STRUCTURAL ANALYSES & EXPERIMENTAL ACTIVITIES SUPPORTING THE DESIGNOF A LIGHTWEIGHT RIGID-WALL MOBILE SHELTER

Paul V. Cavallaro*U.S. Naval Undersea Warfare Center

Newport, RI 02481

Melvin leeU.S. Army Natick Soldier Research, Development and Engineering Center

Natick, MA 01760

ABSTRACT

Lightweight rigid-wall shelters used in mobile militaryoperations are constmcted of sandwich panels comprised ofthin face sheets and thick, ultra light core materials tominimize weight and maximize stmctural integrity.However, such lightweight constmction comes at a cost,often impacting the design and manufacturing of criticaljoints connecting the sandwich panels in a box-likeassembly. Furthermore, joint stiffnesses are often difficultto characterize and their finite values significantly influencepanel deflections and rotations. Although mobile rigid wallshelters must be certified for several transport loadingenvironments, this effort, combines experimental andanalytical approaches at material and sub-structural levels to(l) generate accurate modeling methods, (2) validatematerial- and sub-stmctural models and (3) virtuallyevaluate the shelter's structural performance whileminimizing costly physical testing. Material-level testsfocused on the mechanics of the assembled constituentsforming the sandwich panel and the benchmarking ofappropriate finite elements to predict displacements, stressesand strains. The sub-structural level tests focused on loadinga stmcturally representative shelter section to determine thejoint behaviors and stiffnesses for model benchmarkingpurposes. Finally, a complete rigid-wall shelter model wasconstmcted for evaluating future static and dynamic loadcases as required for certification.

1. INTRODUCTION

A lightweight rigid wall shelter was constructed formounting on various military vehicles such as the HighMobility Multi-Purpose Wheeled Vehicle (HMMWV) asshown in Fig. (1). The shelter was primarily constmcted ofthin aluminum face sheets (skins) adhesively bonded to arelatively thick paper honeycomb core forming a sandwichpanel constmction (SPC) interconnected with variousaluminum extmsions, weldments,. mechanical fasteners andadhesive bonds. Stmctural advantages result from thedecoupling of bending and shear behaviors between facesheet and core materials. That is, the bending stiffness ismaterially dependent upon the face sheets while the shearstiffness is materially dependent upon the core. No

appreciable bending strain energies develop in the core and,likewise, no appreciable shear strain energies develop in the facesheets. The decoupling of these behaviors enables the designerto achieve a level of tailorability unmatched by homogeneousmaterials but analogous to fiber reinforced laminates.

Fig. I Lightweight rigid-wall shelter shown mounted on aHMMWV.

The face sheets were 0.025" thick 6061-T6 aluminum exceptin the wheel well region where the face sheets were 0.0 IS" thick2024-T4 aluminum. Strength properties for these alloys[lj areshown in Table (I). Hexcel® WR-II Kraft®-coated paperhoneycomb core[2] was used. This grade had a 3/8" cell size anda 2.5 Ibs/ft3 weight density. Compression and transverse shearproperties are shown in Table (2). The face sheets and core werebonded together using a structural film adhesive.

Table I Strength properties of aluminum face sheets.

Table 2 Properties of 1.210 inch thick WR II honeycomb.ComDression Properties (DSI Transverse Shear Properties (PSI

WR II Honeycomb Bare 11: Stabilized Across ~ibbon Directionl~ong R~ibon DirectionCore Strenmh SlrenQthTModulus StrenQth Modulus SlrenQth Modulus

3!il" cell. 2.51blft' 2a:J1 340 133.000 141 I 13,000 I B3 I 7{m

Two primary sandwich panels, folded into 3 planar sections(one panel formed the floor, front end wall and roof and thesecond panel formed the road side wall, door end wall and curbside wall), generally comprised the shelter. This enabled thepanels to be aligned to net shape for subsequent joining usingmechanical fasteners, closeout extmsions and weldmentsforming the closed box configuration. Using two primarysandwich panels reduced the number of joints and allowed forease in manufacture.

A critical step enabling rapid shelter development was theintegration of experimental validation tests with analytical

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4. TITLE AND SUBTITLE Structural Analyses & Experimental Activities Supporting The Design OfA Lightweight Rigid-Wall Mobile Shelter

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methods. A test plan was developed to characterize loadsharing mechanisms, panel shear and flexure stiffnesses,deflections and stress states at both the material andstructural levels.

2. MATERIAL LEVEL TESTING

The material level tests focused on the mechanics of anadhesively bonded aluminum/honeycomb sandwich panelsubjected to flexure loads. Four-point flexure tests wereperformed on six 20" long by 4" wide by 1.26" thicksandwich beams as shown in Fig. (2). Four-point rather thanthree-point loading was selected to mitigate failures bylocalized core crushing prior to the onset of face sheetyielding and to provide a gage section region subjected toconstant bending stress with no transverse shear stresses.

1-----2\1.1lXI-----------------..<Cltr.\

point of interest, and area moment of inertia, I, for asandwich beam having total thickness 11010/:

The maximum face sheet longitudinal stress, O'lIIan isobtained at locations of maximum moment and distancefurthest from the neutral axis. For the case of four-pointflexure loading of symmetric sandwich beams, Oil/atbecomes:

(6)

v= shearing force-le /2 :0; Ye:O; +le/2

The transverse shear stress, rtze , in the core as a function of thethrough-thickness variable Ye is then given as:

rxzc(ye>:= 2~s [EJI Jd + EcCl- y~ J]where:

If Ec « Ef , then the transverse shear stress of the honeycombcore is constant through the thickness as the product involving Eebecomes negligible. The transverse shear stress distribution ineither face sheet (both face sheets assumed to be of equalthickness and material) as a function of the through-thicknessvariable yfis:

r _~[&c2+4ICIJ+I/-4y/)EJ] (7)xzJ(YJ) - D 8

s

(I)

(2)

-llo/al/2 .sY .s+llOla/2

MyO'.u :=T

where:

For an isotropic material, the face sheet longitudinalstrain, Clllan corresponding to Oil/at is obtained by assuming auniaxial stress field such that:

Obel/d = 0.198 in ~olal = 0.291 inrue = 131 psi

Osheor = 0.093 in,Cyield = 0.004 in/in,

The maximum value of r')loccurs at the inner surface of the facesheet. To determine the applied load required to initiate yieldingof the face sheets, the yield stress of aluminum 6061-T6 (O'yield =40,000 psi) was substituted into Eq. (I) and solved for Plllat uponwhich Pyield = Plllat.

where:

The corresponding deflection terms, axial face sheet strain[Ef = lOx10

6 psi] and honeycomb transverse shear stress [forWR-II-3/8-2.5 honeycomb along the ribbon direction, Ge =13,000 psi and rue ollowable = 141 psi (thickness correction factorincluded on transverse shear strength] at yield are:

Pyield = 1,295.85 lbs.

The transverse shearing deformation contributed 32% toward thetotal mid-span deflection. The above strain represents themagnitude of strain in both the tensile (lower) and compressive(upper) face sheets since the sandwich is symmetric and linearelasticity (small deformation) is assumed. The continuity of ruacross the face sheet/core interface is checked by solving (7) at Yf= 0 and comparing to the value of rtye'

(3)

(4)

Exx = Young's Modulus in x-direction

Ie = core thickness,If = face sheet thicknessW = width of cross sectiona = distance between load & support pointPlllat = maximum applied machine load

& == (Ymaxmax E

:ex

where:

where:

Castigliano's Second Theorem was used to derive anexpression for maximum lateral deflection based onconsidering the combined strain energies due to bending ofthe face sheets and transverse shearing of the honeycombcore. This is equivalent to a Timoshenko beam solution inwhich the effect of transverse shear deformation on lateraldeflection is included[61. The maximum deflection, ~olal, isthe sum of the mid-span contributions from the face sheetbending deformations and the core transverse shearingdeformation:

Pa 3b 2 - a 2 3Pa()/o/al := E +---

JW Ic +21J 5Gcwlcwhere: b = distance between load points

Ge = core transverse shear modulus

The flexural rigidity, D" of the sandwich is computedusing the parallel axis theorem and Young's Modulus foreach layer as:

The effect of core transverse shear stiffness, Ge, on the midpointdeflection is shown in Fig. (3). Note that Obelld equals thedifference between the two curves shown.

3E/WIJ EJWIJ 2 E W 3D := ---+--d +_c_I

s 6 2 12 e(5)

where: d = distance between face sheet mid-planes

3

........ Slt.UfO.tedlOhal'"~

-6- T~lill M:;f.~ltlO~tbe"

G.=13.ooo PSI a!onll ribbm dnc1io1 dWR-Il...J.'8-25Hooe)'Comb

Hont'Ycolllb Cor. TrUlsveru She",Modulul. Gel (Pll)

Fig. 3 Effect of honeycomb core transverse shear modulus,Ge, on mid-span deflections.

The NISA[7j finite element program was used to modelthe four-point bend test specimen. Two models weregenerated using namely the sandwich element (type-33) andthe solid composite element (type-7). The sandwich elementwas an 8-noded laminated shell element that supportedmembrane-bending coupling and transverse sheardeformations. Each node had 6 degrees of freedom(DOF's), 3 translations (u], 1t2 and U3) and 3 rotations (aj,a2, and a3)' However, the rotational DOF normal to theelement surface, a3, commonly referred to as the drillingDOF, possessed no rotational stiffness. This element wasdeveloped for thick shells with two or more face sheets andone or more cores. Figure (4) shows the stress componentsfor a general 3-D laminated composite.

Fig. 4 General 3-D stress components for laminatedcomposites.

The face sheets were restricted to states of plane stress(o:~" O)?" and O:,y) and the core supported only transverseshear stresses (axz and a yz). At the interface between theface sheets and the honeycomb core, the interlaminar shearstresses O"z-t and 0:"), were of particular interest. Thesestresses represented the in-plane shear stresses within theadhesive layer and, when compared to the allowable shearstrength of the adhesive, can be used to predict adhesiveshear failures. The solid composite element was a laminatedbrick element formulated using general 3-D states of stress(0"", O)?" O;z' O"\)" 0"", and 0),,) with each node having 3translational DOF's only. The layers of both the sandwichelement and the composite solid element were assumed

perfectly bonded. Although the sandwich element did notsupport through-thickness deformations, which were expected tobe negligible for global shelter models, it was consideredsignificantly more computationally efficient than the solidcomposite element. Use of the solid composite element for a fullshelter model would not be practical due to the mesh refinementrequired. The solid composite element was preferred in localizedmodels where 3-dimensional stress states and through-thicknessdeformations were prevalent such as in joints, weldments,adhesive bonds, and fastener regions.

Displacement and stress results of the sandwich elementmodel subjected to a total load of 1,295.85 Ibs are shown in thecontour plots of Figs. (5-7). At maximum load, the totaldisplacement of 0.278 in. occurred at the mid-span and varied byonly 4.5% of the analytical solution provided by Equation (4).The magnitudes of the maximum axial stress,o:w in the tensileand compressive face sheets were identically equal to 40,400 psi,which varied by only 1.0% of Equation (2). The maximumtransverse shear stress,O"", in the honeycomb core, remote fromlocalized effects of the load points, was 138.0 psi, which variedby 5.4% of the predicted value from Equation (6).

Fig. 5 Transverse displacement contour plot.

Fig. 6 Axial stress, am plot for the lower face sheet.

Fig. 7 Transverse shear stress, a", plot in the honeycomb.

The solid composite element model shown in Fig. (8)captured the through-thickness deformations at the load andsupport points. For the same maximum load used in the

4

02

Fig. 8 Vertical displacement plot from the composite solidelement model at 1,295.85 Ib total load.

RcwF(106)

loading Bar

loading Bar

RowD(I"'1

Ends of Floor"O"·(nusionc,""""

A

Row'"ll.')

Fig. 11 Section view of representative shelter test.

IFace Sheet Tensile FailuresHate: S1r.il oages#1,).,3 .ebondedto1hesllekr Merf.ceshedsInd str.'" glge5 #-1,5,6 a'e bMded kJ the sheltr i'lner f.c:eshed:s.

S1.... lngages .. R.owM"-Ue"taltedmfyfo'"tesb~s2,.1.

5. SUB-STRUCTURAL TESTING

Fig. 10 Shelter section test and strain gage map.

RO'lifBt1-')

RcwCtt·GI

The test specimen shown in Figs. (10-11), along with itsrestraints and four-point loading arrangement, was designed togenerate flexure stresses along the shelter's hoop direction sinceflexure is generally the dominant mode ofloading. It representedone lateral half of the shelter without the forward and rear walls.

The honeycomb ribbon direction was oriented parallel to theshelter hoop axis. As with the beam tests, four-point rather thanthree-point loading was selected to mitigate failures by localizedcore crushing prior to face sheet yielding. Two clampedrestraints, located at each end of the floor "D"-extrusion (foreand aft of the wheel well), were necessary for generatingsubstantial flexure stresses at the side wall/wheel well interface.Absent these clamped restraints, analysis demonstrated that: (1)flexure stresses within the side wall/wheel well region were

....•.•...•.29t .•....... o.m ....0.15 0.319 0.293 Ut20-'"

0.05

4

Specimen Number

0.1

-c~. ----::~D.3~~:::f-~---~---~-·t.. ·=···~···~-·~···~···:j:::t=~~~···~···~··~···=···=···~··~···~···~···~··~···~···~···I=~Deme.nt 0.25 ....

sandwich model and closed fonn solution, the maximumdisplacement of the composite solid element model was0.282 in. at the mid-span. However, this model did notinclude the effects of plasticity (stress softening) due tolocalized yielding of the compressive face sheet, which wasevident in the tests. Post-test inspections revealed that thepennanent indentations made at either the load or supportpoints ranged in depth up to 0.06 in. The maximum axialstresses, O:\:

Total Applied Load (fbS)

Fig. 12 Displacement vs. load for section test #3.

1I:',.... $tnln Corn·blion . Rol.lt Awoo

'2 .FEA

J~~~1illJ~ u

w U LU" _rest'lDTEsr12

~clES1'~

OAVO TESt

"'~ ..., , . .C'yV5nln Correl~tion·Row B

--~.W.rEST"oT£SUJi~2_ orEST.)Q.4.VO fEST

Ilrl Stnin eon~iJtIcn • ftow C.- CHQ.T5T'1OT61"I~ OTESTf'lliA'''') TESt, z. ••ttY Str;1r. COfTd,uon· RowD_._- DFEA

a;eST"

olEStf2

i~=:~ o TUU)SAW iE$14;1ClU

1 1 ) 4 ~ •

,,+-----------------~o 2,00) J.JXXJ 6»):) 8J)X1 '0.000

95

Total/l,pplied load lib, I

Fig. 13 Plot of rotation angles vs. load for test #3.

Fig. 14 Example comparisons of experimental and FEA Eyystrains for rows A-D.

The first section test included 72 strain gages located in rowsA through L. The maximum applied load was 9,935 Ibs. Strainsat rows D and F were mostly dominated by bending but, bycontrast, were less than 25% of Row E. This suggested that the"D"-extrusion joints along the upper and lower side wall edgestransferred less moment than expected for rigid connections.This counter intuitive observation rendered the restraining effectsof these joints to be rotationally limited. That is, the "D"-extrusion joints behaved closer as pinned rather than rigidconnections. The rotational stiffness at the wheel well-to-floorjoint was comparable to those of the side waIVroof and sidewall/wheel well joints. Displacement readings from the roof andside wall LVDT's at maximum load were 2.493" inward and0.104" outward, respectively. A localized tensile fracturethrough the floor inner face sheet and the blast hoop stiffeneroccurred inches forward of the wheel well region as shown inFig. (10). Although some plastic deformation of the side wallwas evident, there were no face sheet/core delaminations orfractures of the honeycomb core.

12pal'Opal8,000

ROOF

6.0004 pal2.000.Q5+----~--~--~--~---~-~

o

minimal and (2) the floor panel would fail by excessivebending deflections, which were not prevalent incertification tests because the floor contacts the vehicle'scargo area floor. The edges of the floor and roof panelswere bolted to the test frame with clamping plates to providenear rigid boundary conditions. Using a test frame andhydraulic actuator, two line loads were quasi-staticallyapplied to the side wall using the loading bars as shown inFig. (10). The section was oriented in the test frame so thatthe side wall was positioned horizontally. A sphericalbearing connected the loading assembly to the test frame.Since the bearing was not capable of resisting torques orbending moments, the loads applied along the two load barswere equal. Stiffness of the overall loading assemblyprevented deformations of the bars along their lengths.Therefore, all points of the side wall in contact with aspecific loading bar deflected the same amount as thatparticular load bar. However, this does not infer thatdeflections under both load bars were equal. A preliminaryFEA model was used to identify locations for positioning upto 76 strain gages. Strain gages were bonded on the innerand outer face sheet surfaces as shown in Fig. (10). Straingages I & 2 were located at the forward end of the testsection, 3 & 4 were located at the middle of the test sectionand 5 & 6 were located at the aft end of the test section.Strain gages 1,2 and 3 were located on the outer face sheetsand 4, 5 and 6 were located on the inner face sheet surfaces.All gages were uniaxial and oriented along the shelter hoopdirection. The center deflections of the side wall and roofpanels were recorded using linear variable displacementtransducers (LVDT's). Results of the three section tests areshown in the deflection, joint rotation and strain gage plotsof Figs. (12-14). Generally noted, behavior was linear withrespect to load. All strain gage rows except A and Hdemonstrated bending dominance. Rows A and H weresubjected to appreciable membrane strains in comparison totheir bending strains. The presence of membrane strainsindicated load stiffening effects that introduced geometricnonlinearities at approximately 5,000 Ibs.

LVDTNI

15 ~I LVDTN2l

I-+-LVDTi12 --+-LVDT#1125

05

6

Instrumentation of the second test section was modifiedto include additional strain gages located at row M. Themaximum load during the second test was 10,422 Ibs. with atensile fracture of the inner face sheet and hoop stiffeneroccurring directly through strain gage M3. This failurelocation was symmetrically opposite that of the first test.Strains were consistent with those obtained from the firsttest. The maximum roof and side wall displacements were2.859" inward and 0.051" outward, respectively. Plasticdeformations were evident in the "D"-extrusions of theroof/side wall joint, side wall and floor panels. There were,however, no delaminations between the face sheets andhoneycomb core or fractures of the honeycomb.

Two modifications were made to the testing of the thirdsection specimen. Dial indicators were used to measurerelative motions at the five locations shown in Fig. (13) sothat changes in joint angles could be calculated. Themaximum load was 10,441 Ibs. with a tensile failure of theinner face sheet and hoop stiffener occurring adjacent tostrain gage M3. Deflection readings for the LVDT's were2.574" inward for the side wall and 0.058" outward for theroof. Post-test inspection revealed no delaminationsbetween the face sheets and honeycomb core and no corefractures. Joint and panel rotation angles were plotted inFig. (13) as a function of load. The maximum-recordedchange in joint angle was -6.3°, which occurred at theroof/side wall joint at 6,950 Ibs. This was not, however, themaximum load but the load at which the dial indicatorslipped and no further measurements at this location werepossible. A linear extrapolation predicted that the roof/sidewall joint rotated by -7.0° had the dial indicator remained inthe proper position. The floor panel rotation was 19.6° atmaximum load.

Failure modes for each of the 3 section tests wereidentical, occurring at symmetric positions across the floor.Failure loads varied by less than 5% of their average. Straindistributions were repeatable. Deflection readings from theroof and sidewall LVDT's were linear up to the onset ofyielding at the fracture regions. Strain results of eachsection test counter intuitively revealed that the "D"-extrusion joints were compliant rather than rigid asoriginally anticipated. The observed limited rotationalresistance was expected to have an impact on the strain anddeflection behavior of the full shelter.

A linear elastic model of the sub-structural testspecimen of Fig. (10) was developed to simulate the testbehavior and to match the strain gage and LVDT deflectionresults. The model included 2nd-ordered sandwich elementsfor the panels, beam elements for the extrusions and springelements for calibrating the joint rotational stiffnesses. Anominal 10,000 lb. total load was applied and the jointrotational stiffnesses were adjusted until the model matchedthe experimental strain, deflection and joint rotation results.

Correlation of the experimental versus FEA strains is shownin the bar charts of Fig. (14). The LVDT and FEA deflectionresults were tabulated in Table (4). The joint rotationalstiffnesses, KROT, used in the FEA model are shown in Table (5).These stiffnesses provided the necessary calibration to bestmatch strains and deflections between the model and tests.

Table 4 Experimental & FEA deflections.a.nection Results At 5,000 lb. Total Load

(Negative indicates extward)

SKle'MIl Roof

FEA 0.878 -0.121Testt#1 1.009 -0.079

Test *2 0.809 -0.105

Test 13 1.014 0.055

However, comparisons between experimental and FEA resultswere shown for a load of up to 5,000 Ibs. Beyond this load, thenonlinearities observed at strain gage rows A and H could not bereflected by the linear elastic model. The bar charts ofexperimental and predicted strains shown in Fig. (14) generallydemonstrated a high level of correlation.

Table 5 Section test joint rotational stiffnesses.Joint Location Description KROT (In-lbJrad)

RoorTo Side Wan 2.500

Side Wal To 'Mleel Well 2.500

Wheel We8 To Floor 2.500

Side Wal To Floor 5.000WheelW~ Horizontal To floC)( 2.500

Side Wal To Vertical lMleel Well 5.000

Floor To V\lheel Wei Fore & Aft 2.500

Clamped Edges 55.000

5. SHELTER MODEL FOR MODAL & DYNAMIC CASES

A complete structural model of the shelter was developed asshown in Figs. (15-16). This model included all sandwich panelsections, extrusions, joint rotational stiffnesses, and rigidboundary conditions that simulated the shelter-to-vehiclemounting attachments.

Fig. 15 Complete shelter FEA model.

Each of these was required for predicting the response of the fullshelter to various loading events when mounted to the HMMWV.Rigid link elements represented the hinges and strikermechanisms connecting the door panel to the door end wall. Thedemarc panel elements were connected to the sandwich elementsand closeout extrusions by using rigid links. The second orderedsandwich and linear beam elements used in the section modelwere used for the full shelter model.

7

BEHTCOR",ERROTATlOIW.

SPRlNOB.ENEHfS -..........

EXTR161Ct1fOftfUTURE

nJHNaoro"""

~~~:iiTsIOOIC. AROUtD DEMAAC

PANas

HNCE & LATCHBARR:OIO..'"

(12)

The eigenvalue analysis was performed using the LanczosMethod[8] with an upper cut-off frequency,!c, of 100 Hz. Thecomputed mode shapes (eigenvectors) were flexure dominantwith symmetric and anti-symmetric deformations. Thefundamental natural frequency,j], was a symmetric flexure modeof the roof at 30.45 cycles/sec (Hz). Plots of the first 4 modeshapes are shown in Fig. (17). Using WI = 191.32 rad/sec and W2= 227.33 rad/sec corresponding toj] andfi, respectively, a and ~were computed as a = 10.389 , fJ = 2.389xl 0-4 .

Fig. 16 Description of beam, spring and rigid link elements.

6. MODAL & DAMPING ANALYSES

For dynamic events, the time-based equations of motionshown in equation (8) included effects due to damping. Thedamped behavior of the shelter was determined byconducting an eigenvalue analysis that established naturalfrequencies of vibration, /;, and corresponding mode shapesso that proper damping values could be established.

Rayleigh damping[8J (also known as proportionaldamping) was assumed in accordance with equation (9),which reflected contributions from both mass-baseddamping, a, and stiffness-based damping,~. In general,stiffness-based damping results from hysteretic effectsobserved during cyclic loading of elastic materials.Additional sources of structural damping were expectedfrom mechanically fastened joints and face sheet/honeycombcore adhesive layers. Mass-based damping affects thedynamic response at lower frequencies while structural-based damping affects dynamic response at higherfrequencies. A recommended critical damping ratio, ~, of5% was used for all flexure modes.

[M]k +[ck+[K]K ={j(t)}where: [M] = global mass matrix

[C] = damping matrix{f(t)}= vector of nodal forces[K] = global stiffness matrixX vector of nodal displacements

X = vector of nodal velocitiesX = vector of nodal accelerations

where:

(8)

(9)

(10)

(11)

Fig. 17 Modal analysis results of first four mode shapes.

7. CONCLUSION

A modeling approach, incorporating minimal multi-leveltesting, was described and demonstrated for validating the designof a lightweight, mobile military shelter manufactured usingsandwich panel construction (SPC) methods. This approachprovided a rapid, cost-efficient alternative to the traditional"build-test-build" method. The end product of the current workis a validated, full-featured shelter model.

8. REFERENCES

1. American Society of Metals, "Metals Handbook", 1985.2. Hexcel Corporation, "Mechanical Properties of HexcelHoneycomb Materials", TSB #120,1992.3. Vinson, I.R., "The Behavior of Sandwich Structures ofIsotropic and Composite Materials", Technomic Pub, 1999.4. Allen, H. G., "Analysis and Design of Structural SandwichPanels", Pergamon Press, Oxford, 1969.5. Ugural, A.C., Fenster, S.K., "Advanced Strength and AppliedElasticity", Elsevier, 1977.6. Timoshenko, S., "Strength of Materials", Nostrand 1953.7. NISA, Ver. 11.0, Engineering Mechanics ResearchCorporation, Troy, MI.8. Bathe, KJ., "Finite Element Procedures in EngineeringAnalysis", Prentice-Hall, Englewood Cliffs, NI, 1982.

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Structural Analyses & Experimental Activities Supporting The Design Of A LightweightRigid-Wall Shelter

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Author I Originator

Paul V. Cavallaro, C70T

./' Sponsor Approved? (Check one)

./~~ 6//A7 ~Yes ONo DNATechnical Reviewer / A .f ,//ld .If/,V~ •• ,All M. Sadegh, ONR Summer Faculty, C70T fA Vf/f( ';