MTL TR 92-65 AD-A256 506 JADI11i111,1111 111111111 il !111i1111,111111

EVALUATION OF THE MODIFIED FORWARDHMMWV LIFT PROVISION IN DUAL SIDE-BY-SIDEAIRLIFT CONFIGURATIONS DTIC

H-L ECr -rr4 OCT 2 0 1%92

ROBERT B. DOOLEY, PAUL V. CAVALLARO,

KRISTEN D. WEIGHT, and CHRISTOPHER CAVALLAROMECHANICS AND STRUCTURES BRANCH

September 1992

Approved for public release; distribution unlimited.

.4092-27318/y

Sponsored by

U.S. Natick Research, Developement, andEngineering Center

Natick, MA

US ARMYLABORATORY COMMAND U.S. ARMY MATERiALS TECHNOL(t,,jY LABORATORYWrERIALs TICHNOLOGY LABOAAMO0Y Watertown, Massachusetts 02172.0001

The findings in this report are not to be construed as an official

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REPORTDOCUMENTATIONPAGEBEFORE COMPLETING FORMI, REPORT NUMBER 2, GOVT ACC9SSION NO: 2, REGIPIENfli CATALOG NUMBER

MTL TR 92-654. TITLE (and Subtitle) 6, TYPE OF REPORT & PERIOD COVERLD

EVALUATION OF THE MODIFIED FORWARD HMMWV LIFT Final ReportPROVISION IN DUAL SIDE-BY-SIDE AIRLIFTCONFIGURATIONS C PERFORMING ORn, REPORT NUMBER

7. AUTHOR(a) 47 CONTARACT OR GRANT NUMBER(s)

Robert B. Dooley, Paul V. Cavallaro,Kristen D. Weight, and Christopher Cavallaro

9. PERFORMING ORGANIZATION NAME AND ADDRESS 10. PROGRAM ELEMENT. PROJECT. TASK

U.S. Army Materials Technology Laboratory AREA& WORK UNIT NUMBERS

Watertown, Massachusetts 02172-0001 D/A Project IL162105AH84ATTN: SLCMT-MRS

It. CONTROLLING OFFICE NAME AND ADDRESS 12. REPORT DATEU.S. Army Natick Research, Development, and September 1992Engineering Center ,I. NUMBER OF PAGESNatick, Massachusetts 01760-5010- 35

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IS. SUPPLEMENTARY NOTES

19. KEY WORDS (Continue on reverse side it necessary and identIfy by block number)

High mobility multipurpose wheeled Field testsvehicles (HM5WV) 7inite element analy-sis

External airlift trinsport (EAT)Airlift configuration

20. ABSTRACT (Contlnue an reverse side IH necessary and identify by black number)

(SEE REVERSE SIDE)

FORM

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Block No. 20

ABSTRACT

Field modifications to U.S. Army External Airlift Transport (EAT) operat-ing procedures became critically necessary during Operation Desert Storm. Tomeet mission essential objectives, the U.S. Army 101st Airborne Divisiondeveloped a new dual side-by-side (DSS) airlift sling configuration for air-lifting two high mobility multipurpose wheeled vehicles (HMMWV) simultaneouslywith a CH-47D helicopter. The U.S. Army Materials Technology Laboratoryperformed experimental and finite element analysis (FEA) activities to evaluatethe performance of the forward outboard ILM1WV lift provisions subjected toDSS sling configurations. These activities were performed at the request ofthe Natick Research, Development, and Engineering Center in an attempt to-,.rtifv two DSS airlift configurations under consideration for further us,.

R,.sults of both the experimental and analytical analyses were obtained, dis-":'issed, and subsequently correlated. Conformity of the provision to MilitaryStandard 209-G, "Slinging and Tiedown Provisions for Lifting and Tying DownMilitary Equipment" was evaluated for both configurations.

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CONTENTS

INTRODUCTION ................................................ 1

DESCRIPTION OF HHV AIRLIFT PROVISION ........................ 3Lift Provision Construction............................... 3Lift Hook Loading Design Criteria ......................... 4

SIMULATED AIRLIFT TESTS ..................................... 5Simulated Airlift Test Procedure .......................... 5

CHARACTERIZATION OF DSS AIRLIFT CONFIGURATION ............... 5Results of Configuration Characterization .................. 7

STRAIN DATA ACQUISITION ..................................... 9Strain Data Processing and Results ....................... 11

FINITE ELEMENT ANALYSIS .................................... 13Linear-Elastic Finite Element Hook Models ................. 14Results of Linear-Elastic Finite Element Hook Models ..... 16Correlation of Linear-Elastic FEA Model Results toExperimental Results ..................................... 16Elastic-Plastic Full Provision Finite Element Model ...... 19Elastic-Plastic Model Results ............................ 19

MODEL SENSITIVITY OBSERVATIONS AND ANALYSIS ................. 24

CONCLUSIONS ................................................ 27

ACKNOWLEDGEMENTS ........................................... 28

BIBLIOGRAPHY ............................................... 29

APPENDIX I ................................................. 30

S.. T I S

r

* T

• .'' ' '-'• C,4 ¢ OS

INTRODUCTION

An Army External Airlift Transport (EAT) procedure forground vehicles consists of a CH-47D helicopter transportingtwo HMMWVs in the tandem configuration depicted in Figure 1.The CH-47D employs a pair of fore and aft lift points. Eachlift point supports an individual vehicle using foursling/chain assemblies. Opposite ends of the sling/chainassemblies are attached to the two front airlift provisionsand the two rear bumper shackles of each HMMWV.

VP

Photograph, courtesy of ARMED FORCES JOURNAL INTERNATIOAPL, April, 1992.

Figure 1. CH-47D lifting 2 HMMWVs in tandem configuration.

The tandem configuration was found to interfere withthe operation of the CH-47D altimeter. Signals emitted fromthe altimeter system reflected off the forward positionedHMMWV rather than the ground resulting in erroneous altitudereadings. Interference of this type poses a serious threatto the safety of both the pilot and crew, particularlyduring low altitude night operations.

In an effort to alleviate the altimeter interferenceproblem, the 101st Airborne Division designed and employed anew, uncertified Dual Side-by-Side (DSS) lift configurationas illustrated by Figure 2. Sling rigging procedures forthis configuration applied a significantly increased out-of-plane bending load to the forward outboard (FO) HMMWV lifthooks as compared to both the single and tandem airliftconfigurations. Airlift provisions in all cases are requiredto conform to MILITARY STANDARD 209G (MIL-STD-209G)"Slinging and Tiedown Provisions for Lifting and Tying DownMilitary Equipment"'. Specifically, no lift provision mayplastically deform (yield) as a result of a sling forceapplication up to and including 3.2 times the effectiveworking static load of the individual sling (3.2g).

Fft

Figure 2. DSS airlift configuration.

Two DSS configurations, referred to herein as the 1 0 1 stand NRDEC configurations, were evaluated to determine itconformity of the FO lift provision to MIL-STD-209G wasmaintained. Static suspension tests were performed tocharacterize sling force vectors required for Finite ElementAnalysis (FEA) of the provision. Strain gages were mountedat various locations to report the strain response of theprovision to static 1.0g loads. Experimental strain gageresults were correlated with the FEA model results and wereused to qualify the models for provision evaluation.

DESCRIPTION OF HHV AIRLIFT PROVISION

Since first introduced to military service, the HMMWVhas been upgraded and modified as specific applicationsrequire. In recent years, heavier payloads requiringincreased lift provision strength have resulted in designmodifications tc critical airlift components.

The Heavy HMMWV Variant (HHV) is an up-weighted HMMWVfor light tactical, utility, combat, and scoutingoperations. Among the many design improvements to this HMMWVversion are the forward airlift provisions. As a result ofincreased Gross Vehicle Weight (GVW), original airliftprovisions were rendered insufficient for certain air mobileunit operations. HHVs with their modified forward lift hooksare appropriate for heavy external air transport operationsand were used during all testing and analysis activities inthis investigation.

Lift Provision Construction

As illustrated in Figure 3, the HHV modified forwardairlift provision is constructed of a 0.75" diameterinverted "U" shaped rod welded at both ends to oppositecorners of a 3.00" X 4.00" X 20.375" main bracket box beam.A 0.25" thick reinforcement plate is welded to the straightsections of the rod to provide additional stiffness, therebydiminishing the effect of in-plane bending stresses withinthese sections. Adjacent to the main box beam are two end-beveled "C" channels. The channel beams are welded adjacentto each other and finally to the box beam. The three beamsections together are referred to as the bracket section ofthe provision and are mechanically fastened to the HHV frameby ten (10) 5/8" diameter bolts. All components of theassembly, (bracket and hook) are made of 1020 cold drawnsteel.

In comparison to the modified HHV forward liftprovision, the original consists of a 0.625" diameter rodand does not include the 0.25" thick reinforcement plate.

HOOK

A

R PEINFORCEMENT

PLATE

MAIN BRACKET

"C" CHANNELS BOX BEAM

0 0 0 75 0ap

50 00s

0 SeCTION A-A

Figure 3. Forward HHV lift provision.

Lift Hook Loading Design Criteria

An assessment of the in-plane and transverse normalbending stiffnesses, EIii and EItt respectively, for theHHV provisions was performed. Referring to Figure 3, 0 isdefined as the ratio of the in-plane to the transversenormal moments of inertia along section A-A for the in-planeand out-of-plane directions. The following value of 1 wascomputed for the forward HHV provision:

0 HHV = (Iii / Itt)HHV = 172.54 (1)

Since 1 is much greater than 1.0, the flexuralstiffness in the in-plane direction is much greater than thestiffness in the out-of-plane direction (0 for the standardHMMWV provision is 248.34). These results indicate that thehook was designed primarily for in-plane loads.

As a consequence of the position of the FO provisionrelative to the vertical configuration centerline, worstcase (highest combination of bending and axial loads)provision sling loading was anticipated for the outboardprovision. Sling loading vectors collinear with thelongitudinal axis of the main bracket box beam minimizebending stresses within the "U" shaped rod and are therefore

4

at desirable orientations. Slings loading the FO provisionsin DSS configurations are oriented at angles which tend tobend the hooks inwards towards the vehicle center. Inaddition, FO sling forces require higher magnitudes toachieve a vertical force component equal to the verticalcomponent of the forward inboard sling.

These findings are consistent with hook loadingprocedures as outlined in MIL-STD-209G. For the 10,000 lbGVW class HHV test vehicles used, loading of the hook shouldbe primarily in the in-plane direction and as close to thebox beam centerline as possible.

SIMULATED AIRLIFT TESTS

Simulated DSS airlift tests were performed tocharacterize the FO sling force vector (directions andmagnitude) and to record the strain response of theprovision to static 1.0g loads.

Simulated Airlift Test Procedure

DSS simulation tests were performed by suspending two10,000 lb HHVs from an "I" beam fixture supported by anoverhead crane (Figure 2). A 160 inch long "I" beam was usedto simulate the distance between the fore and aft liftpoints of the CH-47D. Shackles were fixed to both ends ofthe beam to support the forward and rear vehicle slings.

Prior to testing, both HHVs were positioned parallel toeach other, facing the same direction, and were separated byapproximately six (6) inches. Two 3 inch thick corrugatedcardboard honeycomb panels were placed between the HHVs todistribute vehicle contact forces and to protect the HHVside panels. Load cells were mounted in series between thebeam shackles and the lift slings to measure forwardoutboard and forward inboard sling forces throughout thetests.

Sling force and strain data acquisition systems w3retriggered to record prior to lifting. Ti2se sy-tcms remainedactive through the suspension period while anygemeasurements were recorded. Termination of data recordingoccurred only after the vehicles returned to the ground andthe lift slings supported no weight.

CHARACTERIZATION OF DSS AIRLIFT CONFIGURATIONS

Differences between the two DSS configurations areprimarily the result of the user preferred vehicle nose downangle, as illustrated in Figure 4. Vehicle symmetry, withrespect to the longitudinal axis ot the helicopter, is

- CH-47D FORWARD -,LIFT POINT

231.30"195.34"

18.0o0

15. 50,

6.77' 78~

85.00 85.00'

r 1'-~~- -1

j-~3 -- 300"

"r- 3.00"

A. NATICK RD&EC FORWARD 1. I01"t FORWARDSLING ANGIES SiL " ANGLES

Main BracketBox Beam BSMain Bracket

B em SLING SLING-

700' /8.00

-0-

. _0'30.001

C. NRDE( Sidei Angly.'s D. 101 't Side Angles

Figure 4. Front and side view of sling angles relative to vehiclesin both the 1 0 1 8t and NRDEC configu~'ations.

6

required for weight distribution (flight stability) and iscommon to both configurations. Nose down angles and vehiclesymmetry are both achieved by adjusting each sling/chain topredetermined lengths prior to attachment to the helicopter.For each configuration, the required sling/chain assemblylengths vary according to vehicle lift point location.Length adjustments are made secure by passing the slingchain through the hook and sliding the appropriate chainlink into the yoke of a link keeper (slotted shackle).

Results of Configuration Characterization

Results of angle measurements taken during testing areillustrated in Figure 5. Forward outboard sling angles weremeasured with an inclinometer placed collinearly along thesling, and were recorded relative to both the horizontal andvertical. Forward inboard angles were calculated based on FOsling angles and vehicle geometry. Together, thelongitudinal and transverse angles uniquely characterize theorientation of the forward outboard sling relative to acoordinate axes system originating at the sling/hook contactpoint.

Sling force data, as reported by the load cells, wereused in conjunction with angle measurements to resolve theFO sling force into equivalent sets of three mutuallyorthogonal components as shown in Figure 5 (see Appendix Ifor sling force resolution calculations). Tables 1 and 2below summarize the component results.

Table 1. DSS RESOLVED FORCES

(Forward Outboard Sling)

101ST NRDEC

FO Sling force (lb.) 2913 2435Fx (lb.) -617 97Fy (lb.) 2746 2330Fz (lb.) 751 700Angle front (deg.) 72.0 74.5Angle side* (deg.) 8.0 -7.0Nose down angle (deg.) 20.0 5.0

*(Side angle includes -12° provision mounting angle)

7

N+ -

00

4.4

- -~ ------

x 4

0

4-J

00

41

41

ot

44

=P

xO

Table 2. DSS RESOLVED FORCES(Forward Inboard Sling)

101ST NRDEC

FI Sling force (lb.) 1772 2363Fx (lb.) 301 199Fy (lb.) 1739 2329Fz (lb.) 162 345Angle front (deg.) 82.2 83.2Angle side* (deg.) 8.0 -7.0Nose down angle (deg.) 20.0 5.0

*(Side angle includes -120 provision mounting angle)

Expressing the FO sling force components in parametricform yields the following equations;

For 1.0g forces

1 0 1 st Config. F = - 617xi + 2746y• + 7 51ZkNRDEC Config. F = 97xi + 2330yj + 7 0 0 Zk

At 3.2g, the force vectors become

101st Config. F = - 1974xi + 8787y• + 2 4 0 3 zkNRDEC Config. F = 310xi + 7456yj + 2 2 4 0zk

STRAIN DATA ACQUISITION

A total of fifteen Micro-Measurements (TM) type EA-13-125EP-350 uniaxial strain gages were mounted to the hook andbracket sections of the forward outboard lift provision.Figure 6 illustrates the approximate locations of each gagerelative to the provision. Gages number 1 through 5 weremounted on the inner web surfaces of the two "C" channelbeams of the provision bracket. The remaining ten gages(gages No.s 6-15) were reserved for the hook and werepositioned to coincide with predetermined FEA modelintegration point locations.

A MEGADAC Series 2000 (TM) multi-channel dataacquisition system was used to record strain gage signals ata scan rate of 120 samples per second, per channel. Eachgage occupied an individual system channel with additionalchannels reporting elapsed and cumulative time.

9

1110 121_ F

15 13

9 6 F R 1

8 R

4

F = f ront

R = rear ( 180 degrees from front

Figure 6. Strain gage map.

System recording was triggered by means of the hostcomputer using OPUS (TM) software. A "zero load" test filewas recorded prior to lifting the vehicles to obtain straintransducer offset values after balancing and calibrating alldata acquisition system channels. These values were used tocompensate for the offsets of each strain gage on a channelby channel basis as necessary.

The tests performed are listed in Table 3 according tothe sling configurations tested and the order in which theyoccurred.

Table 3. STRAIN DATA TEST FILES

TEST #/FILE EXT. DESCRIPTION

1 / R01 Zero Values2 / R03 101st3 / R04 101st4 / R05 NRDEC

10

Strain Data Processing and Results

Strain gage results obtained from suspension tests wereindividually plotted (60 plots = 4 tests with 15 channelsper test) with microstrain as the ordinate (Y-axis), andelapsed time in minutes as the abscissa (X-axis). Theseplots represent the axial strain reported by each gage as afunction of time and show the strain response of the hookand provision through the lifting, static suspension, andunloading portions of each test.

To determine the average digital values of the plottedstrain, five samples of data were extracted from each gagefile. The first and last samples consisted of approximately100 data points each and correspond to pre and post vehiclelifting. These samples were used to determine residualstrains (6e) in the provision by comparing initial sampleaverage to final sample average strain values.

Sample numbers 2, 3, and 4 consisted of approximately500 points each and were taken at the beginning, middle andend of vehicle suspension respectively. Statistical analysisof these three samples were used to obtain the averagemaximum (tension) or minimum (compression) values of strainfor each gage in both test configurations. The "+/-" columncomputed for both configuration tests reports the average ofstandard deviations from samples 2, 3, and 4 of each straingage. Results of the statistical analysis are summarized inTable 4.

Table 4. STRAIN GAGE RESULTS.

Descrip. Zero Values 101st Config. NRDEC Config.

(4)(40 Ca (/E)

Gage # Ave. +/- Ave. +/- 6E Ave. +/- SE

1. -20 22 -137 3 -16 -105 21 42. -18 22 -109 3 -16 -80 21 43. -18 22 22 3 -20 10 23 -44. -19 22 -95 3 -7 -89 21 45. -18 22 -102 3 -7 -95 21 56. -18 22 922 5 -5 794 44 17. -18 22 522 5 -9 388 33 -238. -35 22 -772 11 1 -661 17 389. -23 22 151 3 -4 -91 21 3

10. -21 22 260 3 -12 319 31 -2211. -19 22 -64 3 3 -123 20 912. -21 22 883 3 -6 752 46 -4413. -19 22 113 3 -5 18 28 114. -14 22 -671 3 0 -593 17 3815. -17 22 312 3 0 281 30 -14

II

TRI I LIFT TESTIBIST CONF E248GAGE *6967•

ve 767

346

97. 2,41 4.9h 7.46 9.95 Y2.44TIME (MIN) •223

a i440i =t,%t10 1 Value

I Cnit CH) 498. 00086o2 sum 457818.2359603 "l.wn 917.4725424 SLEi ca.e. e4 awai) O.5664185 neladw% 537s865886a Vat- Is-VO96.9.7 StDww (act) 12.65282861 etA~a V . 949.411800

36 R.~geb5.23538811 Sunw~'"SL G.GeA86412 IC&&wtom Iu .1*26413a

Figure 7. Strain versus elapsed time plot for strain gage #6,test #2 (statistics of sample 1 3 included).

A sample plot of gage number 6, test file number 2 isshown in Figure 7 with the results of a statistical analysisof one of five samples (in this case, sample #3) taken perplot. Figure 8 shows the five locations at which the sampleswere taken. Statistical average values obtained from samplesnumber 2, 3, and 4 were again averaged to yield a singlerepresentative value of strain for each gage during thestatic suspension portion of each test.

CO ma×D Cma×xmaxD

m

S

t

r

n

Si sampl~ing location

(i = 1-5)

CO'O Cmax. OD

time CmInD

Figure 8. Typical sampling locations of strain plot.

Strain gage data from test file number 3 (correspondingto the 2nd 101st lift test) was not processed in the mannerdescribed above. This file was recorded after the datastorage media tape for the first two tests was removed fromthe test system and replaced by a second new tape. A brieflift test (test number 3, data storage tape number 2) wasperformed during which time the suspended vehicles swungback and forth in a pendulum type motion. The vehicles werereturned to the ground before the swinging motion subsidedand was therefore not a true suspension test. The purpose ofperforming this test was primari'y to ensure that the newtape was properly formatted and that the system wasfunctioning properly.

FINITE ELEMENT ANALYSES

In order to determine the response of the outboard HHVprovision to the loading cases shown in Figure 5, finite

J3

element analyses of both the hook and bracket wereperformed. Preliminary linear elastic analyses predicted theexistence of plasticity within the U-shaped hook componentof the outboard provision. Stresses obtained from thesemodels exceeded the material yield strength (a = 87,500psi) at 3.2 times the magnitude of the static loading vectorfor both the 1 0 1st and NRDEC load cases. These models areshown in Figure 9.

Modeling efforts were then focused on the developmentof a 3 dimensional, elastic-plastic finite elementrepresentation of the complete provision to detect possibleplastic deformation throughout the hook and bracketsections. The elastic-plastic model was generated by usingthe PATRAN 2 (TM) preprocessor and analyzed with the ABAQUS 3

(TM) commercial code. This model revealed that stresseswithin the bracket section of the provision were relativelylow in comparison to the hook component.

Linear Elastic Hook Finite Element Models

Hook models of the 1 0 1 st and NRDEC FO provision shown inFigure 9 were restricted to the portion of the hook locatedabove the reinforcement plate (refer to Figure 3). Sectic isbelow the reinforcement plate were not modeled since (1)stresses resulting from the applied loads were distributedlocally between the hook and the vertical welds of thereinforcement plate, (2) significant increases in axial,bending, and shear stiffnesses existed at the weld regionsand (3) stresses within the hook section located below thereinforcement plate were minimal.

Hook models were constructed of ABAQUStm Type B31 beamelements. These two noded, 3 dimensional isoparametricelements assume a linearized displacement field in {u,v,w,}with integration points (I.P.,finite points, the locationsof which report results) located as shown in Figure 10.

Boundary conditions fixing the translational androtational degrees of freedom (DOFs) were imposed on nodeslocated at the top of the welds connecting the hook to thereinforcement plate as shown in Figure 9. Justification forthese boundary assumptions were based on the significantincreases in axial, bending, and shear stiffnesses at thewelds. Loading vectors representing the outboard slingswere described for each DSS case by the parametric equationsindicated.

The vectors were applied to the models as concentratedforces to simulate the worst case condition as in previousinvestigations4 . Contact at the chain/hook interface wasmodeled as only one edge of one link, in contrast to thecombination of one edge and one face of two individual linksas shown in Figure 11.

14

x !!9-

51 NODE

NTE DATWTIN P.I IN.

OCA IT I ON- 290

133

E=95 0 A'

Figure 9. Linear elastic FEA mesh of provision hook.

NODE

INTEGRATION POINT2

LOCAT IONS

C T10Abaqus type 831 beam element

Figure 10. Integration point locations of beamelement.

F FslIng sl Ing

SELECTED LOADING ACTUAL LOADING

METHOD FOR F.E.A. CONTACT POINTS

MODELS

Figure 11. Sling/chain hook contact points.

Results of Linear-Elastic Hook Finite Element Models

Results for the 1 0 1st and NRDEC cases were obtainedusing a sling force load of 3.2 times the measured 1.0gsling force (indicated force vectors in Figure 5) and aredisplayed in Figures 12 and 13 for the 1 0 1 8t and NRDECconfigurations respectively. Axial strains, e , areplotted over the curved section of the hook as Alfunction ofthe angle 0, where 0 is defined ds the angle between theline connecting the element midpoint to the center of thehook's curved section. Each strain plot consists of fourcurves, one curve for each of the four integration points (#3,7,11,15) on the midpoint cross section of each element.

Correlation of Linear-Elastic FEA Model Results toExperimental Results

Strain gage results obtained from the simulated DSSairlift tests were multiplied by a factor of 3.2 andpositioned on the corresponding locations of the curves inFigures 12 and 13. A yield strain value of 2.966E-03 in/inobtained from previous material characterization tests was

16

J z

x ZU)

a: ( -4 a. a. M -

*4 P4 -u > Z

4j~

14-4 ( 1

.-4 U4

SU Wit: 14u 0iS L0

0 za 44 4

*$ -- 4

ZU Lu

W 13

*j .4 0- (U

o

14 w*~@ ~ .0X4LUU

> M. -4IV V

17(a.,

-J z

0~ .4

F. In C

T :-TI- I*Wr) 0 f-4

_________ 4j~4

0 S. 0

0* ~'0

to u

-u 0 4

LU 44C 0.,0

Lu4-) -- 4

'0 (i

IL -

-LU

4 4 z

Iz 040%-4 -*ý4

>1

aa RrII I

4~4 II'fu4 -

'-4

z U) I-- M -4 Z

used to indicate model yielding5 . This constant value wasplotted for both tension and compression in bothconfiguration cases from 0 = 0.00 to • = 1800. Model strainresults of a magnitude greater than E = +2.966E-03 or lessthan e = -2.966E-03 indicate material yielding at 3.2g slingforces.

Linear elastic FEA strain result plots indicate thatboth configurations yield below the required 3.2g slingforce limit. Strain gage data plotted on the curves inFigures 12 and 13 demonstrate the degree of correlationbetween experimental and F.E.A. results. These plotsindicate that the magnitude of strain within the hookexceeds acceptable limits.

Elastic Plastic Full Provision Finite Element Model

An elastic plastic finite element analysis of the fuliprovision (bracket and hook) was performed on the modelshown in Figure 14.

The model is constructed of ABAQUS (TM) Type B31 twonoded beam elements throughout the hook portion of the modelwith ABAQUS (TM) S4R Type four noded shell elementsthroughout the bracket section of the provision (withisotropic hardening, plastic strain data obtained fromprevious tests).

Attachment of the hook to the bracket was accomplishedby means of multiple point constraints between neighboringnodes of the hook and bracket interface. This methodprovides rigid beam attachments between the nodes of theý,ook and nodes of the bracket resulting in equivalentdisplacemdnts and rotations to approximate the weldinterface. Rigid body displacement of the model wasprevented by fixing translational and rotational degrees offreedom at the 10 node locations corresponding to thevehicle frame bolt holes.

Sling force loading of the model was performed with1.0g loads determined during testing and resolved in amanner similar to the linear elastic model case. Applicationof the force to the hook occurred at a single representativenode for each of the two cases (chain hook interfaceassumptions similar to the linear elastic model).

Elastic-Plastic Model Results

Results of the elastic-plastic FEA model wereevaluated, and correlated with experimental results, in amanner identical to the linear elastic model method.

.Mn

AMn

-. U1"m

Figure 14. Elastic-plastic tull provision FEA mesh.

20

~ I L HU O n

0 >-)},4 U)

(D 0 >(D 4.J C (

0 n-

t~ 419-4 Wn

(IV)0~

o *

0 00 •

% o4

(/) 4lU -) 0

o 0

-I I- 4I I Io

0 0L 0- 1w0

0 0 0 (0 C 0

I4 I ID

NIUdIS

21

nEnm~: t- -' 4 :

O >-

-4

0 0(00

014-4 -4 -,

0 44 4.)

l .,- r--4. 0 aJ :

U Z-4 Z 4J

101ST. 3.20. HOOK TIP.00330-

.00220- -.- LlfOl 101?

-o.-ELOU 1013

-.-*- EEflN 1015

-. 00110-

-. 00220-

-. 003300. 3.00 6.00 9.00 12.0 15.0 16.0

INTEGRATION POINT

Figure 17. Elastic-plastic 3.2g FEA results for forward portion ofhook in 1 0 1 8t configuration. Corresponding strain gageresults superimposed all at 3.2g.

NATICK. 3.2G. HOOK TIP.00288 -

.00160-

E--[LE•NT 1012.000960 -*"-ELEMENI 1013

*.-ELEMI NI 1014-*- ELEMENT 1015

0. -0 M11111111. Flo

-. 0009o -

-. 00190

-. 0028850. 3.00 6.00 9.00 12.0 IS.0 18.0

INTEGRATION POINT

Figure 18. Elastic-plastic 3.2g FEA results for forward portion ofhook in NRDEC configuration. Corresponding strain gageresults superimposed all at 3.2g.

23

Figures 15 and 16 shows the four curves generated bythe beam element integration points as a function of theelement number along the curved portion of the hook. Figure15, representing the 101st configuration shows a largenumber of element integration point strains whose absolutevalues exceed the 2.966E-03 in/in yield strain failurelimit. The NRDEC case (Figure 16), to a lesser degree thanthe 1 0 1st, shows failure at the forward hook/bracket weldinterface and chain/hook contact point.

Finite element analysis results of the forward straightsection of the "U" shaped rod were compared to the frontstrain gage band, consisting of gages no. 12 - 15. Figures17 and 18 show the correlation betweeo experimental and FEAresults for the 101st and NRDEC configurations respectively.For these plots, integration point identification numbers(3,7,11,15) are plotted along the abscissa versus strain.These beam element integration point locations were theintended locations of the forward strain gage band. Slightslippage occurred during gage bonding resulting in gagelocal-ion changes both circumferentially and longitudinally.

Contour plots of the bracket section of the provisionwere generated for both configurations. No attempt was madeto correlate these results with data obtained via straingages number 1 through 5. Photographs taken of the bracketsection lacked the necessary detail and clarity required toaccurately estimate the coordinates of the gages in thatsection. Results do however predict that failure by yieldingwould occur at bracket locations close to the frame boltholes at 3.2g loads in both configurations. Contour plotresults for both configurations are shown in Figures 16.

MODEL SENSITIVITY OBSERVATIONS AND ANALYSIS

Strain gage daLa files, processed prior to thestatistical sampling and averaging procedure, indicated adiscrepancy between the two 101st configuration lift tests.Graphical methods of strain value determination for thesetwo tests resulted in the values shown in Table 4.

Close examination of the data on a gage-by-gage basisrevealed an apparent trend in magnitudes of strain betweenthe two files. Repeatability of strain readings for mostout-of-plane sensing strain gages was relatively high forthe two 1 0 1 st configuration lift tests. This trend indicatesthat the sensitivity of the hook to small changes in out-of-plane loads (F.) is small. Repeatability of in-plane sensingstrain gages on the other hand was poor. This trend wouldindicate that the sensitivity of the hook to small changesin the vectorial sum of in-plane (Z Fx + Fy) forces is high

24

Table 4. STRAIN GAGE RESULTS FOR TWO TESTS OF THE 101stCONFIGURATION.

GAGE # R03 R04 161 locationME ju 6 E ip/op **

1. -134 -135 na2. -105 -102 na3. 20 20 na4. -95 -100 na5. -100 -100 na6. 900 910 1Q op7. 515 398 117 ip8. -767 -770 3 op9. -150 -81 69 ip10. 255 343 88 ip11. -64 -131 67 ip12. 880 930 50 op13. 112 65 47 ip14. -611 -670 59 op15. 315 390 75 ip

** lip" indicates gages sensing in-plane bendingstrains, "op" indicates gages sensingout-of-plane bending strains.

In light of these apparent trends, and the fact thatthe second 1 0 1st configuration lift test (R04) was notintended to yield high quality (stabilized vehicle motion)results, an FEA model sensitivity analysis was performed.

Table 5 lists, in a columnwise fashion, the integrationpoint strain results for element number 44 of the linearelastic 101st configuration FEA model. This model was loadedwith the appropriate 3.2g force, and executed five times(once for each of the five loaded node numbers identified).The total horizontal span between the extreme nodes, #16 and#20, is 0.7454 inches.

Percent variation calculations were then performed todetermine the percent change in axial strain between thehighest and lowest reported strain values for each of thefour integration points. The results of these calculationsare displayed in Table 6.

25

Table 5. FEA STRAIN RESULTS OF ELEMENT NUMBER 44 AS AFUNCTION OF LOADING NODE.

Loaded I.P. #3 I.P. #7 I.P. #11 I.P. #151Node # (AE) (A0 (A) (ME)

16 -1325 1645 1798 -1172

17 -1340 1823 1878 -1294

18 -1318 2052 1935 -1435

19 -1278 2201 1951 -1529

20 -1204 2377 1943 -1638

Table 6. PERCENT VARIATION OF FEA STRAIN RESULTS PERINTEGRATION POINT.

Integration Force Application Distance 6%Point Nodes (in)

3 17/20 0.565 11.3

7 20/16 0.745 44.5

11 19/16 0.554 8.5

15 20/16 0.745 39.8

In the worst case, integration point #7 results vary byas much as 44.5 % when adjusting the loading position of themodel by a corresponding 0.745 inches. Both integrationpoints number 7 and 15 report bending induced axial strainsfor the out-of-plane direction. Referring to Table 4,the differences of 50 and 59 microstrain corresponding toout-of-plane gages #12 and #14 (element #44 vicinity)respectively, could be attributable to inconsistentchain/hook contact positions between the two tests. To alesser degree, and as predicted by the model, in-planestrain discrepancies would also result from adjusting thecontact position of the sling/chain to the hook duringtesting.

26

CONCLUSIONS

Simulated DSS airlift tests were performed tocharacterize the forward outboard sling force vector forboth DSS configurations. Experimental 1.0g FO sling forceswere found to be 2,913 lbs for the 1 0 18t configuration and2,435 lbs for the NRDEC configuration. Sling force vectorresolution calculations were performed using two methods.Results of the calculations were used to resolve the slingforce into mutually orthogonal components relative to acoordinate axes system originating at the chain hook contactpoint. Results of component inspection indicated that theNRDEC configuration loads the provision more efficientlythan the 1 0 1st configuration. The excessive rearward in-plane load (FX = - 617 lbs) applied by the 1 0 1 stconfiguration FO sling, induces bending stresses, which whenextrapolated to 3.2g loads and superimposed with axialstresses, create unacceptable levels of stress as stipulatedby MIL-STD-209G.

Results of both FEA models predicted material yieldingbelow the required 3.2g elastic limit as stipulated inMIL-STD-209G. The 1 0 1 5t configuration was found to fail at2.62g by means of interpolation. A significant improvementover the 101st configuration failure load was observed withthe NRDEC configuration, which by interpolation indicatedfailure at 3.18g.

Due to discrepancies found between two 1 0 1 stconfiguration strain gage data files, a model sensitivitystudy was performed where sling load application points wereadjusted over the bounds of a 0.75" horizontal distance. Anelement close to the forward hook/bracket weld was observedas the loading node was moved through the 0.75" range.Results of the study were similar to the behavior of the two101st configuration data files which demonstrated that boththe models and the actual hook itself are complex structuresand are both extremely sensitive to loading points.

It was the intention of this investigation to determineif the two DSS configurations proposed for possible safetyrelease conformed to MIL-STD-209G. FEA models used todetermine material yielding correlated with experimentalstrain results at various locations to varying degrees. Inboth configurations, FEA models indicated that materialyielding would occur at sling force magnitudes lower thanthe 3.2g limit set in MIL-STD-209G. The results imply thatboth configurations do not meet the requirements ofMIL-STD-209G.

27

ACKNOWLEDGEMENTS

The authors would like to express their appreciation toRonald Aghababian and Jashiema Roach of AMTL for theirassistance in processing statistics and test data.

Special thanks to Andrew Mawn, Nancy Harrington and JohnDoucette of NRDEC for their assistance in planning,coordinating and participating in the simulated airlifttests.

28

BIBLIOGRAPHY

1. MIL-STD-209G; Military Standard Slinging TiedownProvisions for Lifting and Tying Down Military Equipment,p. A-14 para 5.1.1.1.

2. ABAQUS, Version 4.7, Hibbet, Karlson, and Sorenson,Inc., Providence, R.I.

3. PATRAN, Version 2.5, PDA Engineering, Costa Mesa, CA.

4. Beatty J. H. 1LT, Cavallaro P.V., Pasternak R.E., et al,Strain Analysis of HMMWV Front Lift Provision, MTL LetterReport.

5. Cavallaro C., Dooley R.B., Weight K.D., Cavallaro, P.,Testing and Analysis of Modified HMMWV Front Lift Provision,MTL Technical Report #92-31.

29

APPENDIZX I

Sling Force Resolution Calculations

30)

Method 1. Graphical determination of resolved forces.

KNOWN: 3. Lift sling suigi at*sat 40.0,0) and 9orni!ates itpoint A.

2. Magnitude of sling force

:i 2.913 pounds." 3. Angle dOe is 72

4. ADSIe bOc ist

1. Iatizate Oboas 100 units2. Angle blOq' Is 83. 02- 100 6o. a - 99.03

Ocl- 100 sin 4 - 13.324. Angle d'Oel 1 73

Od' sin 72 - 99.03Od'l 104.13Oe'- 104.13 Cos 72 - 32.11

2 2 2S. Ca' - SORT ( C' 4 0.2 4 ODa ) * 103.15

Of* * SQRT (Oc' 4 0.' ) - 35.06

4. By siailar triangles,

8a'Of' - Ca/Of

Of - (Os x Of')/Oa' , (2913 x 35.06) /105.15Of - 972.21"-1 "1

7. Angle fe'C tan (OC'/Oe') * tan (13.92/32.18)

"Asle f'Oe' - 23.39

Angle 1 'Of' * 90 - 23.39 6#.&2

•* Ce - OCf c€ 23.39 - 191.46 *Ot*e COc - Of coo 4#.61 - 385.59 ***

5. 41'/oe * og'/o0

Og - (Og' x Oal/O.' -(99.03 z 2913)/105.25

*** 0g 2746 **O

9:. ecall Of - 971.26-1

angle fOc tan (0*/Cc) - #4.41 degreesangle fOo - (90 - #4.41) - 23.39

10. Project coordinate of point f to axis rotated24 degrees clockwise relative to axis 0g.

971.28 coo (64.41 - 14) a 417 (negative Oc direction)971.26 coo (23.39 * 14) = 751 (positive *a direction)

11. The three resolved couponento of Os are:

O0 a 27440O (16 dog. rot.) - 751Oc (1 dog. rot.) - 417

31

Axes Transformation Calculations

Input: Resolved forces in vehicle coordinate system.Results: Resolved forces in hook coordinate symtwm.'*Refer to pages 7,8,31.

Rotate about x-axis by (+ex):

X, - xi a xi'

Y, -ycogBxl +Ysinext

z' - -zsiGjl*j + ZCOSOx~

F 1 0 0 r S

Rotate about y'-axisby(e

x, xcosS ,i1 - x'sine ,h'

y y*5*Y'.

[x cosey1 0 -siney FxI n.tal*I[ ::.1 0 1 0 jF ney . 0Foey z'

Rotate about Z'-axis by (+e01 5 9 h

x"a i~cose ,,i" + x's&inO2 5 .j"- -Y"sine~P 2.8 +yOOCos8z.j"

a Zoe ~*' a k"

Fx' ~ose *, sinOz1 F6

0y -sinta, IIe 0IP

32

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