Experimental and analytical study on MOX fuel behavior under RIA-simulating conditions in the NSRR

Y. UDAGAWA, T. SUGIYAMA, M. AMAYA, M. SUZUKI, F. NAGASE, T. FUKETA Fuel Safety Research Group, Nuclear Safety Research Center Japan Atomic Energy Agency Tokai-mura, Ibaraki-ken Japan Abstract Pulse-irradiation experiments and subsequent code analyses have been made in order to study MOX fuel behavior during a

reactivity-initiated accident (RIA) in light water reactors. The experiments were performed by using pulse-irradiation

capability of the Nuclear Safety Research Reactor (NSRR) of the Japan Atomic Energy Agency (JAEA), and fuel performance

code RANNS developed by JAEA for accident conditions was used for the post-test simulations. The test DW-1 was

performed on an 8x8 BWR-MOX fuel rod with Zircaloy-2 cladding, and the tests BZ-1, -2, and -3 were performed on 14x14

PWR-MOX fuel rods with Zircaloy-4 cladding. The results of these experiments suggest that the failure limit under RIA

conditions is dominated by the cladding corrosion states including oxidation and hydride precipitation, and accordingly, the

same failure criterion is applicable to UO2 and MOX fuels. The computational analysis gives the following insights into the

fuel behaviors during the transients. For all the test cases, the pellet temperature differences of the MOX fuels from UO2 fuels

with similar enrichment and burnup were only about 50 deg C at most. Pellet cladding mechanical interaction (PCMI) loading

on the cladding in rod-axial direction was significantly smaller than in hoop direction in the BWR-MOX case, and resulted in

relatively small PCMI-induced plastic deformation of the cladding. The pellet-cladding (PC) gap reopening was influenced by

the degree of PCMI-induced cladding deformation. Larger PCMI-induced cladding deformation resulted in earlier PC gap

reopening. The pressure of the released fission gas had significant effect on cladding deformation solely in the case that boiling

transition occurred and cladding temperature reached as high as 800 deg C during the film boiling. Even in this case, the gas

pressure was not high enough to facilitate significant plastic deformation, although cladding yield stress level was reduced with

the large temperature rise. High temperature creep deformation was operative during the film boiling. It was shown by the

computational analysis that the pellet temperature evolution in the MOX fuels during the pulse irradiation was similar to that in

UO2 fuels, and that the cladding deformation behaviors observed are explained well by the mechanisms consistent with those

assumed for high burnup UO2 fuels.

1. INTRODUCTION Extensive use of mixed-oxide (MOX) fuels in power-producing light water reactors (LWRs) is being promoted step by step in Japan with a view to using uranium resource effectively. The Nuclear and Industrial Safety Agency in Ministry of Economy, Trade and Industry (METI/NISA) licensed the MOX fuel loading up to one-fourth of their cores in five PWRs (Takahama Units 3 and 4, Genkai Unit 3, Ikata Unit 3, and Tomari Unit 3), that up to one-third in four BWRs (Kashiwazaki-Kariwa Unit 3, Hamaoka Unit 4, Shimane Unit 2, and Onagawa Unit 3), and full-MOX core in the Ohma ABWR, and the Nuclear Safety Commission (NSC) of Japan reviewed those licensings. The licensees have made their application for the licenses conforming to the safety review guidelines, authorized by the NSC, for MOX fuel loadings in LWR cores. The first guideline for MOX fuel loading up to one-third in reactor cores suggests that the framework of the safety evaluation for UO2 fuel cores is applicable to MOX-loaded cores up to assembly burnup of 45 MWd/kg, supposing that differences of MOX fuels from UO2 fuels are adequately taken into account by replacing the neutronics, thermal and mechanical properties, which are the inputs to the safety evaluation. Similar suggestion is made in the second guideline, for full-MOX cores, but up to

assembly burnup of 40 MWd/kg. The 5 MWd/kg lower burnup than that suggested in the first guideline is to suppress the radiation damage to the cladding, because harder neutron spectrum is expected in full-MOX cores than that in UO2 fuel cores. The 45 MWd/kg burnup suggested in the first guideline is determined so that the maximum pellet burnup in MOX fuel assemblies is less than that in the standard 8x8 UO2 fuel assembly which was referred to in the guideline. The technical knowledge on high burnup MOX fuel behaviors, especially under accident conditions is rather limited. For the safety evaluation of high burnup MOX fuel in LWRs, it is important to promote understanding of its behavior under accident conditions in comparison with that of UO2 fuel. In order to assess possible MOX effects in fuel behavior during a reactivity-initiated accident (RIA), high burnup MOX fuels irradiated in power-producing LWRs were subjected to pulse-irradiation experiments in the Nuclear Safety Research Reactor (NSRR) of the Japan Atomic Energy Agency (JAEA) [1]. These tests were analyzed using the RANNS code, the fuel performance code developed by JAEA, for the better understanding of the high burnup MOX fuel behaviors. The present paper describes the test results such as fuel failure limit, fuel rod deformation, and fission gas release (FGR) on the tested fuel rods in comparison with the data obtained for the irradiated UO2 fuels, and the insights obtained by the computational analysis.. 2. TEST CONDITIONS 2.1. Pulse-irradiation in the NSRR The NSRR is a modified TRIGA Annular Core Pulse Reactor (ACPR) whose salient features are a pulsing power capability and a large (22 cm in diameter) dry irradiation space located in the center of the reactor core to accommodate a sizable experiment [2]. At the maximum reactivity insertion of $4.6, the peak power reaches ~21 GW and the corresponding pulse width is 4.4 milliseconds. Figure 1 shows linear heat rate histories of the four MOX fuel rods during the NSRR pulse irradiation tests, DW-1, BZ-1, BZ-2, and BZ-3, which are presented in this paper.

Fig. 1. Linear heat rate histories of four MOX fuel rods during the NSRR pulse irradiation tests, DW-1, BZ-1,

BZ-2, and BZ-3.

The experimental capsule for irradiated fuel rod tests is a double container system which consists of outer and inner capsules made of stainless steel. Figure 2(a) shows a schematic diagram of the capsule for tests at a room temperature (RT) and atmospheric pressure; the RT capsule. The capsule contains an instrumented test fuel rod with stagnant water. A length of the test fuel rod installed into the capsule is approximately 300 mm in total length with a pellet stack of 100 to 120 mm. Cladding surface temperatures are measured by 0.2 mm bare-wire type R (Pt13%Rh/Pt) thermocouples (TCs) spot-welded to the cladding. Coolant water temperature is measured by sheathed type K (chromel/alumel) TC (1 mm in diameter) near the cladding surface at top of the test fuel rod and/or center of the fuel stack. A strain gauge type pressure sensor is installed at the bottom of the inner capsule to measure the increase of capsule internal pressure, and a float-type water column velocity sensor is used in order to evaluate mechanical energy generated by dispersed pellets/coolant interaction.

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In addition to the RT capsule system, a new high temperature (HT) capsule was developed and subjected to the experiments in order to assess effects of initial coolant temperature on fuel behaviors [3]. The HT capsule can be used at a coolant temperature up to 286 deg C. Figure 2(b) shows the schematic of the HT test capsule. The inner capsule is equipped with an electric heater to raise the coolant temperature and pressure up to ~286 deg C and the corresponding saturation pressure. One TC of type R is welded on the cladding surface of the test fuel rod, and two pressure sensors are located at gas and liquid phases of the coolant. Because of the spatial limit in the HT capsule, the test fuel rod is short; approximately 130 mm at maximum in total length, including a pellet stack of ~50 mm. The preparation of the test fuel rods and extensive pre- and post-test fuel examinations are performed in a large hot cell facility, the Reactor Fuel Examination Facility (RFEF), of JAEA.

Fig. 2. Capsules for (a) RT tests and (b) HT tests

Table 1 Test fuel rods and conditions in DW-1, BZ-1, BZ-2, and BZ-3 Test ID

DW-1 BZ-1 BZ-2 BZ-3

Test fuel rod Nuclear power plant Dodewaard Beznau Beznau Beznau

Rod type 8x8 BWR 14x14 PWR 14x14 PWR 14x14 PWR Cladding material Zry-2 with Zr-liner Low-tin Zry-4 Low-tin Zry-4 Low-tin Zry-4

MOX pellets production MIMAS SBR MIMAS MIMAS Initial Pu enrichment, total (%) 6.4 5.5 5.6 5.6

Initial Pu enrichment, fissile (%) 4.6 4.0 4.1 4.1 Rod average burnup (MWd/kg) 45 48 59 59 Average oxide thickness (μm) 10 30 20 20

Cladding hydrogen content (wtppm) 50 340 160 160

Test conditions Coolant temperature (ºC) ~20 ~20 ~20 281 Coolant pressure (MPa) 0.1 0.1 0.1 6.6

Initial fuel enthalpy (J/g)* 0 0 0 70 Maximum fuel enthalpy (J/g)* 497 673 630 594

* : 20 ºC-based enthalpy, calculated using the RANNS code 2.2. Test Fuels and Pulse Operation Conditions The four test fuel rods subjected to a series of the experiments on LWR-MOX are listed in Table 1. The test rod of

(a) RT capsule (b) HT capsule

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the DW-1 was sampled from the segment shipped from the Dodewaard NPP of the Netherlands. The BZ-1 test fuel rod was sampled from the segment transported from the Beznau NPP of Switzerland. The fuel rods tested in BZ-2 and -3 were sampled from another segment shipped from the Beznau NPP. Pure helium was filled into all of these four test fuel rods at a room temperature and a pressure of 0.1 MPa. The three tests, DW-1, BZ-1 and -2 are so-called “cold tests” performed with a coolant condition of a room temperature and atmospheric pressure. Each test rod was installed into the RT capsule shown in Fig. 2(a), and then subjected to the pulse-irradiation of the maximum reactivity insertion. This pulse-irradiation led fuel enthalpies of the test rods to 497 to 673 J/g as listed in the Table 1. On the other hand, the test BZ-3 was conducted with a coolant condition of a high temperature and high pressure by using the newly developed HT capsule shown in Fig. 2(b). The coolant water was heated electrically up to 281 deg C just before the pulse-irradiation, and the pressure reached the corresponding saturation pressure of 6.6 MPa. The fuel enthalpy was elevated at this stage, and became 70 J/g higher than that at a room temperature. During the reactivity insertion, the fuel enthalpy reached 594 J/g at maximum. Accordingly, the enthalpy increase due to the pulse was 524 J/g. The fuel enthalpies were obtained by time integration of the linear heat rate with consideration of heat removal to the coolant. The calculation was performed with the RANNS code, described below. 3. RANNS CALCULATION 3.1. FEMAXI-7 code and RANNS code In JAEA, the FEMAXI-7 code [4], and the RANNS code [5] have been developed. The former is for the analysis of fuel behavior in normal and anticipated transient conditions, and the latter is for the fuel behavior in accident conditions. These two codes are used to experimental and exploratory analyses, and are in the process of validation and improvement of models with irradiation test data. The both codes treat a single fuel rod and give a coupled solution of one-dimensional FEM mechanical analysis and thermal analysis in the radial direction, with a calculation mesh as shown in Fig. 3. In the present calculation, the fuel pellet and cladding were divided into 36 and 10 concentric ring elements, respectively, both in mechanical and thermal analysis. The two outermost ring elements and the remaining eight ring elements correspond to the outer surface oxide layer and the metallic wall, respectively. The pellet stack of the analyzed rods were treated as solely one axial segment since the axial length is small and the axial power profile during the pulse irradiation experiments is nearly uniform.

Fig. 3. Geometry of mechanical and thermal analysis model of FEMAXI-7 and RANNS codes

3.2. Calculation conditions The pre-pulse fuel rod geometry as pellet-cladding (PC) gap was determined based on the measured cladding diameter profile and oxide layer thickness using the FEMAXI-7 code. Calculation by FEMAXI-7 for base-irradiation in the Dodewaard and the Beznau plant was performed along approximate power histories for each test fuel rod. In the calculation, the length of the pellet stack was set to be identical to that of the test fuel rod for the pulse irradiation test. Cladding oxidation rate was adjusted so that the oxide layer thickness at end-of-life (EOL) agrees with the post irradiation examination (PIE) data. Also, fuel pellet swelling rate was adjusted so that the cladding diameter at EOL agrees with the PIE data. Consequently, the FEMAXI-7 calculation result reproduced the PIE result in each test case. The EOL fuel state determined by these adjusted calculation was used as the initial

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condition for the RANNS calculation. The linear heat rate histories during pulse irradiation, shown in Fig. 1, were evaluated based on the NSRR power histories measured with neutron detectors and coupling factor, the ratio of the energy generated in unit mass of a test fuel to the total reactor power of the NSRR. The coupling factor was evaluated for each test rod by using the mass analysis data. The MATPRO-11 [6] and MATPRO-A [7] model were adopted for pellet specific heat and pellet thermal expansion rate, respectively. The MATPRO-09 [8] model was adopted for cladding specific heat, thermal conductivity, and in-pile creep rate. Rosinger’s model was adopted for high temperature Zircaloy creep, which can be operative when departure from nucleate boiling (DNB) occurs and the cladding reaches high temperature. The latest MATPRO model was adopted for the cladding mechanical properties to describe elastic and plastic behavior [9]. Pellet mechanical behavior was assumed rigid, i.e. plastic and creep strains are inhibited. Thermal bonding, which is close to perfect thermal contact, was assumed between pellet and cladding when gap is closed under pellet-cladding mechanical interaction (PCMI) loading. Specific models to MOX fuels were Ohira’s pellet thermal conductivity model [10] and radial power profile which was calculated by using the Monte-Carlo burnup code MVP-burn [11]. For the tests DW-1 and BZ-3, in which cladding surface temperature was measured with TCs, the cladding surface temperature history was used as a boundary condition (BC) at each time step. 4. RESULTS AND DISCUSSON 4.1. Test DW-1; BWR-MOX fuel behavior The test rod remained intact in the test DW-1. DNB occurred, cladding surface temperature reached ~400 °C in local area, and film boiling continued for 1 to 1.5 seconds, as shown in Fig 4. Cladding residual hoop strain was 0.41% on average over the active region. The cladding residual hoop stains measured in BWR fuel experiments are plotted as a function of fuel enthalpy increase in Fig. 5. A larger PC gap results in lower residual strains. BWR fuels tend to have a larger PC gap at EOL conditions compared to similar PWR fuel. Therefore, BWR fuels tend to have a higher threshold enthalpy for measurable residual strain compared to PWR fuels. The strain measured in the test DW-1 agrees well with other UO2 fuels data [12] as can be seen in the figure.

Fig. 4. Histories of cladding surface temperature in tests DW-1 and BZ-3

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Fig. 5. Cladding residual hoop strain in BWR fuels as a function of enthalpy increase According to post-test rod puncture and gas analysis, FGR during the pulse-irradiation, the molar ratio of the fission gas (xenon (Xe) and krypton) released during the test to the total amount of the fission gas generated during operation cycles in the NPP, was 14.0% in the test DW-1. Figure 6 shows FGR in BWR fuels during the pulse-irradiation as a function of that during operation in NPPs. The FGR at the EOL is not available for the DW-1 test rod and the data from sibling rods irradiated together with the DW-1 rod range from 20 to 30%. FGR during pulse irradiation is strongly influenced by base-irradiation conditions, e.g. linear heat rate. It can be seen in the figure that the FGR in the test DW-1 is consistent with the data from experiments on BWR-UO2 fuels [12].

Fig. 6. Fission gas release in BWR fuels during the pulse-irradiation as a function of that during operation in NPPs

Figure 7(a) shows evolution of pellet temperature in the test DW-1, calculated using the RANNS code. The TC data “DW-1 #1” of cladding surface temperature shown in Fig. 4 was used as a BC in all the DW-1 calculations. The figure also shows a hypothetical calculation result which adopts a pellet thermal conductivity model [13] and a radial power profile suitable for UO2 fuel analysis, to compare with the MOX model result. According to the MVP-burn calculation results, the MOX fuel pellet has higher radial power peaking than the UO2 pellet in early stage of burnup, though the difference is smaller at the EOL, 45 MWd/kg. As a result, the MOX model case gives higher pellet surface temperature and lower pellet center temperature than the UO2 case, as shown in the figure. The temperature difference between the two cases is, at most, only 50 deg C. Plastic strain values calculated using the RANNS code are compared with the measured value in Fig. 5. Calculated plastic strain is dependent on cladding stress state during PCMI because RANNS determines yield point with von Mises criterion. Stronger biaxial stress state, i.e. higher axial stress by PC bonding in addition to hoop stress, tends to result in larger plastic hoop strain. The case with weak PC friction, which results in relatively small cladding axial stress, gives better agreement with the measurement than the case with strong PC friction. This suggests that PCMI loading on cladding in axial direction was significantly smaller than that in hoop direction in the test DW-1. Such tendency has been observed also with direct elongation measurements of pellet stack and cladding in the previous BWR-UO2 tests [12, 14].

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The weak PC friction case still overestimates the measured cladding residual strain, slightly. Another possible cause of the overestimation could be underestimation of pre-pulse PC gap. The base irradiation calculation using the FEMAXI7 code, which provided the initial condition of the above mentioned RANNS calculation, predicted almost zero FGR and 4 microns of diameter PC gap in cold shut-down state. With significant FGR in base irradiation, rod internal pressure and EOL PC gap tend to be larger. A calculation with assumption of 30% FGR in base irradiation predicted 13 microns of diameter PC gap, which resulted in only 0.05% decrease in cladding residual strain.

Fig. 7. Evolution of (a) pellet temperature, (b) PCMI pressure, cladding surface temperature, and cladding hoop strain in test DW-1

Figure 7(b) shows evolution of PCMI pressure and cladding strain in the weak PC friction case. The PCMI pressure reaches its peak value just after the pulse irradiation at around 0.23 sec and results in the rapid increase in cladding plastic strain. While the PCMI pressure decreases with cladding temperature rise, it is more than about 30 MPa (20% of the peak value) during film boiling and induces small increase in cladding creep strain. It should be noted that PC gap reopening occurs after the quench, i.e. gas induced cladding deformation is difficult to occur. Although hypothetical calculations with artificial FGRs during pulse irradiation up to 50% were performed to investigate gas pressure effects, no significant change in cladding deformation was observed. Such long PC contact is mainly attributed to the relatively low cladding temperature during film boiling and small PCMI-induced deformation, as confirmed by comparing with the BZ-3 calculations described below. Another possible cause could be underestimation of initial PC gap as described above. More precise evaluation of initial PC gap is important to understand BWR fuel behaviors and it requires further validation of the models for FGR, cladding irradiation creep, and pellet swelling which are used in base irradiation calculation. The test DW-1 results suggest that no discernible difference appears in rod deformation and FGR between MOX and UO2 fuels in BWR, at least, below a burnup of 45 MWd/kg. 4.2. Test BZ-1 and -2; PWR-MOX fuel failure The tests BZ-1 and BZ-2 resulted in PCMI failure at fuel enthalpies of 329 J/g and 554 J/g, respectively. A long axial crack propagated over the pellet stack region as shown in Fig. 8. Results from a series of the NSRR experiments on high burnup LWR fuels show that the heavier corrosion of cladding during operations in NPPs, which in turn give rise to the larger hydrogen absorption in cladding, results in fuel failure at the lower enthalpy. In particular, the thickness of hydride rim that appeared in high burnup PWR fuel cladding, i.e. cladding peripheral layer containing dense hydride clusters, well correlates with fuel enthalpy at failure [15]. Figure 9 shows fuel enthalpy at failure as a function of cladding oxide thickness. Fuel enthalpies at failure in the two MOX tests BZ-1 and -2 are consistent with a tendency derived from a number of tests on UO2 fuels, and in turn indicate that any MOX effects do not appear. The threshold of fuel failure due to PCMI only depends on the cladding state alone.

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Fig. 8. Post-test appearances of the rods in tests BZ-1 and BZ-2

Fig. 9. Fuel enthalpy at failure as a function of cladding oxide thickness

Fig. 10. Visual appearance, X-ray image, residual hoop strain profile, and longitudinal cross section of the

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4.3. Test BZ-3; PWR-MOX fuel behavior The test BZ-3 resulted in no failure. As can be seen in Fig. 4, DNB occurred just after the pulse and the cladding surface temperature reached ~800 deg C at maximum. Film boiling continued for ~3.8 seconds. The cladding residual strain was 4.4% as the circumferentially-averaged and axial-peak value. Figure 10 shows the visual appearance, X-ray radiograph, residual hoop strain profile of cladding, and the longitudinal cross section of the post-test rod. The cladding strain was almost zero at the end pellet, which was fragmented probably by the radial constraint force from the cladding when the pellet thermal expansion occurred. Such localized pellet fragmentation is attributed to the weak axial constraint on the end pellet. The large strain appeared in the posttest BZ-3 rod exceeds the level achievable by PCMI driven with pellet thermal expansion, which is about 1% in the present case, and suggests gas-induced deformation during the post-DNB phase [16]. The averaged value of cladding residual strain over the region with strain > 1% was 3.4%. The post-test rod puncture and gas analysis showed that FGR during the test BZ-3 was 39.4%. The FGRs during PWR fuel experiments are plotted in Fig. 11 as a function of peak fuel enthalpy [17]. Data from the previously-performed NSRR experiments on ATR/MOX fuels [18] and REP-Na experiments performed in sodium loop of the French CABRI reactor [19-21] are included in the figure. It can be seen that the FGRs of PWR fuels correlate with the maximum increase of fuel enthalpy. The ATR/MOX fuels have a homogeneous micro-structure similar to that in SBR-MOX fuels, and the FGRs from the ATR fuels remain in the same level of those from UO2 fuels. The MIMAS-MOX fuels tested in the REP-Na experiments and the test BZ-3, on the other hand, show the larger FGRs. In particular, the FGR of 39.4% in the test BZ-3 looks significantly large in comparison with those in tests on UO2 fuels, even if one takes into consideration the initial fuel enthalpy of 70 J/g (17 cal/g) in the experiment started from a coolant condition of 281 deg C.

Fig. 11. FGR in PWR fuels during the pulse-irradiation as a function of peak fuel enthalpy The highest FGR among each UO2 fuel ranged from 20 to 30% in the previous NSRR experiments [16], and it is generally accepted that the FGR achievable in an RIA-simulating test corresponds to the total amount of accumulated fission gas in grain boundaries. In the MIMAS-MOX fuels, a large amount of fission gas is accumulated in the Pu agglomerates, and could give additional FGR during the RIA transient. Figure 12 shows SEM/EPMA results of MOX pellet sampled from the mother rod of the test rods for BZ-2 and BZ-3. In the element mapping image (Fig.12(a)), localized Pu-rich regions, called Pu-spot, can be seen as white area in the image of Pu. The volume fraction of Pu-spot in the pellet was estimated to be ~20%. The distributions of neodymium(Nd)-rich regions and Xe depressed regions correspond well to that of Pu-spot. Besides, cracks passing through Pu-spot, which are hardly observed in pellets after base irradiation, were found in the pellet after the test BZ-3. Hence, it is likely that fission gas in Pu-spot was released during the test BZ-3. In order to estimate FGR from Pu-spot, the Nd-rich regions were selected so that the total length of Nd-rich regions in the line analysis (Fig.12(b)) is equal to the volume fraction of Pu-spot (~20%). In the selected regions, which are colored and labeled “Pu-spot region” in the figure, local burnup is ~70% larger than the average and local Xe concentration is ~30% lower than the average. Thus, FGR from Pu-spot alone is estimated to be 0.2×(1.7−0.7) / 1.7 ≈ 12% [22]. Such FGR from Pu-spot could account for the large FGR in the test BZ-3. Nevertheless, it cannot be concluded that the large FGR was the result of some MOX effects. In the previous HT tests performed on high burnup PWR-UO2 tests, relatively large FGRs were observed, compared to RT tests [23].

It has been discussed that there could be other causes of large FGR specific to HT test conditions, such as the higher pellet temperature due to the higher initial coolant temperature than those in RT tests and the fragmentation of the end pellet, shown in Fig. 10, which could enhance gas release. Therefore, further investigation is needed regarding the main cause of the large FGR observed in the test BZ-3.

Fig. 12. SEM/EPMA results of MOX pellet sampled from mother rod of BZ-2 and BZ-3 test fuel rods: (a) element mapping image and (b) line analysis

Figure 13(a) shows evolution of pellet temperature in the test BZ-3, calculated using the RANNS code. The TC data “BZ-3” shown in Fig. 4 was used as a BC in all the BZ-3 calculations. The difference of temperature evolution between the MOX model result and the hypothetical UO2 model result is smaller than that in the test DW-1 case. The MOX effect on pellet temperature, which is due to the thermal conductivity and radial power profile, is negligible in the test BZ-3 case.

Fig. 13. Evolution of (a) pellet temperature, (b) PCMI pressure, cladding surface temperature, and cladding hoop strain in test BZ-3

Hypothetical calculations with artificial FGRs during pulse irradiation up to 50% were performed to investigate the effect of gas pressure on cladding deformation. The FGR was assumed to be so early and rapid that the release is completed before the peak fuel enthalpy. The temperature of the released gas is equal to the pellet average temperature, which is about 1800-1900 deg C as can be seen in Fig. 13(a). The cladding deformation is not localized and is axially and circumferentially uniform. Figure 13(b) shows evolution of PCMI pressure and cladding strain in the 40% FGR case. The PCMI pressure reaches its peak value just after the pulse irradiation at

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around 0.24 sec and results in the rapid increase in cladding plastic strain, followed by PC gap reopening before the cladding surface temperature reaches its peak value. The earlier gap reopening, which was commonly found in the cases with FGR > 5%, than that in the test DW-1 case is mainly attributed to the relatively large PCMI-induced cladding plastic strain. Although the cladding yield stress level is reduced with cladding temperature rise, the gas pressure is not high enough to facilitate significant plastic deformation, and thus creep deformation is operative during film boiling as seen in Fig. 13(b).

Fig. 14. Cladding residual strain under BZ-3 test conditions evaluated as a function of hypothetical fission gas release during pulse irradiation

Figure 14 shows calculated cladding residual strain as a function of hypothetical FGR. Since previous studies of TCs fin-effect showed that the welded TCs underestimates the cladding surface temperature depending on the cladding surface oxide layer thickness and temperature range [24], the figure also shows the result from another series of calculations in which a 100 deg C higher cladding surface temperature history (during film boiling) than that shown in Fig. 4 was used as a BC. The addition of 100 deg C is considered large enough to compensate the possible underestimation by the fin-effect. The calculation results show that 3.4% strain, which is equal to the averaged residual strain of the BZ-3 test rod over the region with strain > 1%, occurs at release fraction of about 36-37%, which is comparable with the measured FGR, 39.4%. Namely, the large cladding residual strain observed in the test BZ-3 can be understood as the result of gas-induced deformation mainly dominated by high temperature creep, which is the mechanism consistent with that assumed in high burnup UO2 fuels. 5. CONCLUSIONS High burnup MOX fuels irradiated in commercial LWRs were subjected to RIA-simulating pulse-irradiation experiments in the NSRR, and computational analysis of the experiments using the RANNS code was performed for better understanding. The present results suggest that the same failure criterion is applicable to UO2 and MOX fuels. Specifically, the following features of the MOX fuel behaviors were found.

• The results from the test performed with a coolant condition of room temperature and atmospheric pressure were consistent with those obtained in previously-performed tests with high burnup UO2 fuels regarding the rod deformation, FGR, and failure limit.

• The computational analysis of the test DW-1 indicated that the pellet temperature evolution during the pulse irradiation was similar to that of UO2 fuels. The PCMI loading on the cladding in rod-axial direction was significantly smaller than in hoop direction and resulted in relatively small PCMI-induced plastic deformation of the cladding. The PC gap reopening occurred so late as after the end of the film boiling mainly due to the small PCMI-induced cladding deformation. The pressure of the released fission gas had no effect on cladding deformation because of the rather moderate cladding temperature rise during the film boiling.

• FGR from Pu-spot alone during the pulse irradiation was estimated to be about 12% based on the SEM/EPMA data of the MOX pellet sampled from the mother rod of the test rods for BZ-2 and BZ-3. Such additional FGR from Pu-spot could account for the large FGR in the test BZ-3. Nevertheless, it cannot be concluded that the large FGR was the result of some MOX effects because relatively large FGRs were observed in the previous HT tests performed on high burnup PWR-UO2 tests, compared to RT tests. Further investigation is needed regarding the main cause of the large FGR of 39.4% observed in the test BZ-3.

0

1

2

3

4

5

6

7

8

0 10 20 30 40 50

4 8 12 16

Cla

ddin

g re

sidu

al h

oop

stra

in (%

)

Fission gas release (%)

Peak rod internal pressure (MPa)

FGR in test BZ-3

PCMI-induced strain

Strain in test BZ-3 Average over region with strain > 1%

With TC data

With TC data + 100°C during film boiling

Coolant pressure in test BZ-3

• The computational analysis of the test BZ-3 indicated that the MOX effect on pellet temperature, which is due to the thermal conductivity and radial power profile, was negligible. The PC gap reopening occurred before the cladding temperature reached its peak value mainly due to the large PCMI-induced cladding deformation. Although cladding yield stress level was reduced with the large cladding temperature rise up to about 800 deg C during the film boiling, the pressure of the released fission gas was not high enough to facilitate significant plastic deformation. The observed large cladding residual strain can be understood as the result of gas-induced deformation mainly dominated by high temperature creep, which is the mechanism consistent with that assumed in high burnup UO2 fuels.

ACKNOWLEDGEMENTS

The tests DW-1, BZ-1, BZ-2, and BZ-3 and their analyses have been conducted as part of a contract program sponsored and organized by the Nuclear and Industrial Safety Agency, Ministry of Economy, Trade and Industry. The authors are indebted to the engineers, technicians, and researchers in JAEA working on the NSRR experiments and the pre- and post-test fuel examinations in the RFEF.

REFERENCES

[1] FUKETA, T., et al., “Behavior of LWR/MOX Fuels under Reactivity-Initiated Accident Conditions”, Proceedings of Top Fuel 2009, Paris, France, September 6-10, 2009, pp.465-472 (2009).

[2] MURAMATSU, Y., et al., “Instrumentation techniques in NSRR experiments”, IAEA-TECDOC-CD-1635 (2009).

[3] SUZUKI, T., et al., “Development of high temperature capsule for RIA-simulating experiment with high burnup fuel”, Transactions of RRFM2007/IGORR, Lyon, France, Mar. 11-15, 2007, (2007).

[4] SUZUKI, M., et al., "Light Water Reactor Fuel Analysis Code FEMAXI-7; Model and Structure", JAEA-Data/Code 2010-035 (2010).

[5] SUZUKI, M., et al., “Analysis of pellet-clad mechanical interaction process of high-burnup pwr fuel rods by RANNS code in reactivity-initiated accident conditions”, Nucl. Technol., Vol. 155, pp. 282-292 (2006).

[6] HARGMAN, D., L., et al., “MATPRO-VERSION 11: a handbook of materials properties for use in the analysis of light water reactor fuel rod behavior”, NUREG/CR-0497 (1981).

[7] SIEFKEN, L., J., et al., “SCDAP/RELAP5/MOD3. 3 Code Manual: MATPRO--A Library of Material Properties for Light-Water Reactor Accident Analysis”, NUREG/CR-6150 (2001).

[8] MACDONALD, P., E., et al., “MATPRO: Version 09. A handbook of materials properties for use in the analysis of light water reactor fuel rod behavior”, TREE-NUREG-1005 (1976).

[9] GEELHOOD, K., J., et al., “FRAPTRAN 1.4: A Computer Code for the Transient Analysis of Oxide Fuel Rods”, NUREG/CR-7023 (2011).

[10] KAMIMURA, J., et al., “Thermal and Mechanical Behavior Modeling For High Burnup Fuel”, Porc. Of 2008 Water Reactor Fuel Performance Meeting, Seoul, Korea, Oct. 2008 (2008).

[11] OKUMURA, K., et al., “Validation of a Continuous-Energy Monte Carlo Burn-up Code MVP-BURN and Its Application to Analysis of Post Irradiation Experiment”, J., Nucl., Sci., Technol., Vol. 37, No. 128 (2000).

[12] FUKETA, T., et al., “Behavior of PWR and BWR fuels during reactivity-initiated-accident conditions”, Int. Topical Mtg. LWR Fuel Performance, Park city, Utah, U.S.A., Apr. 10-13, 2000 (2000)

[13] OHIRA, K., et al., “Thermal Conductivity Measurements of High Burnup UO2 Pellet and a Benchmark Calculation of Fuel Center Temperature”, Proc. ANS Topical Meeting on Light Water Reactor Fuel Performance, Portland, U.S.A., Mar., 1997 (1997).

[14] SUZUKI, M., “Present status of the verifications and model development of FEMAXI and RANNS condes in JAEA”, IAEA Technical Meeting on Fuel Behavior under Transient and LOCA, Mito, Japan, Oct. 18-21, 2010 (2010).

[15] SUGIYAMA, T., et al., “Failure of high burnup fuels under reactivity-initiated accident conditions, Annals Nucl. Energy., Vol. 36, pp. 380 (2009).

[16] FUKETA, T., et al., “Behavior of High-burnup PWR Fuels with Low-tin Zircaloy-4 Cladding under Reactivity-Initiated-Accident Conditions”, Nuclear Technology, Vol. 133, No. 1, pp. 50-62 (2001).

[17] FUKETA, T., et al., “Behavior of 60 to 78 MWd/kgU PWR fuels under reactivity-initiated accident conditions”, J. Nucl. Sci. Technol., Vol. 43, pp. 1080 (2006).

[18] SASAJIMA, H., et al., “Behavior of irradiated ATR/MOX fuel under reactivity initiated accident conditions”, J. Nucl. Sci. Technol., Vol. 37, No. 5, pp. 455 (2000).

[19] SCHMITZ, F., et al., “High burnup effects on fuel behavior under accident conditions: the tests CABRI REP-Na”, J. Nucl. Mater., Vol. 270, pp. 55 (1999).

[20] PAPIN, J., et al., “Synthesis of CABRI-RIA tests interpretation”, EUROSAFE 2003, Paris, France, Nov.25-26, 2003 (2003).

[21] LEMOINE, F., “Estimation of the grain boundary gas inventory in MIMAS/AUC MOX fuel and consistency with REP-Na test results”, J. Nucl. Sci. Technol., Vol. 43, No. 9, pp. 1105 (2006).

[22] AMAYA, M., “Fission Gas Release of LWR-MOX Fuel under RIA condition”, Fuel Safety Research Meeting 2010.

[23] SUGIYAMA, T., et al., “Evaluation of Initial Temperature Effect on Transient Fuel Behavior under Simulated Reactivity-Initiated Accident Conditions”, J. Nucl. Sci. Technol., Vol. 47, No. 5, pp.439 (2010).

[24] BESSIRON, V., et al., “Study on Transient Temperature Measurement at Fuel Clad Surface in NSRR Experiments”, JAEA-Research 2006-83, (2007).