Separate effects irradiation testing of miniature fuel specimens

Journal of Nuclear Materials 526 (2019) 151783

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Separate effects irradiation testing of miniature fuel specimens* Christian M. Petrie*, Joseph R. Burns, Alicia M. Raftery, Andrew T. Nelson, Kurt A. Terrani Oak Ridge National Laboratory, P.O. Box 2008, Oak Ridge, TN, USA, 37831

a r t i c l e i n f o

a b s t r a c t

Article history: Received 28 March 2019 Received in revised form 9 August 2019 Accepted 1 September 2019 Available online 3 September 2019

Qualification of new nuclear fuels is necessary for their deployment and requires a thorough understanding of fuel behavior under irradiation. Traditionally, nuclear fuels have been qualified by performing exhaustive integral tests under a limited range of prototypic conditions designed for their specific reactor application. While some integral fuel testing is essential, basic data on behavior and property evolution under irradiation can be obtained from separate effects tests. These irradiations could offer reduced cost, reduced complexity, and in the case of accelerated testing, reduced time to achieve a given burnup. Furthermore, it may be desirable to design test irradiations capable of deconvoluting the myriad effects of burnup, temperature gradients, and other factors inherent to integral irradiation tests. Oak Ridge National Laboratory has developed an experimental capability to perform separate effects irradiation testing of miniature fuel specimens in the High Flux Isotope Reactor (HFIR): the “MiniFuel” irradiation vehicle. The small size (<4 mm3) of the fuel specimens simplifies the design, analysis, and postirradiation examination. By reducing the fuel mass, the total heat generated inside the experiment vehicle can be dominated by gamma heating in the structural materials instead of fission heating in the fuel. This essentially decouples the fuel temperature from the fission rate, allowing for highly accelerated testing (3X 18X the burnup rate of a typical light water reactor for 235U enrichments varying from 0.22 wt% to 8 wt%) and an extremely flexible experiment design that can accommodate a wide range of fuel temperatures (~100  C to >1200  C), compositions, enrichments, and even geometries without requiring detailed analyses for each fuel variant. This paper summarizes the experiment design concept, evaluates potential applications for specific fuel forms, and briefly describes the first set of experiments on uranium nitride kernels that have been assembled and are currently being irradiated in the HFIR. © 2019 Published by Elsevier B.V.

Keywords: Fuel Irradiation Separate effects Burnup

1. Introduction A nuclear reactor core generally consists of a solid fissile-bearing material (the fuel) contained within a non-fissile material that transfers thermal energy generated via fission to a flowing coolant. Most reactor designs use solid fuels, with a few exceptions for designs that employ liquid fuels [1]. A large variety of solid fuels have been produced and tested over the decades; these fuels range from

* This manuscript has been authored by UT-Battelle, LLC, under contract DEAC05-00OR22725 with the US Department of Energy (DOE). The US government retains and the publisher, by accepting the article for publication, acknowledges that the US government retains a nonexclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this manuscript, or allow others to do so, for US government purposes. DOE will provide public access to these results of federally sponsored research in accordance with the DOE Public Access Plan (http://energy.gov/downloads/doe-public-access-plan). * Corresponding author. E-mail address: [email protected] (C.M. Petrie).

https://doi.org/10.1016/j.jnucmat.2019.151783 0022-3115/© 2019 Published by Elsevier B.V.

simple configurations of monolithic ceramics and metals clad in plates or tubes to complex configurations of multilayer encapsulated fissile kernels compacted into various shapes [2]. However, only a few fuel systems are readily available for deployment into advanced reactors. This limited number of fuel systems is due to the long and costly timeline required for fuel development and qualification [3]. Furthermore, once a qualified fuel system is available, its application is limited to the strict operational regime specified by its long and rigorous testing campaign. This means that advanced fuel systems are not readily available for implementation by reactor designers, which stifles innovation towards development of advanced nuclear energy systems. Advanced fuels for the existing light water reactor (LWR) fleet and for advanced reactor systems are actively sought today. A prominent example of the former is the development of accident tolerant fuel technologies [4]. Enhancements to traditional UO2 are also being considered that would increase fuel thermal conductivity and fission product retention through the use of various

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additives [5e11]. For advanced systems that target passive safety and high resource utilization targets, high density monolithic fuels such as U-metal alloys (e.g., UeMo and UeZr), uranium carbide (UC), uranium nitride (UN), and uranium silicide (U3Si2) offer increased uranium density and enhanced thermophysical properties as compared to conventional UO2 fuel all while maintaining an acceptably high melting point [12e18]. Coated particle fuels such as tristructural isotropic (TRISO) fuels are being considered for several advanced reactor concepts and even for some LWR applications [19e24]. However, in all these instances, the information available on the fuels’ irradiation performance (microstructural evolution and stability, swelling, fission gas release, etc.) is limited to a narrow range of temperature and burnup. This lack of information impedes the ability to readily adopt these fuels in advanced reactors. Conspicuous examples from a long list of stalled efforts include: 1. The ability to adopt TRISO fuel particles qualified for hightemperature gas reactors [24] in molten saltecooled reactors that intend to utilize them at higher powers and lower temperatures, 2. The applicability of UeSi compounds for application at temperatures well above those characteristic of their historic research reactor application [25], and 3. Understanding and verifying the suspected effect of impurities on the performance of carbide and nitride fuels [26] and expanding their application beyond the limited experience in sodium fast reactors [27] to a range of liquid metal or gas-cooled systems. Traditionally, irradiation performance data have been acquired through many integral fuel irradiation experiments in which fullsize fuel pellets, compacts, or pebbles are tested under conditions that closely match those of the intended application. While some integral fuel testing is necessary, it is expensive and time consuming, particularly when considering a large test matrix that could include variations in fuel centerline temperature, burnup, and power history or variations in the fuel itself such as composition, enrichment, grain size, impurities, or non-stoichiometries. Furthermore, the large number of variables that affect fuel performance make it difficult to develop fundamental models of various phenomena from integral fuel tests. These tests typically have many independent variables that cannot be well controlled. The process of designing and executing irradiations to isolate a single variable to gain a better understanding of its influence on fuel performance has become known as the separate effects approach. To illustrate the complexity of integral fuel experiments, Fig. 1 shows general trends in fuel fission rate and temperature during LWR integral fuel testing as compared to ideal trends for a separate effects irradiation test. The thickness of the fuel temperature curves indicates the temperature gradients through the fuel. This figure highlights the relatively large temperature gradients present during integral fuel testing, as well as the strong dependence of fuel temperature on fission rate. In addition to temperature gradients, there are also large burnup gradients under thermal spectrum irradiations as a result of self-shielding. Other phenomena that impact fuel temperature during integral fuel testing are indicated in Fig. 1 (a), including fuel densification and swelling, burnupdependent properties such as thermal conductivity, cladding creep-down, pellet cracking and relocation, and fission gas release (FGR). These complexities make it difficult to develop fundamental models for individual phenomena such as fuel swelling or FGR. Separate effects irradiation testing could be used to deconvolute the myriad effects of burnup, temperature gradients, and other factors inherent to integral irradiation tests. The ability to isolate

Fig. 1. Plots showing general trends in fuel fission rate and temperature during (a) LWR integral testing and (b) separate effects testing.

these variables during irradiation testing of a wide range of fuel concepts within a reasonable time and cost would be highly beneficial to the nuclear community. The irradiation tests could be performed with or without intentional temperature gradients while minimizing any changes in the minimum fuel temperature (i.e., boundary condition) over the course of the experiment. The remainder of this paper primarily discusses experiments that target isothermal conditions. However, there are instances where intentional temperature gradients could be introduced to evaluate, for example, the effects of temperature gradients on fuel swelling, cracking, redistribution, or fission gas release. The fission rates could also vary over time due to burnup effects or be held constant over the course of the experiment. Many of these tests could be accelerated to quickly accumulate burnup. In addition, the testing platform should be flexible so that a range of fuel compositions, heavy metal densities, enrichments, and even geometries could be tested without requiring detailed designs and analyses specific to each fuel concept. To this end, this work describes a new “MiniFuel” vehicle for performing accelerated separate effects irradiation testing of small fuel specimens in the High Flux Isotope Reactor (HFIR) at Oak Ridge National Laboratory (ORNL). This paper describes the experiment design concept and the neutronic and thermal analyses, and it

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evaluates potential applications for specific fuel forms. This paper also briefly describes the first set of experiments testing UN fuel kernels at LWR-relevant fuel temperatures.

2. MiniFuel irradiation vehicle architecture The unique design of the HFIR [28] presents challenges and opportunities in the pursuit of fuel irradiation experiments. The HFIR core schematic shown in Fig. 2 indicates the various experimental positions in the flux trap, the removable beryllium reflector, and the permanent beryllium reflector. More details regarding the layout of the HFIR core are provided in literature [29e31]. The HFIR typically operates 6e7 cycles per calendar year, with each cycle lasting 23e26 days. Given its envisioned mission of isotope production, the HFIR's flux trap design provides an extremely high thermal neutron flux of ~2  1015 n/cm2/s in the vicinity of the core. This implies that even when using depleted uranium coupons, the initial power density is significantly higher than the typical LWR application with a thermal neutron flux on the order of 1013 n/cm2/ s. Furthermore and unlike most other reactors, the HFIR's high neutron flux breeds additional fissile isotopes, resulting in prohibitively large increases in power as the irradiation progresses. This renders the HFIR flux trap as a non-ideal location for most fuel irradiation experiments, with a few exceptions for specific tests targeting fast reactor applications. While the HFIR flux trap can provide a relevant fast neutron spectrum using thermal neutron shields such as Gd or Cd shrouds [32], this increases experiment complexity and ultimately limits the total burnup that can be accumulated before the shield is consumed. The neutron flux conditions in the Be reflector (5  1014 nthermal/cm2/s and 7  1013 nfast/cm2/s) are more favorable for irradiating fuels, particularly for thermal reactor applications. The MiniFuel irradiation vehicle design consists of small fuel specimens irradiated in discrete capsules in the inner small vertical experiment facility (VXF) positions to provide an ideal separate effects testing platform. The MiniFuel design was inspired by UO2 disk irradiation experiments at the Halden reactor [33,34] that allow for uniform power and temperature profiles. Threedimensional (3D) computer aided design (CAD) models of the MiniFuel experiment architecture are shown in Fig. 3, including small fuel specimens enveloped in cups contained inside sealed

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capsules that are further encapsulated inside sealed targets. The targets are housed inside a basket through which the reactor's primary coolant flows. The basket assembly may be placed in any of the 11 inner small VXF positions in the HFIR reflector. Each basket contains three vertical channels in different radial positions. This allows for radial positions 2 and 3 to be radially equidistant from the center of the core. These positions are expected to receive essentially the same neutron flux and gamma heating. Radial position 1 faces away from the core. Three targets are stacked vertically in each of these radial positions, so a total of nine targets may be accommodated in each basket. Axial target positions 1, 2, and 3 correspond to the bottom, middle, and top positions, respectively. Each target contains six sealed capsules that contain small fuel specimens inside a cup, a high-Z metal filler, and passive SiC thermometry. The first-generation MiniFuel capsules are constructed from Grade 5 titanium (Tie6Al4V) and the fuel specimens are surrounded by a molybdenum cup similar to that used in previous Halden experiments [34]. The fuel cup is designed to physically hold the fuel while minimizing chemical interaction, and it allows for future testing of potential fuel chemical interaction with a nonfissile constituent (such as fuel cladding in the targeted application). The high-Z filler generates most of the heat in each capsule due to heating from gamma photons emitted by the HFIR fuel. Finally, inclusion of SiC specimens allow for post-irradiation dilatometric evaluation of the temperature during irradiation [35e37]. While the dilatometric technique does not require that the irradiated SiC temperature monitors reach saturation swelling, the magnitude of the swelling does affect the confidence in the estimated temperature. However, because SiC reaches swelling saturation at low dose, the relatively low dose rate (~0.07 dpa per HFIR cycle) for the MiniFuel experiments is not expected to be an issue. For example, at 500  C and 0.21 dpa (3 HFIR cycles), the volumetric swelling in SiC is calculated to be 1.1%, or 83% of the saturation swelling [38]. Active temperature instrumentation could also be incorporated in future tests. As shown in Fig. 3, the capsules may contain a variety of fuel specimen configurations: (1) a single nominally ø3 mm  0.3 mm thick monolithic fuel disk, (2) six nominally ø800 mm bare fuel kernels, (3) four nominally ø1.2 mm coated TRISO fuel particles, and (4) 21 nominally ø425 mm bare fuel kernels. Other configurations can be accommodated as needed for experiments with specific geometric requirements. A small hole in the capsule's end cap is used to perform the final seal weld after the capsule is filled with helium. After the capsules are assembled inside the target, a similar seal weld is performed on the target housing. Fuel temperatures are controlled by varying the composition and size of the gas gap between the capsules and the target. The size and composition of the gas gap required to achieve a given fuel temperature depends primarily on gamma heating in the capsule components. Thimbles on either end of the capsules ensure that the capsules remain centered within the target. Compression springs at both ends of the target keep the capsules stacked. 3. Description of the first MiniFuel experiment

Fig. 2. HFIR core schematic showing experiment positions and horizontal beam (HB) tubes. The MiniFuel experiments will be placed in the inner small vertical experiment facility (VXF) positions.

This first MiniFuel experiment tested uranium nitride fuel with low amounts of solute carbondU(C,N)din the form of bare kernels [39e41] and TRISO particles [42]. The goal is to evaluate the basic performance of the sol-gelederived U(C,N) with low C content in the bare and coated conditions and compare its performance with prior computational results. Although irradiation of U(C,N) microspheres has been conducted before [43], this is the first irradiation of U(C,N) TRISO since it was first produced [42]. The irradiation temperatures of 500e600  C are within the range of the expected normal operating conditions of fully ceramic microencapsulated

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Fig. 3. Irradiation design concept.

fuel with UN TRISO particles in a LWR [19]. An image of a polished cross-section of high density (96.8% TD) UC0.20N0.80 kernels is shown in Fig. 4, and the test matrix is summarized in Table 1. The higher density kernels were densified using a hot isostatic press [40]. For the remainder of this paper, the U(C,N) kernels tested in this work are generally referred to as “UN” kernels. All fuel specimens consist of either natural or depleted uranium. Assessment of varying enrichment levels is discussed in Sections 5.1 and 5.2. Several different carbon impurity levels and kernel densities are included. Some kernels include burnable absorbers (Gd) [44] with varying contents. This test matrix highlights the utility of the MiniFuel capability: the effects of impurities, burnable absorbers, and density are being evaluated in a single irradiation target, and the addition of a second target allows for evaluation of two different burnup levels. Irradiation of these kernels started in June 2018 [45]. The low-burnup target completed irradiation in November 2018, and the high-burnup target is scheduled for completion in winter 2020. Initial post-irradiation examination (PIE) of the low-burnup target is scheduled for calendar year 2019. 4. Analysis methodology 4.1. Neutronics analysis Neutronics calculations are carried out to assess neutron and gamma heating rates in the fuel and capsule components to

Fig. 4. Polished cross-sectional image of UC0.20N0.80 kernels with 96.8% TD after densification using a hot isostatic press [40].

provide input to subsequent thermal analyses, as well as to generate predictions of fuel burnup and fission gas inventory over the course of the experiment. This is accomplished using the MCNP5 [46] and SCALE [47] software code packages. The MCNP calculations are based on existing beginning-of-cycle (BOC) and end-of-cycle (EOC) models [29] of the HFIR. The models include the experiments that were inserted during the 400th HFIR cycle (April through May 2004) with additional modifications to include the MiniFuel experiments. The heat generation analyses account for contributions from fission neutrons (prompt and delayed), prompt fission photons, delayed photons from fission product decay, alpha and beta decay heat, and photon heating from local activation product decay. Prompt fission neutron and photon heating is calculated directly from MCNP transport simulations using an established fission neutron source distribution definition with both neutron and photon tracking activated to implicitly yield an appropriate fission photon source distribution [48]. Heat generation from these sources is tallied in all experiment components. To account for heat generated from fission product decay photons that originated in the HFIR fuel, a separate calculation is performed with a fixed photon source distribution reflecting the gamma emission rate and spectrum due to these accumulated fission products [48]. Activation and decay calculations are carried out using the ORIGEN module of the SCALE code package. To achieve consistency between the MCNP and ORIGEN calculations, problem-specific ORIGEN cross section data are generated from MCNP reaction rate and flux tallies in the fuel samples under representative middle of cycle conditions. These effective cross sections are provided to the COUPLE cross section processing module to evaluate burnup and transmutation over many cycles using ORIGEN. The ORIGEN calculations yield local alpha and beta decay heating from activation products, as well as activation product gamma emission rates and spectra. This latter information is used to construct a local activation gamma source distribution for a final MCNP calculation assessing heat generation in the experimental capsule components from activation gammas. All alpha and beta decay heat is assumed to be deposited locally. This methodology is implemented as described above to conduct a thorough assessment of heat generation and burnup accumulation for the first cycle of irradiation. The same ORIGEN models developed for the activation and decay calculations in the single-cycle analysis can be used to extend the calculations of fuel fission heating, burnup, and decay heat for multiple irradiation cycles. This is accomplished by fixing the effective microscopic cross sections calculated at the midpoint of the first cycle and

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Table 1 Test matrix for first MiniFuel irradiation experiment. Target burnup (MWd/kg U)

Target temperature ( C)

0.22%

8.7, 60.0

500 600  C

0.71%

12.0, 63.9

Fuel form

Kernel theoretical density

235

UC0.20N0.80 kernels UC0.20N0.80 kernels UC0.20N0.80 TRISO particles UC0.15N0.85 kernels U0.98Gd0.02C0.15N0.85 kernels U0.89Gd0.11C0.11N0.89 kernels

96.8% 90.9% 87.2% 90.6% 93.6% 92.0%

U Enrichment (wt %)

simulating fuel burnup for a total of 30 26-day HFIR cycles, with 15day decay periods assumed between cycles. The assumption that the effective cross sections in cycle 1 are applicable to later cycles is reasonable, given that the small size of the fuel samples inherently limits changes in the incident neutron flux energy spectra between cycles. Statistical uncertainties in the fuel heat generation rates (HGRs) and isotopic transmutation are minimized by running a highly rigorous (~1010 particle histories) MCNP simulation to calculate the effective cross sections. The ORIGEN output reports the total fission heat generation and accumulated burnup at the beginning and end of each irradiation cycle, as well as decay heat during and between cycles. Future work will investigate other longterm burnup and depletion calculation methodologies to assess the validity of using cycle 1 cross sections to simulate 30 cycles of irradiation. 4.2. Thermal analysis Thermal finite element calculations are performed using the ANSYS software code package with custom macros for determining thermal contact conductance between components and heat transfer through small movable (due to thermal expansion) gas gaps [49]. The 3D CAD model for a single irradiation target is imported directly into ANSYS after applying ¼-symmetry. Minor components such as the target end caps, compression springs, fillets, and welds were removed. The remaining features are meshed with 20-node hexahedral 3-D thermal solid elements with a nominal mesh size of 0.4 mm, except in the fuel, where the mesh size is 0.15 mm. Internal heat generation is applied to all components using the HGRs calculated in Section 5.1. A convection boundary condition is applied to the outer surface of the target housing. A value of 44.8 kW m 2 K 1 is used for the convection heat transfer coefficient, and 58  C is used for the bulk coolant temperature. These numbers were determined using RELAP5 for a previous fueled VXF irradiation experiment using the same facility with identical flow geometries, and therefore identical flow velocities [50]. The only difference between the MiniFuel experiments and the previously analyzed experiments is the heat flux at the target outer surface, which could change the bulk coolant temperature and the target surface temperature. However, the increase in bulk coolant temperature through the channel was <8  C for the previous experiment, which had a much higher heat load. Therefore, the bulk coolant temperature is not significantly affected by the experiment heat load. The convective heat transfer coefficient calculated using the Sieder-Tate correlation is a function of the Reynolds and Prandtl numbers and the water thermal conductivity, which do not change significantly for small variations in coolant temperature. There is also a 0.14 power dependence on the ratio of the coolant viscosity at the bulk coolant temperature to that at the target outer surface temperature. The calculated peak target surface temperature for MiniFuel using the assumed 44.8 kW m 2 K 1 heat transfer coefficient is 61  C compared to the previously calculated surface temperature of 77  C. Evaluating the change in coolant viscosity at the

slightly lower MiniFuel target surface temperature would result in at most a 3% change in the convective heat transfer coefficient, indicating that the assumed heat transfer coefficient is reasonably accurate. The effects of fuel swelling are not considered in the thermal analyses. However, these effects are not expected to be significant for the fuels included in the initial test matrix since the fuel specimens are small (<3 mm3 per capsule), and heat is primarily transferred via conduction from the bottom of the fuel specimens to the cup and then through the capsules. Thermal contact resistance is considered between all contacting components and radiation heat transfer is included. FGR and the resulting effect on heat transfer within the capsules is considered. For simplicity, it was assumed that 23% of fission products (two fission products per fission) are gaseous [51] and that all gaseous fission products are xenon. Krypton isotopes are ignored in the heat transfer analysis because xenon constitutes the largest component of the fission gas, particularly for 239Pu fission, for which the Xe/Kr yield ratio is ~13.6. Even for relatively high burnup (up to 100 MWD/kg U) and large FGR (10%), the gas composition within the capsules remains greater than 83% He for the largest fuel volume (ø3.2 mm  0.3 mm thick disk). The 17% reduction in He content only impacts heat transfer within the capsules. It is later shown in Sections 5.2 and 5.3 that temperature gradients within the capsules are small compared to the much larger temperature difference through the primary gas gap. Therefore, variations in the assumed gaseous fission product yields would not significantly impact fuel temperatures. The heat generated in the capsules passes through the primary gas gap between the capsules and the target housing. The resulting temperature difference between the capsules and the target housing depends on the heat flux and the size and composition (which determines the thermal conductivity) of the primary gas gap. The fill gas composition inside the target and the outer diameter of the capsules are varied to achieve the desired fuel temperatures. The effects of the capsules’ thermal expansion are considered in the thermal contact conductance through the gap. Because the fuel specimens are individually sealed in He-filled capsules, FGR does not affect the temperature difference through the primary gas gap if the capsule welds do not fail. Temperature-dependent material properties are used for all materials. Table 2 summarizes the materials included in the thermal model and the references used for the material properties.

Table 2 Materials and property references for thermal analyses. Material

Components

Property references

Titanium Silicon carbide 304 stainless steel Molybdenum UN UO2 U3Si2 Grafoil

Capsules, centering thimbles Thermometry Target housings Filler, cups Fuel kernels, disks Fuel kernels, disks Fuel disks Insulators

[52e54] [53,55] [53] [53,54] [56e61] [62,63] [10,64e66] [67]

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Dose- or burnup-dependent properties are used for SiC, UN, and UO2. Because this experiment is conducted in HFIR's reflector positions, the neutron flux energy spectrum is highly thermal. Therefore, the displacement damage rates are relatively low (~0.06 dpa per HFIR cycle, or ~0.45 dpa per calendar year in steel), and the effect on the thermal properties of the structural components is not expected to be significant. Some of the fuel properties include porosity dependence.

5. Results

Table 3 Calculated peak HGRs in structural materials at BOC and EOC in radial target position 2. Component

Stainless steel target Titanium capsules and centering thimbles Molybdenum cup Molybdenum filler Silicon carbide thermometry Grafoil

Peak HGR (W/g) BOC

EOC

4.52 ± 0.04 4.87 ± 0.07 6.02 ± 0.18 5.78 ± 0.09 3.99 ± 0.18 3.92 ± 0.14

4.92 ± 0.05 5.46 ± 0.08 6.78 ± 0.19 6.48 ± 0.10 4.35 ± 0.20 4.35 ± 0.14

5.1. Neutronics analysis Fig. 5 shows predicted HGRs in the titanium capsules (also used for the end caps and thimbles) and the molybdenum filler as a function of distance from the core midplane (z). These components, along with the molybdenum cups, make up most of the mass of each capsule. Results are shown at BOC and EOC along with an exponential fit to each data set. The HGRs are shown for radial target position 2, which faces toward the HFIR core. Radial target position 3, which has the same radial distance from the center of the core, has very similar HGRs. Gamma heating rates are reduced by ~10% in radial target position 1 compared to radial target positions 2 and 3 because position 1 is further from the center of the core. The increase in HGRs from BOC to EOC is due to the withdrawal of the HFIR control plates throughout the cycle and the resulting radial shift in the location of the peak fission density within the HFIR fuel. As the peak fission density moves radially outward, gamma heating in the experimental positions within the reflector also increase. The structural material HGRs are calculated during the first cycle of irradiation with 0.35% 235U UO2 disk fuel. While variations in fuel geometry, composition, and enrichment could have some impact, the small fuel size limits its contribution to structural heating, which is dominated by heating from neutrons and photons emitted by the HFIR fuel. Scoping calculations indicate that the fuel samples' net contribution (via fission neutrons and photons) to heating in any structural component is less than 10 3 W/g. Table 3 summarizes the peak (i.e., at the core midplane) HGRs for all materials at BOC and EOC in radial target position 2. All thermal analyses described in this paper are performed in radial target position 2 in

Fig. 5. Calculated HGRs vs. axial distance from the core midplane (z) at BOC and EOC in radial position 2. Exponential fits to the calculated data are shown as solid (BOC) and dashed (EOC) lines with coefficient of determination R2.

the center axial target. Uniform fuel temperatures can be achieved in all six capsules within a given target, even in the top and bottom axial target positions, despite the large gradients in structural HGRs. This is accomplished by varying the capsules’ outer diameters, thus varying the individual capsule-to-target gas gaps. Structural HGRs generally increase by 10e15% from BOC to EOC. These increases are accounted for in the thermal analyses. Structural material HGRs are dominated (~85e90%) by prompt gamma heating from photons emitted from the HFIR fuel. Delayed gamma heating from photons emitted from radioactive decay of fission products within the HFIR fuel account for ~5e10% of the total heat generation. Other minor contributions to the structural material HGRs include prompt neutron heating and decay heat from beta and gamma emission resulting from neutron activation of the structural materials. Fuel HGRs change significantly over time due to burnup of the initial 235U and breeding of fissile Pu isotopes. Fig. 6 shows fuel fission heating rates and accumulated burnup in UN fuel vs. irradiation time and the number of HFIR cycles for five different enrichment levels. As mentioned previously, this analysis assumes 26-day HFIR cycles. The data in Fig. 6 are for fuel loaded in radial target position 2 at the core midplane (axial target position 2). Gamma heating in the fuel at this location contributes an additional ~8e10 W/g, and local decay heat within the fuel is generally ~5% of the fission heating. These contributions are not included in the fission heating rates shown in Fig. 6, but they are included as inputs to the thermal analyses performed in Sections 5.2e5.4. The fuel fission rates increase over the course of each HFIR cycle, even after reaching equilibrium Pu isotope concentrations due to variations in neutron flux caused by movement of the HFIR control plates. The fission rates vary by approximately 10 W/g after reaching equilibrium. To compare the calculated fission rates with those of an LWR, additional curves are shown for a constant linear heating rate of 20 kW/m, assuming 95% TD UO2 fuel with an 8.4 mm pellet diameter. The fuel enrichment can be selected by balancing accelerated burnup accumulation with fission rate fluctuations and feedstock availability. If the goal is to minimize fluctuations in fission rate (consistent with the separate effects approach), then naturallyenriched fuel should be used. Natural uranium provides enough initial heating from fission of 235U to nearly offset the delay in breeding of Pu isotopes. The result is a fuel fission heating rate that is nearly constant at ~150 W/g over 30 HFIR irradiation cycles. With depleted uranium the fission rate increases significantly over the first 3 or 4 HFIR cycles due to the time it takes to breed 239Pu. However, even the initial fission rate with depleted uranium is similar to that of an LWR operating at 20 kW/m. For low-enriched fuel (235U wt% > 0.73%) the fission rate decreases over time due to burnup of the initial 235U. For new fuel systems with a low technology readiness level, there could be some incentive to increase the enrichment to rapidly accumulate burnup and provide initial data on irradiation performance at the expense of fission rate

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[68]. The general trends are the same for all fuel forms. For a given enrichment, burnup evolution is identical for all fuel compositions because it is specified per unit mass of uranium. Fuel fission rates are specified per unit mass of the fuel compound (e.g., per unit mass of UN). Therefore, the fuel fission rates vary depending on the uranium density of the fuel form. The data in Fig. 6 can be scaled by uranium density to determine fuel fission rates for any uraniumbearing fuel form. Table 4 summarizes the minimum and maximum fuel fission rates for various fuel forms and enrichments. For 235U enrichments of 2.5 wt% or greater, the maximum fission rate occurs at the beginning of the first cycle due to the large contribution from 235U fission. For depleted (0.22% 235U) and natural (0.73% 235U) uranium, the contribution from 235U fission is much lower, and the maximum fuel fission rate occurs after ~200 days of irradiation due to breeding fissile Pu isotopes. Beyond 200 days, the fission rates are largely unchanged except for a very slow decline due to gradual burnup of 239Pu after conversion from 238U. The minimum fission rate for fuels with depleted and natural uranium occurs at the beginning of the experiment before fissile Pu isotopes have been bred. 2. Thermal analysis of UN kernels

Fig. 6. Calculated fuel fission heating (Q, solid lines) and burnup (BU, dashed lines) vs. irradiation time and number of HFIR cycles for different enrichment levels. Results are for UN fuel located at the core midplane in radial target position 2. The plot in (b) shows the same data in (a) with a reduced range of Q to show more detail after the fuel fission rates approach an equilibrium. The “LWR” curves show Q and BU for a light water reactor operating with a linear heating rate of 20 kW/m.

fluctuations over the first 100e200 days of irradiation. For all fuels, fission rates do not change significantly after ~200 days of irradiation (6 HFIR cycles), at which point most of the 235U has been burned, and an equilibrium has been reached between breeding and burning of Pu isotopes. Based on these results, it is anticipated that LWR discharge burnups of 60 MWd/kg U can be achieved within 15 HFIR irradiation cycles or less, even for fuel fabricated from depleted uranium. Lesser burnups of 10 MWd/kg U can be achieved within 3e4 HFIR cycles. While the fuel heating per unit mass is more than an order of magnitude greater than the heating rates in the structural components (see Table 3), the fuel mass is much smaller than the masses of the structural materials. For example, the DU UN kernels that are currently being irradiated in the HFIR generate only 7.8% of the total heat in one capsule at EOC 15. This is particularly important in decoupling the fuel temperature from the fuel fission rate. The fuel fission rate and burnup calculations are repeated for other fuel forms besides UN, including UC, U3Si2, and UO2. A special case is also run for two-phase uranium oxide, uranium carbide (UCO) fuel enriched to 14 wt% 235U, which is representative of existing AGR-2 fuel that could be tested using the MiniFuel vehicle

The UN kernels currently being tested (see test matrix in Table 1) are evaluated using ANSYS thermal simulations to predict temperature evolution over the course of the experiment. For this, simulations are performed at BOC and EOC for the first (cycle 1) and last (cycles 3 and 15) HFIR cycles during which these kernels will be irradiated. Analyses are performed for dense (95% TD) UN kernels with 0.22 wt% 235U and lower density (90% TD) UN kernels with 0.73 wt% 235U. FGR from the UN kernels to the free volume inside the capsules is assumed to be 3% for these analyses. This assumption is based on the model proposed by Storms for UN using the Thomas data set [69]. The Storms model predicts 1.2% release for UN with 95% TD, a burnup of 6% fission of initial metal atoms, and a temperature of 800 K. Adding the expected recoil release of 1.8% for this geometry [70] gives a total release of 3%. For UN kernels with a TD of 87% (lowest TD in the current experiment), the predicted FGR increases to 3.7%. Varying FGR from 0 to 10% results in a maximum deviation of 21  C in peak fuel temperature, indicating that the assumed value for FGR does not significantly impact fuel temperatures for this experiment. Fig. 7 shows temperature contours for a target containing capsules, each of which is filled with 95% TD UN fuel kernels with depleted uranium (0.22 wt% 235U). The temperatures are calculated at EOC 15, at which point the fuel HGR is 180 W/g and burnup is 60 MWD/kg U. The irradiation targets are filled with a 40.5% HeeAr balance mixture, and the nominal radial gap between the capsule's outer diameter and the target's inner diameter is 284 mm. Despite the accelerated nature of the testing (i.e., high fuel fission rates), the temperature gradients in the fuel are relatively low (22  C) because of the small fuel size. The passive SiC temperature monitors have an average temperature of approximately 483  C (83  C lower than the fuel temperatures), with temperature gradients of only 2  C. The passive SiC temperature monitors will be used to confirm the irradiation temperatures. Fig. 7 shows that the capsule's outer surface temperatures are approximately 440  C, which means that the primary temperature difference within the experiment occurs between the capsules and the ~63  C irradiation target (not shown in Fig. 7), which is directly exposed to the reactor coolant. This temperature difference depends on the total heat flux passing through the capsule-to-target gap. If the gamma heating in the structural components continues to be much greater than the total fuel heating, then this

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Table 4 Maximum and minimum fuel fission heating for various fuel forms and enrichments. 235

U enrichment (wt %)

0.22% 0.73% 2.50% 5.00% 8.00% 14.00% a b

Maximum/minimum fission heating (W/g) UN

UC

U3Si2

UO2

UCO

163/36 163/109 361/139 716/137 1143/135 1,994/128b

163/36 163/109 363/140 721/137 1152/135 2,010/129b

159/35 159/106 354/136 703/134 1122/131 1,958/126b

151/33 152/101 337/130 669/128 1068/125 1,861/120b

154/34a 154/103a 342/132a 679/129a 1084/127a 1,889/122

Calculated from fission heating in UO2 scaled by uranium density. Calculated from fission heating in UCO with 14% 235U scaled by uranium density.

Fig. 7. EOC 15 temperature contours (in  C) obtained from a model with ¼ symmetry for 95% TD UN fuel kernels with depleted uranium (0.22 wt% components (top), a single center sub-assembly (middle), and a fuel kernel from the center sub-assembly (bottom).

235

U). Results show all target

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temperature difference will not change significantly as the fuel fission rate evolves. However, the temperature difference between the fuel specimen and the cup will depend on the fuel's HGR and on the total FGR, which determines the conductivity of the gas between the fuel and the cup. This temperature difference is approximately 63  C at EOC 15. Placing spherical fuel kernels inside hemispherical holes in the cup results in theoretical point contact, which represents the worst-case geometry for fuel-to-cup contact. Fig. 8 includes circular markers indicating the average kernel temperatures at BOC and EOC for cycles 1, 3, and 15. The solid color regions surrounding each marker indicate the minimum and maximum fuel temperatures at each time. Results are shown for two fuels: (1) UN kernels with 0.22 wt% 235U and 95% TD, and (2) UN kernels with 0.73% 235U and 90% TD. Fig. 8 shows the extreme temperatures to which the fuel will be subjected over the course of the irradiation experiment. All fuels will be subjected to thermal cycling each time the reactor is shutdown and the fuel eventually cools to approximately room temperature before the next HFIR cycle is initiated. The depleted uranium specimens have a lower temperature of 442  C at BOC 1 because there has not been enough time for breeding of Pu to occur. At EOC 1, fuel temperatures increase due to a combination of breeding of Pu isotopes, which in turn leads to increased fuel HGRs, and an increase in structural material HGRs due to movement of the HFIR control plates. By BOC 3, significant breeding of Pu has occurred, and fuel temperatures do not vary significantly between cycles 3 and 15. After cycle 3, the fuel temperature varies by 40e45  C for each cycle. These variations are primarily due to variations in the structural material HGRs caused by movement of the HFIR control plates. The natural uranium specimens show similar behavior to that of the depleted uranium specimens, except that the fuel temperatures at BOC 1 are higher for the natural uranium specimens. This is because of the increased fission heating due to the higher initial 235U content. The result is that the average temperature of the natural uranium samples remains constant to within 73  C over the entire irradiation. 5.3. Thermal analysis of disk fuel The UN kernels evaluated in Section 5.2 are sol-gel derived. The more conventional approach for fabricating ceramic fuels is to press and sinter the fuel into pellets. As such, there is motivation to be able to test fuel in the form of small pellets or disks. Thermal analyses such as those described in Section 5.2 are performed for UO2

Fig. 8. Temperature evolution of UN kernels with depleted and natural uranium over 15 cycles of irradiation. Circular markers indicate the average kernel temperatures at BOC and EOC for cycles 1, 3, and 15. The solid color regions surrounding each marker indicate the minimum and maximum fuel temperatures for each data point. (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.)

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and U3Si2 disks. These analyses assume disk dimensions of ø3.2 mm  0.3 mm thick and 95% TD. FGR is assumed to be 3% for these analyses, an assumption that is somewhat arbitrary. FGR in UO2 will be significantly lower than the 3% determined based on UN kernels with a larger ratio of surface area to volume. However, as mentioned in Section 5.2, the fuel temperatures are not very sensitive to the assumed FGR value for such a small amount of release. Fig. 9 shows calculated temperature contours for a UO2 disk specimen with natural uranium at EOC 15 (63.5 MWD/kg U burnup and 168.4 W/g fuel HGR). This simulation, which has the same target temperature of 500e600  C as the UN kernels, uses the same fill gas and capsule diameters to show the variation in fuel temperatures using a different fuel geometry and composition. Fuel temperatures are 510e539  C, which are slightly lower than those for the UN kernels (549e573  C). The average cup temperature (484  C) is identical for both fuel forms. This is because the total heat generated in the capsules is dominated by gamma heating in the structural components, and the variations in fuel geometry and composition do not impact the cup temperature. The difference in fuel temperature is due to the larger thermal contact resistance for kernels in a hemispherical hole in the cup vs. flat disks in direct contact with the flat surface of the cup. Additional simulations are performed for UO2 and U3Si2 fuel disks. Multiple enrichments are evaluated to determine the effect of enrichment on fuel temperatures. Fig. 10 summarizes fuel temperatures at BOC and EOC for cycles 1 and 15, which covers the extreme temperatures to which the fuel would be subjected. For the same 0.73 wt% 235U enrichment, UO2 and U3Si2 disk fuels show similar temperatures. As expected, U3Si2 has smaller temperature gradients within the specimen due to its higher thermal conductivity. The maximum variations in the average temperature of the naturally-enriched UO2 and U3Si2 disk specimens over 15 cycles of irradiation are 67 and 63  C, respectively. Variations in average fuel temperature over 15 cycles of irradiation are slightly higher in UO2 vs. U3Si2 because of the lower thermal conductivity of UO2, which makes the spatial temperature gradients in the UO2 specimens more sensitive to variations in fuel fission rate. The effect of thermal conductivity is most obvious when comparing the spatial temperature variations in UO2 fuel vs. those in U3Si2 fuel, particularly during cycle 15, at which point the thermal conductivity of UO2 is further degraded due to burnup effects. However, even for UO2 with 8 wt% 235U (118 MWd/kg U burnup at EOC 15), the maximum temperature variation within the specimen is only 35  C at EOC 15. The data in Fig. 10 also show how significant the temperature increases are at the beginning of the experiment when using higher enrichments. Naturally enriched UO2 fuel disks show a lower average temperature of 462  C at BOC 1. The average temperature increases to 514  C at EOC 1 and remains within 30  C of 500  C at BOC and EOC 15. For 2.5 wt% 235U, the fuel temperature peaks at EOC 1 with an average temperature of 607  C. Increasing the enrichment to 5 and 8 wt% 235U causes the peak temperature to occur at BOC 1, and the average temperature exceeds 1000  C for 8 wt% 235U. Furthermore, the temperature differences within the specimen become much larger when the enrichment is increased to 5 and 8 wt% 235U. Combining the results of Figs. 6 and 10, one can select an enrichment that offers a balance between accelerated burnup accumulation vs. temperature variations at the beginning of the experiment. For tests that seek to understand the impact of temperature gradients within the fuel, both the fuel thickness and enrichment can be varied to achieve the desired temperature gradient. A significant percentage of the temperature increases at BOC 1 with higher enrichments occur at the contact interface between the fuel disk and the cup. In the case of an 8 wt% 235U disk, the

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Fig. 9. EOC 15 temperature contours (in ⁰C) obtained from a model with ¼ symmetry for 95% TD UO2 disk fuel kernels with natural uranium (0.73 wt% 235U). Results show all target components (top), a single center sub-assembly (middle), and a fuel disk from the center sub-assembly (bottom).

temperature difference between the bottom of the fuel and the top of the cup is 173  C. For some applications, the thermal conductance at this interface could be much higher, which would greatly reduce the fuel temperature. For example, filling the capsules with a liquid metal or salt could increase the gap conductance by more

than an order of magnitude. Alternatively, some fuel kernels could be compacted in a graphite (advanced gas reactor fuel) or SiC (fully ceramic microencapsulated fuel) matrix, which effectively eliminates gap conductance effects.

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Fig. 10. Temperature evolution of UO2 and U3Si2 disk fuel with varying enrichment over 15 cycles of irradiation. Circular markers indicate the average kernel temperatures at BOC and EOC for cycles 1 and 15. The solid color regions surrounding each marker indicate the minimum and maximum fuel temperatures for each data point. (For interpretation of the references to color in this figure legend, the reader is referred to the Web version of this article.)

5.4. Extension to higher temperatures Additional simulations are performed using naturally enriched 95% TD UO2 disk fuel to assess the potential for extending MiniFuel to higher temperatures. The target temperatures are 500e550, 700e800, 900 1,000, and 1100 1200  C. Table 5 summarizes calculated fuel temperatures as well as the target fill gas and gas gap required to achieve these temperatures. Table 5 shows that temperatures as high as 1150  C can be achieved using an argon fill gas. It should also be noted that the gas gaps are relatively large compared to those for traditional integral fuel tests conducted in dry capsules, which makes the MiniFuel experiment temperatures much less sensitive to machining tolerances. For example, previous irradiation testing of sodium fast reactor fuel at the Advanced Test Reactor used a nominal helium gas gap of 50 mm [71]. With typical machining tolerances on the order of tens of mm, the uncertainty in the fuel temperature could easily exceed 100  C in this case. The smallest MiniFuel gas gap of 284 mm is more than an order of magnitude greater than typical diametrical machining tolerances. A 284 mm gap with ±10 mm diametrical tolerances gives an uncertainty of ±3.5% on the gap size, or ±13  C on the temperature difference across the gap for the initial MiniFuel experiments. For testing at temperatures beyond 1150  C, the capsule material could be changed from a titanium alloy to a denser metal, such as molybdenum, which would provide increased gamma heating to drive higher temperatures. Preliminary calculations suggest that temperatures approaching 1400  C could be achieved with naturally enriched UO2 fuel using molybdenum capsules.

6. Discussion 6.1. Comparison with integral fuel experiments The MiniFuel experiments provide a separate effects fuel

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irradiation testing platform that is expected to complement existing integral fuel irradiation testing capabilities. The accelerated nature of the MiniFuel experiments is far from prototypic, but it allows for rapid burnup accumulation and quick screening of a wide range of fuel concepts. HFIR's compact size also limits the number and size of the reflector experiment positions. Therefore, large integral fuel testing campaigns that target prototypic conditions are best conducted in larger cores with increased experiment volume such as the Advanced Test Reactor at Idaho National Laboratory. Within the nuclear fuels community, there are varying ideas for how to best combine separate effects irradiation testing capabilities with integral fuel testing to accelerate the fuel qualification process. However, there is a general consensus that (1) separate effects testing must be leveraged in order to accelerate fuel qualification, and (2) some integral testing will remain necessary. One approach would integrate these testing capabilities with modern modelling and simulation tools to first identify parameters that have the highest impact on fuel performance and a relatively high uncertainty. This would allow for targeted separate effects tests to be conducted to improve fuel performance models. Rapidly iterating between engineering-scale models and separate effects tests could allow the community to gain confidence in fuel performance under both normal operation and design basis accident conditions prior to executing longer term, expensive, integral tests, which would ideally be used only to confirm the fuel performance models and identify any relevant phenomena that were not observed during the separate effects tests. Table 6 provides a quantitative comparison of the performance of MiniFuel conditions to those in a typical LWR integral fuel test. The parameters being compared are the enrichment, fission rates, the time to achieve various burnup levels, the maximum spatial fuel temperature variations, and the maximum temporal variations in peak fuel temperature. Fission rates (in W/g) and peak fission power (in W/specimen) are compared. Specimens are assumed be a ø3.2 mm  0.3 mm thick disk for MiniFuel and a single ø8.4 mm  12.6 mm tall pellet for the integral test. Actual specimen dimensions during integral fuel testing may vary significantly, but these estimates allow for an order of magnitude estimate of total fuel heat load. Table 6 shows that the MiniFuel experiments can provide burnup accumulation at a rate nearly 3  that of traditional integral fuel testing using natural uranium as opposed to low-enriched uranium. This is possible because of the very high neutron flux in the HFIR. The small fuel volume results in a total fission power (in W) that is orders of magnitude lower than that of integral fuel tests. As a result, the temporal variations in fuel temperature are much less sensitive to variations in fuel fission rate. MiniFuel temporal variations in peak fuel temperature are nearly an order of magnitude lower than those of the ATF-01 integral fuel test, although temporal variations can be lower in some integral experiments. The small size of the fuel also limits the spatial temperature gradients, despite the higher fission rates per unit mass. Spatial temperature gradients in MiniFuel are lower than the numbers estimated for a typical integral fuel test by a factor >20. The MiniFuel experiments rely on breeding and subsequent

Table 5 Summary of temperatures and experimental parameters required to achieve fuel temperatures in the range of 500e1200  C for 95% TD UO2 disk fuel with 0.73 wt%

235

U.

Design temperature (⁰C)

Fill gas

Gas gap (mm)

EOC 15 temperatures (Average [min max] in ⁰C)

500e550 700e800 900 1000 1100 1200

40.5% He, Ar balance 40.5% He, Ar balance Ar Ar

284 699 429 1429

529 [510e539] 763 [743e773] 943 [923e953] 1183 [1162 1193]

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Table 6 Performance comparison of MiniFuel vs. a typical LWR integral fuel test. Parameter

MiniFuel UO2 disk, 0.73 wt%235U, 95% TD

Typical integral fuel test

235

0.73% 152 3.8 97 369 29 73

5.44%a 59b 429c 247d 1,081e 667f 698g

U enrichment (wt%) Peak fission rate (W/g) Peak fission power (W/specimen) Effective full power days to achieve 15 MWd/kg U burnup Effective full power days to achieve 50 MWd/kg U burnup Maximum fuel temperature spatial variation up to 60 MWd/kg U burnup (⁰C) Maximum temporal variation in peak fuel temperature up to 60 MWd/kg U burnup (⁰C) a b c d e f g

Taken from Harp et al. [11]. Determined from slope of burnup vs. time for ATF-01 experiment [72]. Estimated from peak fission rate in W/g assuming a 95% TD UO2 pellet with dimensions of ø8.4 mm  12.6 mm height. ATF-01 capsule [72]. Extrapolated from first 300 days of irradiation of ATF-01 capsule [72]. Estimated from temperature contours for ATF-1 rodlet in [73]. Maximum variation over first 300 days of irradiation of ATF-01 capsule [72].

fission of 239Pu to allow for rapid burnup accumulation after the initial 235U content is depleted. The rapid depletion of 235U in the MiniFuel experiments and the shift to 239Pu fission has significant advantages. It is not necessary to fabricate enriched fuel samples for irradiations. In a conventional integral irradiation test, the fuel samples must be fabricated to a specific enrichment level that determines the fuel temperature and burnup rate. The design of the experiment is strongly linked to enrichment, so fuel fabrication efforts cannot begin until the experiment design has been finalized. This can result in delays of multiple years between the time a specific test is first envisioned and knowledge of the required enrichment so that fuel fabrication can begin. Another advantage of using lower enriched fuel in the MiniFuel vehicle is that fabrication activities involving enriched uranium inherently require additional administrative controls that greatly increase fabrication time and cost compared to the time and cost when working with depleted or natural uranium. The use of depleted or natural uranium could also allow for test fuels to be fabricated at universities or other facilities that have restrictions on handling enriched uranium. A subtle impact of the fact that 239Pu (and not 235U) fission dominates in the MiniFuel design is a slight shift in fission product yield curves. Fission yield from 239Pu presents some notable differences compared to that of 235U [74]. For example, ruthenium and iodine yields will increase for 239Pu compared to 235U, while zirconium, barium, and cerium yields will be reduced. The differences are not extreme, but they may become important when higher burnups are analyzed or when a focus is placed on fission product behavior.

6.2. MiniFuel post-irradiation examination A detailed summary of planned MiniFuel PIE activities is provided in a previous report [75], and a brief summary is provided here. Post-irradiation examination will first involve disassembly of the targets and extraction of the capsules. The capsules will be punctured and flushed with helium, and all xenon and krypton isotopes will be trapped using charcoal traps cooled with liquid nitrogen. These charcoal traps will then be gamma-counted using a high-purity germanium detector to determine the total 85Kr activity. Fuel FGR will be determined using the measured 85Kr activity, in combination with neutronics calculations of the total 85Kr produced during irradiation and confirmatory measurements of burnup. Burnup will be measured using destructive wet chemistry techniques by determining concentrations of stable, non-volatile lanthanides with well-known fission yields such as 146Nd [76]. Current analyses estimate that the minimum measurable 85Kr activity is 2.1  10 2 mCi. For comparison, a single ø3 mm  0.3 mm

thick UO2 disk with 95% TD will generate 83 mCi of 85Kr after accumulating a burnup of 10 MWD/kg U, assuming an average energy of 200 MeV per fission, an average 85Kr yield of 0.00286, and a85Kr half-life of 10.756 years. In this case, the minimum measurable fission gas release would be 0.025%, which is much smaller than the expected recoil release of 1.4% for this geometry [51]. The previous estimate used a typical 85Kr yield for 235U thermal fission, which is appropriate for fuels with higher initial 235U enrichment. For fuels with lower initial 235U enrichment, where fission is dominated by 239Pu thermal fission (85Kr yield of 0.00165), the minimum measurable release is 0.043%, which is still much smaller than the expected release. After puncturing, cutting operations will be performed to extract the fuel specimens and the passive SiC thermometry, which will be evaluated using continuous dilatometric techniques to determine the irradiation temperature [35]. Ultimately, the fuel samples will be carefully extracted for further examination. The goal is to collect accurate mass and volume data from the irradiated MiniFuel specimens. The volume data will be obtained via x-ray microtomography (XCT) techniques [75]. Finally, detailed examination of the fuel microstructure will be performed using modern microscopy techniques such as those detailed in Gerczak et al. [77].

6.3. Potential future applications The MiniFuel irradiation concept provides several key advantages as demonstrated above. First, the flexibility and simplicity of the capsule and irradiation design greatly reduces the cost of obtaining data on fuel performance. Second, the irradiation conditions allow for accumulation of burnup at an accelerated timescale compared to conditions in other facilities. Finally, the decoupling of temperature from fission rate allows for neutron irradiation of fuels under conditions not possible using conventional integral testing. The loss of test reactors worldwide has resulted in an increased demand on the remaining irradiation facilities. Irradiation capacity is therefore in high demand from the existing commercial nuclear industry, military, and space agencies with specialized needs that limit opportunities for other research. This increased demand often limits the opportunity for systematic assessment of the effect of lesser understood variables on fuel performance; it is impractical, for example, to simultaneously irradiate fuel specimens possessing varying ranges of processing impurities to study their possible impact on fuel swelling. Contemporary examples of this include the impact of carbon and/or oxygen content on the behavior of UN [78] and the impact of secondary uranium silicide phases on the performance of U3Si2 [11].

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Similarly, the availability of the MiniFuel facility can enable broader assessment of nontraditional fuel forms. The accident tolerant fuel research efforts of recent years have included several proposed concepts that deviate significantly from those of historic experience. Interest in maximizing thermal conductivity or improving other aspects of fuel performance have prompted investigation of high-density composite fuels such as UN-U3Si2 or UN-U3Si5 [79,80], assessment of uranium boride containing fuels [81], or UO2-based concepts that contain secondary phases [82]. Other approaches aim to deploy fuel architectures that were initially developed for other reactor applications to LWRs. One example is the fully ceramic microencapsulated fuel concept in which the advanced gas reactor TRISO particle fuel with a graphite matrix is modified to incorporate a high density kernel in a SiC matrix [83]. Regardless of the properties of these fuels as measured in their unirradiated states, these concepts may have critical vulnerabilities that may not be evident without irradiation. Use of MiniFuel to perform initial screening studies could help identify deleterious behaviors without requiring expensive and timeconsuming integral tests. The precise control afforded by the MiniFuel design also provides the opportunity to collect data on fundamental behaviors of nuclear fuels in the absence of a convoluting temperature gradient. This feature is well suited to separate-effects testing. For example, fission gas behavior is a critical fuel performance metric. Fission gas release has been historically measured by puncture and analysis of irradiated rodlets [84], which has been augmented in the more modern era by electron probe microanalysis [85]. Recent advancements in technology and sample preparation have extended capabilities to the sub-micrometer length scale [77], but even these modern methods are challenged to account for the temperature and burnup gradients present in an operating fuel. The MiniFuel approach provides the opportunity to measure fission gas retention for samples irradiated at a constant temperature and to pair this data with detailed microscopy data. Such data sets have the potential to revolutionize modeling of fission gas transport and release modeling as performed across all length scales, in addition to providing a rapid means to screen the effect of dopants or other modifications to UO2 microstructure on fission gas behavior. The effects of solid fission products can be similarly analyzed. In addition to providing microscopy samples to elucidate the structure and composition of fission product phases for lesser known fuel forms, even the behavior of oxide fuel fission products can be better understood using this approach. The oxygen potential of oxide fuel as a function of burnup is a major benchmark for thermochemical models, yet it has only received limited study due to the difficulty of measurement. Conventional approaches have relied on harvesting small particles of irradiated UO2 from the pellet radius in an attempt to limit the impact of temperature gradients [86]. The availability of fuel samples irradiated at a constant temperature would provide a far greater sample volume for this complicated measurement. The examples described above illustrate the utility of the MiniFuel concept using the existing design. Minor increases in complexity will offer additional capabilities. Introduction of thin cladding samples would provide a means to assess fuel-cladding interactions. Modification of the capsule design may also allow for investigation of metal coolant and/or bond effects (e.g., Pb, Na). Although increasing 235U content or hypothetically incorporating plutonium into as-fabricated MiniFuel samples will greatly increase fission rate early in life before reaching an equilibrium state, as illustrated in Fig. 6, the high thermal flux of HFIR will drive a rapid accumulation of burnup as initial fissile inventory is increased. It would be possible to reach burnups relevant to transmutation fuels or fast reactor driver fuels in a period of several years and would

13

allow for these families of nuclear fuels to be studied in a similar manner. Finally, the material behaviors relevant to lesser studied fuel cycles (e.g., thorium) can be investigated at a far lower cost than would be necessary at larger test reactors. 7. Conclusions This paper summarizes the MiniFuel separate effects irradiation testing capability recently established at ORNL through experiments performed in HFIR's permanent reflector. This new capability provides a flexible, cost-effective vehicle for obtaining fundamental data regarding the irradiation performance of a wide range of nuclear fuels. The key to this experiment design is the miniaturization of the fuel and HFIR's extremely high neutron flux. This combination allows for highly accelerated burnup accumulation using reduced 235U enrichment (even depleted or natural uranium) without prohibitively large temperature gradients within the fuel specimens. The MiniFuel vehicle also allows for a near decoupling of fuel temperature and fission rate. This is possible due to the total internal heat load being dominated by gamma heating in the structural components instead of fission in the fuel specimens. This paper describes the neutronic and thermal analysis for initial irradiation testing of UN kernels and UN-bearing TRISO particles with varying fuel density, carbon impurities, and burnable absorbers. Additional analyses are performed for other fuel temperatures, compositions, enrichments, and geometries to demonstrate the flexibility of the MiniFuel irradiation vehicle. The discussion compares the separate effects MiniFuel testing approach to traditional integral fuel testing and highlights some potential future applications. Data availability The raw and processed data from Figs. 5, Figure 6, Fig. 8, and Fig. 10 are available to download from https://doi.org/10.17632/ zhybxr6yxx.1#file-f00ed9a4-c208-4416-b196-54aa09eb3971. The raw/processed data required to reproduce these findings, including component geometries, time- and spatially-dependent HGRs for all materials, and material property data, as well as 3D nodal temperature results cannot be shared at this time due to technical or time limitations. Acknowledgements This work is supported by the US Department of Energy (DOE) Office of Nuclear Energy, Advanced Fuels Campaign. A portion of this research uses resources at the HFIR, a DOE Office of Science User Facility operated by ORNL. The paper is authored by UT-Battelle under Contract No. DE-AC05-00OR22725 with the US DOE. Robert Morris assisted in the development of the hot cell puncturing system to be used for measuring fission gas release and provided valuable insights and guidance. David Bryant assisted with capsule assembly. Joel McDuffee and Josh Peterson-Droogh performed reviews of the thermal and neutronic calculations supporting this work. Appendix A. Supplementary data Supplementary data to this article can be found online at https://doi.org/10.1016/j.jnucmat.2019.151783. References [1] R.C. Briant, A.M. Weinberg, Nucl. Sci. Eng. 2 (1957) 797. [2] D. Olander, J. Nucl. Mater. 389 (2009) 1.

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