NUREG/IA-0185

International 

Agreement Report 

Development and Validation 

of a Transition Boiling Model 

for the RELAP5/MOD3 

Reflood Simulation 

Prepared by 

V. H. Sanchez-Espinoza, E. Elias*, Ch. Homann, W Hering, D. Struwe 

Forschungszentrum Karlsrube 

Hermann-von-Helmholtz Platzl 

D-76344 Eggenstein-Leopoldshafen 

Germany 

*Technion University 

Haifa, Israel 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washington, DC 20555-0001 

June 2000 

Prepared as part of 

The Agreement on Research Participation and Technical Exchange 

under the International Code Application and Maintenance Program (CAMP) 

Published by 

U.S. Nuclear Regulatory Commission
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NUREG/IA-0185

International 

QO Agreement Report 

Development and Validation 

of a Transition Boiling Model 

for the RELAP5/MOD3 

Reflood Simulation 

Prepared by 

V H. Sanchez-Espinoza, E. Elias*, Ch. Homann, W Hering, D. Struwe 

Forschungszentrum Karisrube 

Hermann-von-Helmholtz Platzl 

D-76344 Eggenstein-Leopoldshafen 

Germany 

*Technion University 

Haifa, Israel 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washington, DC 20555-0001 

June 2000 

Prepared as part of 

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

Published by 

U.S. Nuclear Regulatory Commission
Abstract

The heat transfer model of the RELAP5/MOD3.1 (R5M3) code was extensively 

reviewed and assessed. The most important deficiency of the current version of the 

code was attributed to its treatment of the transition boiling heat transfer regime. 

The current transition boiling model significantly underpredicts the heat transfer 

rate to the liquid phase. Since at low quality conditions the liquid boiling part is a 

major fraction of the total heat transfer, the current model underpredicts the quench 

temperature and the quenching rate under most conditions relevant to LOCA and 

degraded core analysis. 

Therefore, a new model has been developed and implemented in the R5M3 code 

for predicting the transition boiling'heat transfer. The new model is based on an 

extension of the phenomenological formulation suggested originally by Chen. It uti

lizes only local state variables calculated by the R5M3 code and does not require 

other history parameters, such as quench position or CHF and minimum film boiling 

temperatures, which are not available at actual time step. 

A number of separate effect and bundle tests are analyzed with the modified code 

version. The predictions are compared with those obtained by the frozen code version 

and with available experimental data. Several variables, such as wall temperatures, 

vapor and liquid velocities, void fraction etc., are examined in order to evaluate the 

general prediction capability of the code in modeling boil-off and reflood transients. 

In addition, the current and the modified stand-alone transition boiling models are 

tested against a large sample of the available data-base on steady-state post dryout 

heat transfer. 

In all cases, the predictions of the modified model fit the measured data better. 

The temperature curves are physically and conceptually more sound than those 

predicted by the frozen code version. This is achieved by introducing a more realistic 

modeling of the transition boiling heat transfer which affects only one subroutine of 

the R5M3 code. Preliminary results of the FZK transition boiling model validation 

against both the full-length FLECHT-SEASET bundle test 31701 and the integral 

test LOFT-LP-LB-1 have shown a better prediction of the quench process than the 

frozen R5M3 version.

iii
Contents 

1 Introduction 1 

2 The RELAP5/MOD3 Reflood Model 3 

2.1 RELAP5/MOD3 Code Basis .......................... 3 

2.2 R5M3 Reflood Heat Transfer Models ..................... 5 

2.3 Reflood Heat Conduction ............................ 11 

3 Experimental Programs for Code Validation 12 

3.1 The FZK Single-Rod Quench Tests ...................... 13 

3.1.1 Single-Rod QUENCH Rig Description ................ 13 

3.1.2 Test Procedure ............................. 13 

3.1.3 Measurement Uncertainty ....................... 17 

3.2 The NEPTUN Bundle Tests .......................... 18 

3.2.1 Test Description and Procedure .................... 18 

3.2.2 Measurement Uncertainty ....................... 20 

3.3 The Validation Matrix ............................. 20 

4 Test Modeling for Reflood Model Assessment 23 

4.1 The FZK Single-Rod Test Rig ......................... 23 

4.2 The NEPTUN Test Section .......................... 26 

5 Results of Assessmental Calculations with RELAP5 MOD3 26 

5.1 FZK Single-Rod ................................. 28 

5.2 NEPTUN Facility ................................ 31 

6 Stand-Alone Assessment of the RELAP-Reflood Model 31 

6.1 Validation Matrix ................................ 31 

6.2 Discussion of Results .............................. 34 

7 Limitations of the Current 'Transition Boiling Model 37

V
8 The Suggested Transition Boiling Model 40 

8.1 Total Boiling Heat Flux ............................ 40 

8.2 Liquid Contact Area Fraction ............................... 42 

8.3 Implementation of the new Transition Boiling Model ............... 43 

9 Validation of the New Transition Boiling Model 44 

9.1 Stand-Alone Validation ....... ............................ 44 

9.2 Validation within RELAP5 MOD3.1-FZK ....................... 45 

9.2.1 Discussion of Results: FZK Single-Rod Tests: Specimen filled with 

Z r0 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 45 

9.2.2 Discussion of Results: FZK Single-Rod Tests: Empty Specimen . . 59 

9.2.3 Discussion of Results: NEPTUN Tests .................... 63 

10 Parametric Studies 71 

11 Run Statistics and Code Performance 74 

12 General Conclusions and Recommendations 80 

13 Ongoing Activities 81 

14 References 85

vi
List of Figures

1 Flow Regime Maps of the Constitutive Relations in R5M3 ............ 4 

2 Qualitative Representation of the Boiling Curve used in RELAP5 MOD3.1 5 

3 Arrangement of the Single Rod Quench Tests at FZK ............... 14 

4 Rod QUENCH Rig: Cross Section and Temperature Measurement ..... .15 

5 FZK Single-Rod Quench Test Conduction ....................... 16 

6 Cross Section of the NEPTUN Heater Rod ...................... 19 

7 Cross Section of the NEPTUN Bundle ......................... 21 

8 Axial and Radial Nodalization of the Single-Rod Quench Test Rig ..... .25 

9 Nodalization Scheme of the NEPTUN Test Section ................ 27 

10 Comparison of the Original R5M3 Predictions with the Measured Temper

atures in the FZK Single-Rod Test No. T02106 ................... 29 

11 R5M3 Predictions of a. Surface Heat Flux and Critical Heat Flux, and b. 

Surface HTC in the FZK Single-Rod Test No. T02106 .............. 30 

12 Measured and Predicted Clad Outside Temperatures and Steam Coolant 

Temperature for the Steam-Cooled FZK Single-Rod Test No.T16046 . . 32 

13 Measured and Predicted Clad Temperature in the NEPTUN Test 3056 . 33 

14 Measured vs Stand-Alone R5M3 Model Predictions of Transition Boiling 

Data ........ ....................................... 35 

15 Stand-Alone R5M3 Model Prediction of Run 5251 in Bennett et al. (1967) 36 

16 Heat Flux versus Wall Temperature predicted by RELAP5 MOD3 for G = 

20kgsm2 , p = 0.IOMPa and saturated conditions . ................ 38 

17 Schematic description of the Suggested Transition Boiling Model ...... .41 

18 Comparison with Data of a Stand-Alone Formulation of the New Transition 

Boiling Model ......................................... 46 

19 Stand-Alone New Model Prediction of Run in Bennett et al. (1967) . . .. 47 

20 Measured and Predicted Temperatures in the FZK Single-Rod Test No. 

T16106 .............................................. 49 

21 Measured and Predicted Clad Temperatures in the FZK Single-Rod Test 

No. T31085 ....... ................................... 50 

22 Measured and Predicted Clad Temperatures in the FZK Single-Rod Test 

No. T24066 ........ ................................... 51

vii
23 Predicted Heat Flux Parameters at the Mid-Section FZK Single-Rod Test 

No. T16106 .......................................... 52 

24 Comparison of Current and Modified R5M3 Predicted Boiling Curves in 

the FZK Single-Rod Test No. T16106 .......................... 53 

25 Predicted Flow Parameters in the FZK Single-Rod Test No. T16106. . .. 55 

26 Measured and Predicted Clad and Pellet Temperatures in the FZK Single

Rod Test No. T02106 .................................... 56 

27 Measured and Predicted Clad Temperatures in the FZK Single-Rod Test 

No. T21066 .......................................... 57 

28 Measured and Predicted Clad Temperatures in the FZK Single-Rod Test 

No. T24085 .......................................... 58 

29 Measured and Predicted Clad Temperatures in the FZK Single-Rod Test 

No. ITE20115 ......................................... 60 

30 Measured and Predicted Clad Temperatures in the FZK Single-Rod Test 

No. ITE22115 ......................................... 61 

31 Measured and Predicted Clad Temperatures in the FZK Single-Rod Test 

No. ITE29115 ......................................... 62 

32 Measured and Predicted Clad Temperatures and Void Fraction in the NEP

TUN Test 5036 ........................................ 64 

33 Predicted Heat Transfer Parameters in the NEPTUN Test 5036 ...... .65 

34 Measured and Predicted Clad Temperatures in the NEPTUN Test 5050 . . 67 

35 Predicted Heat Transfer Parameters, 0.746 m Elevation in the NEPTUN 

Test 5050 ........ .................................... 68 

36 Predicted Heat Transfer Parameters, 0.978 m Elevation in the NEPTUN 

Test 5050 ........ .................................... 69 

37 Predicted Heat Transfer Parameters, 1.21 m Elevation in the NEPTUN 

Test 5050 ........ .................................... 70 

38 Effect of Gap Conductance on Quench Velocity in the FZK Single-Rod 

Test No. T16106 ....... ................................ 72 

39 Gap Temperature Curves in the FZK Single-Rod Test No. T16106 .... . 73 

40 Temperature Difference Across the Gap at Different Heights in the FZK 

Single-Rod Test No. T16106 ................................ 74 

41 Gap Width Behavior During Reflooding in the FZK Single-Rod Test No. 

T16106 ............................................ 75

viii
42 Code Performance: Rod Test No. T16106 ........................ 76 

43 Code Performance: NEPTUN Test 5036 ..... ................... 78 

44 Code Performance: NEPTUN Test 5036 ..... ................... 79 

45 Measured and Predicted Cladding Temperatures of the FLECHT-SEASET 

Test 31701 ............................................ 83 

46 Measured and Predicted Cladding Temperatures of the LOFT-LP-LB-1 Test 84

ix
List of Tables

1 Comparison of predicted and measured temperature difference for cooldown 

tests ........ ....................................... 18 

2 NEPTUN measurement errors and scattering of the data ............ 22 

3 Test matrix for the validation of the R5M3-reflood model ............ 23 

4 Test Matrix for Stand-Alone Validation of R5M3-Reflood Model ...... .33 

5 Comparison of RELAP5 and the Original Chen Model for Transition Boil

ing Heat Transfer ....................................... 39 

6 Comparison of the Original -and Modified Chen Model for Transition Boil

ing Heat Transfer ....................................... 44

x
Nomenclature'

Latin Letters 

CP [kg--] heat capacity at constant pressure 

D [m] Diameter 

D. [i] Hydraulic diameter 

Al Fraction of wetted wall surface 

G [kg/rn2 s] Mass flux 

Gr Grashof number 

h [w--T] Heat transfer coefficient 

H [j] Specific enthalpy 

Hfg [1L] Specific latent heat of evaporation 

k [I-] Heat conductivity 

L. [m] Mean path length 

n Droplet number density 

p [Pal Pressure 

Pr Reduced pressure 

Pr Prandtl number 

q" [iI•] Heat flux 

Re Reynolds number 

T [K] Temperature 

v [!] Velocity 

Xtt Martinelli flow parameter rN m21 

g -[-9 Gravitational constant 

M Multiplier 

We Weber number 

r [im] Surface roughness 

1

Variables and Symbols not included here are explained there where they appear first.

xi
Greek Symbols 

a Void fraction 

E Emissivity 

A [m] Characteristic wave length 

[-'-'] Dynamic viscosity 

p [m] Density 

a [E] Surface tension 

Mole fraction 

Accomodation coefficient 

ýo [Mkj] Molecular weight 

Subscripts 

a Absorption 

c Conduction 

C Cladding 

CHF Critical heat flux 

e Equivalent 

F Fuel 

FB Film boiling 

g Vapor 

gap Gap 

i Index 

I Liquid 

mac Macroscopic 

mic Microscopic 

nc Number of circumferential segments 

NB Nucleate boiling 

pg Constant pressure-gas phase 

p1 Constant pressure-liquid phase 

q Quench 

r Radiation 

s Saturation 

stf Stratified 

tot Total 

TB Transition boiling 

TP Two-phase 

w Wall 

web Weber

xii
Abbreviations

AMM Accident Management Measures 

CAMP Code Assessment and Maintenance Program 

CHF Critical Heat Flux 

CORA Out-of-pile-bundle test facility, FZK 

CSNI Committee on the Safety of Nuclear Installations 

DBA Design Basis Accident 

DNB Departure from Nucleate Boiling 

EPR European Pressurized Water Reactor 

ECC Emergency Core Cooling 

FZK Forschungszentrum Karlsruhe, Germany 

HTC Heat Transfer Coefficient 

ICAP International Code Assessment and Applications Program 

IMF Institute of Material Research, FZK 

INEEL Idaho National Engineering and Environment Laboratory, USA 

LOCA Loss of Coolant Accident 

LOFT Loss of Fluid Test Facility, INEL 

LWR Light Water Reactor 

MFB Minimum Film Boiling 

NPP Nuclear Power Plant 

NEPTUN Thermalhydraulic test facility, PSI 

PBF Power Burst Facility, USA 

PSI Paul Scherrer Institute, Switzerland 

PWR Pressurized Water Reactor 

RELAP Reactor Excursion and Leak Analysis Program 

RPV Reactor Pressure Vessel 

R5M3 RELAP5/MOD3 

SCDAP Severe Core Damage Analysis Package 

S/R5 Coupled Version of SCDAP and RELAP5/MOD3 

LP Low pressure 

FP Fission Products 

SFD Severe Fuel Damage 

TC Thermocouple

xiii
1 Introduction 

An important accident management measure (AMM) for mitigating severe accidents in 

Light Water Reactors (LWRs) consists of core quenching by the injection of cold water into 

the overheated, partially oxidized or even severely degraded core. In special cases the fuel 

temperature at the time of water injection may be larger than the rewetting temperature, 

which is the highest temperature at which a direct contact is possible between the coolant 

and a heated surface. A complicated flow and heat transfer pattern, known as post dryout 

two-phase flow, exists under such conditions, characterized by the formation of a wet patch 

on the hot surface which eventually develops into a moving quench front. The rate of 

quenching depends on the heat transfer regimes along the heated surface. At high surface 

temperatures corresponding to film boiling, the quench proceeds rather slowly as the liquid 

is separated from the surface by a continuous vapor blanket. As the cladding temperature 

is decreased to the minimum heat flux temperature (sometimes called Leidenfrost or 

quench temperature), a transition boiling regime is encountered where the intermittent 

wetting of the surface begins and the heat transfer rate increases with decreasing surface 

temperature. At a surface temperature corresponding to critical heat flux, most of the 

surface becomes available for wetting and intense nucleate boiling ensues, causing the fuel 

cladding to cool rapidly until the liquid saturation temperature is reached, below which 

the fuel cladding is cooled by single-phase liquid convection. 

Since this quenching process plays an important role in defining the consequences of a 

severe accident, full characterization of the core conditions during the accident requires a 

detailed consideration of the coupled phenomena of heat and mass transfer in the coolant 

channel and on the clad surface together with reliable modeling of possible chemical and 

material phenomena. 

The thermal, hydraulic and chemical phenomena taking place during quenching of 

degraded fuel bundles by water were recently considered by the Committee on the Safety 

of Nuclear Installations (CSNI) to be poorly understood and unsatisfactorily modeled 

by severe accident codes [17]. For instance, analysis of the TMI-2 accident [29] andsev

eral bundle experiments such as LOFT LP-FP-2, PBS SFD-ST and CORA [17] have 

established that quenching of a degraded core may enhance the clad oxidation process, 

increase hydrogen production and in turn, significantly increase the fuel rod temperatures. 

These trends are not always predicted by the major transient reactor codes. Moreover, 

in some cases even the surface temperature history during the preheating and quench

ing of nonoxidized fuel rods are not always predicted in the desired accuracy [20], [41]. 

Therefore, several experimental programs have been initiated to complement the present 

state of knowledge on degraded core quench. In parallel, under the International Code As

sessment and Applications Program (ICAP) and the Code Assessment and Maintenance 

Program (CAMP) [14], there is an international effort to review, assess and improve the

1
treatment of the quenching process in the major reactor transient analysis code systems. 

The code review process is typically based on utilizing data from separate-effect and 

integral test experiments for validating either individual physical models in the codes 

or their integral behavior. The reflood model in R5M3 [22] has been recently assessed by 

Analytis [1], [2]. It was shown that the predicting capability of the code could be improved 

by introducing two different heat transfer packages to account individually for reflood and 

general heat transfer conditions. A similar approach is used also in other codes such as 

RELAP5/MOD2 (R5M2) [23] as well as in the codes ATHLET [4], [5] and REFLA/TRAC 

[34] which use correlations to predict the quench front velocity. Nevertheless, it should be 

recognized that the need to rely on empirical correlations or to assume a special reflood 

heat transfer package stems mainly from the current lack of understanding of the physical 

processes controlling the reflooding phenomena and from inadequate modeling of the 

boiling curve near the quench front. Therefore further effort must be spent in order to 

arrive at a unified heat transfer model suitable for boil-off, reflood and non-reflood heat 

transfer calculations at low and high initial surface temperatures. 

A large amount of heat transfer data on quenching of fuel rods at relatively low initial 

temperatures is currently available [32], [30], [38],[39],[35], [27], [26], [43], [8], [47], [31], 

[45]. This information which was mainly generated for developing a reliable heat transfer 

model for LOCA analysis, may also be relevant for degraded core analysis. Many of the 

theoretical and experimental studies on the heat and mass transfer mechanisms during 

top and bottom reflooding have been reviewed by Yadigaroglu [50], Yadigaroglu et al. 

[52], Elias & Yadigaroglu [13], Collier [12] and Olek [40]. Recently, data on high tem

perature quenching become available upon completion of several programs on degraded 

core quenching, [17], [20]. The main difference between the low and high temperature 

quenching data is that in the latter case the post dryout heat transfer region extends over 

a longer section along the heated surface and, thus, becomes an important mechanism 

for cooling the cladding prior to complete quenching. Therefore, improved modeling of 

the film and transition boiling heat transfer regimes is essential for accurate degraded 

core analysis at high initial surface temperatures. It must also be recognized that at ele

vated core temperatures the thermal behavior of the fuel rods strongly affects the rate of 

clad oxidation. Thus the prediction of the core conditions in that case calls for accurate 

modeling of both the heat transfer and the chemical problems. 

The main objective of the present study is to develop and assess a unified heat transfer 

model which can be applied for both reflood and general heat transfer computations. In 

order to effectively benefit from the extensive testing and validation of the Chen transition 

boiling model [11] , the present study extends the original Chen formulation to improve 

its prediction capability for low pressures and low qualities. A stand-alone version of the 

new model was validated against measured transition boiling data over a wide range of

2
temperatures and pressures. This modified heat transfer model was implemented in the 

R5M3 code and further tested by comparing its predictions with data obtained in the FZK 

single rod test program [191 and in the PSI-NEPTUN bundle experiments [41]. In this 

report, the limitations of the current R5M3 heat transfer model are listed and discussed 

and a full description of the new transition boiling model is provided. 

2 The RELAP5/MOD3 Reflood Model 

2.1 RELAP5/MOD3 Code Basis 

The R5M3 is being developed at the Idaho National Engineering and Environment Labo

ratory (INEEL), for the analysis of both operational and accidental situations (transients 

and LOCAs) in LWRs within the range of Design Basis Accidents (DBA) [22], [24]. The 

coupling of R5M3 with the core meltdown code SCDAP, being developed by INEEL [22], 

allows the prediction of the main phenomena of core meltdown accident scenarios (beyond 

DBA) that take place within the reactor pressure vessel (RPV). In the coupled version 

SCDAP/RELAP5/MOD3 (S/R5) the quench thermalhydraulics are mainly modeled in 

the R5M3 part. 

The code R5M3 uses a one-dimensional, non-homogeneous and non-equilibrium ther

modynamic two-fluid approach. Accordingly, a six-equations formulation is used to repre

sent the phasic continuity, momentum and energy conservation equations. The two-phase 

two-component model accounts for the presence of a non-condensable (hydrogen, helium, 

air, etc.) component in the gas phase and for the presence of boron in the liquid phase. 

A set of constitutive relations is included to close and solve the system of equations, 

whereby the unknown terms (exchange terms and virtual mass term) in the conservation 

equations are defined as functions of the state variables. Since the closure laws depend on 

the flow or heat transfer regime, modern codes like R5M3 update the closure relationships 

at each time step according to the prevailing flow regime [51]. Figure 1 lists the different 

flow regime maps considered in R5M3. In these maps different sets of criteria are used to 

define the various regime boundaries [22], [24]. 

Since this work concentrates on the analysis of reflood situations in vertical pipes and 

bundles, the emphasis is focused mainly on the vertical flow map and wall heat transfer 

mechanisms (see Fig. 1). Relevant parameters are void fraction, wall conditions (wetted 

or unwetted), mass flux, pipe dimensions, and wall superheating [22], [24]. 

In DBA and beyond DBA scenarios the reflood quenching process covers over a wide 

range of flow and heat transfer regimes for certain times both single phase (vapor, liquid) 

and two phase (film, transition, and nucleate boiling) flow and heat transfer mechanisms.

3
Figure 1: Flow Regime Maps of the Constitutive Relations in R5M3 

These heat transfer modes for both reflood and non-reflood conditions are described in 

R5M3 in terms of a unique boiling curve which is shown in Fig. 2. 

Generallyr thetotal wall heat flux (q",) is given as: 

totq~t = h.9 (T. - 7q) + hi(T. - TI) 1 

where h is HTCrT is temperature and the subscripts Irg and w refer to the liquidi 

vapor and wallirrespectively. The liquid reference temperaturerTr1is typically taken as 

the saturation temperatureIT,. 

The total HTC is the sum of both h. and hL. R5M3 utilizes an energy partitioning 

logic to determine the heat transfer coefficients and the rate of vapor productionr EJ' for 

each heat transfer mode [22]. Detailed description of the heat transfer models is given in 

subsection 2.2 

The selection logic used in subroutine htrcl for defining the boundaries of the pre-and 

post-CHF modes that are of interest in reflood situations is given below in terms of the 

wall superheat: 

* Post-CHF 

SKif qFB > qTB then film boiling 75 K <AT g 600 K if qFB 5 qTB then transition boiling

4
Pre-CHFRegion • Post-CHFRegion

ATCHF -TCHF-T S: ATMFB =TMFB - TS

ATFB ATs - TW -s

Figure 2: Qualitative Representation of the Boiling Curve used in RELAP5 MOD3.1

if AT, > 600 K then film boiling is assumed.

* Pre-CHF { if qNB > qcHF then transition boiling 

ATs < 75 K if qNB •-qcHF then nucleate boiling 

where the subscripts FB, TB and NB denote film, transition and nucleate boiling, respec

tively. Note that the R5M3-approach does not require specific models for the temperature 

at the CHF, TCHF and the minimum film boiling temperature, TMFB, (see Fig. 2) to 

model the transition from one heat transfer mode to the other. 

The description of the reflood process requires the modeling of both the thermalhy

draulic mechanisms and the heat conduction process within the solid structure. While 

the reflood thermaihydraulics are embedded in the general wall heat transfer models, 

R5M3 uses a two dimensional heat conduction approach with a "Fine-Mesh-Rezoning" 

method only for the reflood structures in order to represent the physics of the process 

(see subsection 2.3, page 11). 

2.2 R5M3 Reflood Heat Transfer Models 

During quenching the hot surface passes from an essentially dry state to full contact 

with liquid [17]. This is done through a series of heat transfer modes ranging from free5

U)

A
or forced-convection cooling by steam, dispersed-flow film boiling, inverted annular film 

boiling, transition boiling, nucleate boiling, and finally single-phase convection to liquid. 

Corresponding to the boiling curve, the quenching process involves a shift from a film 

boiling heat transfer mode through transition boiling to a nucleate boiling regime. The 

code R5M3 has no special heat transfer reflood model. The description of the wall-to

fluid heat transfer in reflood situations is modeled within the general heat transfer logic. 

The most important reflood-relevant correlations used in R5M3 for vertical channels are 

briefly described below. 

*Nucleate boiling 

In the nucleate boiling regime heat is transferred mostly to the liquid (h. = 0). 

The liquid HTC, hl, is derived from Chen [10] as a sum of two terms: 

hNB = hi = hic + h. (2) 

where hm.c and h,,.c represent microscopic and macroscopic convection terms defined as: 

0O.79 ['045e,.49 

h,,ic = 0.00122- ki- 9• .• 4 .p. 9 AT4 pp0 .

7 5 S, (3) Or. 0.2 9 .5 Ai542 (Hf9 pg)0.

24 

-D P r°4 ReI' F S0.023 (4) 

The suppression factor, S1, is defined as: 

+ 0.12R eT *)' S(1 1 : R eTP < 32.5 

= (1 + 0.42Re0.)-' : 32.5 < ReTp < 70 

0.0797 : ReTP > 70 

with the Reynolds number for the two-phase region, ReTp, defined as: 

PReTP = min {70, [vI P1 al D.e] F1 .

25 10-4} (5) 

The Reynolds factor, F, is correlated as a function of the Martinelli flow parameter, 

F1.0 tX- 1 < 0.10 (6) 

2.35 (Xut 1 + 0.213)°0736 : X 1t > 0.10 

X;'0= min(lO0, [0.5] .1

6
0 Transition boiling 

As stated above the formulation is based on (but is not completely identical to) the 

Chen model [11]. Differences with the original work of Chen are discussed in section 7. The total transition boiling wall heat flux is modeled as a weighted sum of two components 

corresponding to vapor convection and liquid boiling, respectively: 

qTB (1- A ft Mstf)q Ma + fA Mtf q, (8) 

The wall-to-vapor HTC used to calculate q in eq. (8) is the largest value determined 

by the correlations for single phase convection in laminar, turbulent or laminar natural convection. In R5M3 the liquid heat flux, q7, is taken as the critical heat flux, q"HF, calculated by a look-up table (Groeneveld approach [16]). 

The multiplier for stratified flow, Mtf, and the void fraction multiplier, M,, in eq. (8) are correction factors applied to smooth the transition to stratified flow regime and 

to avoid heat transfer to vapor at low void fractions. Thus Ma is defined as follows: 

1 1.0 : a > 0.5 

M={=linear ramp : 0.0 <a <0.5 

0.0 : a = 0.0 

The fraction of wetted wall surface, fl, in eq. (8) is determined as: 

f, = exp[-v'ra-a(a)g'(G) min(15, LIT5 )] (9) 

where a'(a) and g'(G) are functions of the void fraction, a., and mass flux, G, defined by: 

a'(a) -- 0.05 4 + 0.075ag (10) 

with ag _< 0.999, and 

g'(G) = rnax(g,, g2) (11) 

with 

G 

g= 2.4- 135.6 (12) 

and 

g= 0.2-G 

g 135.6 (13) 

The transition boiling HTC to the liquid, hi, is then calculated by:

7
hi fI qH__._F .(14) AT, 

The vapor convective heat flux, q", is given by the relation 

q" = h9 (T. - Tg). (15) 

The convective heat transfer coefficient to the gas hg is determined by the Dittus 

Boelter correlation. T. and T. are the surface and superheated vapor temperatures, re

spectively. 

The main differences between the original Chen model and the transition boiling model 

implemented in the frozen R5M3 version is exposed in the Section 7. 

oFilm boiling 

Film boiling may occur during several flow patterns, namely inverted annular flow, 

slug flow and dispersed flow. In R5M3 [22] both conduction and convection to the vapor 

film and to the liquid (droplets) as well as radiation to the liquid (continuous blanket 

or dispersed mixture of droplets and vapor) are taken into account in the calculation of 

the total heat transfer coefficient. Convection between the vapor film and the liquid is 

included in the interfacial heat transfer model. 

In general the total heat flux, qFB, due to film boiling is given by: 

it 11(16) It 

qFB = qFB,g + qFB,"

The first term in eq. (16) is the heat flux to the vapor, calculated in R5M3 by using 

I, 

a suitable single-phase heat transfer coefficient. The heat flux to the liquid, qFBL, is a 

sum of the Sun's radiation model [46] and of the modified Bromley's conduction heat flux 

model [24], [211: 

I, f 

q,, (17) 

qFBj = qFB,r + qFB,c"

The liquid HTC due to conduction, hi, c, is expressed by the modified Bromley

Pomeranz correlation [24] as follows: 

hk2[g P, k(P - P 9)Hi Cp,° 0'25 

hlc= 0.62 P L r (P 9 

L Prg t AT, 3 J M.i (18) •(8 

The enthalpy factor, Hl9 , includes the energy absorbed by the superheated vapor 

surrounding the tube as:

8
Hl'g = Hig + 0.5Cpg (T. - T,). (19)

The parameter L in eq. (18) is expressed by the minimum critical wave length as: 

L ( )0.5 (20) = 9 ( Pi - p() where a is the liquid surface tension and g is the earth gravitation constant. 

The void fraction factor, M,, is given by the relation: 

S1.0 a <_ 0.2 

M,,= spline :0.2 < a<0.999 

0 : a = 0.999 

The property Cp9 is evaluated at T. while p., 1L, and k9 are evaluated at the film 

temperature (Tfijm =•2 

The radiation term in eq. (21) describes the radiation energy transfer between a wall 

and a mixture of vapor-liquid-droplet. The radiation heat flux is expressed by: 

qFB, = B(TF- T3) (21) 

where F' is a gray-body factor and B is the Stefan-Boltzmann constant (B = 5.678 x 10-8 W). 

The corresponding HTC, hi,,, is given as: 

/I 

qFB, =F' B (T. + T)(T•,2T) (22) h T, . T T, - (22) 

The gray-body factor, F', depends on the emissivities, ' and ef, and absorption coefficients, a. and af, of vapor and liquid. It is defined as follows: 

Pi= 1 (23) [(2(1 + !; + •)]"(23 

The factors (1, (2 and C3 in eq. (23) are given by the relations: 

(1 = [%(1 -69E)] (24) 

2 el(- ) (25)

9
1 (6 (3 = 0.111 + (1 - )(26) 

The emissivities, c and fi, are given as: 

E = 1- exp(-ag L.) (27) 

=l = 1 - exp(-aL Lm) (28) 

where Lm is the mean path length of the medium between the structures in radiation 

exchange. In R5M3 the following assumptions are made: 

* E = 0.02, 

* Lm = 0. 9 D, and 

c, = min (1 - exp(-O.9D, at), 0.75) 

The liquid absorption coefficient, al, is approximated in [22] as: 

aL = 1.11 aid. (29) 

where d is the droplet diameter. 

The maximum liquid cylinder diameter, d,., which may be formed in a tube with 

diameter D, assuming that all the available liquid forms a cylinder in the center of the 

tube, is given by: 

S= Da0.5 (30) 

On the other hand, the average droplet diameter, dwb, based on a Weber number 

criterion of 7.5 for the post-CHF, is calculated in R5M3 as: 

dwebWe a (31) 

Pg maX((Vg - vi)2, 0.005)( 

Consequently the droplet diameter, d, is determined in R5M3 from the relation: 

d = min(deb, d..). (32)

10
2.3 Reflood Heat Conduction

As mentioned in earlier sections the reflood process must be described by both (axial 

and radial) heat conduction within the heat structure and heat transfer mechanisms at 

the wall-fluid interface. In reflood situations a two-dimensional heat conduction equation 

is solved only for the reflood-structures. Additionally a fine-mesh rezoning scheme at 

the quench front position is activated. By this way the axial node, where the quench 

front is located, is subdivided into smaller meshes depending of the prevailing heat

transfer conditions. According to a set of predefined rules, the number of fine nodes is 

doubled or halved automatically by the code during the calculation, in order to adequately 

describe the physical processes of the quench phenomena. This is necessary because some 

parameters like heat flux and heat transfer coefficient are changing dramatically within 

very narrow regions, of the order of milimeters, whereas the mesh size normally used is 

typically 10-20 cm [221. 

If the reflood-structure is a typical fuel rod, R5M3 takes into account a two

dimensional heat conduction within the pellets and the cladding. The heat transfer over 

the gap is also modeled in R5M3 considering not only the gas conductance in the gap but 

also the radiation exchange between the pellet and the cladding [22]. According to [22] 

the total effective gap conductivity, htt,gap, is given as: 

htot,gap = hgap + hr (33) 

where hgap is the gas gap conductivity and h, is the radiation contribution. 

The conductance through the gas in the gap is given by: 

N 1k9 1 h -a= N •, [tt,, + 3 .

2 (rF + rc) + (771 + 772)] (34) 

where: 

nc number of the circumferential segment, 

N total number of circumferential segments (=8), 

kg thermal conductivity of the gas, 

tnc width of gap at the midpoint of the n-th circumferential segment, 

rF surface roughness of the fuel, 

rc surface roughness of the cladding, and 

m7, '72 temperature jump distance terms for fuel and cladding. 

These terms are defined as follows: 

0.024688 kg 7

771 + '7 = [B E e.po.5] (35)

11
where: 

Pg gas pressure, 

Tg gas temperature in gap, 

Qj mole fraction of i-th gas component, 

& accomodation coefficient, and 

o molecular weight. 

The constant 0.024688 in eq. 35 has the following units [(J kmoi K) 0 5

]. 

The radiation part h, is expressed as: 

1 

h, = B [ I + (RE)a (-L - 1)] •(F+T) CT (36) 

where: 

TF outer pellet temperature, 

Tc inner cladding temperature, 

RF outer pellet radius, and 

RC inner cladding radius. 

The gap conductance model of R5M3 is also able to consider different gases filling 

the gap, for example He, Ar, Kr, Xe, N, H, and 0, as well as the cladding deformation 

(ballooning) for the calculation of the effective gap conductivity. The proper modeling of 

the gap in codes like R5M3 is important for a correct prediction of quench front movement, 

as will be shown in the following sections. 

It must also be mentioned that the axial and radial heat removal of highly oxidized 

and overheated fuel rods becomes more complicated due to changes in material composi

tion and geometry during the course of severe accidents. The appearance of an oxide layer 

(ZrO2) due to cladding oxidation changes dramatically the axial and radial heat conduc

tion, which in turn influence the quench front progression, because the heat conductivity 

of ZrOT2 is almost ten times lower than that of metallic Zircaloy. This oxide layer is also 

very porous and wavy, which also may alter the flow conditions (turbulence, nucleation, 

etc.) near the wall. 

3 Experimental Programs for Code Validation 

Two experimental programs, the FZK single-rod program [36], [19] and the NEPTUN 

tests [41], were considered for the validation of the reflood model in R5M3. Both programs 

are relevant for degraded core reflooding. The single-rod tests are being conducted at FZK 

while the NEPTUN test program was performed at PSI. The two experiments are briefly 

described in this section.

12
3.1 The FZK Single-Rod Quench Tests 

A series of separate-effect 

segments at 

tests were carried out on Zircaloy PWR fuel rod cladding 

the tests 

the Institute of Material Research (IMF) of the FZK. The main objective of 

during 

is to study the generation of new metallic surfaces by cracking and fragmentation quenching at high temperatures [36], (191. Different unoxidized and preoxidized 

test series were performed with cladding segments (empty or filled with Zr0 2 pellets), where the temperature at the onset of quenching varied from 1273 K to 1873 K. 

3.1.1 Single-Rod QUENCH Rig Description 

Figure 3 shows the test arrangement. It consists of a small scale specimen enclosed quartz tube and heated in a by an induction coil. The outer diameter of the rod specimen the same of a PWR is fuel rod and the cladding thickness is 0.725 mm. The specimen mm long. is 150 In some of the tests the specimen contained ZrO2-pellets of 9 mm diameter, leaving 150 am annular gap. 

during 

As seen in Fig. 3 the quartz tube allows the observation of two-phase phenomena the quench phase (vapor generation, splashing of droplets, collapsed movement, liquid level 

of the single-rod 

etc.). Fast video photographs were taken for each test. Relevant parameters quench tests including the rod specimen cross section with the location of the thermocouples are shown in Figure 4. 

phases) 

The 

were 

temperature 

made 

measurement during the whole test period (preheating and quench by the use of different 

For water 

thermocouples depending on the test conditions. quenching tests without pre-oxidation, 

were used (spot welded). thermocouples welded on a rhenium foil For water quenching tests with pre-oxidation, were the thermocouples 

measured 

fixed 

not 

by a -. 2 mm Zircaloy ring. In several of the tests, the temperature were 

radial positions 

only at different axial positions on the outer clad surface but also in different (see Fig. 4a and 4b). In the quench tests with the cladding segments with ZrO2-pellets filled 

radially over 

the thermocouples were positioned axially along the segments and the fuel rod simulator. The radial temperature measurement 0.075 m consisted at the level of of three thermocouples. The first shielded thermocouple was located the center of in the ZrO2 pellet. The second one was located in the gap between the ZrO2 pellet and the cladding (in a groove on the pellet touching the cladding inner surface) and the last one spot welded on the outer cladding surface. 

3.1.2 Test Procedure 

Figure 5 depicts a typical temperature 

specimen 

curve for the pre-quench and quench phase. The is first heated in an argon (tests with no pre-oxidation) or in an argon oxygen

13
0

0

Induction M 

C,,u _

Quartz Tube

Valve

leh * nder

Hydrogen (to condenser + mass spectrometer) 

Fuel Rod Specimen 

0 

0 Specimen Temperature 

at onset of Quenching 

S100oo- 1600 C 

Specimen Length 

* 150 mm 

r (Ar+Oxygen) 

]==(Ar+Water) 

Water Temperature 

20-90C

Figure 3: Arrangement of the Single Rod Quench Tests at FZK

14

I

I

LL 'n
Rod Specimen 

I

0.122 m 

0.075 m 

0.028 m

(a) Axial Position of the Thermocouples

0.5375 mm 

0.465 nun:

Cladding 

Gap 

- Pellet

(b) Radial Position of the Thermocouples 

Figure 4: Rod QUENCH Rig: Cross Section and Temperature Measurement

15
U 1800 

2 1600 Argon atmosphere (no pre-Oxidation) CIS 

1400 Argon-oxygen atmosphere (pre-oxidation) 

E 

.- 1200 

1000 

800 C 

600 

400 

20O 

* tTime 

Heat-up phase during Pre-oxidation 120 sec to reach Quenching phase 

60 sec. duration dependent: desired quench 

on desired oxide temperature layer scale 

Figure 5: FZK Single-Rod Quench Test Conduction 

atmosphere (tests with pre-oxidation) by an induction coil until a predefined temperature 

is reached. Quenching is then initiated by raising a water filled quartz cylinder with a 

constant velocity of 0.015 m/sec or by flowing of saturated steam from the bottom of the 

heated cladding segment. Under that procedure the quench cylinder reaches the top of 

the specimen within 10 secondsrmuch before a quench front is established at the bottom 

of the test section. The induction coil is positioned outside the quartz cylinder (see Fig. 

3). The unoxidized specimen is heated during the first.10 min in an argon atmosphere 

(constant argon flow rate of 40 1/h) until a desired nominal quench temperaturerToris 

reached at the mid-elevation. The temperature at onset of quenching is kept constant 

for additional two minutesrbefore starting the quench phase. The quench phase begins 

when the quench cylinder touches the lower end of the specimen. At this time the heating 

is terminated. During the quenching phasersubcooled water is injected from the bottom 

by moving the quench cylinder upwards at a constant velocity [181r[35] and [19). From 

the tests performed with unoxidized empty tubesronly the tests listed in Table 3 were 

selected for code validation. In these tests three thermocouples were positioned axially 

along the cladding segment and one thermocouple was fixed on the inside cladding surface 

at a height of 0.075 m. All tests were performed under atmospheric pressure conditions.

16
3.1.3 Measurement Uncertainty

The behavior of the rod specimens during the pre-quench and quench phase at the 

QUENCH rig is determined by measurement of the temperature (at the cladding sur

face, in the gap and at the pellet centerline) as well as of the hydrogen generated during 

the tests. Additional metallographic examinations follow after each test to determine ox

idation scale, cracks formation, etc. In the case of quenching of unoxidized rod specimens 

spot welded thermocouples (bare-wire type) at the cladding outer surface were used in the 

majority of both empty and filled rod specimens. In the gap between pellet and cladding 

and at the cladding inner surface different methods of thermocouple attachement were 

used at FZK [36] as follow: 

"* sleeved thermocouple, in a groove on the outside of the pellet, wedged between the 

pellet and the clad inner surface. 

"* sleeved thermocouple, pressed against the inner clad surface by a disk. 

"* spot-welding of the thin foil, to which the thef-mocouple (bare-wire type) is attached, 

to the inner clad surface. 

The influence of thermocouple attachement methods on the temperature measure

ment during the quench tests at the QUENCH rig was investigated by AEA Technology 

[151 using a finite element heat transfer code aiming to identify the possible cause of 

measurement errors. 

A large number of possible sources of error are mentioned in [151 for the different 

thermocouple attachement methods in the case of water quenched tests. The different 

contributions to whole error are: 

"* fin effects 

"* added mass 

"• added surface area 

"* surface conditions 

"* thermal contact (air gap) 

"* enhanced boiling 

"* false junction

17
a effectiveness of thermocouple thermal contact

"* disk clad heat transfer 

"* thermocouple response 

"* pellet clad temperature difference 

"* pellet tube gap offset 

A comparison of measured clad temperature using different thermocouple types indicate 

that the boiling at the clad surface of the thermocouple position is being affected by the 

presence of the thermocouple and its attachement mechanism. Therefore, Fry at al. [15] 

conclude that the presence of the thermocouple on the outside of the Zircaloy clad can 

significantly affect the heat loss from the specimen surface, and hence its temperature 

history. These effects of thermocouple attachement upon the boiling heat transfer will 

dominate any other error sources of thermocouple error. In [15] error quantification for 

the cool-down tests using different model assumptions were made. Table 1 shows typical 

ranges of temperature differences predictedby theTAU models for the different attache

ment methods in comparison with the measured temperature differences. Of course, larger 

temperature differences are expected for water quenching tests. 

Table 1: Comparison of predicted and measured temperature difference for cooldown tests 

Thermocouple Attachement Measured TAU Predictions 

Outside Spot-welded -11 K (assumed) -11 to -16 K 

Inside Spot-welded +14 to + 28 K +1 to +14 K 

Inside pressed ±23 to +56 K +15 to + 49 K 

3.2 The NEPTUN Bundle Tests 

3.2.1 Test Description and Procedure 

The NEPTUN test facility was designed to study reflooding in bundle geometries [44]. It 

consists of 33 electrically heated rods and four guide-tubes. The layout and dimensions 

of the heater rods are shown in Fig. 6. 

The outer dimensions of the rods are similar to those of a PWR fuel rods except 

being half length in size (1.68 m heated length, 10.7 mm outer diameter and 1.33 pitch 

to diameter ratio). The complete bundle is placed in an insulated octagonal housing. A

18
DImnsions in [mm]

Kanthal ((alloy of Fe, Cr, Al and Co) 

SBoron 

W Inconel 60 

W Copper 

WI A12 03 

Figure 6: Cross Section of the NEPTUN Heater Rod

19
cosine shape axial power profile with a peaking factor of 1.58 is used. Five spacer grids 

are used axially at equal distances along the fuel assembly. The tests (5036 and 5050) 

were carried out under the pressure of 0.41 MPa. The rod power mounted to 2.45 kW. 

Rod cladding temperatures, fluid temperatures and differential pressures are measured 

at eight levels. The lower measurement level is located 50 mm above the beginning of the 

heated zone. The other measurement points are distributed at equal axial distances (232 

mm) along the rods. Each heater rod is instrumented with four to eight thermocouples, 

placed inside the cladding at the eight measurement levels. The five central heater rods 

are additionally supplied with external thermocouples (see Figure 7). Furthermore the 

flooding rate, exhaust steam flow rate, water carry over, absolute pressures and heating 

power are measured. The void fractions and the collapsed liquid level are deduced from 

the measured absolute pressure and the pressure difference. The pressure in the upper 

plenum of the test section is held constant during the experiments by a pressure control 

system. 

The NEPTUN facility consists, in general, of an inlet water supply system, a test 

section and an exit section. During the initial test period, the test section is kept in a 

saturated steam environment at a defined experimental pressure. The power in the heater 

rods is then switched on and the heater rod temperatures start to increase. A short time 

before the cladding temperatures reach the desired value, the test section is quenched from 

below with water at given temperature. The bundle power is held constant throughout 

the experiment, which is different to the FZK-tests where the power was shut off prio to 

quench. 

3.2.2 Measurement Uncertainty 

A discussion of errors made by measurements of quench-relevant parameters is given in 

[41]. In that reference there is no detailed quantification of the measurement accuracy in 

the NEPTUN test facility, so that the errors of the measured data were mostly estimated 

[41]. The Table 2 summarizes the estimated measurement errors and scattering of the test 

data of seven NEPTUN tests [41]. 

3.3 The Validation Matrix 

For the assessment of the R5M3-reflood model, different Single Rod Quench tests without 

pre-oxidation were selected. The test matrix includes quenching by water of empty and 

filled tube segments. In addition, one test is selected in which the specimen is cooled by 

steam flow. Although this last case is not a true quenching test, it provides essential data 

for analyzing the measurement uncertainties. Moreover steam quenching is representative 

for the flow conditions in the upper part of the bundle (Table 3).

20
Figure 7: Cross Section of the NEPTUN Bundle 

21

I

I I
Table 2: NEPTUN measurement errors and scattering of the data 

Quantity Probable error Largest scattering 

Flooding water mass flow ±5.3 

Flooding water temperature +0.50C 

Test section pressure ±0.03 bar 

Collapsed water level ±4 cmn 

Void fractions ±0.04 

Rod clad temperature ±50C 480C, exp. 5050 

between all rods without 900C , exp. 5036 

external thermocouples 

Quench times between all +1.2s 2.5s, exp.5050 

rods without external 14.2s, exp. 5036 

thermocouples at 

measurement level 4 and 5

Two NEPTUN tests (No. 5036 and 5050) were selected for the validation task. The first 

is a low inlet velocity test which has been used formerly as a base case in the assessment 

of the R5M2 code [41] and the R5M3 code [2]. The second represents a high inlet velocity 

case (Table 3).

22
Table 3: Test matrix for the validation of the R5M3-reflood model 

Test ToTub Specimen TC TC vin 

(K) Type Type Fixation (cm s) 

FZK Separate-Effect Tests 

ITE20115 1273 363 empty radial spot welded 1.5 

ITE22115 1673 363 empty radial spot welded 1.5 

ITE29115 1866 363 empty radial spot welded 1.5 

T24085 1475 363 filled axial spot welded 1.5 

T02106 1520 363 filled radial spot welded 1.5 

T21066 1673 363 filled axial spot welded 1.5 

T31085 1873 363 filled axial spot welded 1.5 

T24066 1873 363 filled axial spot welded 1.5 

T16106 1873 363 filled radial spot welded 1.5 

PSI NEPTUN Experiments 

5036 1030 411 Heater rod axial inside clad 1.5 

5050 1140 411 Heater rod axial inside clad 15

4 Test Modeling for Reflood Model Assessment 

The numerical models applied to represent experiment setup and test procedure for the 

calculation with the code R5M3 are described here. The results of assessmental calcula

tions related to the reflood model are discussed in the next subsections 4.1 and 4.2. 

4.1 The FZK Single-Rod Test Rig 

In modeling the test system and procedure it is recognized that the main quench phe

nomena take place within the quench cylinder which, upon reaching predetermined initial 

conditions, is pushed upwards at a constant velocity of 0.015 m s. The movement of the 

quench cylinder is modeled by specifying a constant inlet velocity of water at the bottom 

of the coolant channel. 

Radiation losses from the test section to the environment are initially high, mainly 

because of the high initial temperature level of the fuel rods. The inner and outer quartz 

cylinders are almost transparent to radiation. However, when quenching begins, the spec

imen temperature decreases within a few seconds to temperatures below 1000 K, the 

radiation losses to the environment do not strongly influence the quench front progression 

and the accompanying changes of heat transfer mode from film to transition and from

23
transition to nucleate boiling which typically occur at lower values of AT,. The radiation 

exchange option of R5M3 was activated in the simulation throughout the transient. The 

argon flowing through the annular gap around the quench cylinder and the outer quartz 

cylinder does not play an important role in the quench process. Therefore, it was ne

glected in the simulation. As in the experiment where tests are carried out at atmospheric 

pressure, a constant outlet pressure of 0.1 MPa was used in the simulation. 

Figure 8 shows the nodalization scheme for the R5M3 simulation of the FZK separate

effect quench tests. 

The time-dependent volumes (Tmdpvol) and junctions in Figure 8 are used to repre

sent the initial and boundary conditions at the inlet and outlet of the test section (tem

perature, mass flux rates, pressure). The branch elements represent the fluid conditions 

below and above the test section ("Pipe Element"). The quartz cylinder is represented 

as a pipe element with 15 internal volumes, where the initial thermalhydraulic conditions 

(pressure, temperature, mass flux rate, quality etc.) of the fluid channel are fixed. The 

rod specimen (filled or empty) and the quartz cylinder are modeled as heat structures. 

The first heat structure represents the rod specimen considering its material composi

tion. This structure is coupled by convective boundary conditions with the fluid channel. 

It also allows for the specification of the initial distribution of the specimen temperature. 

The second heat structure is coupled by convective boundary conditions with the fluid 

channel on one side and with a constant temperature (representing the environment tem

perature) on the other side. This heat structure represents a heat sink and it is required 

to model the large radial heat losses from the test rig to the environment. The thickness 

of this outer tube was found to have no effect on the results. 

Figure 8(b) shows the radial nodalization of the test specimen. The cladding is sub

divided into three rings. More nodes may be necessary to model oxidized clad because of 

the large difference between the thermal properties of oxidized material. The ZrO2 pellet 

(when it exists) is modeled as a single node since the radial temperature variations are 

typically small.

24
Relap Heat 

Structure 

Simulator Rod

Interval Number 

(b) 

Mesh Points =

80 Outlet 

Tmdpvol 

50 Upper Plenum 

(Branch) Relap Heat 

Structure 

1115 

1114 

1113 

1112 

fill0 

1110 

1109 

1108 

1107 

1106 

1105 

1104 

1103 

1102 

1101 (ie 

I Quartz Wall 

10 Lower Plenum 

(Branch) 

3 Water 

I- dvo

I 2 345 iiJ!11L

2 3456 r

M Pellet = Gap M Cladding

Figure 8: Axial and Radial Nodalization of the Single-Rod Quench Test Rig

25

(a)
4.2 The NEPTUN Test Section

The model used to describe the NEPTUN test section is shown in Fig. 9. It is identical 

to the scheme used formerly in the assessment of the R5M3 code at PSI [2]. The model 

consists of a pipe component to represent the fluid volumes in the test section. The 

heater rod is divided into nine radial layers to accurately simulate the radial temperature 

distribution and material composition. 

The inlet water system is simulated by a time-dependent volume connected to the 

coolant channel by a time-dependent junction. The upper plenum is represented by a 

time-dependent volume with constant pressure, connected to the upper end of the channel 

by a single junction. Guide tubes and housing were neglected in the calculation since they 

have only small effects on the flooding behavior of the bundle [41]. 

The effect of different nodalizations on the R5M2 code predictions was studied in [41], 

using 10, 18 and 32 volumes for the bundle for a given number of fine mesh nodes in the 

heat conduction elements. It was found that in the high flooding rate experiment (test 

number 5050) the effect of the nodalization on the cladding temperatures is small. In the 

low flooding rate experiment (test number 5036),the effect of the number of nodes on 

the results was not systematic. The present calculations were performed with 18 volume 

elements in the coolant channel and 16 fine mesh points per heat slab as recommended 

in [41]. 

5 Results of Assessmental Calculations with 

RELAP5 MOD3 

Using the nodalization scheme mentioned above, base case calculations were carried out 

with the code version RELAP5 MOD3.1 (R5M3) for the QUENCH rig tests. These cal

culations started with a steady state calculation to check for input errors and to reach 

quasi steady state conditions before the initiation of transient calculation. The aim of 

the transient calculations (pre-quenching phase) was to met the specimen temperature at 

onset of quenching for each test. Then calculation of the quenching phase followed. 

In the case of the NEPTUN tests, we adopted the same base case as in [41]. Therefore 

a repetition of this work is not necessary in this place. 

Since temperatures are the only data measured in the test program of FZK, the presint 

analysis will based primarily on the predicted temperatures. In addition, other calculated 

parameters will be shown to enhance the physical understanding of the results.

26
(a) 

0 

Radial Intervals = 

(b) 

Mesh Points =

Time Dependent Volume 

Single Junction 

17-Hydrodynamic 

V olume 

1 Heat Structure 

Junction 

14 Pipe 

Time Dependent Junction 

Time Dependent Volume

1 2 3 4 5 6 789

2 3 4 5 6 7 8910 r

Figure 9: Nodalization Scheme of the NEPTUN Test Section

27
5.1 FZK Single-Rod

A typical sample of rod temperature histories at various radial locations in the middle of 

the specimen is shown in Fig. 10. The preheating process lasting for 720 seconds was also 

simulated by R5M3 but is not shown. 

In all the curves of Fig. 10, the temperature changes gradually at the early phase of the 

transient due to early cooling by steam, two-phase mixture produced mainly at the lower 

part of the coolant channel and due to radiation. At a certain time, each curve exhibits a 

'knee' shape when the temperature drops sharply to almost a steady value corresponding 

to the local fluid temperature. This marks the rewetting of the measured point along the 

test rod. During a period of few seconds around this knee, the rod experiences dramatic 

changes of the heat transfer mode from film boiling through transition and nucleate boiling 

to single phase to liquid (see Figure 11). 

The quench temperature, T., corresponding to the knee point can be defined as in 

Barnea et al. [6] at the intersection between the tangent line to the temperature-time 

curve at the point where its slope is the largest, with the tangent to the curve before 

quenching. The quench temperature marks the onset of rapid surface cooling caused by 

an enhanced heat transfer rate without complete liquid-solid contact. The measured T. 

at the inner and outer sides of the cladding are about 800 and 700 K respectively. The 

corresponding predicted Tq are significantly lower (about 720 and 600 K respectively). 

The quench temperature, Tq, corresponds roughly to the minimum film boiling temper

ature, TMFB, at which the film boiling and transition boiling heat fluxes become equal. 

The systematic underprediction of Tq depicted in Fig. 10b, c indicates, therefore, an 

underprediction of the transition boiling heat transfer. This is further demonstrated in 

Fig. 11 which shows the predicted heat flux, critical heat flux and heat transfer coefficient 

at the mid-level of the specimen's outer surface. The rewetting process is shown to be 

accompanied by a sudden increase in the heat flux as the heat transfer mode changes 

from film boiling to the more intense boiling mechanisms characteristic to the transition 

and nucleate boiling regimes. However, the predicted peak heat flux in Fig. 11 a is very 

low relative to the local critical heat flux. This is a direct result of the transition boiling 

model used in R5M3 as described in eq. (9). 

The measured clad outside temperature prior to quenching is shown to be lower than 

the calculated (see Fig. 10c). This is assumed to be caused by local measurement errors: 

the size of the thermocouple and of its welded bead is of the order of the thickness of the 

thermal boundary layer developing along the specimen. The layer is smaller at the bottom 

of the specimen. The rhenium foil acts as a cooling fin and the thermocouple is to a larger 

extent exposed to water droplet in the vapor film surrounding the cladding. The same 

effect is encountered in the case of steam cooling tests. For this case, the Fig. 12 exhibits 

a comparison between predicted and measured surface temperature at three different

28
1600 Single Rod T02106 

0) ":3• 0 Experiment ' -- - Frozen R5M3 . . . 

Z 1200 

E 1000 

(D 

( 800 

2 600 C 

o 400 (a) 

720 736 752 768 784 800 816 

Time (s) 

1600 Experiment 

1 - Frozen R5M3(pellet outside) 

- - Frozen R5M3 (cdad inside) 

S1200 

(D 

to 800 

0.0 

E 600 0 

400 (b) 

200 . 720 736 752 768 784 800 816 

Time (s) 

1 6 0 0~ 0 ..... .. Experime

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

1400 

":3 0 Q .. .. .. 

1200 (D 

E1000 

(D 

V 800 

:3 

0 600 

~~0 

400 (c)0 

200 720 736 752 768 784 800 816 

Time (s) 

Figure 10: Comparison of the Original R5M3 Predictions with the Measured Temperatures 

in the FZK Single-Rod Test No. T02106

29
1.5e+06 

1.3e+06 

1.1e+06 

9.4e+05 

7.6e+05 

5.8e+05 

4.0e+05 

2.2e+05 

4.0e+04 

-1.4e+05 

-3.2e+05 

-5.0e+05 

72

1.2e+04 

1.0e+04 

8.0e+03 

6.0e+03 

4.0e+03 

2.0e+03

u.ueuu 1% . an I

u.ue'+vU v. . . . .. . . . . . . . 720 728 736 744 752 760 768 776 784 

Time (s)

792 800 808 816

Figure 11: R5M3 Predictions of a. Surface Heat Flux and Critical Heat Fluxrand b. 

Surface HTC in the FZK Single-Rod Test No. T02106

30

4i

E 

X 

_=

728 736 744 752 760 768 776 784 792 800 808 816 

Time (s)

() Single Rod T02106 

- -

Froze RSM3HeatFlu 

Frozen RJ

0

E 

0 

U 

S 0= 

U 

S 

I-

'a 

-r

(b) 

Heightoo.075 m, 

.L .- ... .-. .... - • • ... . ...:....... i......
levels along the specimen. In all cases the measured temperatures are lower than the 

calculated surface temperature and their measured values correspond to an intermediate 

values between the surface temperature and the steam coolant temperature. 

The predicted clad inner surface temperature prior to quenching (part b in Fig. 10) 

is generally in good agreement with the measured data. It should be recognized that the 

measured inner surface temperature may be overestimated because of the finite size of 

the inner thermocouple which makes it sensitive to possible radiation and convection heat 

transfer from the Zr0 2 pellets. Therefore, it is expected that the measured gap tempera

ture will lay between the predicted inner clad temperature and outer pellet temperature. 

5.2 NEPTUN Facility 

The results obtained with the frozen version R5M3 for the test 5036 are the same as 

predicted in [45]. Figure 13 shows a comparison between measured and predicted cladding 

temperature. As can be seen in Figure 13 the predicted temperature does not show the 

typical 'knee'-point of a rewetting process. The initial cladding temperature during the 

first 35 s coincides very well with the measured one. Hereafter the code predicts a higher 

heat transfer coefficient that leads to an earlier 'turn-around'-point and to faster cladding 

temperature decrease compared with the measured data. Hence both the quench time and 

the quench temperature are missed. This also indicates that the modeling of the post-CHF 

heat transfer mechanisms in the frozen version of R5M3 is not correct. 

6 Stand-Alone Assessment of the RELAP-Reflood 

Model 

In order to assess the original transition boiling model implemented in R5M3 as part of 

the reflood model, a stand-alone program of this model was coded outside the RELAP

code and validated against steady-state experimental data. In particular, steady-state 

data measured in a heat flux controlled mode, [9], and in temperature controlled mode, 

[49], [28], were considered for the validation of the stand-alone model. 

6.1 Validation Matrix 

The Table 4 gives the test data used for model validation.

31
CU 

CL 

E 

() 

-o 

0

1600 

1400 

1200 

1000 

800 

600 

400 

200 

7:

730 740 750 

Time (s)

730 740 750 

Time (s)

760 770 780

760 770 780

760 770 780

Figure 12: Measured and Predicted Clad Outside Temperatures and Steam Coolant Tem

perature for the Steam-Cooled FZK Single-Rod Test No.T16046

32

740 750 

Time (s)

0 730

Single Rod T16046 

F _oExperiment " 

Frozen R5M31 0' 

0 6 

00 0 

0 O0 0...... .. 

"(a) Height=0.125 m.

-. 1600 

S1400 

9,= 

m 1200 0. 

E 

(D 1000 

:2 800 

o 600 

S400 

200 72

ýD - Frozen R5M 

0o) 

. ~ 00 ~ . ... 0 

000000

to 00000 000

Height=0.075 m

(b)

0

E 

0) 0D 

0 

,a 

0

1600 

1400 

1200 

1000 

800 

600 

400 

200 

72

--- Frozen RSM3 

• o . , • -. . . . . .: 

.0 ..7 - - . . . . . 

0000000 000 00 =~*~0000 0000000 -

"Height=0.025 m 

(C )•

:0

I I T l i I I I I I I I'i'lllTI I
7 

CL 0. 

E 

0 

0 

0

1200 

1000 

800 

600 

400

200 

Figure 13: Meas

0 40 80 120 160 200 

Time (s) 

ured and Predicted Clad Temperature

240

in the NEPTUN Test 3056

Table 4: Test Matrix for Stand-Alone Validation of R5M3-Reflood Model 

Reference Test Type Main Parameters 

Bennett et al.[9] Heat flux controlled 0.0126 m tube diameter 

6.9 Mpa 

300 < G < 5500 kg/m 2 -s 

Weisman et al. [47] Temperature controlled 0.1 < p < 0.4 MPa 

heated by Mercury 16 < G < 46 kg/m 2 -s 

Johannsen et al. [28] Temperature controlled Hollow Cylinder 0.050 m long 

od = 0.032 m, id = 0.010 m 

0.1 <p < 1 MPa 

25 < G < 200 kg/m 2 -s 

3 < AT2, < 30 K

The symbols in the Table 4 have the following meaning: 

G mass flux 

od outer diameter 

id inner diameter 

ATi, inlet subcooling.

33

II
6.2 Discussion of Results

The Figure 14 compares the predicted and measured data. Many of the measurements 

are underpredicted by one or more orders of magnitude. The data shown in Fig. 14 cover 

a wide range of thermalhydraulic conditions: Bennett et al. [91 report measurement in a 

vertical electrically heated 12.6 mm inner diameter tube at 6.9 MPa and mass flux of up 

to 5500 kg m2-s. Johannsen et al. [28] report measurements of CHF and transition boiling 

heat flux in a hollow cylinder of 50 mm length with an outer diameter of 32 mm and an 

inner bore of 10 mm. Data are measured in the pressure range of 0.1 to 1.2 MPa, mass 

flux of 25 to 200 kg m2 -s and inlet subcooling of 3 to 30 K. An electronic feedback system 

was applied to control the temperature of the heat transfer surface and thus stabilize the 

otherwise unstable boiling process. Weisman et al. [49] conducted transition boiling heat 

transfer tests at low pressure (about 0.2 MPa) using a loop in which the test sections were 

heated by hot mercury. The inlet linear water velocities were 0.017 to 0.048 m s which 

correspond to the lower end of the reflood range of conditions during LOCA. 

Careful examination of Fig. 14 shows that the original R5M3 model predicts reasonably 

well mostly the high quality post CHF data in which the convective heat flux to vapor is 

the dominant term. Clearly, the introduction of qCHF instead of the heat flux function, 

q (p, AT.), suggested in [11] effectively reduces the contribution of the boiling term in the 

transition boiling region and causes the R5M3 model to underpredict all the data in the 

low quality range in which the boiling term is important. This is best demonstrated by a 

typical case shown in Fig. 15 which compares the measured and predicted heat flux along 

the pipe wall in run No. 5251 of Bennett et al. [9]. The predicted total heat flux is the 

sum of a vapor convective term, q, and a boiling term, qj'. It is seen that the predicted 

boiling term is negligible throughout the post CHF region. The total heat flux is thus 

underpredicted.

34
1 E+1

1E+O 

"CM 

E 

1E-2 

IE-3 

1 E-3 1 E-2 1 E-1 1E+O

Measured q" (MW/m2) 

Figure 14: Measured vs Stand-Alone R5M3 Model Predictions of Transition Boiling Data

35

1 E+1
1.2

1 .0t

E 

x 

to 

0) 

x

0.8 

0.6

0.4 -

0.2 +

fln 4.0 4.0 4.2 4.4 4.6 4.8 

z (M)

5.0 5.2 5.4

Figure 15: Stand-Alone R5M3 Model Prediction of Run 5251 in Bennett et al. (1967)

36

q"(tot, measured) 

0 0 0 0 0 0 0 0 0000 0

q"(tot, predicted) 

q " (cno nv. , npr e~d i c ted) R5M3 Model 

Bennett Run 5251 

G=1356 kg/(m2 s)

q"(boiling, predicted)

I- .9

5.6
7 Limitations of the Current Transition Boiling 

Model 

The calculational results obtained in subsection 5 with the frozen R5M3 version and 

in subsection 6 with the stand-alone transition boiling model point out the need to 

improve the reflood prediction capability of the code. The systematic deviations between 

the predicted and measured quench temperatures outlined above result mainly from the 

implementation of an inappropriate and physically erroneous interpretation of the Chen 

transition boiling model in the R5M3 code. In order to better understand the limitation of 

the current transition boiling model, Table 5 summarizes the main features of the original 

Chen formulation in comparison with the current R5M3 model. It is noted that some of 

the constants (0.05 and 0.075 instead of 0.005 and 0.0075) in the correlation of the wetted 

area fraction, f, (eq. (9)), are modified in the R5M3 code in a non--consistent manner. 

The original Chen correlation with the coefficients 0.005 and 0.0075 was programmed and 

validated against different tests data obtaining very good results. On the contrary, if the 

coefficients 0.05 and 0.075 are used, the original Chen graphs cannot be reproduced. As 

can be seen in the figures 14 and 15, the Bennett data cannot be well predicted, if the 

R5M3 model with the coefficients 0.05 and 0.075 is used. Therefore the coefficients used 

in the original Chen model and not the ones used in R5M3 as stated in [22] are correct. 

The wall superheat term in fi is arbitrarily limited in R5M3 to v'A1T • 15. This limit 

has a strong effect on the predicted Tq in R5M3. More significantly, the original expression 

suggested in the Chen model for calculating the wall heat flux is altered by using the 

critical heat flux to replace the complicated heat flux function, qj (p, AT,), derived by 

Chen. By doing that, R5M3 practically "killed" the liquid component of the correlation. 

Since qj'(p, AT.) >> qcHF, the boiling part of the R5M3 transition boiling model is 

underpredicted. This affects mainly tests at low quality in which boiling is the dominant 

heat transfer mechanism in the transition boiling regime. Finally, there is a peak in the 

liquid heat transfer coefficient at 271.2 kg m2/s in the R5M3 transition boiling model and 

at 2712 kg/m2/s in the original Chen model. This peak is irrelevant in R5M3 since the 

liquid fraction is multiplying qCHF which is much less than q,,. recommended by Chen. 

Consequently, the R5M3 liquid component is very small and not important (mistake in 

R5M3 neglects the boiling part of the transition boiling model) while in the original Chen 

model the boiling component is important. This is well demonstrated in the figures 15 

and 19. 

The modifications introduced in R5M3 to the Chen model have resulted in several 

physically erroneous trends beside the systematic underprediction of Tq. First the pre

dicted dependence of the quenching rate on the coolant inlet mass flux for values higher 

than 270 kg m2 .s has been shown to be wrong, [2]. Moreover, since f, < 1, the heat flux

37
RUGGLES et al.

1.2

"IE

100

Tw- Tsat

Figure 16: Heat Flux versus Wall Temperature predicted by RELAP5/MOD3 for G 

20kg/sm2Fp = O.IOMPa and saturated conditions

38
Table 5: Comparison of RELAP5 and the Original Chen Model for Transition Boiling 

Heat Transfer

Original Chen T.B. Model

q" =flq' + (1 - fL)q1 

q= q'(p, ATs) >> qCHF 

q9 = hg(T, - Tg) 

h9 from momentum transfer analogy 

f, = exp[- v7a(a) f(G) v/ET] 

a(a) = 0.005/(1 - a40 ) + o.0075a 

f(G) = max(fl, f2) 

fi = 24 - G/135.6 

f2 = 0.2G/135.6 

Limits: x < 1 ,fq' g qCHf

RELAP5/MOD3 Model

q= fiMstfq7 + (1 - fMstf)qMa 

qI = qCHF 

It 

qg = hq(T. - T9 ) 

h9 from Dittus Boelter correlation 

f= exp [-Vfi8 a(a) f(G) min{15, x/aT}] 

a(a) = 0.05/(1 - ao4 0 ) + 0.075a 

f(G) = max(fl, f2) 

fi = 2.4 - G/135.6 

f2 = 0.2G/135.6 

No limits applied

applicable for degraded core reflood and to implement it in the R5M3 code instead of 

the RELAP-transition boiling model. Unlike other models which are based on history 

parameters such as TMFB or TCHF (which are not readily available at each time step by 

the R5M3 method of calculation), the suggested model has the additional advantage of 

utilizing only local channel conditions for predicting the wall heat flux. The new model is 

fully described in the next section 8. Since several versions of the Chen model have been 

used by different investigators and since the original report by Chen et al. [11] contains 

a number of typographical mistakes, the next section 8 is presented in a self-contained 

manner repeating some of the formulations presented in [11].

39
fully described in the next section 8. Since several versions of the Chen model have been 

used by different investigators and since the original report by Chen et al. [11] contains 

a number of typographical mistakes, the next section 8 is presented in a self-contained 

manner repeating some of the formulations presented in [11]. 

8 The Suggested Transition Boiling Model 

The mechanistic approach suggested by Chen et al. [11] for predicting q, and f, in eq. (8) 

is generally adopted with several modifications to facilitate its implementation in large 

system codes such as R5M3. The boiling component, q;', is calculated as the average 

heat flux during the short period of contact between the liquid and the superheated wall. 

Following the Chen model, a three-step process is postulated to describe the mechanism 

of heat removal by a film of liquid at the wall: conduction heating of the liquid film, 

nucleation and bubble growth within the liquid layer and, finally evaporation of a residual 

liquid film at the wall (see Fig. 17). 

8.1 Total Boiling Heat Flux 

The total heat flux is then a sum of three terms: 

- q•=tl 1 + +t12 12 +t2 + t 2 V(p, AT.) (37) 

where the subscripts 1, 2 and 12 denote the three steps involved, and O, and t1 are 

the absorbed heat per unit area and the time associated with each step, respectively. 

P(py, AT,) is an empirical correction function of the reduced pressure, pr = PPerit, and 

the wall superheat, AT,. This function is introduced to improve the model predictions at 

low pressures and low superheat with respect to the original Chen formulation [11]. 

The first heat transfer mechanism in eq. (37) corresponds to the superheating of a 

thermal boundary layer in the liquid in contact with the heated wall. In this period 

heat is transferred from the wall via pure conduction until a layer of liquid equal to the 

nucleation bubble radius is superheated to the nucleation superheat temperature. This 

process can be described by solving the transient conduction equation for semi-infinite 

bodies to determine the heat flux at the heating surface. The solution suggested in [48], 

[25] and [11] is: 

=1 R ap Cp1 T 

0.213 pi Hfg 

t12 1 0.24 0. p, Pa Hfga- Tx PTI pr + AT (38) 

40
Conduction Heating Bubble Growth Mficrolayer Evaporation 

1-12 2 

Figure 17: Schematic description of the Suggested Transition Boiling Model 

where ~R is the ideal gas constant, or is surface tension, C,, is heat capacity, p is pressure, Hf2 

is the specific latent heat of evaporation and al is the thermal diffusivity. The coefficient 

pr is defined by: 

PT (k p C,). (3, (k p C,)1 (I9 

The second heat transfer process in eq. (37) refers to the bubble growth period. Fol

lowing Chen et al. [111 it is assumed that the heat absorbed by the bubbles comes entirely 

from the superheated liquid layer at the wall, thus: 

012 = 0 (40) 

The time required for a bubble to grow to its final size can be readily determined by 

solving the bubble growth equation. However, to avoid the necessary numerical solution 

of the resulting implicit equation relating the bubble size and time, the period of bubble 

growth to a maximum size of 15 scm is correlated as: 

t1 AA TB, (41) 

where the coefficients A and B are given as functions of the reduced pressure, pT:

41
- 12.031 + 72.62p, - 167.29p2 pr < 0.136 (42) -8.1724 + 25.14p, - 26.29p2 pr > 0.136 

and 

B = • -0.4379 - 15.52pr + 49.32pr p, _< 0.136 (43) = -1.3052 - 3.137pr + 3.827p2 Pr > 0.136 

Subsequent to the bubble growth and escape process a thin superheated liquid micro

layer is formed on the wall, [48]. Heat absorbed from the wall by that layer is given in 

[11]: 

8 T•P1 ( 1 GCi•-T2\ (44) 

p9 AT ( Hf9 2) 

The time, t2, required for this evaporation process may be determined by estimating 

the thermal conduction from the wall to the evaporating film. Assuming a linear tem

perature gradient in the liquid film from a value T,,, at the wall to T, at the evaporating 

interface, we obtain: 

32 p, e1- T2 C-1 AT 

32 2 C 1 T (45) k, p Hp2 AT 1 Hfg 2 

The last term in eq. (44) and in eq. (45) accounts for the initial superheat of the liquid 

layer. At low pressures the film evaporation heat, 02, is dominant. The conduction heating 

process, 01 is more important at high pressures and larger degrees of superheat. 

A correction function, Ob(pr, AT.), was introduced in eq. (37) to better fit the data 

at low pressures and low superheat. An examination of a large amount of experimental 

results yielded a correlation for 7P in the form: 

?,(p•, AT,) = 1 + 3 exp (-0.42P32AT) (46) 

Note that IP(p, AT,) approaches unity as the wall superheat or the pressure increase. 

8.2 Liquid Contact Area Fraction 

The model suggested by Chen et al. [11] for estimating the liquid contact area fraction, 

fl, is adopted here with some modifications to improve its prediction capabilities at low 

pressures and low qualities. The present model predicts f, as an exponential function of 

the form: 

= exp [-a(a)g(G)(1.8AT8 )' 1 (47)

42
where a(a) and g(G) are functions of the void fraction, a, and the mass flux, G, given by: 

0.005 

a(a) - 1 - 40 + 0.0075a (48) 

and 

g(G) = max(gl, g 2) (49) 

where 

G 

g, = 20-0.635.6 

G 

g2 = 0.2 G (50) 135.6 

The power coefficient, n, in eq. (47) is taken as 0.5 in Chen et al. [11]. Comparison 

with data at low qualities suggests a quality and pressure dependent function of the form: 

n = 0.6 + 0.12 exp (-pl05 ) - 0.24x (51) 

where x is the local quality and p is the pressure. Equation (51) was found to fit well the 

available steady-state data on transition boiling. Table 6 compares the present model 

with the original transition boiling formulation by Chen et al. [11]. 

Equation (51) concludes the present modified transition boiling model, which can 

be summarized as follows: knowing the local quality, pressure, void fraction and vapor 

temperature one can calculate q,, qg, 0 and f, which, in turn, can be substituted in eq. 

(8) to yield the total heat flux q.". Recognizing the various approximations involved in 

the models for q, and fl, a limit must be applied as follows: 

fq? <- qCHF (52) 

where qcHF is the critical heat flux which can be determined from suitable correlations 

or from a lookup table as done in the R5M3 code. 

8.3 Implementation of the new Transition Boiling Model 

The transition boiling model developed in section 8 was implemented in the frozen 

RELAP5 MOD3.1 code version, getting the new code version RELAP5-MOD3.1-FZK 

(R5M3-FZK). The implementation affects only the formulation of the subroutine pstdnb 

of the R5M3 code where the post DNB-heat flux and the HTC are calculated. The 

new model utilizes only local state variables calculated by the R5M3 code like static 

quality, actual void fraction, mass flux, liquid and vapor temperature, phasic velocities, 

etc. and does not require other history parameters, such as quench position, critical heat

43
Table 6: Comparison of the Original and Modified Chen Model for Transition Boiling 

Heat Transfer

q" fqj + (1 - fl)%g 

q= q' (p, AT.,) 

q= ho(T, - Tg) 

ft = exp [-(1.8)na(a)g(G)AT8 ] 

n = 0.5 

a(a) = 0.005(1 - a40) + 0.0075a 

g(G) = max(g1 , g 2) 

g, = 24 - G135.6 

g2 = 0.2G135.6 

Limits: x < 1 , fjq1 < qCHF

flMstfq' + (1 - fLMstf)lMm .,f q, q 

qt= q '(p, AT,) 

q.= hg (T. - Tg) 

fL = exp [-a(a)g(G)(1.8AT.)n ]V)(pr, AT.) 

n = 0.6 + 0.12 exp (- 22 1.2p,) - 0.24x 

4 = 1 + 3exp [-0.42p, 2ATl 

a(a) = 0.005(1 - a40) + 0.0075a 

g(G) = max-(gi, g2 ) 

g, = 20 - 0.6G/135.6 

g2 = 0.2G/135.6 

Limits: x < 1 , fq qcHF

flux, and minimum film boiling temperature, which are not available at each time step. 

Summarizing, it can be pointed out that the suggested transition boiling model goes 

back to the original Chen's formulation and extends them for better prediction of reflood 

situations characterized by low pressure and low quality. 

9 Validation of the New Transition Boiling Model 

9.1 Stand-Alone Validation 

A stand-alone version of the new transition boiling model was formulated for model 

validation. In this model the same procedure described in section 8 is used to predict

44

Original Model i Modified Model
the total post critical heat flux for given local conditions. Furthermore, non-equilibrium 

effects are also included in the model according to Chen [11] as follows: 

X Hfg(p, TS) T9-TI(53) 

Xe Hg (p,Tq) - HI(p, Ts) T= -1 

where B(p,) is an empirical function of the reduced pressure, given in [11] as: 

B(pr) = 0.26 (54) 1.15 - '6 

Equation (53) is solved iteratively at each point along the test section to yield the 

quality, x, and the vapor temperature, T.

Figure 18 shows the predicted vs measured total heat flux. Considering the wide range 

of parameters covered by the data (362 data points in the range of up to 7 MPa pressure 

and mass flux up to 5000 kg m2.s), the new model is shown to predict the data reasonably 

well with a standard deviation of about 40%. Additionally, the results in figure 18, where 

the predicted heat flux for Bennett fits very good the measured data, demonstrates that 

the coefficients 0.005 and 0.0075 used in the original Chen correlation and not those used 

in RELAP (0.05 and 0.075) are the right ones. 

Figure 19 depicts the new model prediction of run 5251 of Bennett et al. [9]. The 

boiling component of the heat flux is larger than that shown in Fig. 15 and the agreement 

with the measured data is significantly improved. The new transition boiling model was 

implemented in the R5M3 code to predict some of the available reflood data, as described 

in the next section. 

9.2 Validation within RELAP5 MOD3.1-FZK 

The FZK single-rod and the NEPTUN bundle tests listed in the Table 3 of section 3.3 

were used for extensive model validation. These tests were calculated with both the new 

and the frozen code version in order to show the improved prediction capability of R5M3

FZK. The most representative results of such calculations are presented and discussed in 

the following subsections 9.2.1, 9.2.2 and 9.2.3. 

9.2.1 Discussion of Results: FZK Single-Rod Tests: Specimen filled with 

Zr02 

A.1. Quench Temperature: 1873 K 

The tests T16106, T31085, and T24066 were carried out under the same conditions. 

The Figures 20, 21, and 22 give the specimen temperature at different radial and axial

45
1 E+2

1 E+I 

E A-A'A 

:0 1E+O . 

0 x 

16E

0 0',x Johannsen (1990) 

+ 0%-, -- 0 Johannsen (1990); CHF 

a Bennett (1967) 

1 E-2 1 E-1 1 E+0 1 E+I 

Measured q" (MWlm2) 

Figure 18: Comparison with Data of a Stand-Alone Formulation of the New Transition 

Boiling Model

46
1.2

0.8

E 

4

(U 

::

0.6 to

0.4--

0.2-

I,.,

4.0 4.2 4.4 4.6 4.8 

z (M)

5.0 5.2

Figure 19: Stand-Alone New Model Prediction of Run in Bennett et al. (1967)

47

q"(tot, measured) 

0 0 0 0 0 0 0 0 0 0 

q [tot, predicted) 0 W- W.Wý 

000..00, " New Model

. q"(conv., predicted) Bennett Run 5251 

G=1356 kg/(m2 s)

q"(boiling, predicted) 

S- -l - - - - - - - - - - - - - - - - - -•- --------- - iI---------- -

7 I I

5.4 5.6

1.0 -1
positions, predicted with both the frozen and the modified R5M3 versions, in comparison 

with the measured data. The centerline temperature (picture a of Fig. 20) predicted by 

the modified model is in very good agreement with the measured one. The measured gap 

temperature (Picture b of Fig. 20) coincides with the predicted pellet surface temperature 

throughout the transient. 

The measured clad outside temperature at the middle of the specimen agrees well 

with the predicted temperature of both code versions for the tests T31085 , T24066, 

and T16106. From the Figures 21 and 22 it can be seen that the measured temperature 

oscillates very much especially during the transition boiling regime near to the quench 

front position, where an oscillatory two-phase flow prevails. This and the stochastic nature 

of the reflood flow makes difficult not only an accurate temperature measurement but 

also the repeatability of the temperature measurement. The quench time (770s) is well 

predicted by the modified code in the case of the tests T16106 and T24066. The measured 

quench time for the test T31085 is few seconds later (776). The quench temperature is in 

good agreement with the measured value (about 800K) for the three tests. 

In order to gain a better physical understanding of the complex heat and mass transfer 

phenomena taking place during quenching, Figs. 23 to 25 present calculated results for 

various parameters which cannot be readily measured. 

Heat transfer data at the mid--section of the specimen are shown in Fig. 23. The 

rewetting process is shown to coincide with changing heat transfer mode from subcooled 

film boiling (mode 47) to subcooled transition boiling (mode 45). This change is now earlier 

than by the frozen version, since the transition boiling heat flux is no more underpredicted. 

The change from subcooled transition boiling (mode 45) to subcooled nucleate boiling 

(mode 43) is also initiated earlier when the wall superheat decreases below about 50 K, 

as expected. The changes in heat transfer mode are accompanied by large variations in 

the heat flux (see Figure 23a). 

The heat flux and the heat transfer coefficient predicted by the modified code version 

is much higher than the one predicted by the frozen version as a consequence of the model 

improvements. It nearly reaches the predicted critical heat flux (see Figure 23a), which 

is in agreement with the known shape of a boiling curve. The deviation of the maximum 

heat flux from CHF is a result of the fine-mesh scheme used in R5M3 to describe the 

reflood process. The CHF value is reached in one of the fine-mesh points while values 

of the average heat flux is plotted in Fig. 23. In Fig. 24 the heat flux as a function of 

the wall superheat is represented. The maximum heat flux point occurs at AT, ; 40 

K. The minimum film boiling superheat predicted by the modified model is about 500 

K compared to about 200 K in the frozen R5M3 code. This is a result of the modified 

transition boiling model. 

Figure 25 shows the void fraction and vapor velocity at the mid-section of the spec48
W 

E 

0) 

I

0) 

C

CL 

1)I 

E 

W 

CO 

0 

= 

(.9 

E 

C) 

0. 

E 

0 

-o

I 736 752 768 784 800 816 

Time (s)

1800 

1600 

1400 

1200 

1000 

800 

600 

400 

200 7, 

1800 

1600 

1400 

1200 

1000 

800 

600 

400 

200 

72 736 752 768 

Time (s)

784 800 816

Figure 20: Measured and Predicted Temperatures in the FZK Single-Rod Test No. T16106 

49

20 736 752 768 784 800 816 

Time (s)

1800 

1600 

1400 

1200 

1000 

800 

600 

400 

Ontn

7,

S.. . .. .... ' 'I o Experiment 

Frozen Rt5M 

. .. . . . . R.M.- ~. F .. ..

0 

S. . .: .. .. . ... . . . . .T . .. . .. . . . . . . . ... . . . . . . . 

00 

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

HeIhtO.05 

ip

o Experiment 

- - Frozen R5M3 (clad inside) 

- - R5M3-FZK (cdad inside) % . .Frozen RSM3 (pellet outside) 

-'- R5M3-FZK (p~t outside) : .• ...... ..• : .. . . .... ! s....i . :....... : : 

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

.. .. Height=O.075-m . 

S. b. .......... .. .... ... .... .. .. ... ... .. 

. ... .. .. .. . . . . _

0

I.

0

Singte.1RQdT16j06.* 

" - Frozen R5M3 

(a) Height=O.075 m S. . .... . :.. . ... .. i . . . . . .... . . . ... . ... ... .. ... .

i 3

I q

I
S1600 

o S1400 

L 12000 -- Fo=e 5M 

a), 

i-- 100 i •RSM3-FZK 

100 

- 800 

0 

S6 0 0 

400 Height=0.125.m 

200 

720 728 736 744 752 760 768 776 784 792 800 808 816 

Time (s) 

1800 

0 - Frozen R5M 

1600 

1400 . 

1200 

E 

T81000 

(D 

3-: 

0s 

CO 600 

400 ()Height=O0.07.5.m 0 

200.............. 

720 728 736 744 752 760 768 776 784 792 800 808 816 

Time (s) 

Figure 21: Measured and Predicted Clad Temperatures in the FZK Single-Rod Test No. 

T31085

50
7 1600 

0) 0 Experiment 

1400- - Frozen R5M 

cc. - R5M3-FZK 

C. 1200 

E 

(D 

S1000 

_0 

• 800 0

• 600 -. 

400 (a) Height-O.075 m 

200-

720 728 736 744 752 760 768 776 784 792 800 808 816 

Time (s) 

1800 

£ 1600 t•.• - - Frozen ,R5M , 

0 

:. 1400 0 

0 CO 

0.1200 -

E 

(D 

1000 

CD 80 HeighttO.025 m " 8 0 0,, % . ... ... . . . . . . . . . . . . . 

0 

- 600 A 

o 2 0 . .. . . . . . .... ..... ... 

200 

720 728 736 744 752 760 768 776 784 792 800 808 816 

Time (s) 

Figure 22: Measured and Predicted Clad Temperatures in the FZK Single-Rod Test No. 

T24066

51
728 736 744 752 760 768 776 784 792 800 808 816 

Time (s)

0 728 736 744 752 760 768 776 

Time (s)

784 792 800 808 816

728 736 744 752 760 768 776 784 792 800 808 816

Time (s)

Figure 23: Predicted Heat Flux Parameters at the Mid-Section FZK Single-Rod Test No. 

T16106

52

4

E 

LL

5.0e+01 

4.9e+01 

4.8e+01 

4.7e+01 

4.6e+01 

4.5e+01 

4.4e+01

a 

0 

"I-

(b) Fi¶ mn R

4748 satureted film boiling, 

..... 48 subcooled film boiling 

46 saturated transition boiling 

45 -subcooled transition boiling 

44Isaturated nucleate boiling 

4 supcooled nucleate boiling 

Height=-0.075 m

4.3e+01 

4.2e+01 

7;

A 

0 

0 

0 

I

U "0-
6E+05

5E+05 

4E+05

E Cr3E+05

2E.+05

1 E+05 

Original R5M3 

OE+O0 0 i I I I I 

0 100 200 300 400 500 600 700 800 

(Tw - Ts) (K) 

Figure 24: Comparison of Current and Modified R5M3 Predicted Boiling Curves in the 

FZK Single-Rod Test No. T16106

53
imen and the collapsed liquid level in the channel. The rewetting process is shown to be 

accompanied by a sharp drop in void fraction and vapor velocity. Since the transition 

boiling heat flux in the original model is lower, these trends are not as obvious in the 

results of the frozen R5M3 code. The collapsed liquid level reflects the variations of the 

average liquid void fraction in the channel. Both void fraction and vapor velocity show an 

oscillatory behavior near the quench front due to oscillations of the heat flux and therefore 

of the evaporation rates. 

A.2. Quench Temperature: 1673 K 

Figure 26 shows the predicted and measured temperatures for run No. T02106 of 

the FZK single-rod program. This test is similar to the test No. T16106 except for its 

lower initial temperature which results in a higher quench velocity. In the test T16106 

the mid-section quenches in about 50 s while in T02106 quenching occurs after 35 s. 

The improvements in the results obtained with the modified R5M3 code over the current 

version of the code are obvious. As expected, the measured gap temperature lies between 

the predicted inner clad surface temperature and the outer pellet temperature. The test 

T21066 is similar to the tests T31085 and T24066, where the temperature was measured 

on the clad outside along the specimen. The predkcted clad outside temperatures at the 

middle and at the upper thermocouple positions agrees much better with the measured 

values than at the lower part. At the lower part the discrepancy is larger. On the other 

hand, the quench time predicted with the modified code version is comparable with the 

measured values. 

A.3. Quench Temperature: 1473 K 

In this case, the clad outside temperature and the quench time at the middle and at 

the upper part of the specimen, calculated with the modifidd code version, is much closer 

to the experimental results than in most of the other tests (see Figure 28). Merely at the 

lower part the discrepancy between measured and predicted temperature is larger, but 

the quench time is close to the measured one. Quenching of the mid-section (h = 0.075 

m) occurs in this case at t ; 755 s (35 s after reflood initiation). The quench front velocity 

is decreasing as the initial wall temperature increases.

54
o0.6-FmzenR5M3-FZK 

U- 0.60 ~F~ 

0 

> 0.40 

Height=0.075 m 

0.20 

0.00 - . 

720 728 736 744 752 760 768 776 784 792 800 808 816 

Time (s) 

1.20 • =.vFr.zen : : i 

S1.00 R5M3-FZK 

S0.80 

o 

> 0.60 Height=O.075 m 0 

C. 

> 0.40 

0.20 

0.00 

720 728 736 744 752 760 768 776 784 792 800 808 816 

Time (s) 

0.20 

0 

o2 : (C 0.16 ....

S0.12 . I------ --- ---. i; 

0.0.08 o.o4 CL) 

0.04 

0.00 , 720 728 736 744 752 760 768 776 784 792 800 808 816 

Time (s) 

Figure 25: Predicted Flow Parameters in the FZK Single-Rod Test No. T16106

55
1 600 1Single Rod T02106 

( 1400 Exprment S•w,• • - -- Frozen R5M3J 

S1200 RM-Z 

C) 

E 1000 

a 800 

S6 0 0 

C) 400 (a) Height'=0.075 m 

200 

720 736 752 768 784 800 816 

Time (s) 

1600 o Expenment - - Frozen RSM3(pellet outside) 

- - Frozen R5M3 (dad inside) 

7 1400 0o- R5M3-FZK (pellet outside) 

"- R5M3-FZK (clad inside) 

ca 1200 (.9 

.G_ 1000". . ... , ".. W . 

O 800 . 

0~0 

E 600 

400 (b) Height=O.075 m 

200 

720 736 752 768 784 800 816 

Time (s) 

1600 

1= 400 - .R_5 M3-FZ K.. . .. ... .. ...

E 1000 

S[ , . ; ....... ~i..iHeight=O; .O7 • -. .. 

1800 

00V 

0)0 

o 400 (C) 

200 

720 736 752 768 784 800 816 

Time (s) 

Figure 26: Measured and Predicted Clad and Pellet Temperatures in the FZK Single-Rod 

Test No. T02106

56
£ 

(D 

CL 

0~ 

E 

(D 

0) 

0

V 

cc 

CO, 

£ 

C) 

C) 

E 

(D 

C0 

00 

£

2UU 720 728 736 744 752 760 768 776 784 792 800 

Time (s)

1600 

1400 _0 

1200 Q 

1000 

800 

600 

400 

200 

720 728 736 744 752 760 768 776 784 792 800 

Time (s)

728 736 744 752 760 768 776 784 792 800 

Time (s)

Figure 27: Measured and Predicted Clad Temperatures in the FZK Single-Rod Test No. 

T21066

57

1600 

1400 

1200 

1000 

800 

600 

400

Single.Rod T21066. 

0 

4 o Expenimern 

-- Frozen R5M 

0 R5M3_F 

0(a) Height=O0.125 m

~~~~ .... ". .... ---. -- o Exp rient ..... : ....... 

- - Frozen R5M 

"• R5M3-FZK 

.Heig9htO-;.075 m-

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

|
to 

a) 

E 

a) 

"I

C) 

-. 

E 

(D 

*0 

a) 

"D 

CU 

E 

I

0 

*0 

CD a)

1600 

1400 

1200 

1000 

800 

600 

400 

"inn

Single Rod T24085

720 728 736 744 752 760 768 776 784 792 800 

Time (s) 

•6oO4oo : : i ::-:......: : .... :- ..:.. . . ' . ..: 

16000 

-- Frozen RSM3 

1400 

1200 

1000 

. HeightfO.075. m , 

200 

4 0 0 . . . . . .. .. . . . . .-. . .-. . . . . . ..- 

720 728 736 744 752 760 768 776 784 792 800 

Time (s) 

1600 

1400 . ... .Frozen R5Mz 

1200 

1000 

800 " . : 

800 !Height=0.025 m 

600 

400- (C) 

200 

720 728 736 744 752 760 768 776 784 792 800

Time (s)

Figure 28: Measured and Predicted Clad Temperatures in the FZK Single-Rod Test No. 

T24085

58

o Experiment 

- - Frozen R5M3 

- R5M3-FZK 

..(a) . . .. . .

Heig . . . ht=O0.125 . . . . . .m ... . 0 . . . - . -. :.. 

S. . . . . . . . . . . . . . . . . . . . i i i ] 1 1 1 1 - - . 1 1 1 1 1 1 I
9.2.2 Discussion of Results: FZK Single-Rod Tests: Empty Specimen 

The calculational results obtained with the frozen and modified R5M3 code for the FZK 

empty tube tests listed in Table 3 are given in Figs. 29 to 31. These tests differ among each 

other mainly in their initial temperatures. The predicted inside and outside temperatures 

at the middle of the specimen are almost identical and. It lies between the measured inside 

and outside temperature as shown in Figs. 29 to 31. The cladding temperature predicted 

by the modified model is generally closer to the measured values. The measured clad inside 

temperature of test 29 exceeds the outside temperature by as much as 400 K. This trend, 

which persists for a few seconds, is believed to be due to a measurement error. The inner 

thermocouple is typically mounted on a rhenium foil which is welded to the inner surface. 

It is believed that the inner thermocouple in this case was disjointed during the test, thus 

remained at higher temperature for a longer period. In contrast the measured difference 

between the inner and outer temperatures are small in the two other tests considered in 

this study (ITE22115 and ITE29115). The measured clad outside (tests 29) and clad 

inside (tests 30 and 31) temperature curves show a clear quench temperature of about 

600 - 700K. The measured quench velocity is higher compared to the values obtained 

by the tests with filled tubes. For the range of initial wall temperatures studied in this 

program (1300 to 1800 K), the measured quench times for the mid-section of the empty 

tubes vary between 10 to 18 s while it lies between 52 and 70 s for the filled tubes. Because 

of the fast transient nature of these tests, temperature measurements are typically scarce 

and less accurate. The oscillatory behavior of the clad outside temperature can be observed 

in the Figs. 29 to 31. In addition, the characteristic shape of a temperature curve with a 

definite 'knee' point is not always observed. 

In general, the quench times and quench temperatures predicted with the modified 

code version are closer to the measured data than those obtained by the frozen version. As 

for the filled tube tests the measured outside surface temperatures are consistently lower 

than the predicted surface temperatures. As mentioned before, this is mainly caused by the 

inability to accurately measure the outer surface temperature. The external thermocouples 

experience an enhanced cooling due to their exposure to the steam flow in the channel. 

The effect of the thermocouple cooling is, however, smaller at higher elevations where the 

steam temperature approaches the wall temperature.

59
724 728 732 736 

Time (s)

740 744 748

Time (s)

Figure 29: Measured and Predicted Clad Temperatures in the FZK Single-Rod Test No. 

ITE20115

60

I 1400 

1200 

1000 

800 

600 

400

£ 

CU 

CL 

E 

CU I

-o

Single Rod ITE20115 

o o Experiment (clad inside) 

0• a . Experiment (clad outside) 

a -- -- Frozen R5M3 (dad inside) 

9 0O I 1 R5M3-FZK (clad inside) 

0 0 6b 

.0.75.1 '

0

200 

721

1400 

"-' 1200 

S1000 

E 

0 800 1

600 

0 

a 400 

0
Single. Ro.d_!TýE221.15 

0 Experiment (clad inside) 

SExperiment (clad outside) 

Frozen R5M3 (clad inside) 

R5M3.-FZK (dad inside) 

a) ......

20 724 728 732 736 740 744 748 

Time (s) 

L% ..... : ~~ ~ oF....i ~ ... .. ! ............. -- .? . .. . .. .. .. ... .. .... .. . 

. ....... ~ ki3FZX 

• c•) " i .... i .. .. ! ... . ... ... .....r ze 3 3 ""'" . . !. . .. . . .!. . 

-A... . ... ..... :" .... ]' ... .! ..... • .... 

S. ~. :..... •..... f. . . ..... ..... ." ....... i...... •.. .. !..... I .. .. i...... 

ii~l i ii~iil~ .HelghI.o0751 m i i iili 

" ... ....... . .... .. .. ...... 

.... .. • ' .. i ..- .. . .....

724 728 732 736 

Time (s)

740 744 748

Figure 30: Measured and Predicted Clad Temperatures in the FZK Single-Rod Test No. 

ITE22115 

61

1800 

1600 

v 1400 

*Z 1200 

E 1000 

800 

-o g 600

I

I

400 

200 

7

1800 

"1600 

S1400 

C- E 1200 

F- 1000 

CD 

S800 

o 600 

"D0 

L 400

200 7,'20
2000 

1800 Single Rod ITE 29115 

SExpement (clad outside) 

1600 0 Experiment (clad inside) 0 - Frozen RSM3 (clad inside) 

v OT 1400 140000 R5M3-FZK (clad inside) "• a °5 

(a a1200 HeightO0.075 m 

I-- . . . . 

• "o 800 .. .. . 

0 600 (a 0. 

. 

4000 200 . 

720 724 728 732 736 740 

Time (s) 

2000 

1800 0 peri_. mentn ". I-- Frozen R5M3" ID 16 00• -% 0 . . . . . ; • 1 M 3-FZK . . . . . . . . 

E- 1600 0•0 .3 

D 0 

" 14000 

aD 0 CL 

E 1200 

"a) 1000 V0 

00 

VO 600co 

S400 Height-0.075m mo , 

200 

720 724 728 732 736 740 

Time (s) 

Figure 31: Measured and Predicted Clad Temperatures in the FZK Single-Rod Test No. 

ITE29115

62
9.2.3 Discussion of Results: NEPTUN Tests

Figures 32 and 33 present some of the results obtained for the NEPTUN low flooding 

rate run 5036 along with the measured heater temperature and void fraction. Since a 

constant electric heating is applied throughout the transient, the temperature at a given 

axial position rises initially until a turn-around temperature is reached but decreases 

eventually due to increasing convective heat removal. Results by the frozen version of 

R5M3 predict early turn-around and show no clear rewetting phenomena. The modified 

code results agree well with the measurements. Quench times are also well predicted at 

the two heights plotted in Fig. 32(b and c). As can be seen from Fig. 32(a) the reflood 

process is accompanied by intense void fraction oscillations. The void fraction drops from 

its initial value of 1 and, after reaching a minimum value of about 0.5, stabilizes at a 

level of about 0.8, which is somewhat above the measured value of 0.7. Rewetting at level 

0.746 m (at t : 120 s) occurs at relatively high void fraction. The void fraction oscillations 

result from oscillations in the heat flux and boiling rate at lower elevations. 

Figure 33(a) shows frequent variations of the heat transfer mode between transition 

and film boiling before a transition boiling regime (mode 45) is established at t ; 130 s. 

This is probably due to the large variations of void fraction experienced between t = 100 

and t = 120 s. The void fraction affects the predicted transition boiling heat flux through 

its effect on the fractional wetted area (cf. eq. (47)). The rewetting process is accompanied 

by a significant increase of the HTC at the surface, Fig. 33 (c), characteristic to nucleate 

boiling. 

Figures 34 to 37 describe in detail the NEPTUN high flooding rate test (test 5050). 

Predicted and measured temperatures are shown at three elevations along the heater rod. 

Unlike the low flooding rate test, the temperature curve in test 5050 does not rise at the 

beginning of the transient. Enhanced boiling and water entrainment at the bottom of the 

test section provide ample amount of coolant along the channel to balance the electric 

power input. The modified R5M3 calculations indicate distinct quench behavior with a 

clear 'knee' temperature. At the 0.746 m elevation, the calculated void fraction at the 

rewetting time is less than 50 % (Fig. 35), compared to about 80 % void fraction observed 

in test 5036 at the same elevation (Fig. 32(a)). The transient in test 5050 is so fast, that 

the vapor void fraction histories do not show the distinct oscillations as in the previous 

case. At lower elevations (up to 0.978 m), the new results are in good agreement with 

the experimental data. At the higher elevation (1.21 m), the predicted quench velocity is 

generally overpredicted (early quenching). Although the measured initial temperature at 

1.21 m is low, test results show a late quenching, probably due to the relatively high void 

fraction at that elevation, Fig. 37. 

The early quenching predicted by the modified R5M3 code may point out a possible 

deficiency in the prediction of the void fraction downstream of the quench front. The void

63
1.00 

0.80 

.0 

0.60 

LL 

0 0.40 

0.20 

0.00 0 

1200 

S1000 

S800 

E 

F

S6 0 0 

0 

D 400 

C) 

200 0 

1200 

( 1000 0 

S800 

E 

I

0 600 

0 

V 400 

C) 

200 

0

Time (s)

40 80 120 160 200 240 280 

Time (s)

40 80 120 160 200 240 280

Time (s) 

Figure 32: Measured and Predicted Clad Temperatures and Void Fraction in the NEPTUN 

Test 5036

64
0 

ao 

CU

5.0e+O1 

4.9e+01 

4.8e+01 

4.7e+01 

4.6e+01 

4.5e+01 

4.4e+01 

4.3e+01 

4.2e+01 

9.0e+05 

8.0e+05 

7.0e+05 

6.0e+05 

5.0e+05 

4.0e+05 

3.0e+05 

2.0e+05 

1.0e+05 

O.Oe+00 

1.8e+04 

1.5e+04 

1.2e+04 

9.0e+03 

6.0e+03 

3.0e+03 

O.Oe+00

0 40 80 120 160 200 240 280 

Time (s)

120 160 200 240 280 

Time (s)

Figure 33: Predicted Heat Transfer Parameters in the NEPTUN Test 5036

65

Time (s)

E 

x 

LL 

Co 1"

E 

OD 

0 

Co 

Ia) 

-I
fraction calculations are affected mostly by the interfacial momentum transfer and by 

the vapor generation models of the code. Therefore, further investigation of these models 

may be necessary before a clear conclusion can be drawn about the disagreement of the 

quenching rate at high elevation. In addition, refinement of the numerical scheme of the 

code is recommended in order to account for possible multiple quenching along the fuel 

rod.

66
0000 

0 

"CL CL, 

600 600 D 

00 4 46 

( ... 203 

ch 0 xpermentCD 

400 Modified R6. .. 400 { o 

L 

)(D 

(a) Height=0.921 m 

200 . . . . . ; . . . , i , 200 

0 10 20 30 40 50 60 

Time (s) 

100er•0ent o 1000 0 

S" • ./--R5M3 SD. 

0 . . Modified R5 .... 

4 00 800 

00 

r O, . . .. , .. : . .. . .. ... ... .. ..... ..... . . . . . . 

0 C 

600 0 600 0 00' 

00 

40 . .. ... . 400 CCD (b) Height=O7..978 m' 

20o 1 .0 2 .0 30 40 50 6020 

Time (s) 

1200 31200 

0) 1000 1000 . 

0 

CD 

0) -IIig 

0600 

o CD 

-0 400 400 (D 

0 10 20 30 40 50 60 

Time (s) 

Figure 34: Measured and Predicted Clad Temperatures in the NEPTUN Test 5050

67
50 

49 

S48 

'0 

47 

• 46 

a; 

t-45 

"I 44 

43

1 10 20 30 

Time (s)

E.PTUN-5050. 

"Modified R5M3.

REFLOOD TEST 

Height_0O.746 m 

(a) .. . i . .. "

"E 

C 

B C'-, 

W 

M3.6e+04 

3.2e+04 

2.8e+04 

2.4e+04 

2.0e+04 

1.6e+04 

1.2e+04 

8.0e+03 

4.0e+03 

O.Oe+00

0 10 20 30 40 50 60 

Time (s)

3.6e+04 

3.2e+04 

2.8e+04 

2.4e+04 

2.0e+04 

1.6e+04 

1.2e+04 

8.0e+03 

4.0e+03

-1' 

-.

ci 

_o. 

0

0.Oe+00

Figure 35: Predicted Heat Transfer Parameters, 0.746 m Elevation in the NEPTUN Test 

5050 

68

40. 0 6042 42 4 50 6

0 

L.) 

-o

1.00 

0.90 

0.80 

0.70 

0.60 

0.50 

0.40 

0.30 

0.20 

0.10 

0.00

1.00 

0.90 

0.80 

0.70 

0.60 

0.50 

0.40 

0.30 

0.20 

0.10 

0.00

0 

-n 

_o 

0.

SModifed R5M. 

He. .h. .7.46 

(C)...

50 

49 

48 3I 

CD 

47 

46 • 

45 

0 

44 4 C. 

I
10 20 30 

Time (s)

Time (s)

30 

Time (s)

40 50 60

Figure 36: Predicted Heat Transfer Parameters, 0.978 m Elevation in the NEPTUN Test 

5050

69

50

49 

(D 48 o 

0 

47 

M a) 46 

43

42

REFLOOD TEST NEPTUN

........ ......-..3-..K 

Heighl=0.978 m 

(a)..

0

0 

C)i 

U

LL 

_0

1.00 

0.90 

0.80 

0.70 

0.60 

0.50 

0.40 

0.30 

0.20 

0.10 

0.00

3.6e+04 

3.2e+04 

2.8e+04 

2.4e+04 

2.0e+04 

1.6e+04 

1.2e+04 

8.0e+03 

4.0e+03 

0.Oe+00

C 

C._ 

Co 

I

I

.. ...... '..
50 

49 

a) 48 "0 

0 

2 47 

"• 46 

•45 

"1-44 

43 

42

Time (s)

10 20 30 

Time (s)

40 50 60

Figure 37: Predicted Heat Transfer Parameters, 1.21 m Elevation in the NEPTUN Test 

5050

70

C 

0 

LL 

~0

1.00 

0.90 

0.80 

0.70 

0.60 

0.50 

0.40 

0.30 

0.20 

0.10 

0.00

1.8e+04 

1.6e+04 

1.4e+04 

1.2e+04 

1.0e+04 

8.0e+03 

6.0e+03 

4.0e+03 

2.0e+03 

O.Oe+O0

E 

0 

CA 

CU 

I

CO 

I

-(-

......... ~~~~... .. ...... .... .. 

..... ...- FZ 

()Height=-1.21 m,......

0

m
10 Parametric Studies

In section 2.3 the importance of the heat transfer within the fuel rod was pointed out 

considering parameters like gap resistance, gap pressure, gap filler gas, etc. during reflood 

situations. A set of calculations were carried out to study the effect of the gap model in 

R5M3 on the temperature prediction during reflood. Figure 38 shows different clad out

side temperatures predicted under different assumptions. Assuming no gap conductance 

is equivalent to assuming that there is no interaction between the pellet and the cladding 

(a gap with infinite heat resistance). Disconnecting the cladding from the pellet results 

in early quenching as shown in the clad outside and inside temperature curves in Fig. 

38. This is in agreement with experimental results [7] which indicate that decreasing the 

heat conductivity in the radial direction increases the quench front velocity. Neglecting 

the radiation heat transfer between the pellet and the inner surface of the cladding re

duces in principle the gap conductivity and hence, increases the quench velocity (Fig. 38). 

However the effect of the radiation component on the total gap conductivity is shown to 

be small. This is better understood by examining Figs. 39 and 40 which show the temper

ature difference across the gap and the temperatures of the pellet and the inner cladding 

independently. Significant temperature difference exists during rewetting, but since the 

clad and pellet temperatures at that time are relatively low, radiation heat transfer does 

not play an essential role in the total gap conductance. 

The gap conductance depends to a large extent on the gap width which is in turn a 

function of temperature and thermal expansion properties of the clad and pellet materials. 

Figure 41 is a typical plot of the gap width variations during reflooding. It refers to the 

predicted results at the mid-section of the specimen in the base case test (FZK single-rod 

test No. T16106). The gap width drops initially as the cladding is cooled. As the pellet 

cools down the gap width increases. During the rewetting time the gap width decreases 

momentarily in reaction to the sudden drop in the cladding temperature.

71
2, 

(D 

0) 

CL 

E 

CD 

a

0) 

-C 

0

1800 

1600 

1400 

1200 

1000 

600

4 0 0 ... ... . .. .. r eightFO.075 m. 

2 0 0 . . ... . . . .. . ... . . . . . 720 730 740 750 760 770 780 790 800 

Time (s) 

Figure 38: Effect of Gap Conductance on Quench Velocity in the FZK Single-Rod Test 

No. T16106

72

•~~~~. . •. . .. ... .. ... . . 

QUENCH RIG TEST T16106 

- Gap Conductance Model including Radiatio, 

-- Gap Conductance Model without Radiation 

-No Conductance Model • 1-.o• •, •,• 'w °= •' ..... ... ,... 

v, i ...• .. .. . ......... ............. 

RELAP51Mod3.1-FZK 

- I.. . ...
1800 

- 1600 

CC. 1400 

1200 

S1000 

0 S800 

® 600 0. 

E 

0) 400 I

200 

0 

721 750 760 770 780 790 800 

Time (s)

- RELAP5d.

a a 

I .... .. 1... 

Height=0..075 m

III 0 730 740 750 760 770 780 790 800 

Time (s)

Figure 39: Gap Temperature Curves in the FZK Single-Rod Test No. T16106 

73

I 

I

F

QUNECH RIG TEST T161.06 

-RELAP5/Mod.-Z (eltu 

.......... 

i- - RELAP5IMod3.1-FZK OClad~inI 

Heightz;0.0.75.m.:. 

(a. ). . : . .

0 730 740

600 

500 

"- 400 0. 

(D30 

o~0 

0, 

I- 200 

100 

0 

72

S. . . . . . . . . . . . . i i i i i i i i t

3.1-F'ZKiI
44 

'SinOgleRod TI 6106 4 0 ...-... ..... .. . ..... 

• ..--..- - .. [ Height = OA 5 m .. . . .........- ....... : ' ' i 

6 -Height= 0.75 m 36 ..... HeihtO1 

. 0.125 m I ...... •i- -: ...-..--... i • : -• ........ .- .. -....-.. 

"• ~ ~ ~ . ~ ...... ~ ~ ~ ~ .. .8.•.!.i...-.:-: .. 32 

~28 

0

a s '" ' ' : " : "" " . ... . . . '... . . . . .. . .. 

0 2 

2 .. .. ............. . ... . .......- ... ............... ........... 20 -. :... . •..: :.•.:. •-: .. ..... : :...:..-.. ... ..• .:.......------.. .- •-. 

> 

04 - . . .. . . . .. ; : " " ; : " * ' ' . . . 

F- 16 • ........

12 . . 

o .!... : .......--. ! .... N., 

4 . ." .... ' .... 

720 728 736 744 752 760 768 776 784 79, 800 808 816 

Time (s) 

Figure 40: Temperature Difference Across the Gap at Different Heights in the FZK Single

Rod Test No. T16106 

11 Run Statistics and Code Performance 

The calculations presented in this report were performed on both IBM RS/6000 and 

parallel IBM RS/6000 SP machines. The discussion is based on the results obtained on 

the IBM RS/6000 machine. 

A. QUENCH Rig 

The model used for the QUENCH rig consisted of 19 volumes (15 volumes in the test 

section, two time dependent volumes, a lower and an upper plenum), 4 junctions and 30 

heat structures (see Figure 8). To demonstrate the run statistics and code performance 

of the modified RELAP5/MOD3.1-FZK in comparison with the frozen version, the test 

T16106 was chosen. Figure 42 gives the courant time, the CPU time as well as the mass 

error calculated by both code versions. One can see that both CPU and courant time are 

similar. 

B. NEPTUN Facility 

In the case of the NEPTUN tests the facility was modeled by 20 volumes (18 in the test 

section, a lower and an upper time dependent volume), 2 junctions and 16 heat structures

74

S. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ; q • • " I v W . . . .
150 

130 

110 

E Height =0.075 m 

:. 

C, 

90 

70 

50 i.i I I . 

720 745 770 795 820 845 870 

Time (s) 

Figure 41: Gap Width Behavior During Reflooding in the FZK Single-Rod Test No. 

T16106

75
E 

:3 

-C 

C 

0 

EC 

F

a) 

E 

2 

I

C, 

C 

1=.

).800 

).700 

).600 

3.500 

3.400 

0.300 

0.200 

).100 

).000 

.7: 

600 

500 

400 

300 

200 

100

0.025 

0.020 

0.015 

0.010 

0.005 

0.000 -

736 752 768 

Time (s)

784 800 816

Time (s)

720 736 752 768 784 800 816 

Time (s) 

Figure 42: Code Performance: Rod Test No. T16106

76

Single Rod T16106 

F R5M3-FZK I 

0)a 

/ (I

.0
(see Figure 9). Figure 43 shows the time step, CPU-time and mass error calculated 

by both code versions. The time step and mass error predicted by the modified code 

exhibits much less oscillations than that of the frozen version (see Figure 43 a, c). The 

CPU consumption of the modified version is smaller by about 10 %. Figure 44 depicts 

the vapor temperature, void fraction and courant time calculated by both code versions. 

The predictions of the modified code version are in general more stable and oscillate less 

than that of the frozen code version.

77
8.0e-02 Ktz#iLUUU I I Nzb F)IU UN -- 3.3O 

6.0e-02 

o".- Frozen R5M 

/--- RSM3-FZKJ uf 4.0e-02 

CO 

2.0e-02 , 

'(a) 

O.Oe+00 0 40 80 120 160 200 240 280 

Time (s) 

1.2e+03 i : /" •. 

1.0e+03 

-~8.0e+02 

D .... .. _. ..

E 

R 6.0e+02 Frozen R5M3 

O 4.0e+02 

2.0e+02 

0 40 80 120 160 200 240 280 

Time (s) 

3.0e-02 -- Froze R5M,• : --- ~R5M3--FZ J 

"( 2.0e-02 

E 1.Oe-02 

IO.Oe+00 

0 40 80 120 160 200 240 280 

Time (s) 

Figure 43: Code Performance: NEPTUN Test 5036

78
£ 

0) 

() 

0ý 

E 

0) I

0 0U 

E 

0 0 

T 

0 

LL 

0 

U 

(D 

"a, 

0) 

E 

CO 

0 

0 

ITime (s)

79

8.0e+02

IIL REF-LOOD .TEST-N.EPTU-N-ý536

• " " " Frozen RSM3 " 

R5M3-FZKJ 

(a) HeightO0.862 m

6.0e+02 

4.0e+02 

2.0e+02 

0.Oe+00 

0 40 80 120 160 200 240 280 

Time (s)

1.0e+00 

8.0e-01 

6.0e-01 

4.0e-01 

2.0e-01 

0.Oe+00 

3.0e-02 

2.0e-02 

1.Oe-02 

O.Oe+00

Time (s) 

Figure 44: Code Performance: NEPTUN Test 5036
12 General Conclusions and Recommendations 

The heat transfer model of the RELAP5 MOD3.1 code was extensively reviewed and as

sessed. Several reflooding experiments with varying pressure, flooding rate, heating power 

and initial wall temperature were simulated in order to evaluate the general prediction 

capability of the code in modeling boil-off and reflood transients. The most important 

deficiency of the current version of the code was found to result from its treatment of 

the transition boiling heat transfer regime. In this regime the code significantly under

predicts the heat transfer rate to the liquid. Since at low quality conditions, the liquid 

boiling component constitutes a major fraction of the total heat transfer, the current 

model underpredicts the quench temperature and the quenching rate under most condi

tions relevant to LOCA and degraded core analysis. 

In order to further evaluate the transition boiling procedure of the current R5M3 

code, a stand-alone model was formulated and tested against the available data-base on 

steady-state post dryout heat transfer. Analysis of over 360 data points has shown that the 

R5M3 model consistently underpredicts the measured heat flux in the transition boiling 

regime. A new model based on an extension of the original Chen model has, therefore, 

been formulated. The model utilizes only local state variables generally calculated by the 

R5M3 code and does not require other history parameters such as quench position or CHF 

and minimum film boiling temperatures, which are not readily available at each time step. 

Stand-alone validation of the new model has shown acceptable agreement with the data 

(standard deviation of about 40%). The new model was, therefore, implemented in the 

R5M3 code and used to simulate a large sample of the FZK Single-Rod and the NEPTUN 

rod bundle reflood tests. Implementation of the new model in R5M3 is straight-forward 

and requires modifications in one subroutine only. The model was also implemented in 

the SCDAP RELAP5 code and tested against some of the CORA data. The following 

conclusions can be drawn: 

1. The new model yields good predictions of reflood and general heat transfer problems 

using a single logic of heat transfer selection. 

2. High and low flooding rate experiments were predicted relatively well with the mod

ified code. The successful prediction of the high flooding rate tests is particularly 

important for the analysis of LOCA and degraded core accidents where high flooding 

rates are to be expected. 

3. The unphysical behavior of the current R5M3 code such as continuous cooling of 

the node without clear turn-around temperature and no quenching phenomena, is 

eliminated in the modified version of the code.

80
4. At the NEPTUN high flooding rate test, both the modified and current versions of 

the code predict early quenching at the upper section of channel. This may be a 

result of inaccurate modeling of the void fraction downstream of the quench front. 

5. The parametric study has shown that an accurate modeling of the gap conductance 

is important for the analysis of fuel rod or heater rod reflooding. The gap model of 

the code was found to be adequate. 

6. The modifications introduced in R5M3 does not disturb the code performance. 

On the contrary, the modified code runs faster and more stable with consider

able less oscillations of key parameters like void fraction, gas temperature, mass 

error, and courant-time. Typical CPU-times of the IBM RISC 6000 to run the test 

T16016 and NEPTUN5036 are about 400 s and 1000 s, respectively. 

Although the suggested model has been validated as a stand-alone program and as 

a part of the modified R5M3 code, a wider assessment with a large number of different 

transients is recommended before one can make a definite statement about the capability 

of the modified code to correctly predict the physical phenomena in the complete range 

of expected operational and accident conditions. In particular, the model performance at 

high surface temperatures and oscillating flow conditions far downstream of the quench 

front, should be studied. For instance, it has been shown [3] that in some of the NEPTUN 

tests, better agreement between the predicted results and the measured temperature pro

files downstream of the quench front can be achieved by limiting the proposed transition 

boiling model to the range of 0 to 0.2 m from the quench front. Further refinements of 

the model are, therefore, recommended to improve its performance at high temperatures 

and high void fractions, characteristic to top-quenching and spray cooling. 

Several modifications have been introduced in the RELAP5 MOD3.2 version of the 

code. These are mainly related to heat transfer and to the interface closure laws of the 

code. It is recommended to implement the present transition boiling model in the latest 

code version R5M3.2 in order to further evaluate its prediction capability for reflood and 

general heat transfer problems. 

13 Ongoing Activities 

To further extend validation of the improved transition boiling model implemented in 

the RELAP5 MOD3 code (R5M3-FZK) two additional reflood tests are being analysed, 

namely the bundle test FLECHT-SESEAT 31701 [30] and the integral test LOFT LP-LB

1 [37]. They represent more reactor-specific thermal hydraulic(pressure, flooding rate),

81
geometry (core height, large number of fuel rods) and system (primary and secondary 

coolant systems) conditions. 

Since the reflood model is still deactivated in all distributed versions of the code 

RELAP5 MOD3.2, this work is being done in the code version RELAP5 MOD3.1. Further 

model improvements makes sense only within the RELAP5 MOD3.2 frame. 

Preliminary results of the present validation activities are shown in figures 45 and 46 

comparing the measured clad temperature at different axial positions with the predictions 

of both code versions, the frozen (R5M3) and the modified (R5M3-FZK). 

In the case of the high flooding rate (15 cm sec), full-length PWR-bundle FLECHT the 

predicted clad temperatures with R5M3-FZK show a clear knee-temperature marking the 

begin of rewetting and a better agreement with the measured ones than those predicted 

with R5M3. 

Fig. 46 exhibits the measured clad temperature of different thermocouples radially 

positioned in the hot channel (central square bundle) at three core elevations. It can 

be seen that not all thermocouples undergo the same behavior. For better comparison of 

measured and predicted temperatures the procedure used in [33] was followed to determine 

a reference temperature (T-average) by averaging the most representative thermocouples 

values. Generally, quench temperature and quench time predicted with R5M3-FZK are in 

better agreement with the experimental data than the R5M3-predictions. Especially the 

prediction of the quench temperature is much more improved. Finally, it has to be noted 

that there is still an overprediction of the heat transfer coefficient in the dispersed film 

boiling region, which we did not touched. In-deep assessment of the quench-relevant heat 

transfer mechanisms in the film boiling mode is foreseen. 

Acknowledgements 

One of the authors (E. Elias) was privileged to spend six months (from Sept. 1996 to 

Feb. 1997) on Sabbatical leave at the FZK, Germany. This work was performed within 

the framework of the FZK-CAMP agreement. We would like to thank S. N. Aksan and 

D. Lilbbesmeyer from PSI for making available to us his NEPTUN , FLECHT-SEASET 

and LOFT LP-LB-1 input decks.

82
12 0 0 . . . . . . . . . . . . . . . . 

"" 800 \ 

E ,,D 60 0 R M. .. . -', .! .: .. .... -.. ... 

(0 400 . .. . _ •.. . . . .1 : 

(a): Height=2.47 m 

200 6" '; IV 160 

0 20 40 60 80 100 120 140 160 

Time (s) 

1200 .. 

•I5M3-FZKI 1000 "7 : - - • • ,J Expefiment/ . .;.. ..;.... 

8 0 0 ' , ... . ... ... ... ... : ...i.. 

S600 

800 600 !: ) • Height--:.92 m 

(b) 

2 . 0 . 0 . . . . . . . . . .. . . . .. ... . . . . . . . 

0 20 40 60 80 10 10 14 160 

Time (s) 

120 . ... 

.... 1000 :; /• RSM3-FZKI 

1000 [-" Expedment! 

" 800 

E 

a 600 

F---oHeightl,ý-.19 m 

ca 

400 

(C) 

200 

0 20 40 60 80 100 120 140 160 

Time (s) 

Figure 45: Measured and Predicted Cladding Temperatures of the FLECHT-SEASET 

Test 31701

83
- rE-OFG"-24 

1200CC-GI2 

[3-EJRGUU-FZK 

lac 

S.........:14e i -• t= O O .D .n ... ....... . ... ... i....... ...i .. .... . ........... .... .. . 

S a . ... . .... ... . 

;''1*•...1"..........-- - '- l -T- 1 -• 'l"- ' 1 '• - 1E 6 I -} 

-e 1 S a o ...... ...... -- . . . ..--. ..-... ..- .. . . .i.... . . .. . . ".... - ' m 

S' 03 ............ . . . .. . . . . . . .. . . . . .. .. . . . 

V 

".... .. ... ... .. .. . . ... . ....... . .. . 

a 0 

0 20 40 60 so 100 

Time (a) 

-- TE-rlrGB4. i 

1 2 0 ...... .. ...-.. ..... -. 

"")it - -- TE G IS- 2? 

S: ! -- TIE-6 L?-2 1 

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

I

.4 

2 C " . ... ....... ...... ..... .. . 

E H~l~htll.Tm/4 ; m i i .............. 

0 20 40 60 go lCC 

Time (a) 

Fi•e4:Maue -r C -e-r-ed d----e-e-1- -- s the-r -- ,-F-[•-•-1rr - h-EU

8-r w4-3m 

2E- 10CC2 

TrU.ShI?-2? 

go E6C72 

GL 

400 

. K......! gJht--@.I B74 in 

2C.. ... . ... . ......... J....J .4..... 

C 20 40 Go BC IC0 

Timea (a) 

Fiur 4: esuedan redctd lddngTepratresofth - ru-seam-nTes 

12C............................8464I
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89
NRC FORM 335 U.& NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER (2-N) NRCM 1102, (Assigned by KRC, Add VoL, Supp., Rev., 3201,3202 BIBUOGRAPHIC DATA SHEET ndddeum Numbers,. Ifany.) 

(See insbz~iot•o on the revorse) 

2. TITLE AND SUBTITLE NUREG/tA-0185 

Development and Validation of a Transition Boiling Model for the RELAP5/MOD3 Reflood Simulation 3. DATE REPORT PUBLISHED 

MONTH YEAR 

June 2000 

4. FIN OR GRANT NUMBER 

5. AUTHOR(S) 6. TYPE OF REPORT 

V.H. Sanchez-Espinoza, E. Elias*, Ch. Homann, W. Hering, D. Struwe Technical 

7. PERIOO COVERED prnusne Dan) 

8. PE•FORMING ORGANIZATION - NAME AND ADORESS gNRC, povxde Divi,• Oce a, Region, U.S. Nuc/weRgularIb4Cnnsss nehg emdsad=s. if conhc•, provide nem aid mg eddiws.) 

Forschungszentrum Karlsrube 

Hermann-von-Helmholtz Platzl 

D-76344 Eggenstein-Leopoldshafen 

Germany

is~I~s~J¶J'4UIA~/Ul~ I II -NAMEI AND~ ADDJRESS (WN~RC, 4W~ --Sane as ebo-'e ifcon#actor prvide NARC Dh'sin a~lic or Regson. U.S. Nuclea Regftisby Cornwniso, an mwAng aftesa)

Division of Systems Analysis and Regulatory Effectiveness 

Office of Nuclear Regulatory Research 

U.S. Nuclear Regulatory Commission 

Washinaton. DC 20555-0001

ID. SUPPLFLUNTARV I.1OTF

11.. ABSTRACT (2oo w-da oratos) 

The heat transfer model of the RELAP5/MOD3.1 (R5M3) code was extensively reviewed and assessed. The most important deficiency of the current version of the code was attributed to its treatment of the transition boiling heat transfer regime. The current transition boiling model significantly under predicts the heat transfer rate to the liquid phase. Since at low quality conditions the liquid boiling part is a major fraction of the total heat transfer, the current model under predicts the quench temperature and the quenching rate under most conditions relevant to LOCA and degraded core analysis. Therefore, a new model has been developed and implemented in the R5M3 code for predicting the transition boiling heat transfer. The new model is based on an extension of the phenomenological formulation suggested originally by Chen. It utilizes only local state variables calculated by the R5M3 code and does not require other history parameters, such as quench position or CHF and minimum film boiling temperatures, which are not available at actual time step. 

12. KEY WORDS'DESCRIPTORS (Lis words ar shra• toh wi/awst eseen in ocatg the report) 13 AVAILASUL7Y STATEMENT 

RELAP5/MOD3 unlimited Reflood 14. SECURITY CLASSIFICATIOO Transition Boiling 0193 page) 

unclassified 

(rft RV"x 

unclassified 

15. NUMBER OF PAGES 

16. PRICE

NRC FORM 335 I,

S ... .... . ......... ... .ES
Federal Recycling Program
NUREG/IA-0185 DEVELOPMENT AND VALIDATION OF A TRANSITION BOILING MODEL JUNE 2000 

FOR THE RELAP5/MOD3 REFLOOD SIMULATION 

UNITED STATES 

NUCLEAR REGULATORY COMMISSION 

WASHINGTON, DC. 20555-0001

years