on SG heat transfer carried out at Texas A&M by Prof Hassan and his

colleagues (Section 3.14) would clearly provide a good starting point for the study.

The modification to the Chen enhancement factor F derived for a U-tube bundle has

some physical basis and also makes use of the Semiscale data. However, it would

appear that the complete correlation still uses the RELAP5/MOD2 modified

suppression factor, which introduces an unphysical dependence on void fraction. In

addition, the correlation does not account for any scale effects.

The study should also take account of work performed at MEL in which various

boiling heat transfer correlations were compared against the complete series of Wolf

Creek commissioning tests using the NUMAC-PWR code. In particular, Atthey (1989)

found that the standard form of the Chen correlation significantly under-predicted the

total primary to secondary heat transfer, but that the predictions obtained with various

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NP Restricted TECILIO4711R91

forms of the Thom, Jens Lottes and Rohsenow and Bergles correlations were

generally within the uncertainty of the data. Interestingly, it was found that the Chen

correlation gave good agreement with the data when the suppression factor was set

equal to one. In a subsequent study, Atthey (1990) went on to show that the

RELAP5/MOD2 form of the Chen correlation also under-predicted the data by a

significant amount (although it performed better than the standard version), but that

good agreement with the data could be obtained if the enhancement factor F was

modified in the manner suggested by Hassan. Strictly speaking, these comparisons

are not an exact reflection of the performance of RELAP5/MOD2, as the void fraction

in the modified suppression factor was calculated using the EPRI drift flux correlation - which tends to predict a lower void fraction than the RELAP5/MOD2 interphase drag

model and thus a greater boiling heat transfer coefficient.

Although Atthey's comparisons are important in that they are based on

measurements from a full sized Model F SG, the data considered only covered

steady-state operation for reactor powers between 27 and 99% load. This covers a

relatively narrow range of thermal-hydraulic conditions (e.g. SG pressures between

69 and 72 bar and bundle mass fluxes from 295 to 460 kg m-1 s - ). Clearly any boiling

heat transfer adopted in RELAP5 for bundle geometries will have to be valid for the

full range of SG conditions encountered in plant safety studies, and possibly those

present in the core. The basic Thorn, Jens Lottes and Rohsenow and Bergles

correlations considered by Atthey only depend on the pressure and wall superheat,

although a limited mass flux dependence was introduced by combining the

correlations with the Dittus-Boelter correlation. In contrast, the Chen correlation

depends on pressure, wall superheat, mass flux, quality and diameter. Consequently,

there must be a question mark over how well these correlations would perform under

abnormal conditions. More importantly, none of the correlations predict the type of

variation of heat transfer coefficient with elevation that was apparent in the Semiscale

data.Y

Recent papers which discuss mechanisms governing boiling heat transfer in a tube

are those due to Gungor and Winterton (1986), Gungor and Winterton (1987) and

Klimenko (1988). These may also provide a useful input to the study.

5.2 Secondary side inventory under normal operating conditions

5.2.1 Code deficiencies

An assessment of how well RELAP5/MOD2 calculates SG liquid inventory is not

particularly straightforward, as in many cases the true inventory is not known

accurately. Also, as shown in Appendix A.2, SG inventory depends on both the

recirculation ratio and downcomer level, and these parameters are sometimes

artificially adjusted in RELAP5IMOD2 analyses to avoid instabilities in the predicted

separator behaviour.

Despite these difficulties, the review of RELAP5/MOD2 assessment calculations found

clear evidence to support the notion that the code under-predicts SG inventory when

the recirculation ratio and downcomer level are calculated correctly. Specific

examples include:

0 The RELAP5/MOD2 Sizewell 'B' model has been found to under-predict SG

inventory relative to the design value at normal operating conditions (Section

3.2).

32 It is also interesting to note that the dependence on wall superheat is different in each case - the

Thorn correlation depends linearly on wall superheat (as does the Chen correlation effectively); the

Rohsenow and Bergles correlation depends on wall superheat squared; and the Jens Lottes

correlation depends on wall superheat cubed.

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NP Restricted TECIL/O4711R91

* The initial inventory for ISP-22 (SPES loss of feedwater test) was under-predicted

by RELAP5/MOD2 (Section 3.10). Also, in the ISP-26 analysis (ROSA-IV 5% cold

leg break), the initial collapsed liquid level in tile riser was too low, implying an

under-prediction of SG inventory (Section 3.7).

* In the UK and Studsvik analyses of LOFT test L3-6 (2.5% cold leg break with

delayed reactor trip), RELAP5/MOD2 under-predicted the SG level after the

reactor had tripped and the secondary side fluid had settled. Various reasons for

this discrepancy were suggested, but the most likely cause is an under-prediction

of the liquid fraction in the SG riser at full power conditions (Section 3.8.3).

An under-prediction of initial SG inventory may also be inferred from inaccuracies in

the RELAP5/MOD2 predictions of SG level during boildown transients, particularly the

early prediction of SG dryout. However, as discussed in Section 5.3, in order to

obtain the correct SG level trajectory under fast boildown conditions, it is not

sufficient to calculate the correct initial inventory alone.

5.2.2 Models at fault

For a given recir.culation ratio and downcomer level, the calculation of SG inventory

in RELAP5/MOD2 depends, primarily, on the void fraction distribution predicted in the

riser section and the geometric modelling of the SG components (e.g. downcomer

flow area). Although the information needed to model these components may not

always be known accurately and additional errors may be introduced in setting up a

ID representation, the fact that the tendency to under-predict inventory has been

observed on a range of plants and facilities suggests that an over-prediction of the

riser void fraction is the major cause. This could result from an over-prediction of

interphase drag and/or subcooled vapour generation. However, the magnitude of the

inventory errors seen are sufficiently large that they cannot be explained by errors in

the subcooled boiling regime alone.

A detailed assessment of the wet-wall interphase drag models in RELAP5/MOD2 is

presented by Putney (1988; 1989a; 1989b). In particular, it is shown that for low flow

conditions in a PWR-type fuel assemblies, there is considerable experimental and

theoretical evidence which shows that the code over-predicts the interphase drag

force and thus void -fraction. 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 model 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 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 slug model

which neglects the effects of profile slip.

The hydraulic diameter of a PWR U-tube SG is only a little larger than that associated

with the core fuel assemblies. Consequently, RELAP5/MOD2 may be expected to

over-predict interphase drag in the SG bundle for the relatively low flow conditions

typical of normal operation. Note however, that the region of the SG riser above the

U-tube bundle consists of a circular channel having a diameter of around 3.6 m. In

this region, RELAP5/MOD2 will apply a slug flow model to calculate the interphase

drag force, which assumes that the diameter of the Taylor bubbles is 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

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NP Restricted TECIL/04711R91

bubbles) becomes significant - and tile consequences of neglecting profile slip are

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

It is possible therefore, that whilst RELAP5/MOD2 will over-predict interphase drag in

the heated section of the riser, this could be compensated for, to some extent, by an

under-prediction in the unheated section. However, for many cases, it is likely that

annular flow will be predicted to exist in this section of the riser, and thus the

deficiencies in the code's slug flow model will have no effect". Also, the unheated

section of the riser is somewhat smaller than the heated section. Thus, in general the

net effect of the known deficiencies in the RELAP5/MOD2 interphase drag models will

be an under-prediction of SG inventory - although this tendency will clearly reduce

as reactor power reduces.

5.2.3 Implications for PWR safety analysis

The principal consequences of an under-prediction of initial SG inventory are twofold:

" Following a reactor trip and the settling of the secondary side fluid (i.e. when the

recirculation flow in the SG has ceased and the vapour generation rate has fallen

off), the fall in the downcomer water level is likely to be over-predicted. As a

result, the subsequent auxiliary feedwater flow could be over-predicted - although

the net effect might be to make-up for the initial under-prediction of inventory,

thereby mitigating the deficiency.

" For transients involving SG boildown, degradation of primary to secondary heat

transfer and SG dryout could be calculated to occur too early. The situation is

worse for slow transients in which most of the boildown takes place under decay

levels, as the excess inventory on the plant can take a long time to boiloff. For

example, in the RELAP5/MOD2 analysis of the SPES loss of feedwater test

(Section 3.10), SG dryout was predicted to occur more than 1000 s too early.

The second consequence evidently holds the more serious implications for PWR

safety analysis, as it has a dramatic effect on the timings of all subsequent events.

5.2.4 Techniques for compensating for deficiencies

As shown in Appendix 2, the tendency for RELAP5/MOD2 to under-predict SG

inventory may be reduced by artificially increasing the recirculation ratio, either by

raising the downcomer level or reducing the flow losses. Unfortunately, from the

analyses reviewed in this report, it would appear that the adjustments required to

obtain the correct initial inventory are not physically sensible (and as discussed in

Section 5.3, for fast boildown transients it may not be sufficient to calculate the

correct inventory alone). The same may be said if an increase in SG inventory is

sought by increasing the downcomer flow area. In addition, artificial adjustments of

the SG parameters can have undesirable effects. For example, if the reactor and

MSIV trips in a boildown transient occur on low downcomer level, raising this level

may serve to delay the predicted trip time and result in a greater boiloff in the pre-trip

period - thereby exacerbating the effects of an under-prediction of the initial

inventory. This dilemma was encountered in the RELAP5IMOD2 analysis of the SPES

loss of feedwater test. It was also found that, due to the over-estimation of the

frictional pressure losses, the use of a high recirculation ratio prevented the code's

values for the downcomer DP measurement being used to determine the downcomer

level (as is done on the plant).

An alternative method of increasing the initial SG inventory is to use an artificially

large value for the riser hydraulic diameter. This approach will lower the interphase

drag calculated in the slug flow regime, by reducing the interfacial area per unit

33 This was the case for example, in the RELAP5IMOD2 analysis of the Wolf Creek commissioning test

at 99% reactor power, but not at 36% power.

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NP Restricted TECILI04711R91

volume, and was apparently adopted in the KAERI analysis of the KNU#1 loss of

offsite power transient (Section 3.3). However, the actual value of hydraulic diameter

used was around 250 cm, and it is not clear what effect this would have on the

interphase drag calculations for the annular and post-dryout flow regimes - both of

which will be encountered in the SG bundle during a boildown transient.

Consequently, the approach may be difficult to justify, even for sensitivity type

calculations.

Finally, note that any modification which increases the SG inventory by reducing the

riser void fraction will also tend to enhance the primary to secondary heat transfer -

although even if the inventory is calculated correctly, SG heat transfer is still likely to

be under-predicted.

5.2.5 RELAP5/MOD3 considerations

RELAP5/MOD3 incorporates improved interphase drag models for both rod/tube

bundles and large diameter vertical pipes (Putney, 1989b). As a result, the code

should provide a better prediction of SG inventory than RELAP5/MOD2. Early

calculations with MOD3 have shown a 25% increase in inventory over MOD2

(Coney, 1990).

On the LOBI-MOD2 facility, a series of 'characterisation' tests have been carried out

with the specific aim of obtaining accurate data on SG inventory at steady-state full

power conditions. In these tests, designated as BC-01 (Sanders, 1988a; 1988b), the

rig was first stabilised at normal operating conditions, and then the fluid on the

secondary side was trapped by suddenly closing the SG feed and steam valves

(within I s). Tests were performed at four different settings of downcomer water

level, and the secondary side inventories were measured in both the intact and

broken loop SGs.

These LOBI experiments would clearly provide a good test of SG inventory modelling

in a code such as RELAP5 (as well as good data on SG heat transfer). It is therefore

recommended that they are used to assess RELAP5/MOD3 in this area.

5.3 Level trajectories during fast boildown transients

5.3.1 Code deficiencies and models at fault

In the previous section it was shown that RELAP5'MOD2 tends to under-predict SG

inventory at full power conditions and that the primary reason for this is an

over-prediction of interphase drag in the bundle region. One of the consequences of

this deficiency is an early prediction of SG dryout in boildown transients, and this is

particularly evident when the boildown occurs under decay heat conditions. For

transients in which a significant part of the boildown takes place at full power

conditions however, RELAP5/MOD2 has also been found to over-predict the rate at

which the downcomer level falls. Particular examples are the UK analyses of LOFT

tests LP-FW-01, L9-4 and L9-3 (Section 3.8.4), and the INER-Taiwan analysis of MB-2

LOF test no. 1 (Section 3.9.1). Furthermore, although attempts to artificially increase

the initial inventory by increasing the recirculation ratio and downcomer flow area

had the affect of delaying the time of SG dryout (i.e. for LOFT tests L9-4 and L9-3),

they did not have a significant effect on the rate of boildown. This raises questions

over the role played by the code's interphase drag models in determining level

behaviour during the boildown process itself.

In order to investigate this phenomenon, Appendix A.3 develops a simple analytical

model for calculating mixture and liquid level trajectories during the boildown of a

PWR U-tube SG under 'ideal' thermal-hydraulic conditions, following a complete loss

of feedwater. The principal assumptions made are as follows.

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1. All fluid on the secondary side is at saturation and the recirculation flow has

ceased.

2. The mixture level in the riser has fallen to at least the top of the U-tube bundle.

3. Any liquid entrainment above the mixture level is negligible and no liquid

condensate enters the downcomer.

4. The linear power is uniform below the mixture level and does not change with

time.

5. The SG pressure does not change with time and spatial variations are very small.

6. Liquid flow rates are small and a quasi-static approximation may be used for

vapour mass balances.

It is appreciated that the situation envisaged is somewhat idealised, particularly the

restriction to uniform heating conditions. Nevertheless, the general trends indicated

by the model should still be representative, at least qualitatively, of those prevailing

during SG boildown.

As shown in Appendix A.3, under the above conditions, the water level in the

downcomer, zd, may be approximated by the collapsed liquid level in the riser, zc,

which is related to the mixture level in the riser, z,, , by

1 ' [(C 0 - 1)zm + In{1 + BZm}] (27) CO" a CO

where

B= Co q (28)

Pg hgf Ar vgj

Here CO and v,, are the drift flux parameters (Zuber and Findlay, 1965), Ar is the riser

flow area, q' is the linear power and the rest of the notation is defined in the

Nomenclature list.

In addition, the riser mixture level satisfies the differential equation

dzm[ ~ 1 B ]I=-Dzm (29) -z- CO - I + 1 + 1B zmDz,(9

where

CO q' A (30)

pf hgf Ar I + "

which has the analytical solution

= exp( - D t) (31)

1 + Bzm 1 + B Zm0

where zO is the initial mixture level and Ad is the downcomer flow area.

Finally, the total liquid inventory in the SG is given by

Iv = (Ad + Ar) PfZc (32)

The above results show that for the simple boildown transient considered, the

mixture and liquid level trajectories depend on both the initial levels and the drift flux

parameters effective during the boildown process. This implies that in a code such

as RELAP5, the interphase drag models affect the prediction of boildown phenomena

through their effect on both the initial inventory calculation and the level dynamics.

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NP Restricted TECIL/0471/R91

Thus, under certain circumstances, attempts to artificially compensate for the

under-prediction of initial inventory without correcting the interphase models, may

not necessarily lead to the correct prediction of the level trajectories. This is most

likely to be true for fast boildown transients occurring under full power conditions.

If the boildown occurs under decay heat conditions, the errors in the initial inventory

calculation, which is generally established at full power conditions, are likely to

dominate (since deficiencies in the interphase drag models have a progressively

greater effect as the power is raised and the void fraction in the mixture region

becomes more significant). These conclusions are consistent with the results of

RELAP5/MOD2 assessment calculations.

Equation (31) also shows that the drift flux parameters have different roles in

determining the level trajectory. In particular, the exponential decay term only

depends on the distribution coefficient, CO, not the vapour drift velocity, v0i. This

suggests that in order to predict the correct level behaviour during boildown, it is

necessary to calculate both the profile and local slip components of the drag force

correctly. In contrast, in order to obtain the correct initial inventory, it is only

necessary to calculate the total drag force correctly.

Using appropriate definitions for the drift flux parameters and a suitable numerical

solver, equations (27) and (31) may be applied to 'emulate' the modelling of SG

boildown behaviour in RELAP5/MOD2 for the conditions considered, and to compare

the level trajectories obtained against those predicted by a reference model. This

approach was used to study level behaviour in a full size PWR U-tube SG for

conditions characteristic of boildown under full power conditions. The results of this

analysis are discussed in Appendix A.3.7 and illustrate many of the points made

above.

5.3.2 Implications for PWR safety analysis and compensation of deficiencies

The principal consequence of the direct effect of interphase drag on SG level

dynamics is that, for fast boildown transients, an under-prediction of the initial

inventory is not the only cause of inaccuracies in the calculated level behaviour. The

implications of such inaccuracies for PWR safety analysis have already been

discussed in Section 5.2.3, and the existence of this additional source of error does

not really make these implications any worse - particularly as it is a result of the

same fundamental deficiency. It does mean however, that attempts to increase the

initial inventory by artificially increasing the recirculation ratio or downcomer area

may not have the desired effect. As already observed though, the increases

necessary to achieve the correct inventory may riot be physically sensible.

It was noted in Section 5.2.4 that the interphase drag in the bundle region may be

'corrected' by setting a very large hydraulic diameter. This parameter however, only

affects the local drag force. The analysis described above indicates that in order to

obtain the correct level behaviour during boildown, it is necessary to calculate both

the local and profile slip components of the drag force correctly. Thus, imposing a

large hydraulic diameter may also fail to force the correct level behaviour.

For slow boildown transients typical of those occurring under decay heat conditions,

the interphase drag calculation will have less influence on the level dynamics, and

the errors in the initial inventory will dominate the predicted level behaviour.

5.3.3 RELAP51MOD3 considerations

As noted in Section 5.2.5, RELAP5/MOD3 incorporates improved interphase drag

models for bundle geometries and may therefore be expected to provide a better

prediction of level behaviour for fast boildown transients. However, in view of the

different roles played by the local and profile slip components of the drag force and

the importance of obtaining accurate estimates of each one, this type of problem

should provide a good test of the new interphase drag models in RELAP5/MOD3. In

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NP Restricted TECL/U0471/R 91

particular, Putney (1988) has observed that the split between the local and profile

components of the total slip may not be well known for bundle geometries.

One of the disadvantages of using results from integral test facilities to assess SG

modelling in codes such as RELAP5, is that deficiencies in a specific model are often

difficult to identify due to the mutual interaction between different phenomena and the

lack of detailed instrumentation. This is particularly true for the case of SG boildown.

It is therefore suggested that the modelling of this process in RELAP5/MOD3 be

assessed against the loss of feedwater tests carried out on the MB-2 facility. These

tests provide separate effects type data relevant to both fast and slow boildown

transients, as well as data appropriate to the refiling of the SG.

5.4 Wall heat transfer above mixture level

In the RELAP5/MOD2 analysis of LOFT loss-of-feedwater test L9-3 (Section 3.8.4), the

predicted primary to secondary heat transfer remained virtually constant during the

boildown phase down until the SG side was almost empty, at which point a sudden

drop in heat transfer occurred. In the experiment however, a gradual degradation in

SG heat transfer occurred as the level fell 4. Attempts to improve the representation

of SG heat transfer in the calculation by adopting a finer nodalisation at the bottom

of the bundle had little effect. Closer examination of the results indicated that during

boildown, the liquid fraction in the bundle nodes decreased more or less together,

so that instead of a sequentially emptying of the nodes, the liquid was smeared along

the riser. Evidently, this behaviour is indicative of an over-prediction of liquid

entrainment above the riser mixture level. As a result, wet wall conditions were

predicted to occur in all nodes until the SG was virtually empty, whereupon the heat

transfer fell sharply.

An over-prediction of wall heat transfer above the mixture level was also evident in

the RELAP5/MOD2 analyses of the boildown phases of the LOFT LP-FW-01 (Section

3.8.4) and LOBI ST-02 (Section 3.11.3) ioss-of-feedwater tests, and in the analysis of

the blowdown phase of LOBI steam line break test BT-12 (Section 3.11.5). Again an

over-prediction of liquid entrainment above the mixture level appeared to be the most

likely cause.

It is believed that this deficiency in the code is a consequence of the known

deficiencies in the interphase drag model, exacerbated by the absence of a proper

vertical stratified flow model and, possibly, the criteria applied to determine wet/dry

wall conditions. In particular:

* As discussed in Section 5.1.2, RELAP5/MOD2 over-predicts interphase drag in the

bubbly and slug regimes for bundle geometries. If this regime is present at the

mixture level, the liquid entrained above the level will be over-predicted.

* The RELAP5/MOD2 vertical stratification model (which may or may not have been

invoked in the transients noted above), does not alter the liquid fraction at the

junction immediately above the mixture level, which is still assumed to be equal

to the average liquid fraction in the node below (donor cell rule). This effectively

carries the liquid present in the node into the bottom of the node above. (The

same effect will clearly occur if the vertical stratification model is not invoked).

* How far the liquid entrained from the mixture region is carried along the riser

evidently depends on the interphase drag calculated in the nodes above the

mixture level - the higher the interphase drag, the further the liquid would be

expected to be carried. The interphase drag calculated for a node in which wet

wall conditions are predicted is likely to be greater than that for a node under dry

34 Since RELAPS/MOD2 does not explicitly recognise the presence of a mixture level in a node and

effectively assumes that the wall is wet until the node is virtually empty, the code would be expected

to reflect a gradual reduction in heat transfer with level by a step-wise degradation.

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INP Restricted TECIL/0471IR91

wall conditions. In RELAP5/MOD2 however, wet wall conditions are assumed to

occur unless the vapour fraction is above 0.9999 or the critical heat flux has been

exceeded. It is not obvious that the latter will always be true above the mixture

level in SG boildown or blowdown transients when significant entrainment is

present - i.e. the conditions could be such that the code automatically assumes

that any liquid present in a node that has a liquid fraction above 0.0001

completely wets the wall. Thus, a situation could arise in which an

over-prediction of liquid entrainment above the mixture not only results in wet

wall conditions being assumed and a consequent over-prediction of wall heat

transfer, but also creates the conditions for the entrainment to be carried over to

the node above.

The consequences of an over-prediction of SG heat transfer above a mixture level in

a bundle are that, prior to the primary and secondary sides becoming thermally

decoupled on the plant, the primary to secondary temperature difference may be

under-predicted. In addition, once the two sides have decoupled, the primary side

heat-up may not be calculated correctly. In both cases however, it is unlikely that the

implications for PWR safety analysis will be any worse than those already associated

with errors in predicted level behaviour itself.

Few options are available for compensating for the over-prediction of heat transfer

through changes to the input deck. Setting a very large hydraulic diameter in the

bundle region might help by reducing the interphase drag in the bubbly and slug

regimes, but it is not clear what effect this would have in other regimes.

Some improvement in the modelling of SG heat transfer above a mixture level may

be expected in RELAP5/MOD3 due to the new interphase drag and vertical

stratification models. It is suggested that this feature of the code may be assessed

when using the MB-2 Ioss-of-feedwater tests recommended in Section 5.3.3 to

examine the code's modelling of SG level behaviour. The results should help to

establish whether the criteria used to determine wet or dry wall conditions need to

be revised.

5.5 Dynamic level swell

A number of the RELAP5/MOD2 assessment studies reviewed in Section 3 revealed

deficiencies in the calculation of SG dynamic level shrink and swell; viz.

" For the Doel 2 SGTR accident (Section 3.4), RELAP5/MOD2 over-predicted the

downcomer water level swell that occurred in the intact SG when cooldown was

initiated by opening the SRV, and also when steam was discharged to the turbine

driven auxiliary feedwater pumps.

" For the Doel 4 loss of load test (Section 3.5), the code over-predicted the

downcomer level swell that occurred when the steam dump valves opened, and

under-predicted the rapid fall in level (shrink) following the sudden increase in

SG pressure when the turbine admission valves closed.

" In the analysis of the MB-2 100% steam line break test T-2013 (Section 3.9.2), the

initial level swell calculated by the code was too rapid while single phase steam

was being discharged through the break (causing the onset of two-phase break

flow to begin too early).

Although the code failed to predict the small level swell that was observed in the

Tihange 2 plant trip test following the manual opening of the steam dump valves

(Section 3.6), this was believed to be due to an under-prediction of the initial liquid

inventory.

Evidently, the level discrepancies noted above may be attributed to the tendency of

RELAP5/MOD2 to over-predict interphase drag for low flow conditions in bundle

geometries, which was fully discussed in Section 5.2.2. Following a depressurisation

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transient in a SG, this results in an over-prediction of void fraction in the bundle

region and leads to excessive mixture level swell being calculated in the riser. This

in turn causes an over-estimation of the liquid entrained to the separator, which then

falls back into the downcomer and the region where the DP measurement taps are

located - resulting in a high water level indication. Conversely, following a

repressurisation transient, the over-prediction of interphase drag results in an

under-prediction of the void fraction collapse in the bundle, and thus an

under-prediction of the downcomer level shrink.

The principal consequences of deficiencies in the modelling of SG level shrink and

swell are errors in the predicted actuation and operation of the plant protection and

control systems that depend on downcomer level. These are likely to have the

greatest effect when simulating normal operational transients (e.g. loss of load, plant

trips) and SGTR accidents, but should be less important for the more severe intact

primary circuit faults and LOCAs. Note however, that the feedwater systems may act

to compensate for an error in the predicted level behaviour, and this could introduce

errors in the predicted SG inventory.

As noted for other areas, the only way of forcing the code to calculate SG level shrink

and swell correctly, is to artificially reduce interphase drag in the bundle region by

specifying a very large hydraulic diameter. Again, however, this could be difficult to

justify. Also, the analysis of SG boildown transients presented in Appendix A.3

indicated that the profile and local slip components of the drag force played different

roles. This is also likely to be true for depressurisation and repressurisation

transients, and thus adjusting the hydraulic diameter, which only affects the local slip

component, may not necessarily have the desired effect.

The modelling of SG level shrink and swell may be expected to be more accurate in

RELAP5/MOD3 due to the improved interphase drag models - although the comments

made in Section 5.3.2 regarding the calculation of the individual components of the

drag force should be noted. Note also that when testing these models, Putney

(1989b) found that models that gave good agreement with data from tests involving

steam bubbling through stagnant water in large diameter vessels, did not perform

very well against a test involving the rapid blowdown of a 4 ft diameter vessel.

Further investigation indicated that the flow regime below the mixture level in the

blowdown experiment had not developed into the maximum sized cap bubble regime

that characterised the steam bubbling tests, but was closer to a classical churn

turbulent bubbly type regime. This apparent dependence of flow regime on

depressurisation rate is likely to be less evident in a SG bundle, as the narrow flow

channels may be expected to have a greater influence on constraining the flow

regime development. In addition, the depressurisation rates typical during steam

discharge from a SG relief valve (or a steam line break transient), are somewhat less

than that present in the blowdown experiment considered (73 to 20 bar in 20 s).

For plant safety studies, an accurate prediction of dynamic level behaviour is

particularly important for the range of transients typical of those that could be

imposed by the operation of the SG relief or turbine valves. Unfortunately, no

separate effects type tests have been identified that are suitable for assessing

RELAP5/MOD3 in this area. The Doel 2 SGTR accident and Doel 4 loss of load test

would therefore appear to provide the most appropriate data, although being plant

transients the quality and quantity of these data is not particularly good.

5.6 Liquid carryover during steam line break transients

Section 3 of this report discussed results from RELAP5/MOD2 analyses of 100%

steam line break tests performed on the MB-2 (Section 3.9.2), LOBI (Section 3.11.5)

and Semiscale (Section 3.13.2) facilities. In each case, an over-prediction of the liquid

carryover to the break was observed. This was initially attributed to shortcomings in

the interphase drag models; but further analysis of the MB-2 test using a pre-release

version of RELAP5/MOD2 has indicated that this is not the case. The most likely

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NP Restricted TECIL/0471/R91

explanation is that liquid carryover to the break may have been reduced in the

experiment by the accumulation of liquid on structures in the upper part of the SG -

an effect that was not represented in the RELAP5 analyses. In the MB-2 and

Semiscale tests, some of this liquid may have drained back into the top of the riser

later in the tests, and been responsible for a period of good heat transfer that was

not captured in the calculations.

The main consequence of an over-prediction of liquid carryover in a steam line break

transient is an under-prediction of the primary side cooldown 35 (although for the LOBI

test, the overall primary side was actually over-predicted due to an under-estimation

of the metalwork heat). Under certain circumstances, this might result in the code

failing to predict a return to power after a steam line break due to excessive

overcooling and, as such, is therefore nonconservative.

Preliminary RELAP5/MOD3 analysis of the MB-2 steam line break indicates no

significant improvement in the modelling of liquid carryover to the break. Indeed if the

smaller carryover in the tests is due to the accumulation of liquid on structures in the

upper SG, it is difficult to see how this effect could be represented without introducing

a special model.

5.7 Separator performance

Several RELAP5/MOD2 analysts have reported difficulties in obtaining satisfactory

steady-state conditions in SG models due to instabilities in the separator component.

Typical problems encountered are the periodic emptying and filling of the separator

volume and the steam carryunder to the downcomer. As explained in Section 3.8.2,

these problems are probably caused by an over-prediction of the void fraction at the

top of the riser, and may therefore be attributed to the known deficiencies in the

interphase drag models applied in the bundle region. In fact, many analysts have

removed the instabilities by artificially increasing the recirculation ratio, which as

shown in Appendix A.2 has the effect of reducing the void fraction at the riser outlet.

Since this change also effects the DP level measurement and the SG inventory (albeit

in the right direction), a more acceptable approach may be to specify a very small

value for the carryunder parameter VUNDER.

Some analysts have also -seen instabilities in separator performance during the

analysis of SG transients. It has also been noted that results can be sensitive to the

separator nodalisation, particularly the value of loss coefficient at the liquid fall back

junction. Whilst some of these problems may well be a result of the interphase drag

deficiencies, others are more likely to be due to the non-mechanistic nature of the

separator model.

As far as can be ascertained, the deficiencies in the RELAP5/MOD2 modelling of the

SG.separation process do not appear to hold serious implications for PWR safety

analysis. In general, either the errors produced are relatively minor, or can be

avoided by appropriate adjustments to the input deck.

It is expected that RELAP5/MOD3 will provide a better prediction of separator

performance due to the new interphase drag models. Although additional

improvements could probably be achieved by the development of a more mechanistic

separator model, this is not recommended as a high priority task.

35 In theory, the calculation of a low initial SG inventory could also contribute to an under-prediction of

the primary side cooldown. However, this effect was not present in the tests examined as they were initiated from very low power conditions.

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NP Restricted TECIL/0471IR91

5.8 CCFL and liquid hold-up on primary side

A number of RELAP5/MOD2 analyses of LOCA tests have indicated a tendency for the

code to predict excessive liquid hold-up in the SG U-tubes or inlet plenum:

In the UK analysis of LOFT test L3-6 (2.5% cold leg break with delayed pump trip

- Section 3.8.3), the SG U-tubes drained significantly later than they did in the

experiment. This resulted in a late prediction for the decoupling of the primary

and secondary sides, causing the secondary side pressure to be systematically

under-estimated. A similar effect was apparent in the Studsvik analysis of the

test and in the KAERI analysis of test L2-5 (200% cold leg break).

For the ROSA-IV 5% cold leg break test (Section 3.7), the code predicted a period

of liquid-up in the SG inlet plena that was observed in the experiment. Use of a

modified version of RELAP5/MOD2 which allowed for the existence of stratified

flow in the hot leg riser improved the calculation, but the time of draining was still

over-predicted - leading to an under-prediction of the core liquid inventory during

the boildown phase.

Excessive liquid hold-up in the SG U-tubes was calculated in the analysis of LOBI

SGTR test BL-21 (Section 3.11.4). The analyst attributed this discrepancy to an

over-prediction of interphase drag in the vertical section of the hot leg preventing

liquid from draining back (but gave no evidence to support this view).

This behaviour of the code indicates a tendency to under-predict the vapour flow at

which CCFL occurs at some point in the pipework between the hot inlet and the top

of the SG U-tubes, which may be attributed to an over-prediction of the local

interphase drag force. It is also consistent with an analysis of UPTF test 11 performed

by Dillistone (1990), in which RELAP5/MOD2 was found to under-estimate the steam

flow necessary to prevent liquid drainback from the SG inlet plenum by more than a

factor of 3. The author shows that this is primarily due to the use of an over-simplified

flow regime map which does not allow the possibility of stratified flow in the hot leg

riser. As noted above, removing this restriction also helped to improve the results in

the ROSA-IV test analysis.

Other deficiencies in the modelling of CCFL and liquid-up on the SG primary side

evident from the results of RELAP5/MOD2 assessment studies, that are not totally

consistent with the above, are as follows.

* In the analysis of LOBI test BL-02 (3% cold leg break - Section 3.11.2), there was

some evidence that the code was calculating excessive draining from the SG inlet

plena. Further investigations indicated that this was due to poor modelling of

CCFL along the horizontal part of the hot leg.

* In the analysis of the SPES loss-of-feedwater test (Section 3.10), the primary side

natural circulation mode following core recovery was incorrectly predicted due

to shortcomings in the modelling of CCFL in the SG U-tubes. The failure to

calculate CCFL in the U-tubes was also evident in the analysis of LOBI SGTR test

BL-21.

* Errors were present in the RELAP5 predictions of liquid hold-up in the SG tubes

for the Semiscale small break LOCA tests S-LH-1 and S-LH-2 (Section 3.13.1),

although no clear systematic trends were evident.

These results reflect a general problem in modelling flooding type phenomena in

RELAP5/MOD2 for the range of geometries present in the PWR. Some improvement

may be expected with RELAP5/MOD3 as a result of the new interphase drag and user

controlled CCFL models.

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NP Restricted TECILI0471/R91

5.9 Secondary side condensation effects during repressurisation

transients

Two instances have been uncovered where RELAP5/MOD2 appears to have

over-estimated vapour condensation rates on the SG secondary side under

repressurisation conditions:

1. In the analysis of the Doel 2 SGTR accident (Section 3.4), RELAP5/MOD2 failed to

calculate the increase in secondary side pressure that occurred after the faulted

SG had been completely isolated, and resulted from the piston effect of the rising

level on the vapour in the steam dome. This discrepancy was traced to the

calculation of an excessive vapour condensation rate in the dome, and was

believed to be due to the use of constitutive equations appropriate to slug flow

under stratified flow conditions. Further calculations however, showed that an

increase in pressure could be obtained by subdividing the steam dome into

several volumes.

2. In the analysis of the Tihange 2 plant trip test (Section 3.6), the SG pressure was

significantly under-predicted in the period immediately following the turbine and

reactor trips when the pressure was increasing rapidly (although the final

pressure was well predicted). Further investigations indicated that this

discrepancy was due to excessive condensation being calculated in the regions

where the feedwater was injected, which had the effect of driving the vapour to

saturation, leading to a temporary stagnation in the pressure evolution. Since the

plant data indicated a similar pressure plateau but at a higher pressure, the

analysts deduced that the code's bubbly flow models cannot maintain strong

thermal disequilibrium under fast pressurisation conditions, and that due to

premature condensation the vapour temperature returns to quasi saturation

conditions too quickly.

These two examples of the calculation of excessive vapour condensation are

evidently caused by different modelling deficiencies. In addition, the effects seen

have not been observed in any other studies and thus it is not possible to say

whether they represent general trends. Nevertheless, they are recorded here for the

sake of completeness.

6 CONCLUSIONS

1. 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 is 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. For the Sizewell 'B' SG

operating at full load conditions, the error in secondary side pressure is around

3.5 bar. The deficiency can be attributed to the application of the 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. Although 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. Various methods are available for compensating for the

deficiency, but in most cases the most defensible approach is to artificially

increase the U-tube thermal conductivity. In general, the errors in SG heat

transfer behaviour do not hold serious implications for PWR safety analysis,

although they are likely to be worse in RELAP5/MOD3 as the Chen correlation is

applied in its standard form.

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NP Restricted TECILIO471/R91

2. At steady-state normal operating conditions, RELAP5/MOD2 under-predicts the

SG secondary side liquid inventory if the recirculation ratio is calculated

correctly. The deficiency can be attributed to an over-prediction of void fraction

in the SG bundle, caused by an over-prediction of the interphase drag force

(which is a known deficiency in the code). The error in inventory that results,

which tends to reduce as reactor power reduces, affects the prediction of the

downcomer level following a reactor trip and the calculated time of SG dryout in

loss of feedwater accidents. The latter holds the more serious implications for

PWR safety analysis and is particularly severe for transients in which most of the

boildown occurs under decay heat levels - as the excess inventory on the plant

takes a long time to boiloff. As a result, the predicted dryout may be advanced

by around 1000 s, which has a dramatic affect on the timings of all subsequent

events. Although the initial inventory may be artificially increased by increasing

the downcomer flow area or the recirculation ratio (through raising the initial SG

level or reducing the flow losses), the adjustments required to compensate for

the void fraction errors may not be physically sensible. The initial inventory can

also be 'corrected' by specifying a very large hydraulic diameter on the

secondary side of the U-tube bundle (t-250 cm). However, this may be difficult to

justify. The calculation of SG inventory is expected to be better in RELAP5/MOD3

due to the new interphase drag models.

3. For fast SG boildown transients typical of those occurring under full power

conditions, RELAP5/MOD2 tends to over-predict the rate at which the downcomer

level falls, in addition to under-predicting the initial inventory. Moreover,

increasing the initial inventory by increasing the downcomer flow area or

recirculation ratio, does not have a significant effect on this rate - it only serves

to delay the time at which SG dryout occurs. These effects arise because, in

addition to determining the initial inventory, the interphase drag models in the

bundle have a direct effect on the mixture and downcomer level trajectories

during the boildown process itself. The effect is less noticeable for slow boildown

transients where the vapour generation rate is relatively low, as the errors in

initial inventory tend to dominate. Theoretical analyses also indicate that in order

to predict the correct level behaviour during boildown, it is necessary to calculate

both the local and profile slip components of the drag force correctly. In contrast,

in order to obtain the correct initial inventory, it is only necessary to calculate the

total drag force correctly. This implies that attempts to correct the initial inventory

by specifying a large secondary side hydraulic diameter, which only affects the

calculation of the local drag force, may not necessarily produce the correct level

behaviour. It also means that this type of problem should provide a good test for

the new interphase drag models in RELAP5/MOD3.

4. During SG boildown and blowdown transients, RELAP5/MOD2 tends to

over-predict wall heat transfer above the mixture level. The evidence available

indicates that the code predicts wet wall conditions for the whole of the bundle

until the SG has virtually dried out, and that this is due to the calculation of

excessive liquid entrainment from the mixture region. It is believed that this

problem is a consequence of the known deficiencies in the bundle interphase

drag model, exacerbated by the absence of a proper vertical stratified flow model

and, possibly, the criteria applied to determine wet/dry wall conditions. Some

improvement may be expected in RELAP5/MOD3 due to the new interphase drag

and vertical stratification models.

5. The deficiencies in the bundle interphase drag model are believed to be

responsible for a tendency of RELAP5/MOD2 to calculate excessive dynamic level

swell in response to steam discharge from the SG. They may also be the cause

of instabilities seen in the calculated separator performance, although the

non-mechanistic nature of the separator model may also'be a contributing factor.

In the past, a tendency to over-predict liquid entrainment to the break in steam

line break transients have been attributed to shortcoming in interphase drag

modelling. Recent studies however, have indicated that this is not the case, and

the most likely explanation is the failure to model the accumulation of liquid on

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NP Restricted TECILI0471/1R91

structures in the upper SG. Other deficiencies in SG modelling in the code

include the modelling of counter current flow limitation (CCFL) and liquid hold-up

in the U-tubes and inlet plena. These reflect a general problem in modelling

flooding type phenomena in RELAP5/MOD2 over a range of geometries and may

be linked to the interphase drag models and flow regime maps - although in many

cases the primary source of the deficiency lies outside the area of SG modelling.

Nevertheless, some improvement may be expected in RELAP5/MOD3, primarily

as a result of the new user controlled CCFL model.

6. In practice, the above deficiencies in SG modelling do not diminish the value of

RELAP5/MOD2 for performing safety case calculations for the majority of PWR

LOCAs and intact primary fault sequences of interest. The only real exception

concerns the modelling of transients in which the SG boils dry very slowly, when

the under-prediction of initial inventory will result in the SG being calculated to

dry out too early. A typical example is a total loss of feedwater accident with

normal reactor trip and closure of the MSIV. Due to the inaccuracies in the timing

of major events, it may be difficult to use the results of the calculation for safety

assessment purposes.

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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

Keevill M B, 1988, RELAP5/MOD2 analysis of LOFT experiment L9-4,

PWR/PKWG/P(88)372

Keevill M B, 1989, RELAP5/MOD2 analysis of a commissioning test involving a 50%

load reduction from 100% power in the Woif Creek PWR, GD/P-EN/759,

PWR/PKWG/P(89)421

Klimenko V V, 1988, A generalized correlation for two-phase forced flow heat transfer,

Int. J. Heat Mass Transfer, Vol. 31, No. 3, pp. 541-552

Kumamaru H, Koizumi Y and Tasaka K, 1987, Investigation of pre- and post-dryout

heat-transfer of steam-water two-phase flow in a rod bundle, Nucl. Eng. Des., Vol.

102, pp. 71-84

- 69 -
NP Restricted TEC/L/0471/R91

Lee E-J, Chung B-D and Kim H-J, 1990, ICAP assessment of RELAP5/MOD2, cycle

36.05 against LOFT small break experiment L3-7, NUREG/IA-0031, April

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

Model 'F' steam generator from plant measurements, CERL report no.

TPRD/L/3079/R86

Preece R J, 1987b, Uncertainty bands applying to the Westinghouse Model 'F' steam

generator thermal-hydraulic performance measurements, CERL report no.

TPRDIL/3151/R87

Proctor A R, 1990a, RELAP5/MOD2 analysis of a commissioning test involving a plant

trip from 100% power in the Wolf Creek PWR, TD/SPB/REP/0089

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/HTWGIP(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/HTWGIP(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

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Rogers J M, 1989a, An analysis of Semiscale Mod-2C S-FS-1 steam line break test

using RELAP5/MOD2. AEEW-R2476, March

Rogers J M, 1989b, An analysis of MB-2 100% steam line break test T-2013 using

RELAP5/MOD2. AEEW-R2477, March

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NP Restricted TEC/L/0471/R91

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systems, J. Heat Trans., Vol 87, pp. 453-468

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NP Restricted TECIL/0471/R91

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 1D (channel averaged) flow. SI units are assumed

throughout.

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

ak, Volume fraction (a =a- )

Pk Density

Pks Density at saturation conditions

Vk Specific volume

V kj Specific volume at saturation conditions

hk Specific enthalpy

hk. Specific enthalpy at saturation conditions

p Pressure (equal for both phases)

Tk Temperature

T, Saturation temperature

Tý Wall temperature

A .T., = T, - T,

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

pressure at T,

q" . Wall heat flux

/Ak Dynamic viscosity

k, Thermal conductivity

CP, Specific heat at constant pressure

Pr , Prandtl number (= C,1, Pk / kk)

vk Velocity

j , Superficial velocity (j, = Cav,)

p Mixture density (p = cp. + ap,)

G Mixture mass flux (G = p~j, + prf)

x Quality

rg Vapour generation rate per unit volume

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NP Restricted TEC/L/0471/R91

Ap Phase density difference (Ap = pf - p9)

h91 Latent heat of vaporisation (hif = hg, - hr,)

Vr Relative velocity (vr = v. - VF

g Acceleration of gravity

a Surface tension

vgj Vapour drift velocity (in drift flux model)

C, Distribution coefficient (in drift flux model)

A Flow area

D. Equivalent heated diameter

- 73 -
NP Restricted TECILI04711IR 91

Appendix A CONCEPTUAL STEAM GENERATOR MODELS

This appendix develops simple, analytical models of some of the main phenomena

which control SG performance during normal and abnormal operating conditions.

The particular areas addressed are

* Primary to secondary heat transfer.

" Steady-state operation and effect of recirculation flow.

" Steam generator boildown.

Results from these models are used in the main body of this report.

Although the models developed are very simple and somewhat idealistic, they are

still believed to be representative. In the context of the present study, they are

particularly useful for interpreting the results of RELAP5 analyses and for identifying

modelling deficiencies responsible for apparent code deficiencies.

A.1 Primary to secondary heat transfer

Consider a short length, 6L, of a single SG tube having internal radius Rp, wall

thickness t and thermal conductivity k. If the wall and bulk fluid temperatures on the

primary side of the tube are denoted by T,; and Tp, the heat flux at the wall, q"p may

be written as

q"p = Hp (Tp - Twp) (A1.1)

where H, is the primary side heat transfer coefficient.

Likewise, the wall heat flux on the secondary side of the tube is given by

q"s = Hs (Tws - Ts) (A1.2)

Under steady-state conditions, continuity arguments demand that

27RpR6Lq"p = 27r(Rp + t) 6Lq"s = Q (A1.3)

where Q is the total heat transfer rate across the length of tube (in W m-1).

Integrating the steady-state heat conduction equation across the tube wall and

neglecting effects due to axial conduction and the variation of tube thermal

conductivity with temperature, leads to

Tws- Twp - Rpq"p In 1 + - (A1.4)

Since t / RP is small, the In term may be approximated by the first term in its Taylor

series expansion, giving

= -- q,(A1.5) Tws - Twp =-.-F q p (15

Combining equations (A1.1), (A1.2), (A1.3) and (A1.5) to eliminate q"p , q"s, TwP and

Tws, results in the relationship

27r Rp 6L (Tp - Ts)

SRP e = (A1.6) 1 + 1 +t

Rp +t I H 5 H k k

This equation shows how various SG parameters affect the heat transfer across a

short length of tube. It is used in Section 2.3 to demonstrate the sensitivity of SG heat

transfer to changes in these parameters.

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NP Restricted TECILIO471/R9I

A.2 Steady-state operation and effect of recirculation flow

A.2.1 Conceptual model

At the simplest level, the secondary side of a PWR SG may be modelled using three

basic components: Riser, Separator and Downcomer. Under ideal normal operating

conditions, the thermal-hydraulic conditions in each component may be described as

follows.

Riser

Subcooled (or saturated) water enters the bottom of the riser, boils as it flows

upwards through the U-tube bundle, producing a saturated two-phase mixture

which enters the separator.

Separator

Assuming perfect separation (no water carryover or steam carryunder), saturated

steam passes from the separator to the SG steam dome and saturated water

returns to the top of the downcomer (via the recirculation flow).

Downcomer

In the downcomer, the recirculation flow mixes uniformly with the feedwater flow,

which is sufficienf to maintain a constant water level.

The situation described is shown schematically in Fig. A.1. The notation used is as

follows.

WF Feedwater flow rate (kg s 1)

Wc Recirculation flow rate

WF1 Total riser flow rate

W, Steam outlet flow rate

h,, Saturated liquid enthalpy (J kg'')

h•. Saturated vapour enthalpy

hF Feedwater enthalpy

hR, Riser inlet enthalpy

XRo Equilibrium quality at riser outlet

O(RO Void fraction at riser outlet

Q Total primary to secondary heat transfer rate (W)

q' (z) Linear power in tube bundle at elevation z (W m 1)

4 (z) Normalised axial heat flux profile in bundle at elevation z

ZDC Elevation of water level in downcomer (m)

Z..I Elevation at beginning of saturated boiling

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NP Restricted TEC/L/0471/R9g

Ze Elevation at top of tube bundle

ZRO Elevation at riser outlet

ZsP Elevation at top of separator

AR Flow area of riser

KRI Effective loss coefficient at bottom of riser

Other notation used in this appendix is defined in the Nomenclature list.

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NP Restricted TECIL/O4711R91

zSP -

Seperator z DC

xRO cRO hs•h

ZRO•-

- Downcomer

Riser

Q -4-*

Zsat

m WR • •Z=O

W R

hRI KRI

Figure Al Conceptual model of SG secondary side

JMP/CR(29.8.91)047I/A1

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NIP Restricted TECIL/04711R91

A.2.2 Steady-state conditions

Imposing steady-state mass conservation for the model described above gives

WS = WF (A2.1)

and

WR = WF + WC (A2.2)

i.e. the feedwater flow must be equal to the steam flow and the total riser flow must

equal the feedwater flow plus the recirculation flow.

From an energy balance across the downcomer, the liquid enthalpy at the riser inlet

is given by

Wc hfs + WF hF

hR; = WC + WF (A2.3)

while an energy balance across the riser and separator gives

Q

WF = Q (A2.4) hg5 -lip

which is a condition for steady-state operation.

The mass flowing quality at the riser outlet is defined by

WC+WFXRO -(A2.5) = WF 

To model the recirculation flow, it is convenient to represent all the frictional

resistance around the recirculation loop using an effective loss coefficient, KF1, defined at the junction between the downcomer and riser. Then, ignoring acceleration

and inertial terms, and the vapour head between Zoc and Zsp in the downcomer, a

momentum balance around the loop gives

f ZD ZRO 2S

Prgdz = J pgdz + J pgdz + KRIWR (A2.6)

*'0 ZRO

In the above equations, WF/Q and hF are usually specified as boundary conditions for

a RELAP5 simulation. The recirculation flow, Wc, however, is sometimes specified independently from the plant value (if known) to achieve the desired separator

performance and/or the correct SG liquid inventory. The following sections show

how these effects come about.

A.2.3 Effect of recirculation flow on RELAP5 separator performance

Consider the case when Q, WF and hF are specified and satisfy equation (A2.4).

Equation (A2.6) shows that an increase in the recirculation flow may be obtained by

raising the liquid level in the downcomer or by artificially reducing the frictional

losses in the recirculation loop (by specifying a low value of KRI)". The equation also

shows that an over-prediction of interphase drag and thus voidage in the riser will

demand a lower SG level or a greater friction loss to achieve a particular recirculation

flow.

The liquid enthalpy at the riser inlet (equation (A2.3)) may be expressed as

3s In Appendix (A2.4), it is shown that an increase in the recirculation flow will lead to an increase in the

liquid fraction in the riser and separator regions. This will result in an increase in the local mixture

density, which will tend to check the initial increase in mixture level.

- 78 -
NP Restricted TECIL/0471/R91

hF

hRI = hfs Lf + ± (A2.7) + WF hfs

Under normal operating conditions hF < h,5 (i.e. both terms in equation (A2.7) will be

positive), and thus an increase in the recirculation flow will result in an increase in

the riser inlet enthalpy. (If hF = hr,, hR, = hf, for all values of Wc.)

Equation (A2.4) on the other hand, shows that an increase in the recirculation flow

will result in a decrease in the riser outlet quality. Furthermore, from the drift flux

model (Zuber and Findlay, 1965), the void fraction at the riser outlet is given by

[ XRO WC + WF

OCRO = 1X O1- +g WR 1 (A2.8)

C I[gs-- XRO 1 -Pfs XRO+ vg I[ AR 4 g

which after eliminating xto using equation (A2.4) may be written as

C(RO = [ Vgs + giAR (A2.9)

+ Vfs WF

Assuming that CO and v0j are not strongly dependent on Wc or a, this result shows that

an increase in the recirculation flow will result in a decrease in the void fraction at the

riser outlet.

The average void fraction in the SG separator is approximately equal to the void

fraction at the riser outlet. In RELAP5/MOD2, the separation process is represented

using the following non-mechanistic model:

0 The vapour fraction of the outlet 'steam' flow is set to one if the separator vapour

fraction is greater than a user supplied parameter VOVER (i.e. no carryover), but

ramped to zero if the separator vapour fraction falls below this.

* The liquid fraction of the outlet 'water' flow is set to one if the separator liquid

fraction is greater than a user supplied parameter VUNDER (i.e. no carryunder),

but ramped to zero if the separator liquid fraction falls below this.

Consequently, for a given value of VUNDER which is greater than zero, RELAP5 may

predict vapour carryunder to the downcomer if the void fraction calculated at the riser

outlet is sufficiently high. The above analysis shows that this carryunder may be

eradicated by increasing the recirculation flow, either by raising the liquid level in the

downcomer or reducing the friction losses in the recirculation loop.

Conversely, any liquid carryover to the steam dome may be removed by reducing the

recirculation flow.

A.2.4 Effect of recirculation flow on SG inventory

In the following analysis, the effects of subcooled boiling are ignored and it is

assumed that no vapour is present in the riser until after the bulk fluid has become

saturated at elevation Z,.

Denoting the height of the tube bundle as ZR, the linear power (W m`) at elevation z

may be written as

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NP Restricted TECIL/04711R91

q' (z) = -2- 46 (z) (A2.10)

where 0 (z) is the axial heat flux profile.

Integrating the steady-state energy equation between the riser inlet and the

beginning of saturated boiling gives

WF(hrs - hF) = --a- 0 (z)dz (A2.1 1)

or, from equation (A2.4)

(hf --hF) -- hgs -- hr f'Za,

(hfo - hF) = (z) dz (A2.12)

Since any dependence of k (z) on the recirculation flow is likely to be small, these

equations show that, to a first approximation, the beginning of saturated boiling is not

affected by changes in the recirculation flow.

Thus, the region of the SG riser over which changes in the liquid fraction may result

from variations in the recirculation flow, viz. Z,; to Ze, is independent of the

recirculation flow itself. Integrating the steady-state energy equation over this region,

and neglecting variations of the fluid properties with pressure (which are very small),

gives

WF

X (z) + F ( (A2.13)

where

(I)(z = hgs --hF "/[ r

D(z) hg= Z J (z) dz (A2.14)

Substituting this result in the drift flux equations and rearranging then gives

c (z) = Vgs (A2.15)

Co vgs + vrs WC + WF + R W

Assuming again that any dependence of 0 (z), C0 and v,1 on Wc or a are small,

equation (A2.15) indicates that an increase in the recirculation flow results in a

decrease in the local void fraction at all elevations above Z,,, (including the unheated

region of the riser above the top of the tube bundle). Since Z,•, was shown to be

largely unaffected by changes in the recirculation flow, this implies that, for a fixed

water level in the downcomer, an increase in recirculation flow will result in an

increase in the total SG inventory. Of course, if the increase in recirculation flow is

obtained by raising the downcomer level (rather than reducing the frictional losses),

the increase in inventory will be even greater.

It is appreciated that the above analysis is rather crude in that it completely neglects

the effects of subcooled boiling and makes a number of assumptions regarding the

dependence of certain parameters on the recirculation flow and local void fraction.

Also, no account was taken of the affect of changes in the recirculation flow on the

secondary side heat transfer coefficient and thus pressure. Nevertheless, none of

these effects are likely to alter the overall conclusion regarding the relationship

between recirculation ratio and SG inventory.

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NP Restricted TEC/L0471/R91

A.3 Steam generator boildown

This section presents an analytical model for calculating the mixture and collapsed

liquid level trajectories during the boildown of a PWR SG. Although the derivation

of the model equations involves a number of simplifying assumptions, the result is

sufficiently representative to allow the key parameters to be identified and for

carrying out sensitivity calculations.

The model presented is based on a two-phase level theory developed by Sun, Duffey

and Peng (1981). However, in order to obtain explicit analytical solutions for the

mixture and collapsed liquid level trajectories, the authors introduced a number of

assumptions that are not valid under SG boildown conditions. The present model

avoids these assumptions at the expense of obtaining analytical solutions that are not

explicit, although they are relatively easy to evaluate using a standard numerical

solver.

A.3.1 Modelling assumptions

The boildown model developed here applies to the situation following a complete

loss of feedwater in a PWR U-tube SG and when the following conditions exist

" The liquid in the system is at saturation and the recirculation flow has ceased (so

that the liquid flow from the downcomer to the riser is very low).

* The mixture level in the riser has fallen to at least the top of the U-tubes (i.e. the

fluid in the riser is heated at all elevations).

The situation envisaged is shown schematically in Fig. A.2, where the symbols used

are defined as follows:

z Elevation

Zd Downcomer level

z,_ Riser mixture level

zc Riser collapsed liquid level

z. Riser equilibrium level2 7

Ad Downcomer flow area

A, Riser flow area

q' Bundle linear power (W m-')

WfJ Riser inlet liquid mass flow rate (kg s-1)

MAO Riser outlet vapour mass flow rate

Other notation used in this appendix is defined in the Nomenclature list.

37 The equilibrium level is defined as the elevation at which the vapour flow rate equals the inlet liquid

flow rate, and above which the liquid flow rate is zero. In practice, the equilibrium level is the lowest

level of uncovery for a given inlet flow and subcooling.

-81 -
NP Restricted TEC/L/0471/R91

Wg, out

9

000

Heater riser

Zm

Downcomer 0

0

0

0

0 0 0

0 0

0

0

0 0

0

Z e

zC U

0 00

Zd

0

0 0 0

wf. i.l

Ar Ad

Figure A.2. Simple model of SG boildown

JMP/CPR(S.7.91)047 I/A2

- 82-
NP Restricted TECILI0471IR91

The principal assumptions made in the model are as follows

* Thermal equilibrium conditions exist below the mixture level.

* The SG pressure does not change with time and spatial variations may be

neglected in the calculation of fluid properties.

* Any liquid entrainment above the mixture level is negligible and no liquid

condensate enters the downcomer.

* The drift flux parameters Co and v9j are independent of flow rate and void fraction.

" The boildown is sufficiently slow that a , L, so that the vapour volumetric

flow may be obtained from a quasi-stat•t vapodT mass balance.

• Flow rates are sufficiently small for the liquid flow rate to be ignored when

calculating void fraction from the drift flux equation, and for momentum balances

to be dominated by gravitational forces.

• The linear power is uniform below the mixture level and does not change with

time.

It is appreciated that the restriction to uniform heating conditions is somewhat

idealised and is not particularly representative of the boildown of a PWR SG.

However, the general trends indicated by the model are still likely to be appropriate.

A.3.2 Vapour generation rate

In the riser region, the enthalpy of the fluid below the mixture level satisfies the

equation

a {a h8 + afpf hf) + {pv h + af p, vf hf}

aFt g Pgh + + azph~ {pgvhG 9 + p~rt

_ p + q' (A3.1)

at Ar

where the small contributions due to viscous dissipation and the j 8p / az term have

been ignored. Expanding the derivatives on the left hand side and introducing the

vapour and liquid mass balance equations:

a a (32 - {Pg} g + a7{%Pg pg Vg) =-g (A3.2)

af () + az T(A3.3) g 

then gives {099 ahg ahg I + {, vh} a +A34

atPg t + Vg + a t +A3.4 ± A 3.4)

+ (hg - h,) 1-'g

Under thermal equilibrium conditions and the constant and uniform pressure

assumptions, this equation reduces to the following expression for the vapour

generation rate

rg -

(A3.5) hgf Ar

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NP Restricted TECIL/0471/R91

A.3.3 Vapour volumetric flux equation

For slow boildown transients

a-' a." (A3.6)

so that the vapour mass balance equation (A3.2) may be approximated by the

quasi-static expression

apg jg

8z Fg (A3.7)

Introducing equation (A3.5) and noting that the modelling assumptions imply uniform

properties, leads to an equation for the vapour volumetric flux of the form

=g (Z) q' z (A3.8)

ig ( ' -- Pg hgt Ar

A.3.4 Void fraction equation

From the drift flux model (Zuber and Findlay, 1965), the void fraction below the

mixture level is given by

a - g (A3.9)

•g ýC 0 j + Vgj

For low liquid flow conditions typical of those encountered during SG boildown, it is

reasonable to set j, = 0 in the above equation" to obtain

=g = jg + (A3.10) CO ig +4*g

A.3.5 Level and inventory equations

The collapsed liquid level in the riser, z,, is usually defined as the equivalent

hydrostatic head; i.e.

jZmpgdz = Jprgdz (A3.11)

Under the modelling assumptions in use, this reduces to

zc = 1 - 0. + a _ýg dz

f - - +o ag P(A3.12) dz 

SJ"(1 - ag) dz

as p, / p, is small for the conditions of interest39 . For a uniform flow channel, the latter

equation also represents the liquid column height that would result if all the vapour

were removed.

The downcomer and riser level are related by system momentum effects. However.

since flow rates considered are small, it is reasonable to neglect the inertial,

acceleration and loss terms. Integrating the momentum equations over the

downcomer and riser mixture regions then gives

38 This will always be true above the equilibrium level.

39 pg / pr = 0.05 at 70 bar pressure and reduces with decreasing pressure.

- 84-
NP Restricted TECILIO4711R91

fod jZdpfgdz = fpgdz (A3.13)

which reduces to

Zd = zC (A3.14)

i.e. the collapsed liquid levels in downcomer and riser are equal.

The total liquid inventory in the system is defined by

IV = J pfAddz + Jo (1 - ag)pfAr dz (A3.15)

=(Ad + Ar) pfZc

from equation (A3.14)

An equation describing the transient level behaviour may be obtained from the total

mass balance equations for the downcomer and mixture regions; viz.

d 'zd d f pf Ad dz = -Wi, in (A3.16)

for the downcomer and

d Z"z

-d-t fpArdz Wt',in - Wgout (A3.17)

for the mixture region. Using equations (A3.11) and (A3.14), these equation may be

combined to give

+ A 1 dzc -Ad- Wg'out (A3.18) [ Ar] I tpfAr

Finally, introducing the volumetric flux equation (A3.8) leads to the level differential

equation:

7Ad dzC q-Zm (A3.19) 1 A-r I --- pg hgf Ar

A.3.6 Model solution

The final boildown model derived above is defined by equations (A3.19), (A3.12),

(A3.9) and (A3.8), which are repeated below for convenience

+ Adl dzC q'Zm (A3.19)

[+ -7J.. -- pghgfAr

Z':= f(- ag) dz (A3.12)

S jig (A3.9)

- + Vg

jg (z) - pg hgtAr (A3.8)

These equations constitute a system of four equations in four unknowns:

ZC, z.,, ag and j, . To obtain solutions for zc and z.,, it is necessary to eliminate the

parameters j, and •,. This can be achieved by combining equation (A3.8) and (A3.9)

- 85-
NP Restricted TECIL/04711R91

to obtain an expression for a, as a function of z, and then inserting the result into

equation (A3.12). Integration then leads to the following relation between z, and z,

1( - 1)z, 1 Inf1 + BZm}] (A3.20) zc -- CCO -B)mt BC°

where

q'

B = Co (A3.21) Pg hgf Ar Vgi

and the drift flux parameters have been assumed to be independent of flow rate and

void fraction.

Taking the derivative of this equation with respect to time and introducing equation

(A3.19) to eliminate z,, leads to the following expression for z,

dzm[ [ - +1 B I = Dzm (A3.22) dt Io- + B ZM

where

D = (A3.22)

pf hgf Ar 1 + ± d

Equation (A3.22) has the analytical solution

M -- exp( - D t) (A3.23) 1 + B Zm I + B ZmO

where zo is the initial mixture level.

Using a standard numerical solver (e.g. NAG routine), the above result may be used

to determine to mixture level trajectory during the boildown of a PWR SG. The

corresponding collapsed liquid level (in the riser and downcomer) may then be

obtained from equation (A3.20) and the SG inventory from equation (A3.15).

The basic model equations (A3.19), (A3.12), (A3.9) and (A3.8) were presented by Sun

et al. (1981) as a special case of a more general two-phase level theory. The

derivation presented in this appendix clearly shows the assumptions and

simplifications inherent in this model. In order to solve the equations, Sun et al. (1981)

introduced the additional assumptions that

BZm > 1 and Co-1> Bz (A3.24)

and, as a result, were able to obtain an explicit solution for z,, . Unfortunately, the

above assumptions cannot be justified in general for a SG boildown situation, and

thus the more exact (but implicit) solution is preferred for the present study. For

example, in Appendix A.3.7 below, the boildown model is used to emulate the

behaviour of RELAP5/MOD2 for which the appropriate value of CO is one, so that the

second condition could never be satisfied.

A.3.7 Sample calculations

For low flows in a PWR U-tube SG, the interphase drag models in RELAP5/MOD2 are

approximately equivalent to the use of a drift flux model with the following parameters

(Putney, 1988).

Co = 1.0 vgi = 0.43[ g PfI pDh ]/2 (A3.25)

-86-
NP Restricted TECIL/O4711R91

Bestion (1985) has shown however, that theoretical and experimental results support

the parameters

Co = 1.2 vgj = 0188[g Pg JP' ],2 (A3.26)

Using these definitions and a suitable numerical solver, equations (A3.20) and (A3.23)

may be applied to 'emulate' the modelling of SG boildown behaviour in

RELAP5/MOD2 for the conditions considered, and to compare the level trajectories

obtained against those predicted by a reference model.

This approach was used to study level behaviour in a full size PWR U-tube SG for

conditions characteristic of boildown under full power conditions. Geometric data for

the analysis were taken from the UK RELAP5/MOD2 deck for Sizewell 'B'. The SG

pressure and heat flux were taken as 60 bar and 15 Wcm-2 , respectively, and the

calculations were begun with the mixture level nominally set at the top of the U-bend.

Solutions were generated using the following models:

1. Drift flux parameters equivalent to the RELAP5/MOD2 interphase drag models

(i.e. equation (A3.25)). This solution will be referred to as the 'RELAP5 equivalent'

calculation.

2. The Bestion drift flux model (i.e. equation (A3.26)) - the 'reference' calculation.

3. The RELAP5/MOD2 equivalent drift flux model with the downcomer flow area

adjusted to give the same initial liquid inventory as the Bestion model.

4. The RELAP5/MOD2 equivalent drift flux model with the initial mixture level raised

to give the same initial liquid inventory as the Bestion model.

Note that calculations I and 2 do not provide an exact representation of the situation

typically encountered in a RELAP5/MOD2 analysis of a SG boildown transient. In

particular, the RELAP5 equivalent and reference calculations are begun with the

same mixture level (at the top of the bundle) and different downcomer levels. In

reality, a RELAP5 analysis of SG boildown would normally begin with the downcomer

level equal to the plant value and a different volume of liquid above the bundle. For

the same reason, the effect of increasing the downcomer area in calculation 3 is not

an exact representation of its effect in a plant or rig calculation. Unfortunately, the

type of situation normally encountered in a boildown test cannot be simulated with

the simple model, due to the requirement for the whole of the mixture to be heated

(at a uniform heat flux). Consequently, calculation 4 is also somewhat

unrepresentative as the effective heated length is extended to cover the liquid added

to make-up for the under-prediction of the initial inventory. Despite these

uncharacteristic features, the calculations should still exhibit effects that are relevant

to the modelling of boildown phenomena in RELAP5/MOD2.

In calculations 3 and 4, it was found that in order to force the RELAP5/MOD2 models

to calculate the same initial inventory as the Bestion models:

(a) the downcomer flow area had to be increased from 0.732 ml to 5.363 in,

or

(b) the initial mixture level had to be increased from 8 m to 56.3 m.

Clearly, these adjustments are not physically sensible and should be regarded only

as a device for compensating for errors. They also give some indication of the

magnitude of the initial void fraction error in the bundle region.

The level and inventory trajectories obtained from the four calculations are shown in

Fig. 15 to Fig. 17. Examination of these results reveals the following effects.

* Compared with the Bestion model, the RELAP5/MOD2 models under-predict the

initial liquid level (in riser and downcomer) and inventory by a significant amount.

- 87-
NP Restricted TECIL/O471/R91

As a result, the SG is predicted to empty earlier. Closer examination of the

results reveals that the liquid level and inventory also fall at a more rapid rate in

the RELAP5/MOD2 calculation, although the rates of fall in the two calculations

tend to converge as the transient proceeds40 . The initial mixture is the same for

both calculations but falls more rapidly in the RELAP5/MOD2 calculation.

indicating an earlier degradation in primary to secondary heat transfer.

0 If the downcomer area in the RELAP5/MOD2 calculation is increased to obtain the

same initial inventory as the Bestion calculation, the initial mixture and liquid

levels are unchanged. During the boildown however, the levels fall more slowly

than those from the unmodified RELAP5/MOD2. Eventually, the levels are

over-predicted with respect to the Bestion calculation. indicating a possible late

prediction of SG dryout. The inventory obtained from the modified RELAP5/MOD2

calculation is always greater than that associated with the Bestion calculation.

The above behaviour occurs because, although the under-prediction of the initial

inventory arises from an under-prediction of the liquid fraction in the bundle, the

liquid added to compensate for this does not change the liquid content in the

bundle (viz. the initial collapsed liquid level). Consequently, the extra liquid is

held away from the boiling region, but provides a greater reservoir from which

liquid can be drawn to replenish the liquid boiled off. A similar effect would result

if the downcomer area were artificially increased in a RELAP5/MOD2 analysis of

a real boildown transient.

* If the initial mixture in the RELAP5/MOD2 calculation is increased to obtain the

same initial inventory as the Bestion calculation, the same liquid level is also

obtained. During the boildown however, the liquid level, mixture level and

inventory fall very rapidly and tend to converge to the trajectories obtained in the

unmodified calculation - which is controlled by the equivalent RELAP5/MOD2 drift

flux parameters.

This effect occurs because the liquid added to compensate for the initial

under-prediction of inventory increases the effective boiling length and is

therefore boiled off very quickly. Evidently, this type of situation could be

encountered in a RELAP5/MOD2 analysis of a plant boildown transient, if the

initial inventory is artificially increased by increasing the recirculation ratio. When

the recirculation flow has ceased, this additional liquid will exist above the bundle

causing the bundle to uncover later in the calculation. When it does uncover

therefore, the true mixture level will be lower, and thus the excess liquid in the

calculation will boiloff rapidly leaving the subsequent boildown behaviour to be

controlled by the interphase drag models.

40 This was confirmed by normalising the RELAP5/MOD2 trajectories to coincide with the Bestion curves

at time zero.

88 -
NP Restricted TEC/L/0471/R91

Appendix B MEASUREMENT UNCERTAINTIES FROM

COMBINED ERRORS

The general equation for estimating the steam generator restrictor nozzle outlet

steam pressure is given by Preece (1987a):

P - P 1301.31 1 + 8.451 2(B1.1)

10 105

where the steam flow,w, = (wf - wb) and v, is the specific volume of saturated water

vapour at the steam line pressure P,. The constants 1.311 x 10-1 bar m3 kg-1 and

8.451 x 10-1 bar s2 kg-Im I represent the hydrostatic and dynamic head loss

components respectively. Similarly, the pressure drop across the steam flow

restrictor nozzle, P, ,is found from:

Pn = 0.299 ws 2v (B1.2)

The steam dome pressure at inlet to the restrictor nozzle is thus:

Pd = Pe + Pn (B1.3)

For the general case of:

Pd = Pd (PlI, V. ws) (B1.4)

the unknown uncertainty, Xdin Pd can be found in terms of the known uncertainties,

Xp1, X) and X,, to the components represented in equations (81.1) and (B1.2). This

leads to the following general expression:

P dPd 2 v, dPd 2 Ws dpd 2

IXpdI = X + (_•.dXv __-_) + (_Xw -w )} (B1.5) PPdP' d dv, Pd dws

or more specifically:

__ 1 XPI)2 + 1.1441 W 2 1.311 2

IXPd+ vw2 wv1) Xvl) +

10e 10 (81.6)

.+ (2.2882 viw 2 X }o.5 0S) 

10 4

Values of Xp,, X., and X,,, can be found in Preece (1987b) for the Wolf Creek steam

generator plant.

- 89-
NP Restricted TEC/L10471/R91

Nominal OD of Average

Plant/ Ref Thermal No. of No. of Length SG

Facility Plant Scale Power Loops Tubes Tubes of SG

(MW) per SG (cm) Tubes

Sizewell 'B'

W SNUPPS Full 3425 4 5626 1.748 16.756 U-type

Wolf PWR Model F

Creek

W PWR U-type

KNU #1 WB C TE Full 1724 2 3386 2.2 22.831 Mode

(BECHTEL) .81 Model 51

Doel 2 W PWR Full 1192 2 3260 2.22 18.16 U-type

(TRACTEBEL) Model 44

W PWR U-type

Doel4 (TRACTEBEL) Full 3000 3 Model E

(preheater)

Framatome U-type

Tihange 2 PWR Full 2785 3 3361 2.22 20.4 Series 51M

(TRACTEBEL)

10

4-loop 1 0(4%o

ROSA-IV W-type 1 (14% of 2 141 2.54 19.71 U-type

PWR 48 scaled

ref power)

Typical 1 1845

LOFT 4 -loop 6 50 2 1.27 4.578 U-type

PWR 68 0

MB.2 Single 1 10 1 52 1.75 13.905 U-type W SG 125 Model F

SPES W 312 type 1 6.5 13 1.746 16.564 U-type PWR 427

4-loop 1 24

LOBI KWU Biblis 712 5.3 2 8 2.2 14.279 U-type

PWR

Raised 1 1 - 5% of

OTIS Loop -16- scaled 1 19 1.587 15.9 OTSG

B&W PWR ref power

Semi- 4-loop 1 6 U-type

Mosca2l2e scale PWR 1706 2 2.22 18. Model U-yp51

Table 1. Summary of PWR plants and facilities considered in review

- go -
NP Restricted TECIL/0471IR91

Average Error in Uncertainty in

era Calculated Measured

Plant/ Test SG Hea Organisation Code SG Modelling

Facility Flux and Ref Version

(W cm- 2

) HTC-A HTC-A

(%) (%)

Wolf SS at 36.1% 5.9 -4.9 -4.9 to 5.2 NP TEC 36.04 . 00

Creek Load (1) IBM CR= ,/

SS at 51.5% 8.3 -8.5 -3.8 to 3.9

Load

SS at 73.3% 12.1 -6.6 -2.7 to 2.8

Load

SS at 81.4% 13.4 -8.6 -2.4 to 2.5

Load

SS at 98.9% 16.2 -11.5 -2.2 to2.2

Load

SS at D0= gap

98.9% 16.2 -8.2 -2.2 to 2.2

Load CR =

LR from NETD 36.05 D.= OD

100 to 16,1 -9.6 -2.5 to 2.5

50% (2) UK E04 CR = H

Trip from D - OD

100% 16.1 -5.1 . -3.3 to 3.5 NE TD 36.05

Power (3) UK E05 CR='J

9% D, Dh

KNU #1 LOP 16.2 increase KAERI

(4) 3.5 C

In HT area Drag Dh VL

ROSA-IV ISP-26 2.4 -7.4 -12.1 to 13.2 NE TD/HSD 36.05 D= Dh

5% CLB 2.3 -6.6 -12.4 to 13.5 (5) UK E03 CR = H

LOFT L3-7 15.3 -17.7 -8.5 to 8.8 KAERI 36.05 09 = 0

I' CLB (6) CR= ?

L2-5 10.4 -2.7 -10.3 1o 10.7 KAERI 36.04

200% CLB (7) CR3=0 ,

2%2-5 10.4 -8.9 -10.3 to 10.7 INER-China D. =7

200% CLB (8) CR ?

LP-SB-03 0ETD D. = Gap

0.4% CLB 15.2 -12.4 -5.9 to 6.2 (9) 36.01

LP-S-01 ED, = Gap 1% HLB 14,1 0.0 -8.7 to 9.0 NETO 36.02

(10 CR =

0. = Gap

LP-SB-02 14,1 -6.1 -8.8 to 9.1 NE TD 36.04

1% HLB (11) CR = ?

09 = Gap

LP-FW-01 14.9 -1.4 -5.1 to 5.3 NE TD 36.04

LOFW (12) CR,= ?

L9414.7 -5.9 -8.3 to 8.7 NE TD 36.05 D0 = Gap

LOP-ATWS (13) UK E03 CR = 5.1

L9-3 14.3 -7.6 -8.6 to 9.0 rC 36.05 D. = Gap

LOFW-ATV4S (14) UK E03 CR = ?

L9-3 WTC 36.05 D. = Gap

LOFW-ATWS 14.3 -6.2 -8.6 to 9.0 (Revised) UK E03 (14) CR = >

L3-5, L3-6 Good NPTEC 36.05 D. = Gap

2.5% CLB 15 Agreement (15) UK E03 CR =

L.3-6 Studsvlk D. 0h

2.5% CLB 14.7 -7.4 -14.5 to 15.4 (Case A) 36.04

(16) CR = 4.7

2.5% 1_3-6 CL3 Studsvlk D.= Dh 14.7 -9.6 -14.5 to 15.4 (Case B) 36.04 2.5% CLE3 I I CR = 4.7

-91 -
NP Restricted TECIL/0471/R91

Error in Uncertainty in Average Calculated Measured

Plant/ SG Heat Value of Value of Organisation Code SO Modelling

Facility Test Fux and Ref Version

(W cm- 2 _______-__)_(%/) ) HTC-A HTC-A (%)

M- 2 3% D.

MB-2 LOFW ) increase INER-Taiwan 36.00

in HT area (17) C, -

18.6 2.9 -6.4, to 6.5 NE PPG 36.05

SPES ISP-22 17.8 4.9 -6.8 to 6.9 (18) UK E04 - LOFW 17.7 4.9 -6.5 to 6.7 CR H

A2-90 16.1 10.3 ? WTC 36.05 De = Gap

LOBI LONOP-ATWS 17.8 -1.7 7 (19) UK E01) CR =

D.==

OTIS LOFW 0.5 -10.0 INEL (20) 36.04 CR=

Semi- S-LH-1 19.9 5.1 -6.1 to 6.3 NE TD 36.05 Dscale 5% CLS 19.9 -3.9 -6.2 to 6.5 (21) UK E03 CR =

S-L H-2 20.2 5.9 -6.2 to 6.4 NE TD 36.05 =

5% CLB 16.6 10.3 -6.2 to 6.4 (22) UK E03 CR _ ?

Table 2. Errors in total SG heat transfer coefficient HTC-A

The parameter HTC-A is defined in Section 2.7

For nonsymmetrical systems, the first entry refers to the intact loop.

SG Modelling parameters

D., Equivalent heated diameter on secondary side

CR Circulation ratio

OD Tube outer diameter

Dh Hydraulic diameter on secondary side

Gap Irter-tube gap

/ Calculated correctly

H Calculated high

VL Set very large

? Unknown

> Unknown. but greater than entry above

Other symbols used in the table are defined on the next page.

- 92-
NP Restricted

Organisations

Abbreviation

NP TEC

NE TD

NE HSD

NE PPG

WTC

KAERI

INER-China

INER-Taiwan

Studsvik

INEL

References

TEC/L/O471IR91

Organisation

National Power Technology and Environmental Centre

Nuclear Electric Technology Division

Nuclear Electric Health and Safety Division

Nuclear Electric PWR Project Group

Winfrith Technology Centre

Korea Advanced Energy Research Institute

Institute of Nuclear Energy Research, China

Institute of Nuclear Energy Research, Taiwan

Studsvik Engineering AB, Sweden

Idaho National Engineering Laboratory

Number

(1)

(2)

(3)

(4)

(5)

(6)

(7)

(8)

(9)

(10)

(11)

(12)

(13)

(14)

(15)

(16)

(17)

(18)

(19)

(20)

(21)

(22)

Reference

This report (Section 4)

Keevill (1989)

Proctor (1990a)

Chung et al. (1990)

Cooper and Froushan (1990)

Lee et al. (1990)

Bang et al. (1990)

Kao (1988)

Harwood and Brown (1990)

Hall and Brown (1990)

Hall (1987)

Croxford et al. (1988)

Keevill (1988)

Birchley (1988)

Scriven (1990)

Eriksson (1990)

Wu and Wang (1988)

Willmott (1990)

Whittall (1990)

Willmott (1990)

Hall and Bull (1989)

Brodie (1989)

- 93 -
NP Restricted TECIL/0471/R91

Average Error in Uncertainty in

Ant/eag Calculated Measured Plan Test SG Value of Value of r SG Modelling Facility Flux HTC-Oand Ref Versio

(W cm_2 ______ ______ ) HTCMO (%) (%) a e Vri ____

Wolf SS at NP TEC 36.04 0D OD

Creek 36.1% 5.9 -12.9 -11.9 to 17.4

Load IBM C ,

SS at

51.5% 8.3 -23.2 -10.2 to 14.3

Load

SS at 73.3% 12.1 -18.1 -7.2 to 8.9

Load

SS at

81.4% 13.4 -23.6 -6.9 to 8.6

Load

SS at D, = gao

98.9% 16.2 -23.3 -6.3 to 7.7

Load CR =

SS at

98.9% 16.2 -30.4 -6.3 to 7.7

Load

LR from NE TD 36.05 D= OD

100 to 16.1 -24.4 -6.3 to 7.6

50% (2) UK E04 CA H

Trip from NE TD 36.05 D. OD

100% 16.1 -13.3 -7.6 to 9.8

Power (3) UK E05 CR= "

9% D.= Dh

KNU #1 LOP 16.2 increase 4KAERI 36.05 CR \

in HT area Drag Dh VL

ISP-26 2.4 -63.7 -44.4 to inf. NE TD/HSD 36.05 D0. D, ROSA-IV 5% CLB 2.3 -57.4 -43.3 to Inf. (5) UK E03 CR = H

LOT1CBKAERI 36.5 =0D

LOFT L137 15.3 -28.3 -12.3 to 14.8 36.05

1' CLB (6) CA ?

L2-5 10.4 -6.9 -18.5 to 28.8 KAERI 36.04 = D

200% CLB (7) CR =

L2-5 IE-hn 200% CLB 10.4 -22.6 -18.5 to 28.8 INER-Ch(na ? =

20%CB(8) CR=

LP-SB-03 NE TO D 9= Gap

0.4% CLSB 15.2 -20.7 -9.2 to 10.9 (9) 36.01C =

LP-SB-01 NE TD D, = Gap

1% HLB 14.1 -0.5 -11.9 to 13.7 (10) 36.02 CR = ?

LP-SB-02 14.1 -10.5 -12.2 to 14.2 NE TD De = Gap

1% HLB (11) 36.04 CR = ?

LP-F'-01 N TD0D = Gap

LP-FW-01 14.9 -2.4 -8.3 to 9.6 NETD 36.04 D. = ?

LOFW (12) CR = 1

L9-4 NE TD 36.05 D, = Gap

LOP-ATWS 14.7 -10.4 -11.71t 13.7 (13) UK E03 CR = ?

L9-3 WTC 36.05 D0= Gap

LOFW-ATWS 14.3 -13.2 -12.2 to 14.4 (14) UK E03 CR = 5.1

L9-3 w'c 36.05 0. = Gap

LOFW-ATWS 14.3 -10.9 -12.2 to 14.4 (Revised) UK E03 (14) CR = >

2.5% LB 15 Good NPTEC 36.05 D ap

2.5% CLB Agreement (15) UK E03 CR

Studsvik D0 , Dh 2.3-6 14.7 -12.2 -18.2 to 22.1 (Case A) 36.04 2.5% CLB (16) CR = 4.7

Studsvik D .. Dh L3-6 14.7 -15.7 -18.2 to 22.1 (Case B) 36.04 2,5% CLB (16) C__= 4.7

-94 -
NP Restricted TEC/L/0471/R91

Error in Uncertainty in Average Calculated Measured

Pln/SG Heatl acltd Maue Organisation Code

Planty Test FG Value of Value of andsvtion SG Modelling Facility Flux H OHT and Ref Version

(Wc-) HTC.O HTC-O (wcm-') (%) (%/) _____

23% 1 INER-Taiwan

M-2 LOFW Increase (17) 36.00

in HT area CR ?

ISP-22 18.6 9.5 -11.4 to 14.7 NE PPG 36.05 D, =Dh

SPES 17.8 14.1 -11.7 to 15.0 (U LOFW 17.7 14.2 -11.5 to 14.7 (18) UK E CR = H

LOW A2-90 16.1 18.7 ? WTC 36.05 D. = Gap

LONOP-ATWS 17.8 -4.1 7 (19) UK E0 CR =

0D, '

OTIS LOFW 0.5 -52.5 ? INEL 36.04 (20) CR =

Semi- S-LH-1 19.9 12.8 -9.7 to 11.4 NE TD 36.05 D. =

scale 5% CLE 19.9 -9.6 -10.4 to 12.6 (21) UK E03 CR =

S-LH-2 20.2 13.8 -9.7 to 11.5 NE 7D 36.05 D=

5% CLE 16.6 25.7 -9.8 to 11.6 (22) UK E03 CR =?

Table 3. Errors in total SG heat transfer coefficient HTC-O

The parameter HTC-O is defined in Section 2.7

For nonsymmetrical systems, the first entry refers to the intact loop.

Other symbols used in the table are defined under Table 2.

- 95 -
NP Restricted TEC/L10471/R91

Test Run Number WC046 WC047 WC049 WC051 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 4. 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. The primary

fluid flow is that deduced using the method outlined in Section 4.1.

- 96 -
NP Restricted TEC/LIO471IR91

Reactor Power, %

Parameter

36.1 98.9

best estimate 1.14 1.74

Tube thermal

conductivity lower bound 1.10 1.70

upper bound 1.18 1.78

best estimate 1.08 1.28

Heated tube

length lower bound 1.06 1.27

upper bound 1.11 1.29

best estimate 118.3 11164

Heated

diameter lower bound 1/12.8 1/179

ubper bound 1/5.7 1/149

Table 5. SG heat transfer adjustment factors for Wolf Creek analysis

* Expressed as a fraction of the inter-tube gap

- 97 -
NP Restricted TECILI0471/R91

Fig. 1. Sensitivity of increasing SG heat transfer

2.00

1.SO

Increase in Heated Lent

-- Decrm in Wal Thictne3s

- - - increase in Thermal Conducily

........ Increase in Secondary HTC

Increase in Primary HTC

1.60 -1 - -

• ................................................. .......

1.0

1 r2 3 4 5 6 7 8 9 1e Factor of Incrase or Deerease in Parameter

Fig. 2. Sensitivity of decreasing SG heat transfer

.... Decre in ItWed Length

-- Incrse in Wall Thclm

-- D-ec. in Thermal Con~uctWty

........ Decrease in Secondary l-C

- Decrease in Primay HTC

4-

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

V - Or ..........

.3 ... .. ....

oi

1 o2 3 I 4 5 7 8a 9 10 Facor of Decrease or Incree in Parameer

98-
NP Restricted TEC/L/04711R91

Fig. 3. Variation of boiling heat transfer coefficient with heated diameter

P=68bar G=33•kg/m2-s DIS=7DegC X=0.17

-.- Standard Dien RIC

- - - RELA/002 HTC

........ REIa O/M02 Hmic .o ..........- , .................

..... REAP5/U02Hmc

C. E 4=s 0 . -. -=. . .............. ......... .'z-. . ..................................................... .......................... ...........

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

" i ..... .....;...............

I

a-

.00 .50 1.00 1.50 2.00 2.50 300

Heated Diameer (cm)

Fig. 4. Variation of boiling heat transfer coefficient with mass flux

P=68bar D1S=70egC X=0.17 OH= 1.7cm

E ------------------------------------------

1i5=

- .Standard Chan HmC

--- REIMP/M002 HiT

I....... M /M002 Hrnfic

................ M-O2Ka

................... I I I

340 360 380 400

Moss Flux (k/m2-s)

420 440 460

- 99 -
NP Restricted TECIL/O471/R91

Fig. 5. Variation of boiling heat transfer coefficient with quality

P = 68.0 br G = 400.0 kg/m2-s DTS =7.0 Deg C DH= 1.7 cm

--- Standard Chen HTC

5o: - - - RBAP5/MO02 HTC

"L M /........ 002 Hmic

(> - lAP-/Mo02 Hmrc

10,0M ..

.OO . .100 .150 .200 .20

QuarY

-100-
NP Restricted TEC/L/0471/R 91

Steam restrictor nozzle - inlet plane

!VTh

ro 1-Pressure tapping PT7, points on the

primary moisture separator

C

O E

Pressure/temperature probe

- Tube sheet upper face

• i

Figure 6 Locations of special instrumentation in Wolf Creek steam generator

JMP/•cPR(8..91)0471/6

- 101 -
NP Restricted TEC/LI0471IR91

67131 2

Mid deck plate .1 601 612

Intermediate 608 610 641

deck plate 606

Lower deck plate .... 60

603

6 1

32320

602

5 6 5 2

4 F6~ 602

3 8 3 4

2 A92 5

6

Primary fluid - pump suction leg Piayfud-h

Figure 7 Nodalisation of the model 'F' steam generator

Main feed

leg

JMP/CPR(8.8.91)0471i7

t

- 102 -
NP Restricted TECILlO4711R91

Fig. 8. Errors in the calculated SG dome pressure, Wolf Creek

DIFFERENCE BETWEEN MEASURED AND CALCULATED

STEAM PRESSURE, DELTA P, IN SG DOME

4.00 .................................................

3.50 ............ ......... ................ ....

3.00-.......

-Z2.50-.........

2.0 .. .. ..

1'2. 0 0 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

01.50-.

0.00-

30 4!0 5o 610 70 80 90 100

REACTOR POWER.(%)

Fig. 9. Component pressure drop below mid-deck plate level, Wolf Creek

PRESSURE DIFFERENCE BETWEEN BO'nTOM OF

DOWNCOMER AND MID-DECK PLATE LEVEL

1.g00. ..............................................

0.580...... .... ......... :......_-.. .. 0.900-

u b dp Measured (Pla-PTe) l W

0O.900...._..................... •........ :................

ILI

.1C - .----

WO.B20-.

0. -. . ; • .

-: • C .l:ula:ed

0.1320 ...............- . . . . . .

O.BO0 ; ; ;

30 40 50 60 70 80 90 100 REAORPO)

- 103 -
NP Restricted TECILIO471/R91

Fig. 10. Component pressure drop above mid-deck plate level, Wolf Creek

PRESSURE DIFFERENCE BETWEEN MID-DECK

PLATE AND INLET TO SG RESTRICTOR NOZZLE

0.0140........ ...........

Measured (PT8-Pi)

0.0120. . . . . .

n--"- ---- -------- ----- . -.- -. 

o:0.0100. ...-. ý............... ted....-..........

0- .o0o8........ ;..............................

0.0080-.

0.0060............................... ........

I I I I I •

30 40 50 60 70

REACTOR POWER (1)

80 gO 100

Fig. 11. Component pressure drop across the mid-deck plate, Wolf Creek

PRESSURE DROP ACROSS THE

MID-DECK PLATE

0.0350 .................................................

0.0300 -

0.0250........ . ............. alculated. ....... ............ 0

cO0.0200 ...

IJ Measured (PT7'PT8)

0.0150 . ........ :. . . . . . . . . . . .. :. .. . . ... ... . ... .... ... :

0.0050-

30 80 90 10

30 40 50 60 70

REACTOR POWER,(I)

80 90 100

- 104 -
NP Restricted TECIL/O471IR91

Fig. 12. RELAP5/MOD2 SG secondary side HTC for Wolf Creek at 99% load

Incremse in HTC

3-

90M

80= El ........ .. 2-

7000 ........ ..

-G40=0

011

5000

P;= 0,0

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

\!!!Ll E De=Tube 00

20,000 I I I I I i

0 2 4 6 B 10

Dtonce from Hot Side Tube Sheet (m)

I I I 12 14 16 1

- 105 -
NP Restricted TEC/LIO471IR 91

Fig. 14. Comparison of Chen correlation with data from Semiscale

SG: Semiscale results taken from broken loop long tube

W - RU.AP/MOD2 (Wolf Creek)

a- Cen Coffeotir

a -Data 60,' -

5g'4o~

T31

4000 3.-

a.... ........ ......................... ........ . . ...............

lox I I I I

1 2 3 4 Dvon(m)

I I 1 1

5 6 7 a

Fig. 15. Liquid level trajectories for simple SG boildown transient

P= 60.0bar Q= 15.00W/cm2-tube

3.50

2.50

2.00 -

1.50

0.50-

I - REUP/mnO Madeb

- - - KMAP/10002 w'di Adjusted D/C Aree

-R- - ao/=00 wbt Adjused ZffO

I

C

I

- 106 -
NP Restricted TEC/L10471/R91

Fig. 16. Inventory trajectories for simple SG boildown transient

P = 60.0 bar Q= 15.00 W/cmN2-tube

10,00 -Il

14,000 -

-6,0, n

4.00 -M

I %

•5 -, -

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

50 100 150 200 250 300 350 400 450 5

The (s)

Fig. 17. Mixture level trajectories for simple SG boildown transient

P= 60.0 bar 0= 15.00 W/cm2-tube

8

8 1 - RELAP5/M002 Models

7-lki• -...... Bs ato,nCm 

A3 IN\':.

0 50 100 150 200 250 300 350 40 450

Tirm (s)

0

- 107 -
NP Restricted

20

15 -

10

TEC/L/0471/R91

m Wolf Creek (NP TEC)

0 Wolf Creek (NE TD)

X ROSA - IV

A LOFT (UK)

LOFT (non-UK)

SPES

+ LOBt

OTIS

* Semiscale

C,

wJI.-

5

0

-5

-10

-15

-20

U 0 - De=gap

FD

m

0 2 4 6 8 10 12 14 16 18 20 22

Average SG heat flux, W/cm 2

Figure 18 Errors in HTC-A vs SG heat flux

JMP/CPR(29.7.91)0471118

- 108-
NP Restricted TEC/L/0471/R91

U

D

x

A

A

'N

+

Wolf Creek (NP TEC)

Wolf Creek (NE TD)

ROSA- IV

LOFT (UK)

LOFT (non-UK)

SPES

LOB!

OTIS

Semiscale

60

40

20

*

C~

0

LU

0

U

-20

-40

U -De = gap 0 m

0

-60 X

x I I I I 1 1

0 2 4 6 8 10 12 14 16 18 20 22

Average SG heat flux, W/cm 2

Figure 19 Errors in HTC-O vs SG heat flux

JMP/CPR(29.7.91)0471/19

- 10g -
NP Restricted TEC/L/O471IR91

U

A

Vw

Wolf Creek (NP TEC)

Wolf Creek (NE TD)

ROSA- IV

LOFT (UK)

LOFT (non-UK)

SPES

LOBI

OTIS

Semiscale

20

15

10 x *

0

I.-

=

CI-.

L.

w

5

0

-5

-10

-15

-20

U

U

* De=gap A

10 15 20 25 30 35

DT-A, 0C

Figure 20 Error in HTC-A vs primary to secondary temperature difference

JMP/CPRC29.8.91)0471/20

- 110-
NP Restricted

Distribution

Copies to:

TECIL/0471IR91

Addressee:

Mr P Lightfoot

S Dr C. Lopez-Cacicedo

Dr R.S. Adrain

Dr I.Milne

Mr M.W.E. Coney

Dr J.M. Putney

Mr A.H. Scriven

Document Retrieval Centre

Dr P.R. Farmer

Dr N.E. Buttery

Mr K.H. Ardron

Mr P.C. Hall

Mr I.L. Hirst

Dr G. Ahmed

Dr D.R. Atthey

Mr B. Chojnowski

Mr D. Withrington

Mr K.T. Routledge

Mr I. Brittain

Information Services

Nuclear Electric, PPG

NP TEC

NE HSD

NE PPG

NE TD

NE ITD

NE BNL

NNC

AEA Winfrith

NP Oakdale, Harrogate

- 111 -

NRC PORM 335 U.S. NUCLEAR REGULATORY COMMISSION I.REPORT NUMBER 12-891 (Abtolud bV NRC. Add VoL. Ijpp. Rev RCM 1102. & Adde• d N um ryI If eayj 5201.3202- BIBLIOGRAPHIC DATA SHEET

s*e ins;t,,uctons on the •,vve. NUREG/IA-0106

2. TITLE AND SUBTITLE TEC/L/0471/R91

Assessment of PWR Steam Generator Modelling in RELAP5IMOD2 3. DATE REPORT PUBLISHED

MONTH y E AR

June 1993

4. FIN OR GRANT NUMBER L2245

5. AUTHOR (SI 6. TYPE OF REPORT

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

7. PERIOD COVERED hc.lve Dares)

National Power, Technology and Environmental Centre

Kelvii Avenue, Leatherhead

Surrey, KT22 7SE

United Kingdom

9. SPONSORING ORGANIZATION - NAME AND ADDRESS (If NRC. tvprma " Dwi• co,•M• a er provide NRC D'Won. Ofi or 'Poon. U.S. AhdwA AWo,v .Cowm~d,,,. vw mamoaddrmj

Office of Nuclear Regulatory Researh

U.S. Nuclear Regulatory Commission

Washington, DC 20555

10. SUPPLEMENTARY NOTES

'10. NOTES

11. ABSTRACT M23W- o-mi

An assessment of Steam Generator (SG) modelling in the PWR thermal-hydraulic code RELAP5/MOD2 is

presented. The assessment is 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 are identified and their implications for PWR safety analysis are discussed - including methods for

compensating for the deficiencies through changes to the input deck. Consideration is also given as to whether the

deficiencies will still be present in the successor code RELAP/MOD3.

12. KEY WORDSiOESCR:PTORS (List woonk orim oMar wmsrmaktn,•u , in locarag rohe ftp.1 I13. AVA#LABLIT'Y STATEMENT

ICAP Pro m Tnlmt P R o14. PWR SECURITY CLASSIFICATION

RELA/5MOD2 Thi, Pa•e•

Unclassified

IThýs Repo"I

Unclassified

15. NUMBER OF PAGES

16. PRICE

NRC FORM 335 12491
Federal Recycling Program

UNITED STATES

NUCLEAR REGULATORY COMMISSION

WASHINGTON, D.C. 20555-0001

SPECIAL FOURTH-CLASS RATE

POSTAGE AND FEES PAID

USNRC

PERMIT NO. G-67

OFFICIAL BUSINESS

PENALTY FOR PRIVATE USE, $300