NUREG/IA-0209

NSI RRC KI 3067

IRSN-2002/33

International

Agreement Report

Adaptation of

USNRC's FRAPTRAN and

IRSN's SCANAIR Transient Codes

and Updating of MATPRO Package

for Modeling of LOCA and

RIA Validation Cases with

Zr-1%Nb (VVER type) Cladding

Prepared by

A. Shestopalov, K Lioutov, L. Yegorova

Nuclear Safety Institute of the Russian Research Centre "Kurchatov Institute"

Kurchatov Square 1, Moscow 123182, Russian Federation

April 2003

Prepared for

U.S. Nuclear Regulatory Commission, Institute for Radioprotection and Nuclear Safety (France),

and Ministry of Science and Technologies of Russian Federation

Published by

U.S. Nuclear Regulatory Commission
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NUREG/IA-0209

NSI RRC KI 3067

-s IRSN-2002/33

International

Agreement Report

Adaptation of

USNRC's FRAPTRAN and

IRSN's SCANAIR Transient Codes

and Updating of MATPRO Package

for Modeling of LOCA and

RIA Validation Cases with

Zr-1%Nb (VVER type) Cladding

Prepared by

A. Shestopalov, K Lioutov, L Yegorova

Nuclear Safety Institute of the Russian Research Centre "Kurchatov Institute"

Kurchatov Square 1, Moscow 123182, Russian Federation

April 2003

Prepared for

U.S. Nuclear Regulatory Commission, Institute for Radioprotection and Nuclear Safety (France),

and Ministry of Science and Technologies of Russian Federation

Published by

U.S. Nuclear Regulatory Comnission
ABSTRACT

The report presents the results of analytical work on updating and supplementing Zr-1%Nb (ElIO) cladding

material property models as part of MATPRO package intended for joint use with thermal mechanical codes

analyzing high burnup fuel transient behavior. The scope of work also included adapting U.S.NRC's

FRAPTRAN code to the behavior analysis of fuel with El 10 cladding (VVER type) and carrying out of the

code assessment using selected experimental data. Satisfactory compliance of the results calculated by

modified FRAPTRAN version with in-pile RIA and out-of-pile LOCA simulated test results is obtained. The

obtained calculation data have been also compared with the predictions of the FRAP-T6 (U.S.NRC) and

SCANAIR (IRSN, France) codes modified earlier.

i i i
TABLE OF CONTENTS

Page

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

2. ZR-1 %NB CLADDING DATA BASE IMPLEMENTED IN MATPRO ............................................ 2

2.1. Zr-I %Nb cladding thermal properties ................................. 2

2.2. Zr-I %Nb cladding mechanical properties . . . 4

2.3. High temperature oxidation kinetics . . .8

3. FRAPTRAN MODIFICATIONS FOR VVER FUEL ROD SIMULATION . ...... 10

3.1. Volumetric heat generation rate (HGR) calculation for fuel pellet with central hole .10

3.2. Radial power profile calculation for fuel pellet with central hole . . .10

3.3. Calculation of the radial displacement of fuel due to thermal expansion for pellet with central hole I I

3.4. Alternative Post-Critical Heat Flux heat transfer model .12

3.5. Cladding rewetting model for IGRIRIA tests .12

3.6. Zr-I %Nb high temperature creep model for MATPRO and SCANAIR .13

3.7. Description of modified routines and coding aspects .14

3.7.1. Description of the new global variables .14

3.7.2. Description of the new and modified subroutines. 15

3.7.3. Description of the new input variables .18

3.7.4. Description of additional output information .19

4. CODE ASSESSMENT AGAINST IN-PILE IGR/RIA AND OUT-OF-PILE RIAR/LOCA TEST DATA .... ....... 20

4.1. IGR/IRIA assessment .21

4.1.1. 5OF-13 IGR assessment case .21

4.1.2. 5OF-16 IGR assessment case .22

4.1.3. 96F-09 IGR assessment case .23

4.1.4. High-burnup fuel simulation of IGRIRIA test. 25

4.1.5. Calculated results of high-burnup fuel simulation obtained with modified FRAP-T6, SCANAIR

and FRAPTRAN codes .26

4.1.6. Discussion of principal thermal parameters calculated by the codes .38

4.1.7. Summary of the assessment of the FRAPTRAN (modified) against RIA-simulated test data .42

4.2. RIAR/LOCA assessment .43

4.2.1. Description of experiments modeling thefirst stage of LOCA .43

4.2.2. Calculation procedure and main results .45

4.2.3. Summary of the assessment of the FRAPTRAN (modified) against LOCA-simulated test data.. 46

5. CONCLUSIONS .......... ...... 48

v
LIST OF FIGURES

Page

Fig. 1. Volumetric fractions of a and A-phases depending on Zr-I%Nb alloy temperature [18] ................ 8

Fig. 2. Comparison of test data on Zr-1%Nb steam oxidation with Cathcart-Pawel model predictions ...... 9

Fig. 3. Block diagram for calculating of cladding mechanical response using modified model in

FRAP}RAN code (modified) ................... . 14

Fig. 4. Clad outer temperature in 50F-13 test: comparison between measured and calculated with

FRAP-T6 (modified), SCANAIR (modified) and FRAPTRAN (modified) codes ........................ 22

Fig. 5. Clad outer temperature in 50F-16 test: comparison between measured and calculated with

original FRAPTRAN code ............................................................................ 22

Fig. 6. Clad outer temperature in 50F-16 test: comparison between measured and calculated with

FRAP-T6 (modified), SCANAIR (modified), and FRAPTRAN (modified) codes ....................... 23

Fig. 7. Clad outer temperature in 96F-09 test calculated with Bromley-Pomerantz post-CHF heat

transfer ........................................................................... 24

Fig. 8. Clad outer temperature in 96F-09 test calculated with Labuntzov post-CHF heat transfer .............. 24:

Fig. 9. Comparison of internal rod pressure in 96F-09 test, calculated with FRAP-T6 (modified) and

FRAPTRAN (modified) codes ............................................................................ 25

Fig. 10. Energy characteristics vs. time calculated by modified FRAP-T6, SCANAIR and FRAPTRAN

codes (H5T test case) ........................................................................... 27

Fig. 1. Cladding temperatures and heat transfer coefficients vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (H5T test case) ....................................................................... 28

Fig. 12. Cladding hoop stresses and itemal pressures vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (H5T test case) ....................................................................... 29

Fig. 13. Cladding hoop strains and gap widths vs. time calculated by modified FRAP-T6, SCANAIR

and FRAPTRAN codes (H5T test case) ........................................................................... 30

Fig. 14. Energy characteristics vs. time calculated by modified FRAP-T6, SCANAIR and FRAPTRAN

codes (H7T test case) ........................................................................... 31

Fig. 15. Cladding temperatures and heat transfer coefficients vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (H7T test case) ....................................................................... 32

Fig. 16. Cladding hoop stresses and internal pressures vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (H7T test case) ....................................................................... 33

Fig. 17. Cladding hoop strains and gap widths vs. time calculated by modified FRAP-T6, SCANAIR

and FRAPTRAN codes (H7T test case) ........................................................................... 34

Fig. 18. Energy characteristics vs. time calculated by modified FRAP-T6, SCANAIR and FRAPTRAN

codes (HIT test case) ........................................................................... 35

Fig. 19. Cladding temperatures and heat transfer coefficients vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (HIT test case) ....................................................................... 36

Fig. 20. Cladding hoop stresses and internal pressures vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (HIT test case) ....................................................................... 37

Fig. 21. Cladding hoop strains and gap widths vs. time calculated by modified FRAP-T6, SCANAIR

and FRAPTRAN codes (HIT test case) ........................................................................... 38

Fig. 22. Comparison of gap thermal conductivity models from FRAP-T6 (modified) and FRAPTRAN

(modified) and FRAPTRAN (original) ............................................................................ 40

Fig. 23. Peak enthalpy calculated by FRAP-T6 (modified) and FRAPTRAN (modified) ........................... 41

Fig. 24. Peak enthalpy calculated by FRAP-T6 (modified) and FRAPTRAN (modified) with gap

thermal conductivity model from FRAP-T6 (modified) ................................................................. 41

v i i
Fig. 25. Schematic diagram of electrically heated facility for studying the first stage of LOCA ................ 44

Fig. 26. Simulator cladding temperature and cladding pressure drop .......................................................... 44

Fig. 27. Cladding appearance and cross-section in burst area .................................................................. 45

Fig. 28. Comparison of calculated and experimental values of failure pressure and temperature ............... 46

Fig. 29. Comparison of calculated and experimental values of cladding residual strain .............................. 46

vi i i
LIST OF TABLES

Page

Table 1. Zr-1%Nb cladding thermal properties implemented in MATPRO package ................................... 2

Table 2. Specific heat vs. temperature under slow heat up rate [13] ............................................................3

Table 3. Specific heat vs. temperature under fast heat up rate [13] .............................................................. 4

Table 4. Mechanical properties of Zr-1%Nb cladding ........................................................................... 5

Table 5. Parameters of plastic deformation equation versus temperature and cladding type .......................6

Table 6. Anisotropy coefficients F, G, H versus temperature and cladding type ......................................... 7

Table 7. Conservative oxidation kinetics for Zr-1%Nb cladding ................................................................. 9

Table 8. New global variables incorporated into FRAPTRAN code ............................................................ 15

Table 9. List of modified routines of FRAPTRAN code .......................................................................... 15

Table 10. List of new input data parameters .......................................................................... 18

Table 11. Output information files .......................................................................... 1 9

Table 12. Additional parameters in output data files . .......................................................................... 9

Table 13. Assessment matrix of fuel rods tested in IGR reactor under RIA conditions and RIAR out-ofpile LOCA tests .......................................................................... 20

Table 14. Brief description of fuel rod design parameters and IGR test conditions ..................................... 21

Table 15. Major input data for high-bumup fuel simulation needed for FRAPTRAN code ........................ 26

Table 16. Comparison of fuel-cladding gap thermal conductivity models in FRAP-T6 and FRAPTRAN

codes ........................................................................... 39

Table 17. Comparison of main thermal and mechanical results predicted by the modified FRAP-T6,

SCANAIR and FRAPTRAN codes [3] ......................................................................... 42

Table 18. Main parameters of simulator and initial data for LOCA calculations ......................................... 45

Table 19. Comparison of experimental and calculated data on high-temperature cladding deformation

under conditions modeling the first stage of LOCA ....................................................................... 47

ix
FOREWORD

A world-wide trend to substantially increase nuclear fuel burnup to higher levels has led a

number of countries, including the United States, to evaluate the effects of higher burnup on

fuel behavior. Reactivity-initiated accident experiments, performed in France and Japan,

have shown that fuel damage under these conditions may occur well below the threshold

criteria used by various regulatory bodies, including the U.S. Nuclear Regulatory

Commission (NRC). Questions have also been raised about the adequacy at high burnup of

other fuel damage criteria used in safety analyses.

Consistent with the NRC's mission strategy to evaluate and resolve safety issues, the Office

of Nuclear Regulatory Research is conducting a thorough investigation of high-burnup

effects on fuel behavior. As part of this investigation, we recognized the value of a reactivityinitiated accident test program conducted by the Russian Research Center (Kurchatov

Institute). In cooperation with the French Institute for Radioprotection and Nuclear Safety

(IRSN) and the Russian nuclear industry, NRC now sponsors high-burnup fuel behavior

work at the Kurchatov Institute. This includes in-reactor experiments, measurements of

mechanical properties of irradiated cladding, and modification to the IRSN and NRC fuel

behavior codes used to analyze fuel response to accident conditions.

The NRC participates in several experimental and analytical programs in order to gain a

more complete understanding of highly irradiated fuel behavior under accident conditions.

Among these programs, the work conducted at the Kurchatov Institute is significant.

The ultimate goal of these activities is the development of new regulatory-criteria for

high-burnup fuel under design-basis accident conditions. However, the work has become

even more relevant to safety considerations, in both France and the U.S., due to the

introduction of the niobium-bearing zirconium alloys, which are similar to the alloys currently

used in the Russian program. A portion of the Russian analytical work is described in the

following report

Farouk Eltawila, Director

Division of Systems Analysis and Regulatory Effectiveness

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

xi
LIST OF ACRONYMS

BIGR

CHF

FGR

FRAPCON-3

FRAP-T6

FRAI T RAN

HGR

HTC

IGR

IRSN

LOCA

LWR

MATPRO

NSI RRC KI

NV NPP

PNNL

PWR

RIA

RIAR

SCANAIR

TRIFOB

NRC

VVER

Impulse Graphite Reactor (Russia)

Critical Heat Flux

Fission Gas Release

NRC's steady-state fuel rod behavior code

NRC's transient fuel rod behavior code

NRC's transient fuel rod behavior code

Heat Generation Rate

Heat Transfer Coefficient

Impulse Graphite Reactor (Kazakhstan)

Institute for Radioprotection and Nuclear Safety (France)

Loss-of-Coolant Accident

Light Water Reactor

Library of material properties for LWR accident analysis, version 11, revision 1 (USA)

Nuclear Safety Institute of Russian Research Centre "Kurchatov Institute" (Russia)

NovoVoronezh nuclear power plant

Pacific Northwest National Laboratory

Pressurized Water Reactor

Reactivity Initiated Accident

State Research Centre "Research Institute of Atomic Reactors" (Russia)

French IRSN's transient fuel rod behavior code

Code for calculation of isotopic composition (Russia)

U.S. Nuclear Regulatory Commission

Russian type of pressurized water reactor

xiii
1. INTRODUCTION

-At present, activities related to developing and updating thermal mechanical computer codes for LWR fuel

behavior modeling in the broad range of possible emergency modes and achieved bumup values are being

intensively carried out. Steady tendency toward increasing burnup in commercial reactor sets forth

corresponding requirements for quality of the fuel rod behavior codes intended for high bumup fuel analysis

in design-basis accidents. This issue took on additional actuality in connection with consideration of

alternative types of fuel claddings with high performance up to high burnup values.

Claddings of zirconium-niobium El 10 alloy that are used in Russian pressurized water reactors of VVER

type show high resistance to oxidation and hydriding during base irradiation when bumup of 50 MWd/kgU

and higher are reached [1, 2]. Along with low corrosion, highly irradiated El10 claddings show high

residual ductility, which is verified by the pulse experiment results at IGR and BIGR reactors and by the

mechanical test data [3, 4, 5]. Thus, claddings of such an'alloy type'are of interest as candidates for fuel with

high burnup 'limit. As a result, the main task of the presented work was to expand the domains of

applicability of the specified codes, which were Zircaloy cladding-oriented initially, toward alternative fuel

analysis.

Work with fuel rod behavior codes conducted at NSI RRC KI since 1996 was focused upon implementation

of material property models for Zr-1%Nb cladding in NRC's FRAP-T6 and French IRSN's SCANAIR

transient codes and upon adaptation of the codes for calculating validation cases. The latter include in-pile

and out-of-pile experiments with fuel rods of VVER type.

Providing the codes with the Zr-1%Nb cladding material property package'was the first important stage of

work. Earlier, in the course of adapting FRAP-T6 and SCANAIR codes, a package of the main thermal and

mechanical properties of the E-110 claddings was developed in MATPRO format [6, 7]. It needs to be

emphasized that a special experimental program was initiated jointly by NSI RRC KI and RIAR to measure

mechanical properties of unirradiated and irradiated claddings in the broad range of temperatures and strain

rates [3, 8]. Later on the results of the new mechanical tests allowed to update the mechanical properties

database [5, 9, 10]. Obtaining of the new test data coincided timewise with the issue of the first version of

FRAPTRAN code developed by PNNL per request of NRC [11]. Therefore the updated correlations for

cladding plastic deformation presented in the present report were implemented in MATPRO package to'be

used already with'FRAlPTRAN and SCANAIR codes.

During the next stage, calculations of the selected RIA and LOCA validation cases with fuel rods of VVER

type were conducted. The main objective of these calculations was to'analyze accuracy of predictions by

FRAPTRAN code of the most important thermal and mechanical parameters of the fuel with Zr-1%Nb

cladding, as well as to check performance of the mechanical model with the new characteristics in different

cladding loading modes. In order to model the validation cases correctly, modifications of the. original

FRAPTRAN version were needed aside from the updating of the material property package. Mainly, the

modifications were stipulated for by the design of the fuel rods of VVER type (central-hole in the fuel

pellet) and by specificities of clad-to-coolant heat transfer in pulse tests. In addition to that, a number of

coding problems were resolved including adding of new global and local variables and correcting a few bugs

and inaccuracies in the as-received version of FRAPTRAN code.

Accumulated experience in modeling fuel rods with Zr-1%Nb claddings using FRAP-T6 and SCANAIR

codes, as well as the analysis of the first FRAPTRAN validation calculations allowed us to draw some

generalized conclusions on quality of predictions by the code of behavior of a fuel rod with alternative

cladding. Based on calculation results for high burnup fuel under power pulse conditions, proposals on

future development of FRAPTRAN code were also presented in the report. They have to do with correct

accounting of high bumup effects when analyzing fuel behavior in RIA.

1
2. ZR-1 %NB CLADDING DATA BASE IMPLEMENTED IN MATPRO

Correlations of Zr-1%Nb (El10) cladding material properties, which were incorporated in MATPRO

package, can be attributed to three main blocks:

* basic thermal properties;

* mechanical properties;

* high temperature oxidation kinetics.

Formally, to model fuel rods of VVER type in the most correct manner, large number of properties of

cladding and fuel pellets should be reviewed and implemented in MATPRO as an alternative to the

properties of western Zircaloy cladding and ceramic fuel. However, analysis of full range of accountable

properties (including sensitivity study) conducted earlier [3] allowed to limit the number of key correlations

for Zr-1%Nb necessary for implementation in MATPRO. As to the properties of U0 2 pellets, the difference

between PWR and VVER types of fuel were found to be insignificant. This gave occasion not to duplicate

original MATPRO correlations of material properties of the fuel. It should be mentioned here that this

approach was also used, in particular, for thermal conductivity of fuel in burnup function. Earlier, during the

process of modifying FRAP-T6 code for analysis of high bumup VVER fuel rods [61, limited published data

were used for modeling degradation of thermal conductivity of fuel vs. bumup increase. At that time,

appropriate model in MATPRO-V1 I was not available. Recently developed FRAPTRAN code began to use

new FRAPCON-3 thermal conductivity model [12], which takes bumup effects into account. This model

was decided to be used in calculations of VVER fuel. Thus, within the framework of this activity in the

course of modification of MATPRO package, primary attention was focused on material properties of Zr-Nb

cladding.

2.1. Zr-i %oNb cladding thermal properties

Correlations of thermal properties of Zr-1%Nb cladding, incorporated in MATPRO package were taken

from available publications of domestic analysts. Table I contains a set of applicable correlations and

constants with references to source publications and routines of MATPRO package modified for

calculations of fuel rods with E1O claddings.

Table 1. Zr-1 %Nb cladding thermal properties implemented in MATPRO package.

Parameer, units W . W~ Nome$=$; turerRotiie Re.

Thermal conductivity X =[W/m K] T- temperature (K) CTHCON [13]

X=15.0636exp(0.4618-10 3T)

Specific heat Cp=[J/kg K] T- temperature (K) CCP [13, 14]

Slow heat up rate < 0.02 K/s (able 2)

Fast heat up rate > 1000 K/s (able 3)

Enthalpy H=[J/g] T- temperature (K); CCINP [13, 14]

TCdWT Tur current temperature (K)

E=f JCdT

293

Slow heat up rate < 0.02 K/s (Cp from Table 2)

Fast heat up rate > 1000 K/s (Cp from Table 3)

2
-P;armter, unit;--<4't622xXf- 2'' 9|'l'' '' i'|omenl Routine 4i Ref.? - Lineafthemi~1 expansion ~ jm/m] - - thermal expansion in hoop CTHEXP, [3,6, 13]

.;_ 5 7!. § X X iu direction (ntm)r

c,=0.1338985-10- 2+3.85875-10-6T- Ez - thermal expansion in axial

0.127813365-102. direction (m/m)

= 0.3336985.10T 2+5.65390106T- T- temperature (K)

0.199649865-10-2

883>. *: 77 7: : -

£z=0.13725577-102+5.4 10-(T-573)

cE=0.3336985-10-4T2 + 5.6539-1 0 6T - 0.19965-10-2

Ez =3.0465577.10-3+2.312.10-'(T-883) - 7.358.10 4 (T883)2+1.7211.1 0 (T-883)3

E65.5977. 10'3+2.312.104(T-883)-7.358.10 4(T883)2+1.721 l.10'1(T-883)3

fi=1.076459-103+9.7 10-6(T-1153)

e=3.627600-10- 3+9.7 104(T-1153)

Density,-p=[kg/m 3] p - density (kg/rn3) INTINP [15]

p(T=293) = 6550

Melting point, T,,.,, =[K] T,,,, - melting temperature (K) PHYPRP [15]

T ,l, = 2133 '_-_'_.

Heat of fusion H,, =[l/g] Hf - heat of fusion (J/g) PHYPRP [15]

H f=210 -_ __^_ _ __ _ _ - _ '

Phase transition temperatures T,=[K] - PHYPRP 16

Ta,,,+=883, Tc,, Ap=l153

Meyer nicro-hardness, HM=[MPa] ' T- temperature (K) CMHARD [16, 17]

T<0K : -44.4 .'

HM=2172.1 - 10.7055-T + 0.02765.T' - 3.278 10'5-T3

+

1.4231 07.T 4

44_ ., . 44.4 _.44>444444 -7e 2,iX 5:4 45 i",;r4i3 

Hm = exp(26.034 - 0.026394-T 4+ 4.3502-10 5 T2

2.5621.10'T 3) ' '

Table 2. Specific heat vs.- temperature under slow heat up rate [13].

.. 39 4 7:3~ 3'53' ji3~ 17 873 @88328 '-73. 102 1073< 1153~ 1173 41200............................

1.30 8P0

Specific ^ _ __ 1 1 1 1 

heat 345 360 370 380 383 385 448 680 816 770 400 392 392 393 393

(J/kg K) __ ___ . _ _ _ _ _

3
Table 3. Specific heat vs. temperature under fast heat up rate [13].

Cp = 237.5 + 15.91_10-2T, [J/kg K] 500 <T <1050 K

Cp = 199.7 + 12.364-10 2T, [J/kg K] 1200 <T < 1600 K

2.2. Zr-l %Nb cladding mechanical properties

Due to the difference in chemical composition and heat treatment, strength and ductility properties of asreceived claddings of Zircaloy and Zr-I%Nb type are significantly different, especially at low and medium

temperatures. In the case of highly irradiated cladding, difference in the levels of oxidation and

hydrogenation results in fundamentally different mechanical behavior of these claddings under accident

conditions. That is why incorporation of mechanical properties of both unirradiated and irradiated Zr-l %Nb

claddings was considered the key task in modifying the material property package for transient codes.

In the framework of such a task, the program on development of a modem database on mechanical

properties of the cladding used in the Russian pressurized water reactors of VVER type was initiated in NSI

RRC KI at the end of the 1990s. Direct adaptation of the obtained data to the code algorithm requirements

was an inseparable component of the program.

In the course of the program implementation [3], the main emphasis was made on studying of the properties

important from the fast accidental processes point of view. Reactivity initiated accidents (RIA) and early

stages of loss-of-coolant accidents (LOCA) can be considered among such processes. High strain rates,

intensive stress growth in the cladding and a wide range of cladding temperatures can be considered general

particularities of the cladding loading for these modes. Therefore, the main varied test parameters were the

temperature and the strain rate.

Two types of claddings were tested in the framework of the program: as-received claddings and claddings of

commercial fuel rods with the bumup of about 50 MWd/kg U.

The following kinds of tests were conducted with the cladding specimens:

* uniaxial tensile tests in transverse and rolling direction;

* biaxial tube burst tests with various biaxiality stress ratios.

The uniaxial tests were aimed at obtaining plastic deformation law parameters versus test conditions, and

cladding failure criteria were the objective of the biaxial burst tests. As a result of the initial stages of the

experimental program, the first version of the modified MATPRO package was acquired and used with

FRAP-T6 (modified) and SCANAIR codes for analyzing high bumup VVER fuel [6, 7].

The last result of the mechanical tests (uniaxial tension in rolling direction and low temperature biaxial

tension), which were reported in [5, 9, 10], allowed us to generalize the accumulated data and update

correlations for the plastic deformation law parameters. The most important result of the test program was

the anisotropy factors vs. temperature for unirradiated and irradiated cladding. The anisotropy factors

allowed us to derive deformation laws in terms of effective stress-effective strains, which seems to be more

correct for anisotropic claddings of Zr-based alloys. The other important thing was an extension of the

temperature range for the cladding failure criterion in the form of the true hoop stress at rupture. Thus, this

section presents the updated mechanical property correlations implemented in the MATPRO package (See

Table 4-Table 6). Table 4 contains a set of correlations of Zr-I%Nb cladding elastic properties - Poisson's

ratio, elastic and shear moduli taken from the literature.

4

1120 1134 142 1 1155 116| 1168 1177 180 1090 200

Specific heat, 420 480 600 1000 1600 1400 1000 600 400 360 348

(JIkg K )_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _
Table 4 also shows dependencies of Zr-l%Nb alloy high-temperature creep rate on the temperature and the

stress obtained in the unirradiated cladding samples [18]. These dependencies are incorporated into the

newly developed MATPRO module CREEPS (See Section 4), because the original version of the package

does not contain correlations for high-temperature cladding creep.

The implemented Zr-I%Nb alloy creep model takes into consideration, along with the temperature and the

stress level, volumetric ratio of a and P phases in phase transition area. Corresponding test data obtained for

equilibrium conditions are shown on Fig. 1. It is necessary to note that a literature search for the data on the

effect of heating/cooling rates upon phase transition temperatures of the alloy produced no results. At the

same time it is known that high cladding temperature variation rates typical for accidents shift C-P

transition boundaries. Effect of the heating rate upon annealing dynamics of irradiation-induced damages

should also be taken into account in predicting mechanical response of irradiated claddings. Therefore,

temperature dependencies of the main thermal and mechanical properties obtained mainly in thermodynamic

equilibrium conditions may require updating. Such updating is a subject of future activities related to

incorporating of heating/cooling rate into the set of the key test parameters.

It should repeat that currently the modified MATPRO package contains the cladding mechanical properties

for only two levels of burnup - 0 (as-fabricated) and 50 MWd/kgU.-Obtaining the continuous dependencies

on burnup or fast neutron fluence was out of frame of the test program.-However, basing on limited

literature data [1, 19] one can preliminarily assume that presented here correlations for irradiated El 10

cladding are applicable for bumup higher than 10 - 15 MWd/kgU. Additional work is needed to confirm

such assumption and to obtain quantitative estimations of burnup dependence.

Table 4. Mechanical properties of Zr-1 %Nb cladding.

Parameter, uim No;endature Rui Ref

Elastic modulus, E=[MPa] T- temperature (K) . CELMOD [13]

273 K<TS1073 K E=1.121-105 -64.38T

1073 K<TS1273 K E=9.129.104-45.OT

Poisson's ratio, v = [unitless] T- temperature (K) CSHEAR, [13]

T<1273 v= 0.42628-5.556-10-5T CLADF,

COUPLE:

Shear modulus, G = [MPa] T- temperature (K); CSHEAR

T<1273 K G E v - Poisson's ratio (unitless);

- - - 2-(1+ v) - - - E - elastic modulus (GPa)

Plastic deformation equation a - true effective stress (MPa); CKMN [101 -

K - strength coefficient (MPa);

.o TE ) . . .. . . E - true effective strain (unitless);

(see Table 5) n - strain hardening exponent

(unitless);

C - current strain rate (I/s);

.. - basic strain rate (lis);

Itl0 s;

m - strain rate sensitivity

-__________ -- ___________ ___________ ______ _ exponent (unitless)

5
Par. i eter, unitsV .t - Nom i datre <' 7 R outne' Rd .

Mechanical limits Sut - engineering ultimate strength CMLIMT [3,10] (MPa);gt CMIT 310

Sut Sy B. B. calculated by MATPO equations with SY - engineering yield stress

parameters of plastic stress-strain curve (MPa);

St - total elongation (%);

°u- uniform elongation (%);

True hoop stress at burst aB = [MPa]

293STS723 OaB =2016.268-5.2948 T +

unirradiated 0.00627 T2-2.8233.10-6 T3

423STS723 CTB -4178.356-12.894 T +

irradiated 0.0154 T2 -6.5545.106 T3

973'cT<1190 aB= 116139.02 exp(-0.0065753 T) T- temperature (K);

1190<TS1473 CTB= 7611.82 exp(-0.004283 T)

Anisotropy coefficients of Hill's equation, C, - effective clad stress (MPa); CANISO [10]

F, G, H =[unitless] CS9 - hoop clad stress (MPa);

a, = o(ae - ay)2 + G(a1 - a, )

2 + H(a, - (T)2. 2 - axial clad stress (MPa);

(see Table 6) c, - radial clad stress (MPa)

High-temperature creep strain rate, £ =[m/m] cr, - effective clad stress (MPa) CREEPS [18]

1. T<883 K i - effective creep strain rate (1ts)

* cra=9+32: £ =7.1.1 *0 5 o,22

exp(-28900/T), T- temperature (K)

* ca=32+90: £ = 26 a, 5 exp(-28900/T), f,, fp -normalized volume fraction

* C,>90: i = 2- 1 C exp(O.05a,)exp(-28900/T). of a- and 5- phase respectively (unitless)

2. T>1173 K £ = 0.09 a, 

3 5

exp (-13200/T), a., ca - effective stress in a- and

f- phase respectively (MPa)

3. 883ST<l070K a=f acrx+fDaf , Ca - creep strain rate in aand ,B- phase respectively (ls)

4. 1070STS 1173 K £C= fa£+ffi £_

Table S. Parameters of plastic deformation equation versus temperature and cladding type.

Parameter Type of cladding

Unirradiated Irradiated

293<TS797.9 K 293<Te763 K

K= 898.3710095 - 1.911883946T + K =916.8547193 - 0.6046334417.T -

0.002024675204T 2 - 9.628259856-10-7.T3 0.0002474820043-T2 Strength - :- : i- --- E - :- - ;:g; L''T' 

coefficient 797.9<T:1223 K 763<T:859.4K - (MPa) K = exp(-0.005608069738.T)- 15180.65748 K = exp(-0.00965027547T).491246.9131

859.4<T•1223 K

K = exp(-0.005608069738T) 15180.65748

6
Type of cladding Parameter

Unirradiated Irradiated

.,S_! -.ri - 293c~I •12- : V i 'i -i: -- 293 ' * F --

n= 0.04628421012 +0.000197951907-T- n =-0.1255447757 + 0.001350416112-T -

3.314868215-10-7T 2 + 13913294-10"'0 .T3 3.536814687. 10eT2 + 3.734672258.10 9T3

- .365014312-10-12

.T4

Strain 7<T79K

hardening

exponent n = -0.239614587 + 0.002839248035-T -

(unitless) 8.226160457I0-.T2 + 9.276772204-10 9 T3

- 3.588141876 10-12

.T4

_ .8_.9,1223 K ,-_

n - 0.04628421012 + 0.000197951907-T -

3.31486821510 7 -T2 + 1.3913294-l10'0 OT

. ._ i_;, . "293T•752. K T-7,

m = 0.02280034483 - 3.448275862-10-'-T Strain rate -- --- -T- -r - -

sensitivity - - , 

exponent m -2.534966886 + 0.006626767224-T - 5.303091629-10-T 2 + 1.34653092-10 9

.-T

(unitless) - - -- .,

;: ! : ! :S g;: f ! 02.1q 1223 K . :t : ;;.E

m = -0.1619955889 + 3 .080302048-10-T

Table 6. Anisotropy coefficients F, G, H versus temperature and cladding type.

Coefficient Cladding type

Coefficient

Unirradiated Irradiated

- -- 51293 273 - - 293< T515.9,K

F= 1.39239 - 4.63177-10-3 -T+ F= 4.82048 - 4.21033.10 2 -T + 1.618275-104-T2

1.62105-10-5-T 2 - 2.58537.10-8T+ -. - 2.68661-10-7-T + 1.60548*10 '

0.T4

1.8076.101 '.1 - 4.60713-10". 5 ..-

F v . lG-&--->1273K T-,= - -- 515.9T5823 K

(unitless) - - - -

F=0.5 F= 20.522409 - 1.14701-10-'.T + 2.46179- 10

4T2 _2.33290 10 7.T3 + 8.21321.10 ".T4

-T.823:K;- -

F=0.5

-23 ~~173K 93<eT:560KZ -

G=-6.6085-10 2+4.28093.10-3.T2 - G- 1.39276- 1792591-10-2-T+ 1.19333.14e-T2

1.51357- 1 05-T2 + 2A 818. 10-4 .T3 - - 3.776742- 10-7-T3 + 5.69241.1 0O rT4 -

1.72441- 10- .T4 + 4.49996-10'l5 .T5 3.247347-10'r 5

'Un'itless)' G ~~~ T>~1273 K 560<T83 K ....- ...4!' .' ...i;

G=0n5ess - 0 - G- 1.541960+8.71593610 3-T -

1.17013 10-5-T2 + 5.010771 10-9-T3

~~l K>83 

G=0.5

7
Coefficient Cladding type

Unirradiated Irradiated

293<T•T1273 K:'~-~L? >- 293<17'•823 K;-~

H= 0.173693 + 3.50846 104 -T- H= 0.5178583 - 1.71631-10-3-T+

H 1.074777- 10A6 T2 + 1.67189 10-9-T3 - 5.313208-10-*T 2 - 7.13646- 10-9 T3 +

(unitless) 8.31926. 1013_T 4 + 1.07169.10 o6-T5 3.870678. 1012_T4

T> 1273 K i- - 823 -

H=0.5 H=0.5

A- 

' 0.6 - - Beta

X 0.2 0

800 900 1000 1100 1200

Temperature (E)

Fig. 1. Volumetric fractions of a and P-phases depending on Zr-1%Nb alloy temperature [181.

2.3. High temperature oxidation kinetics

In the FRAPTRAN code, high temperature Zry cladding oxidation calculations were provided for using two

models selected by the user

1. Baker-Just model [20];

2. Cathcart-Pawel model [21].

At that, Baker-Just model is used as the conservative oxidation kinetics, and Cathcart-Pawel model - as the

best estimate model.

The same way as for the Zry cladding, two high temperature oxidation models for Zr-1%Nb cladding can be

suggested. Thus, the conservative kinetics used in Russia for licensed calculations of VVER fuel rods [22] was

implemented in the MATPRO package. The corresponding analytical correlations for weight gain, ZrO 2 and

aZr(O) layer thicknesses are given in Table 7.

As to the best estimate model, the authors of this report currently recommend using the existing CathcartPawel model as such a model for Zr-1%/oNb cladding. This model descnbes the last test data obtained at NSI

a-
RRC KI and RIAR the most closely (see Fig. 2). At that, developing of the separate best estimate kinetics for

El 10 is found to be premature for the time being due to incompleteness of the experimental program, the results of which will be published in 2002 in another NSIIIRSNINUREG reporL

Table 7. Conservative oxidation kinetics for Zr-l%Nb cladding.

iixtio a ra m I ImR;' T-=njeiatureiine . , M Routin'

Am=9.2- l 02exp(-10401 /o), Am - weight gain [mg/cm2] COXTHK

&z02=l.04exp(-2240/T)4h -Z,o2 - ZrO2 oxide layer [cm] COXWTK

&z,(o)=568 2exp(-6790fl)4T &ZO) - c-Zr(O) layer [cm] 1173<T<1773K

r- time (s);

T - temperature (K); ;

25

-- d0.iie side .

.D10 25 A

to; 

0 400 800 1200 1600 2000

Time (s)

Fig. 2. Comparison of test data on Zr-l%Nb steam oxidation with Cathcart-Pawel model predictions.

9
3. FRAPTRAN MODIFICATIONS FOR VVER FUEL ROD SIMULATION

This section gives a description of modifications made to the original version of FRAPTRAN. As was

already noted, in addition to incorporating alternative cladding material properties, certain modifications of

models and algorithms were necessary in order for the code to be able to correctly simulate VVER fuel rod

behavior. Mainly, the modifications were imposed by the following characteristics of the fuel geometry and

conditions of tests, which were selected as the assessment cases:

* presence of a central hole in the fuel pellet. This condition had to be taken into consideration in

calculating of volumetric heat generation rate, normalized radial power profile, and fuel radial

displacement;

• carrying out of power pulse tests of single rods in ampoule conditions (large volume of stagnant water

under normal conditions). These test rod cooling conditions cannot be quite adequately described by the

heat transfer models that exist in the original version of FRAPTRAN. Therefore, alternative models for

film boiling and rewetting models were needed;

* absence of strain component in the cladding mechanical model due to high temperature creep.

Neglecting creep during certain stages of LOCA when stresses in the cladding may not reach the yield

point and/or may decrease with time may result in erroneous predictions of deformation and rupture in

cases when only instantaneous plasticity model is used. That is why incorporating the creep components

into mechanical calculations was considered an important task;

* significant scope of modifications in the material property models and in the calculation modules of the

body text of the code caused introduction of a large number of global and local variables, input/output

variables, as well as writing of a number of new subroutines. Therefore, this section gives a detailed

explanation of the modifications pertaining to coding aspects. This might prove to be useful both. for

developers and users of the code.

3.1. Volumetric heat generation rate (HGR) calculation for fuel pellet with central

hole

In the original FRAPTRAN version, radially averaged volumetric heat generation rate is calculated only for

a fuel pellet without the central hole:

q = q,2 

where q, - radially averaged volumetric HGR (Wtm3);

ql- radially averaged linear HGR (W/m);

rf - fuel outer radius (m).

In this work, the determination of the radially averaged volumetric heat generation rate in the case of

VVER-type fuel pellets with central hole was modified as:

vr.rfir)

where rfi - fuel inner radius (m).

3.2. Radialpower profile calculation forfuelpellet with central hole

Initially, radial power distribution is determined for a fuel pellet without central hole with the following

normalization:

10
K,dr

2 2=1 

where K , - normalized radial power factor (unitless);

qi (r) - linear HGR at current fuel radius (W/m);

q, - radially averaged linear HGR (W/m).

The condition for normalized radial power factor in the case of VVER-type fuel pellets with central hole is

introduced in the following form:

|K,dr

,=.

- r)

3.3. Calculation of the radial displacement offuel due to thermal expansion for

pellet with central hole

The radial displacement of the fuel pellet due to thermal expansion was modified for the case of fuel with

central hole in the following manner:

UT = E(T)dr,

where UT - fuel radial displacement (m);-

e(T) - fuel thermal expansion (n/m);

rf - fuel outer radius (m);

rr - fuel inner radius (m).

Additional term Uc of the pellet radial displacement named as "hour-glassing" effect is calculated using the

expression:

Uc=0.0025rf0 ,

Uc=0.0025ri0(l-P134.5), if O<P<34.5,

U,=O, if P>34.5,

.where P - fuel/cladding contact pressure (MPa).

Besides, the fuel relocation term U,, is accounted for in the fuel radial displacement:

U,r=O.3MAz - for fresh fuel,-

U,,O=.45Agap - for high-bumup fuel,

where A,,p - initial gap width.

At the current stage of the code assessment, the following effects of the pellet radial displacement were

elininated for VVER-type fuel pellets:

* Fuel expansion due to "hour-glassing" effect;

* Fuel relocation at the beginning of a pulse test (U,,=O).

-1 1
3.4. Alternative Post-Critical Heat Flux heat transfer model

The model for post-critical heat transfer coefficient (HTC) calculations based on Bromley-Pomerantz

correlation [23] was replaced with the Labuntzov model. This model was developed for the turbulent

regimes of film boiling and was modified to account for the boiling conditions of the large volume of

subcooled water [24]:

a,?, = 0.25(Xcpg(pf pd 9 )113

where a,, - heat transfer coefficient (W/m2K);

Al- vapor thermal conductivity (W/m K);

cpg- vapor specific heat /kg K);

pf - fluid density (kg/m3);

pg- vapor density (kg/r 3);

g - Gravity acceleration (Mis 2 );

vg - vapor kinematic viscosity (m2Is).

To take into account the initial subcooling of water, the correction factor is introduced [25]:

a; = a,, Q + O. ( Pr '075 A'

p0 hfg

where LFB - corrected heat transfer coefficient (W/m2 K);

Ai - enthalpy of fluid at saturation minus enthalpy at fluid bulk temperature (J/kg).

Thermal-physical properties of vapor are determined at film temperature:

T = T.,, + T

2

where T,, - cladding temperature (K);

T,- saturation temperature (K).

3.5. Cladding rewetting modelfor IGRIRIA tests

The moment when rewetting begins is determined with the model developed by RRC KI [3, 26]. Rewetting

moment calculation option is specified in the input data. Heat transfer coefficient in the transition mode

from film boiling to nucleate boiling in determined by means of linear interpolating between two points on

the boiling curve. The first point corresponds to the film boiling heat transfer coefficient at the moment

when rewetting begins. The second point corresponds to the heat transfer coefficient in the case of complete

wetting, i.e. at nucleate boiling, and is determined per critical heat flux of the second type. 2-D heat

conduction equation with quench front movement boundary condition is resolved analytically:

whr d- sz )ir hat( c i

where C - specific heat capacity;

T- temperature (r,z,v); 12
A- clad heat conductivity;

q - volumetric internal heat source;

p- clad density.

Obtained approximate analytical solution for quench front velocity and time of rewetting is expressed by the

following equation:

z(t) = u(t)dt,

to

where t- time;

z- axial coordinate;

U(t) - quench front movement rate;

t - time when transition boiling begins (Tclad=640 K);

t - rewetting time.

Currently the rewetting model does not implenented into FRAPTRAN code. The time of rewetting is

specified on the base of previous FRAP-T6 (modified) calculations.

3.6. Zr-1%1vNb high temperature creep model forMATPRO and SCANAIR

In the FRAPTRAN code, cladding stress-strain equations include components of thermal, elastic and plastic

deformation [I].

Description of the fuel rod deformation under the loss-of-coolant accident conditions, as compared to the

reactivity increase accident conditions, has a number of peculiarities, the most important of which is that, in

a number of LOCA scenarios, the duration of the high temperature cladding deformation process is on the

order of hundreds of seconds. Besides, cladding deformation over a significant period of time can occur

under the mechanical loading conditions when stresses in the cladding do not exceed the material yield

stress, or at monotonous mechanical load dropping. Therefore, it is proposed to modify cladding mechanical

behavior equations using material visco-plastic deformation model.

Modification of the cladding stress-strain equations consists in that elastic and plastic deformation equation

components are supplemented with viscous deformation components.

Thus, the cladding mechanical model in the FRAPTRAN (modified) code includes description of cladding

thermal expansion effects, elastoplastic deformation effects and cladding material high temperature creep

effects. Additivity assumption for thermal, elastic, plastic and viscous strains is used. In order to describe

cladding mechanical state, Hooke's law and Prandtl-Reuss visco-plastic flow model that can be written down

in strain increments as shown below are used:

{ -V. +ŽE: +e +d4 .+e; fa)

1 - - * ~~~~~~~~~~~~T r.o 

T

E. I-{C - Va-, + E + d + , + de. + adT -- (2) E r

Er fa (3)

. r~~~~~~o

where To - initial cladding temperature;

T - current cladding temperature;

a - temperature expansion factor;

E - modulus of elasticity;

v - Poisson's ratio;

E, , - full cladding deformation in hoop, axial and radial directions;

13
EO ,C e, Er plastic cladding deformation in hoop, axial and radial directions;

dep, dsP, ds,p - plastic cladding deformation in hoop, axial and radial directions at time increment dt;

e% eC _ viscous cladding deformation in hoop, axial and radial directions;

dEc, de-%c dEc - viscous cladding deformation in hoop, axial and radial directions at time increment dt;

Ur,,e ar - radial, hoop and axial stress components, respectively.

At each time increment, plastic and viscous strain increments (dEP and dc) are determined in accordance

with the Prandtl-Reuss flow rule.

Increments of the plastic strain components are determined per effective stress a and effective plastic strain

according to the stress-strain diagram. Creep strain component increments - per effective stress and average

cladding temperature in accordance with experimentally obtained dependence for material creep rate given in

Table 4 of Section 2.2.

Modified technique for fuel rod cladding deformation behavior calculation is implemented in CLADF and

BALON2 subroutines. Block diagram for calculating of cladding mechanical response using the modified

model in FRAPTRAN code is shown in Fig. 3.

Fig Bloc 3. diga , for cacltn of clddn mehnia repos usn !. modi: fied mode -L. in -

FRAPTRANcode (mdiPTed)

3.7.1. Descrip-ion of theMechanical cacatione : I - .. -;E...::. ^-DEFORM-,FCl cm.;:'-. i:V,, .j 4.

The- ne glba variabe ar intoue fo the nel dveloped an - t moifie .- modes. Tab- S 0 cont ;- -l fin the

14~~~~~~7

-~~~~ 0 ;^ <>--.oCOUPLE" gap size oegp -

;;:: f :-; ^ E closedgap- aclto =5_' ,;, , :

*~ ~~ ~ ~ -- ~ 'k2i' ~ , , ~ M ~~~~~N" i t; f Yes;0,;,,AE,,-,3\i 

: ; ----- : ;f - - - :;-;-Axisymrnetric deformation -F- -^ 70 Lcal looning -- ::: ; -.; [ --y ;-:4 "'CLADF" ";.i!i BALON2" _- 

. . iW.'.-,! ' t 0 ,Creep deformnation CREEPS" 

^~~~~~~~~~~~~~~~~~~~~~~~~~i'.~ k' E ' :K;:7i"-'.; j i~:

Fig. 3. Block diagram for calculating of cladding mechanical response using modified model in

FRAPTRAN code (modified).

3. 7. Description of modifed routines and coding aspects

3.7.1. Description of the new global variables

'Me new global variables are introduced for the newly developed and modified models. Table contains the

4

I
description of the new global variables. All new global variables stored in WWER.H header. Currently

IWWER indicator is temporarily used to recognize the cladding type; In the near future, ICM indicator with

option of ICM=6 for Zr-I %Nb cladding type will be introduced in FRAPTRAN and MATPRO subroutines

instead of IWWER indicator.

Table 8. New global variables incorporated into FRAPTRAN code.

Name Type - tra Desc i - -

iwwer INTEGER *8 wwer.h Option for selection of Zr-1 %Nb cladding type

common /wwer/

nfilm INTEGER*8 wwer.h Indicator for Labuntzov or Bromley-Pomerantz models

common /wwer/ - used for film boiling heat transfer calculation

icreep INTEGER*8 wwer.h Indicator to use the cladding creep model:

common /wwer/ 0 - the model is not used

I - creep model is used

jzmax INTEGER*8 wwer.h Axial slice number for printout of time dependent fuel

common /wwer/ rod parameters

rfi - REAL*8 wwer.h Radius of central hole in pellets. rfi is used in HEAT

common /mech/ and POWER routines to correctly compute volumetric

heat generation rate in fuel pellet with central hole

irr REAL*8 wwer.h Indicator for selectibn of Zr-l%Nb cladding mechanical

common /wwer/ properties:

o - unirradiated cladding properties

I - irradiated cladding properties

moxid REAL*8 cobild.f Indicator to select the model for high-temperature

oxidation of Zr-I%Nb cladding:

5 - conservative model

3.7.2. Description of the new and modified subroutines

Table 9 lists the modified subroutines of FRAPTRAN code and presents description of the new and

modified subroutines.

Table 9. List of modified routines of FRAPTRAN code.

I I I IM

Although the name of the header "WWER" is rather arbitrary, in order to avoid contradictions it needs to be said that

in literature one may find double spelling of the abbreviation for the Russian pressurized water reactor - VVER and

WWER. 

1 5

____9St,0> __ inplemenntton o Zr-1L %N b matenai pro rues into MAIrRU ongrai suoroutines

CCP Function for calculation of clad specific Zr-I %Nb clad specific heat. _________heat______________ ______

CCPINP Driver for calculation of clad enthalpy Zr-I %Nb clad enthalpy.

CTHEXP Driver for calculation of clad thermal Zr-1 %Nb clad thermal expansion. __________expansion_________________________

Driver for calculation of -clad Young's Zr-I %Nb clad Young's modulus. CELMOD modulus I__ ___ ___ ___ ___ ___ ___ __I

Function for calculation of clad shear Zr-I %Nb clad shear modulus.

modulus i

I .- .

I I

-

 q : ,
M -I . . I

Driver for calculation of clad thermal

conductivity -

Zr-I %Nb clad thermal conductivity.

CMHARD Function for calculation of clad Mayer Zr-I %Nb clad Mayer micro-hardness.

rnicro-hardness

Driver for calculation of K, i, n clad Calculation of Zr-i%Nb K, m, n clad parameters for fresh

CKMN parameters for clad plastic stress-strain and irradiated type of cladding.

curve

CMLIMT Driver for calculation of clad short-term Calculation of Zr-1%Nb clad uldimate strength parameters

strength and plastic parameters for cladding failure evaluation.

Routine for calculation of cladding Calculation of anisotropy coefficients F, G. H for fresh

CANISO anisotropy coefficients. CANISO is not and irradiated type of Zr-I %Nb cladding.

used in mechanical calculation

Driver for calculation of high-temperature Calculation of weight gain, ECR, Zr- and Zr(O)-oxide

COBILD oxidation layers for Zr-1%Nb clad. Driver calls COXTHK and

COWTK subroutines.

Driver for calculation of fuel and cladding Introduction of thermal-physical parameters for Zr-1%Nb

PHYPRP thermal-physical parameters clad (fusion heat, alpha- and beta-phase temperatures, melt

point etc.).

- -New rndels an Idsuboutines~ f

Driver for calculation of clad-to-coolant Calculation of clad-to-water heat transfer coefficient was

heat transfer coefficients developed for the turbulent regimes of film boiling and

QDOT modified to account for the boiling conditions of the large

volume of subcooled water (Labuntzov model). Input

pararneter nfilm is used as optional variable

New subroutine for clad creep rate Computation of creep rate of Zr-1%Nb cladding versus

CREEPS calculation under high-temperature temperature and effective stress. Subroutine was coded for

conditions the MATPRO package

New subroutine adopted from MATPRO/ COXTHK and COXWTK subroutines contain

RELAP5 [27] and supplied with the Zr- conservative model (moxid=5) of high-temperature

COXTHK I %Nb growth rate constant for oxide oxidation for Zr-I %Nb.

thickness, oxygen-stabilized alpha layer moxid - optional variable for oxidation models

thickness, and thickness of the beta layer

New subroutine adopted from MATPRO/

COXWTK RELAP5 [27] and supplied with the growth rate constant for weight gain for

Zr-I%Nb cladding -

+ . icaionsad orrect s 9.i40 t _ffi.;t---rK--f-i>--@9-<>,;>Su-Mod 

Print output information in listing 1. Implementation of the additional output parameters to

lisdng and time dependent parameters for secondary plot

PRNTOT file 'FORT.DAT

2. Correction of the average fuel temperature calculation

accounting for a central hole in the fuel pellet

Major driver. Routine reads control Reading new input variables for VVER conditions (iwwer,

variables nfilm) and a fuel rod slice number to plot (jzmax)

To account for user-specified fission gas Elimination of the mistake associated with dimensions of

release in calculation relfrac input array is relfrac and FuelGasSwell arrays. To correct reading of the

determined as: time dependent tables of fission gas release and transient

FRAPTRAN relfrac(mfgr) fuel swelling, the dimensions of relfrac and FuelGasSwell

nfgr=25 arrays were set as:

User-specified array of fuel displacement mfgr=50

due to fuel gas swelling is determined as mfs=50

time-dependent displacement table: Thus, all 25 pairs of points can be read from input deck.

FuelGasSwell(mfs)

mfs=25

Driver reads input data in NAMELIST Extension of clad nodalization by implementation of

format and clad nodalization setup. integer variable nmesh and cmesh array for clad radial

IOFILES Default value of clad mesh points coordinates into NAMELIST:

ncmesh=2 namelist /solution/ dtmaxa, dtss, prsacc, tnpacl, soltyp,

maxit, noiter, epshtl, naxn, zelev, nfmesh, fmesh, ncmesh,

cmesh

16

CTHCON
=I IL I S I I 

Recalculation and processing input

parameters of fuel rod. Integration of

radial nower profile in fuel

Correction of calculation of radial power profile integral

in fuel pellet with central hole

INITIA Intitialization of all variables in -transient The same problem as in COMINP

calculations. Integration of radial power

profile in fuel

Calculation of fission gas inventory in Introduction of fuel volume calculation for pellets with

high-burnup fuel. Fuel volume is central hole to obtain generated fission gas products

calculated only for fuel without central

hole -_-_-_-_ _ _ _

Mechanical calculation of fuel-cladding For Zr-I %Nb cladding clad temperature limit equal to

interaction. Clad temperature limit 1030 K was obtained.

COUPLE (tedot=1089 K), at which the cladding tedot=1030 K

strain rate becomes excessive, is set.-In This limit was obtained during checking of the code

this case calculation continues bypassing working with Zr-I %Nb mechanical property correlations

the normal iterative solution -

FRAPTRAN uniform cladding strain for Elimination of uniform cladding strain for Zry adopted

Zry cladding was adopted from from FRAPCON-3 code. Because this correlation is valid

FRAPCON-3 code [28]: only in low-temperature range between 580 and 680 K.

e, =0.096-1.142-10'T+ Under temperatures more than 840 K uniform strain

becomes negative.

CMLIMT 0.01856exp(-F/1025) - So, the dependence of uniform strain is described by

8.05*10 FRAP-T6 (V21) relationship both for Zry and Zr-1%Nb

T-temperature (K); cladding:

F - fast fluence (nm 2); n

h, - excess hydrogen concentration (ppm) e- exp(-) -I 1+m

n - strain hardening exponent (unitless);

m - strain rate sensitivity exponent (unitless)

Subroutine specifies fuel rod power as Correction of the way volumetric heat generation rate is

function of time and axial elevation. determined for fuel with central hole. So,

Determination of volumetric heat I q__

generation rate (HGR) in fuel in following For the fuel without hole q = 2

form:

POWER q =4q , For the fuel with hole q,=

q,- volumetric HGR (W/m3 ); r, - fuel inner radius (m)

q - linear HGR (W1m);

rf,, - fuel outer radius (m)

Driver for global computation of the fuel 1. Correction of average fuel temperature calculation in

rod variables vs. time. Routine calls fuel pellet accounting for central hole (the same correction

COBILD, where clad linear power due to as in PRNTOT routine).

oxidation reaction is computed by 2. Elimination of clad temperature limit for beginning of

equation: high-temperature oxidation (tempcl>1073 K). Time

P= 1.15 lOW Aw Dfit(2 dt), increment less than 0.001s leads to prompt increase of

COMPUT P - linear power (W/m); clad linear power (at the first time increment) and the clad

Aw - weight gain (kg/m2); temperature above the limit of 1073 K This results in

Dfr - fuel rod outer diameter (m); local increase of outer clad temperature. So, the

dt - time increment (s) elimination of temperature threshold provides for the

Clad temperature limit for beginning of contnuity of linear power function vs. clad temperature.

high-temperature oxidation (tempcl) is

equal to 1073_K 

1 7

COMINP
'a I~~~M111111 I I Ie*

In COMPUT routine, determination of the

current value of transient fuel swelling is

done by interpolation:

TranFuelSwell =

POLATE(FuelGasSwell,Time,50,iu)

where:

'FuelGasSwell' - time dependent input

fuel swelling table

'Time' - current time

- Correction of wrong entry to POLATE function:

TranFuelSwell = POLATE(FuelGasSwell,Time,25,iu)

25 - number of pairs of entries in 'FuelGasSwell' array

Computation of the deformed fuel radius Modification of the approach to determine deformed fuel

due to transient fuel swelling: radius due to transient fuel swelling:

RF=R,f *TranFuelSwell Rf-R&r + R, I1d-(TranFuelSwell-l)

Rf- fuel radius (m); RC0,d - cold fuel radius (m);

R&f- deformed fuel radius due to thermal TranFuelSwell - relative change in cold fuel radius (ld)

DEFORM expansion, relocation and steady-state due to transient fuel swelling (unitless).

swelling (m); This approach is more suitable for specifying deformation

TranFuelSwell - relative change in history of fuel radius based on the net transient fuel

deformed fuel radius (Rde) due to transient swelling

fuel swelling (1.0 - no fuel swelling)

(unitless)

Routines: IOFILES subroutine controls the input Extension of 'RodAvePower' array for average linear heat

IOFILES data and specifies the array for average generation rate from 100 points to 1000:

CARDIN linear heat generation rate: npthal=2000

POWER npthal=200 RodAvePower (npthal,100)

FCMI2 RodAvePower(npthal,100) Thus, 'RodAvePower' array contains 1000 pairs of timeHeaders: That is the 'RodAvePower' array that power points

BCDCOM contains 100 pairs of time-power points

POWRD

3.7.3. Description of the new input variables

Temporarily the additional paramneters occupy three first strings in the input deck specification. Description

of the new input parameters is presented in Table 10.

Table 10. List of new input data parameters.

Parameter Description Type.- Value

iwwer VVER/PWR type of cladding INTEGER*8

* Cladding - Zircaloy 0

* Cladding- Zr-1%Nb 1, 2*

nfilm Indicator of the film boiling HTC INTEGER*8

* Bromley-Pomerantz model 0

* Labuntzov model I

icreep Indicator to use cladding creep model INTEGER*8

* Cladding creep model isn't used 0

* Cladding creep model is used

jzmax Axial fuel rod slice number for the INTEGER*8 1-kmax

secondary plot file 'FORT.DAT'

IWWER=l implies using heat capacity of Zr-I %Nb obtained for low heatup rates, whereas under IWWER=2 heat

capacity for high heatup rates (1000 K/s) is used. Other thermal and mechanical properties are identical.

18
3.7.4. Description of additional output information

Generating of the following files is provided for in order to depict information on cladding residual strain

caused by high temperature creep (Table 11).

Table 11. Output information files.

Filename 'Outpqut infrmationi paeter-aibe aeFra

CREEPH.DAT Cladding creep hoop strain creeph REAL*8

-_______________ array(50)

CREEPZ.DAT Cladding creep axial deformation creepz REAL*8

array(50)

CREEPR.DAT Cladding creep radial deformation creepr REAL*8

array(50)

ERESH.DAT Total cladding creep and plasticity hoop eresh, REAL*8

strain array(50)

Relative contribution of plasticity and creep spl REAL*8 constant

deformation into the overall deformation (for screep REAL*8 constant

the maximum stress cross-section) -

ERESZ.DAT Total cladding creep and plasticity axial eresz REAL*8

deformation array(50)

ERESR.DAT Total cladding creep and plasticity radial eresr REAL*8

deformation array(50)

Cladding strain is recorded in the output files depending on time for each axial segment of the cladding. Description of additional parameters in output data files is provided in Table 12.

Table 12. Additional parameters in output data files.

Variable naei Paraneter inito measurement

ts Time s

creephj) Cladding creep hoop strain %

creezh(j) -Cladding creep axial deformation

creerhG) Cladding creep radial deformation %

eresh() Total cladding creep and plasticity hoop strain %

splasta(zmax*) Relative contribution of plasticity deformation into the %

overall deformation

screepazmax) Relative contribution of creep deformation into the %

overall deformation

eresz) Total cladding creep and plasticity axial deformation %

eresr() Total cladding creep and plasticity radial deformation

19

,. . I
4. CODE ASSESSMENT AGAINST IN-PILE IGRLRIA AND OUT-OF-PILE

RARALOCA TEST DATA

The most general principles that called forth the test case selection for the present work are illustrated in

Table 13. There is an assessment matrix of fuel rods tested in IGR reactor under RIA conditions and in

RIAR out-of-pile LOCA test facility presented there.

Three assessment cases were selected to simulate fresh fuel rod behavior (instrumented test rods 50F-13,

50F-16, 96F-09). Test results were compared with calculation data obtained with of FRAPTRAN (original

version), FRAPTRAN (modified), FRAPT-T6 (modified) and SCANAIR (modified) codes. As is shown in

Table 13, the assessment cases for fresh fuel rods present the range of cladding temperatures from 293 to

1840 K and characterize boiling curve from free convection heat transfer to film boiling.

Three assessment cases were chosen to simulate high-bumup fuel rod behavior (H5T, H7T and HIT). High

burnup fuel rod modeling results (H5T, H7T and HIT) were obtained earlier using FRAPT-T6 (modified),

SCANAIR (modified) codes and are given in work [3]. The calculation results obtained using FRAPTRAN

(modified) were compared with the results obtained using FRAPT-T6 (modified), SCANAIR (modified)

codes. The obtained discrepancies in the main fuel rod thermal mechanical parameters were analyzed, and

possible reasons for the differences were discovered.

Table 13. Assessment matrix of fuel rods tested in IGR reactor under RIA conditions and RIAR outof-pile LOCA tests.

sessed f h

c. 7ca/g I h5m a"prameters, - oralculation rults

50F-13 RIA with 69 Nucleate boiling cladding temperature Cladding temperature

fresh fuel __ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _history

50F-16 RIA with 139 Nucleate boiling cladding temperature Cladding temperature

fresh fuel CHF history

Post-CHF heat

transfer

96F-09 RIA with 315 Nucleate boiling cladding temperature, Cladding temperature

fresh fuel CHF intemal gas pressure and internal pressure

Post-CHF heat history

transfer

H5T High burnup 176 Nucleate Fuel and cladding Comparison with

RIA test boiling, CHF, temperature, internal FRAP-T6 (modified)

Post-CHF heat gas pressure, cladding and SCANAIR

transfer stress and strain, gap (modified) calculation

H7T High bumup 187 width etc. data

RIA test

HIT High bumup 151

RIA test

RIAR- Out-of-pile - Cladding cladding strain, time Comparison with test

LOCA2 LOCA test deformation and of failure data on cladding

cladding failure strains and time of

failure

20
In order to verify FRAPTRAN code under LOCA conditions, calculations of Zr-I %oNb cladding deformation

behavior under varying cladding heating and cooling modes and pressure increase rate (RLAR-LOCA2) were

conducted. To study fuel rod behavior during the first stage of LOCA, a special facility with direct electric

heating of fuel rods [1] was built at RLAR. This facility allows to conduct testing of fuel rod simulators with

aluminum oxide filler in argon environment under the conditions modeling the first stage of LOCA. More

detailed description of the installation and test procedures is given in Section 4.2. .

4.1. IGRIRIA assessment

Principal fuel rod design parameters and IGR test conditions are listed in Table 14. The very detailed

description of test data on power pulses, energy depositions, temperatures, and internal gas pressure is

published in [3]. It should be noted that all selected test cases have already been discussed in detail in [3, 6,

7], where verification procedures for modified FRAP-T6 and SCANAIR codes were described. Therefore,

comparison of FRAPTRAN calculations was performed not only with test data, but also with results

obtained using modified versions of FRAP-T6 and SCANAIR codes.

Table 14. Brief description of fuel rod design parameters and IGR test conditions.

P rameter- Units taue

Fuel rod type: - VVER-1000

Cladding - Zr-l%Nb

Fuel U0 2

Burnup of mother rod MWd/kg U 0; -50

Gas composition: He-100%

Coolant H20

Power half pulse width ms 700-900

Cladding outside diameter mm 9.1

Radial gap thickness mm 0.03-0.12

Fuel pellet outer diameter mm 7.57

Fuel pellet inner diameter mm 2.2

Fuel stack height m 0.15

Rod internal pressure MPa 1.7-2.3

Coolant pressure MPa 0.1

Coolant temperature K 293

-Coolant velocity m/s 0.

4.1.1. 50F-13 IGR assessment case

Comparison of the measured and calculated cladding temperature histories is shown in Fig. 4. Predictions of

the cladding temperatures under nucleate boiling conditions demonstrate a satisfactory agreement with the

test data.

21
600 30

500 - - - S emperanir m W u) -Power (rmasuTed)

2 / . t~ 1 _ 20

0. I-~~~~~~~~~~~~~~~~~~~~~1

E 400

J < ~~~~~~~IGR test 50-13

200 - -

0.0 2.0 4.0 6.0 8.0 10.0

Time (s)

Fig. 4. Clad outer temperature in 5OF-13 test: comparison between measured and calculated with FRAPT6 (modified), SCANAIR (modified) and FRAPTRAN (modified) codes.

4.1.2. 50F-16 IGR assessment case

The cladding temperature history during IGR pulse test and calculated cladding temperature obtained with

original version of the FRAPTRAN code are presented in Fig. 5. As shown in the plot, melting of cladding is

predicted with original FRAPTRAN version. -~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~

Fig. 5. Clad outer temperature in 5OF-16 test: comparison between measured and calculated with original FRAPTRAN code. 2 2

_ Temperature (calc. by FRAPTRAN (modified) code)j \ -4- + Temperature (calc. by FRAP-T6 (modified) code)

- 1 Temperature (caic. by SCANAIR (modified) code)

_1 --- olao{^1 I---- I 

-U-- Temperature (calc. by FRAPTRANriginal code)

2000 - --- Temperature (measured) 500

- Power (measured)

1600 - /_ / - 400

mi 1200- 300 i

2 l \ ,* '. IGR test 50-16

3,800 - < '200 ;

-4~~~~~~~~~~~~~~~~~~~~C

400< ='00

0.0 2.0 4.0 6.0 8.0

Time (s)
Fig. 6. Clad outer temperature in 50F-16 test: comparison between measured and calculated with FRAPT6 (modified), SCANAIR (modified), and FRAPTRAN (modified) codes.

After the modifications of heat transfer model described in Section 3 of this report, the FRAPTRAN (modified) rather satisfactory predicts cladding temperature under nucleate boiling, transition boiling, and film boiling conditions (see Fig. 6).

4.1.3. 96F-09 IGR assessment case

Cladding temperature and internal gas pressure histories were measured during 96F-09 IGR pulse test [3].

These data and calculated results are shown in Fig. 7, Fig. 8, and Fig. 9. As the first step of assessment, the

calculation with original version of FRAPTRAN was carried out. Cladding temperature history in this case is

presented in Fig. 7. As shown in Fig. 7, cladding temperature rapidly achieved melting point due to insufficiently high clad-to-coolant heat transfer in post-CHF boiling regime.

Then the post-critical heat transfer calculation model based on Bromley-Pomerantz correlation was replaced

with the Labuntzov model. After modification of the film boiling heat transfer coefficient, the next run was

carried out with FRAPTRAN (modified). Thus, the comparison between measured cladding temperature and

calculated data is presented in Fig. 8. Calculated cladding temperature was obtained previously with FRAP-T6

(modified). As shown in Fig. 8, the cladding temperature calculated with FRAPTRAN (modified) is in better

agreement with experimental data than respective FRAP-T6 (modified) results. As shown in Fig. 9 the time of

cladding failure is predicted accurately enough.

23

1600 - + Temperature (caic. by FRAPTRAN (modified) code) 500 - Temperature (calc. by FRAP-T6 (modified) code)

-*-- Temperature (calc. by SCANAIR (modified) code)

- -S -. Temperature (measured)

Power (measured) 400

5. 1200

i

300

20

0~~~~~~~~~~~~~~~~~~~~~~

400-

IGR test 50-16

0.0 2.0 4.0 6.0 8.0

Time (s)
Fig. 7. Clad outer temperature in 96F-09 test calculated with Bromley-Pomerantz post-CHF heat transfer.

Fig. 8. Clad outer temperature in 96F-09 test calculated with Labuntzov post-CHF beat transfer.

24

2400 - I I

IGR test 96-09

2000-

* ~ ~~~,Temperature (caic. by FRAPTRAN-original code)~

1600 -- -- 4 - -. Temperature (measured)

1200 -

0

-0 800 -

400- ------

0-

0.0 2.0 4.0 6.0 8.0 10.0 12.0

Time (s)

2000-

/ X A 0-itf ~~Temperatre(cal. byFRAPTRAN (modified) ode)

/ ^6 s Q 411 + ~Temperature (calc. by FRAP-T6 (odified) ode)~

1600- - Temperature (measured), -i

2 _W

cl1200 - t

800 0

Ca

400

IGR test 96-09

0.0 2.0 4.0 6.0 8.0 10.0 12.0

Time (s)
3.6

Fig. 9. Comparison of internal rod pressure in 96F-09 test, calculated with FRAP-T6 (modified) and

FRAPTRAN (modified) codes.

4.1.4. High-burnup fuel simulation of IGRIRIA test

Earlier (see [30]), test fuel rod H5T was selected as an assessment case for the FRAPTRAN code.

FRAPTRAN code developers conducted analysis of thermal mechanical parameters of fuel rod 1HST,

influence of fission gas release on thermal and mechanical behavior of the fuel rod. For boundary conditions

for fuel rod calculation, rod-to-coolant heat-transfer coefficient at constant coolant temperature was

assigned. In addition to that, initial data contained simplified fission gas release history, which did not

account for transient gas swelling effect.

In this work, more detailed and refined method for assigning initial data for calculating high-burnup fuel

using FRAPTRAN(modified) code is presented.

To predict behavior of high-bumup fuel rods, two methods for generating initial data are provided:-

1. Calculation procedure providing for use of steady-state FRAPCON-3 code [28] for base irradiation

regime. FRAPCON-3 code generates a calculation data array at the end of the base radiation, which serve

as initial data for FRAPTRAN code.

2. Fuel rod calculation using FRAPTRAN code without FRAPCON-3 code. In this case, parameters of the

fuel rod after the base operation in the reactor are assigned as initial data. Fuel and cladding parameters

shall be obtained on the basis of post-irradiation examination of the fuel rod.

When analyzing calculation results, one should keep in mind that FRAPTRAN code does .not calculate

fission gas release (FGR) and transient swelling of the fuel matrix. To account for effects of fission gas

release and transient swelling of fuel, initial data should include timetables for fission gas release and for the

swelling of fuel pellets. In other words, temperature regime of the fuel rod does not affect fission gas release

and swelling of the fuel. Dependencies of fission gas release and swelling vs. time were assigned on the

basis of previous FRAP-T6 (modified) assisted calculations [3].

RIAR has conducted post-radiation studies of fuel rods irradiated in VVER-1000 reactor ( unit of NVNPP)

to the bumup of 50 MWday/kg U. As a result, data on deformation of fuel and cladding, nuclide

25

IGR test 96-0 Pressure (calc. by FRAPTRAN (modified) code)

- Pressure (calc. by FRAP-T6 (modified) code) j / t7 ' -- U - Pressure (measured)

1.2.

0.0

0.0 1.0 ' 2.0 3.0 4.0

- Time (s)
composition of fuel, cladding corrosion, fission gas release, etc. were obtained [3]. Therefore, to model

behavior of bum-up fuel rods under IGR reactor conditions, calculation method not using FRAPCON-3 code

was selected. For this case, method of initial data generation for FRAPTRAN code is presented in Table 15.

Table 15. Major input data for high-burnup fuel simulation needed for FRAPTRAN code.

Pare eter FRAPTRAN -iFput Variable Sourceofthedata

Cladding design data RodDiameter, roughc, gapthk Post base-irradiation test data of mother

fuel rod

Fuel design data FuelPelDiam, roughf, frden, Post base-irradiation test data of mother

rvoid fuel rod

Volume of upper and lower vplen, volbp Before pulse-irradiation test data

gas plenum

Fraction of gas mixture gfrac Before pulse-irradiation test data

Internal gas pressure gapprO Before pulse-irradiation test data

FGR in transient relfrac User-specified FGR history from FRAPT6 (modified) calculation data

Transient fuel swelling TranFuelSwell User-specified transient fuel swelling

history from FRAP-T6 (modified)

calculation data

Radially average fuel burnup bup Post base-irradiation test data of highbumup fuel

Radial bumup profile butemp User-specified radial bumup profile

from TRIFOB [31] calculation data

Nornalized radial power RadPowProfile User-specified normalized radial power

profile profile from TRIFOB calculation data

Oxide layer thickness odoxid Post base-irradiation test data of mother

fuel rod

Excess hydrogen cexh2a Post base-irradiation test data of mother

concentration fuel rod

Open porosity in fuel OpenPorosityFraction Detemined by FRAPTRAN code

Similar approach was taken for assigning initial data for high-burnup fuel rods H7T and HIT tested in IGR

reactor.

4.1.5. Calculated results of high-burnup fuel simulation obtained with modified FRAPT6, SCANAIR and FRAPTRAN codes

The calculation results for high-burnup fuel rods H5T, H7T, HIT are presented below in Fig. 10-Fig. 21.

The FRAPTRAN (modified) results are compared with results computed by the SCANAIR (modified) and

FRAP-T6 (modified) codes. Considering that the applied SCANAIR version [32] had no model of cladding

deformation of the ballooning type and respective failure models, calculation results obtained by SCANAIR

code are presented to the point of initiation of clad plastic deformation after the re-opening of gas gap.

26
High-burnupfuel simulation of IGR/RL4 test (H5T assessment case)

2.0 4.0 6.0 8.0

Time (s)

a) FRAPTRAN (modified) vs. FRAP-T6 (modified)

2.0 4.0 ; 6.0 8.0

Time (s)

b) FRAPTRAN(modified) vs. SCANAIR (modified)

250

200 _

0

.

CZ

150 

CC_ n\

.Z :

's

100 ='

U

en :

50 w

0

10.0

250

nAA

-

c.V

CO

CZ

150 0

CC

aSC

CO

(U

c

A 0-.

10.0

Fig. 10. Energy characteristics vs. time calculated by modified FRAP-T6, SCANAIR and FRAPTRAN

codes (H5T test case).

27 -

- FRAPTRAN (modified)

-- FRAP-T6 (modified) ...--- 

+*-- Power of fuel rod

S Fuel enthalpy

- Leakage of energy

10

C-

.0

0

.

0.0

500

400

300

0

0

.0

0

IA

100

0.0

_ _ ~~~~~~~~~FRAPTRAN (odified) 

0 ---- -SCANAIR (modified)

_ . . _ + ~~~~~~~~~Power of fuel rd

_ . . . o A Cl~~~~~ Fuel enthalpy_

v .. ._ | | 50~~ Leaka ge of energy 

7(M _ # # ^ _ > _ ZVV - _

1

.

_ >

DU

.
2.0 4.0 6.0 8.0

Time (s)

a) FRAPTRAN (modified) vs. FRAP-T6 (modified)

2.0 4.0 6.0 8.0

Time (s)

36

2

i:

27 _

._

C)

0

0

I18 rCZ

'9 0

0

u

. -

10.0

35

Cl4

E

28 _.

.C

I

CL)

21 o

CZ

14 

C)

-

0

0

*0

-t.0

10.0

b) FRAPTRAN (modified) vs. SCANAIR (modified)

Fig. 11. Cladding temperatures and heat transfer coefficients vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (H5T test case).

28

2000

1500

'w

C)

0

1000

-

500

0

0.0

2500

2000

i

'- 1500

Ca 0.

E 1000

500

0

0.0
.;II

500

-I I 1 I-4nIIA

-'II I ' 

ll,1

am

I..

I*

..,

I.

.,

a

-U--

1

FRAPTRAN (modified)

FRAP-T6 (modified)

Intemal pressure

Clad hoop stress

0.0 -2.0 4.0 .,6.0 8.0 1.

Time (s)

a) FRAPTRAN (modified) vs. FRAP-T6 (modified) * -Mi. ~ ~ RATIANM

_ . , " _ . - - -- - ~~~SCANAIR (modified) _

; . + ~~~Internal pressure

_ . B ~~~~~~~~~~~~~~Clad hoop stress_

,. ~ ~ ~ ~ ~ ~~~~ U

,. ~ ~ ~ ~ ~ ~ ~ ~ ~ 

T_ L _ I 1< -~I 

:2.0 4.0 - 6.0 8.0 10.0

300 ,,,

0.

0

cN

0

200 -

*0

ce

u

100

s

0

500

400

0 .

s 300 -

In

0.

0

200 

-

100

0

Time (s)

b) FRAPTRAN(modified) vs. SCANAIR (modified)

Fig. 12. Cladding hoop stresses and internal pressures vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (H5T test case). I : - .

29

2.5

2.0

1.5

1.0

0.5

_i

,

Ce.

en

-

0.

0.0

2.5

2.0

CE

0.

C)

cC;

c0

1n

ceOD

0

C)

1.5

1.0

0.5

0.0

0.0

. - .

' _

_ :1 .,.s I 

l

-
10

8

6

4

2

0

10

8

6 -

.

0.

0

CZ

U 4

2

0

0.0 2.0 4.0 6.0 8.0

Time (s)

a) FRAPTRAN (modified) vs. FRAP-T6 (modified)

0.0 2.0 4.0 6.0 8.0

- 0.15

- 0.12

_

0.09

0.

o

- 0.06 e

U

- - 0.03

10

10.0

10.(

0.15

0.12

E

0.09 -

0.

0.06 !

0.3U

0.03

0.00

Time (s)

b) FRAPTRAN(modified) vs. SCANAIR (modified)

Fig. 13. Cladding hoop strains and gap widths vs. time calculated by modified FRAP-T6, SCANAIR and

FRAPTRAN codes (H5T test case).

30

eN

0.

0

0

U

_ ( FRAFITRAN (modified) 

- - - -- SCANAIR (modified)

_ + Clad hoop strain __

1B Fuel-clad gap width /

T C _~~~~~~~~~~I,-
High-burn upfuel simulaiion of IGRIRIA test (H7T assessment case)

0 .

I'0.0 2.0 4.0 6.0 8.0 -10

Time (s)

- a) FRAPTRAN (modified) vs. FRAP-T6 (modified)

0.0 2.0 4.0 6.0 8.0

Time (s)

b) FRAPTRAN(modified) vs. SCANAIR (modified)

300

20

-200 t

Ca

100 >s

C4

0

-300

C-

-200 t

c)

4)

-100*- 

r)

-t.0

10.0

Fig. 14. Energy characteristics vs. time calculated by modified FRAP-T6, SCANAIR and FRAPTRAN

codes (H7T test case). .

31

600

.4- I 1 4~~~~~~~~~~~~

- %

400

200

0

..

C-. * 0

S.

4)

0

_~

-n

600

400

200

0

.-

0

-

0

0

a.

1I

0 I

11 ! - .

- 0

- I ,

I

-no-

,0

r~- FRAPTRAN (modified)

--- FRAP-T6 (modified)

+ Powcr of fuel rod

+ Energy dposition

Cl Fuel enthalpy

O Lelcagc of cncrgy

I ''- ,, -IIAs

,. I ...
- FRAPTRAN (modified)

------ FRAP-T6 (modified)

+- Fuel centerline temperature 36

+4- Clad outer temperature

- Heat transfer coefficient

8.0 10.0

Time (s)

a) FRAPTRAN (modified) vs.

36

-8--- :Heat transfer coefficient 

E

2500- *30

2000 24

1500- 18

2 Ca

1000 1

0

0

500 -______ 6 6_____________ 

0 ------------ 0_

0.J 2.0 4.0 6.0 8.0 10.0

Time (s)

b) FRAPTRAN(modified) vs. SCANAIR (modified)

Fig. 15. Cladding temperatures and heat transfer coefficients vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (H7T test case).

32

3000

2500

2000

4)

c 1500

4)

0.

2-

4)

1000

500

0

2.0 4.0 6.0
.1 -

I I

: C,Z

500

-t I~-.- - -- :1- - + AIM

----- I

It

lb11

II

II

I.h

4

I

II

i

I

I

I...U

L+

II

I I 

I U ~ ~ -- -- -

0.0

2.0 4.0 6.0

,~ ~ . _

FRAPTRAN (modified)

FRAP-T6 (modified)

Internal pressure

Clad hoop stress

- .,.

± -t 4

.

8.0 10

Time (s)

a) FRAPTRAN (modified) vs. FRAP-T6 (modified)

2.0

Ca

0-

0

-o,

100 =

500

-00

1

-500

+ + ~~~~~~~~~~~~~i- £111

_~~~~ . . I 

1 _ _

'a

S.

ff~~~~~~~~~~~~~~~~a 4 i~~~~I

4.0 - 6.0 8.0

CZ

0.

0;

F-;n

0

O - c: i

100

0

I . .

10.0

- Time (s)

b) FRAPTRAN(modified) vs. SCANAIR (modified)

Fig. 16. Cladding hoop stresses and internal pressures vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (H7T test case).

33

2.5

2.0

ce

a)

S- 0.

Q,

ce

ce

Ce a1.5

1.0

0.5

0.0

0.0

2.5 -r2.0

EC)

II0S.

U,

Ca

aL. en

ce

1.5

1.0

,' I~~~~~~~~~~~~~ I - FRAPTRAN (modified) 

(modified) -SCANAIR 

+U-- Internal pressure

- Clad hoop stressT

- 0(0

ZUi

0.5

0.0

_ .

p

X.0

I

A
2.0 4.0 6.0 8.0

Time (s)

a) FRAPTRAN (modified) vs. FRAP-T6 (modified)

2.0 4.0 6.0 8.0

Time (s)

b) FRAPTRAN(modified) vs. SCANAIR (modified)

0.5

0.4

E

E

0.3 -

._

c

to

-

.

0.1

I 0.0

10.0

0.24

0.18 _

E

:2E

* .12 m.

C

en

0.06

0.00

10.0

Fig. 17. Cladding hoop strains and gap widths vs. time calculated by modified FRAP-T6, SCANAIR and

FRAPTRAN codes (H7T test case).

34

16

12

.

0.

0

-o

8

4

0

0.0

6

4

2

-

.ea

0.

0

*0

I

u

0

0.0
High-burnup fuel simulation of IGR/RIA test (HIT assessment case)

I ____________ .. 1 FRAPTRAN (modified) -

-----. FRAP-T6 (modified)

--- Power of fuel rod

+- Energy deposition

B-- Fuel enthalpy

---- Leakage of energy

. ol~

- -

-- - .0... 1 LIL\-_______ _______ -- - -

V

I -T i I I -

0.0 2.0 4.0 6.0 8.0 10.0

Time (s)

a) FRAPTRAN (modified) vs. FRAP-T6 (modified)

300

240 -

C.

20

-

C

60 

- c

120 n rJ

0

300

- 24(

_ .,o

U

I I-O_

en

.

Cr

U

a

Co

O -0

en

U

C*

0 

G

I I

0.0 2.0- 4.0 .6.0 - -8.0 10.0

Time (s) ',

b) FRAPTRAN(modified) vs. SCANAIR (modified)

Fig. 18. Energy characteristics vs. time calculated by modified FRAP-T6, SCANAIR and FRAPTRAN

codes (HlT test case).

35

500

400

i

cr0

S.

3

0

a

A.

300

200

100

0

500

400-

300

200

0

I-=

a0

tu

0

- FRAPTRAN (modified).

------ SCANAIR (modified)

I-U-- Power of fuel rod

- Energy deposition

- Fuel enthalpy

Ie- Leakage of energy

100

1-.u

, . ; , , , n

Utr
Fig. 19. Cladding temperatures and beat transfer coefficients vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (HlT test case).

3 6
500

400

Ce

300 ,,

0

0 200 .

100

0

a) FRAPTRAN (modified) vs. FRAP-T6 (modified)

,-B

I.1

I

I"

+ 4 +

-4-

Is

FRAPTRAN (modified)

SCANAIR (modified)

Intemal pressure

Clad hoop stress

r -

1. 1¶i'iL-L p -. _ _I 

I

500

400

r

300

c 

-

Ce.

0.

200 .C

Ca

100

0 

Z.U 14.U b.U OM 10.0

. Time (s)

b) FRAPTRAN(modified) vs. SCANAIR (modified)

Fig. 20. Cladding hoop stresses and internal pressures vs. time calculated by modified FRAP-T6,

SCANAIR and FRAPTRAN codes (HlT test case).

37

C>

0

CZ

:C

-

U,

ce

2.4 -

2.0 -

1.6

2.5

2.0 -

1.5 

1.0 -

s

cli

ce

CZ-

-

Ec.

U,

r0.5

0.0

0.0

I

2 .8

-------- -

I

I

I

I

I

I

IN A n I: n
0.18

2.0 4.0 6.0 8.0

Time (s)

a) FRAPTRAN (modified) vs. FRAP-T6 (modified)

2.0 4.0 6.0 8.0

0.12 

E

.

c03

ci

0.06 u

0.00

10.0

0.18

0.12 c

.0

0.

o

.

0.00

10.0

Time (s)

b) FRAPTRAN(modified) vs. SCANAIR (modified)

Fig. 21. Cladding hoop strains and gap widths vs. time calculated by modified FRAP-T6, SCANAIR and

FRAPTRAN codes (HIT test case).

4.1.6. Discussion of principal thermal parameters calculated by the codes

Comparison of the above results of IGR test calculations of FRAP-T6 and FRAPTRAN codes shows that

38

8

uO

c

~00

~0

6

4

2

0

0.0

10

8

c

0. 0

0.

.0

0

6

4

2

0

0.0

10
peak fuel enthalpy calculated by FRAP-T6 code systematically exceeds the value of peak fuel enthalpy obtained by FRAPTRAN code. 'Performed analysis revealed that the main reason for such discrepancy is quantitative difference of fuel-cladding gap thermal conductivity models, which determine the fuel-to-cladding energy leakage.

Calculations of fuel-cladding gap thermal conductivity can be performed by both FRAP-T6 and

FRAPTRAN codes using Ross-Stoute model. The difference between the models is in the use of different

empiric coefficients (Table 16).

General form of thermal conductivity of fuel-cladding system is given by:

a,P ga, + af 5,id + a,m,d

where a,p, - thermal conductivity through the gas gap between fuel and cladding (W/m2 K);

aSOid - thermal conductivity due to fuel-cladding contact pressure (W/m2 K);

and - thermal conductivity due to radiation (W/m2 K).

Table 16. Comparison of fuel-cladding gap thermal conductivity,models in FRAP-T6 and FRAPTRAN

codes.

i-i-- --4---L; - Model content

fi: y<::^ i. 2 FRL-- P-T6(mod~ed)-: :PTN -- FRA-:- =::-- -- 

A, A,

A +(Rf +R,)+(gf +g¢) A+3.6(Rf +R)+(gf +g,)

A, - gas thernal conductivity (W/m-K);

A - hot fuel-cladding gas gap (m);

Rf - fuel surface roughness (m);

R -'cladding surface roughness (m);

gf,+gc- combined fuel and cladding temperature jump distance (m)

4.5579* 10-3 '-'A 0.4166 Km Pwi-/ (RE), if P.i < 9106

R-E

A = coefficient dependent upon Pm1 0.00125Km / (R-E), if 0.003 > P., > 9104

A-0.01, if 0.01 P,, 0.0001 0 . 4 1 6 6 K P ER, ,,ifP>0 003

A=lJifPn1<0.000lorPmi>0.0l, R-E

P= = ratio of interfacial pressure to cladding Meyer R..k = 3 3 3

.

3 -P , if P,1 < 0.0087

hardness - R.,,i = 2.9, if P,, > 0.0087

i upon ,P,c exponent . dependent = ratio of interfacial pressure to cladding Meyer n = exponent dependent upon P1 - hardness

afiid n=0, if 0.01I < P. 1 0.0001 K, = mean thermal conductivity of fuel and cladding n=-0.5, if P,1 < 0.0001 (/iK

n=lI,ifP,,,>0.0l (WnK

Km = mean thermal conductivity of fuel and cladding Km = 2 K(Kr+K,)

(Wi-K)K.=2fK Kf+)

K. = 2Kf Kc I(Kf +Kc) where Kf and Kc are the fuel where Kf and K, are the fuel and cladding thermal conand cladding thermal conductivities, respectively, ductivities, respectively, evaluated at their respective

evaluated at their respective surface temperatures surface temperatures

R= (Rf + R,2)

0

" where Rf and R& are the fuel and R = (R2 + R) 0 where R and Rc are the fuel and cladcladding surface roughness, respectively (m) ding surface roughness, respectively (m)

E = exp[5.738 - 0.528.ln(3.937-10 7)] E = exp[5.738 - 0.528-1n(Rra)] where a = 3.937-io7 gm

(Tf2+TC2) (Tt+Tc)FeF.CY

a = Stefan-Boltzmann constant = 5.6697-104 W/mkK4

a F. = einissivity factor determined by MATPRO

F. = configuration factor = 1.0

Tf = temperature of fuel outer surface (K)

T, = temperature of cladding inner surface (K)

39

I 
To verify the influence of conductivity models on thermal parameters of fuel rods, gap conductivity model

from FRAP-T6/VVER was incorporated into the FRAPTRAN code. Then comparative analysis of fuel temperature, cladding temperature and peak fuel enthalpy was performed by the example of fuel rod H5T. The

effect of gap thermal conductivity is illustrated in Fig. 22. Note that conductivity of the gap (contact) calculated using FRAPTRAN (original) model exceed values of conductivity obtained using FRAP-T6 model up to

the 4 second of the process. Then, re-opening of the gap occurs, and influence of conductivity effect from this

point on is insignificant.

40000 

,h FRA-T6

-t- FRAPTRANoriginal model

30000 3 s -70- FR APTRAN with FRW-T6 model

nnnn 0 .10

0

0.

i nnnn./4 IG '. iuvuu - - __ _ X 

0 2 4 6 8 10

Time (s)

Fig. 22. Comparison of gap thermal conductivity models from FRAP-T6 (modified) and FRAPTRAN

(modified) and FRAPTRAN (original).

Thus, gas component of fuel-cladding gap thermal conductivity used in FRAPTRAN code gives higher values

as compared with FRAP-T6 code. This results in increase of fuel-to-cladding heat flow (energy leakage). Due

to this, fuel temperature (peak fuel enthalpy) goes slightly down as compared with calculated temperature

obtained using FRAP-T6 code. Level of peak fuel enthalpy decrease due to the usage of two conductivity

models is shown in Fig. 23.

Fig. 24 displays comparison of calculated results obtained using FRAP-T6 and FRAPTRAN codes, which has

the same conductivity model from FRAP-T6. It was shown that in this case, heat flow and energy leakage

from the fuel rod go down, and therefore, peak fuel enthalpy goes up.

40

I '
-4~~~~~ Al l.----I---- 4

''a. Ii, ii,

0.0 2.0 4.0

- - FRAPTRAN (modified)

---.-- FRAP-T6 (modified)

+U- Power of fuel rod

-B Fuel enthalpy

Oe- . Leakage of energy

I,-e

6.0

1 250

- rnn .-.

150- 0.

9

ea

.-c

C

100 n

>0

to

50 w

0 

8.0 o.a

Time (s)

Fig. 23. Peak enthalpy calculated by FRAP-T6 (modified) and FRAPTRAN (modified).

Fig. 24. Peak enthalpy calculated by FRAP-T6 (modified) and FRAPTRAN (modified) with gap thermal

conductivity model from FRAP-T6 (modified).

Moreover, by the example of calculation results for test rods with fresh fuel 50F 6 (see Fig. 6) and 96F09 (see

Fig. 8), it was shown that influence of fuel-cladding gap thermal conductivity on cladding temperature is significantly-higher. Increase of-maximum cladding temperature in the case of modeling using modified

FRAPTRAN code is 80 K for fuel rod 50F16 and 120K for fuel rod 96F09 as compared with results obtained

41

400

~i

*0

0

4.)

0

300

200

100

I 0 

=1

Co

.

C

~0

la

C.)

Co

T

--- l

L . .

500

111--
by the modified FRAP-T6 code.

In the future, selection of either model of gap conductivity under conditions of high-temperature contact of

fuel and cladding (especially for high-burnup fuel rods) will have to be confirmed by appropriate

experimental data. In other words, empirical coefficients have to be selected on the basis of experimental

data, which correspond to the given type and state of the fuel rod, as well as taking into consideration

particular conditions of its operation.

4.1.7. Summary of the assessment of the FRAPTRAN (modified) against RIAsimulated test data

Comparative analyses of experimental and predicted cladding temperatures were carried out in the energy

deposition range of 120-400 cal/g. Assessment cases of fresh fuel rods were compared with in-pile test

measurements and calculated results of cladding temperature at rod outer surface. The comparison of

FRAPTRAN (modified), SCANAIR (modified) and FRAP-T6 (modified) predictions with test data

obtained from IGR/RIA experiments show that (see Table 17):

* The predicted and measured cladding temperature histories are in satisfactory agreement. However, it

is necessary to mention that cladding temperatures predicted with FRAPTRAN (modified) are higher

than computed with FRAP-T6 (modified) and SCANAIR (modified).

* Comparison of fuel and cladding behavioral models was performed in FRAP-T6 and FRAPTRAN. As

a result it was shown that decrease in peak fuel enthalpy in the assessment cases under consideration

resulted from differences in empirical coefficients in fuel - cladding gap thermal conductivity models

of the FRAP-T6 and FRAPTRAN.

* Simulation of high-bumup fuel rod behavior with FRAPTRAN (modified) showed that the main

thermal and mechanical parameters of fuel rods HIT, H5T, H7T are similar to previous calculation

with FRAP-T6 (modified) and SCANAIR (modified). However the overprediction of the cladding

temperature was obtained with FRAPTRAN (modified), as well as for fresh fuel rods.

* FRAPTRAN (modified) has correctly predicted failure of cladding of rods H5T, H7T and preservation

of integrity of test rod HIT. No systematic differences of FRAPTRAN predictions of maximum hoop

strain of the cladding from the calculated data obtained earlier have been detected. Certain discrepancy

of data on defornation can result from correlation differences of mechanical properties in the updated

version of MATPRO, as well as from differences in cladding temperature history predictions.

* In modeling high-bumup fuel rods, the problems with assignment of initial data on radial-axial burnup

distribution (BUTEMP), time dependent transient fuel swelling (TRANFUELSWELL) and time

dependent transient fission gas release (RELFRAC) were identified and resolved.

• In calculating fuel rods behavior in the power pulse conditions, difficulties with solving mechanical

problem in FRAPTRAN code at re-opening of fuel-cladding gap were identified. At that, the problems

could not be eliminated through mere decrease of time step.

Table 17. Comparison of main thermal and mechanical results predicted by the modified FRAP-T6,

SCANAIR and FRAPTRAN codes [3].

Parameter Unit Experinent SCANAIR lFRAP-T6 FRAPRAN Parameter | Vmt | | ~~~~~~(modified) |(modiied) |(modiied)

Peak fuel enthalpy Cal/g 176 172

Peak fuel temperature K 2817 2834

Peak clad temperature K 1224 1220

42
SCANAIR FRAP-T6 FRAPTRAN

Parameter Unit Experiment (modified) (modified) (modified)

Cladding failure yes - yes yes

Residual clad hoop strain 6.5 . 7.28 5.75

(ballooning region) - . .

.~HI(47. 41kk:S <W

Peak fuel enthalpy Cal/g - 187 172

Peak fuel temperature K . 2876 2797

Peak clad temperature K 1231 1222

Cladding failure yes m yes yes

Residual clad hoop strain % 10.1 11.9 7.71

(ballooning region) _ _ _ .. _ _

HIT (49.2 M Wd/g U) _ _ _ _ _ _ _ _ _

Peak fuel enthalpy Cal/g - 147 151 146

Peak fuel temperature K . 2601 2681 2676

Peak clad temperature -. K . 1192 . 1148 1207.

Cladding failure no - no no

Residual clad hoop strain* . 1.4: 1.83 2.82 2.94

Maximum clad hoop strain" % 2.24 3.43 3.58

4.2. RIARILOCA assessment

4.2.1. Description of experiments modeling the first stage of LOCA

Special facility with heating of a fuel rod by means of direct transnission of electric current through

cladding (Fig. 25) has been developed at RIAR for testing simulators of unirradiated and irradiated fuel rods

under conditions, which model initial stages of LOCA [29]. This facility permitted to conduct experiments

necessary for verifying that behavior of cladding of real fuel rod with complicated geometry, which is

subjected to a complex set of loading factors, would correlate with the results achieved in significantly less

,complicated biaxial burst tests.

43

Local clad hoop strain along fuel rod height

Axially averaged clad hoop strain
Fig. 25. Schematic diagram of electrically heated facility for studying the first stage of LOCA.

Scenarios of temperature change and pressure drop

(RLAR-LOCA2), are shown in Fig. 26 [29].

?2*_ I I -I '_ ~i

UU( 

0

900

600

300

0

at the cladding of fuel rod simulator tested at RIAR

0.

0.

0

-o

0

:nI

VZ

5 10 15 20 25 30

Time (s)

6 _

4 

2 _ -

0

-2-

-4-- I--

0 5 10 15 20

Time (s)

25 30

Fig. 26. Simulator cladding temperature and cladding pressure drop.

Fuel rod simulator consisted of cladding (Zr-1%Nb) and fuel stack. Geometry of A12 03 fuel pellets was the

same as for VVER- 1000 fuel. Simulator was filled with argon under constant pressure, which was maintained

at the set level with the help of pressure stabilizer. Required scenario of pressure drop change at the cladding

was ensured by means of coolant (argon) pressure variation.

In the course of each test, the following parameters were measured:

* coolant pressure;

* pressure in the simulator,

I2

1'

I Zuu

/

0.

S

I 7I I___

I -a

7l

* cladding temperature. 44
In addition, during the stage of post-test examinations hoop strain versus axial length and parameters of burst

region of the cladding were measured. Appearance and cross-section of the simulator after the test are shown

in Fig. 27 [29].

Li . .

.Fig. 27. Cladding appearance and cross-section in burst area.

4.2.2. Calculation procedure and main results

FRAPTRAN calculations of these tests have been conducted with the purpose to verify the set of models,

which describe ballooning and burst of Zr-1%Nb cladding. The calculations were conducted using boundary

conditions of 1V type, i.e. using temperature versus time dependencies (as measured). Pressure drop at the

cladding was assigned in the similar fashion (see Fig. 26). At that, in the last case linear extrapolation of the

assigned pressure drop was conducted to the point of predicted burst of the cladding.

Main parameters of the simulators and initial data RIAR-LOCA2 test calculations are given in Table 18.

Table 18. Main parameters of simulator and initial data for LOCA calculations.

. ;PARAMETER-8y' , ' ' - ' :, UNITS

Cladding Zr- I %Nb

; Fuel A12 03

Coolant Argon

Cladding outside diameter mm 9.1

Cladding inside diameter mm 7.6

* Cladding thickness :mm 0.7

Radial gap thickness mm 0.03

Samples height m 0.2

Results of cladding deformation and time to failure calculations are shown in Fig. 28, Fig. 29. Preliminary

analysis- of these results shows that predictions of cladding burst pressure - temperature and time to rupture

were satisfactory. Calculated residual hoop strain of the cladding (43%) is also close to the experimental value

(54%).

45

I
Fig. 28. Comparison of calculated and experimental values of failure pressure and temperature.

Fig. 29. Comparison of calculated and experimental values of cladding residual strain.

4.2.3. Summary of the assessment of the FRAPTRAN (modified) against LOCAsimulated test data

Comparison of experimental and calculated data in terms of parameters (pressure - temperature) of the fuel

rod simulator failure is shown in Table 19.

46

12 - 1400

10 ..--. Clad temperature - -

/-\ :::-. ; IMeasured pressure drop-.

8 \ _ -: = Calculated pressure drop 1200 - .1200, >R .

6 _-o 

-2 /-/ \ _ I

40 600

5 10 15 20 25 30 1 

Time (s)

60 1400

40 1200

.20 1000 

o 0

6

-o

0 - .00

-0- Cla feeratur

;.-;:.: *1 -1 Measured Measurd- ~~~clad hoop strain-

~-! - . Calculated clad hoop strain

-20 _ 600

0 5 10 15 20 25 30

Time (s)
Table 19. Comparison of experimental and calculated data on high-temperature cladding deformation

under conditions modeling the first stage of LOCA.:

Praee RIARLOCA2

Waepermeterc

Cladding rupture Yes Yes

Pressure drop during cladding failure (depressurization) (MPa) 3.88 4.57

Cladding temperature at the time of failure (depressurization) (K) 109 1089

Time to failure (depressurization) (s) 16.5 17.8

Maximum hoop strain (%) 54.4 42.9

As is seen from Table 19, these first calculations of behavior of the fuel rod simulator with unirradiated Zr1 %oNb cladding in the test, which imitates thermal and mechanical impact under LOCA conditions, revealed

*satisfactory iiatch of calculated and experimental'strength paramneters of cladding rupture'. Cladding failure

crterion in the form of hoo stress'at burst, whiich was implemented in MATPRO (see section 2.2), led to

obtaining of reliable predictions of the cladding rupture time. At the same time, some underprediction of

circumferencial strain at burst was obtained. Most likely, this is a reflection of a known problem of

predicting maximum cladding deformation at burst. High sensitivity of 'this paramneter. to the loading

conditions, which is revealed through wide'dispersion of out-of-pile ad in-pile test data, causes serious

difficulties both in developing criterial correlations and in calculation modeling of cladding deformation

process. Therefore, performance capabilities of the BALQN2 subcode, which is used in FRAPTRAN for

calculating local ballooning and rupture of the cladding, sould be analyzed ctically. Such' analysis

presents a separate task for the future and it should be accompanied by attempts to understand all

accumulated test database on circumferencial elongation of cladding under LOCA conditions.,

Time calculated from the beginning of the test to the point of cladding failure (depressurization).

4 7
5. CONCLUSIONS

1. This study is a continuation of the series of works conducted at RRC KI in relation to adaptation of

transient fuel rod behavior codes to analysis of behavior of fuel with Zr-Nb cladding. Due to its high

operational performance, Zr-1%Nb (E110) cladding is considered as one of the most realistic

candidates for high burnup fuel. Therefore, expansion of the area of application of these codes in the

direction of safety analysis of fuel with alternative cladding (with respect to Zircaloy cladding) was

the main objective of such activity. Accumulated earlier experience in modification and adaptation of

NRC's FRAP-T6 and French IRSN's SCANAIR codes to behavioral analysis of fuel of Russian

pressurized water reactor of VVER type under conditions of Reactivity Initiated Accident was the

basis for conducting appropriate activities using new FRAPTRAN code developed by request of

NRC.

2. As a result of intensive program on studying mechanical properties of unirradiated and highly

irradiated Zr-1%Nb cladding initiated in 1996, expanded MATPRO package of material properties

was developed for behavioral analysis of fuel with niobium cladding in joint use with thermal

mechanical codes. Last updated version of the package presented in this report accumulated all testing

and analytical results obtained to this date.

3. In addition to the development of expanded package of material properties, work on modification and

adaptation of FRAPTRAN code was conducted in the scope necessary to perform calculations of

selected RIA and LOCA assessment cases with fuel of VVER type. All models modified earlier for

FRAP-T6 version designed for calculations of high bumup VVER fuel, have been transferred to

FRAPTRAN. Significant scope of coding work due to introduction of new global and local variables

as well as to correction of some inaccuracies and errors in the original code has been performed.

4. Three IGR power pulse tests with unirradiated rods (for which experimental records of cladding

temperature and internal pressure are available) and three tests with high burnup commercial fuel

have been modeled using modified FRAPTRAN code. Obtained results led to the following

conclusions:

* Main thermal and mechanical parameters predicted by the code are in reasonable compliance with

experimental data obtained with instrumented fresh fuel in wide range of peak enthalpy of the fuel

and corresponding range of clad-to-coolant heat transfer modes.

* Main thermal and mechanical parameters of high burnup fuel rods HIT, H5T, H7T predicted by

FRAPTRAN are also in satisfactory compliance with predictions of modified FRAP-T6 and

SCANAIR codes obtained earlier and available post-test data.

* Certain systematic underestimation of peak enthalpy and overestimation of cladding temperature by

the FRAPTRAN code has been observed as compared with previous data of modified FRAP-T6.

Performed analysis revealed that these differences could be explained by different empirical

coefficients in fuel-cladding gap thernal conductance models used in these codes.

5. Calculations of the selected LOCA assessment cases have been conducted using the modified

FRAPTRAN code. Comprehensive out-of-pile test of the fuel rod simulator with unirradiated Zr1%Nb cladding, during which thermal and mechanical cladding impacts simulated conditions of

initial stage of LOCA, has been performed. Good agreement between calculated and experimental

values was achieved for such parameters as pressure-temperature at burst, and time to failure.

Satisfactory predictions of maximum hoop strain at burst (43% versus actual 55%) were obtained.

6. Very limited LOCA assessment does not allow for unambiguous conclusion about quality of

prediction of maximum hoop strain at burst. However, taking into consideration known contradictions

in the accumulated database on circumferential elongation of cladding under LOCA conditions and

difficulties in modeling of ballooning, it should be mentioned that critical analysis of the entire

existing test database needs to be performed and new approaches to prediction of maximum hoop

strain need to be considered.

48
7. In general, based on experience with the above codes accumulated at RRC KI some general

observations can be made in relation to the first version of FRAPTRAN, which was modified for

high-bumup VVER fuel:

* the code is a new important step in development of the FRAP-T family codes. The issued first

version of the FRAPTRAN code can serve as a basis for achieving up-to-date level of high-burnup

fuel behavior modeling;

* from the user's standpoint, advantages of the new code include friendly service and convenient postprocessing of the output data. Ease of using FORTRAN source and transparent procedure for

preparation of initial data should also be noted;

* disadvantages of the analytical model include the requirement for the user to assign time step and

unavailability of automatic selection of integration step in the process of calculation, which makes

use of the code more difficult and time consuming;

* check of mechanical model of the code in analyzing behavior of the fuel with Zr-Nb cladding is of an

interest for continuation of the code validation. For instance, results of the recent tests of VVER fuel

(burnup 50-60 MWdlkg U) under conditions of narrow power pulses at BIGR reactor [1, 2] can be

used for this purpose. Potential value of the results of IGR pulse tests with initial negative pressure

drop on the cladding (16 MPa) should also be noted [2], since these initial conditions are more

prototypical for full-scale reactor case.

8. From the standpoint of further enhancement of FRAPTRAN models of high-burnup fuel behavior,

the following tasks are believed by the authors of the report to be a priority:

* re-assessment of boiling curve as applied to transient conditions of clad-to-coolant heat transfer

typical for RIA;

* development of models of fission gas release, transient fuel swelling and formation of heterogeneous

fuel structure (density, grain size) along the pellet radius;

* development of the model of mechanical interaction of fuel and cladding with consideration for

irregularity.of temperature fields and material properties.

9. The following priority tasks in development of models of Zr-l%Nb cladding material properties,

implemented in MATPRO package should be considered:

* obtaining of continuous dependence of cladding mechanical properties vs. burnup;

* expansion of the range of studied burnup to 60 MWd/kg U and more;

* taking into account the effect of heating and cooling rates on phase transition temperatures and.

therefore on temperature'dependencies of mechanical properties;

* adjustment of obtained correlations to account for annealing of radiation damages as a function of

heating rate;.

* development of approaches to possible extrapolation of obtained stress-strain data beyond the

uniform elongation for adequate prediction of cladding deformation up to the failure.

49
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1997.

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Moscow, Energoizdat, 1981 (rus). 5 0
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[22] Yu.K. Bibilashvili, N.B. Sokolov, L.N. Andreyeva-Andrievskaya, A.V. Salatov, A.M. Morozov, Hightemperature Interaction of Fuel Rod Cladding Material ( Zr-1%Nb alloy)- with Oxygen-containing

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1997.

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[33] Yu.K.Bibilashvili, L.A.Yegorova, O.A.Nechaeva, I.G. Smirnov, V.P.Smimov et.al., "Experimental

Study of VVER High Bumup Fuel Rods at the BIGR Reactor under Narrow Pulse Conditions."

Intemational Topical Meeting on Light Water Reactor Fuel Performance, Park City, Utah, April 10-13,

2000, pp.306-314.

[34] Yu.Bibilashvili, A.Goryachev, O.Nechaeva, A.Salatov, V.Sazhnov, I.Smirnov, N.Sokolov, Yu.Trutnev,

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part), Aix-en-Provance, France, May 13, 2002. 5 1
NRC FORM 335 US. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

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

NRCM 1102, and Addendum Numbers, If any.)

3201,3202 BIBLIOGRAPHIC DATA SHEET

(See instruxtons on the reverse) NUREG/IA-0209

2. TITLE AND SUBTITLE NSI RRC KI 3067

IRSN-2002/33

Adaptation of USNRC's FRAPTRAN and IRSN's SCANAIR Transient Codes and Updating of

MATPRO Package for Modeling of LOCA and RIA Validation Cases with Zr-1 % Nb (VVER 3. DATE REPORT PUBLISHED

type) Cladding MONTH YEAR

April 2003

4. FIN OR GRANT NUMBER

Y6723

5. AUTHOR(S) 6. TYPE OF REPORT

A. Shestopalov, K. Lioutov, L. Yegorova

7. PERIOD COVERED (Inclusive Dates)

2001-2002

8. PERFORMING ORGANIZATION -NAME AND ADDRESS (If NRC, prvideDivision, office orRegion, US. NucearRegulatory Commission andmaifing address; if contractor,

provide name and mailing address.)

Russian Research Center - Kurchatov Institute

Moscow 123182

Russia

9. SPONSORING ORGANIZATION - NAME AND ADDRESS (IfNRC, tpe Sameas above; if contractor, provide NRC Division, Office orRegion, US. NuclearRegulatory Commission.

and mailing address.)

Division of Systems Analysis and Regulatory Effectiveness

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-001

10. SUPPLEMENTARY NOTES

11. ABSTRACT (200 wvrds or less)

The report presents the results of analytical work on updating and supplementing Zr-I %Nb (El 10) cladding material property

models as part of MATPRO package intended for joint use with thermal mechanical codes analyzing high bumup fuel transient

behavior. The scope of work also included adapting U.S.NRC's FRAPTRAN code to the behavior analysis of fuel with El 10

cladding (ER type) and carrying out of the code assessment using selected experimental data. Satisfactory compliance of

the results calculated by modified FRAPTRAN version with in-pile RIA and out-of-pile LOCA simulated test results is obtained.

The obtained calculation data have been also compared with the predictions of the FRAP-T6 (U.S.NRC) and SCANAIR (IRSN,

France) codes modified earlier.

12. KEY WORDSIDESCRIPTORS (List words orphrases that wil assistresearchers in locating the report.) 13. AVAILABILITY STATEMENT

reactivity initiated accident unlimited

cladding-to-coolant heat transfer model 14. SECURITY CLASSIFICATION

SCANAIR computer code (This Page)

material properties unclassified

cladding strain and ballooning code assessment (This Report)

unclassified

15. NUMBER OF PAGES

16. PRICE

NRC FORM 335 (2-89) This forrn was electronically proctucect by Elit Fecteral Forms, Irc.
Jon recycled% R peral p

Federal Recycling Program
NUREG/IA-0209 ADAPTATION OF USNRC'S FRAPTRAN AND IRSN'S SCANAIR TRANSIENT CODES

AND UPDATING OF MATPRO PACKAGE FOR MODELING OF LOCA

AND RIA VALIDATION CASES WITH ZR-1%NB (VVER TYPE) CLADDING

UNITED STATES

NUCLEAR REGULATORY COMMISSION

WASHINGTON, DC 20555-0001

OFFICIAL BUSINESS

APRIL 2003