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Institute for Nuclear Research Pitesti, Romania
IMPROVEMENT OF ROFEM AND CAREB FUEL BEHAVIOUR CODES AND UTILISATION OF
THESE CODES IN FUMEX 3 CRP
G. Horhoianu INR Pitesti,Romania
Final Report of the IAEA Research Project 14974/RO
December 2011
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Contract Number: 14974/RO Title of Proiect: IMPROVEMENT OF ROFEM AND CAREB FUEL BEHAVIOUR CODES AND UTILISATION OF THESE CODES IN
FUMEX 3 CRP Institute where research was being carried out: Institute For Nuclear Research, Pitesti, Romania Chief Scientific Investigator: Dr. Grigore Horhoianu IAEA Project Officer : Dr. John Killeen Time period covered: 2008-09-01 to 2011-12-14
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Introduction. The IAEA Vienna organized a Coordinated Research Project (CRP) on
improvement the computer codes used for fuel behaviour simulation under the name: FUMEX III [1].The major research objective of this CRP would be to test and develop fuel modeling codes against experimental data and cases provided by IAEA and OECD/NEA [1,2].
Institute for Nuclear Research (INR) Pitesti participated at this CRP with ROFEM and CAREB computer codes [3, 4].The main aim of ROFEM code was to calculate fuel behaviour during steady state operating conditions [5].CAREB code was developed for fuel transients analyses such as LOCA and RIA [6].Recently both codes have been improved with new models in order to extend their capabilities [3 ].
References
1. Killeen,J.; Specific Research Objectives of FUMEX-III CRP, First IAEA Technical Meeting-FUMEX III, Vienna, December 10-12, 2008
2. Sartori,E.; CD with IFPE Data Base Selected for FUMEX-III CRP, First IAEA Technical Meeting-FUMEX III, Vienna, December 10-12, 2008
3. Horhoianu,G. at al.; Improvement of ROFEM and CAREB Fuel Behaviour Codes and Utilization of these Codes in FUMEX III Exercise, Progress Report to IAEA Research Contract No: 14974,IAEA Vienna,24 September 2008.
4. Horhoianu,G.at al; Application of ROFEM and CAREB Codes to FUMEX 3 Exercise, IAEA Technical Meeting-FUMEX III, Pisa, Italy,01-04 June,2010
5. Moscalu,D.R.; Aspects Regarding Nuclear Fuel Burnup Increase, Ph.D.Thesis, Institute for Nuclear Research,Pitesti,1997
6. Arimescu,I.; High Temperature Transients Fuel Performance Modelling, Ph.D. Thesis, Institute for Nuclear Research,Pitesti,1987
1. Objective of the INR Pitesti in FUMEX 3 exercise In order to utilize the versatile nature of the CANDU type reactors, INR-Pitesti has started a development program aiming to introduce extended burnup,Recovered Uranium (RU) fuel in the reactors under construction[1,2]. An important aspect that has been taken into account was the need for improvement of the actual fuel element design performance at higher burnup.With a target burnup of about 30 MWd/KgU, which is three times grater than the maximum burnup of the natural uranium fuel,RU fuel design is under development. In order to calculate the performance of the RU fuel at extended burnup, efforts have been focused on the improvement of the existing fuel behavior codes, using information derived from INR experimental date base and from open literature. This activity, which have been started under IAEA research contract 6197/RB is still in progress. In this process, an important stage is the validation of the improved fuel behavior modelling codes in a comparison exercise,like FUMEX 3.
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The design and operating conditions of CANDU PHWR fuel differ from LWR fuels in many respects. The fuel elements in CANDU use a thin wall collapsible sheath. They are short length (0.5m) and have no separate plenum. Because of collapsible sheath, the fuel and sheath come in contact right from the beginning of life. The CANDU fuel elements operate at considerably higher linear heat rating compared with LWR fuel elements. Higher heat rating in CANDU fuel elements causes significant fuel restructuring, which promotes new processes of fission gas release which are not present in LWR fuels. Because of closed fuel-sheath gap, the sheath creeps outwards unlike LWR fuels, which initially experience sheath creep down till gap is closed and subsequently creep out at higher burnup after the gap closure. The fuel modeling codes for CANDU fuel have to take into account the above factors. The main objective of ROFEM and CAREB codes participation in FUMEX3 exercise is the need for verifying code prediction by comparison with experimental data, in order to identify the areas in which further improvements are necessary.
References
1. Horhoianu G., Nuclear Fuel R&D Program at INR Pitesti for the Period 2006-2010, Internal Report No.7212/2005, INR Pitesti.
2. G.Horhoianu, I.Patrulescu, Technical feasibility of using RU-43 fuel in the CANDU-6 reactors of the Cernavoda NPP, Kerntechnik, volume 73, No1-2, (2008) 2. ROFEM Computer Code Main models and code capabilities ROFEM is a FORTRAN-IV computer code to predict the in-pile thermal and mechanical behavior of reactor fuel rods as a function of the reactor operating history. [1,2] The code can not treat the fuel rod behavior during fast transient conditions. The steady-state thermal calculations are used at each time step with an iterative procedure. The code assumes an axisymmetry. The code consists of two major calculation parts, the thermal analysis part and the mechanical analysis part. In the thermal analysis part, the integral behavior of a whole fuel rod is analyzed and the temperature distribution, the dimensional changes of fuel and cladding, the fission gas release, and the associated inner gas pressure are determined. Then, the results of temperature distribution and inner gas pressure are transferred to the mechanical analysis part where the localized mechanical behavior is analyzed by the two-dimensional axisymmetric finite element method. The finite element method (FEM) analysis is applied to a small part of fuel rod which is expected to have the most severe mechanical interaction. The results from the mechanical analysis part is regarded as a subcode for the analysis of localized behavior. As a function of irradiation time and axial position, the thermal analysis part calculates the following:
− The temperature distribution in the fuel and cladding; − The radial deformation of the fuel due to thermal expansion, swelling, densification, and relocation;
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− The radial deformation of the cladding due to thermoelastic,plastic and creep; − The gap or the contact pressure between fuel and cladding; − The fission gas release and inner gas pressure.
As a function of irradiation time, the FEM mechanical analysis part calculates the following: Stress and strain distributions in the fuel and cladding considering elasto-plasticity, creep, thermal expansion, fuel cracking and healing, relocation, hourglassing, densification, swelling, hot pressing, fuel-cladding mechanical interaction, inner gas pressure, and coolant pressure. The basic fuel rod geometry handled in the code consists of fuel in the form of a stack of sintered fuel pellets and a cylindrical Zircaloy cladding tube, closed at each end, with the upper plenum. The fuel consists of uranium deoxide in the form of pellets. The fuel pellets may be solid or annular and may not be dished, or be chamfered. The filling gas is considered as any composition, at any pressure, of the following four gases: helium, nitrogen, krypton, and xenon. The design and characterization data of fuel and cladding such as geometry, density, grain size, filling gas composition and pressure, and cladding type are given as input data. The fuel rod is supossed to be cooled by pressurized water. The coolant temperature and pressure are given as input data as a function of time and are assumed to be constant in axial direction. The power and flux distribution in axial direction is given as input data as a function of time. The major characteristics of ROFEM are the following:
− ROFEM can analyze the integral behavior of a whole fuel rod throughout its life, as well as the localized mechanical behavior at a small part of a fuel rod;
− Localized behavior is analyzed in detail by the two-dimensional axisymmetric finite element method;
− Elasto-plasticity, creep, thermal expansion, fuel cracking and crack healing, relocation, densification, swelling, hot pressing, heat generation distribution, fission gas release, and fuel-cladding mechanical interaction are modelled;
− A quadratic isoparametric element is used to obtain a more accurate finite element solution with fewer elements than that are required when linear elements are used.
− Contact problem between fuel and cladding is exactly treated, where the contact condition, is determined by iterative procedure;
− An implicit algorithm, which necessitate use of iteration, is applied to obtain a accurate and stable solution for non-linear problems;
− Fuel is assumed as a no-tension material. Crack healing under compression is treated as recovering its stiffness gradually to nominal value. The recovery of fuel stiffness is related to initial relocation;
− Finite element analysis is applied only to a region of half-pellet height; − The code can treat a problem of long irradiation history including power
ramps with reasonable running time.
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In the framework of IAEA Research Project 14974 the individual models included in the ROFEM code have been reviewed in order to improve these models and the predictive capabilities of the codes.In particular from the ROFEM code the folllowing models have been reviewed:
pellet cracking and pellet relocation, pellet densification and swelling, UO2 thermal conductivity degradation with burnup, fission gas release, and pellet and cladding mechanical behavior. In order to improve fission gas release predictive capability the ROFEM code has been coupled with DCHAIN5V code[3,4]. In order to analyse the CANDU fuel behavior in transient/accident conditions,the code ROFEM has been coupled with CAREB code[3].
References
[1] D.R.Moscalu ,et al,Validarea codurilor de analiza a comportarii combustibilului nuclear de tip CANDU existente in institut,Raport ICN No. 438,(1994) [2] D.R.Moscalu,Aspecte privind cresterea gradului de ardere in cobustibilul nuclear,Ph.D. Thesis (1997) [3] Horhoianu,G. at al, Improvement of ROFEM and CAREB Fuel Behaviour Codes and Utilization of these Codes in FUMEX 3 Exercise ,Progress Report to IAEA Research Project 14974/RO,INR Pitesti,June,2009 [4] M.J.F.Notley,I.J.Hastings,A Microstructure_Dependent Model for Fission Product Gas Release and Swelling in UO2 Fuel,Report AECL No.5838 (1978) 3. CAREB Computer Code Main Models andCode Capabilities
The CAREB computer code was developed to simulate the thermal-mechanical response of a fuel element during rapid, high-temperature transients [1].The model assumed is a single UO2/Zircaloy sheath element with axi-symmetric properties. Physical effects considered in the code are: - Expansion, Contraction, cracking and melting of the fuel, - Variation of internal gas pressure, during the transient, - Changes in the fuel/sheath heat transfer, - Thermal, elastic and plastic sheath deformation (anisotropic) - Zr/H2O chemical reaction effects - Beryllium assisted crack penetration of the sheath (initiated from Be-brazed appendages) The new release of the code, CAREB. 1B, extends the capability of CAREB to model of sheath failure due to oxygen embrittlement upon rewetting and effect of oxide strengthening on sheath creep [2]. Other new features improved ROFEM-CAREB interface [2]. The equations used in the model to represent these physical effects use transient boundary conditions of coolant temperature, coolant pressure and sheath/coolant heat
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transfer coefficient evaluated by thermal/hydraulic codes. The conditions at the start of the transient are obtained from steady/state fuel performance code ROFEM [2].The CAREB code monitor conditions leading to sheath failure. Several failure mechanisms are explicitly represented in the code:
• Sheath overstrain (>0.15 ) • Localized overstrain under oxide cracks • Excessive sheath creep rate (>10-1s-1) • Low ductility sheath failure (>0.004 strain) • Beryllium-assisted crack penetration • High fuel enthalpy (>838 kJ/kg UO2) • Oxygen embrittlement
Beryllium-assisted crack penetration failure mechanism specific to CANDU fuel elements. Intergranular cracking of the Zircaloy fuel sheath can occur at bearing pad and spacer pad locations of the element brought on by the penetration of a beryllium-braze alloy in the presence of an applied hoop stress[3]. The thermal transient experienced by the sheath represent a complex heat treatment which cause changes in anisotropy, annealing, grain size , phase transformation(α-β) plus other changes in sheath microstructure which effect the plastic creep behaviour of the sheath. An improved transient creep model which accounts for these transient changes in microstructure [4] allow detailed examination of the plastic deformation of the sheath which occur under a variety of postulated LOCA transients. As changes in sheath microstructure are evaluated by CAREB during transient, differences in initial sheath properties can be explicitly represented in output data. CAREB was designed to be used as self-contained code with the minimum amount of input information required for execution. However, the fuel element description at the start of the transient is best obtained directly from the steady-state companion code ROFEM, due to detailed information available. These two codes, ROFEM and CAREB, can be run sequentially to describe any arbitrary pre-transient/transient reactor irradiation history desired. The capability of the CAREB code was extensively verified through the comparison with a large number of in-reactor and out of reactor test [5]. In the framework of IAEA Research Project 14974 the individual models included in the CAREB code have been reviewed in order to improve these models and the predictive capabilities of the codes.From the CAREB code the following models have been reviewed:transient creep,clad/steam reaction, beryllium assisted crack penetration.A new model for beryllium assisted crack penetration has been included in CAREB code.
References
1. Arimescu, I., High Temperature Transient Fuel Performance Modelling, Ph.D. Thesis, Institute for Atomic Physics, Bucharest ,1987 2.Horhoianu, G. at al, Improvement of ROFEM and CAREB Fuel Behaviour Codes and Utilization of these Codes in FUMEX 3 Exercise ,Progress Report to IAEA Research Project 14974/RO,INR Pitesti,June,2009 3. Kohn, E. and Sagat, S., Beryllium Assisted Cracking of Zircaloy “, 6th Canadian Fracture Conference, Harrison Hot Springs, B.C., Canada, 1982 June
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4. Sills, H.E. and Holt, R.A., “Predicting High Temperature Transient Deformation from Microstructural Models,4th International Conference on Zirconium in Nuclear Industry, Stratford-upon-Avon (1978) 5. Ion, S., Ionescu, D.V., CAREB Verification and Validation Manual, Internal Report No.4804, INR Pitesti, March, 1996 4. The IAEA research contract No. 14974/RO In the frame of IAEA research contract No. 14974/RO the following activities have been already performed:
- Analysis of the basic models of the ROFEM and CAREB codes and adaptation of these models in order to meet FUMEX exercise requirements;
- Selection from open literature of various irradiation experiments in order to verify the code response in evaluating the behavior of different types of fuel;
- Analysis of the FUMEX cases input data received from NEA for identifying the aspects that have to be clarified or need additional information;
- Selection of the FUMEX cases that are very close to INR objective in this exercise;
- Preparation of ROFEM and CAREB input files for the preliminary runs; - Preliminary run (with some uncertainties in the input data) on FUMEX-3 case; - Verification of the input data; - Final runs for selected cases; - Analysis of released experimental results; - Comparison between code predictions and experimental results. - Technical specification for LOCA instrumented tests planed to be performed in
TRIGA research reactor. 5. Application of ROFEM and CAREB to the FUMEX 3 exercise In accordance with INR objectives for this exercise we have selected for calculations with ROFEM-2 and CAREB codes four cases: Case1: NR bundle and JC bundle,both irradiated in NRU reactor of AECL. Case 2: EC89 and EC51 in TRIGA research reactor of INR. Case 3: FIO 131 and FIO 130 LOCA tests in NRX reactor of AECL. The principal reason for this selection was the CANDU type fuel elements used in these irradiation tests, requirements of the actual version of the codes. 5.1 Case 1: JC bundle and NR bundle, both irradiated in NRU reactor (AECL, Chalk River Labs., Canada) [1]. The irradiation parameters are summarized in Table 1 [12].
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Table 1. Fuel elements irradiation parameters
Fuel Element No. Parameter
JC NR
Max. element average power (kW/m)
58.6 62.8
Rod average burnup (MWh/KgUO2)
642.5 235
Coolant H2O, CANDU conditions Coolant pressure (MPa) 9-10 Coolant temperature (oC) 300
Loop coolant consisted of light water at nominal conditions of 300C (573 K), and pressures between 9.3 and 10.0 MPa.
Nominal fabrication parameters of JC fuel elements are listed in Table 2.
Table 2 Nominal Dimensions of Bundle JC Outer Elements
Fabrication Parameters For Bundle JC
fuel form single dished pellets
enrichment %U-235 in U 1.55 pellet i/d mm 0 pellet o/d mm 12.116 dish depth mm 0.4 dish void volume (per dish) mL
0.021
pellet land width mm 0.254 pellet chamfer height mm 0.0 pellet chamfer width mm 0.0 pellet length mm 20.89 pellet density g/cc 10.65 grain size µm 9 clad material Zircaloy-4 clad inside diameter mm 12.221 clad outside diameter mm 13.098 clad inside surface coating none diametral gap mm 0.10 axial gap between mm end of fuel stack and end of sheath
1.67
fuel length mm (stack length)
480.39
10
pellets per element 23 fill gas 90% Ar, 10%
He pressure 1 atm. Calculated filling gas mL volume
2.5
Nominal fabrication parameters of NR fuel elements are listed in Table 3. Three different pellet-stack-to-end-cap types, as mentioned above, were used in the outer elements of the bundle.
Table 3 Nominal Dimensions of Bundle NR Outer Elements
No Plenum 8mm Plenum (0.35 cc)
12mm Plenum (0.58 cc)
fuel form single dished pellets
single dished pellets
single dished pellets
enrichment %U-235 in U 1.41 1.41 1.41 pellet i/d mm 0 0 0 pellet o/d mm 12.13 12.13 12.13 dish depth mm 0.66 0.66 0.66 dish volume (per dish) mL 0.033 0.033 0.033 pellet land width mm 0.440 +/- 0.190 0.440 +/- 0.190 0.440 +/- 0.190 pellet chamfer height mm 0.175 +/- 0.125 0.175 +/- 0.125 0.175 +/- 0.125 pellet chamfer width mm 0.440 +/- 0.190 0.440 +/- 0.190 0.440 +/- 0.190 pellet length mm 16.0 16.0 16.0 pellet density g/cc 10.65 10.65 10.65 grain size µm 10 10 10 clad material Zircaloy-4 Zircaloy-4 Zircaloy-4 clad inside diameter mm 12.20 12.17 12.19 clad outside diameter mm 13.11 13.10 13.12 clad inside surface coating graphite
(CANLUB) graphite (CANLUB)
graphite (CANLUB)
diametral gap mm 0.067 0.043 0.057 axial gap between mm end of fuel stack and end-cap, or plenum insert
2.59 2.62 2.58
fuel length mm (stack length)
480.36 472.19 468.37
pellets per element 30 29 29 fill gas 100% He 100% He 100% He pressure 1 atm. 1 atm. 1 atm. Calculated element voidage mL
2.0 2.1 2.7
Note: Pellet chamfer is a bevel edge.
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5.1.1 JC Fuel Bundle Bundle JC was a prototype 37-element fuel bundle for the Bruce-A Ontario Hydro reactors[1].For irradiation in the NRU reactor, the centre fuel element was removed and replaced by a central tie rod for irradiation purposes in the vertical test section. Coolant for the test was pressurized light water under typical PHWR conditions of 9 to 10.5 MPa and 300°C. The fuel elements used 1.55 wt% U235 in U uranium dioxide fuel and were clad with Zircaloy-4 material. The bundles’ elements were coated with a graphite coating. The fuel is somewhat atypical of 37 element-type fuel since the length to diameter ratio (l/d) is large (1.73) due to the pellets being ground down from a OD of 14.3 mm to 12.12 mm. The outer element average measured burnup was 642.5 MWh/kgU on discharge.The standard deviation was 7.9Mwh/kgU,and maximum and minimum values were 654 and 636 MWh/kgU respectively. Burnups were measured at the mid-plane of the four elements. Outer element powers varied between 57 kW/m near the beginning of life and 23 kW/m at discharge.Uncertainty in linear power was ±10%. Due to the long irradiation, the bundle experienced 153 short shutdowns, and 129 longer duration shutdowns. No element instrumentation was used during the irradiation. However, the bundle was subjected to extensive post-irradiation examination (PIE) that included dimensional changes, fission gas release, fuel burnup analysis, and metallography that included grain size measurement. The measured FGR and residual sheath strains are within the range that is expected for similar-operated commercial power reactors and reported in open literature [2,3]. Analysis of JC fuel behavior with ROFEM Code The power history are presented in Figure 1. It can be noticed that in the cases JC bundle the power history were of the ramps type with the maximum power at the beginning of the power history. The average linear power represent the heat rate contributed by the fuel element to the coolant over an irradiation interval. It is the integral of the power as a function of time, divided by the time period. The peak linear power is the maximum instantaneous power that occurred during the irradiation interval. The power history,originaly based only on loop calorimetry and reactor physics calculations,have been corrected for measured chemical burnup.
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0
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70
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0 5000 10000 15000 20000
Time (h)
Line
ar P
ower
(kW
/m)
Figure 1. Power history of JR case.
Figure 2 show the fuel central temperature evolution during irradiation. Figure 3 show the FGR evolution during irradiation. The calculated release show an appreciable increase at the end of irradiation and underestimate the measured data by about 18.5 cm3
(measurements:45.3-56.5cm3 ).
0
500
1000
1500
2000
2500
0 5000 10000 15000 20000
Time (h)
Tem
pera
ture
(C
)
Figure 2. Fuel central temperature vs. time
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0 2000 4000 6000 8000 10000 12000 14000 16000 18000 20000Time (h)
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sion
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ease
(m
m3)
Experimental P.I. Data
ROFEM Results
Figure 3. Fission gas release vs. time.
The prediction and measured sheath deformations are presented in Figures 4 and 5 and summarized in Table 4. The calculated data show a increase at the power ramp followed by the decrease and again slow increase during irradiation. It can be noticed the calculated values for plastic sheath strain are in good agreement with experimental data in the ridge region and superestimate the measurements in the pellet mid plan region by about 0.05 %.
0
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She
ath
Hoo
p S
trai
n at
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e (%
)
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Figure 4. Sheath hoop strain at the pellet midplane region vs. time .
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ath
Hoo
p S
trai
n at
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ge (
%)
Experimental P.I Data ROFEM Results
Figure 5. Sheath deformation at the pellet end (ridge) region vs. time .
Table 4. Comparison between ROFEM code calculation results and P.I experimental data
Fission Gas Release (cm3, STP)
Sheath Hoop Strain (%) Pellet Midplane Pellet End (Ridge)
calculated measured calculated measured1 calculated measured1
JC bundle
26.8 45.3÷56.5 0.8 0.32÷0.69 1.4 1.16÷1.76
1 Average values on axial direction 5.1.2.NR Fuel Bundle
Bundle NR was a prototype 37-element fuel bundle for the CANDU 600 reactor[1]. For irradiation in the NRU reactor, the centre fuel element was removed and replaced by a central tie rod for irradiation purposes in the vertical test section. Coolant for the test was pressurized light water under typical PHWR conditions of approximately 9 to 10.5 MPa and 300°C.The fuel elements used 1.41 wt% U235 enriched UO2 fuel pellets and were clad with Zircaloy-4 material. The inner sheath surface was coated with a graphite layer. Three types of pellet-stack-to-end-cap geometries were used for the outer elements: a 350 mm3 plenum insert (six elements), a 580 mm3 plenum insert (six
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elements), and no plenum insert (six elements). Intermediate and inner element rings had no plenum insert.Outer element burnups reached average measured burnups of 236.4 MWh/kgU (the maximum was 238.0 MWh/kgU and minimum 235.6 MWh/kgU). Outer element powers were steady during the irradiation and average on time interval ranged between 58 and 62 kW/m during the irradiation. No element instrumentation was used during the irradiation. However, the bundle was subjected to extensive post-irradiation examination (PIE) that included dimensional changes, fission gas release, and fuel burnup analysis. Analysis of NR fuel behavior with ROFEM Code The power history is presented in Figure 6. It can be noticed that in the case NR bundle the power history were of the steady state type with the maximum power at the beginning of the power history.
0
10
20
30
40
50
60
70
80
0 500 1000 1500 2000 2500 3000 3500 4000 4500
Time (h)
Line
ar P
ower
(kW
/m)
Figure 6. Power history of NR case.
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0
500
1000
1500
2000
2500
3000
0 500 1000 1500 2000 2500 3000 3500 4000 4500
Time (h)
Tem
pera
ture
(C
)
Figure 7. Fuel central temperature vs. time.
Figure 7 show the fuel central temperature evolution during irradiation. Figure 8 show the FGR evolution during irradiation. The calculated release show an appreciable increase at the end of irradiation. The calculated FGR underestimate the measured data by about 2.5 cm3(measurements:37.3-41.8cm3).Thought this difference is within a reasonable extent considering the errors that would have been involved in the measurement of FGR and the deviation of the calculated power history.
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sion
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Rel
ease
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m3)
Experimental P.I Data
ROFEM Results (group 1)
ROFEM Results (group 2)
ROFEM Results (group 3)
Figure 8. Fission gas release vs. time.
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The prediction and measured cladding deformations are presented in Figures 9 to 14. The experimental data show the decrease of cladding diameter during irradiation and a slight increase at the power ramp. It can be noticed the satisfactory agreement between measured and calculated values in the midplane region at the end of the power history for elements group 1 and group 3.For the elements group 2 the calculated results slowly underestimate the measured data. In the pellet end region (ridge) calculated results are in satisfactory agreement with experimental data for elements group 1 and slowly higher then experimental data for elements group 3.For the elements group 2 the calculated results underestimate by about 0.17% measurements.
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She
ath
Hoo
p S
trai
n at
Mid
plan
e (%
) Experimental P.I. Data
ROFEM Results
Figure 9 Sheath deformation at the pellet midplane region vs. time for group 1 NR bundle elements.
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-0.2
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ath
Hoo
p S
trai
n at
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plan
e (%
) Experimental P.I. Data
ROFEM Results
Figure 10. Sheath deformation at the pellet midplane region vs. for
group 2 NR bundle elements.
-0.3
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Hoo
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trai
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e (%
) Experimental P.I. Data
ROFEM Results
Figure 11. Sheath deformation at the pellet midplane region vs. time for
group 3 NR bundle elements.
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0
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0.6
0.8
1
1.2
0 500 1000 1500 2000 2500 3000 3500 4000 4500
Time (h)
She
ath
Hoo
p S
trai
n at
Rid
ge (
%)
Experimental P.I. DataROFEM Results
Figure 12 Sheath diameter at the pellet end region (ridge) vs. time for
group 1 NR bundle elements.
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1.6
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Hoo
p S
trai
n at
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ge (
%)
Experimental P.I. DataROFEM Results
Figure 13. Sheath diameter at the pellet end region (ridge) vs. time for
group 2 NR bundle elements.
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Hoo
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trai
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ge (
%)
Experimental P.I. DataROFEM Results
Figure 14. Sheath diameter at the pellet end region (ridge) vs. time for
Table 5. Comparison between ROFEM code calculation results and P.I
experimental data
Fission Gas Release (cm3, STP)
Sheath Hoop Strain (%) Pellet Midplane Pellet End (Ridge)
calculated measured calculated measured* calculated measured*
NR bundle group 1
34.8 37.3÷40.7 0.1 0.03÷0.14 0.7 0.49÷0.75
NR bundle group 2
34.1 37.3÷39.1 -0.1 -0.06÷0.12 0.3 0.44÷0.58
NR bundle group 3
33.1÷49.1 37.3÷41.8 -0.1 -0.18÷-0.03 0.6 0.28÷0.58
* Average values on axial direction 5.1.3 Discussion Fission gas release behavior
The FGR model of ROFEM code assumes that the grain boundary bubbles grow with an accumulation of influx of gas atoms and are connected to form tunnels to a free space. When the amount of fission gas retained in the grain boundary exceeds a
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saturation level, an excess amount of gas is immediately released to the free space. This is the gas release criteria in the present model. In such model, a large amount of FGR in a short transient is allowed only in the limited situations where the following two situations are satisfied. (a) When a large amount of fission gas is accumulated at the grain boundary. (b) When a large and instantaneous decrease in the saturation value at the grain boundary takes place due to changes of the thermal and mechanical conditions. In the results of the case JC, however, the amount of accumulated fission gas at the grain boundary is small before the highest power ramp at 85MWh/kgUO2, and the change of the fission gas saturation level caused by power ramps is small, too. In the other words, the rate/determining process of FGR is diffusion of gas atoms from the grain inside to the grain boundary in the present analyses during the whole irradiation period. Hence, the predicted FGR histories show only slow cumulative increases instead of the possible jumps caused by other mecanisms. This suggests that the actual rods subjected to JC tests had other mechanisms of fission gas release than diffusion process of gas atoms. Therefore the difference between predicted and measured pressure history is considered to be caused by such specific gas release mechanisms which are not taken into account in the present FGR model. This suggests that the actual FGR model used by ROFEM code must be improved in order to consider the specific gas release mechanisms which are not taken into account in the present FGR model. Sheath strain behavior CANDU fuel operated at burnups ≤ 400Mwh/kgU typically exhibits midpellet sheath strains up to 0.5% [ 2].Sheath strain is dependent on pellet geometry/density,fuel power and as-fabricated diametral clearance(between pellets and sheath).The maximum strain observed at a given burnup generally increase gradually from 0.5% at 400Mwh/kgU to 1.5% at 750 Mwh/kgU. This increase in strain appears to be related to two effects:
a) Swelling of pellets due to the buildup of fission products in the ceramic matrix and associated PCI.
b) Internal gas pressure above the gas over pressurization threshold.
The JC and NR fuel elements appear to have been primarily strained as a result of PCI. The strains in NR group 2 and NR group 3 elements decreased proportionally with decreasing in as fabricated plenum volume and the code correctly predicted the effect of plenum volume on sheath strains.
5.1.4 Conclusions
1. JC and NR fuel bundles were succesfuly irradiated in X2 loop of NRU reactor. 2. The JC and NR fuel bundle was subjected to extensive post-irradiation
examination (PIE) that included dimensional changes, fission gas release, fuel burnup analysis, and metallography that included grain size measurement.
22
3. The measured FGR and residual sheath strains are within the range that is expected for similar-operated commercial power reactors and reported in open literature.
4. The difference between predicted and measured internal pressure history at JC elements is considered to be caused by such specific gas release mechanisms which are not taken into account in the present FGR model.
5. This difference sugest that the actual FGR model used by ROFEM code must be improved in order to consider the specific gas release mechanisms which are not taken into account .
6. ROFEM calculated sheath strains compared satisfactory well with P.I. measurements.
References
[1] NEA Data Bank Data for FUMEX 3 Exercise .JC and NR fuel bundles irradiated in NRU reactor [2] M.R.Floyd, Extended-Burnup CANDU Fuel Performance,Proceedings of Seventh International Conference on CANU Fuel,2001 September 23-27,Kingston,Ontario,Canada
[3] Floyd M.R. at al: Performance of Two CANDU-6 Fuel Bundles Containing Elements with Pellet-Density and Clearance Variances, Proc.Sixth Int.Conf.CANDU Fuel, Niagara Falls, Canada, 1999. 5.2 Case 2: EC 51 and EC 89 fuel elements irradiated in TRIGA reactor of INR Pitesti
1. Introduction
Currently the INR Pitesti Nuclear Fuel R&D Program is focused on providing
experimental data for development and validation of the fuel performance computer
codes [1].The in-pile fission gas pressure measurements provide a wide data base for the
evaluation of the fission gas release from the UO2 pellet during power change operation.
The results from the two instrumented fuel elements of different pellet microstructure
which operated until 178.9 Mwd/tU at significantly power levels in TRIGA Material
Testing Reactor (TRIGA MTR) of INR Pitesti, are presented. Some analyses regarding
these tests were performed using the ROFEM computer code and the results were
compared with experimental data in the framework of the IAEA research project [2]. This
23
paper describes briefly measuring techniques developed and currently in use in INR
Pitesti, and presents and discusses selected in-pile test results comparatively with
computer code results.
2. Test Fuel Elements
The test fuel elements have a nominal 12.16 mm diameter pellets contained in a
Zircaloy-4 cladding of 12.24 mm inner diameter with 0.38 mm minimum wall thickness.
The total length of the fuel stack is 294.8 mm for EC51 fuel element and 292.1 mm for
EC89 fuel element. Fuel pellets were made by INR Pitesti from UO2 powder enriched to
7.04 wt% U-235 for EC51 and 3.92 wt% U-235 for EC89 [3]. The level of fuel
enrichment was selected to achieve a linear power output from each element higher then
55 KW/m in the TRIGA MTR. Apart from the differences in fuel enrichment and
element length the only significant difference between test fuel elements appears to be in
the fuel density and microstructure. The pellet average grain size was 14.5µm at the
EC51 fuel element respectively 10.2µm at the EC89 fuel element. All of the fuel sheaths
were coated on the inside surface with graphite layer. The fill gas was pure helium at 0.1
MPa pressure. Test fuel elements were instrumented with pressure transducers to measure
the fission gas pressure changes during fuel irradiation. A small diameter 1.5 mm
capillary tube was connected to the transducer to facilitate internal gas pressure
measurement. The total volume of pressure transducer and capillary tube was
approximately 100 mm3. Other relevant pellets and sheath information are presented in
Table 1.
3. Irradiation Conditions
Test fuel elements were irradiated in the C2 capsule of TRIGA MTR and an
average exposure of 178.9MWh/KgU was achieved for EC51, respectively 125.4
MWh/KgU for EC 89[4]. The power output of test elements was controlled by reactor
power and was determined through calibration of the flux detectors as power sensors.
The effective irradiation time was 2915 hours for EC 89 and 3864 hours for EC
51.The linear power varied during the course of irradiation and in the last irradiation
24
period the linear power increased until 55kW/m in EC 89 fuel element and until 59KW/m
in EC 51 fuel element [4].
4. Experimental Results
The inner gas pressure evolution during irradiation is presented in Figures 1 and
2. During the first startup both fuel elements showed an abrupt rise in gas pressure, which
gradually fell to a low equilibrium value after firs few days of irradiation. After this, the
measured inner gas pressure, shown in these figures exhibited a progressive increase
from approximately 0.7 Map to 3.9 Map starting from approximately 57 MWh/kgU but
became more pronounced at the end of irradiation (5.4 MPa for EC89 and 9.9 MPa for
EC 51 at the end of irradiation).
5. Post –Irradiation Examination Results (PIE)
Hot –cell examinations on the test fuel elements were conducted at INR Pitesti.
Visual examination of the elements did not reveal defects or abnormalities [7] .Element
EC89 had a 4-20µm thick oxide layer on the external cladding surface. Both elements
have circumferential ridges.
Elements EC51 and EC89 were gamma scanned after irradiation. The axial
gamma scans are shown in Figures 3 and 4. Intensity dips are seen at the pellet interfaces,
and there is clear distinction between pellet interfaces. The pellet interface dips are
typical for CANDU fuel operated at powers higher than 55 Kw/m. The isotopic activity
of the irradiated fuel elements showed a variation along the length of the element. These
gradients reflect the axial power distribution during irradiation in the C2 capsule.
Diameter profilometry is shown in Figures 3 and 4. Elements EC51 and EC89
were measured on three directions using dual-transducer profilometer. Each diameter
profile was analyzed for ridge diameters at pellet interface locations and pellet mid-plane
element diameters. Values for each of these were averaged for each profile as well as for
all three directions.
The mid-pellet residual sheath strain ranges from -0.1% to 0.6% for EC89
element and from -0.15% to 0.21% for EC51 element. The pellet interface residual sheath
25
strain ranges from 0.35% to 0.9% for EC89 element and from 0.01% to 0.65% for EC51
element. Post-irradiation measurements show that the bow of the irradiated element is
slightly larger than the pre-irradiation bow for EC 51 and slightly lower than the pre-
irradiation bow for EC 89.The observed dimensional performance is similar to that
expected at the CANDU fuel operation in similar power conditions [8].
PIE puncturing analysis showed that fission gas released was10.8 cm3 STP and
measured element void volume 1.22 cm3 at EC 89. The fission gas of EC51 was lost
during puncturing.
Three samples from every element were impregnated with resin and were
subject to optical microscopy in order to observe fuel pellet microstructure. The typical
crack pattern observed in the irradiated fuel pellets shows several radial cracks the
number of which mainly depends on the linear heat generation rate .On the other hand, it
could be confirmed from the longitudinal sectional photosthat the dish-shape of pellets
has been changed from the as-manufactured conditions. Grain growth was present in the
central part of the pellet indicating that the flux peaking was relatively high consistent
with that observed in gamma scans. Metal fission products are still clearly visible in the
columnar grains region. The grain size of the peripheral regions did not seem to have
changed during irradiation. The cracking pattern and crystal grain grow regions are
typical of the CANDU UO2 fuel operating in similar power conditions [8]. The PIE
results confirm a 6.5µm thick ZrO2 layer on the EC 89 inner surface of the cladding. The
PIE didn’t reveal the ZrO2 layer on the EC 51 inner surface of the cladding.
6. Discussion
A possible explanation of the abrupt rise in gas pressure at first startup involves
the water vapors released from fuel matrix and graphite layer and/or hydrogen trapped
during pellets sintering. A similar mechanism has been proposed for a similar in-pile test
performed at CRNL [5].
Fission gas starts to be released from the fuel after some burn-up and not at the
beginning of irradiation. As seen, the measured rod pressure remains virtually unchanged
up to 78 Mwh/kgU in the EC 51 fuel element (Figure 1) and up to 57 Mwh/kgU in the
EC 89 fuel element (Figure 2), at which point a power increase causes a pressure
increase.
26
One of the objectives of this experiment was to compare experimental data with ROFEM
code results [2]. The comparison between the measured and calculated values is shown in
Figures 1 and 2.As seen, in the first part of irradiation ,the ROFEM code gives results
very close to the measured value. The agreement could be considered satisfactorily
considering the errors that would have involved in experimental data and ROFEM input
data. In both fuel elements the measured pressures show an appreciable increase in the
last power cycles, although the calculated increase in gas pressure is markedly lower than
the measured increase. A possible explanation of this abrupt increase in gas pressure
involves the presence of the pellet cracking pattern generated by rapid power increase.
The fission gas has now possibility to be immediately released to the free space through
the pellet cracks. The PI puncturing of the EC89 fuel element gave 10.3 cm 3 STP inner
gas whereas ROFEM estimation at the end of irradiation was 7.2 cm 3 STP .This
discrepancies could be mainly due to the underestimate of the fission gas release (FGR)
by ROFEM code model. This is, in turn, equivalent to overestimation of gap conductance
and has effect on fuel center temperature. Therefore the difference between predicted and
measured pressure history is considered to be caused by such specific gas release
mechanisms which are not taken into account in the present FGR model.
The FGR model of ROFEM assumes that the grain boundary bubbles grow with an
accumulation of influx of gas atoms and are connected to form tunnels to a free space [2].
When the amount of fission gas retained on the grain boundary exceeds a saturation level,
an excess amount of gas is immediately released to the free space. This is the gas release
criteria in the present code model. In such a model, a large amount of FGR in a short
transient is allowed only in the limited situations .The results of the EC51, EC89 show
that the amount of accumulated fission gas at the grain boundary is large before the final
power ramps, and the predicted pressure histories show only cumulative increases instead
of the measured jumps at end of irradiation. This suggests that the actual FGR model
used by ROFEM code must be improved in order to consider the specific gas release
mechanisms which are not taken into account in the present FGR model.
An empirical correlation has been established at Halden for temperature
burn-up threshold for fission gas release [6]. In the first part of irradiation period
the fuel temperature in the EC 89 fuel element were below Halden temperature
27
threshold for most of the time and small amounts of released fission gases are
expected in this period, which will have little effect on gap heat transfer and thus
on the fuel center temperatures.
Fission gas release can be strongly affected by the fuel microstructure.
Performance improvements can be obtained by increasing the size of the sintered UO2
grains, thereby increasing the length of the diffusion path and thus retarding precipitation
at the grain boundaries .In these conditions the fuel element EC 51 with larger grain size
presents a different FGR threshold.
7. Planned Future Experiments
Two experimental fuel elements with larger grains (>30µm) are planned to be
irradiated in C2 capsule of TRIGA MTR and a marked fission gas release reduction is
expected in this case [1].
8. Conclusions
1.The gas pressure developed inside operating UO2 fuel elements were measured
during irradiation up to a maximum burnup of 178.9Mwh/kgU.The performance of
irradiated fuel elements are analised for linear power variation betwen 45 to 60 KW/m.
2. The anomalously abrupt rise in gas pressure during first startup is a direct
consequence of water vapors released from UO2 pellets.
3. The predictions of fuel performance code ROFEM in terms of internal gas pressure
were compared with the experimental data. In both fuel elements the measured pressures
show an appreciable step increase in the last power steps, although the calculated step
increase is markedly lower than the measured ones. This suggests that the actual FGR
model used by ROFEM code must be improved in order to consider the specific gas
release mechanisms which are not taken into account in the present FGR model.
4. The detailed fuel temperature history has been determined by ROFEM code
calculations when no temperature measurements were available.
5. Experimental results presented in this paper represent data relative to two fuel
elements with different UO2 grain size. The irradiation data shows that the release in an
28
larger grain size pellet fuel rod was smaller than in a small grain size fuel. Other two
experimental fuel elements with larger UO2 grains (>30µm) are planned to be irradiate in
C2 capsule of TRIGA MTR and a marked fission gas release reduction is expected in this
case.
6. Experimental results and calculated data showed that the fuel performance
factors (e.g., fission gas release, sheath strain) were within the range expected for
CANDU fuel operating in similar power conditions.
References
Horhoianu G.: Nuclear Fuel R&D Program at INR Pitesti for the Period 2006-2010,
Internal Report No.7212/2005, INR Pitesti.
2. Horhoianu G.,et al: Improvement of ROFEM and CAREB Fuel Behavior Codes and
Utilization of These Codes in FUMEX 3 Exercise, Progress Report of the IAEA
Research Project 14974,June 2009,INR Pitesti
3. Covaci M, et al: Fabrication of the Test Fuel Elements, Technical Report
No.40129/1985, SPEC-INR Pitesti.
4. Cicerone T,et al: Irradiation of the Experimental Fuel Elements in C2 Capsule of
TRIGA MTR, Internal Report No 2375/1987, INR Pitesti.
5. M.J. F.Notley, Measurements of Fission Gas Pressures Developed in UO2 Fuel
Elements during Operation, Report AECL 2662, Chalk River, 1966
6. Vitanza C. et al: Fission Gas Release from In-Pile Pressure Measurements, Paper
presented at the Enlarged Halden Programme Group Meeting, Loen, Norway, June
1978.
7. Popov M. et al: Post-Irradiation Examination Results of EC 89 Fuel Element, Internal
Report No 2537/1988, INR Pitesti.
29
7. Floyd M.R. at al: Performance of Two CANDU-6 Fuel Bundles Containing
Elements with Pellet-Density and Clearance Variances, Proc.Sixth
Int.Conf.CANDU Fuel, Niagara Falls, Canada, 1999.
Table 1: Summary of EC51 and EC89 fuel elements design characteristics:
FUEL ELEMENTS EC51 EC89
1. Pellet Sintered UO2 Sintered UO2 Enrichment U235 (%) 7.04 3.92 Density (g/c.c.) 10.68 – 10.71 10.54 – 10.62 Grain size average (µm) 14.5 10.2 Roughness (µm, RMS) 0.56 – 0.60 0.54 – 0.66 Stoichiometry (O/U) 2.0077 2.0055 Pellet O.D. (mm) 12.16-12.17 12.15 – 12.16 Pellet and geometry Chamfered and both end dished 2. Cladding Zircaloy-4 Zircaloy-4 Cladding I.D. (mm) 12.24 ± 0.04 12.24 ± 0.04 Wall thickness (mm) min. 0.38 min. 0.38 Diametral gap (mm) 0.080 0.084 3. Fuel element Axial gap (mm) 1.70 1.85 Active column length (mm) 294.8 292.1 Number of pellets per column 23 22 Filling gas He He Filling gas pressure (MPa) 0.1 0.1 Graphite thickness (µm)
5.18 6.8 – 9.6
Bearing pads and spacers 3 bearing pads and 2 spacers Instrumentation Pressure sensor Pressure sensor
30
Table2 Comparison between ROFEM code results and measurements.
Internal Gas Pressure (MPa) Before the ramp* Maximum value during the
iradiation time** Calculated Measured Calculated Measured
EC89 3.8 3.7 4.4 5.3 EC51 3.9 3.4 9.8 5.2
* at 115.6 MWh/kgU for EC89 and 184.2 MWh/kgU for EC51 ** at 128.8 MWh/kgU for EC89 and 189.3 MWh/kgU for EC51
0
10
20
30
40
50
60
0 20 40 60 80 100 120 140
Burnup (MWh/kgU)
Line
ar P
ower
(K
W/m
)
0
1
2
3
4
5
6
Inne
r G
as P
ress
ure
(MP
a)
ROFEM Calculated Linear Power
EC89 Measured Inner Gas Pressure
ROFEM Calculated Inner Gas Pressure
Figure 1. Measured and calculated internal gas pressure for EC 89 fuel element.
31
0
10
20
30
40
50
60
70
0 50 100 150 200
Burnup (MWh/kgU)
Line
ar P
ower
(K
W/m
)
0
2
4
6
8
10
12
Inne
r G
as P
ress
ure
(MP
a)
Linear Power
EC51 Measured Inner Gas Pressure
ROFEM Calculated Inner Gas Pressure
Figure 2. Measured and calculated internal gas pressure for EC 51 fuel element. 5.3 Case 4:FIO 131 and FIO 130 LOCA tests in X2 loop of NRX reactor.
Introduction The IAEA Vienna organized a Coordinated Research Project (CRP) on
improvement the computer codes used for fuel behaviour simulation under the name: FUMEX III [1].The major research objective of this CRP would be to test and develop fuel modeling codes against experimental data and cases provided by IAEA and OECD/NEA [1,2].
Institute for Nuclear Research (INR) Pitesti participated at this CRP with ROFEM and CAREB computer codes [3, 4].The main aim of ROFEM code was to calculate fuel behaviour during steady state operating conditions [5].CAREB code was developed for fuel transients analyses such as LOCA and RIA [6].Recently both codes have been improved with new models in order to extend their capabilities [3 ].
The behaviour of the fuel elements during high-temperature transients is of importance to safety and licensing of the power reactor[7]. During the initial depressurization phase of a hypothetical loss of coolant accident in a pressurized water reactor, the fuel element will be subjected to rapid, high-temperature transient. The FIO-131 and FIO-130 tests was performed to increase the knowledge base of CANDU fuel behavior under LOCA conditions and to provide additional data for validation of the transient fuel performance codes.
32
In this paper, the outline of the various models involved in the CAREB code and the results of the calculations on the FIO-131 and FIO-130 in-reactor tests conditions are summarized. The purpose of this study was to investigate the feasibility of using CAREB code to perform a thermal-mechanical analysis of the fuel element response under hypothetical loss of coolant accident conditions. The analysis of the FIO-131 and FIO-130 tests and the simulation of the test with ROFEM and CAREB codes was an important lesson learned about in-reactor LOCA tests. As a result, this lesson will be used in C2-LOCA tests planned to be performed in TRIGA research reactor of INR Pitesti.
1. CAREB Code
The CAREB computer code was developed to simulate the thermal-mechanical response of a fuel element during rapid, high-temperature transients [6].The model assumed is a single UO2/Zircaloy sheath element with axi-symmetric properties. Physical effects considered in the code are: -Expansion, Contraction, cracking and melting of the fuel, -Variation of internal gas pressure, during the transient, -Changes in the fuel/sheath heat transfer, -Thermal, elastic and plastic sheath deformation (anisotropic) -Zr/H2O chemical reaction effects -Beryllium assisted crack penetration of the sheath (initiated from Be-brazed appendages) The new release of the code, CAREB. 1B, extends the capability of CAREB to model of sheath failure due to oxygen embrittlement upon rewetting and effect of oxide strengthening on sheath creep [3]. Other new features improved ROFEM-CAREB interface [3]. The equations used in the model to represent these physical effects use transient boundary conditions of coolant temperature, coolant pressure and sheath/coolant heat transfer coefficient evaluated by thermal/hydraulic codes . The conditions at the start of the transient are obtained from steady/state fuel performance code ROFEM [3].The CAREB code monitor conditions leading to sheath failure .Several failure mechanisms are explicitly represented in the code:
• Sheath overstrain , • Localized overstrain under oxide cracks • Excessive sheath creep rate , • Low ductility sheath failure , • Beryllium-assisted crack penetration, • High fuel enthalpy (>735 kJ/kg UO2), • Oxygen embrittlement
Beryllium-assisted crack penetration failure mechanism specific to CANDU fuel elements. Intergranular cracking of the Zircaloy fuel sheath can occur at bearing pad and spacer pad locations of the element brought on by the penetration of a beryllium-braze alloy in the presence of an applied hoop stress[8]. The thermal transient experienced by the sheath represent a complex heat treatment which cause changes in anisotropy, annealing, grain size , phase transformation(α-β) plus other changes in sheath microstructure which effect the plastic creep behaviour of the
33
sheath. An improved transient creep model which accounts for these transient changes in microstructure [9] allow detailed examination of the plastic deformation of the sheath which occur under a variety of postulated LOCA transients. As changes in sheath microstructure are evaluated by CAREB during transient, differences in initial sheath properties can be explicitly represented in output data. CAREB was designed to be used as self-contained code with the minimum amount of input information required for execution. However, the fuel element description at the start of the transient is best obtained directly from the steady-state companion code ROFEM, due to detailed information available. These two codes, ROFEM and CAREB, can be run sequentially to describe any arbitrary pre-transient/transient reactor irradiation history desired. The capability of the CAREB code was extensively verified through the comparison with a large number of in-reactor and out of reactor test [10].
2. In-Reactor LOCA Tests The LOCA tests FIO-131 and FIO-130 were performed in the X-2 loop of the
NRX reactor at Chalk River,Canada [11,12,13,14]. The experiment FIO-130 was wery similar to the FIO-131 experiment, in which an unirradiated, Zircaloy-sheated UO2 fuel element was subjected to a blowdown. Important differences include the pre-blowdown burnup (40 MWh/kg U for FIO-130, versus 3 MWh/kgU for FIO-131) and the element internal pressure. The internal pre-transient pressure of the element was approximately 1.5 MPa in FIO-130, while it was approximately 8.1 MPa in FIO-131, which would produce a much greater driving force for sheath strain. The maximum sheath temperature attained during FIO-130 was about 100°C higher than of FIO-131 due to the longer delay time to reactor trip (28 seconds versus 26 seconds) and the fact that there was no power reduction before the trip, as was the case in FIO-131. The major effect of pre-irradiation on the FIO-130 element appears to be presence of intergranular bubbles in the UO2 structure. The bubbles are quite promitent at the top and middle location of the fuel element. At the bottom location, where grain size was much larger and there was evidence of cracking along grain boundaries, there appear to be very few large bubbles at the grain boundaries.
2.1 Test Objectives
The objectives of theses test were: i) to study the thermal and mechanical response of CANDU fuel during a LOCA-type transient, and ii) to provide in-reactor ,transient fuel behaviour data for the transient computer codes validation.
34
2.2 Fuel Design The test was performed with UO2 fuel pellets in Zircaloy-4 sheath. The test
elements details are shown in Table 1. Table 1. Test fuel element details
UO2 Fuel FIO-131 FIO-130
Diameter (mm) 18.06±0.01 18.06±0.01 Length (mm) 461.5 461.5
Enrichment (wt% U-235) 1.38 1.38 Grain size (µm) 7.6 7.6 Density (g/cm3) 10.64±0.02 10.64±0.02
Surface roughness (µm) 0.8 0.8
Fuel Sheath Material Zircaloy-4 Zircaloy-4
Outside Diameter (mm) 19.76±0.02 19.76±0.02 Thickness (mm) 0.81±0.06 0.81±0.06
Inside Diameter (mm) 18.15±0.02 18.15±0.02 Internal surface roughness (µm) 0.36 0.36
Element Length of fuel stack(25 pellets,mm) 461.5±0.75 461.5±0.75 Axial gap(mm) 1.87 5.77 Diametral clearance(measured,mm) 0.09±0.03 0.09±0.03 Filling gas volume (mm3) 8200±200 6100±300
Transducer cavity volume (mm3) 1473 400 Filling gas pressure(MPa) 0.101 0.101 Helium used in filling gas 0.01 0.01
The test assembly was instrumented to measure fuel, sheath and coolant
temperatures and internal element pressure and coolant pressure during the entire irradiation. The fuel centerline and three fuel pellet peripheral thermocouples were located at the axial mid plane of the element. Six Zircaloy sheath thermocouples were laser-welded to the external sheath surface in pairs 180o apart at top, bottom and mid plane locations. A short capillary line connected the element internal volume to an eddy current pressure transducer. Fuel element power was calculated from loop calorimetry based on three separate inlet-outlet differential coolant temperature measurements.
The measurement uncertainty of thermocouples for fuel assembly was estimated to be +/-2K at cladding, +/-4.5K at fuel pherifery and +/-1% at fuel centerline. TFS4 and TFS6 thermocouples, both failed in the FIO-131 experiment [13].The cladding thermocouple TFS7 in the FIO-130 test also failed.The error associated with the transducer used for coolant pressure and internal gas pressure measurement was determined as +/-0.02MPa [13].Error/uncertainty limits for fuel linear powers were +/-10%.Significant uncertainties are associated with the boundary conditions during the transient due to the lack of measurements. There were not enough temperature
35
measurements made on the flow tube, and there was not direct measurement of coolant flow over the fuel element during both steady-state and the transient portions of the experiment.
2.3 Test Parameters During the steady-state irradiation period, the average power rating of the
element was about 68kw/m for FIO-131 and 55 to 64 kw/m for FIO-130 test. Steady-state irradiation of the fuel element was only interrupted by tow reactor trips that occurred ,when the assembly was subjected to coolant depressurization transients. The power history for the steady-state conditions is presented in Figure 1. The first transient was performed at low fuel temperatures to check reactor loop systems and data acquisition computer prior to the high-temperature transient.
The second high-temperature transient was performed to study performance of Zircaloy-sheathed UO2 element during LOCA and to provide documented data for computer codes validation. The sequences and timing of events during the second transient are presented in Tables 2 and 3 and plotted in Figure 3. The high temperature transient was initiated with the fuel operating at a linear power rating of 65-68kw/m. The internal pressure of the FIO-131 test element was artificially set, while operating in-situ just prior the transient, to 8 MPa in order to simulate the amount of fission gas pressure expected in high burnup CANDU fuel. Other pre-transient test parameters are shown in Tables 4. The high temperature transient was produced by isolating the test section from the loop coolant supply while the reactor was at power, and allowing the coolant to blow down from the top and bottom of the test section simultaneously through a pre-set orifice valve into the disposal tank. The rate of depressurization and the magnitude of temperature rise in the fuel and sheath was controlled by the amount of fission heat produced between blow down initiation and reactor shutdown. Complete sheath dry out occurred within 30 seconds after initiation of blow down. Voiding of the test section resulted in a fuel power increase of about 10 percent as calculated by AECL reactor physics code [7,8]. The blow down was initiated with reactor at full power. the reactor power was lowered to 23.8MW and then was reduced to zero about 26 second after blow down initiation for FIO-131 test. The blow down transient was terminated automatically by cold water injection (rewet) at a pre-set test pressure. Return to initial pressure loop cooling was 75 seconds after the blow down initiation only for FIO-131 test. One of the two remaining clad thermocouples of FIO-130 test (TFS 8) failed during transient.The elements were intact at the end of the LOCA transient.
Table 2. FIO-131 test. Second transient. The sequence and timing of events
Time into the transient (seconds)
0 17 24 55 75 106
Reactor Power (MW) 30 23.8 0 0 0 0 Element Power (middle) (KW/m)
65.4 72.7 64.1 8.9 6.5 6.5
Coolant Pressure (MPa)
9.5 (blow down
start)
5.7 4.7 2.1 (rewet start)
2.5 9.5 (return to
loop cooling)
36
Table 3. FIO-130 test. Second transient. The sequence and timing of events
Time into the transient (seconds)
0 61 70 110 116 126
Reactor Power (MW) 23 20 0 0 0 0 Element Power (middle) (KW/m)
65.1 72.6 6.5 6.5 5.2 5.2
Coolant Pressure (MPa)
9.2 (blow down
start)
4.7 2.2 1.9 (rewet start)
2.3 2.3
Sheath strains were measured by post-irradiation profilometry, and the post-test
conditions of fuel and sheath were characterized by destructive post-irradiation examination (PIE).Pellet sheath mechanical interaction was observed as distinct ridges in the sheath at pellet interface locations. The maximum measured strain were about 4%, even through internal pressure were up to 8 MPa for the FIO-131 test. This amount of sheath strain was effective in relieving internal pressure because of small volume within fuel element. CANDU power reactor fuel elements contain an even smaller internal volume.
Table 4. In reactor test parameters (Pre-transient conditions)
Coolant FIO-131 FIO-130
Temperature (oC) 260-285 260-288 Pressure (MPa) 9.8 9.7
Flow (Kg/s) 0.9 0.8 Heat transf.coef. sheath to
coolant(kw/m2K) 20 20
Fuel (mid-element) Power (KW/m) 65 64
Burnup (MWh/kgU) 3 42 Central temperature (oC) 1900 1975
Peripheral temperature (oC) 785 860 Sheath temperature (oC) 288-325 300-320 Internal pressure (MPa) 8.1 1.5
37
Table 5. In reactor test parameters (during second transient) FIO-131 FIO-130 Maximum linear power (KW/m) * 71 71 Maximum central temperature (oC) 2200 2275 Maximum periphery. temp (oC) 1398.1 1450 Maximum sheath temperature (oC) 877 971 Top of element 689 704
Middle of element 877 - Bottom of element 874 971
Element internal pressure(MPa) 8.2 2.5 Time at maximum temp (s) 25 45 Heat up rate (oC/s) Fuel center 10 47
Fuel periphery 35 31 Fuel sheath 77 81
Time at temp. 500 oC (s) 25 50 Coolant temperature(oC) 368.6-374.9 374-451 Coolant pressure (MPa) 9.8 9.7 Differential pressure across sheath (MPa) 0-3 0-7.5 Stored energy (KJ/kg) 600* 560 Quench rate at rewet (oC/s) Fuel center 23 14
Fuel periphery 25 53 Fuel sheath 115 128 * Calculated value at mid-element location
3. Comparison between test measurements and calculated results 3.1 Steady State Conditions During steady –state irradiation period, the average power rating of the fuel
element was about 68 KW/m for FIO-131 and 55 to 64 kw/m for FIO-130 test .The measured fuel central temperature of 1894oC, a peripheral fuel temperature of 787oC and sheath temperature of 288-325 oC for FIO-131 element and 300-320 oC sheath temperature and a peripheral fuel temperature of 780oC for FIO-130 element. The centerline thermocouple of FIO-130 failed before transient.The estimated steady-state irradiation element-accumulated burnup was about 3MWh/kgU for FIO-131 and 42 MWh/kgU for FIO-130 element. The linear power history and evolution of fuel temperatures predicted by ROFEM are shown in Figure 1.The evolution of internal element pressure is shown in Figure 2. Good agreement was found between calculated values and experimental data for steady state irradiation period.
38
a) b) Figure 1. Steady state irradiation period. Fuel temperature.Calculated values and measured data.(a)FIO-131 element.b)FIO-130 element. a) b) Figure 2. Steady state irradiation period.Internal gas pressure.Calculated values and measured data. (a)FIO-131 element.b)FIO-130 element. 3.2 Transient Conditions Comparison between experimental data and CAREB code results have been
made for several parameters including fuel and sheath temperatures and strains, element internal pressure and thicknesses of ZrO2 layer.
Fuel and sheath temperatures are shown in Figures 4,5 and 6. Neutronic calculations indicated that voiding of the test section during blow down resulted in a fuel power increase of about 10% and this caused the increase on temperatures after 20 seconds into the transient.
Fuel peripheral temperature war measured by TFP1 thermocouple positioned in a hole at 7.35 mm pellet radius. The results of TFP1 compared with calculation results for FIO-131 at the same pellet radius show a reasonably good agreement (Figure 5).
With the reactor at full power, fuel peripheral temperatures continued to rise for about 10 seconds and then decreased slightly until rising again 17 seconds into the transient. This arrest in fuel peripheral temperature increase is attributed to fuel-sheath mechanical interaction, and coincides with a rapid temperature rise of the sheath into stable dry out conditions. Fuel temperatures continued to rise until the reactor was tripped to zero power. Initiation of blow down resulted in a rapid coolant depressurization to near
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39
saturation pressure and momentary dry out of the fuel sheath at all thermocouples locations. The maximum clad temperatures was 877oC at FIO-131 element and 971 oC at FIO-130 element . The fuel sheath remained at elevated temperature for 25s in FIO-131 test and 45s in FIO-130 test.
a) b) Figure 3. Second Transient. Time sequence. (a)FIO-131 element.b)FIO-130 element.
The maximum fuel centerline temperature was 2198oC at FIO-131 element and
the agreement between calculated and measured centerline temperatures are reasonable.The centerline thermocouple of FIO-130 element failed before transient.
Coolant and internal element pressures together with sheath strains and ZrO2 thickness are shown in Figure 7,8 and 9. CAREB correctly predicted the evolution of internal gas pressure and in reasonable agreement with measurements. Pressure measurements and calculation results show that the internal element pressure exceeds coolant pressure for almost the entire transient. A positive driving force for sheath strain therefore existed for most of the time the sheaths were at high temperature. Both the calculations and measurements show that the internal element pressure decreases rapidly once the sheath begins to strain. The effect of small internal void volume within a CANDU fuel element on limiting sheath deformation look to be crrectly predicted by CAREB code.The small underprediction of gas pressure in the case of FIO-130 element is due primarily to the existence of axial variation of temperature and sheath strain which were not taken into account in CAREB code.
a) b) Figure 4 Second Transient. Fuel central temperature.Calculated values and measured data. (a)FIO-131 element.b)FIO-130 element.
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40
Resulting sheath strains were measured by post-irradiation profilometry[13,14].The distinct ridges on the sheath at pellet interface location indicated that strong pellet-sheath mechanical interaction (PCMI) had occurred.The arrest in the peripheral fuel temperature rise and coincident sharp rise in sheath temperatures as shown in Figures 5 and 6 sugest that much of this PCMI occurred during the early part of the high temperature transient .The test element was intact in spite of this higher strain.
a) b) Figure 5 Second Transient. Fuel periphery temperature.Calculated values and measured data. (a)FIO-131 element.b)FIO-130 element.
a) b) Figure 6 Second Transient. Sheath temperature.Calculated values and measured data. (a)FIO-131 element.b)FIO-130 element. a) b) Figure 7 Second Transient.Coolant and Internal element pressure.Calculated values and measured data. (a)FIO-131 element.b)FIO-130 element.
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41
a) b) Figure 8 Second Transient.Sheath strain.Calculated values and measured data. (a)FIO-131 element.b)FIO-130 element. Figure 9 Second Transient.ZrO2 thichness.Calculated values and measured data. (a)FIO-131 element.b)FIO-130 element. As expected, the average strain experienced in the FIO-130 fuel element with
lower internal pressure was much less than in the FIO-131 fuel element.The maximum measured average strain at the middle axial location of fuel element was 2.5% for FIO-131 and 0.75% for FIO-13o fuel element compared with the calculated values of 2% respectiv 0.57%.Comparisons between calculated and measured thicknesses of ZrO2 layer are shown in Figure 9. Calculated values are in reasonable agreement with measured thicknesses of ZrO2 layer at the axial middle region of fuel sheath. Zircaloy oxidation is time-at-temperature phenomenon and the sheath temperatures are away from thermocouple-to-sheath junction as much as 100K hotter. Consequently the measured ZrO2 layer thickness must be correlated with the real sheath temperature evolution.
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42
Table 6. Comparison between ROFEM calculation results and experimental data (maximum value during the transient at middle axial location of the fuel element)
FIO-131 FIO-130 Calculated Measured Calculated Measured
Fuel Central Temp. (oC) 1903.1 1894.2± 19 1970.0 1975*(calculed by AECL)
Fuel Periphery Temp. (oC) 754.0 787.3± 4.5 776.0 780.5 Sheath Temp.(oC) - 288-325 - 300-320 Internal Element Pressure (MPa)
1.25 1.1± 0.02 1.54 1.5± 0.02
*fuel centerline thermocouple failed
Table 7. Comparison between CAREB calculation results and experimental data (maximum value during the transient at middle axial location of the fuel element)
FIO-131 FIO-130 Calculated Measured Calculated Measured
Fuel Central Temperature (oC)
2148.4 2198.1± 19 1708.6 -
Fuel Periphery Temperature (oC)
1391.8 1398.1± 4.5 1343.8 1449.2± 4.5
Sheath Temperature(oC) 847.3 877.3± 2 923.8 971.2± 2 Internal Element Pressure (MPa)
8.2* 8.2* ± 0.02 1.5 2.5± 0.02
Sheath Strain (%) 2.1 1.8-2.5 0.57 0.3-0.75 ZrO 2 thickness (µm) 2.8 2-8 5.78 6-12
*8.2 MPa artificially set
4. Future Work 4.1 LOCA tests in TRIGA reactor of INR Pitesti
INR Pitesti has initiated a program to extend data from in-reactor experiments on fuel behavior under rapid, high-temperature transients such us Loss of Coolant Accident conditions [15].These data will be used to validate computer codes, and to demonstrate that all important physical phenomena are accounted in the models. The in-reactor experiments program at the INR Pitesti started with rapid power transients tests in Power Pulse Reactor of INR Pitesti[16]. Now this program has progressed with LOCA tests in TRIGA SS reactor [7,17].The blow down section of C2-LOCA test facility is under construction at TRIGA SS reactor (Figure 10). A number of blowdawn tests will be conducted next years in C2-LOCA facility. It is planned to study the thermal and mechanical behavior of fuel elements and local thermal-hydraulic phenomena at the fuel elements where maximum fuel-sheath temperature will range between 800 and 1250oC [7,17].
43
Figure 10. Blow down section of C2-LOCA capsule
The C2-LOCA test assembly will be instrumented to measure fuel, sheath and coolant temperatures, internal gas pressure and coolant pressure (Figure 11). A turbine flowmeter will monitor coolant flow in the vicinity of test fuel. Flux detectors calibrated as power sensors will monitor power output of the test element. The axial neutron flux profile will be measured with silver and cobalt detectors. A special device will be constructed in order to obtain a $ 1.5 reactivity insertion in the region of C2-LOCA facility[18]. Specifically investigated will be the effects on fuel behaviour from the following parameters:
i) the variation in metallurgical structures of Zircaloy ,and hence mechanical properties, along a CANDU fuel element caused by the brazing process of appendages to the fuel sheath,
ii) the presence of beryllium-braze alloy on the fuel sheath, iii) the initial(pre-transient)internal element gas pressure, iv) the maximum sheath temperatures, iv) the magnitude of a power pulse in the UO2 fuel caused by local reactivity changes during coolant depressurization ,and vi) the amount of pre-irradiation
The analysis of the FIO-131 and FIO-130 tests and the simulation of the test with ROFEM and CAREB codes was an important lesson learned about in-reactor LOCA tests. As a result, this lesson will be used in C2-LOCA tests planned to be performed in TRIGA reactor of INR Pitesti.
44
322,4
APROX. 346,4
306 ± 0,21,4
TIGTIG
WELDWELD
PELLET STACK
SPACER
BEARINGPAD
TUBE
T/C : CLADDINGTEMPERATURE
END CAP
T/C : CENTRALTEMPERATURE
PRESSURESENSOR
Figure 11. Experimental fuel element for LOCA tests. 4.1. CAREB code improvements Future work will include improvements to the computational efficiency of the code,
further development of the code models and further validation of the code against experimental results [14]. Conclusions The FIO-131 and FIO-130 in-reactor tests were successfully performed in X2 loop of NRX reactor at CRNL Canada, to study performance of CANDU type fuel element during LOCA and to provide documented data for computer codes validation. The CAREB and ROFEM codes were used to calculate thermal and mechanical fuel element behavior in the FIO-131 and FIO-130 tests conditions. Calculations performed with CAREB and ROFEM codes were compared to measured data and the following conclusions were made:
1. The CAREB and ROFEM codes were able to adequately analyze the test
fuel behavior in the FIO131 and FIO-130 in-reactor tests conditions. 2. Satisfactory agreement was found between calculated and measured fuel
and sheath temperatures for both steady-state and transient conditions and these parameters accurately reflects trends and changes of the linear power and cooling for both elements.
3. CAREB predicted the evolution of internal gas pressure in reasonable agreement with measurements.
4. The calculated sheath strain compared reasonably well with P.I. measurements.
5. The effect of small internal void volume within on sheath deformation evolution is corectly predicted by CAREB code.
6. Calculated ZrO2 layer thicknesses were found to be in satisfactory agreement with measurements in the axial middle position of the sheath.
45
7. Capabilities of CAREB code directly cover the rod failure.The code correctly predicted that the fuel elements would not fail according of the potential failure mechanisms included in the code.
8. The major problems encountered in this study were the enough specification of boundary conditions during the transient due to the lack of direct measurements.
9. The analysis of the FIO-131 and FIO-130 tests and the simulation of the test with ROFEM and CAREB codes was an important lesson learned regarding in-reactor LOCA tests. As a result, this lesson will be used in C2-LOCA tests planned to be performed in TRIGA reactor of INR Pitesti.
10. The blow down section of C2-LOCA capsule is under reconstruction in TRIGA Research Reactor of INR Pitesti.A number of instrumented LOCA tests on RU-43LV experimental fuel elements are planed to be performed in C2 -LOCA capsule.
References
7. Killeen,J.; Specific Research Objectives of FUMEX-III CRP, First IAEA Technical Meeting-FUMEX III, Vienna, December 10-12, 2008
8. Sartori,E.; CD with IFPE Data Base Selected for FUMEX-III CRP, First IAEA Technical Meeting-FUMEX III, Vienna, December 10-12, 2008
9. Horhoianu,G. at al.; Improvement of ROFEM and CAREB Fuel Behaviour Codes and Utilization of these Codes in FUMEX III Exercise, Progress Report to IAEA Research Contract No: 14974,IAEA Vienna,24 September 2008.
10. Horhoianu,G.at al; Application of ROFEM and CAREB Codes to FUMEX 3 Exercise, IAEA Technical Meeting-FUMEX III, Pisa, Italy,01-04 June,2010
11. Moscalu,D.R.; Aspects Regarding Nuclear Fuel Burnup Increase, Ph.D.Thesis, Institute for Nuclear Research,Pitesti,1997
12. Arimescu,I.; High Temperature Transients Fuel Performance Modelling, Ph.D. Thesis, Institute for Nuclear Research,Pitesti,1987
13. Horhoianu,G., Investigation of theRU-43LV fuel behaviour under LOCA conditions in CANDU reactor, Paper presented at IAEA Meeting on Fuel behaviour and modelling under severe transient and LOCA conditions ,18-21 October,2011,Mito-city, Ibaraki-ken, JAPAN.
14. Kohn,E. and Sagat,S.; Beryllium Assisted Cracking of Zircaloy “,6th Canadian Fracture Conference, Harrison Hot Springs,B.C.,Canada, June,1982
15. Sills,H.E. and Holt,R.A.; Predicting High Temperature Transient Deformation from Microstructural Models,4th International Conference on Zirconium in Nuclear Industry, Stratford-upon-Avon,1978
16. Ion,S., Ionescu,D.V.; CAREB Verification and Validation Manual, Internal Report No.4804,INR Pitesti,March,1996
17. Ferenbach,I.J.,at al,Zircaloy- ,AECL-8569, October 1984 18. Ferenbach, P. J. et al, In Reactor LOCA Tests of Zircaloy-Sheathed UO2 Fuel at
Chalk River, report AECL-8974,6th Annual Conference of the Canadian Nuclear Society,June,1985
46
19. NEA-1783 IFPE, CANDU experiment FIO-131. Fuel Behaviour under LOCA Conditions, Directory of Experiment, NEA-1783/01, 12 April 2007
20. Yatube,S.,Exp-FIO-130,Raport for FUMEX 3 CRP, Chalk River,2010 21. Horhoianu,G.; Nuclear Fuel R&D Program of INR Pitesti for the Period 2010-
2015,Internal Report No.8779, INR Pitesti, May, 2010 22. Horhoianu,G.,et al; Behavior of CANDU Fuel under Power Pulse Conditions at
TRIGA Reactor of INR Pitesti,KERNTECHNIK Journal,Vol.74,No.1-2April, 2009
23. Horhoianu, G.et al; Large Break LOCA Simulating Tests on CANDU Fuel Elements in TRIGA MT Reactor of INR Pitesti, Annual Meeting on Nuclear Technology ,Nurmberg, Germany ,10-12 May 2005,.
24. Datcu,A.; A Special Device for Local Reactivity Insertion in TRIGA Research Reactor, Internal Report No.8269,INR Pitesti,November,2008
8. Paper published in connection with the IAEA Research Project 14974/RO
ROFEM and CAREB codes calculation results and comparison with experimental data in the framework of FUMEX 3 CRP and IAEA Research Project 14974 are included in the following papers published and presented at International Meetings and Conferences :
1. In-reactor measurements of fuel centerline temperature variation during power change, paper published in KERNTEHNIK vol.75,No.3 from April 2010(authors: G.Horhoianu,D.Ionescu,E.Pauna).
2. In-reactor measurements of the fission gas pressure and comparison with ROFEM code results, paper published in KERNTEHNIK vol.75,No.4 from August 2010(authors: G.Horhoianu,D.Ionescu,E.Pauna).
3. Comparison between CAREB code calculations and LOCA test results in the FUMEX III project, paper published in KERNTEHNIK vol.76,No.2 from May 2011(authors: G.Horhoianu,D.Ionescu,E.Pauna).
4. CANDU fuel behaviour in the load following tests, paper published in KERNTEHNIK vol.76,No.4 from August 2011(authors: G.Horhoianu,S.Palleck).
5. Application of ROFEM and CAREB codes to FUMEX 3 exercise, Paper presented at IAEA Technical Meeting-FUMEX3,Pisa,Italy,01-04 June 2010(authors: G.Horhoianu,D.Ionescu,A.Paraschiv,E.Pauna).
6. Load Following tests on CANDU Type Fuel Elements in TRIGA Research Reactor of INR Pitesti, paper presented at 11
International Conference on CANDU Fuel,
Niagara Falls,Canada,17-20 October 2010(authors: G.Horhoianu,S.Palleck,D.Ionescu.).
7. Investigation of theRU-43LV fuel behaviour under LOCA conditions in CANDU reactor, Paper presented at IAEA Meeting on Fuel behaviour and modelling under severe transient and LOCA conditions ,18-21 October,2011,Mito-city, Ibaraki-ken, JAPAN.
8 Analitical assessment for stress corrosion fatigue of CANDU fuel elements under load following conditions, Paper sent in 2011 for publication in KERNTEHNIK journal. (authors: G.Horhoianu ,D.Ionescu,E.Pauna)
47
9. Investigation of the RU-43 LV fuel behaviour under LOCA conditions in CANDU reactor, Paper sent in 2011 for publication in KERNTEHNIK journal. (authors: G.Horhoianu,M.Serbanel,C.Diaconu)
10. In-Pile LOCA Tests.Comparison Between Code Results and Test Measurements in FUMEX III CRP, Paper presented at IAEA Technical Meeting-FUMEX3,Vienna 05-08 December ,2011 (author: G.Horhoianu)
8. Final Conclusions - The individual models included in the ROFEM and CAREB codes have been reviewed in order to improve these models and the predictive capabilities of the codes.
- In order to analyse the CANDU fuel behavior in transient/accident conditions,the code ROFEM has been coupled with CAREB code.
- The FUMEX 3 CANDU reactor specific cases have been simulated with ROFEM,and CAREB codes. Satisfactory agreement was found for both steady-state and transient conditions .The calculations results reflects trends and changes of fuel design parameters and linear power history.
- Further work for improving and refining the models of the present codes is necessary but it is strongly dependent of the available experimental data base
- New in reactor instrumented tests aiming to cover the fuel modeling problems are already in progress in TRIGA SS reactor and TRIGA Power Pulse reactor of INR Pitesti.