penetration resistance of cast metal−ceramic composite

Penetration Resistance of Cast Metal�Ceramic CompositeLattice Structures

Manuel E. Umanzor,* Romesh C. Batra, Christopher B. Williams,and Alan P. Druschitz

1. Introduction

Cellular structures, comprising a network of interconnectedtrusses and plates, are advantageous because of their favorablehigh specific strength and stiffness and superior thermal

properties over monolithic material struc-tures. They are increasingly being usedas thermal insulators, packaging materials,and bumpers to resist impact loads andabsorb energy as trusses buckle, yield,crush, and fracture.[1] Stochastic foammaterials can be manufactured by severaltechniques such as either direct foamingof a melt or indirect foaming via a precur-sor.[2] However, due to the lack of techni-ques to control parameters, repeatabilityof their fabrication is challenging.[3]

Ordered cellular arrangements are gen-erated by periodically repeating a unit celland hence can be reliably manufactured.Inspiration to design such periodic struc-tures has been drawn from various sources.Yang et al.[4] used 3D printing to producehierarchical structures based on naturalnacre and assess their fracture toughness.Designing at the atomic scale by depositingan atomic layer each time is an alternativeto periodic structures; the tetrakaidecahe-dron or Kelvin’s cell presented by LordThomson is an old example of suchstructures.[5] Originally, Kelvin’s cell was

intended to solve the problem of dividing space with the mini-mum partitional area. However, designers can use computer-aided design (CAD) tools to generate truss arrangements thatresemble the Wigner�Seitz form of the body-centered cubicstructure. Additional examples are included in Ashcroft andMermin.[6] In a counterexample to Kelvin’s conjecture, Weaireand Phelan generated a unit cell inspired by the atomic positionsof the β-tungsten cell (space group Pm3n). Using moderncomputational tools, they concluded that the Weaire�Phelanstructure surpasses the Kelvin tetrakaidecahedron by �1% whenthe figure of merit is the isoperimetric quotient.[7] Zhao et al.used the Menger sponge fractal to algorithm to manufacturesuperelastic porous solids via selective laser melting.[8] One coulduse topology optimization techniques to design a unit cell forgiven loading conditions and constraints and create a structuretailored for the maximum specific energy absorption due to plas-tic dissipation. An outstanding example is the octet-truss latticewith a face-centered cubic (FCC) nodal arrangement that exhibitssimilar behaviors under tensile and compressive loading.[9] Asshown in Figure 1a, the unit cell comprises an octahedral arraycontained in a tetrahedral cell with both sharing the same adjoin-ing nodes.

M. E. Umanzor, A. P. DruschitzDepartment of Materials Science and EngineeringVirginia Polytechnic Institute and State UniversityBlacksburg, VA 24061, USAE-mail: [email protected]

R. C. BatraDepartment of Biomedical Engineering and MechanicsVirginia Polytechnic Institute and State UniversityBlacksburg, VA 24061, USA

C. B. WilliamsDepartment of Mechanical EngineeringVirginia Polytechnic Institute and State UniversityBlacksburg, VA 24061, USA

The ORCID identification number(s) for the author(s) of this articlecan be found under https://doi.org/10.1002/adem.202100577.

DOI: 10.1002/adem.202100577

A challenging issue in armor mechanics is to optimize the impact resistanceof a target per unit areal mass. Herein, the penetration resistance of anA356 alloy–ceramic lattice structure with ceramic tiles encapsulated in themetal matrix is experimentally and computationally studied to achieve thisobjective. A hybrid additive manufacturing/metal-casting technique is usedto fabricate the structure. The performance is experimentally evaluated byimpacting the tiles at normal incidence with 0.30-cal armor-piercing pro-jectiles and 7.62 M80 rounds traveling at high speed. X-ray imaging andelectron microscopy techniques are used to ascertain the quality of thecastings and the damage caused to the target by the projectiles. The castmaterial is tested following ASTM standard E8/E8M to determine the yieldstress and hardening coefficients in the Johnson–Cook model. Largedeformations of the components are analyzed using the finite elementsoftware LS-DYNA. The penetrator’s computed residual velocities differ fromtheir test values by less than 4% for the 0.30-cal projectiles and 19% for theball rounds. The effects of several design variables (e.g., tile thickness andlocation, among others) are numerically scrutinized. The proposed design is�20% lighter than the solid metal target to achieve the same residual velocityof the penetrator.

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Conventional manufacturing processes are not well suited forthese geometrically complex structures. However, additivemanufacturing (AM) is a viable option as it eliminates the needfor additional tooling other than a printer and allows for completefreedom during the design process.[10] These advantages comewith their own unique set of constraints, particularly for metalsmanufactured via fusion processes because only a few alloyscan currently be used for AM methods. These limitations arefurther affected by the desired microstructure, preferredorientation, absence of defects (lack of fusion, cracking, and lossof low-melting-point alloying elements), high residual stresses,anisotropic mechanical properties, and the inherent perilsof handling fine metallic powders.[10–12] Recently, anAM�metal-casting hybrid process was used to additively manu-facture sand molds for casting aluminum cellular materials thathave mechanical properties, microstructures, and imperfectionsnearly identical to those of pure metal-cast materials.[3,13]

Lattice structures have a wide range of applications that includeclose-range blast mitigation[14] and high-velocity impact resis-tance.[15,16] While these studies have shown advantages of cellularstructures, they have focused on monolithic materials and particu-late-reinforced metal�matrix ceramic (MMC) composites.[17]

However, low-strength metal matrices like aluminum alloys maynot offer sufficient resistance to penetration fromprojectiles designed to pierce hardenedmaterials.Wilkins has sum-marized characteristics needed for a target to resist penetration(e.g., strength, mass density, and thickness), as well as the condi-tions for a projectile to penetrate a target. In addition to high velocity(which translates into high kinetic energy [KE]), a pointed projectileis potentially more efficient than a cylindrical penetrator, as its KE isapplied over a small cross-sectional area.[18] In such a scenario thetarget must break the projectile’s tip as soon as possible to defeat it.

Previous research on cast lattice structures using the AMmolds has helped identify materials for the metal matrix andthe ceramic tiles,[19] limitations, and constraints inherent to thisprocess (feature size, feeding distance, and channel depowder-ing).[20] We first upscaled structures’ dimensions in this workand encapsulated tiles of varying thicknesses to find a good com-promise between weight and performance. The castmetal�ceramiccomposites were tested against high-velocity impact/penetration,

and their nonlinear deformations were analyzed by the finite-element method (FEM). After comparing computed results withthe test findings, numerical tools are used to enhance our under-standing of the impact performance of the cast metal�ceramiccomposite lattice structures. We also numerically explore differentscenarios such as the tile location in the lattice structure, the obliq-uity of impact, and the projectile impacting on multiple ceramictiles through its trajectory.

2. Experimental Section

2.1. Unit Cell Design

As shown in Figure 1b, the unit cubic cell of side length 25 and4mm-diameter trusses was designed to have flat boundaryplanes that facilitated the generation of the complete latticestructures.

2.2. Specimen Preparation

The 225� 225� 50mm section of the lattice structure was gen-erated by creating a pattern of 9� 9� 2 unit cells along the threemutually perpendicular axes. The core mold was generated bysubtracting the lattice structure from this solid box, and afive-piece sleeve was designed to cover only the top and the bot-tom faces of the test specimens with 6mm-thick metallic plates.Provisions to add six 4 , 6 , or 8 mm-thick ceramic tiles were cre-ated on the upper face of the core mold. These tiles werearranged in pairs and fit into cavities such that each tile wasencapsulated by 4mm of cast metal on all sides.

The components of the core mold were additively manufac-tured in furan-bonded silica sand using an ExOne S-Max sand3D printer. A continuous sand mixer, model M50XLD fromPalmer, was used to make the remaining parts of the mold usingtraditional patterns. Figure 2 shows steps followed in themold-assembly process and casting and an overview of the tilelocations on the top surface.

Two aluminum alloy A356 lattice structures were cast with sil-icon carbide tiles. The alloy was prepared in an electrical resis-tance furnace at 760 �C. Prior to pouring, grain refiner (Al5Ti1B,by mass %) and silicon modifier (Al10Sr, by mass %) were addedto the molten metal. Chemical composition, shown in Table 1,was determined using a Bruker Q4 Tasman advanced charge-coupled device (CCD)-based optical emission spectrometer.Before testing, the specimens were solution treated at 540 �Cfor 11 h in an air-circulating furnace. The castings were agedfor 24 h at room temperature followed by 3.5 h at 155 �C to pro-duce the T6 temper.

2.3. Mechanical Testing

Two Y-blocks were also cast and heat treated as described earlier.Following the ASTM standard E8/E8M-16a,[21] test specimenswere machined and tested in simple tension. A typical true stressversus true strain curve is included in Supporting Information,from which the following values of material properties wereextracted: offset yield strength, σyð0.2%Þ ¼ 165MPa; modulus ofelasticity, E ¼ 73GPa; and percent elongation, EL% ¼ 3.36.

Figure 1. (a) Representation of an octet-truss unit cell, and (b) normalview of the unit cell with flat bounding faces. All dimensions are inmm. Dark-grey (light-grey) trusses describe the octahedral site (tetrahe-dral cell). Figure 1a from[4] is reproduced with permission copyright 2001,Elsevier Science Ltd.

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2.4. Penetration Tests

Penetration tests with 0.30-cal caliber armor piercing (AP) M2and 7.62� 51mm M80 full metal jacket (FMJ) rounds from astand-off distance of 7.62m at 0� obliquity and ambient temper-ature were conducted by the National Technical Systems. Theprojectile striking velocity (vs) was determined using infraredchronographs using model 57 from Oehler Research withY.I.S. Cowden Group. An additional set of paper screens wasplaced behind the target to measure striking (vs) and residualvelocities (vr). Projectiles were fired from a universal receiverfit with an appropriate barrel. A graphical description of the testsetup is provided in the Supporting Information.

A total of eight 0.30-cal AP M2 shots at different striking veloc-ities were fired at the first target, such that two sequential roundsstruck tiles of 0 (i.e., no tiles present), 4, 6, and 8mm thickness. Asimilar firing scheme with seven FMJ rounds was used for thesecond target. These tests and the main observations are shownin Table 2 and Table 3.

Side-by-side images of the strike and the back faces of the lat-tice structures after testing are shown in Figure 3. The back faceshould be interpreted as a mirror image of the strike face, that is,shot 1 is at the top right on the front face and at the top left on theback surface. It can clearly be seen that the AP shots 2, 3, and 7

and the FMJ shots 2 and 3 did not penetrate through the targetthickness. The ball round shots 6 and 7 pushed the ceramic tilesout of the target before breaking them catastrophically. It wasspeculated that the severely deformed lead core covered a largearea upon impacting the tile, and the KE was high enough to teara significant area of the metal matrix.

The Bombardier Recreational Products conducted X-ray imag-ing. As shown in Figure 4, and as intended by design, from shot2, the AP core fractured from its tip and some jacket fragmentsremained trapped in the lattice structure.

Sectioning of the test specimens revealed several pieces of theAP rounds similar to those shown in Figure 4b in the lattice

Figure 2. Sequence from CAD of the lattice to final casting. The core top cover has been removed to show provisions for tiles.

Table 1. Chemical composition of the final castings (mass %).

Si Fe Cu Mn Mg Zn Ti Al

Castings 6.83 0.09 0.02 0.004 0.35 0.003 0.14 Bal

Specification[36] 6.5�7.5 ≤0.2 ≤0.2 ≤0.1 0.25�0.45 ≤0.1 ≤0.2 Bal

Table 2. Summary of results of eight 0.30-cal AP M2 0� obliquitypenetration tests.

Shot no. SiC tilethickness

Penetration data

vs [m s�1] vr [m s�1] Result

1 4 mm 744 377 Bullet/Spall

2 6 mm 746 NA NP

3 8mm 750 NA NP

4 No tile 742 568 Bullet/Spall



7 8 mm 814 NA NP

8 None 817 664 Bullet/Spall

a)N/P: no through-the-thickness penetration. b)Bullet/spall: fragmentation of thetarget due to bullet impact.





structure. Furthermore, the SiC tiles were fragmented into dif-ferent size pieces, whose details are included in SupportingInformation.

The projectile impact on the target resulted in several failuremodes shown in Figure 5, petaling of the aluminum lattice uponperforation of the striking face, fracturing of the tile, failure of themetal surrounding it, and plugging of the back face in fully pen-etrated targets. For some tests, the projectile KE was not suffi-cient to penetrate through the target thickness. Fragments ofthe trusses, the tiles, the lattice, and the projectile recovered dur-ing a test were analyzed using the FEI Quanta 600 FEG scanningelectron microscope (SEM). As anticipated, the ceramic tiles hadbrittle failure. A close inspection of the fracture surfaces showedWallner lines that generally result from the interaction of a prop-agating crack front with a shock wave.[22] As well known, the

projectile impacts on targets generate compressive waves in bothstructures, and the interactions between the incident and thewaves reflected from bounding surfaces produce high stresses[18]

that create conditions favorable for the formation of Weller lines.Using energy-dispersive X-ray spectroscopy (EDS) on the SEMimages, several particles were identified on the fractured surfa-ces. Lead was detected on the SiC shards and the aluminumtrusses. Aluminum particles were also found on the broken sur-faces of the projectile’s hardened steel core. The cross-sectionalimages of the fractured trusses showed shrink porosity thatwas enhanced at the nodes due to the geometrical arrangementused.

A visual assessment of the penetration test revealed that thealuminum alloy on the striking face offered little resistance to theprojectile movement, and the projectile’s tip broke upon interact-ing with the ceramic tile. The tile’s shards aided the penetrator infracturing the metal beneath the tile.

Unfortunately, we could not retrieve projectiles exiting the tar-gets. However, the visual evidence suggests that the projectile’strajectory did not change significantly from one test to the otherand the smaller diameter of the exit holes implies that some com-ponents of the 0.30-cal AP rounds remained trapped in the latticestructure.

3. Analysis of Deformations by the FEM

3.1. Modeling Details

The highly nonlinear deformations occurring during theimpact were analyzed with the commercial FE software LS-DYNA to delineate the performance of the lattice structures.A schematic representation of the 0.30-cal AP projectile isshown in Figure 6a. The projectile length and the nominalbasal diameter, respectively, equal 35.6 and 7.84 mm. The coremorphology is taken from technical reports[23,24] and in theabsence of precise dimensions of the USA FMJ rounds, thoseof the NATO 7.61� 51 mm BR6 rounds were used (28.6 mmoverall length and 7.83 mm maximum diameter).[25]

Differences between the two projectiles are believed to besmall (see Figure 6b).

Due to the periodic nature of the test specimens, deformationsof the region comprising two unit cells along each principaldirection were studied. Furthermore, the problem size wasreduced to one-half by applying symmetry boundary conditionson the midplane (see Figure 6c for details).

The target without the tiles was discretized into hexahedralFEs and that with the metal encapsulated around the tiles intotetrahedral FEs (ELEFORM¼ 13). The initial FE mesh had ele-ment sizes of 1 mm for the lattice structure and the ceramic tiles,0.5mm for the projectile core, and 0.25mm for the remainingcomponents of the projectile. The FE mesh was successivelyrefined to obtain converged results.

The automatic surface to surface for projectile-to-projectilecontacts and eroding surface to surface for projectile-to-targetinteractions were used. The lattice to lattice (or lattice self-contact) was defined using the automatic-single-surface optionand setting the stated part as the slave segment. Previous studieshave shown that the hybrid AM�metal-casting approach does

Table 3. Summary of results of seven 7.62� 51mmM80 FMJ 0� obliquitypenetration tests.

Shot no. SiC tile Penetration Data

vs [m s�1] vr [m s�1] Result


2 6 mm 863 NP NP

3 8mm 856 NP NP

4 None 865 599 Bullet/Spall

5a) 4 mm 867 294 Bullet/Spall

6a) 6 mm 764 426 Bullet/Spall

7b) 8 mm 753 – Bullet/Spall

a)Tile located on the back face. b)No residual velocity captured.

Figure 3. Post-test photographs of the front and the back faces. The 7.62M80 FMJ shots 6 and 7 were fired on the back bare face that had noceramic tiles.





not create an interface between the cast metal and the ceramictile.[20] Thus, the lattice-to-tile contact was defined as automaticsurface to surface. An input file for one of the problems studiedis included in Supporting Information.

Preliminary simulations showed that the default penalty(SOFT¼ 0) and the soft constraint formulations (SOFT¼ 1)produced unrealistic contact surfaces. For example, seeFigure 7, in which the core failed to interact with the target,while using the penalty formulation. Consequently, all contact

surfaces were simulated using the segment-based algorithm(SOFT¼ 2).

An ogive-nosed projectile contacts the target throughout the pen-etration event.[26] In the problem being investigated, the contact sur-faces changed along the projectile trajectory. Due to severalcombinations of contact surfaces, using a relative velocity-dependent coefficient of friction for them is very arduous. We used0.2 as the coefficient of dynamic and static friction for all contactingsurfaces.

Figure 4. X-ray images: a) before and b) after testing. The image (b) shows a fractured core and a jacket embedded in the lattice structure.

Figure 5. Failure modes observed on cross sections of the lattice structure and on parts of the projectile. The lattice section on the left (right) is for thecase of incomplete (complete) penetration.





3.2. Material Models

The stress�strain results for the cast aluminum alloy mentionedin subsection 2.3 were used to calculate coefficients A and B andthe strain-hardening exponent n in the Johnson�Cook constitu-tive relation (1).[27] We neglected strain rate and temperaturedependence as aluminum alloys exhibit low strain-rate sensitivityand we do not have facilities for testing specimens at differentstrain rates and temperatures; the software does not considereffects of heat conduction.

σy ¼ ðAþ Bε̃pnÞð1þ C ln ε

: �Þ (1)

Constitutive relations for materials of different componentsare shown in Table 4. Values of parameters A, B, C, and n inEquation (1) for the aluminum alloy and gilding copper areshown in Table 5, and mechanical properties of the cores ofthe 0.30-cal AP and the FMJ projectiles are shown in Table 6.

As shown in Table 4, the SiC ceramic was modeled using theJohnson�Holmquist (JH-2) constitutive relation[28] with valuesof material parameters taken from Cronin et al.[29] and shownin Table 7.

3.3. Comparison of Computed and Test Results

Simulations were conducted using the commercial FE softwareLS-DYNA R11.1.0 shared memory parallel (SMP) double preci-sion executed on an Intel/Haswell Xeon E-2680v high-perfor-mance-computing (HPC) platform with 24 cores per noderunning on Linux Redhat 6.5 Ulm. Prior to carrying out theFE mesh convergence index (GCI) calculations, we determinedthe optimum number of processors to use with the multiple-coreoption built in LS-DYNA. In the region without the tile shown inFigure 6c, initial failure strains (FS) were set, respectively, as0.125, 0.06, 0.6, and 0.12 for the A356 target, the gilding copperjacket, the lead fillers, and the hardened steel core. For initialvelocity, vs ¼ 742ms�1 of the 0.30-cal projectile, the residual

Figure 6. Modeling details: a) 0.30-cal AP M2 projectile, b) 7.62� 51mm M80 FMJ ball round, and c) normal view of the regions with encapsulatedceramic tiles of 0, 4, 6, and 8mm thickness whose deformations are analyzed. The sketches are not to scale.

Figure 7. Unrealistic contacts seen using the penalty and soft-constraint formulations. Simulations (a) and (c) have the same settings except for thecontacting surfaces that resulted in the unrealistic failure of several unloaded elements. This did not occur in simulation (b) that has been used in resultspresented herein.





penetrator velocity, vr, found using a single core, was used todecide the number of processors. As should be clear from resultsshown in Figure 8, a good compromise between the accuracy andthe solution time was provided by 8 and 16 cores for which thecomputed vr differed from the single core value by 1.7%. Thesolution time of the 16-core job was �9% less than that forthe eight-core job.

The GCI and the discretization errors were calculated follow-ing Roache’s method[30] who introduced a constant grid refine-ment ratio, r ¼ h3=h2 ¼ h2=h1, where hi represents the elementsize normalized by the smallest element edge size. It is evidentfrom results for five FEmeshes shown in Table 8 that the relativeerror in vr computed using meshes 3 and 4 is negligible. The FE

mesh 3 was selected for the aluminum target. A similar approachfor the ceramic tiles gave the element edge size¼ 0.5mm.Additional information is provided in Supporting Information.

The element failure due to erosion is known to be sensitive tothe FE mesh.[31] The FS for the aluminum target determined iter-atively using FE mesh 3 and results for shot 4 in Table 2 wasfound to be 0.08 (see Supporting Information for details).Similarly, the FS for SiC tiles using results of shot 6 inTable 2 was found to be equal to 0.5.

Computed results for the target without tiles were first com-pared with the corresponding test findings for shots 4 and 8shown in Table 2. The mathematical and the associated compu-tational models were subsequently validated using results forshots 5 and 6 in Table 2 prior to conducting parametric studies.Results for the target without tiles overestimated vr by nearly 3%.Differences in the two vr values for the target with tiles were 4.6,3.8, and 3.1%, respectively, for the 4, 6, and 8mm-thick tiles(see Table 9 for a summary). For shots using the 7.62mm FMJrounds, the computed vr, respectively, deviated from the experi-mental values by 6.5 and 18.7% for shots 1 and 4. These differ-ences are acceptable in view of the complexity of the problem andthe numerous variables involved. We note that using the plastic-kinematic, the Johnson�Cook, and the piecewise linearly plasticmaterial models for lead did not close the gap between computedandmeasuredvr. Some authors have proposed using the arbitrary

Table 4. Materials for different components and their constitutive relation.

Component Subassembly Material Material model inLS-DYNA

0.3-cal AP M2Projectile

Jacket Gilding copper Johnson�Cook

Fillers Lead Elastic�perfectlyplastic

Core Hardened steel Bilinear plasticity

7.62 mm M80FMJ Projectile

Jacket Gilding copper Johnson�Cook

Core Lead Elastic�perfectlyplastic

Target Lattice Aluminum A356 Johnson�Cook

Ceramic tile Silicon carbide Johnson�Holmquist(JH-2)

Figure 8. The solution time and the penetrator’s residual velocity for dif-ferent numbers of processors. The results indicate that using either 8 or 16processors provides a good balance between the accuracy and the solutiontime.

Table 8. Results of the FE mesh refinement study using the target withouttiles.

Mesh Element size [mm] h Number of FEs r vr [m s�1] tsol [hr]

0 1.333 3.160 249 570 – 530.65 0.64

1 1.000 2.370 354 448 1.333 546.10 1.74

2 0.750 1.778 582 570 1.333 525.40 1.62

3 0.563 1.333 1 028 673 1.333 523.99 4.25

4 0.422 1.000 2 112 759 1.333 523.16 11.73

Table 5. Values of material parameters in Equation (1) for the A356aluminum alloy and gilding copper.

Material FS ρ[g cm�3]

A[GPa]

B[GPa]

C n ν Reference

A356 0.08 2.67 0.149 0.340 0 0.22096 0.33 This work

Gilding copper 0.06 8.94 0.500 0 0.025 1 0.35 [37]

Table 6. Values of mechanical parameters for the cores of the twoprojectiles.

Material FS ρ [g cm�3] E [GPa] σy [GPa] ν Et [GPa] Reference

Hardened steel 0.12 7.85 210 1.400 0.29 15 [38]

Lead 0.6 11.34 16 0.383 0.44 0 [39]

Table 7. Values of the Johnson�Holmquist (JH-2) parameters for SiC.[29]

ρ [g cm]�3 G (GPa) A* B* C* m

3.163 183 0.96 0.35 0 1

n* ε0:(s�1) T (GPa) HEL (GPa) PHEL (GPa) β

0.65 1 0.37 14.6 5.9 1

D1 D2 K1 (GPa) K2 K3 FS

0.48 0.48 204.8 0 0 0.5





Lagrangian�Eulerian (ALE) formulation for lead cores[25] and foraluminum projectiles (e.g., see the study by Schwer andWindsor[32]). We leave this for future study.

3.4. Parametric Studies

We now study the effects on the penetration resistance of thecomposite cast lattice structure of the tile thickness, impact at30� obliquity, bonding of a tile with the casting, tile location withrespect to the casting thickness, and the impact on triple point asseen from the top view.

3.4.1. Effect of Tile Thickness

To quantitatively assess the effect of tile thickness, we setvs ¼ 874ms�1 and consider 0.30-Cal AP projectile. The

simulations revealed that the lattice structure without tilesis not strong enough to defeat the projectile as was foundin physical tests. The AP round is significantly damaged uponcontacting the hard-ceramic tile, as evidenced by resultsshown in Figure 9.

For different tile thicknesses, time histories of the KE of thepenetrator are shown in Figure 10. It is clear that for a fixedtime after impact, the KE is smaller for thicker tiles. At the endof the simulation time, the KE of the projectile striking the8 mm-thick tile encapsulated in the target equaled 640 J, whichis 49% less than that for the target without a tile. This reduc-tion is both due to the decrease in the penetrator speed and itsmass.

3.4.2. Impact at 30� Obliquity

The effect of the 30� obliquity was evaluated only for the8mm-thick tile. Anticipating complete penetration, we includedan additional row of unit cells along the horizontal plane of thetarget (Figure 11).

A slight deviation in the incident angle was noted by inspect-ing the damage to the projectile possibly due to the erosion ofmore elements on the lower portion of the jacket. At the endof the simulation time of 140 μs, the vr of 349m s�1 for the30� obliquity is �5% below 367m s�1 for the 0� obliquity.This difference depends upon the number of truss elementsencountered by the projectile. We note that the simulations

Figure 9. For vs ¼ 874ms�1, deformed 0.30-Cal AP projectiles and the target at 5, 15, and 75 μs after impact and for tile thickness¼ 0, 4, 6, and 8mm.The projectile nose is most severely damaged for the 8mm-thick tile.

Table 9. Comparison of computed and measured residual velocities forselected 0.30-cal AP shots.

Tile [mm]thickness

vs [m s�1] vr [m s�1] Δvr

EXP FEA

0 742 568 591 4.0%

0 817 664 674 1.5%

4 714 390 408 4.6%

6 816 417 433 3.8%





ignored the rotational energy of the penetrator. It has beenreported that the present framework does not accurately predictpitch angles.[33]

3.4.3. Tile Location along the Casting Thickness

The projectile KE at the end of the simulation for the top-mounted, the mid-depth-mounted, and the bottom-mountedtiles, respectively, equaled 795, 828, and 890 J; see Figure 12and Figure 13 that, respectively, show tile’s location and time his-tories of the projectile KE. Note that the projectile’s KE reduction

occurs in part due to its plastic deformations and mass erosion.These results indicate that of the three locations consideredmounting the tile on the strike face of the lattice structure isthe most effective at removing KE from the penetrator.

To shed some light on deformations of the two interactingbodies and decipher the transfer of the impact pressure fromthe projectile to the target, we have shown fringe plots of thevon Mises stress in Figure 14. The top-mounted ceramic tilebears the most impact pressure, severely damages the ammuni-tion upon contact, and retards the penetrator. It is consistent withthe maximum reduction in the KE of the projectile shown inFigure 13.

3.4.4. Bonding between a Tile and the Casting

As mentioned in Section 3, the manufacturing technique used inthis work did not produce a casting-to-tile interface. We hypoth-esized that such an interface would significantly improve theperformance of the cast metal�ceramic composite. Cold-gasspraying was used to create an aluminum�SiC carbide interface,whose tensile strength was measured to be �40MPa.[34]

A simplified version of the ASTM standard C633-13[35] was usedto reproduce this bond strength. The bond without the epoxylayer was simulated using the automatic-surface-to-surface-tie-break contact card (see the Supporting Information for details).The simulation was set by applying fixed constraints at the leftend of the substrate bar and an axial displacement applied to pullthe bar. The computed peak load of 4.5 kN divided by one-fourthof the cross-sectional area of the bar gave the axial stress atfailure¼ 36MPa. The average of the maximum principal stressin all elements abutting the bond surface equaled 39.5MPa,implying that the contact algorithm well reproduced the inputstrength of 40MPa.

Figure 10. Time histories of the KE of the projectile for different tilethicknesses.

Figure 11. 30� obliquity impact on the target with 8mm-thick SiC tile.





The penetration event was numerically analyzed for the targetwith 6mm-thick ceramic tile impacted by the 0.30-cal AP projec-tile at 816m s�1 and 0� obliquity. The lattice-to-tile contact wasmodeled using the automatic-surface-to-surface-tiebreak algo-rithm, and the slave and master side nodes were defined usingsegments instead of part IDs.[28] The simulation predicted about0.5% additional reduction in the final projectile KE. Time histo-ries of the projectile KE for the interfacial tensile strengths of 40,400, and 1000MPa shown in Figure 15 suggest that an increasein the bond strength from 40 to 1000MPa reduces the projectileKE by an additional 7%.

3.4.5. Impact at a Triple Point

Due to the size of casting, several ceramic tiles could beembedded in the mold’s top surface to cover the strike facecompletely. To achieve this in practice, the tiles would needto be distributed in three different planes along with the thick-ness of the lattice arrangement. A theoretical triple-point

Figure 12. Tile position along the lattice structure thickness in the impact direction.

Figure 13. For three tile locations, time histories of the KE of the 0.30-CalAPM2 projectile.

Figure 14. Fringe plots of the von Mises stress: a) target without a tile, b) target with top-mounted tile, c) target with mid-mounted tile, and d) target withbottom-mounted tile. The range for the fringe plots was fixed from 0 to 0.5 GPa, meaning that the von Mises stress exceeding 0.5 GPa is shown as0.5 GPa.





impact was envisaged for this work. A symmetric version ofthe complete casting described in subsection 2.2 was designedto include three ceramic tiles along the impact direction, asshown in Figure 16. The FE mesh size and the contact setswere kept the same as those described in subsections 3.2and 3.3. The complete geometry of the 0.30-cal AP projectilewith vs ¼ 816ms�1 was modeled, and it was positioned to con-tact the three tiles along its trajectory. The FE mesh for thisconfiguration had 5 120 428 tetrahedral elements and only 1simulation was conducted.

At the end of the simulation, the projectile was significantlydamaged with a considerable change in the angle of its trajectory.The time histories of the projectile KE shown in Figure 17 revealthat initially, the single-tile configuration is more effectivebecause of the impact occurring at its center and more of theceramic material damaged. However, after the projectile strikesall three tiles located at different depths, the triple-point impactreduces more KE of the projectile, resulting in the final projectileKE of 484 J that is 39% less than that for the projectile impact on asingle tile.

Figure 15. For different strengths of the casting/tile interface, time histories of the KE of the 0.30-cal AP projectile impacting the target at normalincidence.

Figure 16. Triple-point impact simulation and the state at the end of the simulation time. The arrows indicate boundaries of the planes of symmetry.





4. Conclusion

We have used a hybrid AM�metal-casting approach to fabricate alightweight metal�ceramic composite lattice structure madefrom aluminum alloy A356, subsequently treated to the T6 con-dition. Provisions for tiles were designed in the AM sand molds,and silicon carbide tiles were embedded in the aluminum alloy toadd penetration resistance to the soft aluminum alloy matrix. The225� 225� 62mm castings with 4mm-diameter trusses wereimpacted primarily at normal incidence by 0.30-cal AP M2and 7.62mm M80 FMJ projectiles at about 750m s�1. Real-timeX-ray and SEM imaging were used to evaluate the quality of cast-ings before and after the penetration tests. The SEM images pro-vided information regarding porosity in the cross section of thinportion of the castings that was consistent with the solidificationprocess of metallic alloys. Due to the small sizes of pores, theywere not explicitly considered in simulations using the FEM.However, their effect was incorporated using values of mechani-cal parameters derived from the data of simple tension tests onthe cast metal specimens.

Results of high-velocity impact tests indicated that the castmetal�ceramic composite lattice structures are effective in sig-nificantly reducing the projectile incident KE and hence itsspeed. The targets with 8mm-thick ceramic tiles proved to bemost effective but with a weight penalty of about 3% comparedwith targets with 6mm-thick tiles. Based on the penetration testresults, a casting with 6mm-thick ceramic tiles represents a goodcompromise between the overall weight and performance.

Axisymmetric deformations of the cast metal�ceramic com-posite lattice structure and the projectiles were analyzed usingthe commercial explicit FE software LS-DYNA. The FE meshwas successively refined for one problem to get a converged solu-tion, and this mesh was used for subsequent studies. For mostproblems studied, the computed residual KE is close to that inthe tests. Computed results for the 0.30-cal AP M2 projectileagreed well with the experimental findings and those for the7.62mm M80 FMJ ball rounds had a maximum deviation of19% for the penetrator residual speed.

Effects of the angle of obliquity of impact, the tile thickness,the strength of the bond between a tile and the substrate, andlocating three tiles at different positions along the target thick-ness to provide a triple-point impact have been scrutinized.The triple-point arrangement was found to be very effective indiminishing the KE of the projectile but with a higher weightpenalty. For a ceramic-to-metal interface to significantly improveperformance, bond strengths must be much higher than thatachieved by a cold plasma spray.

The cast metal�ceramic composite lattice structure is mosteffective when the ceramic tile is mounted on the strike faceof the target rather than in the interior, indicating that more vol-ume of the lattice supporting the tile results in higher energyabsorption. Tile arrangement can potentially be improved toenhance the energy-absorbing characteristics of this cast metal–ceramic lattice structure, for example, by creating a pattern sothat the penetrator interacts with several different ceramic tilesthrough its trajectory.

Supporting InformationSupporting Information is available from the Wiley Online Library or fromthe author.

AcknowledgementsThe DLA-Troop support, Philadelphia, PA, and the Defense LogisticsAgency Information Operations, J68, Research and Development, Ft.Belvoir, VA, provided funding for this work. The authors acknowledgeAdvanced Research Computing at Virginia Tech for providing computa-tional resources and the associated technical support. Special thanksto Dr. James M. Kennedy for his kind help and advice in using the softwareLS-DYNA. The authors would also like to acknowledge the contributions ofMs. Anna Schierlmann, Mr. Luke Hall, and Mr. Mike Holmes in success-fully preparing test specimens as a part of their senior design project.Scanning electron microscope imaging was done at Virginia Tech’sNanoscale Characterization and Fabrication Laboratory. In addition, theauthors acknowledge Mr. Glover Kerlin from Bombardier RecreationalProducts for his support with real-time X-ray imaging and NationalTechnical Service for conducting the penetration tests reported herein.

Conflict of InterestThe authors declare no conflict of interest.

Data Availability StatementData available on request from the authors.

Keywordscast aluminum, composite lattice structures, finite element analyses, high-velocity impacts, nonlinear deformations, penetration resistance

Received: May 11, 2021Revised: July 25, 2021

Published online:

Figure 17. KE time-history plots: a) strike at the triple point and b) normalimpact on a single tile, mounted on the strike face of the target.





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penetration resistance of cast metal−ceramic composite

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