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ÚJV 13815-T, M December 2011
IAEA Research Agreement No. 15456
FUEL BEHAVIOUR SIMULATIONS IN FUMEX III CRP AT NRI
Report covers period: May 2009 - December 2011
M. VALACH ∗∗∗∗), J. KLOUZAL, M. DOSTÁL, J. ZYMÁK
∗) chief scientific investigator
Nuclear Research Institute Řež plc
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ÚJV 13803-T, M
Nuclear Research Institute Řež plc
FUEL BEHAVIOUR SIMULATIONS IN FUMEX III CRP AT NRI
Authors: M. Valach, J. Klouzal, M. Dostál, J. Zymák Chief scientific investigator: M. Valach IAEA Research Agreement No.: 15456 Global IAEA Project: FUMEX III
Report covers period 05/2009 - 12/2011
Řež, December 2011
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ABSTRACT Report summarises simulated cases in the frame of FUMEX III Project at the NRI Rez plc.
Acknowledgement: Authors would like to thank OECD NEA and IAEA for kind support of the FUMEX III Project.
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CONTENTS
1. INTRODUCTION ....................................................................................................... 5
2. BRIEF CHARACTERISATION OF COMPUTER CODES USED FOR FUMEX-III CALCULATIONS ......................................................................................... 6
2.1. FRAPCON ............................................................................................................. 6 2.2. FRAPTRAN .......................................................................................................... 6 2.3. TRANSURANUS .................................................................................................. 7 2.4. FEMAXI-6 ............................................................................................................ 7
3. SUMMARY OF RESULTS ........................................................................................ 8
3.1 CASE KOLA/M IR ......................................................................................................... 8
3.1.1. Case description ............................................................................................ 8 3.1.2. PIE DATA description ................................................................................. 10
3.1.3. Kola-3 cases calculated using TRANSURANUS code ................................ 12
3.1.4. Kola-3 cases calculated using FEMAXI-6 .................................................. 33
3.1.5. FRAPCON 3.4 calculation .......................................................................... 40
3.2. AREVA “ IDEALIZED CASE” ................................................................................. 42
3.2.1. TRANSURANUS results .............................................................................. 43
3.2.2. FEMAXI-6 results ........................................................................................ 45
3.3. US-PWR 16X16 LTA .......................................................................................... 47
3.3.1. TRANSURANUS results .............................................................................. 48
3.3.2. FEMAXI-6 results ........................................................................................ 51
3.4. LOCA HALDEN IFA-650.1 AND 650.2 ................................................................. 54 3.4.1. IFA-650.1 calculations ................................................................................ 56
3.4.2. IFA-650.2 calculations ................................................................................ 64
3 CONCLUSIONS ........................................................................................................ 68
4 REFERENCES .......................................................................................................... 69
5 APPENDIX ................................................................................................................ 70
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1. INTRODUCTION NRI Rez plc took part in the previous coordinated research projects focused on fuel behaviour modelling held by the IAEA - FUMEX-I and FUMEX-II. These were very helpful for the development and validation of various codes used in the Nuclear Research Institute Řež (NRI) for the evaluation of the fuel rod thermomechanical behaviour. The NRI participation in the IAEA´s Co-ordinated Research Project “Improvement of Computer Codes Used for Fuel Behaviour Simulation FUMEX-III” is based on the IAEA Research Agreement No. 15456 with the following scope: and programme:
FUMEX III priority cases are as follows
Based on the considerations of our needs related to the modeling for Czech NPPs we have performed basic parametric calculations of two LOCA cases (IFA-650.1 and IFA-650.2) and detailed evaluation WWER related cases Kola MIR ramp rods. The AREVA “Idealized case” and 16x16 LTA cases were also calculated because of the high burnup reached.
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2. BRIEF CHARACTERISATION OF COMPUTER CODES USED
FOR FUMEX-III CALCULATIONS NRI Rez plc. was a part of an agreement with PNNL and GDF-SUEZ for using FRAPCON and FRAPTRAN codes for cases IFA-650.1 and IFA-650.2. The other analysis effort lies in the calculation of WWER and high burnup related cases by the codes used in NRI Rez plc. for thermomechanical calculations (FEMAXI-6 and TRANSURANUS). 2.1. FRAPCON FRAPCON-3 ([1]) is a Fortran 90 computer code that calculates the steady-state response of light-water reactor fuel rods during steady state operation to high burnup. The code calculates the temperature, pressure, and deformation of a fuel rod as functions of time-dependent fuel rod power and coolant boundary conditions. The phenomena modeled by the code include 1) heat conduction through the fuel and cladding to the coolant; 2) cladding elastic and plastic deformation; 3) fuel-cladding mechanical interaction; 4) fission gas release from the fuel and rod internal pressure; and 5) cladding oxidation. The code contains necessary material properties, water properties, and heat-transfer correlations. FRAPCON-3 is programmed for use on Windows-based computers, but the source code may be compiled on any computer with a Fortran 90 compiler. The FRAPCON-3 code is designed to generate initial conditions for transient fuel rod analysis by the FRAPTRAN computer code. The versions FRAPCON-3.4 and 3.4a were used for the calculations. 2.2. FRAPTRAN The Fuel Rod Analysis Program Transient (FRAPTRAN) ([2]) is a Fortran language computer code that calculates the transient performance of light-water reactor fuel rods during reactor transients and hypothetical accidents such as loss-of-coolant accidents, anticipated transients without scram, and reactivity-initiated accidents. FRAPTRAN calculates the temperature and deformation history of a fuel rod as a function of time-dependent fuel rod power and coolant boundary conditions. Although FRAPTRAN can be used in “standalone” mode, it is often used in conjunction with, or with input from, other codes. The phenomena modeled by FRAPTRAN include a) heat conduction, b) heat transfer from cladding to coolant, c) elastic-plastic fuel and cladding deformation, d) cladding oxidation, e) fission gas release, and f) fuel rod gas pressure. FRAPTRAN is programmed for use on Windows-based computers but the source code may be compiled on any other computer with a Fortran-90 compiler. Burnup-dependent parameters may be initialized from the FRAPCON-3 steady-state single rod fuel performance code. The version FRAPTRAN 1.4 was used for the calculations.
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Material properties correlations for oxide fuels and cladding materials are presented and discussed in NUREG/CR-7024 ([3]). Comparisons are made between the material property correlations used in the most recent versions of the codes, FRAPCON-3.4 and FRAPTRAN 1.4. Comparisons are also made with MATPRO, which is a compilation of material property correlations with an extensive history of use with various fuel performance and severe accident codes. 2.3. TRANSURANUS TRANSURANUS ([5]) is a computer program for thermal and mechanical analysis of nuclear fuel rods developed at the Institute for Transuranium Elements (ITU). This code can predict fuel rod behaviour during normal, off-normal and accident conditions. It has clearly defined mechanical-mathematical framework based on finite difference method. The version V1M1J09CEZ01 was used in presented calculations. This version differs from the V1M1J09 version only in the incorporation of few vendor-specific models for E110 cladding, the difference between these versions was very small for the cases analysed within the FUMEX III project. 2.4. FEMAXI-6 FEMAXI-6 (ver.1) ([4]) predicts thermal and mechanical behaviour of a light water reactor (BWR and PWR) fuel rod during normal and transient (not accident) conditions. The solution of thermal and mechanical analysis is fully coupled. It can analyze integral behaviour of a whole fuel rod throughout its life (to high burnup) as well as the localized behaviour of half pellet size part of fuel rod (especially for PCMI studies). The mathematical-physical framework is based on finite element method (FEM). The results from detailed burning analysis can be effectively tranfered from other codes (such as RODBURN and PLUTON). The FEMAXI-6 (ver. 1) was used for calculations.
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3. SUMMARY OF RESULTS 3.1 CASE KOLA/MIR 3.1.1. Case description As a part of the IAEA FUMEX III coordinated research project some data for the transient FGR experiments performed in the MIR reactor with the VVER-440 fuel in 1996 and 1997 were released. The experiments were performed on the well characterized fuel rods irradiated in the 3rd unit of Kola NPP (VVER-440) up to 50 MWd/kgU (FA-198) and 60 MWd/kgU (FA-222). Datasets for three experiments (RAMP, FGR1, FGR2) focusing on the behaviour of the fuel in the transients are part of FUMEX III. Each dataset contains three fuel rodlets. Shorter rodlets were refabricated from the original fuel rods, the refabricated rods were filled with He (1.2 – 2 MPa). The details can be found in [7]. Two refabricated fuel rods (RFRs) were equipped with the pressure transducers (FGR-1), two with thermocouples (FGR-2). The maximum LHR during RAMP experiment RFR-33 is shown in Figure 1, whereas LHR histories and measured gas pressure, resp. fuel temperature for RFR-41 (FGR-1 experiment), resp. RFR-51 (FGR-2 experiment) are plotted in Figure 2 and Figure 3.
0 100 200 300 400 500Time, h
0
100
200
300
400
Lin
ea
r h
ate
ra
te, W
/cm
Figure 1 Max. LHR change of RFR-33 during RAMP experiment.
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0 300 600 900Time, h
0
50
100
150
200
250
300
350
400
450
Lin
ea
r h
ate
ra
te, W
/cm
0
3
6
9
12
15
18
21
24
27
Ga
s p
res
su
re, M
Pa
LHR of the RFR 41
RFR 41 gas pressure
Figure 2 Max. LHR and registration of the pressure sensor of the RFR-41 (FGR-1 experiment).
0 50 100 150 200 250 300Time, h
0
50
100
150
200
250
300
350
400
450
Lin
ea
r h
ate
ra
te, W
/cm
0
200
400
600
800
1000
1200
1400
1600
1800
Te
mp
era
ture
, 0C
LHR of the RFR 51
Fuel temperature of the RFR 51
Figure 3 LHR evolution (thermocouple section) and fuel temperature data of the RFR 51 (FGR-2 experiment).
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3.1.2. PIE DATA description The FUMEX III dataset contains only the information on the FGR (composition of gas mixture in the RFR and volume of the gas in the RFR after the ramp) and limited information regarding the fuel microstructure changes ([7]). However, PIE included also the post-ramp cladding diameter and gap measurements as well as the detailed analysis of fuel microstructure (porosity, grain type and size and Xe concentration along the radius of fuel pellet). Some of this information was presented in [8] and [9]. The open data available to NRI to date are summarized in two following tables (Table 1 and Table 2):
Table 1 Kola3/MIR PIE data.
Experiment: RAMP FGR1 FGR2 FGR1 RAMP RAMP FGR1 FGR2 FGR2
RFR: 33 41 51 32 37 38 48 50 52
Burnup Bu0 [MWd/kgU] 50.8 48.9 49.5 60.2 60.1 60.2 60.5 58.4 58
Dimensional changes
D0 [mm] 9.05 9.07 9.035 9.06 9.04 9.07 9.06 9.06 9.07
∆Dmax [um] 26-33 33-54 G0 [um] 8-30 0-14
Gmax [um] 40-65 15-33 DCH
0 [mm] 1.6 DCH
min [mm] 0 0.61 0 0.5
Fuel Microstructure
RcXe<0.Rel 0.7 (~2.7mm) 0.6 (~2.3mm)
LCol.Grains [um] 50-70 WCol.Grains [um] 20-30
FGR FGR 0.313 0.475 - 0.466 0.169 0.196 0.5 - 0.484
cXe 0.619 0.554 - 0.735 0.555 0.585 0.715 - 0.474 Vgas [cc@STP] 238.5 307.6 - 359.5 171.1 188.4 310.8 - 231.1
Bu0 average burnup of the RFR before the ramp D0 outer cladding diameter before the ramp ∆Dmax increase of the outer cladding diameter after the ramp G0 diametric fuel cladding gap, before ramp Gmax max. diametric (?) fuel cladding gap after the ramp DCH
0 central hole diameter before the ramp DCH
min min. of the central hole diameter after the ramp RcXe>0.
Rel rel. radial position of local Xe concentration exceeding the 0.1 of the Xe created (0 – fuel pellet centre, 1 – fuel pellet outer edge)
FGR integral FGR during the ramp cXe Xe concentration in the gas in the free volume of the RFR after the ramp Vgas amount of gas in the free volume of the RFR after the ramp LCol.Grains max. length of the columnar grains WCol.Grains max. width of the columnar grains
The on-line measurements of the pressure in the RFRs 41 and 48 are also available. However, the final pressure recorded in RFR 41 was 4.28 MPa at 47°C. Considering the amount of the gas measured after the puncturing (307.6 cc@STP), the free volume would have to be approx. 7.85 cc. The initial free volume in the RFR was 5.84 cc ([7]), e.g. the TRANSURANUS calculation gives approx 6 cc before ramp at 50°C. The increment of the free volume of 1.85 cc corresponds to the gap reopening by approx. 0.2 mm (over the diameter), but the PIE results seem to contradict such increase. Therefore it seems that the pressure reading for the RFR 41 underestimates the real pressure, at least by the end of the ramp. 1 newly created during the transient, the as manufactured hole closed completely.
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Table 2 Fuel microstructure of segments.
Bu0
[MWd/kgU]
QlMAX
[kW/m]
RFR / ax. segment
CH CG GS U HBS
r /R r /R r /R P [%] r /R P [%] r /R P [%]
50
0 33,41,51 0 - 0.21 0.21–0.98
3.5-6.5
0.98 – 1.00 5-13 34.3 41/4, 33/7 0.00-0.74 16.0-22.0 0.74-0.98 37.5 41/5,52/1 0.00-0.70 11.0-13.0 0.70-0.98 44.1 41/8,51/7 0 - 0.16 0.16 - 0.51 0.51-0.82 12.5-13.5 0.82-0.98
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0 32,37,38,48,50,52 0 - 0.21 0.21–0.97
0.97-1.00 10-17 25.7 32/3,
0 - 0.20 0.20-0.62
21.6-28.4 0.62-0.97
31.5 32/8, 50/9 0.20-0.66 0.66-0.97
33.2 33/5, 48/9,50/8, 0 – 0.13 0.13-0.41 0.41-0.77 12.5-13.5 0.77-0.97
QlMAX max. of the linear density of the heat generation rate
r /R relative radii of the area (0 – centre of the pellet, 1 – outer edge of the pellet) P porosity CH central hole CG columnar grains GS area with the significant gaseous FP swelling U unrestructured fuel HBS High-Burnup structure
The data indicate that the dynamics of the central hole closure is different in the 50MWd/kgU and 60MWd/kgU rods. This is explained in [8] by the difference in the central temperature at given heat generation rate. At 50MWd/kgU the temperature at 32-35kW/m is not high enough to initiate the growth of the columnar grains and therefore the gas remains initially in the grains, which leads to pronounced gaseous swelling and the central hole closure. At 60MWd/kgU the temperature is higher at the same LHR and the columnar grains are created. The central hole closure during power ramp can be found also in other experiments that are available in the IFPE database. Ramp tests done in Studsvik reactor R2 as a part of multilateral projects and released to the IFPE database are summarized in Table 3. The fuel was UO2 and in some cases with large grain (LG) size and in two cases with central hole. In the SUPER-RAMP (PWR cases) experiment four fuel rods with pellets with central hole failed and the summary was that hollow pellets do not behave during ramp better in with respect to cladding failure than solid pellets. As fabricated hollow pellets length was 13.6 mm and the pellets had no chamfer. During holding at 400 W/cm all four fuel rods failed after 26 – 118 minutes. State of the central hole after the test was assessed from axial cuts and no changes were observed. From five standard rods with solid pellets (with the same length as hollow pellets, with dish, without chamfer) failed three (after 12-22 minutes on cca 400 W/cm). From other 19 rods with solid pellets (length 11 mm, with chamfer and with dish) failed at powers 400 – cca 500 W/cm only two. In the DEMO-RAMP I (BWR) experiment five rods with hollow pellets (with Nb2O5 dopant and 50 microns grain size) were tested. Ramp terminal level was cca 500 W/cm in two cases and 600 W/cm in other two cases, one ramp was with several steps (see Table 3). None of the fuel rods failed. The central hole after the test was observed smaller due to swelling in some rods –Table 4. In one rod (494 in Table 4) the hole remained the same as fabricated, in the second rod (496) was the central hole completely locally (roughly the size of half pellet height) filled/closed and was almost completely filled, in the other rods (495 and 497) the hole changed moderately.
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Thus it was experimentally observed that sufficiently high linear heat rate with long hold can cause significant central hole changes (even closure).
Table 3 Ramp tests characteristics available in IFPE.
Dataset Reactor No. of Burnup Ramp type
type fuel rods MWd/kgU Fuel Grain Cladding failed intact
INTER-RAMP BWR 20 9-22 UO2 8-11 Zry-2 1 step - 24 h hold 11 9
SUPER-RAMP (BWR) BWR 16 28-38 UO2 8-18 Zry-2 1 step - 12 h hold 7 9
DEMO-RAMP I BWR 5 14-17 UO2, dopant, hole 10-50 Zry-2 1 step(24 h hold)/steps (1h hold each) 0 5
DEMO-RAMP II BWR 8 26-29 UO2 7,6 Zry-2 1 step - 0,16 min - 24 h hold 1 7
TRANS-RAMP I BWR 5 10-14 UO2 7,6 Zry-2 1 step - 24 h hold 2 3
SUPER-RAMP (PWR) PWR 28 33-45 UO2, Gd, LG, hole 5-22 Zry-4 1 step - 12 h hold 9 19
TRANS-RAMP II PWR 6 31-32 UO2 5-7 Zry-4 1 step - 26 - 415 s hold 3 3
TRANS-RAMP IV PWR 7 15-16 UO2 11-15 Zry-4 1 step (or 2 steps) - 38-2270 s hold 6 1
OVER-RAMP PWR 39 12-31 UO2 4,5-22 Zry-4 1 step - 24 h hold 14 25
Fuel rod Number of
Table 4 DEMO-RAMP I test rods characteristic with selected metallographic examination data. ([15])
Rod No. CPL RTL Hold time Type of section Distance from Local LHR Central hole diameter Fuel swelling
W/cm W/cm h rod bottom; mm W/cm mm494 275 460 24 Longitudal 259-286 450 3,32 almost none495 275 515 24 Longitudal 249-276 510 1,1 large496 275 610 24 Longitudal 335-362 535 0,4 (at local filling); 2,5 very large; moderate
24 Transverse 238 605 1,5-3 moderate to large497 275 605 24 Transverse 228 605 2,7 moderate498 300 600 6 x 1
3.1.3. Kola-3 cases calculated using TRANSURANUS code As the fuel rods for the Kola3/MIR experiments RAMP, FGR1 and FGR2 were refabricated, the restart option of the TRANSURANUS code is used. The refabricated rods are 40 – 90 cm long segments of the mother rods. The calculation of the base irradiation history was performed for both the full length mother rod and for a rod with a length and power history set to correspond to the segment used in the refabrication. At the end of the base irradiation, following quantities were compared between the short rod and the corresponding segment of the mother rod:
• average burnup
• average fuel/cladding gap
• average outer cladding and pellet radii
• amount of the fission gas
o in the grains
o on the grain boundaries
o in the HBS pores
o released
The comparison was done by an automated script; in case the results differed the length of the plenum in the short rod was changed until matching results were obtained. A plenum length of 70-75 mm worked well for all the rods. In order to automatically prepare the restart conditions of the refabricated fuel rods the TRANSURANUS code has been modified to:
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• read the restart file
• open the file with data of refabricated fuel rods
• read following values from the file:
o free volume in the RFR
o composition of gas mixture within the free volume of the RFR
o gas pressure after the refabrication
o temperature of the refabrication
• modify the gas mixture composition and amount of gas and adjust rod length in such way that the total free volume within the rod was equal to the prescribed free volume.
A set of scripts was also prepared to allow for the parametric calculations (rod fabrication parameters changed within the given tolerances, uncertainty factors applied to linear heat rate and fast neutron flux).
Standard NRI Rez plc input models for WWER fuels with E110 cladding were used (some of them are listed in Table 5 below). In order to assess the uncertainties in linear heat rate two sets of power histories were created – one using the nominal data, the other using a global 1.1 modifier for both the base irradiation and the ramp. Parametrical calculations were done in order to assess the influence of selected models that are specified in Table 5 where are listed changes from the standard (k00).
Table 5 TRANSURANUS input model changes for parametrical calculations.
FGR model Fuel
cracking (no. of cracks)
Grain growth model
Fuel relocation model
Cladding yield strength
k00 URGAS model Constant grain boundary concentration, no burst release
15 Ainscough and Olsen model
Mod. Frapcon model
Default E110 correlation (unirradiated)
k01 Burst release from grain boundaries on power ramp
k02 Burst release from grain boundaries and HBS on power ramp
k03 FORMAS model
k04 FORMAS model
Burst release from grain boundaries on power ramp
k05 Zry-4 correlation
k06 Not modelled
k07 0
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k08 Burst release from grain boundaries on power ramp
0 Eberle & Stackmann, 1997
k09 Burst release from grain boundaries on power ramp
4 Gapcon-Thermal 3 model
k10 4 Gapcon-Thermal 3 model
k11 Burst release from grain boundaries on power ramp
4 Gapcon-Thermal 3 model
Zry-4 correlation
k12 Burst release from grain boundaries on power ramp
4 Mod. Frapcon model
k13 4 Mod. Frapcon model
k14 Burst release from grain boundaries on power ramp
4 Mod. Frapcon model
Zry-4 correlation
k15 FORMAS model
Burst release from grain boundaries on power ramp
4 Gapcon-Thermal 3 model
3.1.3.1. Results of the calculations
The calculations with the nominal power history have following features:
• FGR is highly underestimated for the short RAMP experiments. Comparison of the on-line measured and calculated pressure within RFR-41 and RFR-48 indicates that the FGR is similarly underestimated for the first ramp of the FGR1 experiment.
• For RFRs 33-38 about 10% of actual release is predicted
• For RFRs 32 and 48 about 20% of actual release is predicted
• For RFRs 41 and 52 about 50% of the release is predicted
• The results are similar for URGAS and Forsberg-Massih models (k01, k04)
Table 6 shows the availability of additional fission gas in RFR-41 (calculation k09 nom) compared with the measured amount of the gas released. It can be clearly seen that the “missing” fission gas must be released from the grains, not the grain boundaries or high burnup structure (some 3 mmols). It can be also seen from the calculations with increased LHR that even 10% in the LHR increase over the nominal values does not shift the calculated FGR near the experimental one for the short RAMP experiments and for the first power ramps of FGR1 experiment (see Figure 10 and Figure 11), the FGR in short ramps is generally underprediceted in the TRANSURANUS.
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Figure 4 Ratio of calculated and measured burnup (RFR average) after the base irradiation. Horizonatal axis = RFR number, Vertical axis = case designation, “kx” marks the selection of code models (see section 3.1.3), suffix “nom” marks the nominal power histories, suffix “1.1x” marks the calculation with LHR increased by 10% for both base irradiation and the ramp.
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Figure 5 Ratio of calculated and measured FGR.
17
Figure 6 Ratio of calculated and measured Xe content in the free gas.
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Figure 7 Concentration of Xe in the grains, RFR-41, k01 nominal, after ramp (please note that the visualisation is not correct at the pellet centre – the values should be zero at r < rcentral hole).
Table 6 Gas concentrations – free gas in RFR-41, case k09_nom.
Calculated concentration [/] Measured concentration [/] He 0.496 0.380 Kr 0.053 0.058 Xe 0.451 0.554 N - 0.008
Table 7 FG distribution in RFR 41.
Exp. Data
Calculation k09 Nominal, after ramp
Nominal, before ramp
1.1xLHR, before ramp
Free gas total [cc@STP] 307.6 [umol] 12711 9443
He [umol] 4826 4683 4683 FG in grains [umol] 15297 19346 21244 FG on grain boundaries [umol] 12 62 58 FG in HBS pores [umol] 44 17 77 Two further interesting points can be noted when the calculated average burnup after the base irradiation is compared to the experimental values (Table 1). Firstly, the burnup is consistently underestimated with the nominal power history – this might be caused by the overall underestimation of the LHR in the base irradiation data or it can be caused by the methodology used to covert Cs-137 counts to burnup – it would therefore be interesting to obtain more information on the Bu determination. Note that there are slight differences in the calculated burnup between the calculations with different models, although the initial fuel mass and the power history were exactly the same. This is because in
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TRANSURANUS local quantities, which should be related to a mass of fuel volume element considered in calculation (fission rate, fission products generation and depletion, burnup, …) are in fact related to the volume of the element only (i.e. the units are xx/mm3). As the volume is not conserved, some small numerical glitches may occur, Figure 13 and Figure 14 show the calculated distribution of Xe along the fuel pellet radius. When compared with the PIE results (Table 1), one difference is obvious - the Xe concentration should remain low until ~2.7mm, but according to the calculation it starts to rise already at ~2.0mm. This confirms that the “missing” gas should be coming from the grains and that the diffusion from the fuel matrix is undrepredicted. Comparison of measured and calculated temperature is ambiguous – the temperature is underestimated in the case of RFR-50 by some 350°C, but overestimated in the case of RFR-51. With the LHR increased by 10% the temperature remains underestimated in the case of RFR-50. It is interesting to note that both thermocouples failed at similar measured temperature - about 1500 °C (Figure 8 and Figure 9).
Figure 8 Comparison of measured and calculated temperatures, RFR-50, k09, 1.1xLHR (right) and nominal power history (left).
Figure 9 Comparison of measured and calculated temperatures, RFR-51, k09, 1.1xLHR (right) and nominal power history (left).
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Figure 10 Comparison of measured and calculated pressure, RFR-41, k09, 1.1xLHR (right) and nominal power history (left).
Figure 11 Comparison of measured and calculated pressure, RFR-48, k09, 1.1xLHR (right) and nominal power history (left).
Calculated porosity distribution is shown in Figure 13 and Figure 14 (the pre-ramp porosity was the same in all axial segments). Comparison with the PIE data reveals that the calculation predicts the final porosity quite well.
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Figure 12 Concentration of Xe in the grains, RFR-41, k01 nominal, after ramp.
Figure 13 Post-ramp porosity, RFR-41, k01 nominal (please note that the visualisation is not correct at the pellet centre – the values should be zero at r < rcentral hole).
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Figure 14 Porosity, RFR-41, k01 nominal.
However, there is one major discrepancy between the calculations and experiment – none of the calculations predicts significant closing of the central hole in the fuel pellet. Different combinations of the fuel creep/number of cracks options have been tested, but the central hole radius remains essentially the same (Figure 15). As a next step we have tried to disable the central void formation model in order to confirm that the generated swelling strain is not transferred to the central hole instead of causing its closure. The results for RFR-41 are presented in Figure 17 - the effect of this option is negligible. However, it can be seen that the time step length effects the predicted central hole closure. Further calculations with different time step length were performed, the results are shown in Figure 18. Note that the only difference between the depicted calculations was the time step length. The different time step length was obtained by:
• Setting by the limitation of maximum change of LHR bwteen the steps – in this case the time step length is determined by the code to comply with prescribed limit (cases labeled “dQl” in Figure 18.
• Using short time steps in the input file, the cases are labeled dT in Figure 18. In this case, whole MIR irradiation history was manually input with steps of fixed length (LHR, flux and temperature values were interpolated between the values in the original file). One drawback of the current coding of the transient FGR model in TRANSURANUS can be seen in Figure 19. Since the LHR is input in very short time steps, the condition for the burst release, whis is defined as an absolute change in LHR during the time step regardless of its length, is never met, although the power ramp is physically the same as in the other cases.
Apparently, the central hole closes from the diameter of 1.6mm to just over 1 mm, but the code behaviour is highly unstable. The results do not seem to be converging to any
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“physical” value. This is caused by the fuel creep model, once disabled the results are stable, but the central hole dimensions do not change much. At the same time it can be seen that higher number of the fuel cracks might result in slightly higher predicted FGR, but the calculated cladding diameter and gap changes during the ramp are wrong – the fuel strain is used up to heal the cracks. In result, there is no residual cladding deformation (Figure 16). The overall results related to the pellet and cladding dimensional changes during the transient are shown in Figure 20 - Figure 27. The gap and cladding outer diameter changes can be predicted quite well (k08, k09, k10). However, one should bear in mind that the internal pressure is underpredicted – in order to make a final conclusion regarding dimensional changes, a calculation with higher internal pressure would have to be performed.
Figure 15 Central hole radius, RFR-41, 8th axial segment (LHRmax=44.1kW/m), k09 nom, different fuel creep / number of cracks combinations.
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Figure 16 Fuel-cladding gap, RFR-41, k09 nominal with different fuel creep / number of cracks combinations.
Figure 17 Central hole radius, RFR-41, 8th axial segment (LHRmax=44.1kW/m), k09 nom, effect of the central void model option.
25
Figure 18 Central hole radius, RFR-41, 8th axial segment (LHRmax=44.1kW/m), k09 nom, effect of the time step length.
Figure 19 FGR, RFR-41, k09 nom, effect of the time step length.
26
Figure 20 Diameter of the central hole, cold state, before ramp.
27
Figure 21 Diameter of the central hole, cold state, after ramp.
28
Figure 22 Fuel – cladding gap (diametric), FA-198 rods, cold state, before ramp.
29
Figure 23 Fuel – cladding gap (diametric), FA-222 rods, cold state, before ramp.
30
Figure 24 Fuel – cladding gap (diametric), FA-198 rods, cold state, after ramp.
31
Figure 25 Fuel – cladding gap (diametric), FA-222 rods, cold state, after ramp.
32
Figure 26 Max. change of RFR diameter during the ramp, FA-198 rods.
33
Figure 27 Max. change of RFR diameter during the ramp, FA-222 rods.
3.1.4. Kola-3 cases calculated using FEMAXI-6 We have simulated irradiation history of fuel rods from FA-198 and FA-222 irradiated in Kola-3 NPP using fuel performance code FEMAXI-6 (the code was used as is, some parameters and models are slightly tuned (but kept the same for all WWER calculations!) to get reasonable results for WWER fuel based on Kola-3 (FA-198 and FA-222), Novovoronezh (FA-4108) and Zaporozhye (FA-0325) IFPE data sets). 16 fuel rods from both FAs (FA-198 and FA-222) were calculated. Some segments of the rods were consequently irradiated in the MIR reactor. Calculation strategy was used as follows:
- base irradiation of full length fuel rods (16 fuel rods from FA-198 + 16 fuel rods from FA-222) – results shown in figures Figure 28 to Figure 31;
- base irradiation + ramp irradiation of ramped segments (9 segments) – results shown in table 1 and figures Figure 32 and Figure 33.
The results of calculations are compared with measured values after base irradiation and after the ramp and presented in the following figures and table. Figure 28 – Figure 33 compares calculated and measured values after base irradiation of FA-198 and FA-222. Average burnup shows relatively good agreement – the difference is less than 5 % – Figure 28. The difference between measured and calculated cladding creepdown (Figure 29) is higher and the calculated values do not differ so much as
34
measurement due to very similar irradiation histories. Cladding elongation is simulated with similar differences (Figure 30). Plenum pressure comparison (Figure 32) shows very good agreement for lower FGR values (Figure 33). Pellet cladding gap is calculated as closed after base irradiation for all segments – Table 8. After ramp it is modelled as open for some cases with the values smaller than measured, but the overall trend of calculated gap sizes is comparable to measurement. The reason of underestimation also lies in the FGR underprediction – Table 8 – and consequently in inner gas pressure underprediction (Figure 34 and Figure 35). FGR is calculated in the range of 0,17 to 0,67 (Table 8) of the measured values thus roughly only the half of the measured FGR is predicted. Calculated fuel temperatures are higher (of about 300 °C) for rod 51 (Figure 36) and lower (of cca 150 °C) for rod 50 (Figure 37). The trend of these results (with respect to temperatures and FGR/inner pressure) is in good consistency with TRANSURANUS results presented in the previous chapter (3.1.3).
40
45
50
55
60
40 45 50 55 60
Measurement [MWd/kdU]
Cal
cula
tion
[MW
d/kd
U]
FA-198 Ave bu - F6
FA-222 Ave bu - F6
1:1
+10%
-10%
Figure 28 Calculated (FEMAXI-6) and measured average burnup for fuel rods in FA-198 and FA-222 after base irradiation.
35
20
40
60
80
100
20 40 60 80 100
Measurement [microns]
Cal
cula
tion
[mic
rons
]
FA-198 Clad creep - F6
FA-222 Clad creep - F6
1:1
+15%
-15%
Figure 29 Calculated (FEMAXI-6) and measured cladding creepdown for fuel rods in FA-198 and FA-222 after base irradiation.
10
12
14
16
18
10 12 14 16 18
Measurement [mm]
Cal
cula
tion
[mm
]
FA-198 Elongation - F6
FA-222 Elongation - F6
1:1
+15%
-15%
Figure 30 Calculated (FEMAXI-6) and measured cladding elongation for fuel rods in FA-198 and FA-222 after base irradiation – measurement are taken as a linear fit from the Figure 31.
36
Figure 31 Measured elongation of WWER fuel rods. Figure taken from [12].
0,7
0,9
1,1
1,3
1,5
0,7 0,9 1,1 1,3 1,5
Measurement [MPa]
Cal
cula
tion
[MP
a]
FA-198 Inner pressure - F6FA-222 Inner pressure - F61:1-15%+15%
Figure 32 Calculated (FEMAXI-6) and measured inner gas pressure for fuel rods in FA-198 and FA-222 after base irradiation.
37
0
1
2
3
4
0 1 2 3 4
Measurement [%]
Cal
cula
tion
[%]
FA-198 FGR - F6FA-222 FGR - F61:1-15%+15%
Figure 33 Calculated (FEMAXI-6) and measured fission gas release for fuel rods in FA-198 and FA-222 after base irradiation.
Table 8 Measured and calculated (FEMAXI-6) diametral gap afer base irradiation and FGR and diametral gap after ramp.
Unit
Rod number - 33 37 38 41 32 48 51 50 52Number of mother rod/FA - 76/198 5/222 3/222 99/198 6/222 2/222 20/198 25/222 46/222Instrumentation - No No No PT No PT TC* TC* No
Before rampMeasured pellet-cladding gap µm4-15 0-7 0-7 4-15 0-7 0-7 4-15 0-7 0-7Calculated gap (diametral) µm 0,3 0 0 0,3 0 00 0 0
After rampFGR measured % 31,3 16,9 19,6 47,5 46,6 50 - - 48,37FGR calculated % 7,3 4,5 3,3 25,9 13,9 26,2 42,1 23 32,3Calculated/measured - 0,23 0,27 0,17 0,55 0,30 0,52 - - 0,67Measured pellet-cladding gap µm40 - 65 15 - 33 15 - 33 40 - 65 15 - 33 15 - 3340 - 65 15 - 33 15 - 33Calculated gap (diametral) µm 2 - 13 0 0 12 - 36 0 - 9 0 - 8 32 - 40 0 - 8 1,5 - 9
Test RAMP Test FGR-1 Test FGR-2
38
0
2
4
6
8
10
12
14
16
18
0 100 200 300 400 500 600 700 800 900 1000
Time (hours)
Inne
r ga
s pr
essu
re (
MP
a)Experiment
Calculation FEMAXI-6
Figure 34 Measured and calculated (FEMAXI-6) inner gas pressure during ramp FGR-1 rod 41.
0
2
4
6
8
10
12
14
16
18
0 100 200 300 400 500 600 700 800 900 1000
Time (hours)
Inne
r ga
s pr
essu
re (
MP
a)
Measured
Calculated FEMAXI-6
Figure 35 Measured and calculated (FEMAXI-6) inner gas pressure during ramp FGR-1 rod 48.
39
0
500
1000
1500
2000
2500
0 50 100 150 200 250 300 350 400
Time (hours)
Tem
pera
ture
(°C
)Calculated temperature - FEMAXI-6
Measured temperature
Figure 36 Measured and calculated (FEMAXI-6) fuel temperature during ramp FGR-2 rod 51.
0
500
1000
1500
2000
2500
0 50 100 150 200 250 300 350 400
Time (hours)
Tem
pera
ture
(°C
)
Calculated temperature - FEMAXI-6
Measured temperature
Figure 37 Measured and calculated (FEMAXI-6) fuel temperature during ramp FGR-2 rod 50.
40
3.1.5. FRAPCON 3.4 calculation FRAPCON 3.4 was used for the calculations Kola Mir experiment RFR-41 only because the results obtained were similar to TRANSURANUS and FEMAXI ouputs and, as discussed above, the correct prediction of FGR cannot expected without modeling of the closure of the pellet central hole. Short segment corresponding to the RFR rod modeled for the base irradiation with the plenum volume and fill pressure adjucted to gove the correct values after the end of the base irradiation. Fuel centerline temperature calculated with this approach was again compared with a full-length rod calculation (correct initial fill pressure and free volume) in order to validate the approach. Fill gas composition change after the refabrication was not modeled – ramps start with ~ 5% Xe. M5 was used instead of E110, which is not available in FRAPCON 3.4. Massih model was used for FGR calculations. Calculated results compared to experimental values after base irradiation are listed in Table 9 and after the RFR – 41 MIR FGR-1 experiment in Table 10. It can be seen the agreement after base irradiation is quit satisfactory, but after the FGR-1 experiment the calculated FGR is underpredicted more than of one half (19,8 % to 48,9 %). The calculated maximum temperature in the fuel during base irradiation is plotted in Figure 38. During the experiment the internal gas pressure was measured and its comparison to calculated one is plotted in Figure 39 that reflects the underprediction in FGR in underprediction inner gas pressure, at least the calculated dynamics corresponds to measurement.
Table 9 Experimental values and calculated (FRAPCON 3.4) results of RFR41 after base irradiation results.
Experiment Frapcon - M5/Massih
Bu [MWd/kgU] 48.9 48.5 FR diameter [mm] 9.07 9.08 Radial gap [um] (structural) 4-17 15
Base irradiation FGR [%] 0.5-1*[1] 1.2
Refabrication pressure [MPa] 2.0 (He) 2.03
He fraction 100% 95% Free volume [cc] 5.84 5.86
*Whole rod value
41
0
100
200
300
400
500
600
700
800
900
1000
0 5000 10000 15000 20000 25000 30000
Time [h]
Max
.Tem
pera
ture
[C]
Figure 38 Calculated (FRAPCON 3.4) max. temperature vs. time.
Table 10 Measured and calculated (FRAPCON 3.4) results of RFR – 41 MIR FGR-1 experiment.
Experiment Frapcon M5/Massih
FGR [%] 47.5 19.8 Xe fraction [%]*[1] 55.4 42.2 Max radial gap [um] (structural) 20-33 16
Max DD [um] 26-33 13.1 EOL cladding hoop strain [%] (-0.078)
(-0.153) -0.28
Central hole diameter [mm] 0.6 (re-opened)
No change modeled
EOL cold pressure [MPa] 4.17
*Not corrected for the pre-ramp content
42
0
2
4
6
8
10
12
14
16
18
0 100 200 300 400 500 600 700 800 900
Time [h]
Inte
rnal
Pre
ssur
e [M
Pa] Measured
Predicted
Figure 39 Calculated (FRAPCON 3.4) and measured internal pressure RFR-41.
3.2. AREVA “ IDEALIZED CASE” AREVA "Idealised Case" for high burnup fuel ([16]) is an idealized version, which is based on measurements for three rods operated for 3, 4 and 7 cycles in a commercial french PWR reactor. Those rods experienced very similar power histories, i.e. the 3 and 4 cycle rods match very closely the first 3 and 4 cycles, respectively, from the 7 cycle rod, allowing three FGR 'measurements' for a single power history. The maximum fuel rod burnup is about 81.5 MWd/kgHM with an FGR of about 9 %. The given FGR uncertainties take care for measurement and fabrication unertainties (including the 'idealization' of the case). The given characteristics of fuel rod, cladding and pellet are listed in Table 11.
43
Table 11 Given fuel rod characteristic for AREVA “idealized case”. [16]
Fuel rodLength of active zone 3650.0 mmPlenum volume 8.04 cm3
Fill gas HeFill gas pressure 1.6 MPa (absolute)Flow area of rod 87.8 mm2
CladdingMaterial Zy4 (stress-relieved)Inner diameter 8.25 mmOuter diameter 9.5 mmPelletMaterial UO2
Enrichment 4.5% U-235Density 95.0% of theoretical densityDiameter 8.085 mmLength 13.25 mmChamfer height 0.27 mmChamfer width 0.5425 mmDishing radius 3.0 mmDishing depth 0.31 mmDishing Volume 8.8 mm3 (both pellet sides)
3.2.1. TRANSURANUS results The power history was used as submitted for this AREVA “idealized case”. Another time history of linear heat rate as constant value 100 W/cm (this corresponds to FUMEX-II case 27-2a) was calculated with the same models and characteristics as for the AREVA “idealized case” for comparison – both LHR histories are shown in Figure 40. It is interesting that calculated FGR for both LHR histories (Figure 41) meets the expected value at ~80 MWd/kgU although the calculated maximum temperatures differ significantly (Figure 42). This confirms that the FGR predicted is purely athermal release from.
44
Figure 40 AREVA idealized LHR (black) and uniform LHR 100 W/cm (red) vs. burnup.
Figure 41 Calculated (TRANSURANUS) FGR.
45
Figure 42 Calculated (TRANSURANUS) maximum temperature.
3.2.2. FEMAXI-6 results FEMAXI-6 was again used as is, i.e. without any modification, with models optimized for PWR calculations. The calculated FGR, inner gas pressure and average burnup as a function of time are depicted in the Figure 43. It can be seen that final calculated FGR (13 %) is slightly higher than upper bound of expected FGR value (11,5 %). On the other hand, the calculated average burnup is lower than given value (79,3 vs. 81,5 MWd/kgU). Constant LHR history (100 W/cm) was also calculated for comparison and resulting FGR and inner gas pressure as a function of burnup are shown in the Figure 44. The final FGR value lies under 2 % that much lower than in the calculation with TRANSURANUS code (previous chapter 3.2.1), although the temperatures seem to be consistent – Figure 45 – between 900 and 1100°C for the idealized history and 600 – 700 °C for constant 100 W/cm power.
46
0
2
4
6
8
10
12
14
16
18
0 500 1000 1500 2000 2500
Time (days)
FG
R (
%),
Inne
r ga
s pr
essu
re (
MP
a)
0
10
20
30
40
50
60
70
80
90
Ave
rage
bur
nup
(MW
d/kg
U)
Fission gas release (%)
Inner pressure (MPa)
FGR - expected
Average burnup (MWd/kgU)
Figure 43 Calculated (FEMAXI-6) FGR, inner gas pressure and average burnup vs. time.
0
2
4
6
8
10
12
14
16
18
0 10 20 30 40 50 60 70 80 90 100
Average burnup (MWd/kgU)
FG
R (
%);
Inne
r ga
s pr
essu
re (
MP
a)
Fission gas release (%)
Inner gas pressure (MPa)
FGR - expected
FGR - 100w/cm
Inner pressure - 100 W/cm
Figure 44 Calculated (FEMAXI-6) FGR and inner gas pressure vs. average burnup.
47
0
200
400
600
800
1000
1200
0 10 20 30 40 50 60 70 80 90 100
Average burnup (MWd/kgU)
Tem
pera
ture
(°C
)
0
50
100
150
200
250
Ave
rage
LH
R (
W/c
m)
Pellet center temperature (°C)
Pellet center temperature - 100W/cm
Average LHR (W/cm)
Average LHR - 100 W/cm
Figure 45 LHR and calculated (FEMAXI-6) pellet center temperature (in the middle of rod length) vs. average burnup.
3.3. US-PWR 16X16 LTA US-PWR 16x16 LTA (lead test assembly) extended burnup demonstration program [17] was conducted during the 1980s with the objective to demonstrate improved nuclear fuel utilization through more efficient fuel management and increased discharge burnup. The use of the 16x16 LTAs with Zr-4 cladding in this program demonstrated the capability to achieve peak fuel rod average burnups of ~ 60 GWd/MTU. Both poolside (non-destructive) and hot cell (destructive) post irradiation examinations (PIE) of selected rods from the two LTAs were conducted. These examinations included rods irradiated for 3 and 5 cycles. The irradiation of two 16x16 LTAs was completed in a US commercial PWR. The standard fuel rod design consists of enriched UO2, solid cylindrical pellets. In addition to the standard design fuel rod, three additional design concepts were included in a limited number of rods in the two LTAs. These were: 1) an annular fuel pellet design, 2) large grain size pellets (35 micro-m as opposed to the nominal 7 to 12 micro-m standard pellet design), and 3) cladding with graphite coating (~ 8 micro-m thickness) on the interior surface. [17] We have chosen two fuel rods (priority cases) TSQ002 (standard fuel rod) and TSQ022 (fuel rod with hollow pellets) for our calculations. Basic characteristics of fuel rods and selected EOL data are summarized in the Table 12. The calculation results (with comparison to measured data) are summarized in the following chapters (3.3.1 and 3.3.2).
48
Table 12 BOL and EOL characteristic data of fuel rods TSQ002 and TSQ022. [17]
Fuel rod No. TSQ002 TSQ022
Inner fuel diameter, mm 0 2.34Outer fuel diameter, mm 8.255 8.255Pellet length, mm 9.91 9.91Chamfer and dish yes yesFuel stack length, mm 3810 3810Enrichment, % 235U 3.48 3.48Cladding inner diameter, mm 8.43 8.43Cladding outer diameter, mm 9.7 9.7Fill gas pressure, MPa 2.62 (He) 2.62 (He)Initial free volume, ml 25.42 37.22Rod average burnup, GWd/MtU 53.24 58.12EOL free volume, ml 17.8 31∆ free volume, ml * -7.62 -6.22∆ gas volume (EOL-BOL), % 5.7 2.6
* in the report [17] the unit of ∆ of void volume is in % but the numbers correspond to pure diference (in ml) 3.3.1. TRANSURANUS results The same models as in previous AREVA “idealized case” were used. Calculated maximum tempertures as a function of time for both rods are shown in Figure 46. Calculated FGR time evulotion for both rods is depicted in Figure 47. The influence of some input parameters: in Figure 48 diffusion coefficients according to Matzke or Turnbull are used, and the the zero capacity of grain boundaries for FGR (igrbdm=0) is considered in Figure 49. The temperatures are higher for the solid fuel pellets (TSQ002), but the FGR is higher for fuel rod with hollow pellets (TSQ022). The calculated gas volume change is for rod TSQ002 predicted well, but for the rod TSQ022 overestimated. Calculated average burnup for both rods was slightly lower than measured values (51.6 GWd/MtU vs. 53.24 GWd/MtU for rod TSQ002, resp. 56.8 GWd/MtU vs. 58.12 GWd/MtU for rod TSQ022). Calculated maximum oxide layer thickness is 50 microns that corresponds very well to measured values (measured profiles are shown in Figure 52 and Figure 53).
49
Figure 46 Calculated (TRANSURANUS) maximum fuel rod temperature of rods TSQ002 and TSQ022.
Figure 47 Calculated (TRANSURANUS) fission gas release of rods TSQ002 and TSQ022.
50
Figure 48 Calculated (TRANSURANUS with two different diffusion coefficients) fission gas release of rods TSQ002 and TSQ022.
Figure 49 Calculated (TRANSURANUS with different igrbdm values) fission gas release of rods TSQ002 and TSQ022.
51
Figure 50 Calculated (TRANSURANUS) and measured gas volume change of rods TSQ002 and TSQ022.
3.3.2. FEMAXI-6 results The input models were set in the same way as in the AREVA “idealized case” calculation. Calculated FGR was 5,56 % (TSQ002) and 6,2 % (TSQ022) that is significantly higher than calculated FGR with TRANSURANUS (Figure 47). Calculated average burnup values are close to measurements (52.68 vs. 53.24 for rod TSQ002 and 57.96 vs. 58.12 for rod TSQ022). The difference of calculated void volume change for both rods is very small and the calculated EOL values are between the measured ones - Figure 51. Comparison of calculated and measured oxide layer thickness for both rods is shown in Figure 52 and Figure 53 and can be concluded that the trend is catched well but the maximum is overestimated of 15 – 20 microns. Calculated fuel centre temperatures at the middle of the rod length are shown in Figure 54.
52
-9
-8
-7
-6
-5
-4
-3
-2
-1
00 200 400 600 800 1000 1200 1400 1600 1800
Time (days)
∆ v
oid
volu
me
(ml)
Void volume - TSQ002 - FEMAXI-6 calculation
Void volume - TSQ022 - FEMAXI-6 calculation
Measurement TSQ002
Measurement TSQ022
Figure 51 Calculated (FEMAXI-6) and measured void volume change of rods TSQ002 and TSQ022.
0
10
20
30
40
50
60
70
0 5 10 15 20 25 30
Axial segment (-)
Oxi
de la
yer
thic
knes
s (m
icro
ns)
Calculated oxide thickness TSQ002
Measured oxide thickness TSQ002
Figure 52 Calculated (FEMAXI-6) and measured cladding oxide thickness of rod TSQ002.
53
0
10
20
30
40
50
60
70
0 5 10 15 20 25 30
Axial segment (-)
Oxi
de la
yer
thic
knes
s (m
icro
ns)
Calculated oxide thickness TSQ022
Measured oxide thickness TSQ022
Figure 53 Calculated (FEMAXI-6) and measured cladding oxide thickness of rod TSQ022.
0
200
400
600
800
1000
1200
1400
0 200 400 600 800 1000 1200 1400 1600 1800
Time (days)
Pel
let c
entr
e te
mpe
ratu
re (
°C)
Temperature - TSQ002
Temperature - TSQ022
Figure 54 Calculated (FEMAXI-6) fuel centre temperatures (in the middle of rod length) of rods
TSQ002 and TSQ022.
54
3.4. LOCA HALDEN IFA-650.1 AND 650.2 The LOCA experiments in Halden are integral in-pile tests on nuclear fuel behaviour under simulated LOCA conditions. These LOCA tests can be used for validation and development of fuel behaviour codes and for planning of further LOCA tests. The first LOCA trial test runs IFA-650.1 were carried out in the Halden reactor in May 2003 using a fresh unpressurised PWR rod with Zr-4 cladding. The main objective was to gain experience and to determine how to run the later experiments with the pre-irradiated rodlets. The details are summarized in the Halden report [18]. The second LOCA trial test was carried out in the Halden reactor in May 2004 using a fresh pressurised PWR rod with Zr-4 cladding. The main purpose of IFA-650.2 was to practice the test case with ballooning and fuel failure to find out how to run the later experiments with the pre-irradiated rods. The details of the experiment are summarized in the Halden report [19]. Basic characteric data of both tests are summarized in the Table 13. The rod is located in the centre of the the IFA-650 test rig and surrounded by an electrical heater inside the high-pressure flask. The heater is part of a flow separator, which divides the space into a central channel surrounding the fuel rod and an outer annulus. The heated length equals the fuel length, 500 mm in both cases. The heater is used to simulate the isothermal boundary conditions, i.e heat from adjacent fuel rods during a LOCA. Cladding temperature is influenced by both rod and heater power. The rod power can be controlled by changing the reactor power and by the He3-coil. The rod plenum volume (free gas volume) was made relatively large to be able to maintain stable pressure conditions also during ballooning. Larg part of total free gas volume was located outside the heated region. Locations of free gas volumes inside the rod as well as the schema of the whole rig are shown in Figure 55. In the first LOCA test IFA-650.1 there were measured several variables using three cladding surface thermocouples, one cladding extensometer, one fuel thermocouple, one pressure sensor, two heater surface thermocouples, three vanadium neutron flux detectors, two cobal neutron detectors and thermocouples at the inlet and outlet of the rig. In the second LOCA test IFA-650.2 the rig instrumentation consisted of four cladding surface thermocouples, a cladding extensometer, a fuel pressure sensor, two fast response cobalt neutron detectors and three vanadium neutron flux detectors at three elevations, two heater surface thermocouples and thermocouples at the inlet and outlet of the rig. We have done calculations using FRAPTRAN 1.4 with default (recommended) input parameters and the results with comparisons to measured values are described in the following subchapters (3.4.1 and 3.4.2).
55
Table 13 Basic characteristics of LOCA tests IFA-650.1 and IFA-650.2.
Fuel rod No. 650-1 650-2Fuel stack length, mm 500 500Burnup, MWd/kgU 0 0Fuel diameter, mm 8,29 8,29Pellet length, mm 8 8Cladding Low Tin Zr-4 Low Tin Zr-4Cladding outer diameter, mm 9,5 9,5Cladding thickness, mm 0,57 0,57Fill gas pressure, bar 2 (He) 40 (He)Plenum/Free volume, cc 15 17,4Number of runs 6 1Peak PCT, °C 843-1142 1050Hold time, min - ~4,5Max. rod pressure, MPa 0,36 7,35Temperature at failure, °C - ~800
Figure 55 Schematic LOCA test rig with instrument levels for IFA-650.2. [19]
56
3.4.1. IFA-650.1 calculations The results of calculation „IFA 650.1 Run 1“ using code FRAPTRAN 1.4 with different prescribed cladding temperatures are compared in this section. The input files were the same (using default/standard models; 18 axial nodes) differing only in prescribed cladding temperature as boundary condition:
– measured value from thermocouple TCC1 is prescribed along the whole cladding length; the FEM (mechanical response of the cladding is solved using FEM model) and FD (standard Finite Diference solver) models are used and compared; in the graphs designated as “FEM” and “FD”;
– measured value from thermocouple TCC3 is prescribed on the cladding length except the last 16 mm of active length (part above the heater) where is applied low temperature (290 °C); FD (standard Finite Diference solver) models used (in graphs designated as “frpgas_2”).
The comparisons are presented in relation to measured values that are designated as “TFDB“ in the figures.
3.4.1.1. Calculated (prescribed) cladding temperature
Cladding temperature is prescribed based on the measured values by thermocouples, the difference between TCC1 a TCC3 thermocouples is about 50 – 60°C. Two calculations were done with TCC1 temperature prescribed on the whole cladding length (FEM and FD curves). Third calculation with TCC3 temperature prescribed on the 0 to 50 cm of the cladding length and 290 °C (constant) during LOCA on 50 – 51.6 cm of cladding length, i.e. plenum (“frpgas_2” curves). The following figures (Figure 56, Figure 57 and Figure 58) confirm the agreement of the measured and prescribed cladding temperatures for calculated variations and the difference between three cladding thermocouples (TCC1, TCC2 and TCC3).
Figure 56 TCC1 and calculated (= prescribed) cladding temperature for the whole cladding length –
FEM and FD runs.
57
Figure 57 TCC3 and calculated (= prescribed) cladding temperature for the whole cladding length –
FEM and FD runs.
58
Figure 58 TCC1, TCC2 and TCC3 and calculated (= prescribed) cladding temperature for the whole
cladding length - run “frpgas_2”.
3.4.1.2. Calculated fuel temperature
Calculated fuel temperature using FEM model is significantly higher than measured fuel temperature (the reason is explained below) - Figure 59. The difference between measured temperature and calculated with FD model is 20 – 40°C and can be due to simplified prescribed boundary condition (constant temperature along the whole cladding length). Fuel temperature calculated with cladding temperature prescribed in accordance with TCC3 thermocouple lies lower than measured one (Figure 60) – this confirms that the difference is caused by the boundary condition.
Figure 59 Calculated and measured fuel temperature – runs FEM and FD.
59
Figure 60 Calculated and measured fuel temperature – run “frpgas_2”.
3.4.1.3. Calculated inner gas pressure
The inner gas pressure was calculated significantly higher than measured in all cases (0.76 MPa vs. 0.37 MPa) - Figure 61. When using FEM model, the cladding straining is faster. When the pressure gets 0.65 MPa the ballooning begins very quickly resulting in large calculated hoop strain (20% - see Figure 64) along the whole cladding length (cladding temperature is the same in all axial segments), rod free volume increase (Figure 66) and gas pressure decrease. Cladding ballooning is not evident from the measured inner gas pressure. Opening of fuel-cladding gap impedes the heat transfer and thus the fuel temperature is overprediceted with the FEM model. When the low plenum temperature (290°C) is prescribed (run “frpgas_2”), the calculated inner gas pressure gets closer to measured values – Figure 62 and Figure 63. Calculated cladding hoop strain is very low (0.5 %) – Figure 65. Calculated gas pressure change is still slightly higher than measurement, but prescribed plenum temperature seems to be chosen reasonably. The input of correct plenum temperature is thus essential for HRP LOCA calculations.
60
Figure 61 Calculated and measured inner gas pressure – runs FEM and FD.
Figure 62 Calculated and measured inner gas pressure – run “frpgas_2”.
61
Figure 63 Calculated and measured inner gas pressure – run “frpgas_2” – the same as in previous
figure but with an offset in order to get zero inner pressure at 2000 s.
Figure 64 Calculated cladding hoop strain (for all axial segments) – runs FEM and FD.
62
Figure 65 Calculated cladding hoop strain (for all axial segments) – run “frpgas_2”.
Figure 66 Calculated free gas volume in the rod – runs FEM and FD.
3.4.1.4. Calculated cladding elongation
From the comparison of calculated and measured values it can be seen that FEM model yields in unrealistic cladding elongation (20mm) – Figure 67. On the contrary, results from FD model correspond to measured values very well – Figure 68. Lower cladding
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temperature was prescribed for the run „frpgas_2“ and the phase transition is not achieved – Figure 69.
Figure 67 Calculated cladding elongation – runs FEM and FD.
Figure 68 Calculated cladding elongation – runs FEM and FD – detail of FD run.
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Figure 69 Calculated cladding elongation – run “frpgas_2“.
3.4.2. IFA-650.2 calculations The input file for LOCA IFA-650.2 case was the same (using default/standard models) as in previous LOCA 650.1 calculation with FD (standard Finite Diference solver) – 18 axial nodes used. Cladding temperature was prescribed according to measured cladding temperature from thermocouples TCC1 and TCC3 along the cladding half lengths – the profile just before the cladding burst is shown in Figure 73. The prescribed plenum temperature was lower, its time dependence in relation to measured cladding temperature is shown in Figure 71 and to measured coolant outlet temperature in Figure 72. The plenum temperature was adjusted “by hand” to obtain the desired internal pressure evolution. Therefore the results do not show how well the internal pressure is predicted. The time dependency of cladding temperature in all (18) axial levels is shown in Figure 75 and calculated fuel centerline temperatures in all axial segements during the LOCA are in Figure 76. Calculated and measured internal pressures are plotted in Figure 70 that can also be used as an indicator of beginning of cladding ballooning and cladding burst. Measured time of rod burst is 68 s and calculated burst time 54 s, i.e. the calculation predicts burst some 15 seconds before the measurement. We can conclude that the calculation with used models and boundary conditions is conservative in respect to burst time. Prescribed cladding temperature was set without any maximum (see Figure 73) and the balloon (and burst) was in the lower axial level than measured – the view on the burst region is shown in Figure 74. The real cladding temperature profile could be calculated by thermohydraulic code and after its input to thermomechanical code the burst prediction would be improved.
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Figure 70 Calculated and measured internal pressure (calculated time of rod burst 54 s and measured
burst time 68 s).
Figure 71 Calculated (= prescribed) plenum temperature and measured cladding temperature
(TCC3).
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Figure 72 Calculated (= prescribed) plenum temperature and measured coolant outlet temperature
(TOA).
Figure 73 Calculated cladding hoop strain just before rod burst (the elevation of rod burst is in grey) and calculated (= prescribed) cladding temperature (in 100°C).
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Figure 74 Photograph of the burst area. [19]
Cladding Outside Temperature (K ) stripf.
0,00E+00
2,00E+02
4,00E+02
6,00E+02
8,00E+02
1,00E+03
1,20E+03
1,40E+03
1,60E+03
3500 3700 3900 4100 4300 4500 4700 4900 5100 5300Time (sec)
Cla
ddin
g O
utsi
de T
empe
ratu
re (K
)
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
Figure 75 Prescribed cladding outside temperature in all axial segements vs. time.
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Fuel Centerline Temperature (K ) stripf.
0,00E+00
2,00E+02
4,00E+02
6,00E+02
8,00E+02
1,00E+03
1,20E+03
1,40E+03
1,60E+03
1,80E+03
3500 3700 3900 4100 4300 4500 4700 4900 5100 5300Time (sec)
Fue
l Cen
terli
ne T
empe
ratu
re (K
)
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
Figure 76 Calculated fuel center temperature in all axial segments vs. time.
3 CONCLUSIONS This report documents calculations and comparisons of main output results to experimental data performed in the frame of FUMEX III project at the NRI Rez plc. The central hole filling due to swelling can influence the fuel rod behaviour significantly, but the used codes are not capable in modeling this effect correctly according to observation. This is obvious in Kola Mir cases. In this field we see the future need for development. For LOCA calculations we do not have thermohydraulic models that are essential for appropriate modelling of cladding mechanical response. Several calculations with different plenum temperature showed its significant influence. The cases released in the frame of FUMEX-III project helped to validate NRI Rez plc. thermomechanical codes.
69
4 REFERENCES
[1] K.J. Geelhood, W.G. Luscher, C.E. Beyer: FRAPCON-3.4: A Computer Code for the Calculation of Steady-State, Thermal-Mechanical Behavior of Oxide Fuel Rods for High Burnup (NUREG/CR-7022, Volume 1, PNNL-19418, Volume 1), published March 2011.
[2] K.J. Geelhood, W.G. Luscher, C.E. Beyer J.M. Cuta: FRAPTRAN 1.4: A Computer Code for the Transient Analysis of Oxide Fuel Rods (NUREG/CR-7023, Volume 1, PNNL-19400, Volume 1), published March 2011.
[3] W.G. Luscher and K.J. Geelhood: Material Property Correlations: Comparisons between FRAPCON-3.4, FRAPTRAN 1.4, and MATPRO (NUREG/CR-7024, PNNL-19417), published March 2011.
[4] Suzuki, M., Saito, H. (2005). Light Water Reactor Fuel Analysis Code FEMAXI-6 (Ver.1) JAEA, Japan, JAEA-Data/Code, 2005-003.
[5] K. Lassmann, TRANSURANUS: a fuel rod analysis code ready for use, Journal of Nuclear Materials 188 (1992) 295-302
[6] T. Tverberg: OECD HRP CD with data and documentation for the two first LOCA tests that took place in May 2003 and May 2004. A version submitted to the FUMEX-III, prepared in June 2010.
[7] Federal State Unitary Enterprice A.A. Bochvar All-Russia Research Institute of Inorganic Materials, The tests of VVER-440 refabrication rods in reactor MIR (working materials), 2005 (Included in the IFPE Kola3-Mir dataset).
[8] A. Smirnov, et al. Fission Gas Release From High Burnu-up VVER-440 Fuel under Steady State and Transient Operation Conditions, Fission Gas Behaviour in Water Reactor Fuels, Seminar Proceedings, Caradache, France, 2000.
[9] A. Smirnov, et al. VVER-440 Fuel Behaviour Under Transient Conditions, HRP-349/43, EHPGM, 1998
[10] Smirnov A.V. et al. Experimental Support of VVER-440 Fuel Reliability and Serviceability at High Burnup. In: Proc. of an international seminar "VVER Reactor Fuel Performance, Modelling and Experimental Support". 7-11 Nov. 1994, St.Konstantine, Varna, Bulgaria. Pp. 141-146.
[11] Lemechov S.E., Smirnov A.V, and Tsibulya V.A., "Kola-3 High Burn-up Fuel Validation Tests FA-198 and FA-222." Paper F-1.1 presented at the Enlarged Halden Programme Meeting on High Burn-up Fuel Performance, Safety and Reliability and Degradation of In-core Materials and Water Chemistry Effects, Loen, Norway, 19 - 24 May 1996.
[12] Smirnov A.V. e.a. The Peculiarities of the WWER-440 Fuel Behavior at Higher Burnups. In: Proc. of the Second International Seminar "WWER Reactor Fuel Performance, Modelling and Experimental Support" (299 p.). 21-25 April 1997, Sandanski, Bulgaria. Pp. 58-65.
[13] Smirnov A.V. e.a. Behaviour of WWER-440 and WWER-1000 Fuel in a Burnup Range of 20-40 MWd/kgU. "WWER Reactor Fuel Performance, Modelling and Experimental Support" (299 p.). 21-25 April 1997, Sandanski, Bulgaria. Pp. 40-46.
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[14] Solonin M. e.a. WWER Fuel Performance and Material Development for Extended Burnup in Russia. "WWER Reactor Fuel Performance, Modelling and Experimental Support" (299 p.). 21-25 April 1997, Sandanski, Bulgaria. Pp. 48-57.
[15] Final Report The DEMO-RAMP I Project Studsvik-STDRI-18 (August 1983).
[16] J. Killeen, AREVA idealized case no. II, M. Sperlich, E-mail to FUMEX-III participants, 9.5.2011.
[17] W. F. Lyon: US-PWR 16x16 LTA Extended Burnup Demonstration Program Summary File, Revision 1, March 2005.
[18] V. Lestinen, “LOCA testing at Halden, first experiment IFA-650.1”, HWR-762, OECD Halden Reactor Project, March 2004.
[19] M. Ek, “LOCA testing at Halden, second experiment IFA-650.2”, HWR-813, OECD Halden Reactor Project, August 2005.
5 APPENDIX Tabulated results (FGR, pellet center temperature, burnup, inner gas pressure) of Kola 3-Mir cases, Areva idealized case and US-PWR 16x16 LTA.
