2

Table of contents

1. Introduction .......................................................................................................... 3

2. AREVA idealized case. ........................................................................................ 5

2.1. Irradiation conditions ...................................................................................... 5

2.2. Simulation by means of the code TRANSURANUS ....................................... 6

2.2.1 Densification of fuel ................................................................................. 6

2.2.2 Swelling of fuel......................................................................................... 7

2.2.3 Fuel relocation ......................................................................................... 7

2.2.4 Cladding irradiation growth ...................................................................... 8

2.2.5 Fission gas release .................................................................................. 8

2.2.6 Cladding corrosion ................................................................................... 9

2.2.7 Recent extensions of the TRANSURANUS fuel performance code....... 10

2.3. Results and discussion ................................................................................. 11 3. US-PWR 16x16 LTA Extended Burnup Demonstration Program ....................... 13

3.1. Irradiation conditions and modeling .............................................................. 15

3.2. Results and comparison ............................................................................... 16

3.2.1 Rod burnup. ........................................................................................... 16

3.2.2 Cladding dimension change .................................................................. 17

3.2.3 Cladding corrosion. ................................................................................ 18

3.2.4 Fuel densification and swelling .............................................................. 20

3.3. Summary ...................................................................................................... 21

4. GINNA reactor experiment. ................................................................................ 22

4.1. Irradiation conditions and simulation ............................................................ 22

4.2. Results of TRANSURANUS simulation ........................................................ 25

4.2.1 Fuel Central Temperature and Fission Gas Release ............................. 25

4.3. Cladding creep down.................................................................................... 27

4.4. External Oxide thickness .............................................................................. 28

4.5. Pellet morphology ........................................................................................ 30

4.6. Summary ...................................................................................................... 33 5. Conclusions ........................................................................................................ 34

References ................................................................................................................ 35

3

1. Introduction

The present final report presents a part of the research work, done in the

Research Contract RC 15164, at the Institute for Nuclear Research and Nuclear

Energy (INRNE), Sofia in the frame of the CRP “Improvement of Computer codes

used for Fuel Behaviour Simulation-FUMEX-III” supported by IAEA.

According to the working program of the project the INRNE had to analyse the fuel

performance of WWER rods included into the CRP FUMEX-III as well as the Gd

doped fuel and the transient experiment Riso3, (rods II5 and GE7) and Kola3-MIR.

The rod H09 from experiment OSIRIS irradiated for 4 cycles in the EDF Cruas PWR

to a final burnup of 46 MWd/kgU was simulated too. The US-PWR 16x16 LTA

Extended Burnup Demonstration Program and the Siemens Corporation 14x14 lead

fuel assemblies, irradiated in the Ginna PWR for 5 cycles up to 58 MWD/kgU were

analysed too. The goal of these two programs is to demonstrate and evaluate the

potential of annular pellets and barrier cladding for resisting fuel failures due to pellet-

cladding interaction up to burnups of 60 MWd/kgU.

During the first year of the contract the team has worked on MIR ramp tests on Kola3

rods (Kola3-MIR experiment) by using the latest TRANSURANUS-WWER version

v1m1j09 [1] on the basis of the IFPE-OECD/ IAEA-NEA database [2]. Fuel rods

included in the tests have been operated under normal conditions at Kola NPP up to

maximum burnup of about 50 – 60 MWd/kgU. Nine re-fabricated rods have been cut

from selected FA‟s and carried out under single ramp conditions (RAMP test) and 2

step-by-step power increase tests with instrumented rods.

The attention was concentrated on:

Fuel central temperature during the ramp irradiation (FGR2-test).

Comparison between TRANSURANUS assessments and thermocouple records;

Pin pressure (FGR1-test). In-pile pressure measurements for two rod

with different accumulated burnup during base irradiation;

Fission gas release and gas mixture (from PIE);

Mcrostructure changes of the fuel-central hole closing and porosity

distribution.

4

The calculations were done for base irradiation as well as for the test irradiation,

using restart option of TRANSURANUS code for account rod cutting and refilling.

According to the working program of the project for the second year of contract

INRNE had to simulate and analyse the fuel performance of Gd doped fuel rods

(Experiment GAIN), PCMI cases (Riso3- GE7; OSIRIS-2, H09 ) as well as the

transient experiment Riso3, rod II5, included into the CRP FUMEX-III. The fuel rods

included in the tests have been operated under normal conditions up to average

burnup of about 40MWd/kgU (GAIN rods and Riso3-rods) and local burnup of about

50 MWd/kgU in OSIRIS-2 experiment. One rod Riso3-II5 rod was re-fabricated,

instrumented with thermocouple and pressure transducer, refilled up to 0.64MPa and

carried out under ramp conditions.

Main results of the study was:

Two rods with 3% and 7% Gd2O3 doped fuel have been simulated and

analysed. The fuel stack length changes and the axial distribution of the

cladding diameter as well as the fission gas release in the case of Gd doped

fuel were predicted with reasonable agreement.

The pin geometry changes (length and diameter) in the PCMI conditions and

outer clad oxide layer were simulated and compared.

The radial distribution of the retained Xe and other products after irradiation

(EPMA and micro gamma scanning data) were reproduced by

TRANSURANUS code and comparison with measured data revealed good

agreement.

The TRANSURANUS assessments of the fuel central temperature during

the ramp irradiation (Riso3-II5 irradiation) is very close to the thermocouple

data.

The present final report presents the TRANSURANUS prediction of the fuel&cladding

properties for the two USA experiments with different pellet design as well as the

idealized AREVA case from the FUMEX-III priority cases programme.

5

2. AREVA idealized case.

According to the working program, INRNE had to analyze an idealized case (AREVA

case) to illustrate the functional dependence of fission gas release predictions up to

high burnup conditions. In the present material, the TRANSURANUS analysis of the

AREVA idealised case is reported.

2.1. Irradiation conditions

The AREVA idealized case comprises data on linear heat rate and fast neutron flux

for 14 axial nodes and data of coolant temperature, coolant pressure and coolant

mass flow rate for every time step. The linear heat generation rates are defined in 54

steps, (entire time of irradiation is 2142 days) and ranged between 10 and 25 kW/m

without sharp transitions. Fast neutron flux is proportional to LHR with coefficient of

6.3*1012 n/(cm2s)(per kW/m).

Table 1: Input parameters used for definition of the initial state of the fuel rod in the simulated idealized case.

Fuel rod AREVA

Fuel stack length mm 3650

Number of axial slices 14

Plenum length mm 150

Eff. plenum volume cm3 8.06

Total free volume cm3 19.3

Fill gas He

Initial pressure (20 OC) MPa 1.6

Mean diam. gap µm 165

Fuel pellet

Dishing mm3 12.52

Surface roughness µm 1.0

Initial porosity % 5

Open porosity % 0.02

Porosity at end of

densification

% 3.98

Grain diameter (3D) µm 16.5

Initial content of 235U wt.% 4.5

Cladding material Zr-4

Outer diameter mm 9.5

6

Inner diameter mm 8.25

Surface roughness µm 1.0

Reactor type PWR

Fast neutron flux

(per linear heat rate)

n/(cm2s)

(per kW/m)

6.3*1012

Coolant pressure MPa 15.5

Distance between 2 rods. mm 12.6

2.2. Simulation by means of the code TRANSURANUS

The latest TRANSURANUS version v1m2j11 has been used to run the idealized

case AREVA – seven cycles of irradiation under normal operational conditions up

to.the rod average burnup of 80 MWD/kgU. The calculation has been performed by

standard TRANSURANUS models and options. The most important of these models

are briefly described below:

2.2.1 Densification of fuel

The fuel densification after start of irradiation and fuel swelling are two independent

physical processes that cause the density change. The fuel densification depends on

the fuel temperatures and started with irradiation and process stop up to burnup of

10MWd/kgU, depending of loaded power (temperature). The fission product swelling

of the fuel depends on burnup.

TU option IDENSI=2 starts an empirical densification model where the sinterable

porosity ∆P (i.e.the maximum densification) is derived from the initial grain size d

(µm) and the fractional fabrication porosity P0 [3].

∆P=P0 (2.23/d)

By selecting the option IDENSI=7, two MATPRO models FUDENS and FHOTPS [4]

are started and these correlations take into account both thermal and irradiation

induced densification as well as sintering contribution due the fuel hot pressure. This

MATPRO LWR model requires the average fabrication porosity of the fuel, the

maximum density change determined by a re-sintering test of 24 hours and the fuel

fabrication sintering temperature.

The simulation of the idealised AREVA case was performed by using the standard

UO2 fuel densification, maximum densification depends on grain size (grain size =

11μm).by simple empirical correlation.

7

2.2.2 Swelling of fuel

The standard swelling option (ModFuel(4)=20) uses the original MATPRO-11 model

[4], adopted for the TRANSURANUS code system by K. Lassmann. The standard

TRANSURANUS option for UO2 swelling treatment is the modified MATPRO LWR

model, which considers both the contributions due to gaseous fission products

(gaseous swelling) and to solid fission products (solid swelling).

According to this approach, the solid swelling is simply proportional to the Bu:

solidV b bu where b = 0.06%/MWd/kgU is model parameter, depending on

the density of heavy metals in the fuel.

As concerns the gaseous swelling, TRANSURANUS calculates the fuel volume

increment due to fission gases through a temperature and burn-up dependent simple

formula - (∆V/V)gas = c a(T)/k (1-e-kBu)

The second tested fuel swelling model is selected by ModFuel(4)=18. It is simple

empirical correlation that gives the total swelling rate for LWR, including swelling due

to solid and gaseous fission products. The increment of fractional volume is

proportional to the increase of the burnup.

(∆V/V) = S∆Bu where ∆Bu is the burnup increment during time step ∆t = tn-1-tn

S = 7.0x10-4 per MWd/kgHM is the swelling rate

2.2.3 Fuel relocation

Pellet cracking already occurs at start up due to the difference in the thermal

expansion of hot pellet centre and cold periphery. Pellet fragments moves outwards

because of the fuel rod vibration induced by the coolant motion. This pellet

“relocation” has a strong impact on the thermal behaviour. It reduces the pellet-

cladding gap size, thereby reducing the temperature levels in the fuel at the

beginning-of-life (BOL). Relocation of the fuel fragments induces the largest

contribution to the gap closure (approximately 30-50%). Because of the stochastic

nature of the cracking the relocation is subject of the largest uncertainty by gap

closure determining.

The recommended fuel relocation model (code option ireloc=8) is modified

FRAPCON 3 model, that assure strain increment due to relocation with linear heat

rating increasing.

8

The second model applied (ireloc=5) treated the relocation of fuel as a single (lump)

fuel volume increasing [5]. It is a simple model that calculates an equivalent

deformation, which increases the fuel radius and height just after going on power

2.2.4 Cladding irradiation growth

The standard for TRANSURANUS cladding swelling model (ModClad(4)=20)

calculates the irradiation growth of Zircaloy-4 cladding (in an annealed state)

according to correlation that determined strains in radial, tangential and axial

direction with specified texture coefficients [6]. The strains due to irradiation growth

are given as a function of the fast neutron fluence. Second tested correlation

(ModClad(4)=18) calculated the swelling of stress relieved Zircaloy-4 cladding [5].

Only axial component is defined and depend on the fast neutron fluence. Radial and

tangential components are set to zero.

2.2.5 Fission gas release

The standard approach for modelling fission gas release in the fuel performance

code TRANSURANUS can be described by three mechanisms that are treated in

parallel [7] [8]:

a) thermal release

Gaseous fission products are generated in the grains, migrate to the grain

boundaries and are released along the grain boundaries when these become

saturated. This leads to a concentration gradient towards the grain boundaries and a

diffusion process to the grain boundaries of the fission products are started. It is

modelled, applying an effective diffusion coefficient that depends on the local

temperature [9]

The concentration of the fission products at the grain boundaries is assumed to be

limited by a saturation value. When exceeding this saturation limit, the supplementary

fission gas reaching the grain boundaries is released to the free volume. The TU

code offers different values for the saturation limit (option igrbdm).

igrbdm=1 -- cgb ≤ 1x10-4 µmol/mm2 the limit of saturation coefficient is constant.

igrbdm=2 –Saturation concentration coefficient cgb depends on the fuel temperature

as 1/T.

9

igrbdm=3 - burst release of the gas grain boundary inventory due to micro-cracking

[10]. The complete release from grain boundaries is considered when thresholds for

local temperature (Tloc >1500(1-buloc/80)) and change of LHR

(Δq‟>3.5kW/m) in one time-step are fulfilled.

b) – athermal release during the whole irradiation

A temperature-independent (a-thermal) component of fission gas release is

calculated, applying the empirical relation:

a thermalf a bu

Here bu denotes the local burnup [MWd/kgHM] and a is an empirical coefficient

(6.17×10-5).

c) – athermal release from the high burn-up structure (HBS).

When the local accumulated burn-up in the fuel is over 60-75 MWd/kgHM, a

High-Burnup Structure (HBS) is formed. The TRANSURANUS model of HBS

implementation assumes a transfer of a fraction of the fission gas from the grains into

the HBS, driven by a burn-up dependent rate equation [11]. The fission gas is at first

retained in the HBS and as soon as the local burnup exceeds an additional empirical

threshold, the HBS is assumed to be saturated, i.e. all additionally arriving fission gas

is immediately released to the free volume. The present standard burnup value for

this threshold in TRANSURANUS code is 85 MWd/kgHM but there are a number of

experiments (High-burn-up Rim Project) [12,13] which show 100% retention of fission

gas in the HBS up to a burn-up of 100 MWd/kgHM. This “saturation approach” is still

a subject of discussions.

2.2.6 Cladding corrosion

Cladding corrosion rate mainly depends on outer cladding temperature, reaction

activation energy and fast neutron flux. Other parameters that can enhance or delay

corrosion are the coolant chemical environment (in particular dissolved lithium and

boron quantities) and the cladding chemical composition. Depend on reactor used

these parameters varied in large boundaries and cladding oxide layer prediction is

subject of the largest uncertainty.

Different outer cladding corrosion models available in the code were applied during

this investigation.

icorro=3 MATPRO BWR conditions

10

icorro=4 MATPRO PWR conditions

icorro=14, EPRI/C-E/KWU model for PWR conditions14

2.2.7 Recent extensions of the TRANSURANUS fuel performance code

The extended TRANSURANUS version comprises new model for the kinetics of grain

growth in both UO2 and MOX fuels and models for the production, transport and

release of the He.

The grain growth model was refined on the basis of the recently executed

experiments for UO2 fuel and MOX fuel as well. New kinetic coefficient was derived

from the series of experiments and a comparison has been made with earlier

simulations that had applied a kinetic coefficient derived from mean linear intercepts.

Only at high fuel temperatures, there is a small influence of the grain growth

simulation on operational quantities.

A preliminary model for production, transport and release of He has been included in

TRANSURANUS code [15]. Its approach is analogous to that taken for simulating the

behaviour of Xe and Kr, i.e. assuming diffusion to the surface of spherical grains,

followed by release to the free volume. The average value of the diffusion coefficients

given in [16,] 17,18] was applied. The release of He from the grain boundaries to the

free volume is treated according to the findings in 19

The last TRANSURANUS version was applied for AREVA case simulation with

standard code options and models (see table).

Table: The main option of the AREVA case simulation.

Model Description

Fission gas release URGAS algorithm, with thermal diffusion

coefficients of Matzke and constant athermal

diffusion coefficient. (fgrmod=6)

Saturation limit for concentration

of grain boundary gas igrdbm=1

Constant value, standard for the code.

Threshold burnup for FGR from

the HBS (MWd/tU)

Standard for the PWR-RRR1=85MWd/kgU,

Fuel densific. at BOL idensi=2 Simple empirical model.

Cladding creep rate

ModClad(7)=20

Lassmann-Moreno model of Zircaloy effective

creep rate calculation.

Irradiation growth of the cladding

ModCladl(4)=20

Standard for the code

UO2 material properties Standard TU models for the UO2

11

Standard for PWR fuel swelling

model ModFuel(4)=20

Swelling rate for LWR fuel

Thermal conductivity of the fuel,

ModFuel(6)=21

The standard TU correlation for UO2,

accounting for the local porosity.

Fuel relocation model

ireloc=8

Modified FRAPCON 3 model

Corrosion model icorro=4 MATPRO model (PWR conditions)

2.3. Results and discussion

The fuel centre temperature calculated according data of irradiation power and code

models shows normal conditions without any rapid perturbation and it is lower than

1000oC. After fourth cycle of irradiation (Bu>45 MWd/kgU) nevertheless linear heat

rate decreasing the fuel centre temperature arises up to 1100oC. The main reason for

fuel temperature Increasing is fuel thermal conductivity degradation at higher burnup.

Fig. 2.1. Irradiation power and predicted fuel centre temperature.

Results of the idealized case simulation illustrate the code prediction for FGR as a

function of burnup up to 80 MWd/kgU. The calculations (Error! Reference source

not found.) demonstrate that the TRANSURANUS code can predict a smooth

development of FGR up to 80 MWd/kgU without any instability.

200

400

600

800

1000

1200

0 20000 40000 600000

5

10

15

20

25

Fuel Centre TemperatureLinear Heat Rate

LHR

Time (h)

Te

mper

atur

e (o C

)

Line

ar H

eat

Rate

(kW

/m)

AREVA idealized case

12

Fig. 2.2. Irradiation power and fission gas release v/s burnup

The black line presents code predicted gas release after third, fourth and seventh

cycles of irradiation. The results are compared with the expected results (red points),

according the authors of this idealised case.

The calculated and expected burnups of this case are compared in the next figure.

.Fig. 2.3. Evolution of the averaged rod burnup, compared with expected values.

0

5

10

15

20

25

0 20 40 60 800

4

8

12

Data ExpectedFGR predictedLHR

Burnup, MWd/kgU

Line

ar H

eat

Rate

, [k

W/m]

Fiss

ion

gas

rele

ase,

[%]

AREVA idealized case

0

20

40

60

80

100

0 20000 40000 60000

Data ExpectedTU predicted

Time (h)

Ro

d Av

erag

e Bu

rnup

, [M

Wd/k

gU]

AREVA idealized case

13

Growth of the fuel grains does not observed, most probably by reason of relatively

low fuel temperature. The calculated RIM zone width reaches 1.26 mm and

according the TRANSURANUS model high burnup structure development starts

above local burnup of 40MWd/kgU. Local burnup is averaged along pellet radius.

The result of high burnup structure for the middle of the rod (7th slice) is presented

below.

Fig. 2.4 Thickness of the fully develop HBS as is predicted by the TU model.

The idealised case, provided by AREVA is a good opportunity to test code models

and compare the results with other fuel performance codes.

3. US-PWR 16x16 LTA Extended Burnup Demonstration Program

The objective of this program was to demonstrate improved nuclear fuel utilization

through more efficient fuel design and barrier cladding and increased discharge

burnup (up to 58MWd/kgU). The standard fuel rod design consists of enriched

(3.48%) UO2 solid cylindrical pellets, stainless steel compression spring and Al

spacer disc at each end of the fuel column. In additional to the standard design fuel

0

0.4

0.8

1.2

1.6

0 10 20 30 40 50 60 70 80 90

TRANSURANUS prediction

Local Burnup in slice 7 [MWd/kgU]

T

hick

ness

of

HBS,

ful

ly d

evel

oped

(m

m)

AREVA idealized case

14

rod two different design concepts were included in a limited number of rods in the

test assembly:

a) an annular fuel pellet design,

b) cladding with graphite coating (~ 8 μm thickness) on the interior surface.

The test matrix included in IFPE data base consisted of 8 full length fuel rods with

different combinations of solid and annular pellets as well as standard and inside

coated cladding. One segmented rod (9segments) is presented in the database but

the segment's power is not clearly defined and it is not included in this investigation.

The full length rods were irradiated to an assembly and lead rod average burnups of

52 and 58 MWd/kgU.

Rod

identifier

Rod type

Pellet/cladding

Rod average burnup

[MWD/kdU]

TSQ002 Standard/Standard 53.2

TSQ004 Standard/Standard 50.5

TSQ022 Annular/Standard 58.1

TSQ024 Annular/Standard 54.7

TSQ044 Standard/ID Coated 52.5

TSQ053 Standard/ID Coated 58.1

TSQ061 Annular/ID Coated 55.5

TSQ064 Annular/ID Coated 55.6

The rods of interest (FUMEX-III priority cases) are TSQ002 and TSQ22 - full length

fuel rods with solid and annular pellets of enriched (3.48%) UO2, standard Zircaloy-4

cladding and the same cladding-fuel gap, fuel stack length and filling gas pressure of

2.62 MPa. The linear heat generation rates ranged between 10 and 25 kW/m without

sharp transitions. The rod power and burnup histories were determined for 25 slices

and included in the database. Both poolside (non-destructive) and hot cell (destructive) post irradiation

examinations (PIE) of selected rods have been done. Poolside examinations of the

LTAs included visual inspection, dimensional measurements, eddy currant testing

(ECT), and waterside corrosion thickness measurement. Hot cell fuel rod PIE

included void volume measurements, fill gas analyses, cladding visual inspections,

dimensional measurements, neutron radiography, and gamma scanning.

The IFPE database comprises PIE data on:

void volume measurement and gas release volume assessment;

15

cladding outer diameter measurements;

outer cladding oxide thickness;

Grain size measurements for two rods are presented too but there are

not pre-irradiation measurements of grain size and the comparison with

calculations without the initial exact grain size is meaningless. In the report

there is data that grain size is about (7-12) μm and for the purposes of

simulation the value of 10 μm was taken.

The bulk density assessment after irradiation for two rods afforded an

opportunity to assess the fuel swelling. There is no measured data about fuel

stack length and this is a serious obstacle for reasonable assessment of the

fuel swelling.

3.1. Irradiation conditions and modeling

The rods were simulated using standard for code options and standard fuel and

cladding models. Instead of the modified FRAPCON 3 relocation model (standard)

the incorporated in the code KWU relocation model was applied.

The TRANSURANUS simulation of the FGR includes standard options: URGAS

algorithm with thermal diffusion coefficient of Matzke [9]; thermal FGR simulated by

linear dependence of burnup.[7]; the intergranular gas release model with standard

saturation limit for the gas boundary concentration was applied.[8]; the burnup

threshold for enhanced gas release from HBS was set to 100 MWd/kgU.

Outer oxide layer was simulated with MATPRO corrosion model for PWR conditions.

In addition the same model for BWR conditions was tested for comparison only.

Dish volume and end chamfers corrections were assessed on the base of data from

the Appendix 3 and table 4.2.2.of the report included in the database. Fraction of dish

volume+chamfer = 0.0228(/) for the solid pellets, and fraction of chamfer= 0.014(/) for

the annular pellets was applied. The plenum length for each test rod was fitted to

meet the measured initial rod void volume.

For two rods TSQ002(std\std) and TSQ022(ann/std) the pre-irradiated fuel density

was measured (95.3%TD and 94.8%TD) and it is somewhat different from the value

in the file 'Pre-Characterization' (95%TD). The exact value of pellet density of

standard and annular pellets was taken.

16

The power evolution for two rods with different pellet design is very similar and FCT

of the rod TSQ002 with solid pellets is higher.

Fig. 3.1LHR comparison of two rods Fig. 3.2. Predicted FCT for two rods

3.2. Results and comparison

3.2.1 Rod burnup.

Results of burnup prediction by means of the TRANSURANUS code for eight full

length rods are compared with the results of measurement.

Fig. 3.3. TRANSURANUS predicted average fuel rod burnup v/s measured.

The burnup is systematically under-predicted by the code less than 5%. The possible

non detected variation of local LHR as well as the manufacturing uncertainties on the

0

10

20

30

0 10000 20000 30000 40000

TSQ022, annular pelletsTSQ002, solid pellets

Time (h)

L

ine

ar

He

at

Ra

te (

W/m

m)

16x16 US PWR exp. Linear heat rate at the midlle of the rods ( sl.13)

0

100

200

300

400

500

600

700

800

900

1000

1100

1200

0 10000 20000 30000 40000

TSQ022, annular pelletsTSQ002, solid pellets

Time (h)

F

ue

l C

en

tra

l T

em

pe

ratu

re (

oC

)

Fuel temperature in the pellet centre (13 sl.)

30

40

50

60

70

30 40 50 60 70

-5%

TSQ053

Burnup measured (MWd/kgU)

Bu

rnu

p c

alc

ula

ted

( M

Wd

/kg

U)

US-PWR-16x16 LTA Program

17

pellet density, dimensions and enrichment result in observed difference predicted v/s

measured burnup.

3.2.2 Cladding dimension change

The cladding diameter change was applied as a reference by choosing optional set

of models. The difference in pellet design (solid and annular pellets) leads to different

FCT and as a consequence different thermal expansion, FGR and gap conductance.

It might be expected that the relocation of pellet column would be different for the

solid and annular pellets. Two TRANSURANUS relocation models (modified

FRAPCON and KWU-LWR model) were tested and corresponding affect the fuel rod

dimensions was compared with experimental data. The comparison revealed that in

both cases fuel stack of annular pellet as well as of solid pellet the fuel relocation

model KWU-LWR is preferable. With standard FRAPCON relocation model the

pellet-cladding gap closures earlier and cladding creep down is smaller.

The experimental data on cladding outer diameter change along the fuel stack,

corrected for oxide layer for the rods with different pellet design and standard

cladding are compared with code predictions. The results are presented in the Fig.

3.4 (solid pellets) and Fig. 3.5 (annular pellets).

Fig. 3.4.Comparison of the outer cladding diameter predicted by

TRANSURANUS code and rod dimensional measurements of rod TSQ002.

9.56

9.58

9.60

9.62

9.64

9.66

9.68

9.70

9.72

9.74

0 1000 2000 3000 4000

r02, exp. dataTU prediction, solid pelletsinitial

Axial Position (mm)

Cla

d O

ute

r D

iam

ete

r (m

m)

US PWR 16x16, Rod TSQ002, solid pellets

18

Fig. 3.5. Outer cladding diameter change of the rod TSQ022 with annular

pellets, compared with dimensional measurements.

The TRASURANUS code predicted well cladding creep of rods with solid and

annular pellets and standard Zircaloy cladding.

3.2.3 Cladding corrosion.

Cladding corrosion depends strongly of the coolant temperature, time of irradiation,

cladding composition and method of preparation as well as coolant chemistry.

The IFPE database comprises data on cladding corrosion from poolside

examinations, eddy currant testing (ECT) and waterside oxide measurement. Hot sell

fuel rod PIE provided data on oxide layer by metallographic measurements. The

oxide layer thickness for 7 rods included in the IFPE database is corrected based on

these two types of oxide measurements.

The cladding corrosion for the tested rods was simulated by the standard MATPRO

corrosion model of Zyrcaloy-4 cladding for PWR conditions. The MATPRO corrosion

model for BWR conditions was tested too.

9.56

9.58

9.60

9.62

9.64

9.66

9.68

9.70

9.72

9.74

0 1000 2000 3000 4000

exp. dataTU prediction, annular pellets

initial

Axial Position (mm)

Cla

d O

ute

r D

iam

ete

r (m

m)

US PWR 16x16, RodTSQ022, annular pellets

19

Fig. 3.6 Fig. Measured and calculated oxide thickness along fuel rod length.

Fuel rod TSQ002 with solid pellets and standard Zircaloy cladding.

Fig. 3.7 Fig. Measured and calculated oxide thickness along fuel rod length.

Fuel rod TSQ022 with annular pellets and Zircaloy cladding.

0

20

40

60

80

100

0 1000 2000 3000 4000

BWR-MATPRO corrosion modelPWR-MATPRO corrosion modelexp

MATPRO-BWR

MATPRO

Axial position (mm)

Ou

ter

Cla

d O

xid

e T

hic

kn

ess (

m)

TSQ002

0

20

40

60

80

100

0 1000 2000 3000 4000

BWR MATPRO corrosion model PWR MATPRO corrosion modelexp.

Axial Position (mm)

Ou

ter

Cla

d O

xid

e T

hic

kn

ess (

m)

TSQ022

20

Pellet type (annular or solid) is not expected, to influence water side corrosion

because it depends strongly on coolant temperature. The small difference in

measured oxide thickness for the rods with solid and annular pellets is most probably

caused with difference in Burnup at EOL.

The TRANSURANUS MATPRO models for Zircaloy corrosion were calibrated for the

PWR or BWR conditions, but the corrosion rate of the Zr allow depends on many

other cladding properties as differences in composition and/or fabrication processing

or hydrogen pick-up. Depend on reactor used these parameters varied in large

boundaries and cladding oxide layer prediction is subject of the largest uncertainty.

3.2.4 Fuel densification and swelling

The fuel densification after start of irradiation and fuel swelling are two independent

physical processes that cause the density change. The fuel densification depends on

the fuel temperatures and started with irradiation and process stop up to burnup of

10MWd/kgU, depending of loaded power (temperature). The fission product swelling

of the fuel depends on burnup. The measured as well as the predicted density

difference ( Δρ = ρEOL(%TD) – ρBOL(%TD)) is negative for both solid and annular

pellets.

TU simulation of density change included standard for LWR fuel swelling model and

empirical rate of the densification [3].

At 3 different axial positions (3, 13 and 14 slices), the density of solid fuel pellets

(Rod TSQ002) was measured both - pre-irradiation and at EOL. Annular pellets from

rod TSQ022 were examined at two axial position and data are included in the IFPE

database (Appendix 7of the 20)

The experimental density was determined by weighing. The predicted by means of

TRANSURANUS code density was assessed from pellet radial density distribution by

averaging over 60 radial coarse zones using weight for different zones.

The measured and the TU predicted data of the pellet volume change are displayed

in the Error! Reference source not found..

21

Fig.3.8. Pellet volume change v/s local burnup. Comparison with the measured

data.

To obtain the trend of the curve of fuel swelling with burnup, the calculations were

performed in the wide burnup range and this calculated data are presented in the

figure. The change of pellet volume ΔV/V (%) as an assessment of the densification

and swelling was calculated from the changes of the pellet radius as well as pellet

height, after extracting increment due to relocation. The presented data correlate with

the several measured points.

3.3. Summary

Fuel rods with both solid and annular pellets are simulated by means of

TRANSURANUS code without any significant differences in the results.

The rod average burnup predictions are underestimated by less than 5%.

The cladding creep simulation is in good agreement with experimental

measurements for both rods with annual as well as solid pellets.

The peak and averaged oxide thickness is underestimated with standard

MATPRO corrosion model for PWR conditions.

-3

-2

-1

0

1

2

3

4

5

0 20 40 60

exp.data, annular pelletsTu calcexp.data, solid pellets

annular pellets

solid pellets

Burnup, (MWd/kgU)

V/V

(%

)

TSQ002 (stand/stand.)&TSQ022(annular/stad.)

22

4. GINNA reactor experiment.

4.1. Irradiation conditions and simulation

Siemens Power Corporation 14x14 Lead fuel assemblies were irradiated in PWR

GINNA reactor up to 60MWd/kgU. The goal of the program was to demonstrate and

evaluate the potential of annular pellets and zirconium barrier cladding for resisting

fuel failures due to pellet-cladding interaction. The annular pellets were designed to

operate at lower FCT and to better accommodate pellet restructuring and FGR. Since

the hole represents about 10% of the total pellet volume, annular pellet are U-235

enriched about 10% higher than comparable solid pellets (3,71% and 3.35%

respectively). Two segmented rods differ only in pellet diameter and that were used

to vary the pellet- cladding gaps. Barrier cladding was designed with a Zr inner layer to minimize PCI and to improve

resistance of the cladding to stress corrosion cracking. The inner Zr layer comprises

about 10% of the total wall thickness, the overall cladding dimensions are the same

as standard cladding. The Zr barrier is fully bonded to the Zircaloy to ensure good

heat transfer properties.

Each segmented rod consisted of 4 segments, located symmetrically at the centre.

Power and burnup accumulated in the two central rodlets were nearly the same for

all segmented rods. The EOL Burnup and Fluence were nearly uniform along length

of the central rodlets. The data base comprises 8 segmented rodlets with different

combinations of solid and annular pellets, standard and barrier cladding and

controlled variations in pellet to cladding gap. 3 rodlets were irradiated for 4 cycles up

to average burnup 42.5 MWd/kgU, and 5 rodlets -for 5 cycles up to average Bu=52

MWd/kgU. The FUMEX-III priority cases program covers two rods, rod2 with fuel

solid pellets and rod4 with annular pellets and standard Zyrcaloy-4 cladding which

were irradiated 5 cycles up to 52MWd/kgU. Linear Heat Rate at central rodlets

ranged from 8 to 34kW/m.

The PIE included data on fuel column elongation, cladding creep down and fission

gas release. The burnup was determined by LHR distribution and FGR was

measured by puncturing the cladding at the plenum level and measuring the gas

pressure in a chamber of known volume sealed around the plenum region of the

rodlets. Gamma scanning was perform on 5 of the irradiated rodlets to measured fuel

column height and to visualize the presence of pellets and to obtain information of

23

any axial pellet gaps. Six axial profilometry scans, each 30 degrees apart radially,

were made in the hotcells. Rod average diameter and rod ovality were determined

from the scans.

In order to select models for code simulation the following TU parameters have been

chosen for testing: Fuel relocation model and axial PCMI condition.

The difference in pellet design (solid and annular pellets) leads to different FCT and

as a consequence different thermal expansion, FGR and gap conductance. The

friction forces between pellet and cladding suppose to be different and fuel column

relocation has to be accordingly modelled. The experiment offers data for fuel column

elongation and cladding outer radius and they are taken as reference to chose

appropriate relocation model. TRANSURANUS owns different optional models to

describe the phenomena relocation affecting the fuel rod dimensions. The main

models are modified FRAPCON-3 model (the contribution from relocation is

accounted by circumferential strain component as a linear function of LHR) and

modified KWU-LWR model (simple model treating the relocation as a single (lump)

fuel volume increasing at the BOL [5]. Fuel column increasing determined by gamma

scanning of rodlets was used as a test of these two relocation models.(see Fig.4.1).

Fig.4.1. Comparison of the measured fuel axial deformation and code

predictions with two relocation models in the case of solid as well as annular pellets.

0

4

8

12

0 4 8 12

std.reloc. modelKWU reloc.model

annular solid

Measured Fuel Stack Elongation (mm)

Pre

dic

ted

Fu

el S

tack E

lon

ga

tio

n (

mm

)

GINNA. Rod 2 (solid pellets) and Rod 4 (annular pellets)

24

The fuel column elongation is under-predicted by the code if standard relocation

model (FRAPCON-3 modified model) is applied in the cases with the different pellet

design. With modified KWU relocation model the annular pellet fuel column is over

predicted (21%) and solid pellet fuel stack is under-predicted (27%). Pellet relocation

influence on gap size and consequently on fuel temperature and FGR as well as the

cladding diameter change.

PCMI conditions arises after gap closing and one should supposed different

friction forces for the two type of pellet design. The TRANSURANUS code owns two

possibilities to assess and include axial friction forces in the case of pellet cladding

contact.

-'No slip'[21] conditions for the axial PCMI

-URFRIC model for the calculation of the axial friction forces [22].

The simulations with and without URFRIC model revealed that there is no significant

influence of friction forces modelling on the cladding dimensions, and consequently

the temperature and FGR. Both cases are simulated without URFRIC model

implementation. The TRANSURANUS models and options chosen for the GINNA

rods simulation are listed in the next table.

Table. Set of models used in TRANSURANUS simulations of the GINNA experiment

Model Description

Fission gas release URGAS algorithm, with thermal diffusion coefficients

of Matzke (fgrmod=6).

Saturation limit for gas

concentration of grain boundary

The temperature dependence of the grain

boundary saturation concentration (igrbdm=2)

Burnup threshold for FGR from

the RIM zone [MWd/tU]

RRR1=100000 MWd/kgU,

Fuel Densification model Simple empirical densification model (idensi=2).

Cladding creep rate ModClad(7)=20, standard for Zyrcaloy

UO2 material properties Standard TU models for the UO2

Thermal conductivity of the fuel The standard TU correlation for UO2 and (U,Gd)O

2 ,

accounting for the local porosity ModFuel(6)=21.

Fuel relocation model Modified KWU-LWR model(ireloc=5)

Axial friction force model Axial friction forces are not calculated (ixmode=0).

Corrosion model EPRI/C-E/KWU model (icorro=14)

25

4.2. Results of TRANSURANUS simulation

The results of TRANSURANUS code calculations with above set of options are

presented in the table.

Table. Comparison of the experimental data for the FGR, pressure, fuel column

elongation and diameter change with outcome of code simulations.

R2-solid pellets R4-annular pellets

Predicted Exp.data Predicted Exp.data

Bu averaged

MWd/kgU

50.7 51.7 56.6 57.0

FGR [%]

3.06 2.36 0.91 0.92

Inner Press

[MPa]

3.00 2.88 2.47 2.32

Void Volume

at EOL [cc]

4.18 3.8 7.04 7.00

∆D [mm]

-0.056 -0.091 -0.066 -0.097

∆Lclad [mm] 1.80 -

1.9 -

∆Lfuel [mm] 5.23 7.3

5.7 4.7

Oxide layer

[µm]

16.7 14 ÷ 41 23.9 16 ÷ 39

4.2.1 Fuel Central Temperature and Fission Gas Release

The fuel central temperature of two rods as predicted by the code indicates a

temperature difference of up to 300oC. Experimental data for the real temperature

difference is not available.

26

Fig. 4.2. Predicted Fuel Centre Temperature for two rods with different pellet

design.

The measured by rodlet puncturing fission gas release and FGR calculated by

means of TRANSURANUS code are pictured in the next figure.

Fig. 4.3. Comparison of the Fission Gas Release as a function of burnup with

puncturing results.

0

500

1000

1500

0 10000 20000 30000 40000

R2-solid pelletsR4-annular pellet

Time (h)

Fu

el C

en

tra

l T

em

pe

ratu

re (

oC

)

GINNA exp. Rod 2(solid pellets) and Rod 4 (annular pellets)

0

1

2

3

4

0 20 40 60

r2, exp.datar2, solid pelletsr4, exp.datar4, annular pellets

Average Burnup (MWd/kgU)

Fis

sio

n G

as R

ele

ase

(%

)

GINNA exp. Annular and solid pellets rods

27

4.3. Cladding creep down

The outcome of this experiment 23 revealed for all rod design variations the cladding

creep down was essentially the same. After 4 cycles of irradiation diameter decrease

over fuel region of cladding of 90-101 μm was detected. During 5th cycle the solid

pellet rodlet cladding experienced a creep down reversal of ~10 μm., suggesting for

solid contact between the pellets and cladding. The creep rate of cladding in the

fuelled sections of annular pellet rodlets diminished during the fifth cycle of

irradiation. A comparison of creep down in the plenum sections and the fuelled

sections indicates that creep reversal took place during fifth cycle for annular pellets

as well, and that hard pellet-to-clad contact in the annular pellet rods was

established.

These two rods Rod2 and Rod4 are with identical cladding material and initial gap

(180µm). There is no noticeable difference between the measured cladding creep

down for two pins with different pellet design.

Fig. 4.4. Comparison of the cladding creep down for two rods.

The measured cladding creep down is about 100 µm and the predicted one is 60-

70µm. The diameter decrease of the rod with annular pellets is slightly higher than

the cladding decrease of rod with solid pellets as it is reproduced by the code.

10.4

10.5

10.6

10.7

0 200 400 600

initialr4, calculatedr4, exp.r2-expr2-calculated

calc.

exp.

Axial position (mm)

Clad

Out

er D

iame

ter

(m

m)

GINNA exp. Rod 2 (solid pellets) and Rod 4 (annular pellets)

28

4.4. External Oxide thickness

Cladding corrosion rate mainly depends on outer cladding temperature, reaction

activation energy and fast neutron flux. Other parameters that can enhance or delay

corrosion are the coolant chemical environment (in particular dissolved lithium and

boron quantities) and the cladding chemical composition. Depend on reactor used

these parameters varied in large boundaries and cladding oxide layer prediction is

subject of the largest uncertainty.

One of the conclusions made in this experiment is that the corrosion rate of the

standard cladding appeared to be higher than the corrosion rate of barrier cladding.

Small differences in composition and/or fabrication processing or differences in

hydrogen pick-up and hydride distribution could be possible causes.

Different outer cladding corrosion models available in the code were applied during

this investigation and the EPRI/C-E/KWU water side corrosion model 24 for PWR

conditions was chosen for the GINNA rods simulation. Calculated outer cladding

oxide layer of different rods compared with measured is presented on next two

figures.

Fig. 4.5. Oxide outer cladding thickness of the Rod2 with solid fuel pellets –

measured and predicted.

In order to facilitate the experiment-calculation relationship, the experimental data

was fitted by linear function (blue line).

0

0.01

0.02

0.03

0.04

0.05

0 200 400 600

KWU model, 5 cyclesfit of exp. data: y = +3.79E-5x

1 +0.016

exp. data after 5 cycles

Axial Position (mm)

O

xid

e O

ute

r C

lad

din

g T

hic

kne

ss

(mm

)

GINNA exp, Rod2

29

Fig. 4.6. Comparison of the calculated end measured oxide layer for Rod4 with

annular pellets

In the vicinity of FA spacers, where coolant mixing tends to lower the clad

temperature and the oxidation by the coolant is reduced. This tendency is expressed

in the case of Rod4 – the oxide layer is thicker in middle of the height of the segment

(Fig. 4.6). The similarity in oxide thickness among the different fuel rod types is not

surprising since the waterside surface in all cases is Zircaloy 4. Pellet type (annular

or solid) is not expected to influence waterside corrosion (oxide thickness), because

it depends strongly of the coolant temperature and therefore of the axial localization

of segmented rodlet (upper or lower central part of the full length rod). For instance

on the next picture the prescribed coolant temperature or two five cycles rods is

compared. Rod2 is located in lower central part of the segmented rod and the rod4 is

located in upper part and respectively the averaged coolant temperature is different

(see Fig. 4.7). The thicker oxide layer was observed for the Rod4 with higher

temperature.

0

0.01

0.02

0.03

0.04

0.05

0 200 400 600

r4, calc. KWU corrosion model r4, exp.data

Axial Position (mm)

Oxi

de

ou

ter

cla

dd

ing

th

ickn

ess

(m

m)

Annular pellet, Standard Cladding, Normal gap.

30

Fig. 4.7. The coolant temperature v/s time of irradiation, prescribed in the IFPE

database

4.5. Pellet morphology

Pellet morphology was studied in detail for both solid and annular pellets. Cracking

was observed in the pellets of all ceramographically examined rods. The central hole

in the annular pellet tended to promote the formation of radial cracks, while cracking

in the solid pellets seemed to occur in a random pattern. The morphology of solid and

annular pellets was essentially the same. The hole of the annular pellets was fully

preserved and free of fuel fragments and the dishes of the solid pellets remained fully

visible.

Ceramographs with a magnification of up to 1000X were used to characterize the

patterns of fission gas precipitation within, and at the boundaries of, the UO2 crystal

grains. The zones have been differentiated by the UO2 grain size and porosity

distribution within and between grain boundaries.

A pellet RIM zone – 50 -70 microns wide (r/r0=0.99) characterised by indistinct grain

boundaries and very high porosity were established. In this zone high Pu

concentration and relatively high burnup, were revealed too.

A band of about 200µm wide corresponding to the transition from zone with not very

numerous intragranular pores and fission gas at the grain boundaries to the zone

250

270

290

310

330

0 10000 20000 30000

average, r2average, r4

Cumulative Time (hours)

Co

ola

nt

Te

mp

era

ture

[o

C]

Coolant Temperature [oC] as given in IFPE data base for the rods r2(SSN5) and r4(ASN5)

31

where fine and numerous fission gas bubbles appears is named transition and it

radial location was determined experimentally.

The radial location of corresponding transition zone in the pellets simulation was

determined from the calculated gaseous porosity along the pellet radius. The

calculated gaseous porosity does not include the porosity in high burnup structure

because it is too high and transition zone gets invisible.

The location of transition zone was assessed as a centre of zone of transition

between two zones with different rate of porosity growth and present in the table.

The radial distributions of the gaseous porosity for two pellet designs are pictured on

Error! Reference source not found.

Fig.4.8. The predicted radial profile of the gaseous porosity of Rod2 (solid

pellets). and experimentally observed border of starting enhanced gaseous swelling

of the pellet

0

0.002

0.004

0.006

0.008

0.010

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

exp.data, r/r0 =0.62

gaseous porosity

Radial Position (mm)

Ga

se

ou

s P

oro

sity

( /

)

GINNA exp., Rod2, solid pellets

32

Fig.4.9. Predicted radial profile of the gaseous porosity for the annual pellet and

experimentally determined radial position of the transition zone.

The experimentally assessed outer radial border of region with copious intragranular

fission gas bubbles in the case of annular pellet is pointed on Fig.Error! Reference

source not found. for comparison.

The radial location of the transition zone determined from the calculated porosity

distribution as well as experimentally observed border of the transition zone are

represented in the table.

Table. Radial location (r/r0) of the transition to intragranular fine fission gas bubble

formation. Last column present the wide of calculated zone with high burnup

structure.

Rodlet

number

r/r0 –

exp

r/r0-

calc

Bu

calc.

RIM

[μm]

R2 - SSN5 0.62 0.62 50.6 67

R3 – ASN4 0.58 0.59 47.2 33

R4 – ASN5 0.60 0.62 56.6 124

R5 – AZN4 0.58 0.59 46.6 36

R6 – AZN5 0.64 0.59 56.2 120

R8 – AZW5 0.62 0.65 57.2 131

R9 – AZM5 0.60 0.59 54.9 100

0

0.001

0.002

0.003

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 1.1

gaseous porosityexp.data, r/r

0=0.60

Radial Position (mm)

G

ase

ou

s

Po

rosity

(/

)

GINNA exp, Rod4-annular pellets

33

The width of RIM zone (fully developed) was calculated and presented in the table

below. Experiment gave RIM zone width of 50-70 μm.

4.6. Summary

Eight segmented rodlets with different combinations of solid and annular pellets,

standard Zircaloy-4 and barrier cladding with inner pure Zr layer were examined after

irradiation of four or five cycles. These rodlets were modelled and results of

TRANSURANUS code simulation are compared with experimental ones.

The results of TRANSURANUS code simulation of all eight rods are consistent with

the expectation concerning pellet design and different gap size.

The measured cladding creep down is about 100 µm and the predicted one is 60-

70µm. The diameter decrease of the rod with annular pellets is slightly higher than

the cladding decrease of rod with solid pellets as it is reproduced by the code

External oxide thickness was found to reach a maximum of 40 to 42 μm for standard

cladding after five cycles of irradiation. The corrosion rate of the standard cladding

appears to be higher than the corrosion rate of barrier cladding but the lower

maximum external oxide thickness of barrier clad cannot be directly attributed to the

presence of the Zr layer. In TRANSURANUS simulations the KWU corrosion model

was applied in the both cases standard and barrier cladding and hence very similar

results were calculated. The comparison with the not uniform measured oxide layers

shows good coincidence.

The pellet morphology was studied in detail for both solid and annular pellet and it

was found that fuel structure at EOL is essentially the same. Experimentally was

determined the radial location of border of enhanced porosity growth and start of

significant fuel swelling. Radial location of this border was compared with calculated

by code radial distribution of gaseous porosity and a good coincidence was pointed

out.

The annular-pellet design performed well throuth five cycles of irradiation and an

assembly birnup of 52MWd/kgU. The simulations by the means of TRANSURANUS

code reproduced the main features of new design and confirmed its advance

possibilities for greater margins of operation at higher burnup.

34

5. Conclusions

The TRANSURANUS code ability to predict the fuel performance was demonstrated

based on the IFPE database of the OECD/NEA within the CRP“Improvement of

models used for fuel behaviour simulation” (FUMEX-III). The last TRANSURANUS

code version was applied to run simulation of fuel rod performance under steady

state operation condition, transient power ramps as well as a variety type of fuel and

cladding. From the present status of the TRANSURANUS computations the following

general conclusions can be drawn when analyse the fuel performance of WWER

rods as well as the Gd doped fuel and the transient experiments Riso3, (rods II5 and

GE7) and KOLA3-MIR.

The TRANSURANUS fuel performance code is mature and can deal with fuel

behaviour up to high burnup,no numerical problems arose in the simulations of

the LWR fuel behaviour.

The TRANSURANUS rod averaged burnup predictions are in very good

agreement with data of measurement. In general, all the calculation fairly falls

into, or quite close, to the ±5% lines. The non-detected variations of LHR

during operation as well as the manufacturing uncertainties on pellet density,

dimensions and enrichment result in spread of the measured burnup.

The agreement between calculated and measured fuel centre temperatures in

ramp tests of re-fabricated fuel rods is very satisfactory for all re-irradiated

cases, except for the Rod 50 from Kola3-MIR experiment, where the

thermocouple of the re-fabricated rod broked down during the first stage of

test irradiation. The data recorded before the failure noticeably overestimated

the calculated temperatures.

Fission gas release, predicted by means of TRANSURANUS code after

steady-state operation is in very good agreement with measured by rod

puncturing FGR at EOL for all cases. Two USA experiments with rods with

different fuel pellet design, with central hole and without it demonstrated that

rod with solid pellets release more gases than the rod with central hole in

pellet. TU reproduced these results. The FGR data after power ramp test (both

rod II5 and rod GE7) are not well predicted, but the results, in general, fall into

the „licensing-acceptable‟ sector. Fission gas relelease more then 40%,

reported in the Kola3-MIR experiment does not reproduce at all.

35

Fuel stack elongation was measured in several cases and well reproduced by

optional set of models. Axial fuel changes predicted by TRANSURANUS for

GAIN rods as well as for the two GINNA rods with simple empirical relocation

model are closer to the measured data in comparison with predictions with

standard model (modified FRAPCON 3).

As concerns the calculated cladding creep down (cladding outer diameter

change) at EOL, the simulation with optional selection of models (model of fuel

relocation and simple model for the swelling) results better agreement with

measured data (GAIN experiment and .RISO3 rods) The cladding creep down

of GINNA rods is under-predicted, for annular pellet rod by 30% and for solid

pellet rod by 50%. Rod GE7 has a special cladding Zircaloy-2 with bonded Zr

liner and most probably there is a need of special cladding creep correlation.

The calculated grain size changes in the center, middle and periphery of the

pellet are in good agreement with the experimental data.

Acknowledgements

The work, described in the present report, was completed in close collaboration with

the Modelling Group of the Institute for Transuranium Elements in Karlsruhe,

Germany. The authors would like to express their gratitude to P. Van Uffelen, A.

Schubert, and J. van de Laar for their valuable help and productive discussions.

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37

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