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Nuclear Engineering and Design 254 (2013) 43– 52

Contents lists available at SciVerse ScienceDirect

Nuclear Engineering and Design

j ourna l ho me page: www.elsev ier .com/ locate /nucengdes

evelopment of the hot isostatic press manufacturing process for monolithicuclear fuel

ustin Crapps, Kester Clarke, Joel Katz, David J. Alexander, Beverly Aikin, Victor D. Vargas,oel D. Montalvo, David E. Dombrowski, Bogdan Mihaila ∗

aterials Science and Technology Division, Los Alamos National Laboratory, Los Alamos, NM 87545, United States

r t i c l e i n f o

rticle history:

eceived 15 November 2011ccepted 5 September 2012

. Introduction

The United States Department of Energy, Office of the Nationaluclear Security Administration established the Global Threateduction Initiative (GTRI) program in the Office of Defense Nuclearonproliferation to reduce and protect vulnerable nuclear and

adiological material located at civilian sites worldwide by provid-ng support for countries’ own national programs. An importantomponent of the GTRI important is the need to convert researcheactors from the use of highly -enriched uranium (HEU) to low-nriched uranium (LEU), with less that 20% uranium enrichment.hese efforts result in permanent threat reduction by minimizingnd, to the extent possible, eliminating the need for HEU in civilianpplications.

Dispersion fuels made from LEU are inadequate for many high-ower test reactors (Clark et al., 2006). To achieve the needed

oading for high-power test reactors, monolithic U–Mo alloy fuelsave been proposed (Clark et al., 2006), posing the challenge ofonding an aluminum cladding to the monolithic fuel foils by pro-esses other than the roll bonding traditionally used for dispersionuels (Clark et al., 2005, 2006).

Clark et al. (2003) proposed that the fuel foils should be pro-uced by a hot rolling process, followed by bonding the cladding tohe fuel foils. To bond the cladding, these authors investigated sev-ral manufacturing processes, such as: high temperature rolling,ransient liquid phase bonding, diffusion bonding via hot isostaticressing (HIP), and friction bonding. Of these processes, the HIPonding process produced the most consistent results (Robinsont al., 2009; Clark et al., 2005) and was chosen as the preferred man-

facturing method, even though HIPping requires that the material

s held under vacuum or reduced atmosphere to prevent the for-ation of an oxide layer on the material surface.

∗ Tel.: +1 505 664 0488; fax: +1 505 667 8021.E-mail address: bmihaila@lanl.gov (B. Mihaila).

029-5493/$ – see front matter © 2012 Elsevier B.V. All rights reserved.ttp://dx.doi.org/10.1016/j.nucengdes.2012.09.002

The HIP process was developed by Battelle Memorial Insti-tute for diffusion bonding of nuclear fuel element assem-blies (Bocanegra-Bernal., 2004; Yamada et al., 2004). The mech-anisms of HIP, residual stresses, HIP diagrams and densificationmaps, and sensing and modeling were studied extensively (Arztet al., 1983; Helle et al., 1985; Li et al., 1987, 1991; Swinkels et al.,1983; Wadley et al., 1991; McCoy, 1985; Nair and Tien, 1987). Anoverview of the fundamental aspects of the HIP process is found inthe review paper by Atkinson and Davies (2000).

Investigations of HIPping applicability for the LEU monolithicfuel plates fabrication focussed primarily on the chemical com-position of the U–Mo fuel foil and the diffusion barrier neededto separate the aluminum cladding and the U–Mo alloy fuel. Sev-eral Mo contents were considered between 8 and 10 wt%. HigherMo compositions showed minimal reaction layer formation atthe fuel-cladding interface, whereas a lower concentrations ofMo had the reverse effect (Clark et al., 2006). Irradiation testingof small fuel plates was performed by Robinson et al. (2009) tostudy fuel swelling, fuel-cladding interaction, blister formation, andfuel-cladding bond effectiveness. Ultimately, U–10Mo (wt%) wasselected as an acceptable compromise between uranium densityand phase stability benefits.

Initial research concerning the diffusion barrier needed to pre-vent the interaction between fuel and aluminum cladding reliedon a 25-�m-thick niobium foil introduced between the U–Mo foiland aluminum cladding. Unfortunately, experiments showed largevoid areas at the interface between the U–Mo and the niobium,despite the fact that bonding between the aluminum and niobiumshowed no apparent reaction layer. Jue et al. (2010) showed thata thin zirconium diffusion barrier between the fuel and the alu-minum cladding eliminated direct fuel-cladding interaction duringthe bonding process.

Production requirements for LEU monolithic fuel plates fabri-cation were discussed by Wachs et al. (2008). Burkes et al. (2008)considered possible characterization methods for the mechanicalperformance of the fuel plates. Recently, Ozaltun and Medvedev

4 eering and Design 254 (2013) 43– 52

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Table 1Room and high temperature tensile properties for 6061-T6 Al and 304 stainlesssteel (SS). Here, E denotes the elastic modulus, Sy is the yield strength, and Su is theultimate tensile strength.

Material Temp (◦C) E (GPa) Sy (MPa) Su (MPa) % el

304L SS RT 193 503 752 42

4 J. Crapps et al. / Nuclear Engin

2010) studied the structural behavior of the monolithic fuel platesuring HIP and annealing. They found that large residual stressesre induced by fabrication. These are mainly due to the differ-nce in the coefficient of thermal expansion in the U–Mo fuel foilnd the aluminum cladding. Similarly, Miller et al. (2010) studiedhe thermo-mechanical coupling in the fuel-cladding interface. Theffects of different delamination shapes were addressed. The pres-nce of a gap due to a delamination increases fuel temperature, buthe increase is relatively insensitive to the delamination size.

In this paper we develop a finite-element model (FEM) of theIP manufacturing process for the fabrication of LEU-10Mo nuclear

uel plates. We use simulations and experiments to characterize thetress distributions corresponding to the baseline HIP can designatz et al. (2011). We study the model sensitivity to material prop-rties and propose modifications of the original HIP can design tomprove the uniformity of the normal stress distribution relative tohe fuel-cladding interface.

The paper is organized as follows: in Section 2 we review the HIProcess and the geometry of the original HIP can design. In Section 3e introduce the FEM and discuss our simulation results for the

riginal HIP can design. In Section 4 we discuss the implications ofwo possible modifications in the geometry of the original HIP canesign. We discuss our experimental results in Section 5 and showhat these results are consistent with our simulations. We concluden Section 6.

. HIP process and geometry

The HIP process requires that the fuel and cladding be enclosedo keep the pressurized media (argon gas) from interfering with theonding process. The enclosure used is a 304 stainless steel can thatan be evacuated, removing gases that may interfere with the bond-ng process (Clark et al., 2006). Preparation of the can components

as discussed elsewhere (Jue et al., 2010; Moore et al., 2008).The HIP can is assembled from two cover plates, two side plates,

bottom plate, and a top plate. Once assembled, the plates areoined together via tungsten inert gas welding (TIG) to create a can

ith one cover plate unattached. Then, the HIP can is loaded withlternating layers of rectangular steel plates (strongbacks) and fuellates composed of U–10Mo foils and 6061-T6 aluminum cladding.he final cover plate is TIG welded in an argon-filled glovebox. Thessembly is evacuated of all internal gasses and sealed with a crimpeld.

The fuel foils consist of U–10Mo co-rolled with a thin layer ofirconium on each face. The fuel foil is encased in the aluminumladding during the HIP process. When inserted into the HIP can, theuel plate stackup consists of two flat sheets of aluminum and one–10Mo foil with dimensions (width and length) approximately

in. less than the aluminum. In order to reduce the relative move-ent of the plates before HIPping, the U–10Mo fuel foil fits into

pocket machined in one of the aluminum plates. During the HIProcess, the goal is to create aluminum–aluminum bonds aroundhe perimeter of the fuel foil and aluminum–zirconium bonds else-here. The strongbacks are intended to help minimize distortion

f the fuel foil and cladding. To prevent bonding of the strongbacksnd aluminum, we use a parting agent to separate the strongbacksrom the cladding and can. The preparation process for the U–10Mooils and cladding is described elsewhere (Moore et al., 2008, 2009).

The assembled HIP can is placed inside a HIP chamber and iseated to 320 ◦C. Next, the inside of the HIP chamber is pressur-

zed to 104.4 MPa (15 ksi) using argon gas, while the temperature

s raised to 560 ◦C (Jue et al., 2010; Ozaltun and Medvedev, 2010).he assembly is held at this temperature and pressure for 90 min,fter which the pressure is reduced to atmospheric and the can isooled for approximately two hours to room temperature.

304L SS 560 152 117 310 586061-T6 Al RT 70 206 476 206061-T6 Al 371 47 12.4 20.6 98

In Figs. 1 and 2 we show the HIP can and fuel plate originalgeometries (Katz et al., 2011). In Fig. 3 we show the fully-assembledHIP can before and after the HIP process, as well as the positioningof the HIP can inside the HIP chamber. We note that the depressedsides in Fig. 3(c) are an indication of a successful HIP-bonding cycle.

3. Finite-element modeling

We use ABAQUS CAE to create and execute the FEM for thisstudy. Here, we consider a two-dimensional model correspond-ing to a transverse cross section located at the center of the HIPcan (long axis of can is normal to the plane of analysis). Wemodel the HIP can, strongbacks, fuel foil, and cladding as individualdeformable components, assembled into the arrangement shownin Fig. 4.

To simplify the calculations, we used symmetry considerationsto justify modeling one quarter of the HIP can assembly andwe applied the appropriate symmetry constraints to the HIP canassembly about the horizontal and vertical axes. We apply isostaticpressure to the outside surfaces. A single node on each componenthad to be fixed at the beginning of the simulation to assure numer-ical stability while the HIP can deforms to initiate contact withthe strongbacks. All fixed nodes are located on the vertical sym-metry line. Once contact has been established, the fixed boundarycondition on the nodes was removed. Fig. 4 illustrates the initialboundary conditions.

The simulation loading is performed in multiple steps to aid incontact detection. The isostatic pressure is incremented by smallamounts until initial contact between the can and strongbacksis established. Afterward, the pressure is incremented by largeramounts to the desired pressure of 104.4 MPa (15 ksi).

The FEM involves two materials: Al 6061-T6 and 304L stain-less steel. Table 1 summarizes elastic and plastic properties for bothmaterials (ASD, 1997; Rothman, 1987). For simplicity, we used 304stainless steel properties as a surrogate for the Zr co-rolled U–10Mofuel foil. Because 6061 Al material data is not available at the processtemperature of 560 ◦C, we used 6061 Al data at the highest temper-ature available (371 ◦C). Because our study is intended to provide aqualitative understanding of the HIP process, these simplificationsdo not significantly affect our conclusions and this assumption wassupported by our experimental results.

For completeness, we note that the solidus temperature for Al6061 is at 580 ◦C and it is expected that the aluminum is at orbeyond an O-temper state after 90 min at 560 ◦C (Jue et al., 2010).The hardness of the U–10Mo has been shown not to change dur-ing the HIP process (Jue et al., 2010). In our simulations we used theABAQUS Von Mises plasticity model and large displacement theory.All simulations discussed here are static.

The HIP assembly was meshed initially such that we hadtwo elements through the fuel and cladding components. Subse-quently, we conducted a mesh refinement study considering anelastic–plastic material behavior in all components at room tem-

perature. Simulations were run with all elements in the analysis at1/2 and 1/4 the initial element size. The simulation results con-verged with a square element size of 0.032 mm (1.25×10−3 in.)in the aluminum cladding and U–10Mo and a square element

J. Crapps et al. / Nuclear Engineering and Design 254 (2013) 43– 52 45

inless

sAtmc

Fig. 1. Dimensions and assembly of 304L sta

ize of 0.206 mm (8.125×10−3 in.) in the can and strongbacks.t convergence, the mesh consists of eight elements through

he fuel and through the cladding. In Fig. 5 we illustrate ouresh refinement results in both the assembly and individual fuel

omponents.

Fig. 2. Dimensions and assembly of fuel plat

steel HIP can components (Katz et al., 2011).

The two-dimensional model of the HIP can is composed ofelements using a plane strain formulation. We use the ABAQUS ele-

ment type CPE4H, which is a plane strain with hybrid formulationinvolving an additional variable relating to constant pressure. Thehybrid formulation is recommended for improved convergence of

es and strongbacks (Katz et al., 2011).

46 J. Crapps et al. / Nuclear Engineering and Design 254 (2013) 43– 52

Fig. 3. Fully assembled and evacuated HIP can before the HIP cycle (a), within the HIP chamber after the HIP cycle showing welded base and evacuation tube (top) (b) andafter the HIP cycle (c), showing depressed sides indicating a successful HIP bonding cycle.

Fig. 4. Assembled HIP can model and boundary conditions.

Fig. 5. In this mesh refinement study, we show that the S22 contours associated with the stress normal to the bonding line between U–10Mo and Al do not change appreciablybetween the medium and fine mesh simulations. This is indicative of a converged solution.

J. Crapps et al. / Nuclear Engineering

Fig. 6. Simulation speedup as the number of processors is increased. The simulationctf

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onsidered elastic and plastic deformation in both the aluminum and steel at roomemperature. We find that the optimal computational efficiency is achieved usingour processors.

imulations where the material is likely to extrude, but remainsonstrained.

A FEM considering elastic-plastic deformation at room tempera-ure and consisting of approximately 80,000 elements takes about0 min wallclock time for solution using a single core of an Inteleon X5570 processor at 2.93 GHz. The computational time can beeduced by executing ABAQUS on parallel processors. Fig. 6 illus-rates the speed improvement vs number of processors relationor this model. The parallel scalability of this ABAQUS model isublinear, and we compromised to run these simulations on fourrocessors.

.1. Ideal case

The purpose of the HIP can assembly is to provide anvacuated container for application of normal stress along theuel-cladding and cladding-cladding interfaces, creating a reliablyonded aluminum-clad U–10Mo fuel plate while allowing easyeparation of fuel plates from the strongbacks and HIP can. Thedeal situation would be to process the fuel and cladding with-ut a stainless steel can. Simulations were performed of this idealase considering plasticity at high temperature in both the steelU–10Mo surrogate) and the 6061 Al cladding. Fig. 7 shows the can-ess HIP simulation S22 (stress normal to the bonding line betweenhe fuel foil and Al) stress contours. The stresses throughout the fuelnd cladding are very uniform and are the best possible conditionsor creating a cladding–cladding and fuel-cladding bond.

.2. Model sensitivity to material properties

Several cases involving the elastic or elasto-plastic models forhe aluminum and steel were explored to investigate the impor-ance of plasticity modeling at room and high temperature. Figs. 8nd 9 illustrate the S22 contours and displacements for simulationsonsidering only the elastic response, plasticity in the aluminum,lasticity in the steel, and plasticity in the aluminum and steel foroth room temperature (20 ◦C) and process temperature (560 ◦C).hese simulations were performed in the low-friction limit.

Fig. 8 shows that modeling the steel plasticity in the HIP can atoom temperature is important, while modeling plasticity in theluminum cladding has a small effect. As the isostatic pressure isncreased to 104.4 MPa, the steel in the corners of the HIP can yields

ue to bending stress and affects the stresses along the bonding linef the U–10Mo and the aluminum cladding. At room temperature,he aluminum and fuel do not yield and remain in the elastic rangefter the full isostatic pressure of 104.4 MPa was applied.

and Design 254 (2013) 43– 52 47

At elevated temperature, the yield strength (Sy), ultimate ten-sile strength (Su), and elastic modulus (E) decrease in both thealuminum (cladding) and the steel (can, strongbacks, and fuel sur-rogate). Fig. 9 shows that modeling plasticity in both the aluminumand steel at elevated temperature (560 ◦C) has a strong effect onthe stresses normal to the bond line. For obtaining reliable resultsin this model, we need to consider the elastic and plastic materialproperties in all components close to the processing temperature.

3.3. Friction effects

The effect of friction was explored using a simple Coulomb fric-tion model. Friction was modeled for every contacting interface.Simulations were performed for the baseline geometry using fric-tion coefficients of 0.1, 0.3, and 0.6 and high temperature materialproperties. Fig. 10 displays the calculated normal stress (S22) con-tours and displacements. The simulation results show that frictionhas a weak effect on the HIP process. Thus, controlling friction isnot expected to influence significantly the uniformity of the normalstress distribution in the U–10Mo and Al bonding region. However,a low friction coefficient (Fig. 10(a)) is consistent with the experi-mental results, which indicate that significant aluminum extrusionoccurs into the gaps between strongbacks in the case when the canis not completely filled.

4. Modified HIP can design

Based on the simulation results discussed above, we concludethat the original HIP can design leads to three areas of concern inthe HIP fuel manufacturing process:

1 Considerable aluminum extrusion into the gap between thestrongbacks and the can will cause insufficient cladding coverageof the fuel plate.

2 The uneven distribution of normal stresses in the fuel-claddingbonding region will lead to uneven deformation (waviness) ofthe fuel plate.

3 The uneven distribution of normal stresses along the aluminumcladding interface is expected to give rise to uneven bonding offuel to cladding.

These numerical findings are confirmed by results of results ofexperimental trials discussed in Section 5.

The FEM allows us to explore design changes in the HIP cangeometry in order to improve the uniformity of the normal stressesalong the fuel-cladding interface. This sensitivity study can reducethe time and cost that would be required to optimize the geometryexperimentally via a trial-and-error approach.

In Fig. 11 we illustrate the normal-stress distributions andexpected displacements corresponding to two proposed modifi-cations in the HIP can geometry. These changes in the originalconfiguration are intended to reduce the aluminum extrusion andimprove the uniformity of the normal stress distribution along thefuel-cladding interface. For comparison, in Fig. 11(a) we show theresults corresponding to the original HIP can design. In Fig. 11(b),the HIP can geometry was modified by extending the length of thealuminum cladding to fill the space between the strongbacks andHIP can. In Fig. 11(c), we extended the width of the can, strongbacks,and aluminum cladding to an internal width of 5 in. In all cases, thedimensions of the U–10Mo foil are left unchanged. For clarity, inFig. 12 we illustrate the normal stress distributions and expected

displacements in the actual fuel elements corresponding to each ofthe designs shown in Fig. 11. We conclude that the extrusion of thealuminum cladding into the gap between the strongbacks and cancan be significantly reduced by extending the aluminum cladding

48 J. Crapps et al. / Nuclear Engineering and Design 254 (2013) 43– 52

Fig. 7. Ideal case for fuel plate manufacturing without a HIP can. In this case, stress contours in the fuel and cladding are very uniform.

Fig. 8. Stress contours and expected displacemets (arrows) for baseline simulations at room temperature considering (a) fully elastic materials, (b) aluminum plasticity, (c)steel plasticity, and (d) steel and aluminum plasticity. Considering plasticity in the aluminum cladding does not influence the normal stress (S22) distributions.

Fig. 9. Stress contours and expected displacemets (arrows) for baseline simulation at high temperature considering (a) fully elastic materials, (b) aluminum plasticity, (c)steel plasticity, and (d) steel and aluminum plasticity. Considering plasticity in both the steel and the aluminum cladding has a strong effect on the normal stress (S22)distributions.

J. Crapps et al. / Nuclear Engineering and Design 254 (2013) 43– 52 49

Fig. 10. Normal stress contours and expected displacemets (arrows) for simulations considering friction coefficients of 0.1, 0.3, and 0.6. The friction coefficient has a weakinfluence on the HIP process.

Fig. 11. Normal stress contours and expected displacemets (arrows) for two HIP can design modifications, compared with the original design (a). First, the aluminum claddingwas extended to fill the original HIP can to aid in preventing aluminum extrusion (b). Secondly, the HIP can and cladding were both extended to improve the distribution ofnormal stress along the fuel-cladding interface (c).

Fig. 12. Fuel plate normal stress (S22) contours and expected displacements (arrows) for the two HIP can design modifications (b) and (c) shown in Fig. 11, compared withthe original design (a). Simulations indicate that the extrusion of the aluminum cladding can be reduced by extending the aluminum cladding to fill the HIP can.

50 J. Crapps et al. / Nuclear Engineering and Design 254 (2013) 43– 52

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ig. 13. Base and tube ends of plate stackups (strongbacks and fuel plates) are showtrongback, showing uneven distribution of fuel plates from side to side, and fromladding and strongbacks have the same width, showing flat fuel plates filling the s

o fill the can. Fig. 11(b) and (c) shows that extending the aluminumo fill the gap between the strongbacks significantly improves theniformity of the normal stress along the fuel-cladding interface.his in turn translates into a flat clad-fuel product. Extending theidth of the can as shown in Fig. 11(c) offers only a limited improve-ent in the uniformity of the normal stress distribution along the

uel-cladding interface.

. Experiments

Several HIP cans of the design described in Figs. 1 and 2 wererocessed following the procedure discussed in Section 2 to deter-ine the effects of strongback and fuel configuration within the

an. Even though at LANL we tested extensively manufacturing

ig. 14. Transverse metallographic cross-sections of clad fuel after the HIP process in thladding (b) and (d) show the corresponding ends of the fuel plate inside the cladding. Inn this case the aluminum cladding did not fill the gaps between the strongbacks and the

ith even cladding coverage and minimal extrusion. In this case the aluminum clad was

removed from the HIP cans for: (test 1) – aluminum cladding width is less than theo end; the fuel plates slide down toward the base of the can; (test 2) – aluminum

between strongbacks. Shifting is eliminated.

process variables such as the material and shape of the strong-backs, the width of the fuel plates, and the parting agent used for aneasy removal of the HIPped fuel plates, in this paper we only focuson the experiments performed using aluminum cladding widths of64.2 mm and 76.2 mm.

Results of experimental trials reinforce the three major effectsdiscussed in Section 4. Experiments performed using the origi-nal HIP can design characterized by fuel plates with width andlength dimensions significantly less than the strongback dimen-sions showed all three effects inferred from our finite-element

simulations: significant lateral aluminum extrusion, uneven distri-bution of aluminum cladding around the fuel, and uneven bondingof the cladding to the fuel. In addition, experiments demonstratedsignificant plastic deformation in the strongbacks, which precludes

e two tests depicted in Fig. 13. Panels (a) and (c) show the edge of the aluminum (test 1) metallography shows the aluminum extrusion at the cladding edge (left).

can and cladding thickness is uneven. In (test 2) metallography shows a fuel plateextended to fill the can.

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J. Crapps et al. / Nuclear Engin

eusing the strongbacks in subsequent HIP cycles. The aluminumladding extrusion and the plastic deformation of the strongbackseinforce the importance of modeling the HIP process within anlasto-plastic framework. Furthermore, friction is important for theeliability of the FEM predictions, but the direct comparison of sim-lation and experimental results favors a small value of the frictionoefficient.

In Fig. 13 we show the base and tube ends (i.e. the bottom andop extremities of the HIP cans depicted in Fig. 3) for two HIP trials.ere, (test 1) corresponds to a trial performed using the original HIPan design, whereas (test 2) corresponds to a modified geometryith the aluminum cladding being extended to fill the gap between

he strong backs and the can. The latter geometry corresponds tohe simulations depicted in Figs. 11(b) and 12(b).

Increasing the size of the fuel plate relative to the size of thetrongbacks dramatically improves the overall uniformity of theompression of the inserted stackup. We show in Fig. 13(b) the as-emoved base end of a HIP can stackup (five strongbacks and fouruel plates) with fuel plates that were 0.5 in. (12.7 mm) narrowernd 1 in. (25.4 mm) shorter that the strongbacks. In Fig. 13(d), thes-removed base end of a HIP can stackup (four strongbacks andhree fuel plates) with fuel plates the same width and length as thetrongbacks are shown. The strongbacks depicted in Fig. 13(a) showignificant deformation. In the (test 1) experiment, the fuel plateshifted both vertically and horizontally in the HIP can. The extrapace in the HIP can allows for the deformation of the strongbacksnd movement of the fuel plates after assembly. In addition, signif-cant extrusion of the aluminum cladding changes nonuniformlyhe overall plate thickness. The latter results in final machiningifficulties. As shown in Fig. 13(c), filling the HIP can by extend-

ng the aluminum cladding to match the strongback’s dimensionseduces the uneven deformation of the strongbacks. Filling the canoes not eliminate entirely the aluminum extrusion between thetrongbacks and the can, but the volume of the extruded materials negligible.

The uneven cladding and relative aluminum shifting in theriginal HIP can design is captured via transverse metallographicross-sections and is illustrated in Fig. 14. Here, Fig. 14(a) and (c)hows the edges of the as-removed fuel plates, corresponding to theube ends in the two tests depicted in Fig. 13(b) and (d). Fig. 14(b)nd (d) shows the transition region between aluminum-aluminumonding at the edges and aluminum–zirconium bonding in the fuellate regions at the same ends. When the HIP can is not filled,he final dimensions (overall and cladding thicknesses) of the fuellate are not uniform. In contrast, when the aluminum cladding

s extended to fill the gaps, dimensional accuracy is dramaticallymproved and metallography shows that the aluminum-aluminumnd the fuel-cladding bonds are consistently sound.

. Conclusions

Finite element simulation and experimentation were used tonderstand and improve the HIP manufacturing process for mono-

ithic fuel plates. As expected, we find that the simulation results areensitive to the temperature dependence of the material properties.or the steady-state simulations considered here, it is importanthat we use material properties close to the relevant process tem-erature. Furthermore, plasticity must be taken into account inrder to obtain accurate predictions. Our simulation results for theIP process are largely insensitive to the value of the friction coef-cient, but a small friction coefficient is required to assure good

ualitative agreement with experiment.

Most importantly, our simulations indicate that extending theluminum cladding to fill the space between the strongbacks pre-ents the aluminum cladding extrusion into the space between

and Design 254 (2013) 43– 52 51

strongbacks and allows for a more uniform distribution of normalstresses along the fuel-cladding interface. We verified experimen-tally that this change in geometry leads to flat clad-fuel products.Furthermore, the plastic deformation of the steel strongbacks isdramatically reduced when the aluminum cladding is extended tofill the HIP can, and these strongbacks can be reused in subsequentHIP runs.

Acknowledgement

Work performed in part under the auspices of the U.S.Department of Energy, Office of the National Nuclear SecurityAdministration, under the Global Threat Reduction Initiative Reac-tor Convert program.

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