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June 2002
NASA/TM-2002-211733ARL-TR-2735
Occupant Responses in a Full-Scale Crash Testof the Sikorsky ACAP Helicopter
Karen E. Jackson, Edwin L. Fasanella, and Richard L. BoitnottU.S. Army Research LaboratoryVehicle Technology DirectorateLangley Research Center, Hampton, Virginia
Joseph McEntire and Alan LewisU.S. Army Aeromedical Research LaboratoryFort Rucker, Alabama
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June 2002
NASA/TM-2002-211733ARL-TR-2735
Occupant Responses in a Full-Scale Crash Testof the Sikorsky ACAP Helicopter
Karen E. Jackson, Edwin L. Fasanella, and Richard L. BoitnottU.S. Army Research LaboratoryVehicle Technology DirectorateLangley Research Center, Hampton, Virginia
Joseph McEntire and Alan LewisU.S. Army Aeromedical Research LaboratoryFort Rucker, Alabama
Available from:
NASA Center for AeroSpace Information (CASI) National Technical Information Service (NTIS)7121 Standard Drive 5285 Port Royal RoadHanover, MD 21076-1320 Springfield, VA 22161-2171(301) 621-0390 (703) 605-6000
The use of trademarks or names of manufacturers in the report is for accurate reporting and does not constitute anofficial endorsement, either expressed or implied, of such products or manufacturers by the National Aeronauticsand Space Administration or the U.S. Army.
1
Occupant Responses in a Full-Scale Crash Test of the Sikorsky ACAP Helicopter
Karen E. Jackson, Edwin L. Fasanella, and Richard [email protected], [email protected], [email protected]
US Army Research Laboratory, Vehicle Technology DirectorateHampton, Virginia
Joseph McEntire and Alan [email protected], [email protected]
US Army Aeromedical Research LaboratoryFt. Rucker, Alabama
Abstract
A full-scale crash test of the Sikorsky Advanced Composite Airframe Program (ACAP) helicopterwas performed in 1999 to generate experimental data for correlation with a crash simulation developedusing an explicit nonlinear, transient dynamic finite element code. The airframe was the residual flighttest hardware from the ACAP program. For the test, the aircraft was outfitted with two crew and two troopseats, and four anthropomorphic test dummies. While the results of the impact test and crash simulationhave been documented fairly extensively in the literature, the focus of this paper is to present the detailedoccupant response data obtained from the crash test and to correlate the results with injury predictionmodels. These injury models include the Dynamic Response Index (DRI), the Head Injury Criteria (HIC),the spinal load requirement defined in FAR Part 27.562(c), and a comparison of the duration and magni-tude of the occupant vertical acceleration responses with the Eiband whole-body acceleration tolerancecurve.
Introduction
In 1999, a full-scale crash test of a pro-totype composite helicopter was performed atthe Impact Dynamics Research Facility (IDRF)that is located at NASA Langley Research Cen-ter in Hampton, VA. The IDRF is a 240-ft. highgantry structure that has been used for con-ducting full-scale crash tests of light aircraft androtorcraft since the early 1970's [1]. The heli-copter was the flight test article built by SikorskyAircraft under sponsorship by the U.S. Armyduring the Advanced Composite Airframe Pro-gram (ACAP). The purpose of the ACAP was todemonstrate the potential of advanced compos-ite materials to save weight and cost in airframestructures while achieving systems compatibilityand meeting Army requirements for vulnerability,reliability, maintainability, and survivability. In1981, the US Army awarded separate contractsto Bell Helicopter Textron and Sikorsky AircraftCorporation to design and fabricate helicoptersconstructed primarily of advanced compositematerials. Each company manufactured threeairframes, and one helicopter airframe fromeach company was equipped to become a flyingprototype. Crash tests of the Bell and Sikorsky
ACAP static test articles were conducted in 1987at the NASA IDRF by the US Army to demon-strate their crash performance [2, 3].
In 1997, the US Army Aviation AppliedTechnology Directorate (AATD) established aScience and Technology Objective (STO) incrash modeling and simulation. The Army Re-search Laboratory's Vehicle Technology Direc-torate (ARL-VTD) was selected by AATD as theprimary performing organization for the STO.The purpose of the STO was to establish astandardized and validated structural crash dy-namics modeling and simulation capability froma commercial computer code that would satisfythe need for a crashworthy performance anddesign evaluation tool. As part of the STO, afull-scale crash test of the Sikorsky ACAP resid-ual flight test article was planned to generateexperimental data for correlation with the crashsimulation. In 1998, AATD cancelled the STO;however, the original plans for the crash test andmodel validation were continued under the sup-port of the NASA Aviation Safety Program [4].For the test, the aircraft was outfitted with twocrew and two troop seats, and four anthropo-morphic test dummies. The crash simulation
2
was performed using the nonlinear, explicit tran-sient dynamic code MSC.Dytran [5], and theresults of the validation study have been pub-lished in References 6 through 9. Experimentalresults from the crash test have been reported inReference 10 including the crash sequence ofevents, an assessment of structural deformationand fuselage damage, and the dynamic re-sponse of the airframe and large mass itemssuch as the engines and rotor transmission.The seat damage was described fairly exten-sively; however, only a limited amount of occu-pant response data was reported.
The objectives of this paper are to pre-sent the occupant test data obtained during thefull-scale crash test of the Sikorsky ACAP heli-copter and to correlate the results with injuryprediction models. These models include theDynamic Response Index (DRI), the Head InjuryCriteria (HIC), the spinal load requirement de-fined in FAR Part 27.562 (c), and a comparisonof occupant vertical acceleration responses withthe Eiband whole-body acceleration tolerancecurve.
Full-Scale Crash Test
Pre- and post-test photographs of theSikorsky ACAP helicopter are shown in Figure 1.The planned test conditions were 38-ft/s verticaland 32.5-ft/s horizontal velocity, with 5° nose-uppitch and no yaw or roll angle. The measuredimpact conditions were 38-ft/s vertical and 32.5-ft/s horizontal velocity, with 6.25° nose-up pitchand 3.5° left-down roll. Also, a 9.6°/secondnose-up pitch angular velocity was induced as aresult of the pendulum-swing drop test proce-dure [10, 11]. During the crash test, over eightychannels of data were acquired at 10,000 sam-ples per second using a 12-bit resolution digitaldata acquisition system (DAS).
Right and Left Crew Dummies and Seats
A 50th percentile male Hybrid II anthro-pomorphic test dummy, weighing approximately170-lbs., was used to represent the pilot in theright front crew position. The dummy was in-strumented with tri-axial accelerometers locatedin the head, chest, and pelvis. A lumbar loadcell was installed to measure the spinal com-pressive force response. Additional load cellswere used to measure the force in the lap andshoulder belts. The pilot was placed in a used
commercial military-qualified helicopter seatprovided by the US Army Aeromedical ResearchLaboratory (USAARL). This seat contained twoinvertube energy absorbers. New MA-16 inertialocking reels were used with lap and shoulderbelts to restrain the dummy occupant during thetest. Accelerometers were mounted to the seatpan to measure forward and vertical accelera-tion responses. Tri-axial accelerometers werelocated on the floor near the pilot seat attach-ment to the outboard seat rail.
(a) Pre-test photograph.
(b) Post-test photograph.
Figure 1. Pre- and post-test photographs of theSikorsky ACAP helicopter.
The USAARL supplied a fully instru-mented modified Hybrid III 50th percentiledummy with a self-contained DAS for the leftfront crew (copilot). This dummy weighed ap-proximately 198-lbs. and was the only dummy tobe outfitted with a helmet for the test. Themodifications to the dummy included the incor-poration of an EME Corp. internal data acquisi-tion system. In addition, the rigid spine box ofthe Hybrid III was replaced with a flexible spineconsisting of rubber disc segments betweeneach rib as well as two torsional joints. Thehead of the Hybrid III dummy was replaced witha Hybrid II head because of the more represen-tative anthropometry features and helmet fitcompatibility.
Data collected for the copilot dummyusing the self-contained DAS included head ac-celerations in three directions; T-1 thoracic ac-celerations in three directions; head pitch rate;C-1 cervical head/neck forces and moments; T-1thoracic forces and moments; torso sternum ac-celerations in two directions; and lumbar forcesand moments. The T-1 thoracic accelerometerswere located at the base of the neck, and thetorso sternum accelerometers were located inthe front of the chest. In all, 29 channels of data
3
were collected at 10,000 samples/second usingthe self-contained DAS, including three addi-tional accelerometers to record tri-axial floor-level accelerations. These accelerometers werelocated slightly behind and between the crewseats.
The copilot dummy was secured in aused commercial military-qualified helicopterseat of a different design than the pilot seat.This seat contained six "torshock" energy ab-sorbers. Likewise, an MA-16 inertia locking reelwas used in conjunction with the restraint sys-tem to limit the displacement of the dummy oc-cupant during the test. Two accelerometerswere mounted to the seat pan to measure for-ward and vertical accelerations, and tri-axial ac-celerometers were located on the floor near thecopilot seat attachment to the outboard seat rail.Load cells were installed to measure lap andshoulder belt forces. Because the fore-aft ad-justing pins for the crew seats used in this testwere located at the front leg, holes were drilledin the front rail at the average longitudinal loca-tion. The seats were also adjusted to the middlevertical positions to allow a maximum verticalstroke of approximately 14.5 inches. The seatpan would contact the flexible seat well floor at astroke of 13.5 inches.
Following the test, measurements weremade indicating that the pilot seat stroked ap-proximately 9 inches of the total 14.5 inchesavailable. It was determined that the outboardpin attaching the pilot seat to the seat rail waseither not properly engaged in the seat rail holeor disengaged during the impact. Without theoutboard pin restraint, the seat rotated inwardabout the remaining inboard pin and the seatpan struck the front seat well frame instead ofstroking down into the well. A post-test photo-graph of the pilot dummy is shown in Figure2(a). The final position of the pilot seat is shownin Figure 2(b). The pin moved approximately 4inches forward of its original location.
A post-test photograph of the copilotdummy and left crew seat is shown in Figure 3.During post-test examination of the copilot, afragment of cloth was found on the right side ofthe dummy's forehead indicating contact with theknee. This discovery provides evidence that thenatural slack in the restraint system allowed thedummy’s head to rotate over and eventuallystrike its knee. The MA-16 inertia reels were not
manually locked before the test. After the testboth inertia reels were found to be locked. Thecopilot seat essentially "bottomed out" with ap-proximately 14.5 inches of stroke. Post-test vis-ual inspection of the seat indicated that the seatpan accelerometer punched through the flexibleseat well floor.
(a) Pilot dummy.
(b) Forward motion of the unpinned pilot seat.
Figure 2. Post-test photographs of the pilot(right crew) and seat.
The filtered vertical acceleration time-histories of the pilot floor, seat pan, and chestare given in Figure 4(a). The data were filteredusing a 2-pole Butterworth low-pass digital filterwith a cut-off frequency of 60-Hz. The datawere filtered both forward and backward in timeto eliminate any phase shift. The same low-pass digital filter was used to filter the test datashown in all of the plots in this paper, exceptwhere noted. The pilot's energy absorbing seatreduced the vertical acceleration peak from 93-gon the floor to approximately 40-g at the seatpan. The chest acceleration peak is slightlylower at 36-g and is delayed in time by .01 sec-onds. Likewise, the filtered vertical accelerationtime-histories for the copilot floor, seat pan, and
Pin and latchingmechanism
Pin hole Tri-axialaccelerometer
4
chest (torso sternum) are plotted in Figure 4(b).The copilot’s energy absorbing seat reduced thevertical acceleration peak from 85.6-g on thefloor to approximately 33-g at the seat pan. Thechest acceleration peak is of the same magni-tude as the seat pan response, but is delayed intime. Thus, both the pilot and copilot seats re-duced the floor-level acceleration peak by 53-g.
Figure 3. Post-test photograph of the copilot (leftcrew) dummy.
The crew seat pan accelerations arehigher than desired for occupant survivability. Afactor that likely contributed to the high seat panacceleration responses was the fact that theoriginal energy absorbing nose gear designedfor the Sikorsky ACAP helicopter was not avail-able for this test. Instead, a standard non-crashworthy commercial nose gear was retrofit-ted to provide a nominal level of energy absorp-tion. However, the retrofitted nose gear couldonly absorb a small percentage of the kineticenergy that the original gear was designed todissipate.
Right and Left Troop Dummies and Seats
Ceiling-suspended troop seats, eachwith two wire-bender energy absorbers, weremounted in the rear cabin area. New wire-bender energy absorbers were installed in thetroop seats; however, these seats were usedand the seating material was in poor condition.The seat pan consisted of a nylon mesh cloththat was worn and oil-stained. Two Hybrid II50th percentile anthropomorphic dummies, eachweighing approximately 170-lbs., were used torepresent the right and left troop occupants.Both dummies were instrumented with tri-axialaccelerometers located in the chest. One accel-
erometer was mounted to the rear frame of theseat, and two accelerometers were attached tothe floor near the troop seats.
-20
0
20
40
60
80
100
0 0.05 0.1 0.15 0.2
FloorSeat panChest
Acceleration, g
Time, s(a) Pilot acceleration responses.
-20
0
20
40
60
80
100
0 0.05 0.1 0.15 0.2
FloorSeat panChest
Acceleration, g
Time, s(b) Copilot acceleration responses.
Figure 4. Filtered vertical accelerations of thecrew floor, seat pan, and pelvis.
Post-test photographs of the right andleft troop occupants are shown in Figure 5. Theseat pan cloth was torn in both troop seats, al-lowing the buttocks of both dummy occupants todisplace downward through the seat frame. Asa result, the wire-bender energy absorbers ex-hibited minimal stroking. The inboard wire-bender of the left troop seat stroked approxi-mately 1 inch while the displacements of theother wire-benders were considerably less.During impact, a large downward deflection ofthe helicopter's roof at the wire-bender suspen-
5
sion location was observed in the high-speedfilm coverage. This roof deflection may havelimited the stroking of the wire-bender mecha-nism, as well.
(a) Right troop.
(b) Left troop.
Figure 5. Post-test photographs of the right andleft troop occupants.
Seat, Occupant, Restraint System, and Struc-tural Responses
The unfiltered head acceleration timehistories of the pilot and copilot dummies areplotted in Figure 6 in the forward, side, and ver-tical directions. Only the copilot forward headresponse is shown in Figure 6(a) due to anoma-lies in the pilot test data. The acceleration re-sponses indicate that the copilot experienced ahigh-magnitude, short-duration spike at 0.188seconds during the pulse, as a result of headstrike. High-speed film coverage shows that at0.188 seconds the copilot's head comes in con-tact with his knees. A head strike is further indi-cated by a bit of green flight suit fabric that was
found embedded in the forehead of the dummyslightly above the right eyebrow. The dummy'shead most likely contacted the right knee. Thefilm shows the head still moving downward be-tween the knees for an additional .017 secondsbefore rebounding upward. The spike at 0.188seconds is observed in all three copilot headacceleration responses with peaks of 450-, 190-,and 42-g's in the forward, side, and vertical di-rections, respectively.
The filtered T-1 thoracic acceleration re-sponses of the copilot dummy are shown in Fig-ure 7 for the forward, side, and vertical direc-tions. The pilot dummy was not instrumented atthis location; consequently, no comparison withthe copilot test data is possible. The effect of thehead strike experienced by the copilot dummy isevident in the large 48-g peak in the forward ac-celeration response shown in Figure 7. In com-parison, the vertical acceleration peak is 28-g.
Plots of filtered vertical chest accel-eration responses of the pilot and copilot, andthe right and left troop dummies are shown inFigure 8. In general, the pilot and copilot dum-mies experienced higher peak accelerationsthan did the troop occupants. Plots of filteredchest acceleration responses of the pilot andcopilot, and the right and left troop dummies areshown in Figure 9 in the forward direction.Again, the pilot and copilot experienced signifi-cantly higher peak acceleration responses thandid the troop dummies.
The filtered acceleration responses ofthe pilot pelvis are shown in Figure 10 for theforward, side, and vertical directions. The copi-lot dummy was not instrumented with pelvic ac-celerometers; consequently, no comparison withthe pilot test data is possible. As expected, thepeak vertical acceleration (40-g) is significantlyhigher than either the forward (22-g) or side (13-g) peak accelerations.
The filtered vertical seat pan accelera-tion responses of the pilot, copilot, and troopseats are shown in Figure 11. It should benoted that the data for the troop seats was ob-tained from accelerometers placed on the rearseat frame, since the troop seats contained acloth seat pan. The acceleration pulse shapesare similar for the pilot and copilot seat pans andfor the right and left troop seat frames, respec-tively.
6
-500
-400
-300
-200
-100
0
100
0 0.05 0.1 0.15 0.2 0.25
Copilot head
Acceleration, g
Time, s
(a) Forward acceleration.
-200
-160
-120
-80
-40
0
40
0 0.05 0.1 0.15 0.2 0.25
Pilot headCopilot head
Acceleration, g
Time, s
(b) Side acceleration.
-60
-40
-20
0
20
40
60
0 0.05 0.1 0.15 0.2 0.25
Pilot headCopilot head
Acceleration, g
Time, s(c) Vertical acceleration.
Figure 6. Unfiltered pilot and copilot head accel-eration responses in three directions.
-10
0
10
20
30
40
50
0 0.05 0.1 0.15 0.2 0.25
ForwardSideVertical
Acceleration, g
Time, s
Figure 7. Filtered copilot T-1 thoracic accelera-tion responses in three directions.
-10
0
10
20
30
40
0 0.05 0.1 0.15 0.2 0.25
Pilot chestCopilot chest
Acceleration, g
Time, s
(a) Pilot and copilot.
-10
0
10
20
30
40
0 0.05 0.1 0.15 0.2 0.25
Right troop chestLeft troop chest
Acceleration, g
Time, s(b) Right and left troop.
Figure 8. Filtered vertical chest acceleration re-sponses of four dummies.
7
-10
0
10
20
30
40
50
0 0.05 0.1 0.15 0.2 0.25
Pilot chestCopilot chest
Acceleration, g
Time, s
(a) Pilot and copilot chest.
-10
0
10
20
30
40
50
0 0.05 0.1 0.15 0.2 0.25
Right troop chestLeft troop chest
Acceleration, g
Time, s
(b) Right and left troop chest.
Figure 9. Filtered forward chest acceleration re-sponses of four dummies.
-10
0
10
20
30
40
50
0 0.05 0.1 0.15 0.2 0.25
ForwardSideVertical
Acceleration, g
Time, s
Figure 10. Pilot pelvis acceleration responses.
The restraint systems of both the pilotand copilot dummies were instrumented tomeasure lap and shoulder belt loads. Theseresponses are plotted in Figure 12. Note thatonly the pilot shoulder belt response is shown inFigure 12 due to anomalies in the copilot testdata. The FAR 27.562 (c) specifies that whereupper torso straps are used for crew members,tension loads in individual straps must not ex-
ceed 1,750 pounds, and, if dual straps are usedfor restraining the upper torso, the total straptension loads must not exceed 2,000 pounds[12]. The restraint loads measured in the crewdummies did not exceed these limits. The trooprestraint systems were not instrumented.
-20
-10
0
10
20
30
40
50
60
0 0.05 0.1 0.15 0.2 0.25 0.3
PilotCopilot
Acceleration, g
Time, s
(a) Pilot and copilot.
-20
-10
0
10
20
30
40
50
60
0 0.05 0.1 0.15 0.2 0.25 0.3
Right troopLeft troop
Acceleration, g
Time, s
(b) Right and left troop.
Figure 11. Filtered seat pan vertical accelerationresponses.
Accelerometers were installed to recordfloor-level acceleration responses in three direc-tions using the self-contained DAS in the copilotdummy supplied by USAARL. In addition, floor-level acceleration responses were measurednear the outboard seat attachment points for thepilot and copilot using the NASA data acquisitionsystem. The filtered acceleration responses forthese three locations are shown in Figure 13.The similarity in the floor-level acceleration re-sponses is surprising given that the acceler-ometers were not mounted in exactly the samelocation on the floor and the signals were re-corded on two different data acquisition sys-tems. These results provide confidence in thequality and validity of the experimental data.
8
-200
0
200
400
600
800
1000
0 0.1 0.2 0.3 0.4
Pilot lap beltPilot shoulder beltCopilot lap belt
Force, lbs.
Time, s
Figure 12. Filtered restraint system loads for thepilot and copilot.
Injury Prediction
Dynamic Response Index (DRI)
One commonly used injury predictionmodel is the Dynamic Response Index (DRI)[13]. The DRI is derived from a simple one-dimensional lumped-mass spring damper sys-tem, as depicted in Figure 14. This model wasdeveloped by the Air Force's Wright Laboratoryto estimate the probability of compression frac-tures in the lower spine due to acceleration in apelvis-to-head direction, as might be experi-enced by aircrew during ejection seat opera-tions. Unfiltered vertical acceleration responsesof the seat pan of the pilot and copilot, and theright and left troop were used as input to com-pute the dynamic DRI.
The continuous DRI time histories foreach occupant are shown in Figure 15. Themaximum DRI values for the crew and troopdummies are noted in the legend descriptions inFigure 15 and they range from 22.3 to 30. Op-erational data from actual ejection seat incidentsindicate that the spinal injury rate for maximumDRI values between 20 and 23 range from 16 to50 percent, see References 14 and 15.
Occupant acceleration data are com-pared with the continuous DRI responses in Fig-ure 16. For the pilot, both the chest and pelvisvertical acceleration responses are comparedwith the DRI in Figure 16(a). For the copilot, thevertical acceleration response of the chest (torsosternum) is compared with the DRI in Figure16(b). For the troop dummies, the vertical chest
acceleration responses are compared with theDRI in Figures 16(c) and (d).
-30
-20
-10
0
10
20
0 0.05 0.1 0.15 0.2
Pilot (NASA)Copilot (NASA)Copilot (USAARL)
Acceleration, g
Time, s
(a) Forward acceleration responses.
-10
-5
0
5
10
15
0 0.05 0.1 0.15 0.2
Pilot (NASA)Copilot (NASA)Copilot (USAARL)
Acceleration, g
Time, s
(b) Side acceleration responses.
-40
-20
0
20
40
60
80
100
0 0.05 0.1 0.15 0.2
Pilot (NASA)Copilot (NASA)Copilot (USAARL)
Acceleration, g
Time, s
(c) Vertical acceleration responses.
Figure 13. Filtered acceleration responses of thepilot and copilot seat floor as measured by theNASA and USAARL data acquisition systems.
9
Mass
Dampingratio =0.224
a(t)
Spring natural frequency = 52.9 rad/s
Figure 14. Schematic of the DRI injury model.
-20
-10
0
10
20
30
40
0 0.1 0.2 0.3 0.4 0.5 0.6
Pilot (DRI=28.6)Copilot (DRI=22.3)
DRI
Time, s
(a) Pilot and copilot.
-20
-10
0
10
20
30
40
0 0.1 0.2 0.3 0.4 0.5 0.6
Right troop (DRI=28.6)Left troop (DRI=30)
DRI
Time, s
(b) Right and left troop.
Figure 15. Continuous DRI responses of the fourdummy occupants.
-20
-10
0
10
20
30
40
50
0 0.05 0.1 0.15 0.2 0.25
Pilot chestPilot pelvisDRI
Acceleration, g
Time, s
(a) Pilot chest and pelvis data with DRI.
-20
-10
0
10
20
30
40
50
0 0.05 0.1 0.15 0.2 0.25
Copilot chestDRI
Acceleration, g
Time, s
(b) Copilot chest data with DRI.
-20
-10
0
10
20
30
40
50
0 0.05 0.1 0.15 0.2 0.25
Right troop DRI
Acceleration, g
Time, s
(c) Right troop chest data with DRI.
-20
-10
0
10
20
30
40
50
0 0.05 0.1 0.15 0.2 0.25
Left troopDRI
Acceleration, g
Time, s
(d) Left troop chest data with DRI.
Figure 16. Comparison of occupant responseswith the continuous DRI.
In general, the continuous DRI modelunder predicts the peak accelerations of the pilotand copilot responses, and over predicts thepeak accelerations of the troop responses.Also, the time of occurrence of the peak accel-eration is delayed for the DRI model in compari-son with the test data. The continuous DRI data
10
for the troop occupants must be viewed withcaution, since the cloth seat pan was torn duringthe test, allowing the buttocks of the troop dum-mies to displace through the seat frame. Conse-quently, the vertical response of the seat frameis not likely to be a good indicator of occupantresponse or injury potential for the troop dum-mies.
The one-dimensional DRI model hasobvious limitations for application to impact sce-narios involving multi-directional accelerationcomponents. A more comprehensive methodwas developed to account for acceleration com-ponents in the three orthogonal axes on the hu-man occupant [16]. A FORTRAN programDYNRESP [17], obtained from NASA JohnsonSpace Center, was used to calculate the dy-namic response and injury risk assessment of aseated occupant by analyzing the measured x,y, and z linear accelerations of the seat. Thesedirections are defined in Figure 17. The dy-namic response of the occupant is modeled by amass, spring, and damper system attached tothe seat. Each orthogonal axis is modeled witha different spring-damper representation. Thegeneral risk of injury is determined based on thecombined dynamic responses of the three axesand the defined limits in these directions usingEquation 1:
[ ]1/2
β = ( DRXDRXL
)2
( DRYDRYL
)2
+ +( DRZDRZL
)2
(1)
where DRX, DRY, and DRZ are the dynamicresponses for the x-, y-, and z-axes; DRXL,DRYL, and DRZL are the limit values defined forlow, moderate, and high risk, and β is the injury-risk criterion. Different dynamic response limitvalues, listed in Table 1, are used for low, mod-erate, and high risk [16]. Risk levels for injuryare considered satisfactory if β is less than 1.
This model was applied by inputing theforward (x) and vertical (z) components of ac-celeration obtained from the pilot and copilotseats into the dynamic response model de-scribed by Equation 1. No side component ofacceleration was input since it was not meas-ured in the test. Data from other locations indi-cate that the side accelerations were minimaland, therefore, the omission of this componentshould not significantly affect these computa-tions. The results of this injury risk assessmentare shown in Figure 18. For the pilot, all three
of the risk assessment curves exceed thethreshold value of 1.0, indicating a high risk ofinjury. For the copilot, the high-risk curve isclose to 1.0, but does not exceed the threshold.Thus, the results indicate a moderate risk of in-jury for the copilot. s
Ys
Yaw
Pitch
Roll
Figure 17. Axis system used to calculate thecombined DRI response.
Table 1. Dynamic Response Limit Values forLow, Moderate, and High Risk
DRXL DRYL DRZL
DRX>0 DRX<0 Conventional Siderestraint panels
DRZ>0 DRZ<0
Lowrisk
35 28 14 15 15.2 9
Moderaterisk
40 35 17 20 18 12
High risk 46 46 22 30 22.8 15
Spinal Force
A second injury assessment criteria,defined in FAR Part 27.562 (c), is that spinalload should not exceed 1,500 lbs [12]. Both thepilot and copilot dummies were instrumentedwith lumbar load cells to measure force alongthe spine. A plot of lumbar load versus time isshown in Figure 19 for the pilot and copilotdummies. Initially the load is tensile perhapsfrom the forward motion of the dummy as thehelicopter slows down horizontally. The maxi-mum compressive loads measured for the pilotand copilot dummies are 1,912 and 1,921 lbs.,respectively. These loads occur during fuselagefloor impact and about .005 seconds after thepeak pelvis vertical acceleration. They exceedthe 1,500-lb. threshold for spinal injury which isthe maximum load for civil seat certification [12].
y
x
z
11
-0.5
0
0.5
1
1.5
2
0 0.05 0.1 0.15 0.2 0.25 0.3
Low riskModerate riskHigh risk
β
Time, s
(a) Pilot.
-0.5
0
0.5
1
1.5
2
0 0.05 0.1 0.15 0.2 0.25 0.3
Low riskModerate riskHigh risk
β
Time, s (b) Copilot.
Figure 18. Injury risk assessment for the pilotand copilot dummies.
-2000
-1500
-1000
-500
0
500
1000
0 0.05 0.1 0.15 0.2 0.25
PilotCopilot
Lumbar load, lbs.
Time, s
Figure 19. Filtered pilot and copilot verticallumbar load responses.
Head Injury Criteria (HIC)
The Federal Motor Vehicle Safety Stan-dard 208 (FMVSS 208 [18]) includes a head im-pact tolerance specification called the Head In-jury Criteria (HIC) [19]. The HIC was originallydeveloped as a modification of the Wayne StateUniversity Tolerance Curve [20] and is calcu-lated by the following equation:
(t2 - t1){ }[ t2 - t1
1 ∫t1
t2
a(t) dt]2.5
HIC =
(2)where t1 is the initial time of integration, t2 is thefinal time of integration, a(t) is the resultant ac-celeration in g's measured at the center-of-gravity of the head. The FMVSS 208 estab-lishes a maximum value of 1000 for the HIC,which is associated with a 16 percent risk of se-rious brain injury. It also specifies that the timeinterval, i.e. t2 minus t1, used in the integrationshould not exceed .036 seconds. Limitations inusing the HIC have been documented in the lit-erature [21], including a study that found that thecritical time duration used in the HIC calculationshould be equal to or less than .015 seconds[19]. For this evaluation, both time intervals areused in the HIC calculation.
The resultant head acceleration re-sponse was calculated for the copilot dummyoccupant, as plotted in Figure 20. For the pilot,only the vertical and side components of accel-eration were available, since the forward accel-eration channel was lost after 0.15 seconds.Consequently, no HIC calculation was attemptedfor the pilot. As shown in Figure 20, the copilotexperiences a peak acceleration of almost 500-gat 0.188 seconds. The duration of this accel-eration spike is approximately .003 seconds.The data shown in Figure 20 were used to cal-culate HIC values for time intervals of 0.036 and0.015 seconds. The initial time, t1, used in theHIC calculations was varied systematically fromthe beginning of the pulse to determine themaximum value of HIC.
The results of the HIC analysis, shownin Table 2, indicate that the copilot experienceda head impact during the crash test. The copilotHIC values were 713 and 1185 for time intervalsof 0.036- and 0.015-seconds, respectively.These results also confirm the need to evaluateHIC for different time intervals. In this case, the
12
higher value of HIC (greater than 1000) wasfound using the shorter (0.015 second) time in-terval.
0
100
200
300
400
500
0 0.05 0.1 0.15 0.2 0.25 0.3
Resultant acceleration, g
Time, s
Figure 20. Resultant acceleration responses ofthe pilot and copilot head.
Whole-Body Acceleration Tolerance
The crew and troop occupant accelera-tion responses are compared with the whole-body acceleration tolerance curve establishedby Eiband [22]. The Eiband acceleration toler-ance levels were determined from sled impacttests on human volunteers, pigs, and chimpan-zees that were conducted for a single input ac-celeration pulse in the lateral, longitudinal, andvertical directions. Since the ACAP helicoptercrash test was performed under combined ve-locity conditions, the results of this comparisonmust be viewed with caution. In addition, theEiband curve was determined for a trapezoidal-shaped input acceleration pulse consisting ofthree phases: a ramp up phase to a uniform ac-celeration phase followed by a ramp downphase, as illustrated in Figure 21. The durationand magnitude of the uniform phase of the ac-celeration pulse, shown as the cross-hatchedarea in Figure 21, is then plotted on the Eibandcurve. However, the vertical acceleration re-sponses of the crew and troop occupants aresinusoidal in shape, not trapezoidal, as indicatedin Figure 21. Consequently, the peak accelera-tion and the pulse duration of the sinusoidal-shaped acceleration response are plotted on theEiband curve. This approach was used for thepilot, copilot, and troop results and it provides aconservative estimate of injury prediction, i.e. itlikely over predicts the severity of injury.
Table 2. Summary of head injury assessment.
CopilotPeak resultant head accel., g 486.7Time of peak acceleration, s 0.188
HIC for time duration of .036 s 713HIC for time duration of .015 s 1185
The magnitude and duration of the pilotchest and pelvis vertical acceleration responsesare plotted on the Eiband curve in Figure 22.These data fall on the border between areas ofmoderate and severe injury. Likewise the mag-nitude and duration of the copilot chest (torsosternum) vertical acceleration response areplotted on the Eiband curve in Figure 23. Forthe copilot, the results fall slightly more into thearea for moderate injury. Finally, the magnitudeand duration of the vertical chest accelerationresponses of the right and left troop dummiesare plotted on the Eiband curve in Figure 24.The troop data also fall into the area for moder-ate injury.
-5
0
5
10
15
20
25
30
35
0.05 0.1 0.15 0.2 0.25
Time, s
Acc
eler
atio
n, g
Duration
Magnitude
Figure 21. Vertical acceleration of the pilot pelvisfitted to the Eiband trapezoidal pulse.
1
10
100
1000
0.001 0.01 0.1 1
Acceleration, g
Duration, s
Area of severe injury
Area of moderate injury
Area of voluntary human exposure
Pilot pelvis
Pilot chest
Figure 22. Pilot chest and pelvis vertical accel-eration data plotted on the Eiband curve.
13
1
10
100
1000
0.001 0.01 0.1 1
Acceleration, g
Duration, s
Area of severe injury
Area of moderate injury
CopilotchestArea of voluntary human
exposure
Figure 23. Copilot chest vertical accelerationresponse plotted on the Eiband curve.
1
10
100
1000
0.001 0.01 0.1 1
Acceleration, g
Duration, s
Area of severe injury
Area of moderate injury
Left troopchest
Area of voluntary human exposure
Right troopchest
Figure 24. Vertical chest acceleration data of thetroop dummies plotted on the Eiband curve.
Injury Assessment Reference Values for Re-strained Occupants
Injury Assessment Reference Values(IARVs) are provided in Reference 23 for re-strained Hybrid III dummy occupants. Injuryguidelines are specified for head and neckforces and moments; head and chest accelera-tions; neck force and moment; and femur loads.These injury guidelines are listed in Table 3along with corresponding measurements fromthe modified Hybrid III anthropomorphic dummyrepresenting the copilot, where applicable. Theresults shown in Table 3 are for a mid-size (50thpercentile) male occupant. The IARVs havebeen suggested as guidelines for assessing in-
jury potentials associated with measurementsmade with Hybrid III-type adult dummies. TheIARVs refer to a human response level belowwhich a specified injury is considered unlikely tooccur for the given size individual. The data forthe copilot dummy are less than the IARVs listedin Table 3 with the exception of head/neck ten-sile force, neck compression force, and HIC.These results indicate a high probability of headand neck injury as a result of head strike.
Table 3. Comparison of copilot data with InjuryAssessment Reference Values (IARVs).
Units Mid-sizemaleIARV
Copilotdummy
dataHead/neck Flex moment in-lb. 1684 518.1 Extension moment in-lb. 505 215.2 Shear (fore/aft) lb. 247 245.9 Tension lb. 247 617.5 Compression lb. 247 217.5Chest acceleration(fore/aft)
g 60 47.7
Head (t2 - t1)≤ 15 ms HIC 1000 1185Neck Neck moment in-lb. 1684 1201.8 Neck compression lb. 247 444.4
Discussion of Results
The crew seat pan accelerations, shownin Figures 4 and 11(a), are higher than desiredand may be attributed to the fact that the originalenergy absorbing nose gear that was designedfor the Sikorsky ACAP helicopter was not avail-able for this test. Instead, a standard non-crashworthy commercial nose gear that was ret-rofitted to provide a nominal level of energy ab-sorption was installed on the helicopter. Theretrofitted nose gear did not provide sufficientenergy attenuation and failed early during thetest. If the original crashworthy nose gear hadbeen available, the combination of nose gearcrushing plus seat stroke might have loweredthese acceleration levels.
The chest acceleration comparisonsbetween the crew and troop dummy occupantsshown in Figure 8 appear to be counter-intuitive.These plots indicate that the troop occupantsexperienced lower forward and vertical chestacceleration responses than did the crew dum-mies, in spite of the fact that the crew seats ex-hibited 9- and 14.5-inches of stroking and thetroop seats had minimal stroking. The explana-tion for this behavior is that the troop dummies
14
benefited from the fact that they were located inclose proximity to the center-of-gravity of thehelicopter and were placed almost directlyabove the main landing gear.
To illustrate this point, the raw data ob-tained from accelerometers located on the floornear the pilot and right troop, and near the copi-lot and left troop were integrated to obtain thevertical velocity responses shown in Figure 25.The initial velocity was determined for each lo-cation as the sum of the initial vertical velocity ofthe center-of-gravity plus the rotational compo-nent which is the pitch angular velocity times thedistance from the accelerometer location to thecenter-of-gravity. As shown in Figure 25, themagnitude of the velocity response of the pilotand copilot floor increases slightly for the first0.1 seconds of the pulse, whereas the magni-tude of the velocity response of the right and lefttroop floor is decreasing during the same timeperiod. At the time of fuselage contact with theground at 0.098 seconds, the velocity at the pilotand copilot floor is approximately -460 in/s. Thecorresponding velocity of the right and left troopfloor is -300 in/s at the same time. Thus, theright and left troop dummies benefited from theenergy attenuation achieved from crushing ofthe aluminum honeycomb stage of the mainlanding gear. In contrast, the original energyabsorbing nose gear was missing from theACAP helicopter and a modified civil helicopternose gear was retrofitted to the airframe. Thereplacement gear was not designed to provide ahigh-level of energy absorption and failed earlyupon impact after providing 9 inches of stroke[11].
It is interesting to note the differences inthe shape, magnitude, and duration of the pilotand copilot chest acceleration responses in boththe vertical and forward directions, shown inFigures 8(a) and 9(a). In general, the copilotdata exhibits more oscillations than the pilotdata. In contrast, the vertical and forward accel-eration responses of the right and left troop,shown in Figures 8(b) and 9(b), have the sameoverall shape, duration, and magnitude. Thevariations in crew occupant responses cannotbe attributed to differences in the seat pan orfloor-level acceleration pulses, which are shownin Figures 11 and 13, respectively. Therefore,the varying responses must be attributed to thedifferences in the Hybrid II (pilot) and the modi-fied Hybrid III (copilot) dummies.
-500
-400
-300
-200
-100
0
100
200
0 0.05 0.1 0.15
PilotRight troop
Velocity, in/s
Time, s
(a) Pilot and right troop floor.
-500
-400
-300
-200
-100
0
100
200
0 0.05 0.1 0.15
CopilotLeft troop
Velocity, in/s
Time, s
(b) Copilot and left troop floor.
Figure 25. Vertical velocity responses.
It is evident from the copilot head and T-1 thoracic acceleration responses, plotted inFigures 6 and 7 respectively, and the head andneck forces listed in Table 3 that the copilotdummy experienced a head strike during thecrash test. As mentioned previously, a smallpiece of cloth matching the fabric in the copilot'sflight suit was found on the right side of thedummy's forehead that appeared to originatefrom the dummy's knee. The actual head strikewith the dummy's knee could be seen in thehigh-speed film, which showed that the copilot'shead and upper torso experienced a largedownward translation and forward rotation dur-ing the test. The copilot displayed more forwardmotion of its upper torso and head than did thepilot dummy. The forward pitching of the copilotseat back and the added weight of the helmetcontributed to the excessive forward motion ofthe copilot.
15
The HIC values determined from the re-sultant head acceleration of the copilot rangedfrom 713 to 1185 depending on the time intervalused in the calculation. For the mid-size HybridIII dummy, HIC values greater than 1000 areindicative of a 16 percent risk of serious braininjury. The maximum one-dimensional DRIvalue of 22.3 for the copilot was the lowest ofany of the four dummy occupants, yet still indi-cates a 40% risk of spinal injury. The spinalforce measured for the copilot was 1,921 lbs,which exceeded the 1,500-lb. threshold. In ad-dition, the magnitude and duration of the copilotchest acceleration response falls into the regionfor moderate injury on the Eiband whole-bodyacceleration tolerance curve.
Without the forward component of headacceleration, it was not possible to perform aHIC assessment for the pilot occupant. How-ever, the other injury criteria indicate that thepilot experienced moderate to severe injury.The maximum DRI for the pilot was 28.6, indi-cating a greater than 50% risk of spinal injury.The risk assessment for the pilot using the com-bined dynamic response model shown in Equa-tion 1 indicated that each of the low-, moderate-,and high-risk curves exceeded the thresholdvalue. The maximum lumbar load measured forthe pilot was 1,912 lbs, exceeding the 1,500-lb.limit for spinal injury. Finally, the magnitude andduration of the pilot chest and pelvic vertical ac-celeration responses fell on the border betweenmoderate and severe injury on the Eibandwhole-body acceleration tolerance curve.
Less detailed occupant response infor-mation was available for the troop dummies ascompared to the crew dummies since they wereless heavily instrumented. The maximum DRIvalues for these occupants ranged from 28.6 to30, indicating a greater than 50% risk for spinalinjury. However, these values must be viewedwith caution, since the cloth seat pans in thetroop seats failed during the test allowing thebuttocks of the dummy occupants to displacethrough the seat frame. Consequently, the verti-cal response of the seat frame is not likely to bea good indicator of occupant response or injurypotential for the troop dummies. The magnitudeand duration of the vertical chest accelerationresponses of the left and right troop dummiesfell into the area of moderate risk on the Eibandwhole-body acceleration tolerance curve.
As mentioned previously, several differ-ent criteria were applied to each occupant todetermine the risk of injury for this crash test andconsistent results were obtained. The overallassessment of occupant injury indicates that theACAP helicopter crash test resulted in a moder-ate to high level of risk for injury. Although someinjuries would likely have occurred in this crash,the probability of a fatality is considered small.
Concluding Remarks
A full-scale crash test of the SikorskyACAP flight test helicopter was performed at theImpact Dynamics Research Facility at NASALangley Research Center in June 1999. Thepurpose of the test was to generate experimen-tal data for correlation with a nonlinear, explicittransient dynamic crash simulation developedusing the MSC.Dytran finite element code. Forthe test, the helicopter was outfitted with twocrew and two troop seats, and four anthropo-morphic test dummies. While the results of theimpact test and crash simulation have beendocumented fairly extensively in the literature,the objective of this paper is to present the de-tailed occupant response data obtained from thecrash test and to correlate the results with injuryprediction models. These injury models includethe Dynamic Response Index (DRI), the HeadInjury Criteria (HIC), the spinal load requirementdefined in FAR Part 27.562(c), and a compari-son of the duration and magnitude of the occu-pant vertical acceleration responses with theEiband whole-body acceleration tolerance curve.
The pilot was a 50th percentile Hybrid IImale dummy that was placed in a used com-mercial military-qualified helicopter seat pro-vided by the US Army Aeromedical ResearchLaboratory (USAARL). This seat contained twoinvertube energy absorbers. New MA-16 inertialocking reels were used with lap and shoulderbelts to restrain the dummy occupant during thetest. USAARL also supplied a 50th percentilemodified Hybrid III male dummy with a self-contained data acquisition system (DAS) for theleft front crew (copilot). The copilot dummy wassecured in a used commercial military-qualifiedhelicopter seat of a different design than the pilotseat. This seat contained six "torshock" energyabsorbers. Likewise, an MA-16 inertia lockingreel was used in conjunction with the restraintsystem to limit the displacement of the dummyoccupant during the test. Two 50th percentile
16
Hybrid II male dummies were used for the rightand left troop occupants. These dummies wereseated in ceiling-suspended troop seats withwire-bender energy absorbers that weremounted in the rear cabin area. New wire-bender energy absorbers were installed in thetroop seats; however, these seats were usedand the seating material was in poor condition.
The head and chest acceleration re-sponses of the copilot indicate a strong likeli-hood of head strike. In fact, the HIC values forthe copilot ranged from 713 to 1185, dependingon the time interval used in the calculation. AHIC of 1000 is associated with a 16 percent riskof serious brain injury. The maximum one-dimensional DRI value of 22.3 for the copilotwas the lowest of any of the four dummy occu-pants, yet still indicates a 40% risk of spinal in-jury. The spinal force measured for the copilotwas 1,921 lbs, which exceeded the 1,500-lb.threshold. In addition, the magnitude and dura-tion of the copilot chest acceleration responsefalls into the region for moderate injury on theEiband whole-body acceleration tolerance curve.The maximum DRI for the pilot was 28.6 and thelumbar load was 1,912-lb., both of which arestrong indicators of spinal injury. Less detailedoccupant response information was available forthe troop dummies as compared to the crewdummies since they were less heavily instru-mented. However, the magnitude and durationof the vertical chest acceleration responses ofthe left and right troop dummies fell into the areaof moderate risk on the Eiband whole-body ac-celeration tolerance curve.
Finally, several different criteria wereapplied to each occupant to determine the risk ofinjury for this crash test and consistent resultswere obtained. The overall assessment of oc-cupant injury indicates that the ACAP helicoptercrash test resulted in a moderate to high level ofrisk for injury. Although some injuries wouldlikely have occurred in this crash, the probabilityof a fatality is considered small.
Acknowledgements
The authors wish to acknowledge thesupport of the US Army Aeromedical ResearchLaboratory (USAARL) in providing the modifiedHybrid III anthropomorphic dummy with self-contained data acquisition system and seats foruse in the ACAP helicopter full-scale crash test.
In addition, we would like to acknowledge H.Koch and Sons for providing new MA-16 inertialocking reels and restraint systems for the test.
References
1. Vaughan, Victor L., Jr. and Alfaro-Bou, Emilio,“Impact Dynamics Research Facility for Full-Scale Aircraft Crash Testing,” NASA TN D-8179,April 1976.
2. Cronkhite, J. D., and Mazza, L. T., "Bell ACAPFull-Scale Aircraft Crash Test and KRASH Cor-relation," Proceedings of the 44th Annual Forumand Technology Display of the American Heli-copter Society, Washington D.C., June 16-18,1988.
3. Perschbacher, J.P., Clarke, C., Furnes, K.,and Carnell, B., “Advanced Composite AirframeProgram (ACAP) Militarization Test and Evalua-tion (MT&E) Volume V- Airframe Drop Test,”USAATCOM TR 88-D-22E, March 1996.
4. Jones, L. E., "Overview of the NASA’s Sys-tems Approach to Crashworthiness Program,"Proceedings of the American Helicopter Society58th Annual Forum, Montreal, Canada, June 11-13, 2002.
5. MSC.Dytran User’s Guide, Version 2002,MSC.Software Corporation, Santa Ana, CA, No-vember 2001.
6. Fasanella, E.L., Jackson, K.E. and Lyle, K.H.,“Finite Element Simulation of a Full-Scale CrashTest of a Composite Helicopter,” Proceedings ofthe American Helicopter Society 56th Annual Fo-rum, Virginia Beach, Virginia, May 2-4, 2000.
7. Lyle, K. H., Jackson, K.E., and Fasanella, E.L., “Development of an ACAP Helicopter ImpactModel,” Journal of the American Helicopter So-ciety, Vol. 45, No. 2, April 2000, pp. 137-142.
8. Fasanella, E. L., Boitnott, R. L., Lyle, K. H.,and Jackson, K. E., “Full-Scale Crash Test andSimulation of a Composite Helicopter,” Interna-tional Journal of Crashworthiness, Vol. 6, No. 4,November 2001, pp. 485-498.
9. K. H. Lyle, K. E. Jackson, and E. L. Fasanella,"Simulation of Aircraft Landing Gears with aNonlinear Transient Dynamic Finite Element
17
Code,” Journal of Aircraft, Vol. 39, No. 1, Janu-ary-February, 2002.
10. Boitnott, R. L., Jackson, K. E., Fasanella, E.L., and Kellas, S.: “Full-Scale Crash Test of theSikorsky Advanced Composite Airframe Pro-gram Helicopter,” Proceedings of the AmericanHelicopter Society Forum 56, Virginia Beach,VA, May 2-4, 2000.
11. Boitnott, R. L., Jones, L. E., "NASA LangleyResearch Center's Impact Dynamics ResearchFacility Full-Scale Crash Test Procedures," Pro-ceedings of the 3rd KRASH Users Seminar, Ari-zona State University, Tempe, AZ, January 8-10, 2001.
12. Code of Federal Regulations, Federal Avia-tion Regulations for Aviation Maintenance Tech-nicians FAR AMT, Part 27 Airworthiness Stan-dard: Normal Category Rotorcraft, 27.562Emergency Landing Dynamics.
13. Stech, E. L. and Payne, P. R., "DynamicModels of the Human Body," AAMRL-TR-66-157, Aerospace Medical Research Laboratory,Wright-Patterson Air Force Base, Ohio, 1969.
14. Brinkley, J. W. and Shaffer, J. T., "DynamicSimulation Techniques for the Design of EscapeSystems: Current Applications and Future AirForce Requirements," Aerospace Medical Re-search Laboratory; AMRL Technical Report 71-292, Wright-Patterson Air Force Base, Ohio,December 1971, AD 740439.
15. Coltman, J. W., Van Ingen, C., Johnson, N.B., and Zimmerman, R. E., "Crash Survival De-sign Guide, Volume II - Aircraft Design CrashImpact Conditions and Human Tolerance,"USAAVSCOM TR 89-D-22B, December 1989.
16. Brinkley, J. W. and Mosher, S. E., "Devel-opment of Acceleration Exposure Limits to Ad-vanced Escape Systems," Implications of Ad-vanced Technologies for Air and Spacecraft Es-cape, AGARD-CP-472, April 24-28, 1989.
17. Mosher, S. E., "DYNRESP Six Degree-of-Freedom Model for Injury-Risk Evaluation User'sManual," NASA Johnson Space Center, April 29,1993.
18. Federal Motor Vehicle Safety Standard No.208, Occupant Crash Protection, Code of Fed-
eral Regulations, Title 49, Part 571.208, U. S.Government Printing Office, Washington D. C.,1987.
19. Anon., "Human Tolerance to Impact Condi-tions As Related to Motor Vehicle Design -SAEJ885," APR 80, SAEJ885, Society of Auto-motive Engineers, Inc., Warrendale, PA, April1980.
20. Gadd, C. W., "Use of a Weighted-ImpulseCriterion for Estimating Injury Hazard, " Pro-ceedings of the Tenth Stapp Car Crash Confer-ence, Society of Automotive Engineers, NewYork, 1966.
21. Nusholtz, G. S., et al., "Critical Limitations onSignificant Factors in Head Injury Research,"Proceedings of the 13th Stapp Car Crash Con-ference, Society of Automotive Engineers, Inc.,Warrendale, PA, October, 1986.
22. Eiband, A. M., "Human Tolerance to RapidlyApplied Accelerations: A Summary of the Lit-erature," NASA Memorandum 5-19-59E, Na-tional Aeronautics and Space Administration,Washington D.C., June 1959.
23. Mertz, H. J., "Injury assessment values usedto evaluate Hybrid III response measurements,"NIHTSA Docket 74-14, Notice 32, Enclosure 2of Attachment I of Part III of General MotorsSubmission USG 2284. March 22, 1984.
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REPORT DOCUMENTATION PAGE
June 2002 Technical Memorandum
Occupant Responses in a Full-Scale Test of the Sikorsky ACAP HelicopterWU 728-50-10-01
Karen E. Jackson Joseph McEntire Edwin L. Fasanella Alan LewisRichard L. Boitnott
L-18189
NASA/TM -2002-211733ARL-TR-2735
Presented at the AHS International - The Vertical Flight Society's 58th Annual Forum and Technology Display,June 11-13, 2002, Montreal, Canada.
A full-scale crash test of the Sikorsky Advanced Composite Airframe Program (ACAP) helicopter was performedin 1999 to generate experimental data for correlation with a crash simulation developed using an explicit nonlinear,transient dynamic finite element code. The airframe was the residual flight test hardware from the ACAP program.For the test, the aircraft was outfitted with two crew and two troop seats, and four anthropomorphic test dummies.While the results of the impact test and crash simulation have been documented fairly extensively in the literature,the focus of this paper is to present the detailed occupant response data obtained from the crash test and to correlatethe results with injury prediction models. These injury models include the Dynamic Response Index (DRI), theHead Injury Criteria (HIC), the spinal load requirement defined in FAR Part 27.562(c), and a comparison of theduration and magnitude of the occupant vertical acceleration responses with the Eiband whole-body accelerationtolerance curve.
Hybrid II 50th percentile male ATD, Anthropomorphic Test Dummy (ATD), full-scaleaircraft crash testing, crashworthiness, prediction methods
22
NASA Langley Research Center U.S. Army Research LaboratoryHampton, VA 23681-2199 Vehicle Technology Directorate
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andU.S. Army Research LaboratoryAdelphi, MD 20783-1145
Unclassified–UnlimitedSubject Category 05 Distribution: StandardAvailability: NASA CASI (301) 621-0390
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