USAAVRADCOM-TR-82-D-42

NADC-82268-60

FAA-AM-83-3 U.S. Department of Transportation

Federal Aviation Administration

'. DESIGN AND TEST CRITERIA FOR INCREASED

ENERGY-ABSORBING SEAT EFFECTIVENESS

"~ •kJoseph %&Coltman

., SIMULA, Inc.

2223 S. 48th Street

Tempe, Ariz. 85282 S' 

March 1983

Final Report for Period June 1979 -February 1983

distribution unlimited.

Prepared for

~ APPLIED TECHNOLOGY LABORATORY

U. S. ARMY RESEARCH AND TECHNOLOGY LABORATORIES (AVRADCOM)

'Cý Fort Eustis, Va. 23604

"w w w w w w l w "r w w w w w w
APPLIED TECHNOLOGY LABORATORY POSITION STATEMENT

This report documents one of several efforts the Applied Technology Laboratory, US Army

Research and Technology Laboratories (AVRADCOM) has initiated jointly with the Naval Air

Development Center (NADC), the FAA Civil Aeromedical Institute (CAMI) and the Air Force

Aerospace Medical Research Laboratory (AFAMRL) to define human tolerance levels to vertical crash impact and to optimize crashworthy crewseat design. Results of this program represent a significant advance in the understanding of dynamic/structural/anatomical interactions

which affect occupant injury in severe but survivable crashes (over 75 percent of all Army

helicopter mishaps are survivable). These findings will be integrated with those of ongoing

and future efforts in human tolerance definition and crashworthy seating design with the end

objective being a more precise set of design and qualification criteria leading to less costly and

more weight-efficient crashworthy seats.

"Mr. Kent F. Smith of the Aeronautical Systems Division served as project engineer for this

effort.

DISCLAIMERS

The findings in this tport are not to be construed as an official Dew4rtment of the Army potition unlgU so

deslgtsistd by other authoorked documents.

When Government drawings. qiacff,:ations, or other data sit used for nv pu.rpose other then in connection

with a dfit*Iy rWat•4i G(.•lenm•nt (nocsvrement o•praion. (he United Status Govetnment thereby incurs no

responsibillty not any obligatlion whiatoevet: and the fact that the Government may have formulated, furnished.

or nl asny way supplied the soid drawings. specifications, or other data is not to be rogattded by imolicalion or

otherwie as in any manner licensng the holder or any other person or corporation, or conveying any rights or

Vermiuio•n, to menufacturo. use. O0 Yell py patented Inveition that may in any way be relotqd thereto.

Trade names cited in this report do not constitute an official endotrsmen"t or Approval of the us* of such

commercial hardware or sofrwo..

DISPOSITION INSTRUCTIONS

Destroy this report when no longer netded. Do not return it to the originator.

4W
UNCLASSIFIED

SECURITY CLASSIFICATION OF THIS PAGE (•omn D.a Entett

SREPORT DOCUMENTATION PAGE READ INSTRUCTIONS

BEFORE COMPLETING FORM i. REPORT NUMBER 2.GOVT ACCESSION NO. 3. RECIPIENT'S CATALOG NUMBER

USAAVRADCOM TR 82-D-42.

4. TITLE (and Subtitle) 5. TYPE OF REPORT A PERIOD COVERED

F IN"AL REPORT

DESIGN AND TEST CRITERIA FOR INCREASED 1979 - February 1983

ENERGY-ABSORBING SEAT EFFECTIVENESS 6. PERFORMING O/'. REPORT NUMBER

TR-82411 %/

7. AUTHOR(.) S. CONTRACT OR GRANT NUMBER(&)

Joseph W. Coltman DAAK51-79-C-0026

-. PERFORMING ORGAN.KATION NAME ANMC ADDRESS 10. PROGRAM ELEMENT. PROJECT, TASK ,"Simula In(. AREA A WORK UNIT NUMBERS

2221 S. 48th Street 612209 1LI62209AH76 00 271EK

Tempe, Arizona 95282

III. CONTR'LLING OFFIC. NAME ANO ADODI-SS 12. REPORT OATE

"":)pplied Tzchnoloqy liboratory, U.S. Army Research March 1983

and- Techrology Laboratories 'AVRADCOM) 13. MUMMER OF PAGES

*"Fort Eustis, Virginia Z3604 229 IT • - *- ORIMG AOENC' NAMi C AOOIwki o $S(U Ue uittgmi Chow fll Oll0 ) IS. SECURITY CLASS. (at tl. rec*I)

Unclassified

IS CUOECLASSlFICATION/N5owNGRAOING

10. INSPIUTO STACHUENT (011M lAS. 00e

Approred for public rolea3e; distribution unlimited.

17. 0D1ISTPUIUOu STATEMNT (014M l.b.*. m40g a*&WA. 30, Ii d how4 A. Re

11SJJWLCMEMTANV MOTUS

Aircraft Restraint Systems

Crastivorthinona SUats

Design Criteria

Energy Absorption

M AinTUAZVM .1ft ,.MWm a#iffM.Ui W d#&ItPr MpNet& MWINEj

"IMTia report documentA a research effort to increase the effectiveness of enerqyabsorbing seats, through improved design and qualification test criteria. Contained herain are descriptions of a parametric tent program and analysis of seat

and occupant rosponse sensitivity to design and test variables Reco~mwidations

for improving military specifications and criteria, such as coht-Ainod in NIL-S5809S(AV), HIL-STD-1290(AV) and USARTL TP-79-22A, to aid in procurement of

optimum systems are also provided.

Joe1~5 O&V00 S Nva 30546 0mTe UNICLASS IF! ED

* sctjaTv CLAWIRPCAYIOII Or TSUI PABE (1m10w' Do 'm W4

-% *
PREFACE

This report was prepared by Simula Inc. under Contract DAAK51-79-

C-0026 for the Safety and Survivability Technical Area of the Applied Technology Laboratory, U.S. Army Research and Technology

Laboratories (AVRADCOM), Fort Eustis, Virginia. Mr. Kent Smith

of ATL has acted as the Technical Monitor. When the program was

initiated in 1979, Mr. George T. Singley III, then of ATL, served

in that capacity.

The Program Manager for the Naval Air Development Center contribution to the program has been Mr. Mark Katzeff of the Aircraft

and Crew Systems Technology Directorate. Data acquisition and

reduction at NADC was conducted under the direction of Mr. Leon

Domzalski of the same Directorate. The Federal Aviation Administration effort was managed by Mr. Richard F. Chandler, Chief of

the Protection and Survival Laboratory, Civil Aeromedical Institute.

"2 This report has been prepared by Mr. Joseph W. Coltman of

* Simula Inc., with contributions from Dr. David H. Laananen and

1Mr. Stanley P. Desjardins.

Ilk

. , •,~ . 1

S° ----4

3
TABLE OF CONTENTS

PREFACE. . . . . . . . .. . .. . . . . . . . . . . . . 3

LIST OF ILLUSTRATIONS ........ ......... 8

LIST OF TABLES . . . . . . . . . . . . . . . . . . . . . 14

1.0 INTRODUCTION. . . . . . . . . . . . . . . . . . . . 17

2.0 HUMAN TOLERANCE TO ACCELERATIVE LOADING ...... 22

2.1 HUMAN TOLERANCE INVESTIGATIONS . . . . . . . . 22

2.2 TOLERANCE-BASED CRITERIA FOR SEATS . . . . . . 25

2.3 SEAT/OCCUPANT DYNAMIC RESPONSE . ... . . . . 29

"3.0 PARAMETRIC STUDY. .......... ....... 32

3.1 VARIABLES EXAMINED ...... . . . . . . . . 32

3.1.1 Impact Conditions . . . . . . . . . . . 33

3.1.2 Occupant Characteristics. . . . . . . . 33

3.1.3 Seat Design Parameters. . . . . . . . . 34

3.2 TEST FACILITIES ...... ..... .. . 36

3.2.1 FAA Civil Aeromedical Institute (CA,,) ........... . ...... 3

3.2.2 Naval Air Development Center (NAX) . . 36

3.2.3 Simula Inc. . . . . . . . 36

3.2.4 Wayne State University (WSU). . . . . . 36 -.4

3.3 PARAMETRIC TEST MATRIX. . . . . . . . . . . 43

"3.4 CADAVER TEST PROGRAM . . . . . . . . . . . . . 43

"3.5 MODIFIED ANTHROPOMORPHIC DUMMY PROGRAM . . . 48

4.0 EQUIPMENT AND TEST PROCEDURES. . . . . . . . . . . 55

4.1 EQUIPMENT - SEATS. . .... ....... . ... 55

4.1.1 Rigid Seat . . ...... 55

4.1.2 Energy-Absorbing Helicopter Seat. 55

4.2 ANTHROPOCORPS4IC DUMMIES. .......... 62

4.3 ANALYSIS OF ACCELERATION PULSE . . . . .. 63

9;22:
"TABLE OF CONTENTS (CONTD)

4.4 STANDARDIZED TEST PROCEDURES . . . . . .*. . . 65

4.4.1 Seat Orientation. . . . . . . . . . . . 65

4.4.2 Seat Preparation...... . . . . . . 68

4.4.3 Dummy Positiohiing. . . . . . . . . . . 68

4.5 INSTRUMENTATION. * . . . . . . . . . . . . . . 69

5.0 ANALYTICAL METHODS . . . . . . . . . . . . . . . . 72

5.1 CURRENT CRITERION. . . . . . . . . . . . . . . 72

5.2 PEAK ACCELERATION. . . . . . . . . . . . . . . 74

5.3 COMPRESSION OF BODY SEGMENTS. . . . . . . . . 75

5.4 DYNAMIC RESPONSE INDEX . . . . . . . . . . . . 75

"5.5 INSTRUMENTED ANTHROPOMORPHIC DUMMY . . . . . . 77

5.6 SEAT/OCCUPANT SIMULATION .. . .. .. . . . . 79

5.7 LINEAR REGRESSION ANALYSIS . . . . . . . . . . 95

6.0 RESULTS OF SENSITIVITY ANALYSIS . . . . . . . . . . 96

"6.1 SENSITIVITY ANALYSES FOR EXPERIMENTAL TEST

SERIES . . . . . . . . . . . . . • . • • . .96

6.1.1 Peak Input Acceleration Series . . 97

6.1.2 Velocity Change Series . . . . . . . . 98

6.1.3 Rate of Onset Series . , . . . . . . . 106

"6.1.4 Energy Absorber Limit-Load Series. 112

6.1.5 Test Facility Impact Conditions. . . . 117

6.1.6 Dummy Type Series. .. , . . . . . 125

"6.1.7 Dummy Percentile Series. . . . . . . . 126

6.1.8 Cadavers Versus Anthropomorphic

6 Dummies. . . . . . . . . . . . . . . . . 129

6.1.9 Seat Orientation Series. . . . . . . . 133

6.1.10 Ramped Energy Absorber Series... 134

"6.1.11 Movable Seat Weight Series . . . . . 136

6.1.12 Seat Frame Spring Rate Series .... 141

6.1.13 Seat Cushion Stiffness Series. . . . . 142

. 6.2 ADDITIONAL PERFORMANCE VARIABLES . . . . . . 144

6.2.1 Dummy Performance Degradation . . . 147

6.2.2 Footrest Location and Angle . . . . . . 148

6.2.3 Landing Gear Acceleration Ramp. . . . . 150

6.2.4 Test Orientation to Gravity Vector. . 152

6

- - - - - - 0
TABLE OF CONTENTS (CONTD)

Page

6.2.5 Tolerances on Input Deceleration

Pulse Allowed by MIL-S-58095(AV) . . . . 154

6.2.6 Method of Transducer Calibration .... 155

7.0 DISCUSSION OF RESULTS . . . . . . . . . . . . . . . 159

7.1 INPUT ACCELERATION PULSE. . . . . . . . . . . . 159

7.2 PERFORMANCE CRITERION . . . . . . . . . . . . . 159

7.3 SEAT DESIGN PARAMETERS. . . . . . .. 160

7.4 TESTING PROCEDURES. .............. 161

CONCLUSIONS. . . . . . S8.0 . . . . . . . . . . . . . . . 163

9.0 RECOMMENDATIONS. .................. 164

9.1 FOLLOW-ON RESEARCH. .......... ... 164

9.1.1 Energy Absorber Limit-Load Threshold . . 164

9.1.2 Standardized Test Dummy. . . . . .. . . 164

9.1.3 Refinement of Qualification Test

Methods. . . . . . . . . . . . . . . . . 165

9.2 SUGGESTED MODIFICATIONS TO CURRENT ENERGYABSORBING SEAT SPECIFICATIONS . . . . . . . .. 165

10.0 REFERENCES. . . . . . . . . . . . . . . . . .. . . 167

APPENDIXES

A - CAMI TEST A81-121. * ... . 170

B - CAMI TEST A81-122. .... . . . . .... . . . . 179

C - CAMI BASELINE TEST A8O-053 .. •......... 190

D - NADC BASELINE TEST N-189 . .......... 201

E - SIMULA BASELINE TEST SEAC-1 . . . . . . .. . 210

F - WAYNE STATE UNIVERSITY TEST 159. . . . . . . . . 219

7

- C;P:Tg:: ~ w
LIST OF ILLUSTRATIONS

Figure Page

1 Overall program elements leading to improved

seat design and test criteria. . . . . . . . . 20

2 Duration and magnitude of spineward acceleration endured by various subjects . . . . . . . 23

3 Duration and magnitude of headward acceleration endured by various subjects. . . . . . . 24

4 Maximum acceptable vertical seat

acceleration . . . . . ....... ..... 27

5 Typical response of seat pan, dummy chest,

and dummy pelvis to vertical crash loading . , 28

6 Spring-mass representation .of seat-occupant

system . . . . . . . . . . . . . . . . . . . . 30

7 Triangular deceleration pulse describing

the nominal test condition . ........ 35

8 CAMI aled. . . . . . . . . . . 37

9 CAMI wire-bending decelerator mechanism.... 38

10 Typical baseline deceleration pulses

for the four test facilities . . . . . . 39

11 NADC drop tower ...... . . . . . . . . . 40

' 12 NADC decelerating mechanism . . . . . .... 41

13 Drop tower used in Simula tests. . . . . . . . 42

14 P.I-.t 572 pelvic stI~ment with lumbar spine , 1.ad coil assembly . . . . . . . . . . . 52

15 CAMI rigid seat (with VIP-95 dummy Sin place) .. .. . . . .. . . . .. . . . . . 56

"16 UH-60A crewseat assembly . . . . . . . . 57

17 Inverted-tube energy absorber. .. . . . . . . 59

18 Nominal load-stroke characteristics of ranped

energy absorbers used in parametric test

program... . . . . . 60

19 Aircrew restraint system ........... 61

* 8

". -. ' "-W " -" 
LIST OF ILLUSTRATIONS (CONTD)

Figure Page

20 Rigid block in Black Hawk crewseat . . . . . . 64

21 Graphic approximation example for rate

of onset . . . . . . . . . . . . . . . . . . . 65

22 Vertical test, drop tower configuration .... 66

23 Vertical test, horizontal sled

configuration (CAMI) ............. 67

24 Tensiometer placement on occupant restraint

system . . . . . . . . . . . . . . . . . . . . 70

25 Vertical component of seat pan acceleration

measured in two similar tests. . . . . . . . . 73

26 Seat pan vertical acceleration measured in a

qualification test of a production energy- absorbingseat t............... 74

27 Spinal-injury model. . . . . . . . . . . . . . 76

28 Probability of spinal injury estimated from

laboratory data compared to operational

experience......... .... . . . . 78

29 Eleven-segment (symmetric) occupant

30 Generalized coordinates for symmetric

occupant model .... .81

31 SON-LA input properties for Black Hawk seat

cushion. . . . . . . . . . . . . . . . . . 82

32 SON-LA input properties for polyester restraint system webbing ........... 83

33 Actual sled deceleration pulse for CANI.

test A81-124 ... . . . . . . .... 84

34 Approximate pulse input ror SOW-LA

simulation of test A81-124 . . . . . . .... 84

35 CAMI test A81-124, dummy pelvis

z acceleration ..... .. .g..... 85 "4 :

36 CAMI test A81-124, dummy chest

z-acceleration ......... . . . . . . . 86

9
LIST OF ILLUSTRATIONS (CONTD)

Figure Page

37 CAMI test A81-124, dummy head -

z-acceleration . . . . . . . . . . . . . . . . 87

38 CAMI test A81-124, seat z-acceleration .... 88

39 CAMI test A81-124, seat stroke . . . . . . . . 89 4

40 CAMI test A81-124, energy absorber force . . . 90

41 CAMI test A81-124, footrest x-force. . . . . . 91

"42 CAMI test A81-124, footrest z-force. . . .. . 92

43 CAMI test A81-124, lumbar spine axial load . . 93

44 CAMI test A81-124, lumbar spine moment . . . . 94

45 Rigid seat series parametric test reference. . 96

46 Energy-absorbing seat series parametric test reference . .. .. .. . .. . . . . . . 97 ['

47 Inpit acceleration pulses for peak

"acceleration series. , . • . . . .... . . . 98

48 Peak input acceleration series, maximum

seat stroke. . . . . . . . . . . . . . . . . 100

49 Peak input acceleration series, maximum

seat pan z-axis acceleration . . . . . . . . 100

50 Seat pan z-axis acceleration measured in

peak input acceleration series . . . . . . . . 101

51 Pelvic z-axis acceleration measured in

peak input acceleration series . . . . . . . . 102

52 Peak input acceleration series, maximum

pelvic a-axis acceleration . . . . . .. . . . 103

53 Peak input acceleration series, duration of

"seat pan z-axis acceleration at 23 G . . . . . 103

54 Peak input acceleration series# Dynamic

Rosponse Index ... . .. . .. . 104

55 Peak input acceleration series, predicted

. maximum lumbar spine axial load

(compression). . . . . . . . . . . . . . . . . 104

1*. • ,,•• • m •• 10i.g•Am•,- 

- . *- ~ .w ~ *w w '' 10* ~ *~' - w 

. . . .. . . . . . . . . . . . . .
LIST OF ILLUSTRATIONS (CONTD)

Figure

56 Peak input acceleration series, predicted

maximum lumbar spine moment (flexion). . . . . 105

57 Comparison of input acceleration pulses

for CAMI velocity change series. . . . . . . . 105

58 Input velocity change series, maximum seat stroke . . . . . . . . . . . .. . .. .. . 107

59 Input velocity change series, maximum seat

pan z-axis acceleration. . . . . . . . . . . . 107

60 Input velocity change series, Dynamic

Response Index . . . . . . . . . . . . . . . . 108

61 Input velocity change series, predicted maximum lumbar spine axial load (compression). . . . . . . ......... 108

62 Input velocity change series, predicted

lumbar spine moment (flexion).. . . . . . . . 109

63 Rate of onset series, maximum seat stroke... 110

64 Comparison of input acceleration pulses

for rate of onset series . . . . . . . . . .. 11

65 Comparison of seat pan z-axis acceleration

for rate of onset series .. ..... . . . 111

66 Energy absorber limit load series, maximum

seat stroke................ 114

67 Energy absorber limit load series, maximum

seat pan z-axis acceleration. . . . . . . . . 114

68 Energy absorber limit load series, Dynamic Response Index ................ 115

69 Energy absorber limit load series# prdicted

maximum lumbar spine load (compression).... 115

70 Energy absorber limit load series, predicted

maximum lumbar spine moment (flexion) ..... 116

71 Energy absorber limit load series, maximum

lumbar spine moment versus seat stroke . . . . 117

72 Nominal baseline input acceleration pulsed.. lie

""'77-%7o

.44.
LIST OF ILLUSTRATIONS (CONTD)

Figure Page

73 Baseline acceleration pulses for the four

test facilities. . . . . . ... ... ............ 119

"74 Comparison of pelvis and seat pan z-axis

acceleration for baseline test conditions

at three facilities.. ......... ... 120

"75 Test facility impact condition series,

maximum seat stroke. ............ . 122

76 Test facility impact condition series, maximum

seat pan z-axis acceleration . . . . . . . . . 161

77 Test facility impact condition series,

duration at 23 G . of 

. .seat . . pan . . z-axis . . . .acceleration . .. . .. ... 123 '

78 Test facility impact condition series, DRI . . 123

79 Test facility impact condition series,

maximum lumbar spine axial compressive

load 124

80 Test facility impact condition series,

maximum lumbar spine y-axis moment

(flexion) ... . . . . . . . . . . . . . 124

81 Sled deceleration, rigid seat test A80-026 125

* 82 Comparison of seat pan force measured in

rigid seat tests with three 50th-percontile

dummies . . . . . . . . . .. . . . . 127

83 Seat pan vertical acceleration for a Part 572

dummy mode and two cadavers measured in vertical tests with 14.5-G energy absorbers .... 131

84 Seat pan vertical acceleration for a Part 572

dummy and two cadavers measured in combined

mode tests with 14.5-G energy absorbers. . . . 131

"85 Comparison of normalized seat stroke for

cadavers and dummies tested with the Wayne

State University input pulse shape ...... 132

86 Seat orientation series, maximum seat

stroke . . . . . . . . . . . . . , . . . .. . 135

.87 Comparison of energy absorber force-time

histories for ratped EIA series. ........ 136

12

S-. *- ,' . ,- , -¢ . - . - - . ,
LIST OF ILLUSTRATIONS (CONTD)

Figure Pag

88 Movable seat weight series, maximum

seat stroke . . . . . . . . . . . . . . . . . . 139

-8!s Movable seat weight series, maximum seat pan

z-axis acceleration. . . . . . . . . . . . . . 139

9" Movable seat weight series, Dynamic Response Index . . . . . . . . . . . . . . . . . . . . . 140

91 Movable seat weight series, maximum lumbar

spine axial load (compression) ........ 140

"92 Movable seat weight series, maximum lumbar

spine moment (flexion) . . . . . . . . . . .. 141

93 Test configuration (for stiffened seat frame

test) . . . . . . . . . . . . . . . . . . . . . 142

.94 Comparison of response with stiffened seat Sframe . . . . . . . . . . . . . . . . . . . . . 143

95 Measurement of seat cushion force-deflection

characteristics. . . . . . . . . . . . . . . . 144

96- Load-deflection curves for cushions used in

cushion spring rate series . . . . . . . . . . 145

.97 Comparison of seat pan z-acceleration for

various cushion stiffnesses. . . . . . . . . . 146

98- Comparison of pelvic z-acceleration for

- various cushion stiffnesses. . . . . . . . . . 146

99 Comparison of seat pan z-axis force for dummy - performance degradation tests. . . . . . . . . 147

: 100 Comparison of pelvic resultant acceleration

* for dummy performance degradation tests... 148

N

S 101' Test conditions for SOM-LA simulations to . determine the effect of footrest angle S-ndlocation ......... o.....,. 149

102 Tnput acceleration pulses for simulations -

..examining the effect of landing gear •I accelerition on seat and occupant response .. 151

10.03 Conditions for aircraft orientation

simulation . . . . . . . . . . . . . . .. . . 153

13

-w ~ -7
LIST OF ILLUSTRATIONS (CONTD)

Figure Pag

104 Dynamic test requirements for Test No. 1

of MIL-S-58095(AV) ... 9.... .. .. . 155

1⁄4>' 105 Nominal acceleration pulse for Test No. 1,

MIL-S-58095(AV). . . . . . . . . . ............ . 156

106 Minimal and maximum rate of onset conditions

for simulations of Test No. 1,

MIL-S-58095(AV). . . . . . . . . . . . . . . . 156

"LIST OF TABLES

Table

1 Program responsibilities ..... ...... 21

2 Variables examined during parametric study . . 32

3 Test matrix for CAMI rigid seat series .... 44

4 Dummy identification . ........ . ... 45

5 Test matrix for CAMI energy-absorbing

seat series. . .. . . . . . . . . . 46

: 6 Description of cu,.tioK, types . . . . . . . . . 48

7 Test matrix for NADC energy-absorbing seat

8 Test matrix for Simula energy-absorbing

seat series. . .... . . . . 49

9 Summary of cadaver test program. . . . . . . . 50

"10" Test matrix for modified anthropomorphic

dummy series . . . . . . . . . . ..... . . 54

* 11 Restraint system webbing .... . . ..... 62

12 Comparison of test results for peak input

acceleration series ...... . ....... 99

VA Comparison of results for velocity change Sseries 106

14 Comparison of results for rate of onset

series . . . .. . . . . . . . . . . . . 110

"14

-W *W
LIST OF TABLES (CONTD)

Table Page

15 Comparison of results for energy absorber

limit-load series . . . . . . . . . . . . . . 113

16 Measured spinal loads and moments for Wayne

State University modified dummy test series

K' with combined orientation. . . . . . . .. 117

17 Comparison of test facility baseline impact

conditions ............. ..... 121

18 Average peak accelerations and seat pan loads

for rigid seat comparison tests of 50thpercentile dummies (average of three tests). 126

19 Comparison of response for 95th-percentile

dummies in the baseline test configuration . . 127

dummy Summary percentile of the effect of anthropomorphic on peak acceleration

values . . . .I . . . . . . . . . . . . . . . 128

21 Summary of the effect of anthropomorphic

dummy percentile on spinal loads and

moments. I .I I . . . I I I . . . . . . . 1 128

22 Comparison of seat and occupant response for

dummies and cadavers . I I . . I I . . . .. 130

23 Comparison of seat stroke for cadavers and

dummies in the UH-60A crewseat . * . I I . . . 130

24 Comparison test results for seat orientation

series, rigid seat . . . . . . ... . . , . 133

25 Comparison of test results for seat orientation series with energy-absorbing Sseat I . I I I I I I I . I I I I 1 6134

26 Effect of seat orientation on lumbar forces and mcments I I I I I I 135

27 Comparison of results for ramped energy

absorber load series . ...... ..... 137

28 Comparison of test results for seat movable

weight series. . .. . . . I. . . I . .I. 138

29 Comparison of test results for stiffened seat

frame series . . .. . . .. . . . .... 143

is•

6 -

4-, ', ' • * "'>°• ,,•, -w' %- 2, . o w ' 'w,% 

",' *,5. .. ,'*•, ." '6.I %,',,.," , , , ,,.% ..- '6,. '. ,, 6~ " . ." " , ". .* ,' **. ..
LIST OF TABLES (CONTD)

Table Page

"30 Comparison of test results for cushion spring

rate spries. . . . . . . . . . . . . . . . . . 145

31 Comparisou of SOM-LA simulations to determine

the effect of footrest angle and location on

seat and occupant response . . . . . . . . . . 149

32 Comparison of predicted seat and occupant rebponse for input deceleration pulses with

and without landing gear ramp. . . . . . . . . 152

33 Conparison of seat and occupant response for

vertical and horizontal test condition

s..mulations. . . . . . . . ... . . . . . . . 153

34 Matrix of SOM-LA simulations for evaluation

of sensitivity to test pulse tolerances

'-: .t 11.5- f'nd 14.5-G energy absorber limit

load settingss. . ... .. . .. .. . .. .. 157

S35 Comparison of predicted seat and occupant

response to MIL-S-58095(AV) Dynamic Test 1

with 14.5-G E/A liwIt load . . . . . . . . . . 157

"*" 36 Comparison of predicted seat and occupant

response to MIL-S-58095(.AV) Dynamic Test 1

with 11.5-G E/A lin.it load .. . . . . . . . . 157

..44

" 16

.#,,

, , -. . . . .. , , . . . W

7 4 v, 4,I **
1.0 INTRODUCTION

The latest generation of U.S. Army helicopters possesses unprecedented crashworthiness, as pointed out in Reference 1. These

aircraft are equipped with many crashworthy features, including

seats designed to provide efficient restraint in all loading directions and energy-absorbing stroke in the vertical, or +G

direction. The seats are designed to comply with existing sriteria that were developed and documented in 1971 (References 2,

3, and 4); although these seats are far superior to any prior

systems, there are several areas of uncertainty in the design

criteria that require additional research to enable further prog- %i ress to be made in the hardware.

Knowledge concerning human tolerance to +G acceleration is extremely limited and constitutes the greateit uncertainty in seat

design. In fact, little new information concerning human tolerance to acceleration in this direction has been developed in many

years. Although extensive effort has been expended on the critical

areas of human head and neck response (U.S. Navy) and the effects

of restraint system variables on acceleration loads in the longitudinal and lateral directions (U.S. Air Force), essentially no

effort has been directed in this other very critical direction.

The need for additional knowledge of +G response is affirmed when it is considered that aircraft occupants can withstand the full

95th-percentile survivable crash acceleration conditions in the

lateral and longitudinal directions with no energy absorption,

only proper restraint (a complex problem when related to the head),

but cannot tolerate the 95th-percentile vertical crash pulse without energy absorption.

0.,

1. Singley, G. T., III, and Desjardins, S. P., "Crashworthy

Helicopter Seats and Occupant Restraint Systems," in Operational Helicopter Aviation Medicine, AGARD Conference Proceedings No. 255, North Atlantic Treaty Organization, Advisory Group for Aerospace Research and Development, Neuilly

sur Seine, France, May 1978.

2. Crash Survival Design Guide, Dynamic Science, A Division of

Marshall Industries, USAAMRDL Technical Report 71-22, Fort

Eustis, Virginia, Eustis Directorate, U.S. Army Air Mobility

Research and Development Laboratory, 1971, AD 733358.

3. Military Standard, MIL-STD-1290(AV), Light Fixed - RotaryWing Aircraft Crashworthiness, Department of Defense, Washington, D.C., 25 January 1974.

4. Military Specification, MIL-S-58095(AV), Seat System: Crashworthy, Non-Ejection, Aircrew, General Specification For, Department of Defense, Washington# D.C., 27 August 1971.

17

. I - .7.

S 4 ~ 4 4
The last program of significance to investigate the variables associated with vertical (+G ) acceleration of aircraft occupants

was sponsored by the U.S. Irmy Air Mobility Research and Development Laboratory, Fort Eustis, Virginia (now the Applied Technology Laboratory) in 1969 and 1970. The results of this program

are presented in Reference 5. The goal of this earlier program

was to evaluate the meager information then available and to develop an achievable criterion to guide the design of crashworthy

crewseats for U.S. Army aviation. This task was accomplished,

and the resulting criterion has now been in existence for the

past ten years.

Techniques used for the design and evaluation of energy-absorbing

seating systems are explained in detail in References 1, 2, 4,

and 6; however, in summary, the Eiband human tolerance data with

upper level ejection seat criteria superimposed (as presented in

the Aircraft Crash Survival Design Guide, Reference 7) was taken

for the upper limit of tolerable acceleration in the +Gz direction. Tests and analyses were then conducted to establish the

force level necessary for the vertical energy-absorption system

"to limit the acceleration excursions of the seat pan to magni-

"tudes of less than 23 G for time durations in excess of 0.006 sec,

as dictated by the tolerance data. The test results indicated

that if the energy-absorbing mechanism were set to proviJe controlled deformation at a force computed using a 14.5-G load factor,

the desired result could be achieved. Test data supporting the

conclusions of this analysis are presented in Reference 5.

In seat tests conducted during the ensuing years, a characteristic

seat pan z-axis acceleration response was observed. In this characteristic curve, the seat pan acceleration rises sharply during

the onset of the input pulse, then drops rapidly, sometimes pass-

- ing through zero. It then rises sharply again and forms a secondary spike before damping out around the load factor used in the

3T Desjardins, S. P., and Harrison, H., The DesignL Fabrication,

and Testing of an Integrally Armored Crashworthy Crewseat,

Dynamic Science# Division of Marshall Industries; USAAMRDL

Technical Report 71-54, Eustis Directorate, U.S. Army Air

Mobility Research and Development Laboratory, Fort Eustis,

Virginia, January 1972, AD 7421733.

6. Desjardins, S. P., and Laananen, D. H., Aircraft Crash Sur-

". vival Design Guide, Volume IV - Aircraft Seats, Restraints,

"•Litters, andPadding, Simula Inc., USARTL-TR-79-22D, Applied

Technology Laboratory, U.S. Army Research and Technology

Laboratories (AVRADCOM), Fort Eustis, Virginia, June 1980,

AD A088441.

7. Laananen, D. H., Aircraft Crash Survival Design Guide, Volume II - Aircraft Crash Environment and Human Tolerance,

Simula Inc., USARTL-TR-79-22B, Applied Technology Laboratory,

U.S. Army Research and Technology Laboratories (AVRADCOM),

Fort Eustis, Virginia, January 1980, AD A082512.

' -. 18

,w w w w vt~ ~.
design of the energy-absorbing system. In most of the tests conducted during the time period between 1971 (when the criteria were

4:-1 established) and the present, the secondary spikes have exceeded

the criteria limit of 23 G and have been a source of concern. One

* question is whether the secondary spike is a natural response of

the seat and occupant spring-mass system, or if it is caused by

some external source. Also, it is not known whether the acceleration spike is hazardous to the seat occupant.

The need for answering these questions was confirmed in August

1978, during qualification testing of the U.S. Army's Black Hawk

helicopter crewseat, when the secondary acceleration spike once

again exceeded the criteria limits. At that time it was decided

to research the data available and to attempt, through analysis,

to determine whether the secondary spike did, in fact, increase

the hazard to the occupant. The results of this analysis indicated that the secondary spike is a natural response of the seating system and, in itself, does not increase the hazard to the

occupant. However, the analysis also indicated that the crash

pulse might be hazardous to the occupant at times other than during the secondary spike. It was further concluded that the criterion based on seat pan acceleration is not sufficiently comprehensive, and that additional research should be immediately

initiated to both establish the effects of system variables upon

seat and occupant response, and expand knowledge in the area of

human tolerance to decelerative loading in the +G direction.

The data gathered from these programs constitute I necessary prerequisite to establishing a more comprehensive set of criteria

controlling the design of crashworthy crewseats.

Under the leadership of the U.S. Army Applied Technology Laboratory (ATL), Fort Eustis, Virginia, and the U.S. Army Aeromedical

Research Laboratory (USAARL), Fort Rucker, Alabama, a multiservice effort was initiated with goals of performing the research necessary in these areas. The overall program is illustrated in Figure 1. The research efforts were designed to provide at least a minimum level of information in each area and to

meet the following objectives:

9 Establish the sensitivity of seat and occupant response

to system variables.

* Determine the effect on system performance of the type of dummy being used for testing and establish an appropriate standardized dummy for seat system evaluation.

4' * Investigate the performance of the seat with an occupant

more nearly representative of the operational occupant than are anthropomorphic dummies. Cadavers were to be

used for this investigation.

S* Establish, through dynamic testing, additional information concerning human tolerance to acceleration loads

in th*, G direction.

19

.' *•F• 4. . .; '-.' .•• . . ... . . .......... . ... . .....
SEAT AND DUMMY VARIABLE TESTING

4", I II •iSENSITIVITY

SEAT AND OAIMPROVEDA I

I CADAVER IITSEAT U DESIGNE "'" / TESTING iAND TEST CRITERIA

... • ~HUMAN TLRNE'

I TOLERANCE TOLEANC

":" TESTING LIMITS.

•'•"..<. JDETERMINATION OF [•-

• STANDARDIZED

-J | STANDARD AIRCREWMAN,, MILITARY ANTHROPOMORPHIC

I PROPERTIES

Figure 1. Overall program elements leading to improved

seat design and test criteria.

The effort to establish a standardized test dummy is being coordinated by USAARL and involves an ad-hoc committee consisting

of representatives from interested Government agencies and the

private sector. Research to increase the knowledge of human tolerance to acceleration loads in the +G direction is being per- formed by the Naval Biodynamics Laborafory, Michoud, Louisiana,

with support and sponsorship of USAARL. The other research is

being coordinated by Simula Inc., Tempe, Arizona, with test and

advisory support by many of the other involved organizations, and

"with the sponsorship primarily of the ATL and USAARL. The overall

"program is summarized in Table 1.

This report describes the program of seat and dummy testing, as

well as subsequent analyses, to establish the sensitivity of the

system to a number of variables. Chapter 2 presents background

information on the need for energy-absorbing seats and the response of the combined seat/occupant system. Chapter 3 outlines

" the research program in detail. Procedures for testing and analysis are described in Chapters 4 and 5, respectively. Results of

the sensitivity analysis are described in Chapter 6, and conclusions and recommendations, particularly with respect to improvements in seat criteria, are presented in Chapters 8 and 9.

"20

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TABLE 1. PROGRAM RESPONSIBILITIES

Program objective Agency/Organization Responsibility I Sponsor '

Establish the sensitivity Simula Inc., Overall responsibility Applied Technology Laboratory

of vertically (z-axis) Tempe, Arizona for program direction. US Army Research and Technology

stroking seating systems and Frovide all seat hard- Laboratories (AVRADCOM), Fort

improve the criteria for ware, data analysis, Eustis, Virginia

their design. data synthesis and

interpretation, and re- port results. Also,

perform testing,

data acquisition, reduction, and analysis.

Civil Aeromedical Perform dynamic testing Federal Aviation Administration

Institute, Oklahoma supportz provide inCity, Oklahoma strurentation and test

dummies, record, reduce, and submit data; .0;

and advise on other

programs.

"Naval Air Develop- Provide dynamic test U.S. Navy

ment Center, War- support, instrumentaminster, Pennsyl- tion, data acquisition,

vania and reduction,

Determine the tolerance Simula Inc., Overall program respon- U.S. Army Applied Technology

threshold limit-load setting Tempe, Arizona sibility. Provide test Laboratory, Fort Eustis,

for the energy-absorbing seat hardware, estab- Virginial U.S. Army Aeromedical 7.

mechamism that would not lish instrumentation Research Laboratory, Fort Ruckor,

cause spinal injury in ca- requirements, perform Alabama, Federal Aviation

davers. Also, determine the final analysis of data, Administration Technical Center,

response of an energy- Interpret and report Atlantic City, New Jerveyi U.S.

absorbing crews'tat with a results. Air Force, 657Uth Aerospace

human cadaver as an occupant Mladical Pesearch Laboratory,

for comparison with the re- Wright-Patterson Air Force

sponse using an anthropomor- Bo, Ohio

dumOhio ahic 

Bioangineering Acquire cadavers, in- Sam as above

Center, Wayne State strument the cadavers,

University perform the kiynai-ic

tests; a&quire, record

and reducs test data.

Standardize the seat test Ad-hoc cacmefttee Acquire and select por- U.S. Army heromedical Research

Ndumm. consisting of rep- tinent data for use in Laboratory, Fort Itacker,

resentatives from establishing the plyei- Alabama

industry and Gov- cal parameters of the

ernznt military aircrawman

for use in subsequent

development of a otandarditzd test dummy,

Improve knowledge of the U.S. Naval 8io- Overall pro.ram room|n- U.S. Army Acromadical Wa•arch

tolerance of human aubjeeta dynamics Laboratory, sibility including per- Laboratory, Fort Ruckor,

to * accelerative loadinq. kihoktd, Louisiana forming tests of live Alabasma U.S. Naval Biodynamics

human volunteers, tel- lAboratory-, ichoud, Louislana

lowed by surrogates.

Acquilsition of data,

analysis and prsen•tation of results.

,..1

21

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L' 2.0 HUMAN TOLERANCE TO ACCELERATIVE LOADING

The tolerance of the human body to impact forces depends on a

number of variables, including characteristics of the individual

such as age, sex, and general state of health. Military systems

can be expected to be used by personnel who are primarily younger

and in better physical condition than the general population for

which much tolerance data has been obtained. Thus, in some cases,

a degree of conservatism may be built into the application of tolerance criteria in designing Army aircraft. However, whole-body

tolerance criteria have been based on experiments involving subjects seated with "correct" upright posture. Because a helicopter

pilot is unlikely to maintain such posture in flight, particularly

when near the ground, tolerable levels of such variables as +G

acceleration may be significantly reduced under actual crash canditions.

2.1 HUMAN TOLERANCE INVESTIGATIONS

The magnitude and duration of the applied accelerative force have

definite effects on human tolerance, as shown in Figure 2. As

indicated by this curve, a spineward chest-to-back (-G ) accelerative force of 45 G has been tolerated voluntarily by 9ome subjects when the pulse duration is less than 0.044 sec. Under similar conditions, when the duration is increased to 0.2 sec, the

tolerable magnitude is reduced to about 25 G. Accordingly, Figure 2 (Reference 8) shows that the tolerable limits on acceleration loading are a function of duration.

The whole-body tolerance data displayed in Figure 2 were collected

for a variety of full-torso restraints and, in some cases, head

restraint. With less optimum restraint, some debilitation and

injury will occur at this acceleration level, or, in other words,

the tolerable level will be reduced.

The human body is able to withstand a much greater force applied

perpendicular to the long axis of the body in a forward or backward (G ) direction than applied parallel to the long axis (G),

as demonstrated by a comparison of the curves in Figures 2 ang 3.

"A primary reason for the significantly lower tolerance to headward (+G ) loading is the susceptibility of the lumbar vertebrae,

which mult support most of the upper torso load, to compression

fracture. Also, the skeletal configuration and mass distribution

"of the body are such that vertical loads cannot be distributed

over as large an area as can loads applied forward or backward

(G.). These vertical loads, therefore, result in greater force

peJ unit area than do sternumward or spineward loads. Finally,

along the direction of the long axis, the body configuration

allows for greater displauement of the viscera within the body

8. Eiband, A. M., Human Tolerance to Rapidly Applied Accelerations: A Summary of the Literature, NASA Memorandum 5-19-59E,

National Aeronautics and Space Administration, Washington,

.4 D.C., June 1959.

22

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0444

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441

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0

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5* 23

044

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-

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tp (A 44

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cavity. Forces applied parallel to the long axis of the body,

headward or tailward (G ), place a greater strain on the suspension system of the viscira than do forces applied sternumward or

spineward (Gx), thereby increasing the susceptibility of the

viscera to injuries.

Human tolerance to lateral (±G ) impact acceleration is not well

defined, as a rather narrow raXge of pulse durations has been explored in tests. The U.S. Air Force conducted a series of tests

with human volunteers in which several orientations were used

(Reference 9). In 32 tests to evaluate ±G impact vectors, there

were no adverse subjective reactions to acgeleration magnitudes

up to 22 G with a rate of onset of 1350 G/sec and velocity changes

"to 19.3 ft/sec. Subjects were restrained by lap belts, torso harness, and leg restraints and supported by a contoured pad. In a

series of tests conducted by the Royal Air Force (Reference 10),

a conventional military aircraft seat and five-point restraint

system was used with human subjects tolerating up to 17.1 G with-

'1 out irreversible injury.

No end points for human tolerance to lateral impacts were proposed in the reports of these experiments, but a reasonable conclusion from these data is that the human survival limit with

conventional military aircraft restraint is probably on the order of 20 G applied for a duration of 0.1 sec.

M..2 TOLERANCE-BASED CRITERIA FOR SEATS

The desired function of crashworthy crewseats is to protect occupants from the decelerative crash load hazards with severities up

to and including those of a 95th-percentile survivable crash pulse.

Reference 7 reports that 95 percent of survivable accidents involved average accelerations of less than 15 G in the longitudinal

direction, 24 G in the vertical direction, and 16 to 17 G laterally, depending on aircraft type. For a triangular pulse, the

peak values would be twice these levels or 30 G longitudinal, 48 G

vertical, and 32 to 34 G lateral. Comparing these values with

the tolerable levels reported above, it appears that a properly

restrained human can withstand the lateral and longitudinal loads

associated with the respective 95th-percentile survivable crash

"pulses. Therefore, no load attenuation is necessary, or desirable for crewseats due to space limitations of most cockpits, in

9". IriNley, J. W:; Weis, E. B., Clarke, N. P., and Temple,

W. E., A Study of the Effect of F1ive Orientations. of the

Acceleration vector on Human Response, AIRL Memo M-28, Aerospace Medical Resarch Laboratory* Wright-Patterson Air

Force Base, Ohio, 1963.

10. Reader, D. C., The Restraint Afforded by the USAF and Proposed RA.P IAM. Se;atHarnesses for the F-111 'nder High Forward

and Lateral Decelerations, IAt1 Report'421, IStitute of AviaEtion Aedic;ne, Farnborough, Hampshire, United Kingdom, 1967.

25

4W

S. %*~* .a6.,S ~S
those directions. In the vertical direction, however, crash energy absorption and force attenuation are required since the human spine can withstand only a fraction of the load associated

with the 95th-percentile survivable vertical crash pulse.

Various criteria for design of the vertical energy absorption system have been established. Basically, they involve establishing

the force level at which the energy-absorbing system should stroke

to limit the load in the human spine to a tolerable magnitude.

This is accomplished by determining some "effective" weight for

the occupant (effective weight is less than the total body weight

because part of the body weight - the weight of the feet and part

of the legs - is carried by the floor of the aircraft rather than

by the seat), adding to it the movable weight of the seat, and

multiplying the sum by the load factor (G) determined to be tolerable. The criteria included in the 1970 edition of the Crash

* : Survival Design Guide (Reference 11), Paragraph 3.3.3.1, state

that *This effective weight plus the weight of the movable portion of the seat should be decelerated at an average level of

18 G or less." Average G is further defined as "The total velocity change of the seat pan (Av) divided by the total pulse duration of the seat pan." (Notice that no mention was made of the

shape of the deceleration-time history but, rather, only the aver- A

age deceleration established by dividing the area under the curve

, by the duration.) However, some researchers felt that perhaps the

"18-G average value was too high asid would still allow occupant injury. Consequently, in 1969 and 1970, a limited research program

was sponsored by the U.S. Army Aviation Materiel Laboratories and

conducted by Dynamic Science (AvSER Facility) to try to reevalute,

verify, and refine the criteria for the design of crashworthy

* armored crewseats. One recommendation of this program, which is

documented in Reference 5, was the use of a limit-load factor of

14.5 G for integrally armored crewseats.

Again, the objective of the criteria was to provide design guidance to help in the development of seats that would limit occupant spinal loads to tolerable levels during crash pulses of

severities up to, and including, the 95th-percentile survivable

"vertical crash pulse. The acceptable acceleration levels speci- K

fied by Reference 4, which had been based on the Eiband tolerance

"data and the ejection seat design band (Reference 9), were selected for use. As shown in Figure 4, this criterion limits the

headward (&G ) acceleration imposed on the occupant to 23 G for

"time durations in excess of 0.006 sec, where time durations are

defined as the length of the plateau between the onset and the

- offset portions of the accelerotion-time histories experienced by the seat pan. The research program was structured to determine and establish the stroke distance and load for the energy-

,• absorbing mechanisms to limit the accelerations imposed on the

11: Crash 9Survival Design Guide, Dynamic Science, A Division of

9arshal rindustries, USAA6fRDL Technical Report 70-22, Fort

Eustis, Virginia, Eustis Directorate, U.S. Army Air Mobility

* Research and Development Laboratory, 1970, AD 695648.

26

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-- '"DO 40-AE OF I ' ' [ I l

UJ

23 M ;AXIMUM

ACCEPTABLETPULS

,,; UU 20 - _ AREA OF "A CPAL

DURACCEPTABLE PULSE DURATION-MAGUNITUDI

. ,0 - , , , , , , u 8 ' -P-LA T EA U

• 'D 6 - DURATION

• U, PLATEAU

b•.. O 4 -IG MAGNITUDE

Fiur 4. MaiuMcAthl etclsa

4 7 .. . . ,

":':0IT1 T2 T

;- • ! T T= TIME

.001 .002 .004 .006 .01 .02 .04 W0 .1 .2 .4 .6 .8 2 4 6 8 10

DURATION OF UNIFORM ACCELERATION - SEC

.- •Figure 4. Maximum accept,ý,hle vertical seat

':" acceleration (from Reference 4).

S occupant to values less than those estimated to produce injury

according to this criterion. It was thought that if the load at-

*: tenuation system on the seat would provide such performance, then,

within the limits of the seat stroke, the occupant would have as

good a chance of avoiding injury as an individual using an ejection seat:,

The 1970 research program consisted of analysis followed by a

series of eight dynamic tests to develop the criteria for determining the limit load that would produce the desired performance

within the prescribed minimum practical stroking distance of

12 in. Because of funding limitations, the prograt., conducted was

not sufficient to incluie a comprehensive investigation of the

influence of many of the variables, such as energy absorber load- :7

versus-deformation characteristics, anthropomorphic dummy variables, input pulse rate of onset and shape, bottom cushion loaddeflection characteristics, and seat spring rate. However, the

tests that were run indicated that it should be possible to maintain the seat pan acceleration within the limits of the criteria

used for ejection seats.

In the years fo' lowing development of the cri,;eria, several crashworthy seats were developed. As these seats were dynamically

tested, a characteristic acceleration-time history was displayed

(see Figure 5). This characteristic shape had been evident in

4. 27

U Uq 40 -W W %* WA -,
45 "-1 'Seat pan

.-4 o-o-o-o--o Dummy chest - .. - Dummy pelvis

-34 30

__ •4J

__'

0 /

15 •--

e-4

0.00 0.05 0.10 0.15

Time (sec)

Figure 5. Typical resp•onse of seat pan, dummy chest, and dummy pelvis to vertical crash loading.

-.. the criteria development test data as well, but the magnitudes of the peaks and valleys had been lower, As mentioned in the Intro- duction, the shape of the seat pan acceleration-vers..s-time his- tory includes a hiyh initial spike followed by a deep notch that sometimes passes-.through zero, :actually changing sign. This notch is followed by a second high pea), followed, in turn, by various wavefokms damping out and.usually centering around the design limit-load factox of the system. This characteristic shape is apparent in all dynamic tests of the various crashworthy seat de- signs, including the Dynamic Science criteria development seat . (Reference 5), the Simula Inc. UTTAS prototype seat (Reference 12),

" " .2. mzalski; L., and-'singley, G. T., III, Joint Army/Navy Test Progam for Black Hawk Seat.Systems, Report NADC-79229-60 (Draft) aval Air Development Cepter, Warminster, Pennsyl- vania, 1979.

28'

* VýI4P9r~~ # '
the Simula Inc./Norton Co. Black Hawk production seat (Reference 13), and the ARA Inc. JA/N SEAT (Reference 14).

":" 2.3 SEAT/OCCUPANT DYNAMIC RESPONSE

The explanation of the formation of the characteristic accelera-

"tion waveform is associated with the inherent dynamic response of

the seating system and its occupant. Total coupling of the seat

and its occupant is not achieved since the occupant consists of

distributed masses connected by nonrigid body members, such as

the vertebral column and neck in humans or elastic structural members in anthropomorphic dummies. Further, the interface between

the occupant and seat is buttocks flesh and a cushion which do

not form a rigid connection between the occupant and the seat pan.

While stroking toward the floor, the energy absorber exerts a constant upward force on the moving seat mass. (Reference 4 requires

that the energy absorber force be equal to the sum of the effective weight of the occupant and movable part of the seat multiplied by the load factor of 14.5 G.) The acceleration measured

on the seat is then determined by the magnitude of the downward

force applied by the occupant.

Idealizing the dynamic system of seat and occupant, all springs

in the system are initially unloaded prior to impact, as shown in

Figure 6(a). During the initial phase of the impact, the downward inertial load applied by the occupant to the seat pan is low.

The cushion must be compressed before this load can become significant and a force acts to slow the occupant. It is during this

time period that the seat pan acceleration characteristically grows

from zero to a relatively large value as the energy-absorbing systerm acts to slow the small mass of the seat pan. As a result, the

seat pan peak acceleration can reach magnitudes as great as the

input floor acceleration.

* In the second phase, each of the body parts is compressed, applying & downward force on the seat as shown in Figure 6(b). This

downward force decreases the seat acceleration while the upward

reactive force increases the body acceleration. It is possible

that the downward inertial force applied by the body can equal

the force required to stroke the energy-absorbing system. The

"net acceleration on the seat pan can drop to zero, or, in some

S cases, actually reverse direction.

g,

4 13. Dummer, R. J., Qualification Test Report, 613-1787 COOL

"Qualification Testing of Armored Crashworthy Aircrew Seat,

RA-305252-1, for Sikorsky Aircraft Contract 576344, Norton

Company, Worcester, Massachusetts, October 1978.

" 14. Chandler, R. F., Dynamic Test of Joint Army/Navy Crash-.4orthy

Armored Crewseat, Protection and Survival Laboratory, Memor-

"andum Report AAC-319-80-2, Civil Aeromedical Institute, Mike

Monroney Aeronautical Center, Federal Aviation Administration,

Oklahoma City, Oklahoma, 25 April 1980.

29

.4I , 0 "0 0 . U V ~ '

• " ': :' " ? "• " • " , " '•" ." .. .' .. . ."'-... . .. . . . . .. . ... ... . . .. ... . ."- . . .. .... . . . . .... i , i .i
1 F 1 HEAD

-S* .•..t

(a) INITIAL CONDITON. UNLOADED

M 2 CHEAD

•= ~-COMPRESSED SPRINEG

M- PELVIS

SSEAT

44M

E IONITIS RESONDINTODCLRTION, L LOAADD

HEAD

~+

.13T'LL

Cc) HIGH-DECELERATION LOAD, STEADY-,STATE CONDITION

4 *I,,

Figure 6. Spring-mass representation of seat-occupant system.

30

¶.4. **• t:" r S '• *IP' . ... " * * ...... l ..... . ***. *f..... qt... •. , " , ' "" . . . .. .. , ... . 

ft~~~ 

4 ~ M44.. S -4'

4'S~~~CRS LOAD,~ .; 4 ,.

(b) ONSE OF DEEERTO LODWHRIPLI
V.o

The accelerations of all components in the final phase eventually

approach a common steady-state level, and the seat and occupant

move together at the design acceleration level for the energy absorber setting (Figure 6(c)).

Referring to Figure 5, it is important to note that the peak accelerations of the seat pan do not necessarily coincide with peak

accelerations of the occupant pelvis or chest, and thus are not

necessarily hazardous to his safety. This deficiency was apparent

when the existing criteria were developed, but they were adopted

since, from an overall standpoint, the criteria had limited the

incidence of injury in ejection seats. However, the Eiband tolerance data used in the criteria do not consider the seat pan acceleration excursions from the average. These data were smoothed

in much the same manner as used for calculating average G described in the second paragraph of Section 2.2.

Consequently, the more comprehensive study, described in the Introduction, was undertaken to develop the data necessary to better

understand the dynamic response of the seat and occupant systems.

This understanding can then be quantified and used to establish

more comprehensive design criteria to enable the design of seat- ing systems that will reduce the crash hazard to the occupant.

31

.' .....
3.0 PARAMETRIC STUDY

.* The goal of the program described in this report was to provide

the necessary data to more effectively use the available stroking - distance for energy-absorbing seats, thus reducing the probability of occupant injury due to vertical decelerative loading. At

the initiation of the program, it was determined that the available computer simulation methods were not sufficiently validated

". for this type of work; therefore a parametric study had to be

"completed using dynamic impact tests. This chapter describes the

"development of the parametric test matrix to provide the data base

for optimizing energy absorber design.

- "3.1 VARIABLES EXAMINED

Thirteen variables were chosen to determine their effect on seat

"and occupant response. These variables were chosen based on results from the original Criteria Development Test Program (Reference 5) and experience gained in the development and testing of

S/: the UTTAS prototype armored crewseat (Reference 12). The 13 var-

"* iables, listed in Table 2, fall into three major categories as

*! described in the following sections.

TABLE 2. VARIABLES EXAMINED DURING

"*" PARAMETRIC STUDY

* Test facility impact conditions

* Magnitude of input acceleration

- *9 Velocity change

, Rate of onset of input acceleration

, 4$ Dummy type

"" Dummy percentile

"" Cadavers versus anthropomorphic dummies

* Energy absorber limit load

". Ramped energy absorbers

"e Movable seat weight

, Seat frame stiffness

* Seat cushion stiffness

* Seat orientation to impact vector

32

• ,,. * - . . - . * ,. 3 *-
3.1.1 Impact Conditions

"* The impact conditions to which the cockpit area of an aircraft is subjected-are controlled by the velocity and attitude at impact,

as well as the design characteristics of the fuselage section impacting the ground. Seat orientation to the impact vector is a

function of the aircraft orientation at impact and the design angle of the seat, which is usually pitched backward in the aircraft to enhance seated comfort. The velocity change is controlled mainly by the input velocity; however, the restitution

characteristics of the fuselage sections crushed during impact

can allow the aircraft to attain a significant rebound velocity. The velocity change is equal to the sum of the magnitudes of the

initial velocity and the velocity achieved during rebound. The

load-deformation characteristics of the fuselage sections crushed

during impact will also determine the rate of onset, which is a

measure of the rate at which the impact loads can build up in the

structure, and the maximum impact load in the structure which is

characterized by the peak input deceleration.

In a dynamic impact test of an occupant/seat system, crash forces

are simulated by inertial loading that is induced either by acceleration from rest or deceleration from an initial velocity. Although it was originally thought that these conditions were equiv- alent, Hearon, et al. (Reference 15) have recently shown that

there is an inherent difference in the two methods due to dynamic

preload. In the parametric study, all four test facilities used

the deceleration technique to simulate inertial loading. Each

facility had a unique method of simulating the load-deformation

characteristics of a crushing structure. Seat and occupant response for the various test facilities was the final variable in

this area.

3.1.2 Occupant Characteristics

Energy-absorbing seat development and testing is currently conducted using anthropomorphic dummies. The applicable military

standards do not specify a standard dummy, and the available test

dummy types can differ significantly in their response characteristics depending on the original design goals for the dummy. None of the existing dummies were designed or optimized for vertical impact testing. Therefore, an obvious area of interest in

this program was to compare the characteristic responses for anthropomorphic dummies used in dynamic testing. Also of interest

was a determination of the effect of dummy weight (described by percentile of the U.S. male population) on seat and occupant response.

15. Hearon, B. F., Raddin, J. H., and Brinkley, J. W., "Guidance

for the Utilization of Dynamic Preload in Impact Injury Prevention," in Impact Injury: Mechanismsg Prevention and Cost,

North Atlantic Treaty Organization, Advisory Group of Aerospace Research and Development, Neuilly sur Seine, France,

April 1982. 4

33

:%:
The use of anthropomorphic dummies as human surrogates for dynamic

testing implies that the response of dummies and humans is similar. However, this is not a commonly held belief among the re-

"searchers in crash testing. The U.S. Army has undertaken an effort to examine human occupant response and tolerance to vertical

impact loading through the use of cadavers. This work is presently being conducted under Contract DAAK51-79-C-0016 (Reference 16). In the parametric study, preliminary results from the

cadaver testing program will be used to compare the effects of

cadavers and anthropomorphic dummies on energy-absorbing seat response.

3.1.3 Seat Design Parameters

A number of seat design parameters directly affect response of

the occupant and seat. Although development of flight-weight

pcototype seats has shown some variation in these parameters,

this study attempted to examine the anticipated design range to

determine the sensitivity of occupant response.

The energy absorber limit load is of particular interest because

it directly affects the magnitude of inertial loading in the body.

For a constant-load energy-absorbing system, the limit load is

normally expressed as a multiple of the effective weight of the

occupant and movable seat weight, which is called the limit-load

"factor. This limit-load factor is set to prevent the accelerative

loading in the spine from exceeding tolerable levels. Ideally, the

limit load should be set at the highest tolerable level to maximize

the use of the available stroke distance in severe crashes. Data

gathered from the cadaver test program described earlier and from

field experience with production energy-absorbing seats should

allow the limit-load factor to be optimized. In this study the

• j effect of energy absorber limit load on seat stroke and occupant Al response will be examined. Also, analysis has suggested that,

under certain conditions, an energy-absorber load that increases

throughout the impact may actually reduce the dynamic overshoot

in the body. This variable was examined by building special

"ramped" absorbers with the desired load characteristics.

Ir. Three additional seat design parameters were examined to determine

.' their effect on body dynamics during the crash sequence. These

parameters were the weight of the stroking portion of the seat,

the longitudinal stiffness of a vertically stroking seat, and the

load-deflection characteristics of the seat pan cushion.

3.2 TEST FACILITIES

Four test facilities were used to complete the parametric test

matrix: the FAA Civil Aeromedical Institute (CAMI), the Naval

"i[i 16. Contract No. DAAK51-79-C-0016, "Hardware, Test Support, and

Enginering Services," Applied Technology Laboratory, U.S.

Army Research and Technologies Laboratory (AVRADCOM), May 9,

1979.

34

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Air Development Center (NADC), Simula Inc., and Wayne State University Bioengineering Center (WSU). The majority of tests were

conducted at CAMI due to their ability to conduct laboratory-type

testing in a controlled, indoor, horizontal decelerator facility.

NADC and Simula, both vertical drop tower facilities, were used to

repeat key tests to determine the effect of input pulse shape and

facility type (horizontal decelerator versus drop tower). Wayne

State University conducted the cadaver test program and a series

of dummy tests using a horizontal decelerator.

The four test facilities and sample input pulses for the nominal

input condition, shown in Figure 7, are described in the following

sections.

50--

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3.2.1 FAA Civil Aeromedical Institute (CAMI)

The CAMI facility uses a guided sled, shown in Figure 8, pulled

along a set of horizontal rails by a falling weight. The sled

achieves the desired initial velocity approximately 5 ft from the

impact point. During travel through the final 5 ft, the sled

moves at a constant velocity and passes through a velocity trap

"placed 2 to 4 in. before the impact point. At the impact point the sled encounters multiple stages of 1/4-in, steel wires which

are plastically deformed as they are pulled between dies. The

decelerating mechanism is shown before and after a test in Figure

9. Up to four wires can be placed in 42 successive stations, and

the numbers and locations of wires can be varied to obtain a wide

range of input pulse shapes. A -..

The CAMI deceleration pulse is shown in Figure 10(a). It is characterized by a smooth shape with only about five percent of the

velocity change occurring in rebound.

3.2.2 Naval Air Development Center (NADC)

The NADC facility shown in Figure 11 uses a drop cage, guided in

a vertical tower, which is accelerated to the desired velocity by

gravity. The initial velocity is measured as the average velocity of the sled during the 6 in. of cage travel prior to impact.

The decelerating mechanism is a Van Zelm-type which uses steel

straps pulled over rollers to absorb the crash energy, as illustrated in Figure 12.

The baseline input pulse shown in Figure 10(b) has a distinct oscillation with a short duration peak followed by a reduction below zero and then the major deceleration pulse. Approximately

ten percent of the velocity change is due to rebound.

3.2.3 Simula Inc.

Tests conducted by Simula used the Arizona State University drop

tower shown in Figure 13(a). The decelerating mechanism is a

pyramid-shaped paper honeycomb stack designed by Simula to provide the necessary deceleration characteristics by varying the

shape and qualities of each honeycomb layer, and is shown in Fig- ure 13(b). A comparison of the Simula and CAMI baseline pulses

in Figures 10(c) and 10(a), respectively, shows that they have

similar features. Inherently, a significant rebound occurs when

air is trapped as the honeycomb stack crushes, and in the baseline

pulse approximately 25 percent of the velocity change is due to

rebound.

3.2.4 Wayne State University (KwU)

The WSU WHAM IV test facility uses a horizontal sled accelerated

by pneumatic pressure to within 5 ft of the impact point. During ! the ensuing constant-velocity phase and prior to impact, a magnetic proximity sensor is used to estimate average velocity over a

1-ft distance. The decelerating mechanism is a hydraulic cylinder

"36

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"ation pulse has a smooth shape, it is characterized by a rate of

onset significantly higher than the CAMI or Simula facilities. A

minimal amount of the velocity change is due to rebound.

3.3 PARAMETRIC TEST MATRIX

A series of 62 dynamic tests was conducted to provide data to determine the effect of the 13 variables. Fifty tests were conducted at CAMI, of which 23 used a rigid seat and 27 used a pro- duction energy-absorbing crewseat. NADC and Simula conducted

nine and three tests, respectively, all of which utilized the

energy-absorbing seat. Tables 3 through 8 describe the test conditions for all of the tests conducted at CAMI, NADC, and Simula.

3.4 CADAVER TEST PROGRAM

The cadaver test program was conceived to provide a design base

of human body tolerance data for optimization of energy-absorbing seating systems. The main objective is to determine the highest

energy absorber limit-load setting that can be tolerated by the

human body. The limit-load threshold is being evaluated by conducting dynamic tests with cadavers and assessing the spinal condition following the test. A major technical difficulty yet to

be overcome in the program is to relate spinal fracture threshold

between the cadaver and Army aviator populations.

The test program was designed to initiate testing at the operational limit-load setting of 14.5 G and gradually reduce the limit

load until a threshold was achieved. Originally, it was planned

to conduct a purely vertical dynamic test (+13 degrees pitch of

the seat vertical axis) and, if no spinal fracture occurred, then a combined-mode test (+30 degrees pitch) would also be conducted.

However, it became apparent after the first test series at 14.5 G

that the spinal condition could not be reliably assessed with

posttest x-rays. The remainder of the test program followed the

protocol of one test per cadaver in the combined mode, which was

"believed to be more severe, followed by a spinal autopsy to determine condition.

As mentioned above, the test series began with a 14.5-G energy ab- sorber limit-load setting. Subsequent series were conducted at

11.5 G and 8.5 G; the latter level appears to be near the spinal injury threshold for the cadaver population. The typical fracture

encountered in this test program was an anterior wedge fracture

of the vertebral body due to a combination of compressive loading and bending.

Bone samples were taken from each cadaver to provide a baseline for

comparison of bone strength with the aviator population. It is

4• interesting to note that early accident experience in Black Hawk

helicopters with the same seat design and 14.5-G energy absorber

settings shows that this level provides a significant degree of

protection from spinal injury for Army aviators.

43

*~~~~~- .,. . . . = 7 .

.. 1. 

,. ,% . * . * A .. . .
TABLE 3. TEST MATRIX FOR CAMI RIGID SEAT SERIES

Rate of Dummy Seat

Primary Test Onset Av G Dummy* Percen- Pitch

-Variable no. (G/sec) (ft/sec) Tye tile (deg)

Basic Dummy _Type

Hybrid II A80-025 510 42.1 14.2 1 50 0

A80-026 510 42.1 14.4 1 50 0

A80-027 550 41.9 14.4 1 50 0

VIP-50 A80-022 460 42.0 14.0 2 50 0

A80-023 520 42.3 14.7 2 50 0

A80-024 510 42.4 14.1 2 50 0

CG-50 A80-032 520 42.0 14.6 3 50 0

A80-033 510 41.9 14.6 3 50 0

A80-034 510 41.8 14.5 3 50 0

Dummy

Percentile

95th A80-028 520 41.6 13.8 4a 95 0

A80-030 510 41.8 14.0 4a 95 0

A80-031 540 42.1 14.1 4b 95 0

A80-029 530 41.8 14.3 5 95 0

5th A80-035 510 42.2 14.2 6 5 0

Seat A80-038 570 42.6 15.6 1 50 +30

Orientation A80-037 580 42.3 14.7 4 95 +30

A80-036 490 42.7 14.9 6 5 +30

Dummy

Performance

Verification Tests

Following A80-083 510 41.4 13.8 1 50 0

CAMI A80-084 480 42.0 14.1 1 50 0

Tests A80-085 510 42.3 14.1 1 50 0

Following A81-004 500 42.5 14.4 1 50 0

NADC and A81-005 520 42.5 14.6 1 50 0

Simula A81-006 500 42.4 14.4 1 50 0

Tests

*See Table 4 for description of anthropomorphic dummy types.

44

4,147
TABLE 4. DUMMY IDENTIFICATION

Test

Matrix

Dummy Serial Providing

Code Dummy Description Number Organization

- 1 Alderson Hybrid II (50th 870 NADC

percentile - Part 572)

2 Alderson VIP-50A (50th 314 CAMI

percentile)

3 Alderson CG-50 (50th 501 NADC

percentile)

4a Alderson VIP-95 (95th 133* CAMI

4b percentile with elasto- 153 ATL

meric spine)

5 Alderson CG-95 (95th 500 NADC

percentile)

6 Alderson CG-5 (5th 719 NADC

percentile)

SRigid Block (50th N/A Simula

percentile effective

weight, epoxy with seat

used as mold and seat

cushions removed)

*Alderson VIP-95 S/N 133 had been reworked to add an Ensolite pad

under pelvis in buttocks flesh.

In addition to the human tolerance data that are being collected,

each cadaver test provides information about the characteristics

of human body response in energy-absorbing seats for comparison

with anthropomorphic dummy response. Although the human body is

difficult to instrument, significant comparisons between cadaver

and dummy response can be made by examining seat response such as

seat pan acceleration histories and seat stroke. The sensitivity

analysis discussed in Chapter 6 includes a comparison of seat response for the two occupant types.

Recommendations for an optimal energy absorber limit-load setting

will be presented in the final report at the completion of the

cadaver testing program under Contract DAAK51-79-C-0016 (Reference 16).

The cadaver testing series is summarized in Table 9.

45

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TABLE 6. DESCRIPTION OF CUSHION TYPES

Type

Number Description

1 Simula Black Hawk crewseat cushion designed to minimize rebound

2 Soft polyurethane cushion with same contour and

dimensions as Type 1

3 Rigid, volume-controlled density foam cushion with

same contour and dimensions as Type 1

4 Three layers of 3-in.-thick soft polyurethane foam

without contour

S3.5 MODIFIED ANTHROPOMORPHIC DUMMY PROGRAM

Analyses of the data from the sensitivity testing program demonstrated that the current criterion for seat evaluation based on

seat pan acceleration is not sufficiently sensitive or repeatable

to provide a measure of injury potential. A suggested alternative

was to directly measure spinal forces and moments, which are the

vertebral injury-causing mechanisms, throughout the crash sequence. The U.S. Army Aeromedical Research Laboratory initiated a program

to modify a set of dummies for dynamic testing to assess the feasibility of this method (Reference 17).

Two anthropomorphic dummies, a 50th-percentile Hybrid II (conforming to the Part 572 Specification) and a 95th-percentile Alderson VIP-95 with elastomeric spine, were modified. The modifications consisted of inserting a six-axis load cell at the base

of the lumbar spine projecting into the pelvic accelerometer cavity. A schematic drawing of the Hybrid II lower torso is shown

in Figure 14 with the installed load cell. The VIP-95 dummy also

had sufficient space at the base of the neck segment for installation of a load cell without alteration of existing anthropometry.

The goal of the modified dummy program was to determine the feasibility of measuring spinal loads and moments with simple modifications to an existing anthropomorphic dummy. The performance

S,- verification test program consisted of eight tests divided into

17. Laananen, D. H., and Coltman, J. W., Measurement of Spinal

Loads in Two Modified Anthropomorphic Dummies, Simula Inc.

TR-82405, Final Report, Contract DAMD17-81-C-1175, U.S. Army

Aeromedical Research Laboratory, Fort Rucker, Alabama, May

1982.

48

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52

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two series. The first series repeated four tests from the parametric test matrix conducted at CAMI. In the second series an

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research effort can be found in Reference 17.

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4.0 EQUIPMENT AND TEST PROCEDURES

In order to isolate the effects of the variables being tested at

various facilities and in different tests at those facilities,

standardized equipment and procedures were employed throughout

the program. The seats and dummies used in the program are discussed in this chapter, along with procedures for conducting the

tests and achieving standard input conditions.

4.1 EQUIPMENT - SEATS

Most of the testing was accomplished using an energy-absorbing

helicopter crewseat, but some tests were conducted with a rigid

seat, in order to isolate the effects of dummy type and size.

Both of these seats and associated hardware are discussed below.

4.1.1 Rigid Seat

In order to provide data for dummy comparison, testing was con-

"ducted at the FAA Civil Aeromedical Institute (CAMI) with each of

the dummies in a rigid seat. As shown in Figure 15, the seat pan

and back form a right angle with respect to each other. No cushions were used, and the 1-in.-thick plywood seat pan was supported

by a six-axis load cell. The four-point restraint system used

automotive-type nylon webbing. In order to impose a vertical

(+Gz) acceleration on the dummy, the vertical (z) axis of the

seat was aligned with the velocity vector; in other words, horizontal tor the sled impact.

4.1.2 Energy-Absorbing Helicopter Seat

The majority of testing was conducted with a production UH--60A

Black Hawk crewseat, which has energy-absorption capability in

the vertical direction and is described in Reference 18. As

shown in Figure 16, the seat frame includes two vertical guide

tubes, which serve as races for the linear bearing, bucketcarrier assemblies. Each bearing assembly contains four contoured rollers located at 90-degree increments to surround the

guide tube, thus permitting low-friction translation of the

bucket along the axis of the guide tube.

Two energy absorbers that restrain the seat bucket in its vertical position are attached at the upper crossmember of the frame

and at the yoke mounted on the vertical adjustment mechanism,

which is attached to the s. :t bucket back. Vertical inertial

crash loads force the seat bucket down the guide tubes against

the resistance of the energy absorbers, producing an energyabsorbing stroke in that direction. The energy-absorbing devices

used on this seat are tensile, inversion-tube type.

18. Desjardins, S. P., et al., Crashworthy Armored Crewseat for

the Uh-60A Black Hawk, paper presented at 35th Annual National Forum of the American Helicopter Society, Washington,

D.C., May 1979.

55

--- ---------

""id. ,*4
44*X4

.41

.4,.1

.44

.556

.4 1- 

p

............
'4 w

ol 0 F u 

o-w YZ

or LU _j ()

<' -jz mw

c .4

U) LU >

• 0 4..• 

C4

to

4J4

.4. 0

a I

_ . h. ¢ 0"i Y•<; i.w 4i)#'I

454

- .,.- l1Q

| 57

.4 ,nt lair.

:.-.T ",, "-.

,;.,. .- .,'' '; ,,' -,,- . :,. ". . .,% , : . ., .,, . . .' . . ., ,. -%"., . . ,.. . . • . . . .'

7 I •i • • "l " i •" I . I ' i i i i + " i " ii i i" " • • . i ", •i " . i" •i • i ,i . "i , I . . I . i i ' 4
The interface between the bucket and the yoke is provided by a

T-track mounted on the back of the seat bucket, a sliding fitting

"I with T-slot (through which the T-track passes) attached to the

yoke, and a spring-loaded adjustment locking pin. Withdrawal of

the pin releases the attachment between the seat bucket and frame

so that the seat weight is carried by counterbalance springs,

which were removed for all testing on this program. Five inches

of vertical position adjustment are provided by this mechanism.

Longitudinal adjustment is achieved by releasing spring-loaded

adjustment locking pins in each of the front track fittings, permitting 5 in. of fore and aft adjustment.

The bucket of the production seat, in its operational configuration, provides ballistic protection and structural support for

the bottom, back, and sides of the thighs of the occupant. The

operational bucket is constructed of a lightweight, 13-ply laminate of Kevlar 49 material faced with extremely hard boron carbide ceramic tile in those areas requiring ballistic protection.

7 A layer of nylon spall shield is used to cover the boron carbide

and all cut edges of the Kevlar. Holes for attachment of such

items as frame fittings are drilled through the Kevlar, and proA I visions for attachment are included in the tile. For frame attachment fittings, the boron carbide is cut away, allowing the

brackets to be seated directly on the Kevlar. The lap belt anchors, inertia reel, and adjustment handles are seated against

aluminum standoffs, bonded to the Kevlar through cutouts in the

boron carbide. These support the compressive load of the

attachment without increasing the possibility of cracking the

brittle ceramic. For the parametric test program steel plates

were used instead of the boron carbide ceramic tile on the bucket

to permit adjustment of the seat movable weight. Plates were

bolted to both the back and bottom of the seat bucket, located

so as to maintain the position of the seat center-of-mass at the

same point as in the operational configuration with armor.

4.1.2.1 Cushions. The standard bottom cushion used in the UH-60A

crewseat bucket is a unique design that maximizes comfort and

crash safety. The surface of the cushion is contoured to distri-

"bute load over the greatest buttocks area, in order to minimize

% . any local concentration of loading that would cause discomfort.

rThe cushion base is made from foamed polyethylene, and the contour is lined with a thin layer of temperfoam to distribute local

.* loads and maximize comfort. This material is a loading rate sensitive polyurethane foam to help develop a more rigid, but yet

comfortable, link between the occupant and the seat bucket. A

final layer of reticulated polyurethane foam is provided for both

load distribution and thermal comfort. The cushion is enclosed

in a protective covering of fire-retardant, open-weave nylon material. Also, provisions are made to allow fore and aft circulatior, of air that can pass through the open-celled, reticulated

foam for cooliig purposes. Air circulation is natural, not forced.

58

4 * . q AF
The back and lumbar support cushions are of typical design, using standard upholstery foam and open-weave nylon covers. The head- - -rest cushion is formed of temperfoam for cushioning of head impact.

' -, A new bottom cushion core was inserted into the nylon cover for

each test. Back cushions and headrest cushions were reused

throughout the test series. -44

In the test matrix, three nonstandard cushions, intended to have

"force-deflection characteristics substantially different from the

production cushion were used. Both a rigid bottom cushion, made

of foamed polyurethane, and a soft cushion, contoured from uphol-

"stery foam, were sized to produce approximately the same seating position for the dummy as with the production cushion. Also, three

slabs of furniture foam having an undeformed thickness of nine

inches were used as a cushion to determine the effect of relative

motion between the occupant and seat pan allowed by soft, thick

cushions.

4.1.2.2 Energy Absorbers. Each of the energy-absorbing devices

on the production seat exerts a constant load during stroking.

As illustrated in Figure 17, these devices make use of the inversion of a thin-walled ductile aluminum tube. As required by

Reference 4, the total energy absorber load is determined by multiplying a given load factor times the total moving weight

% of seat and occupant. For the UH-60A Black Hawk, the weight of

the stroking part of the seat is 60.6 lb. Adding the effective

weight of the 50th-percentile Army aviator, 139.0 lb, the total

movable weight is 199.6 lb, which when multiplied by a load factor of 14.5 gives a total dynamic stroking load of 2,894 lb.

* Dividing the dynamic stroking load by a rate sensitivity factor

of 1.2 gives a total design static load of 2,412 lb. The static

stroking load for each of the two energy absorbers is then

1,206 lb.

II Thin walled aluminum tube

Unstroked

Partially str-oked

Figure 17. Inverted-tube energy absorber.

59

. & * % 44,, t• ,4 % ' . 4.... .. O4.W.W 4'" ... 4.. t w .- . - W .
For a given aluminum alloy, the stroking load of the energy ab

sorber is determined by the diameter and wall thickness of the

inverting tube. For two of the tests conducted at CAMI, energy

absorbers with reduced limit loads were needed. Using load factors of 11.5 and 8.5 gave static stroking loads of 956 and 707 lb,

respectively. Devices with these limit loads were manufactured

and used at CAMI in tests A80-055 and A80-075.

For two of the tests at CAMI and one test at NADC, it was desired

to use energy absorbers with an increasing ramp force-deflection

characteristic. The desired characteristics shown in Figure 18

were approximated in the energy absorber design by grinding a

gradual taper on the outside diameter of the aluminum inversion

tube.

251 11

, Test A80-066

20 e Test A80-065

li••!•'•:. - • •ee • tanardTest N-194

S 15 _00 S• 15 - , Baseline.

BaselineStandard E/A used

for baseline tests

* M

10 0, . 2. 

i0

:.,

$ 0

S0 2.5 5 .5 i0 12.5 15 17

i4 stroke (in.)

-• Figure 18. Nominal load-stroke characteristics of ramped

energy absorbers used in parametric test

program.

60

4 , .. j. -. - , t , , --- , .. ,. . . "..
4.1.2.3 Restraint System. The restraint system used in the UH-60A Black Hawk crewseat is a five-point system, which includes

a lap belt tiedown strap. As illustrated in Figure 19, adjusters

are provided in both shoulder straps and both lap belt straps, and the buckle is a rotary-release type, backed by a comfort pad. Characteristics of the low-elongation webbing are presented in

Table 11.

Inertia reel

. - n Inertia reel strap

S1.75-in.-vide

.webbing

4.;:

2.00-in.-wide webbing---:

Shoulder IV V4 S~ harness

hn -- --- Adjuster

"2. 00-in.-wide I .. bbn•ix-fiwebbing

RULUid Lower shoulder , Tiodonstrap

-'Adj -i fitting

•" Lap belt-- Lap belt---

;N A• anchor

webbing Fixed fitting on

•tiedown strap

•t•'• • •Tiodown strap

L-. •I.75-in. -wide webbing

Tiodown anchor

J

Figure 19. Aircrew restraint system.

61

4.41
TABLE 11. RESTRAINT SYSTEM WEBBING

Average

Elongation

at

Design Ultimate Design

Thickness Width Load Load Load

"Component (in.) (in.) (Percent) (Ib) (ib)

Lap Belt 0.057 2.25 7.5 8880 4000

Tiedown 0.057 1.78 7.5 6980 3000

Shoulder 0.057 2.03 7.0 7800 4000

Straps

4.2 ANTHROPOMORPHIC DUMMIES

The Aircraft Crash Survival Design Guide (Reference 7) recommends

that for aircraft seat testing a dummy conforming to Part 572

(Reference 19) specification should be used. All of the modern

anthropomorphic dummies have been designed for automotive impact testing and not for accurate response to vertical impact. The

spinal column, which is a critical region of human tolerance to

-'1.4 aircraft crash loading, has been designed in the Part 572 dummy to simulate response to -G loading rather than the more critical

+G loading seen in airc raft accidents. However, the reinforced

rubber cylinder, used as the lumbar spine in this dummy, permits more consistent positioning than the steel ball-and-socket configuration used in some other dummies. Instability in the latter

type could affect the response of the upper torso with concomitant

penalties on test repeatability. Another advantage of the Part

572 dummy for aircraft seat testing is a humanlike pelvic structure, which should result in load distribution on the cushion :N• close to that for a human. Finally, if the results of tests conducted at different facilities are to be compared, standardization of dummies and test facilities is mandatory, and the Part

572 dummy is the only truly standard dummy.

"For these reasons, an Alderson Hybrid II dummy built to the Part

572 specification was selected for use in the baseline test configuration at all facilities, and it was the only dummy used in the

NADC and Simula tests. At CAMI, tests were conducted using all

six of the dummy designs listed in Table 4. The 95th-percentile dummy most closely resembling the Part 572 design is the Alderson

VIP-95 with an elastomeric lumbar spine, shown as dummy code 4 in

19. Code of Federal Regulations, *Anthropomorphic Test Dummy," Title 49, Chapter 5, Part 572, Federal Register, Vol. 38,

No. 62, April 2, 1973, pp. 8455-8458.

62

,,- "
Table 4. It was one of these 95th-percentile dummies that was

selected along with a Hybrid II for installation of a six-axis

load cell in the lumbar region for the related modified dummy

test program described in Chapter 3.

In order to eliminate the.effects of dummy dynamic response, a

rigid block weighing approximately 140 lb was fabricated of epoxy

with steel weights located to position the center of mass at 10 in.

"above the seat pan and 10 in. forward of the seat back. This corresponds approximately to that of the seated 50th-percentile dummy.

This block, secured by a restraint system, is shown in Figure 20.

4.3 ANALYSIS OF ACCELERATION PULSE

The rate of onset (RO) of the input acceleration pulse at each

facility was determined using a procedure illustrated in Figure 21

(Reference 20) and outlined as follows:

* Locate the calibration baseline.

. Determine the maximum acceleration magnitude (Gp).

iip * Construct a reference line parallel to the calibration

baseline at a magnitude equal to 10 percent of the peak acceleration (G_). The first and last intersections of

this line with ihe acceleration-time plot defines points

1 and 2.

* Construct a second reference line parallel to the calibration baseline at a magnitude equal to 90 percent of

the peak acceleration. The first and last intersections of this line with the acceleration-time plot define points 3 and 4.

"" Construct a second reference line parallel to the calibration baseline at a magnitude equal to 90 percent of

the peak acceleration. The first and last intersections

of this line with the acceleration-time plot define

points 3 and 4.

"* Some logic and practical judgment may be required for

selection of the first and last intersections depending

on the noise, structural or electronic, apparent in the

data. Significant tendencies are important, not noise.

"* Construct the onset line defined by a straight line

through points 1 and 3.

20:. Military Specification, MIL-S-9479, Seat System Upwa'd

Ejection, Aircraft, General Specification for, Department

Dfense, eof Washington, D.C.

63

• ' • ~ =• ..q. 4 "*•'.. ' ::-'I. * ,4•'4*4 .... ,0 ' .. •,•q ........ P.......• ..... " ...... * * '. *q...:.. : .. "

44..4,.4. 4-. . . . . . . . .. . . . .. ,4 ,, ** -*4, ,.%,.4 444 4 4',-,.*, - $- . - ; ..-' - * . .. - . ". "., . ", -" "- . .. k • .. '_", , . , . .*.,* -" 4", ... •
A4)

AV0

a)

it)

Om

w3

64r

Ap.

Z -A. 'k.41 --. U
Gp - , 6

0.9 Gp3

Calibration

baseline 0.1 G2 - 2

to t r

Figure 21. Graphic approximation example for

rate of onset (from Reference 20).

* If desired, construct the offset line defined by a

straight line through points 2 and 4.

* If desired, construct a line parallel to the calibration

baseline, through the peak acceleration. The time interval defined by the intersections of this line with

-'A the constructed onset and offset lines (points 5 and 6)

is the plateau duration.

* Locate the intersection of the constructed onset line

with the calibration baseline (point 7). The time interval defined by points 7 and 5 is the rise time (tr)e

e The rate of onset is the slope of the onset line, G /t

4.4 STANDARDIZED TEST PROCEDURES

Standardized procedures for orienting the seat, maintaining the seat, and positioning the dummy were followed at each facility.

4.4.1 Seat Orientation

"Seat orientation varied depending on the use of a vertical drop

tower or horizontal sled, as describc4 below. The vertical configuration tests conducted by NADC and Simula oriented the seat

so that the seat vertical (z) axis was pitched forward 13 degrees,

as shown in Figure 22. Because for the Black Hawk crewseat the

seat back tangent line is oriented 13 degrees from the vertical

65

4.,* ,* † † N---.. -- .

• .4 * **•**~~* . ~444, *b* *~ 4 4* ** . ~4
.13 Fiur 220etcl et rp oe ofiuain

II

.4l

4'

S~130 ,

axis of the seat, the seat back was parallel to the impact veloc- • vector for the vertical Sity tests. At NADC, one test was con- •l

ducted in a second *vertical* orientation# where the vertical seat..

-axis was parallel to the velocity vector. Also at NADC* one test..

S was conducted in a combined orientation with the seat pitched for- " • '! ward 30 degrees from the velocity vector and rolled 10 degrees. ,1"..

66"''

/ilr'l['-l'l~~... -- i I i .. ... l l i ...... 
4.4.1.2 Horizontal Sled Tests. For the tests conducted in the

vertical configuration at CAMI and Wayne State University, the

*. seat was pitched forward 17 degrees with respect to the velocity

vector. The configuration for CAMI is illustrated in Figure 23.

"The first 13 degrees of pitch was provided to align the back tangent line with the horizontal surface of the sled, or parallel to

the velocity vector. This was done to eliminate initial extension of the elastomeric spine that would be caused by an angle

oriented downward to the seat back. The additional 4 degrees of

was added to counteract Spitch the I G of gravity that reduces the

overturning moment on the dummy. Since during most of the tests

the energy-absorbing mechanism on the seat is set to stroke at

14.5 G, the angle of the resultant acceleration was dezermined by

establishing the tangent of the angle 1/14.5 = 0.069, which is

the arc tangent of 4 degrees. Thus, under stroking loads, the

overturning moment on the dummy should be somewhat corrected for

the 1 G of gravity, and the response should be similar to that

for a seat in the upright-oriented position in a vertical drop.

For the combined orientation tests on the horizontal sled facilities, the seat was pitched forward 34 degrees from the velocity

vector and rolled 10 degrees in the seat coordinate system. The

first 30 degrees of pitch was provided because the present qualification test requirements include a 30-degree forward pitch.

The additional 4 degrees of pitch was added to approximately counteract the 1 G of gravity that would reduce the overturning moment

on the dummy, as described above.

ftt~a letad eelA I

"a°d

0 ,

t 1-0,4d '~

I.

Figure 23. Vertical test, horizontal Sled configuration (CAMI).

67

" aai*i' -. , ** . '.,a. 7.' * ,'6 U "9 : , . .,• a .

•,.•' •";••*"4 P•a : . ..a•P •• ' ""P .. "aP. .. "I . ." - • . .• ... • .... 9 . ."m "i .. .•i -

4* * *
4.4.2 Seat Preparation

The UH-60A Black Hawk crewseat was locked in the top of the verti-

"cal adjustment range for all tests. For the tests using 50thpercentile dummies, the seat was locked in the middle of the fore

and aft adjustment range and was moved forward or aftward for

tests with a 5th- or 95th-percentile dummy, respectively. The

seat was inspected for damage prior to each test, and the four

linear bearing assemblies were replaced after every five tests.

A new set of energy absorbers and a new seat bottom cushion were

installed for each test.

*• 4.4.3 Dummy Positioning

Photographic targets were placed on the outer (lateral) surface

of each shoulder and knee and on each side of the head. These

targets were used to locate the dummy in the seat and to facilitate analysis of motion picture data. A target was also placed

on the side of the seat at the seat reference point. Targets

were placed on the sled and floor fixture to assist in film analysis. Prior to each test, the dummy was positioned in the seat

described as follows:

The dummy joints were tightened until the frictional

resistance in the joint would support the outstretched

.limb.

* Actual installation of the dummy in the seat varied

somewhat among facilties; however, the following general technique was used:

-The dummy's upper torso was folded forward around the hip joints until the torso touched or was as

close as possible to the thS.ghs.

- The dummy was placed in the seat and its buttocks

were pushed back as far as possible.

- The dummy's legs were placed against the footrest .4,. simulating the rudder pedals and taped to the

"footrest .

- The dummy*s body was raised by rotating it again

"around the hip joints and pressing the dummy'l

back against the back of the seat. The rotation

of the dummy's torso tended to push the dummy further back in the seat.

* The restraint system was fastened and adjusted as fol-

•, lows..

- All fittings were inserted into the buckle with

the inertia reel locked.

M

68

**W W. -4W "* , * -.* .. ,• - 4, - -- • -T. . *•.- • • ,' -* " ." . " • . * " ,

%*, . " '. %' •' 4% '% %**4, ''• '. 4. -, ' " "," . " ,. " , .• , • o' .. "-. •" " & " - "
"- The lap belt and shoulder straps were adjusted by

pulling the free end of the straps toward the

buckle. A 50-lb force was applied to the lap belt

straps and a 30-lb force to the shoulder straps.

A 7/32 Allen wrench was placed in the end of the

inertia reel spool and 80 in.-lb of the torque was

applied, tightening the inertia reel strap.

e The location of the targets from a selected reference

point was measured on the seat for the first test and

for each subsequent test using the same dummy. The

position of the dumimy was adjusted until tho target

locations were matched.

4.5 INSTRUMENTATION

Instrumentation used in the tests at CAMI is as follows:

Number

of Channels

AccelerometerE

e Seat bucket, triaxial (aircraft axes).* 3 (2)

e Seat bucket, vertical (parallel to the seat

back tangent line) 1 channel (redundant). 1

* Chest, triaxial.* 3 (2)

* Pelvis, triaxial.* 3 (2)

e Head, triaxial.* 3 (2)

o Platform, vertical (parallel to velocity 1

* vector), 1 channel.

e Platform, vertical (parallel to velocity 1

4 vector), 1 channel (redundant).

e• Fixture, triaxial (aircraft axes). 3 (2)

Loads

* Strain-gaged energy absorber clevises: 2

2 channels.

o Restraint systen. webbing tensiometers, 5

5 channels (as shown in Figure 24).

o Footrest loads (3 channels).* 3 (2)

Displacement

e Vertical, parallel to guide tubes, attached to 1

bucket on centerline, either below seat refer-

"ence point for operation in retraction mode or

near headrest for operation in extension mode.

o Vertical, parallel to guide tubes (redundant). 1

*Lateral measurements were not required for purely vertical (00

pitch, yaw, and roll relative to the velocity vector) tests.

69

"r."
"* Longitudinal, perpendicular to guide tubes, 1

aattached at center of upper crossmember.

" Lateral, perpendicular to guide tubes, 1 (0) attached at top of guide tube.*

Impact Switch

1 channel for each data tape. 1

33 (26)

Inertia reel

= Inertia eel strap

- i.75-in.-wide

webbing

S~Adj us ter

2.00-in.-wide webbing- Adjut-

"2.00-in.-wide

Shoulder webbing

harness strap Lower shoulder strap

a Tensiometer in

Tensiometer adapter •:i i •aater [) Rotary buckle

in adapter

:' 2-1/4-in. Plug-in fitting tensiometer

"Ajut 

Ad j ust er---tens 21--n ione ter/

Lap belt J - Lap belt---

anchor 2.25-in.-wide

"webbing Fixed fitting on

tiedown strap

Tensiometer F Tiedown strap

1-. 75-in.-wide webbing

"Tiodown anchor

"Figure 24. Tensiometer placement on occupant restraint system.

70

) . . . -4 44L\ 4 "" 4 . 4 '4 44 4 4 * 4 * ' 4
-'; For dummy tests with rigid seats the following instrumentation

was used:

Number of Channels

I Accelerometers

9 Chest, biaxial. 2 9 Pelvis, biaxial. 2

: Head (translation and rotation). 3

Platform, vertical (parallel to velocity 1

1' vector).

"" * Platform, vertical (parallel to velocity 1

vector), redundant.

Loads

* Restraint system webbing tensiometers. 5

* Seat pan loads, triaxial. 3

"- Footrest loads, triaxial. 3

Impact Switch

_ 1 channel for each data tape. 1

21

Photographic equipment was as follows:

Cameras

9 High-speed 16 mm color, 500 frames/sec (minimum), 4 each (minimum).

* Real-time 16 mm color, 1 each (selected tests

as required).

e Pre- and posttest 35 mm color, I each.

A triaxial accelerometer mount was located under the seat reference point (SRP) on the bottom of the seat pan. It was mounted

so that the vertical transducer was parallel to the seat pan (which was pitched up 3 degrees from the horizontal axis) and

so that the vertical transducer was located approximately 1.15 in.

forward of the aft edge of the seat pan on its centerline. The

redundant mount (vertical acceleration) was located on the seat 0

back, oriented to measure acceleration in the direction of the

back tangent line, which is pitched back 13 degrees from the $ vertical axis.

A,.

71

S... • . • ... .... .... .. .-. . . . ..
5.0 ANALYTICAL METHODS

Integral to a research effort to improve energy-absorbing seat

design is the development of methods for evaluating those designs

with comprehensive test criteria. In order to develop a basis

for improving test criteria, a number of parameters were examined

to determine their degree of correlation with relative test severity. During the sensitivity analysis, each parameter was used to *. -• assess the relative hazard between dynamic tests in a series,

such as total velocity change. It is the objective of this research effort to determine which parameter has the greatest potential for reliably predicting relative test severity, and in

the future it will be necessary to determine values for this

parameter that provide an absolute measure of spinal injury potential.

The parameters examined in the sensitivity analysis are discussed

below.

5.1 CURRENT CRITERION

The current seat evaluation criterion is based on the magnitude and duration of the seat pan vertical acceleration. As discussed

in Section 2.2, the measure of absolute test severity is based on

the work of Eiband (Reference 8). The tolerable region of acceleration magnitude and duration, shown in Figure 4, dictates that

the seat pan acceleration not exceed 23 G for a duration greater than 0.006 sec. There are several significant deficiencies with

this evaluation method. As noted in Section 2.3 on the interaction of the body and seat pan, the maximum seat pan acceleration

occurs when the body is experiencing the least transmitted loads.

Also, the seat pan acceleration is totally unrelated to upper

• .. torso flexion, which appears to be a major contributor to the

typical anterior wedge compression failures of the spine. Finally, the current interpretation of MIL-S-58095(AV), as opposed

to the intent of Eiband's work, assesses the pass/fail performance of a seat on the acceleration spike with the greatest duration at the 23-G level; but it is possible that multiple acceleration spikes raise the average acceleration well above the tolerable level. For example, in the parametric test program two

baseline tests were conducted at CAMI with nearly identical input

conditions. The seat pan accelerations for the two tests are

shown in Figure 25. The performance of the seat in Test A80-053

would be judged as a borderline case while the same seat used in

test A80-054 would mosL likely be judged as not meeting the requirements of MIL-S-58095(AV). The seat tested in ABO-053 would

be evaluated as a safer design because a higher-frequency oscil1 lation occurred at the precise time to allow the seat pan acceleration to drop below 23 G. A seat with a characteristic high frequency seat pan oscillation, shown in Figure 26, can be rated

as acceptable even though it is obvious that the average acceleration exceeds 23 G for a long duration.

72

V #0 ta V V* W ... *° .,'.. • . . . ..@ D g . , - g , -t • "t "" "" •

SA,."i.?.';.< ... 'i.'•°.i° . .i". 4. .
10

0 0

0 0 -10 0.O•007 sec

-P S • 0.006 sec

D -20

a) o -25. 5G 23 G

30 - 0.005 sec

-40 I i I

0. .050 .100 .150 .200

"Time (sec)

CAMI test A80-053

"10

"0 A

"0 -i0

$ 0. 016 see

S-20 •i,,o• - --- 23 G

: -25.8 G

-30 0. 006 sec -29.4 G

-40 A

0. .050 .100 .150 .200 *4 9

.Time (sec)

CAMI test ASO-054

Figure 25. Vertical component of seat pan acceleration

measured in two similar tests.

S" 73

* *. '777 7.-7:.. 7.
60

-45-

0

30

S23 G

V4

,15 -

-. ,, 4.

'U)

.,-2 0.' -•'15

.- •....'Time (sec)

SFigure 26. Seat pan vertical acceleration measured }• in a qualification test of a production I',L•.energy-absorbing seat.

0ACC LERATION

*The maximum magnitudes of seat and body accelerations were tabu-

•,:;;. lated extensively in the sensitivity analysis. The parameter is •"".. tests. easily calculated For example2 and in provides the peak qualitative input acceleration comparison between seria s deeration scribed magnitudes in Section show 6.1.1, a qeat both direct correlation pan and pelvic z-axis accel- to the peak values

474

5.- .2 • C.O• 

-, •C....
of input acceleration. This example points out the positive and

negative aspects of using peak accelerations as indicators of relative test severity. The parameter does show a definite trend

in the input acceleration series; however, this may be an inconclusive trend because it fails to include the other very important "I : F quantity related to injury, i.e., time duration. It is possible

that. the measured peak acceleration on the seat pan or in the

body occurs during an extremely short duration spike, and that

"the constant energy absorber limit load produces an average acceleration that is invariant to the magnitude of input deceler-

"ation. Without including additional information with this param-

"eter, caution must be exercised in drawing conclusions. However,

in combination with other indicators of test severity, as used in

the sensitivity analysis, peak acceleration values can be very

effective in predicting severity, especially in comparing tests

"where the input pulse, seat design parameters, and occupant type

9-4 can be expected to produce similar response characteristics.

5.3 COMPRESSION OF BODY SEGMENTS

Axial loads in the spine are directly related to the amount of

compression that occurs during the impact sequence. An attempt

was made to calculate lumbar spine deflection by double integra- tion of the relative z-axis accelerations in the pelvis and chest , as described in the following equation:

r t

.9 6 spine = - )dt dt 0 0 %I

"I Unfortunately, after several trials with the data from the three

Simula energy-absorbing seat tests described in Table 8, it was

apparent that there were insurmountable problems with this method.

Any systematic errors in the data due to transducer calibration

and data processing tend to be propagated during the integrations,

"-ký and rotation of the body segments causes the accelerometer axes

.- to become misaligned, resulting in a rapidly diverging function.

"Also, the only way to verify this parameter (without instrumenting the dummy to actually measure spinal deflection) was by comI parison with film analysis, which has its own set of accuracy

problems. Therefore, this parameter was not used in the sensitivity analysis.

• :, 5.4 DYNAMIC RESPONSE INDEX

A weighted impulse function of the seat or body segment acceleration was originally suggested as a candidate for determining

test severity. This type of function has significant advantages

over peak accelerations because it combines the effects of magnitude and duration to a certain degree depending on the weight- ing function.

75

V- a 4,W "WA . 9 9
The U.S. Air Force has done considerable work in developing a

damped, single-degree-of-freedom model of the human upper torso

for study of impact applied parallel to the spinal column (Reference 21). The model, shown in Figure 27, represents the spinal

column (spring element) and body mass acting on it. The singledegree-of-freedom system is used to predict maximum spinal deflection and associated force within the spinal column for a given impact environment. Properties of model elements have been derived

from existing data. Spring stiffness and breaking strength have

been determined from cadaver vertebral segments, and the damping

ratio has been calculated from measurements of mechanical impedance during vibration tests.

2

r- - -- - , m = mass (lb-sec /in.)

6 = deflection (in.)

M c = damping ratio

m k = stiffness (lb/in.)

Z = acceleration input

(in./sec )

4 k

2 S

.*DRI - n max

Wn = natural frequency of

the analog = Vk/im

(rad/sec)

*Dynamic Response Index g 386 in./sec2

Figure 27. Spinal-injury model (from Reference 21.).

The equation of motion for this system is the second-order differential equation shown below.

d2 d 2

d 6 ~+ 2tw ý+ Wn 26

dt

21. Stech, E. L., and Payne, P. R., Dynamic Models of the Human

Body, Frost Engineering Development Corp. AMRL Technical Re-

-* port 66-157, Aerospace Medical Rescarch Laboratory, WrightPatterson Air Force Base, Ohio, November 1969, AD 701383.

76

Ap* -0 W . •1 

4•'- .• ", ' , " , . . . . , • . . • .' " " , , , . . . • . . . . . -- .
The solution of this equation, 6, is the spinal deflection of

the simple, single degree-of-freedom model. This solution is an

impulse-type function in that it takes into account the time period in which the acceleration is acting. The Dynamic Response Index (DRI) is proportional to the maximum value of spinal deflection and has the form:

DRI = max

The Air Force has been successful in quantifying the injury potential for ejection seat injuries with the DRI. The probability of

injury at various DRI levels is shown in Figure 28 (Reference 22).

For evaluation of ejection seat injuries, the seat acceleration

is used as input for the DRI equation. In general, ejection seats

have an extremely rigid bucket compared with production energyabsorbing seats; therefore, both the seat pan and pelvic acceleration were tried as input to the DRI model for the sensitivity

"analysis. However, only the seat pan DRI appeared to be a consistent indicator of test severity, and because of the problems

associated with instrumenting cadavers it was the only form of

the DRI that could be calculated for the cadaver test series.

5.5 INSTRUMMNTED ANTHROPOMORPHIC DUMMY

"Initial work in the sensitivity analysis indicated that all of

the candidate parameters for correlation to test severity and injury potential had significant deficiencies. It was concluded

that the only expedient way to determine injury potential would

be to actual'ly measure the injury-causing loads and moments during the impact sequence. A test program was undertaken, as described in Section 3.5, to determine the feasibility of instrumenting a test dummy to measure these parameters. During the

research effort, six-axis load cells were installed at the base

of the lumbar spine of a Hybrid II dummy and the spine and neck

of a VIP-95 dummy. The only possible installation scheme, with- . out extensive redesign, placed the load cells in the accelerometer

cavity below the spine.

.A limited number of test results were collected in the program

and proved the validity of this evaluation method. However, a

number of technical difficulties were encountered which will

have to be overcome before a dedicated test evaluation dummy can

be built. Lumbar forces and moments measured in the two dummies

differed from those predicted by the mathematical model described

2. 2 Brlnkley, J. W., and Shaffer, J. T., Dynamic Simulation

"Techniques for the Design of Escape Systems: Current ApplM-

"cations and Future Atr Force Requirements, AMRL Technical

"Report 71-29-2,,Aerospace Nedical Research Laboratory;

Wright-Patterson Air Force Base, Ohio, December 1971,

AD 740439.

"77

• , 9 ... .. . ... 9 .... - .". *.... . . * . . . ". , . . , . .. . . ".. . " [ '[["• ' ""
50 . Operational data

"40 - Cadaver data

30- /

h•./

0 20 /

10 E4-) 5/O.

* 40.d 5 / r4

S0.2-

- . .I I I . . . I 'I I

10 12 14 16 18 20 22 24

Dynamic Response Index

Aircraft type Nonfatal ejections

A* 64

B* 62

C 65

"D* 89

E 33

•.,4 F 48

*Denotes rocket catapult.

"Figure 28. Probability of spinal injury estimated from

laboratory data compared to operational

experience (from Reference 22).

"in Section 5.6, particularly during significant forward bending

(flexion) of the torso. A secondary load path, such as the *gut,*

is suspected to account for the differences.

"I f the concept of an instrumented anthropomorphic dummy is adopted

as the criterion for seat performance, and there are many positive aspects to warrant this choice, two tasks will have to be

completed before construction of the test device. A design evaluation of the deflection characteristics of the candidate dummy

must be undertaken to ensure that the response imitates the human

body, and that the load cell installation does nat affect body

78

"" %

4 4.
dynamics. Also, benchmark dynamic tests must be conducted, com-

"--.1 paring cadavers and living human subjects with the standardized

test dummy, to provide a quantitative assessment of the levels of

loads and moments that represent significant injury potentials.

5.6. SEAT/OCCUPANT SIMULATION

"Computer simulations of seat and dummy response were made using

the SOM-LA program during the sensitivity analysis. As described

in References 23 and 24, SOM-LA (Seat/Occupant Model-Light Aircraft) includes a lumped-parameter model of an aircraft occupant

and an optional finite element seat model. Interface between the

seat and occupant is provided by seat cushions and a restraint

system consisting of a lap belt and, if desired, a single-strap

or double-strap shoulder harness. A lap belt tiedown strap, or

crotch strap, can be added for simulation of a full, five-point

restraint system like that used on the UH-60A Black Hawk crewseat. . JThe response of the occupant and seat can be predicted for any

given set of aircraft impact conditions, including the initial

velocity and attitude and the input acceleration.

"Program SOM-LA includes a three-dimensional occupant model which

may, at the option of the user, be restricted to symmetric plane motion. The three-dimensional model is comprised of 12 rigid

links, connected by ball-and-socket and hinge joints and has a

total of 29 degrees of freedom. The symmetric model was used in

the simulations for the sensitivity analysis because it has deformable beam-type elements in the spine and neck, as shown in

Figure 29. As illustrated in Figure 30, this model has 11 degrees of freedom, which include both axial and flexural deformation of the two spinal elements.

The same procedures are used in SOM-LA for calculating forces applied to both two-dimensional and three-dimensional occupants.

That is, the three-dimensional geometry of the restraint system

is used in determining its position and the forces it exerts on

the elipsoidal contact surfaces fixed to the torso segments of

the occupant model. Thus, the two-dimensional model is actually

23. Laananen, D. I.1, •-lukbasi, A. 0., and Coltman, J. W., Conputer Simulation of an Aircraft Seat and Occupant in a Crash

Environment, Volume I - Technical Report, TR-82401, Simula

* Inc., DOT/FAA/CT-82/33-1., Federal Aviation Administration

Technical Center, Atlantic City Airport, New Jersey, Septembar 1982.

24. Laananen, D. H., Coltman, J. W., and Bolukbasi, A. O., Com- - puter Simulation of an Aircraft Seat and Occupant in a Crash

Environment, Volume I1 - Program SON-LA User Manual, TR-81415,

simula Tnc. OOT/IAA/CT-82/3-1, Federal Aviation Administration Technical Center, Atlantic City Airport, New Jersey,

September 1982.

"79

•%*4.,%
Neck element

V3

Spine

.4 6.

"%;•5, 7

* *4.4

*$ 10.

S• ~segment mass cente

0 Joint 8

L X

Figure 29. Eleven-segment (symmetric) occupant model.

a three-dimensional model restricted to plane motion, and no penalty is incurred by its use in simulating impacts for which occupant and seat potion is symmetric, as in most testing conducted

on this program.

For cases where the finite element seat analysis is not desired,

SOM-LA offers two options for modeling the seat. First, a rigid

seat consists of two plane surfaces that remain fixed in the aircraft coordinate system; these surfaces support the cushions. A

second option is the simulation of a guided energy-absorbing seat

like the UH-60A Black Hawk crewseat with a two-degree-of-freedom

model. The energy-absorbing stroke and rotational stiffness of

the seat are represented. Input parameters for this simplified

seat model include energy absorber force-stroke characteristics, A movable seat weight, rotational stiffness characteristics, seat

moment of inertia, and damping coefficients. For the simulations

.made on this program, the actual force-stroke characteristics for

the energy absorbers were used.

Other input data for SON-LA include the properties of the cushions

and the restraint system webbing. The force-deflection characteristics for the UH-60A Black Hawk seat cushion, combined with the

dummy buttocks, are shown in Figure 31. In Figure 32 are shown

the approximate force-strain curves for the low-elongation polyester webbing used in the five-point restraint system. Also, the

sled deceleration is approximated by up to 16 points in acceleration and time. As an example, the sled deceleration for the CAMI

vertical baseline test with the modified 50th-percentile dummy is

shown in Figure 33, and the approximate pulse used as SOM-LA

input, in Figure 34.

•4•, 80

* ~ ~ ~ ~ 4 - ~ '1. 'a ".W IM .4. . . . . . . IP.. ..... " .... a ..... *• a ' I" .P...' . ... W • - - - - ..

' . -. , *4 . k.','..'". '..'L.°,". ",a "-'. . . .." ' ;. " %.. ... : " , ""a,"," .",- .. " " "' ' ' " , " 74 4 a
4:.l

3

'.4ý

" N4 02Generalized coordinates

2

e 2 Z

4 1

3

"4 4 s

52

05 6 p3

S5 8 0 4

9 0

4.1. 10 08

e1 1 (X1 , Z1 ) 11 09

118 9

i.4

•'•.Figure 30. Generalized coordinates for" symmetric -:•._occupant model.

81 "
4000 -

t 3000

0.56 ., ¶.,F -.760 (e -)

'-4

2000

0

4..w. 1000

00 1.0 2.0 3.0 4.0

Deflection (in.)

Figure 31. SOM-LA input properties for

Black Hawk seat cushion.

* Output data from SOM-LA include the following:

e Seat acceleration

, a Occupant pelvis, chest, and head acceleration

* Lap belt# shoulder belt, and tiedown strap loads P

, Energy absorber force

"* Footrest force

e Spinal forces and moments

* Other injury criteria, such as severity indices and the

Head Injury Criterion (HIC)

* Seat stroke

'a. 82

"Vi' 4

•. ,;-; ),*.'•, .% - 9 - - % .. '•I* -.. , o , .9 . ''9 V
10,000 •iiI "

S8,000 -2-1/4-in.-wide (lap belt)

2-in.-wide

(shoulder

6,000 harness)

0

rZ4 4,000

2,000 1-3/4-in.-wide

(tiedown strap)

0 __j

0 0.02 0.04 0.06 0.08 0.10 0.12

Strain (in./in.)

Figure 32. SOM-LA input properties for polyester

restraint system webbing.

As an example, output data for SOM-LA simulation of CAMI test

A81-124 are compared with test data in Figures 35-44. Computer

predictions match the test datd rather closely, with the exception of lumbar force and moment. As shown in Figure 43, the

axial force predicted by SOM-LA is much higher than that measured in the test. The difference is attributed to a second

load path in the dummy which is not present in the mathematical

model. When the dummy torso bends significantly in flexion the

"gut" segment, which is inserted between the pelvis and thorax,

can support a significant compressive load, greatly reducing the

axial force carried by the spin(. Also, the flexural stiffness

of the torso is altered, completely changing the moment carried

by the spine, as evident in Figure 44.

83

i: .. •U ,V • " 9 '.. '9 ' .. " '~ 49" 9U ... 9U.. g .. • ' •. ...
50 I " I I

40

o 30 "-I

$43

H 2020i

d

4)

0

-10

rT 10

I0

0 0.05 0.10 0.15 0.20 0.25

Time (sec)

Figure 33. Actual sled deceleration pulse for CAMI test A81-124.

50 "

Point No. Time Accel.

10 1 0. 0.

"40 2 0.0055 2.88

9 13 3 0.0085 5.75

" ,-. 134 0.0130 11.0

0 8 5 0.0175 14.8

S30 6 0.0250 25.1

o 7 0.0280 28.5

"6 8 0.0315 31.6

$4 20 -9 0.0382 35.0

4) 0 20 10 0.0440 40.1

U0 5 11 0.0480 41.5

0 12 0.0515 39.9 io 4 13 0.0535 35.0

r3 14 0.0590 -1.32

0 15 0.0628 -5.14

0 1 16 16 0.0680 0.

01 14

15

- i0 ' __ I II I , .

0 0.05 0.10 0.15 0.20 0.25

.1fTie (Sec)

Figure 34. Approximate pulse input for SOM-LA

simulation of test A81-124.

"84

. .-. . . .. .. .. . . . . , .. .. . . .

-t-7r-,,'V -,7 7-.*",*-*"---"--. 

a.'-
100

Simulation

e---Test data

80

S3 80"

z

0 40-

5. 20

•:, X

" C: • -20 N

0

> -o

_J .r,

C.

-Joe

-60

-80-

0.00 0.05 0.10 0.15 0.20 0.25

TIME (SEC)

Figure 35. CAMI test A81-124, dummy pelvis z-acceleration.

4 85

i- • , '- ' .. • .. . ... .
100

Simulation

---- Test data ,-o .•80

60 C!D

*.** 40

L.3

H

*-20o X

Li]

0

•:n..:lFigure 36. CAMI test A81-124, dummy chest z-acceleration.

86

0"" . •0•"" .. ' ....... A 3V w . "VW .- M
tO00

"Simulation

Test data 80

80-

CD

z

0

c 20

W

- 0 -

-

A

o/

*.12

oE

S-20-

I

* N

o -40

-60-

-80

-800

0.00 0.05 0.10 0.15 0.20 0.25

TIME (SEC)

Figure 37. CAMI test A81-124, dummy head z-acceleration.

87

i . .•. • t ,w ,•* . .. 9 . . -" P . ... . " • ' . . .. . . ' .* .* 4..9"4.. " .. . U . . ." • . .. - .. . . ., -. . . -
50

Simulation

' ------- Test data

40-

" 30

z

S-10

Li

Z/ 0-4

Cc 0.0 .5O.t .5 .00

10 -1

z

~-210

LJ-30-

-50

4;0.00 0.05 0.10 0.15 0.20 0.25 TIME (SEC)

Figure 38. CAMI test A81-124, seat z-acceleration.

:'1 888 8

. .4• , I .. . • . . .• I " " . . . . . • ' ' . I I . . • . . . . . .•, , " I . . - "•->-• ,..:.,,-,. . • .•. :..I•. • , . . . •. . .. . ... ••-- ........ . ,,*' 1_ _• . : : -: : .. '_: : ':_ -_ : ; : :.: _: : . . _ : : : :< • .; : : : ! ; .: .: , .: .:,.. ', ,: . ,: -, : : : ' :• " .: - - , . .: " .
Simulation

Test data

18

12-

8

S4

.4-0

•: --

*341 Q ~E--e

-12-

-18-

-20 . ' 0.00 0.05 0.10 0.15 0.20 0.25

TIME (SEC)

Figure 39. CAMI test A81-124, seat stroke.

AN

89

-4.p...4.
2000-

Simulation

Test data

1600

- 1200\

, 800

0 -

0 400

rL) U!

M 0--O

Q Cr) .00

a-: -400

*0 * CD

SELI

-1200-

-2000- w-I 1 1 I I I

.4 0.00 0.05 0.10 0.15 0.20 0.25

STIME (SEC)

Figure 40. CAMI test A81-124, energy absorber force.

* 9

AV 0.

S Sil . °

IE.'-• ..... ""IP • ... .. g . ...... P .. .. ,,,p," • • .. ...Lq '''t . .. g . "- -- • ' . ... ..

r .4- '' • '''''' . ' ' ' , , .' ' " " ' ' ,i: .

" a•, *'', ''',,.'''""' ' . ,
Simulation

---- Test data

3200

2400

S 1600

6-4

r,0

' 800

I -o-

;-.4 -1600

-2400

-3200

-4,0001 - i .....

0.00 0.05 0.10 0.15 0.20 0.25

TIME (SEC)

Figure 41. CAMI test A81-124, footrest x-force.

91

:. . . . . . . .. " " -1 . ." •..
4000

S- Simulation

----- Test data 3200-

2400

1600 ELI

* 0

U) 1 0 :

Xi 0 -- ' /

S-800

E~J

o-1600-

-2400-

-3200-

-4 0 - ... - " " ' * .•--i*, ,-. . , . .... .

0.00 0.05 0.10 0.t5 0.20 0.25

TIME (SEC]

Figure 42. CAMI test A81-124, footrest z-force.

92

".....
S2000-

Simulation

Test data

'_. 1200-

-3w

_ 400

0- t

L " I

J'4 •O\

0.00 0.l 0.10 0.15 0.2:00.5[

TIME (SEC) '•

Figure 43. CAMI test A81-124* lumbar spine axial load.

93

-. ...

.. . ... . . ..•. 0-. . . . . •. - .. .. _. • •1.60 • • - ' . - : . . ' .. - :•.

•,"/ -".".7. - 2000%,,'.:''.•,- '-•.".' ... •.- .''. .",, r•".' '. ". . ."•'.",'. .
Simulation

Test data

3200

2400

-3

-f--

Z a

-2400

-30

.6-

. 16000..

b0.00 0.05 0.10 0.15 0.20 0.25

TIME (SEC)

Figure 44. CAMI test A81-124, lumbar spine moment.

94

•:.4 .,•,,,, .••.,-•,,-". -•. . m.,, , •p • ••F •• 'P I p. " •

• • --•T' • :'-.V-'-••---- •• •.- , -•--. *- • •V• • :• ':• -- ; 'T"' T• • ....................... ': ... * ..........- • ,.............. ,...

'*-i '*,_•,_'I___•*.•_..'.I•'.•'L.• .':• "•• •.• . ._L _" . .. " _ _° '
5.7 LINEAR REGRESSION ANALYSIS

The least-squares linear regression analysis whose results are

presented in Chapter 6 is employed to minimize the sum of the

squares of the deviations of the actual data points from the

straight line of best fit. In order to approximate a set of

n data points (x, y) by a straight line given by

y =mx + b

the following equations are used;

n n n

E (xiyi) - ( xi yi)/n iPl i=l 1ii 1

Slope, m n

n r (Xi2 2n (£x.) F 2/n

i~li=:l

:'" n 11

y-intercept, b ( yi m z x)/n Si=1 il 1

The correlation coefficient, R, which provides a measure of tlie

quality of the straight-line fit, is given by

R om a/a

where a and a are the standard deviations of the x a,.1 y data

points.X

'95

'4 :' Ir•'••I ' • ... •P "•' .... 1 .. B • ' • W '" .. W • ,P. . "
6.0 RESULTS OF SENSITIVITY ANALYSIS

Th..; chapter presents a discussion of the sensitivity of seat and

occupant response to the 13 experimentally examined variables.

The sensitivity analysis for the experimental test series consists of a comparison of measured time histories of seat or occupant response, a tabulation of maximum values, and, where possible,

graphs of severity indicators showing trends through linear regression of the data. During the course of this research effort, it

became apparent that several other variables not included in the -° original test matrix could have considerable bearing on an updated

energy-absorbing seat specification. The system performance sen-

*1 sitivity to these variables was either implied from test data or

compared with a test matrix of SOM-LA simulations, since the parametric test series could not be expanded within the budgeted fund6.1 SENSITTVITY ANALYSES FOR EXPERIMENTAL TEST SERIES

Sensitivity analyses for each of the 13 variables examined with

experimental testing are discussed in this section. The charts

included in Figures 45 zim 15 provide a reference source for applicable test nin•sm b-h the rigid seat and energy-absorbing

seat test seri&k.

1l 

.

AGO-02) ASO-031

AOO-024 A#0-4i4

A406-03I A40-111ti

0,ým ",fnt:w

After r

Figure 45. Rigid seat series parametric test reference.

.' 96

1` * 4P 40* 9 W a V .10 W W *W- 'w
A80-076

A80-069

Cushion spring rate Dummy percentile

ABO-067 SEAC-3 A80-073

A80-068 A80-074

Rate of onset Mov seat weight

;•A80-057 A80-060 •:•ARO-058 A80-061

"Baseline Tests

Zyo nu us CM of ACPa pulse___ ________input 

AO0-070 N-192 AS0-062 N-188

ABO-071 N-193 A80-053 N-189 A80-063 N-191

A80-077 A80-054 N-190 A80-064

Simula WSU

SEA-1 WSU-159

Seat- orientationWS - 7 j c'EAl mt oa

: A80--.M N--19, A80-06S N-194

" A80-08.1 N-196 . A~n-066

', .Caddyero veraus

•Test facility a•nthropomorphic dummies

Impact cond1itions

$• A*' Ixl ba~e~lin on tin teatt s | AF020o ArO2, AF025 Arne0, AF029, AP031 AFo33, AP035, sFO3l

"., APO-055 AgO-VO• A ,O-07S AGO-014

Figure 46. Energy-absorbing seat series

parametric test reference.

6.1.1 Peak Input AccelSrationSeries

The peak input acceleration series was designed to investigate

seat and occupant sensitivity to the input pulse shape. The input conditions in an actual crash are a function of the landing

year/fuselage deformation -haracteristics, impact velocity, and

accident site terrain properties. For this series, the velocity

change was held constant at 42 ft/sec; the resulting input pulses

are shown in Figure 47. Measured peak values for the seat and

occupant are summarized in Table 12. There appears to be a definite trend in the vertical response of the seat as determined

by the reqtuired stroke distance and the maximum seat pan acceleration plotted in Figures 48 and 49. Recorded seat pan z-axis acceleration time histories are shown in Figure 50.

97

4a, IV -4• '.• "• •T - ' " : . '•-'""•• 7 '.. .. . .. ... . ...... . .. . .• . ... -' .. . . . . . .•.• • - • • • • ,.. .

,,.,4.• S- - . . . i - - ' " " .. " " ' • . ' ' . , . , . .'' " ' , ,
60 I

A80-064

50 A80-062

-A80-053

"""40 A80-063

o 30

$4-

C.)

.:, ,- 20

• •,0 - - C.) -10

0 .050 .100 .150 .200 .250

Time (sec)

Figure 47. Input acceleration pulses for

peak acceleration series.

Vertical accelerations recorded in the anthropomorphic dummy also

show an increasing trend with the peak input acceleration. This

trend can be seen in the pelvic z-axis time histories shown in

Figure 51 and the sensitivity plot of pelvic acceleration versus

inpu~t (Figure 52).

Several test severity indicators, shown in Figures 53 through 56,

were tabulated for this series, and all show the same increasing

trend with increasing peak input acceleration. However, there is

a significant amount of scatter in the measured seat pan acceleration duration data and the predicted spinal moment. The DRI and

predicted lumbar spine axial load both show a strong relationship

to peak input acceleration, as demonstrated by correlation coefficients of 0.92 and 0.97, respectively.

6.1.2 Velocity Change Series

Tmpact energy is proportional to the square of the velocity.

Therefore, it would be expected that seat and occupant response

4.' 98

% . ,' , ~ ~V ' V

,.! l ,'-, •'•, !! e, .... ! ..... ! ... .'!..-.- . . . -.. v"- --.... ' • ,-t - - ... w
TABLE 12. COMPARISON OF TEST RESULTS FOR PEAK INPUT ACCELERATION SERIES

Duration of

peak •'i Pek Total Toal •akAcceeraion(G) Seat Pan z

Input Initial Velocity Seat Acceleration

Accel Test Velocity Change Stj-oke Seat Seat Pelvis Pelvis Chest Chest Head Head at 23 G

(G) No. (ft/sec) (ft/)ec) (in.) Pan x Pan z x z x z x z (sec) DRI

21.9 ABO-064 41.3 42.1 7.0 15.4 22.4 10.7 24.4 12.1 24.1 22.4 28.5 0. 17.0

32.2 A80-062 41.8 43.3 9.5 26.0 24.6 9.7 36.8 31.1 26.5 37.8 26.1 0.005 18.9

35.2 N-191 36.8 37.9 9.4 14.1 33.3 14.0 30.8 9.40 31.5 19.1 32.5 0.010 20.9

40.7 A80-053 41.8 44.8 10.6 25.3 26.4 17.1 30.5 18.1 30.4 31.5 36.5 0.007 21.0

40.2 ABO-054 42.3 44.2 10.5 22.1 29.4 10.1 30.8 24.1 25.8 38.5 28.2 0.016 19.7

4.1.9 N-189 36.8 40.4 10.8 16.6 36.3 15.0 38.8 10.0 36.1 24.7 40.8 0.006 20.8

45.0 N-190 36.8 40.5 11.6 16.8 34.6 18.7 40.1 14.7 37.4 22.7 42.8 0.007 20.7

51.7 N-188 36.8 39.2 11.5 20.4 42.0 15.7 40.2 13.4 37.8 23.4 43.2 0.009 23.2

"50.9 ASO-063 42.1 45.6 11.7 22.3 31.8 36.8 45.9 20.4 26.8 32.5 32.5 0.013 21.4

. Peak peak Peak E/A Peak Lap Peak Shoulder Peak Footrest

Input Load (ib) Belt Load (lb) Belt Load (lb) Peak LWad (lb)

Accel Test -- G Strap

(G) No. Right ,eft Right Left Right Left Load (lb) x z

21.9 A80-064 1350. 1450. 180. 161. 241, 154. 389. 1340. 1640.

32.2 ARO-062 1400. 1440. 281. 154. 228. 388. 1640. 2220.

35.2 N-191 1430. 1440. 362. 348. 221. 255. 496. 1390. 2040.

40.7 ABO-053 1380. 1590. 373. 213. 200. 147. 593, 1760. 2200.

40.2 A80-054 1420. 1550. 300. 233. 87. 250. 560. 1900. 2120.

44.9 N-189 1449, 1457. 342. 335 188. 221. 476. 1670. 2310.

45.0 N-190 1370. 1415. 361. 508. 181. 355. 515. 1890. 2410,

11.? #4-188 1450. 1400. 274. 488. 94. 100. 963. 1993. 2620.

50.9 AO.-063 1400. 1440. 268. 275. 295. - 690. 2310. 3100.

would be a strong function of this parameter. All of the compiled maximum values listed in Table 13 show an increasing trend

with additional velocity change. It should be noted that it was

not possible to maintain the same pulse shape and peak input acceleration as the velocity change decreased; these three param- u', eters are not independent of each other and would be expected to

* •vary in a similar manner for an aircraft fuselage. Four input

acceleration pulses for the CANX test series are compared in Figure 57.

Several measured test parameters show a strong correlation to the

velocity change variable. These parameters, seat stroke, maximum

seat pan z-axis acceleration, and DRI, are plotted in Figures 58,

59, and 60, respectively; and, as would be expected, seat stroke

"is shown to be highly sensitive to changes in velocity. Program SOM-LA was used to simulate each of the nine tests in this series.

99
15 1 | 1 1 1 1 1

O Measured value

O SOM-LA prediction O

12

0 0

0

U) 9 

00

: 6

Least-Squares Approximation

S ý 0.159 Gi + 3.90

R = 0.957

I , I , , , I I 1 I

15 20 25 30 35 40 45 50 55

Peak input acceleration, G (G)

Figure 48. Peak input acceleration series, maximum seat stroke.

5o

a Meteured value

0 SOH-L.A prediction

4.. 40

9 0 0

UU

300

10 .- A

o. . . . .. ,i , , . l .- -....-. , 0 . .I

4,$ 100

0. C$ t 9C

0. A

,.,,

4 .. Is .. s is 40 .. 4% so 

.7 7 .0
40

35-

30

25 - 22.8 G

20

15 - A80-064

21.9 G

10

5

0A A

-5

40 ,

I " 35

30

25 24.4 G

20

is A80-062

32.3 G

10

5

20

Ox

-5

-10 1

30

25

20

is #AUo-05)

40.7 0

10

0

-10as

*10 1

0 0.0%0 0.400 0.11,0 0.200

'I.%

Figure 50. Seat pan z-axis acceleration measured

in peak input acceleration series.

101

__, m • ... • _1•,•T_• • _•,m qmm&•,,l•Im ,,'-ll••m qm " • s- .mw" "
50

40

30 24.4 G

20 A80-064

10 21.9 G

10

-20

50

40 36.8 G

30

20 A80-062

-10 32.3 G

-. 10C _ _

-10

4 40

0 30.50 5

20 A80-053

10 0 40.4 

30

20 AO-05

0.0. u

10 , O . 0 tO .•0

40

IQ

A4004

0.1to 0.300

"Figure 51. Pelvic z-axis acceleration measured in

peak input acceleration series.

Results of the simulations, which compare favorably with measured

test data, are also plotted in these figures. Included in this

section are predictions of peak spinal load and moments plotted

against velocity change, shown in Figures 61 and 62. Sensitivity of the predicted spinal loads and DIV 't- inr.easing velocity exhibit similar characteristics in which there is a defklote trend, although the velocity does not appear to be a strong influence.

This is a predictable trend because both of these parameters, indicating spinal load, should be greatly affected by the magnitude of transmitted crash loads, which are controlled by the energy absorber limit load and are ideally not influenced by the applied

load duration, which is a function of the impact energy. Data 'presented in Section 6.1.4 for the energy absorber limit load

102 449

',I

* ,., 4',•,4 o , . , .. ,,I .... .... • * .... I . . • . .. •l - .. -.y•• " "• .... WP• '"
so

0 Measured value "

0 SOM-LA prediction

40 0 0 0

030

00

S000

.1 30 0 00

uo 20

e Least-Squares Approximations

STest Data; Gp 0.598 G 11.4

10 R 0.804

Predictions: Gp 0.107 Gi + 19.8

R 0.522

0 .I I I

15 20 25 30 35 40 45 s0 55

Peak input acceleration,, (0)

Figure 52. Peak input acceleration series, maximum

pelvic z-axis acceleration.

O.0 - , I ,

0.020 t 0.000.-46 - 010057#

MIS

0

+4 0

0.00$- a

0 . . . a .. . . I. . . . . , . . . . . . . . . . . . I. 

P$ 20 25 30 is 40 4s SO SS

feak input accolertLoa. I to)

Figure 53. Peak input acceleration series, duration of

seat pan a-axis acceleration at 23 G.

N• 103

.44

4i " " l~'-'." ' .:' '...... • ""*• .... 'I .. •• .. .. , . " ' " :,t . ..• .. • . .. • -
25 I I I I I

0

20 0 1 0

0

15

5 Least-Squares Approximation

q Measured value DRI - 0.177 G + 13.3

0 SOM-LA prediction R = 0.918

15 20 25 30 35 40 45 So 55

*-: Peak input acueleratioin, G (G)

N Figure 54. Peak input acceleration series, Dynamic Response Index.

S,2000

1.500 -

1000

$00 , 12.7 1* 100.

0 -tdatc

Is 20 25 30 5 40 4S 55S0 

Pekinpuat ac.istelauioa, Gs(0I

Figure 55. Peak iiput acceleration series, predicted maximum

lumbar spina axial load (compression).

104

• ,. , . . . . . . . . . . . . . .. i .. . -
..1 2500

C; 0

2000 0 -

.4.0

0

1500

1 0 o o

0

S1000

Least-Squa es Apprcoximfation

"M 18.1 G 1100.

- SOM-LA prediction 9 0.419

0

"is 20 25 30 35 40 45 so 55

.P eak Input Aco@ ler@ tion, a1 (G )

:p.

Figure 56. Peak input acceleration series, predicted

maximum lumbar spine moment (flexion).

- : u ... SS0 ..

40 -0 --- --,, , 

ASO-Oi 8 -07

+:./ ---- =--- 07 'OO

",,ID'.r.. 20; 3

-10+

-'•10

i'

*.44

it'0 0.030 0.100 0.150 0.200

TiOm (see)

,*

Figure 57. Comparison of input acceleration pulses

for CAMI velocity change series.

105

5.

°9, 14 zot s- 4

•i•~~ ...•' ~ .! 7 • 7+':+•l' ... ,.... 'I .."• .... ... •.. ..••. .. , .... .. , ..
TABLE 13. COMPARISON OF RESULTS FOR VELOCITY CHANGE SERIES

Durition of

* Total Peak Seat Pan z

Velocity Initial Input Seat Acceleration

Change Test Velocity Accel Stroke Seat Seat Pelvis Pelvis Chest Chest Head Head at 23 G

(ft/sec) No. (ft/sec) (G) (in.) Pan x Pan z z : z x z (sec) DR!

44.8 ASO-053 41.8 40.7 10.6 25.3 26.4 17.1 30.5 18.1 30.4 31.5 36.5 0.007 21.0

44.2 A80-054 42.3 40.2 10.5 22.1 29.4 10.1 30.8 24.1 25.8 38.5 28.2 0.016. 19.7

40.4 N-189 36.8 44.9 10.8 16.6 36.3 15.0 38.8 10.0 36.1 24.7 40.8 0.006 20.8

40.5 N-190 36.8 45.0 11.6 16.8 34.6 18.7 40.1 14.7 37.4 22.7 42.8 0.007 2ij.7

34.9 A80-077 32.4 40.9 6.4 22.4 28.3 23.4 29.1 15.4 24.4 23.5 26.8 0.007 20.1

33.9 A80-070 32.3 31.2 5.6 20.4 20.9 13.7 26.8 10.7 24.8 19.8 29.1 0. 17.0

30.0 N-192 25.4 31.1 5.4 12.5 27.1 13.7 32.2 7.4 32.5 13.1 )4.5. 0.004 11,-

27.1 A80-071 25.6 23.6 2.2 14.9 18.4 9.0 25.1 12.4 Z5.8 8.4 2%.1. 0. 18.9

22.7 N-193 18.8 22.9 2.3 10.4 19.1 6.7 26.4 4.3 Z$5.1 aO. 25.8 0. 1616

PTotaleak Peak Lap Peak Sho'*l1r Peak rootrest

Velocity aej

Change Test la b)Bllod(b betW GStrop Wdjb

(ft/see)' No. Righ t _1!ft Right lAft Right Left 1o4a (16 X

44.8 A8O-4O53 1380. 1590. 37). 213. 200. 147. $93. 1760. 2200.

44.2 A80-054 1420. ISS0. 300. 2)3. $7. 250. $W0. 1900. zuq4.

40.4 N-lb9 1400. 1460. 342. 335. 18e. ..*, 476. 1610. 2310.

40.S 0.-•90 1370, 1420. 361. SW. 181S. . 1 . 2410.

1449 ABO-07 1370. 1440, 148. 107. 141. I ,. 10. 1950. 2640,

31.9 40I-070 1380 1470. 120. 74. 1.54, 174. 361. 1420. 19170,

30.0 U-192 1520. 1540. M, "1. ?48, MV1. 461. 120. 040.

27.1 40-071 1)70, 14M0. 194. 201, 141. 361. .5n. 113t). 1•90.

23.7 #-193 1180. 1490. 120. 310. 310. 390. 470. 9OW 11t0.

series do show that the DRI and spinal load are strongly influenced by the transmitted loads. Another significant relationship

in the data exists between spinal moment and seat stroke, both of

which are strong functions u" velocity change. "t is theorized

that the test procedure of taping the feet to the footrest caused

"this apparent relationship, because a signficant tonsile load in

the upper legs can be produced as the seat strokes. This load,

which increases with stroke due to the load path angle, causes the

pelvic section to rotate and. that must be resisted by a moment

in the lumbar spine which adds to the moment caused by flexion of

the upper torso.

6.1.3 Rate of Onset Series

Results from the rate of onset test series indicate that the effect this parameter has on seat and occupant performance must be

N .

'% '106

V,,",

vf~.q.t.w ~ --------------- -.
15

c measured valuQ

0 SOM-LA prediction

0 O00

00

ZOOO

S w

S+ / 0

I - 0.948

is 20 25 30 35 40 45 so

Velocity C"O'e AV ttt/4•:)

Figure 58. Input velocity change series, maximum seat stroke.

0 fata pa• ft40

U 00

00

i* O.+P4t

J . ,... I AX. t I .i . ...

15 20 5s 30 3S 40 45 50

Velocity chAO.. &V tft~/ae9g

Pigure 59. Input velocity change series, maximum

seat pan z-axis acceleration.

107

".0
25

00 - C 

20 0

o0

i "s

i00

a )easured v4lue DRr • 0.175 v 12.S

0 SA*-LA predicti" . .. 754

10 2L 30 is -40 45 5

Figure 60. Input velocity change series, Dynamic, Response Index.

- 00 - '

,0 2S 0-. -

'74-

,tt 0 :

• ,'-• " . ...I •.'•• .. I '. ...

4'•i I.08 ~

,: 0 ".,

"4 , .-.. _ , . ,,.•-.. _ • o •• " . . . .. • . . ... •.. .. . .. .. .. .. . . . . .. . . . . . . . . ... . . .

Pigure 61 ", ," Input ,.edcd veocity.,hnq-, . ,, ,
2500 1 1 1 I 1 1

0 SOM-LA prediction 0

2000

.0 -

z 1500 /

r 0 0

44) 00

*~1000

0Z 00

0 34 500 Least-Squares Approximation

HS = 67.5 Av - 1080. p

R 0.890

0 II I I

15 20 25 30 35 40 45 50o

Velocity change, &v (ft/sec)

': Figure 62. Input velocity change series, predicted

lumbar spine moment (flexion).

considered secondary when compared to the effect of peak input

acceleration or velocity change. Compiled data shown in Table 14

for the two series, 44 ft/sec and 34 ft/sec, indicate a lack of

Al significant trends although the sample size is small and the range

of rate of onset is limited. Only the seat stroke, plotted in

Figure 63, can be correlated to rate of onset with any statistical

confidence, and it is not certain whether this trend is actually

due to the rate of onset or to the peak accelerations for the input pulses, compared in Figure 64, which cannot be independently

controlled in an impact test. In fact, if the seat stroke measured in this series is normalized according to the sensitivity

curve for peak input acceleration, shown in Figure 48, then the

stroke values for the four tests at 44 ft/sec would be 10.4, 10.6,

10.5, and 10.4, indicating that the variation in this series can

be explained almost entirely by peak G rather than rate of onset.

The vertical acceleration response of the seat pan, shown in FiguZe 65, does not appear to be greatly influenced by rate of onset

of the input pulse. Similar acceleration-time history comparisons

for body segment accelerations also exhibit a relative insensitivity to race of onset changes, compared with the influence of acceleration magnitude and velocity change.

The effects of rate of onset tolerances on seat/occupant response

were investigated by comput: simulation and are discusssed in

Section 6.2.5.

0.

'".. 109
TABLE 14. COMPARISON OF RESULTS FOR RATE OF ONSET SPRIES

Duration of

Peak Total Seat Pan z

Rate of Input Velocity Seat - Acceleration

Onset Test Accel Change Stroke Seat Seat Pelvis Pelvis Chest Chest Head Head at 23 G

(G/sec) No. (G) (ft/sec) (in.) Pan x Pan z X z X z x z (sec) DRI

820. A80-057 35.0 43.6 9.5 23.1 27.3 16.4 31.2 21.4 24.1 28.5 27.1 0.008 18.4

980. A80-053 40.7 44.8 10.6 25.3 26.4 17.1 30.5 18.1 30.4 31.5 36.5 0.007 21.0

1040. ABO-054 40.2 44.2 10.5 22.1 29.4 10.1 30.8 24.1 25.8 38.5 28.2 0.016 19.7

1750. A80-058 44.6 44.3 11.0 24.3 29.8 17.4 37.2 26.8 26.1 32.2 33.5 0.007 20.6

750. AS0-070 31.2 33.9 5.6 20.4 20.9 13.7 26.8 10.7 24.8 19.3 29.1 0. 17.0

1490. A80-077 40.9 34.9 6.4 22.4 28.3 23.4 29.1 15.4 24.4 23.5 26.3 0.007 20.1

Rae Peak EA Peak Lap Peak Shoulder Peak Footrest Rate of •oad (ib) Belt Load (lb) Belt Load (Ib) -G Pf Sap eak Load ad(b (b)

Onset Test - ta

(G/sec) No. Right Left Right Left Right Left Woad (]b) x z

820. ABO-os7 1430. 1380. 348. 18i. 147. 201. 590. 1870. 2420.

980. A80-053 1380. 1590. 373. '113. 200. 147. 593. 1760. 2200,

1040. AS0-054 1420. 1550. 300. 233. 87. 250. 560. 1900. 2120.

1750. A80-058 1380. 1480. 496. 261, 214. 181. 523. 1950. 3210.

750. AB0-070 1380. 1470. 12Q. 74. 154. 174. 361. 1420. 1970.

1490. ALI3-077 1370. 1440. 148. 107. 141. 174. 389. 950. 2640.

0 NHoautrd viabl for 44 ft/ase "ries "a hQeattul valu•e for 34 r'taeo' KeO'LOS 0

ft.4A

f.O¢ ~~L"1 oo l $ $0 (1*1 i• t Lgoil o

Figure 63. Rate of onset series, maximum seat stroke.

110

. . ...... - , -
50

40 -A80-053

I ' - A80-057

36 / 8-5

0 20

4 10

0 Dow __

-10

-20 I

0 0.050 0.100 0.150 0.200

Time (sec)

Figure 64. Comparison of input acceleration pulses

for rate of onset series.

`40

so

40 A80-053

- - !AO-057

30 - A80-058

20 -

C':

10 -

-20

0 0.050 0.100 0.150 0.200

Tit* (see)

Figure 65. Comparison of seat pan -- axis acceleration

for rate of onset series.

"C 111
6.1.4 Energy Absorber Limit-Load Series

The energy absorber limit load is one of the most important parameters influencing the severity of the environment to which an occupant is subjected. The test series conducted in this program

provides only a portion of the data necessary to optimize seat

design: i.e., the relationship of seat stroke and transmitted

accelerative loads to the energy absorber limit load. The opti- . rmal goal in energy-absorbing seat design is a limit load set at

• .the highest possible level for the majority of occupant weights

to assure that the available stroke distance can be used to absorb the maximum possible crash energy before the occupant bottoms out. The other information required to achieve this goal,

namely the tolerable energy absorber limit load for the Army aviator population, will have to be derived from cadaver testing

"programs, human volunteer and primate experiments, and accident

data from aircraft equipped with energy-absorbing seats.

Severdl differing types and percentiles of anthropomorphic dummies were used in this test series; however, this does not seem

to have obscured important trends in seat performance. An effec-

"tive limit load factor was calculated for each test from the equa- tion: S-" 

= [Energy Absorber Load

LL Vertical fctive Occupant Weight + Movable Seat Weight

Results from this test series are compared in Table' 15. Several

of the variables, particularly seat stroke, and seat pan and

" pelvis z-axis acceleration are greatly influenced by energy ab- - sorber limit load. &eat stroke is plotted as a function of limit

load in Figure 66. Linear rsgression was used to provide a leastsquares approximation of the data; however, additional data plotted in Figure 66 indicete that the relationship between stroke and

limit load may follow a highr order function. These data were

calculated from two sources, i.a., the theoretical stroke cauation

for a triangular pulse derived ii Volume IV of the Aircraft t rash

Survival Design Guide (Reference 5):

""9L2 11 K

Stroke, S G m+L +i2 - )

where G peak input acceleration

t half oZ the input accelezation pulse duration

K = limit load factor/Gm.

112

•il•. •-;......ll .... .•l . P.. ......... S l *l• ,il ' - . '•l Si -* ,Ip S • 

5'e'' '= •'...-. .** '. ":.. .'>' ' ";- ,• ' '• = .' ."7 " '• "".5 "• * '7. .w. .. < . '- .". '.**... . . .. . ,. .. ' . ", "- . " ' ,-: - -. " ," +
TABLE 15. COMPARISON OF RESULTS FOR ENERG.Y ABSORBER LIMIT-LOAD SERIES

Duration of Effective

Peak Total Peak Acceleration (G) Seat Pan z Limit

Input Velocity Seat Acceleration Load

Test Test Accel Change Stroke Seat Seat Pelvis Pelvis Chest Chest Head Head at 23 G Factor

Parameter No. (G) (ft/sec) (in.) Pan x Pan z x z x z x z (sec) DRI (G)

50th%

14.5 G A80-053 40.7 44.8 10.6 25.3 26.4 17.1 30.5 18.1 30.4 31.5 36.5 0.007 21.0 14.6

N 50th %

14.5 G 1A0-054 40.2 44.2 10.5 22.1 29.4 10.1 30.8 24.1 25.8 38.5 28.2 0.016 19.7 14.6

50th %

1.1.5 G 180-055 39.2 44.1 14.2 24.6 28.1 17.1 20.7 24.1 20.7 50.5 18.2 0.018 19.4 11.7

5th %

14.5 G A1O-074 41.4 45.0 9.8 25.4 31.1 14.4 30.8 32.2 32.1 69.3*28.1 0.009 19.7 16.2

Sth%

8.5 G 180-075 42.5 45.9 16.1 26.0 26.6 22.2 34.5 47.6 68.7* 70.4* 70.0* 0.008 17.8 10.5

"95th%

14.5 G 180-073 40.2 45.6 12.0 28.1 25.1 19.4 41.2 15.4 29.8 53.6 20.7 0.004 18.0 12.0

"Peak E/A Peak Lap Peak Shoulder - Peak Footrest

Peak

Test Test Load (1b) Belt load (lb) Belt Load (lb) -G Strap load (Ib)

Pnramter No. Right left Right Left g Left Load (lb) x z

50th %

14.5 0 A80-g53 1380. 1590. 373. 213. 200. 147. 593. 1760. 2200.

50th %

14.5 0 180-054 1420. 1550. 300. 233. 87. 250. 560. 1900. 2120.

50th %

11.5 G ASO-055 1210. 1170. 174. 107. 221. 255. 517. 2190. 1950.

5th %

14.5 0 AO0-074 1,60. 1450. 154. 187. 174. 127. 481. 1190. 1820.

5th %

8.5 0 ABO,-075 6B8. 930. 201. 107. 94. 101. 436. 1490. 1890.

95th %

14.5 a AO-073 1390. 1440. 374. 314. 120. 130. 448. 1640. 2620.

*Amp•i•elas saturatod.

and from Program SOM-LA test simulations using the input accel-

"eration from A80-053, with the only variable being the energy

absorber limit load. If the relationship between E/A limit load

and stroke is not linear, then it becomes even more important to

set the limit load at the highest possible level compatible with

human tolerance because even a small reduction in load can result

in a significant stroke increase with a resulting increase in the

hazard of bottoming out. A nonlinear dependence of stroke on E/A

limit load would indicate that variable-load systems would be

quite beneficial, particularly for retrofit applications where

"stroke distance is usually restricted.

A definite trend is apparent in the seat pan z-axis acceleration

and DRI, presented in Figures 67 and 68, although neither parameter appears to be particularly sensitive to limit load. The

SOM-LA simulations described above predict some very interesting

and unexpected trends in terms of spinal load and moment. The

maximum lumbar spine compressive load, presented in Figure 69,

shows a direct relationship to limit load where the spinal load

113

7. =7 .777,777
. 25 I

5 o Measured value

0 SOM-LA prediction

20

0

""j 2 

0

00

S~Stroke equation for

' triangular pulse

Least-Squares Approximation SS - -1.05 L L + 26.1

R ,, -0.93

4 10 I , , I

0

6 8 10 12 14 16 is

Effective limit load factor, L. (G)

Figure 66. Energy absorber limit load series, maximum seat stroke. 4ý 0tt

S0~ ~ U

5 - - - LeastS are Approximatio as 

411

7..

ft. 0 o

8 10 S6 12 14 16 18

v. ~~Figure 66. Energy absorber limit load series,maiu

/• seat maxnu seat stroke.eain

~11

**%-
25

0

20- --f 0

10

5 Least-Squares Approximation

0 Measured value DRI - 0.403 LL + 13.9

0 SOM-LA prediction R 0.736

0."• I I I

6 8 10 12 14 16 is

"Z£tective limit load factor, L. (G)

Figure 68. Energy absorber limit load series.

Dynamic Response Index.

3000

o SOM-LA prediction

2400

~p1

.3 1800 0

'00

I 4-.-- I• ,LL - 1 .'.4

• " A -1200

0 I%

6 ao lat 14 16 10

tftetiw. limit load factor. LL (a) (

Figure 69. Energy absorber limit load series, predicted

maximum lumbar spine load (compression).

%115

il* ', l ..•*,, .. . ..... *..... . ... - 77 "~
"doubles from approximately 1200 to 2400 lb while the E/A limitload factor is increased from 8 G to 16 G. The spinal moment,

shown in Figure 70, predicts, unexpectedly, that the moment actually decreases for an increase in limit load and this function

. is not just linear but rather a higher-order function. After

analysis of this trend, it appears that the SOM-LA predictions of

maximum spinal load are strongly influenced by the amount of seat

stroke that occurs, which in turn is controlled by the energy absorber limit load. Spinal load is plotted against seat stroke in

Figure 71.

"3000

2400

N

LO 1800 cl

1200

DA

W

600

0 SOA-LA prediction

0 I, I 1 i

6 8 10 12 14 16 16

Eftective limit load faotor. L, (0)

Figure 70. Energy absorber limit load series, predicted

maximum lumbar spine moment (flexion).

The only test data available to verify these trends are from the

modified dummy test series conducted at Wayne State University.

Maximum recorded spinal loads and moments for two combined-mode 4., tests are listed in Table 16. The maximum spinal load is in direct proportion to the energy absorber load as predicted by the

SOM-LA simulations; whereas the predicted inverse relationship

between spinal moment and E/A limit load is not nearly as obvious

in the measured test data. It is apparent that further test work

with an instrumented dummy would be beneficial in verifying these

trends.

64,4"

7% 7;- '4 *

• " , 'W V V.V.'V" .
3000 1

2400

SIS00

0 1200

600

0 SON-LA prediction

0 • I , , I . .I I

0 5 10 15 20 25 30

Seat stroke, S (in.)

Figure 71. Energy absorber limit load series, maximum

lumbar spine moment versus seat stroke.

TABLE 16. MEASURED SPINAL LOADS AND MOMENTS FOR

WAYNE STATE UNIVERSITY MODIFIED DUMMY

TEST SERIES WITH COMBINED ORIENTATION

Limit Load Peak Spinal Peak Spinal

Factor Seat Stroke Load Moment

(G) (in.) (ib)- (in.-lb) _

14.5 4.5 1140. 913.

11.5 6.2 850. 938.

6.1.5 Test Facility Impact Conditions

The sensitivity of seat and occupant response to a particular input pulse shape is an important variable in terms of qualification testing. Each facility evaluating energy-absorbing seating systems will have a characteristically shaped deceleration pulse representative of the type of deceleration mechanism used. All facilities will be attempting to duplicate an idealized triangular

test pulse within some set of reasonable tolerances, such as those

specified in Reference 6; although MIL-S-58095(AV) does not presently set acceptable deviations for the input pulse. In addition

117 14.
to the input pulse shape, another factor that can affect seat and

occupant response is the decelerator orientation, i.e., the relationship of the deceleration direction to the gravity vector.

The parametric test plan for the input pulse shape test series

called for each participating facility to approximate the nominal

baseline deceleration pulse shown in Figure 72. The closest approximation of each facility to the prescribed conditions within

the constraints imposed by their deceleration mechanism is shown

in Figure 73. Seat and occupant response, as represented by the

seat pan and pelvis z-axis accelerations, are presented in Figure 74 for the four test facilities.

50--

42 G

40--

S30-

3

I.' KAv =42 ft/sec

U 20--

4

*.', IC

:: lO-_

".0

"". 0.062

- 0 .025 .050 .075 .100

Time (sec)

Figure 72. Nominal baseline input acceleration pulse.

.2

118

-9 AV- w 9 V .1** * '
s0o ____________ 50

40- 40-

I•::I

~30- 30--o ' .

*+ S

I- ' z

0 0

"20 - " 20

0- -J

w w

-to-. 10.

-10-- -10.

--2 I - I I

0 .026 .050 .ore .100 .126 .1O 0 .026 .060 .075 .100 .126 .160

TIME (SEC) TIME (SEC)

CAMI SIMULA

50 so_0_,_,

40.A 40.

30 3I0•\

3 0

A. to.. 10 1.-

0 10 I0,

I.+.

0+ olat .o6o .ore .too .1t, IS~O o .o24 .060 .o06 too .46, 1#* :

i• ~Time t logo:) time talc)+,;

NADC WSU

Figure 73. Baseline acceleration pulses for".

the four test facilities.

119 :

,"A PIA''
30.5 G

CAMI A80-053

Pelvis z acceleration

26.4 G

Seat pan z acceleration

38.8 G

NADC Drop 189

J. Pelvis z acceleration

36.3 G

4'

Seat pan a acceleration

31,0 0

Simula E/A Criteria Test 1

* .q Pelvis a acceleration

22.1 a

•I' Seat pan it acceleration

Figure 74. Comparison of pelvis and seat pan

" * z-axis acceleration for baseline test

conditions at three facilities.

120

ft* * ~ 4 * ft f- * " •1 ' ' 4•"•+ *,: "ft. ** ' .. ..... O •. .:• • 4 • • • " " ..
There is often a tendency to conclude that certain test facilities are better able to simulate the crushing characteristics of

aircraft by their ability to duplicate the triangular test pulse. The triangular pulse is merely representative of the waveforms

that have been observed in the major impact during aircraft crash

tests. The repeatability of a test pulse among facilities is far

more important than the ability to reproduce a specific crash

pulse.

Measured test values for the baseline tests at the four facilities are compared in Table 17. It is important to note that

there are significant differences in seat and occupant response

among facilities. Six parameters relating to test severity are ý.J plotted in Figures 75 through 80 for comparison to baseline tests

"conducted at the four facilities. There are obvious differences

in magnitudes of seat and occupant response that cannot be explained by the sensitivity curves presented in Sections 6.1.1

through 6.1.4. It is sufficient to conclude from this series

that facilities differ, that the effect of these differences on

seat performance is significant, and that it would be very difficult to adjust test data in order to compare tests conducted

at various facilities. If it is important to compare tests, then

they must be conducted at the same facility,

TASLI 17, COKFAa40" OP MS? PACIUTV BAutuN IFAt'T M 4W4IITIONIS

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121

* + •- i ,IV AP

,, .+ . .*- , ,,' .,-, ,, .+ .,. . .+ . .+ +. .. .+. ...... .. ..... . .. :.- ., ..... . "7. ,'..... . .. .,.
16

12 0

00 -. 0O0

-. 0 5 

0 8

4 .2 0

00

o Measured value

0 SOM-LA prediction

-"' CAMI NADC Simula WSU

Figure 75. Test facility impact condition series,

"maximum seat stroke.

40

0 0

3. 0 '--- - - -0... ... . .

00 0

S 0 00

0 0

I 20 S.. 

.9 10  C M easured value

0 SON-LA prediction

w0

CAJQ wAC Simula wSU

cigure 76. Test facility impact condition series,,

umaximum seat pan z-axis acceleration.

122
. 0.020

o Measured value

* Difference explained

0.015 0* in Section 5.1 and Figure 25.

0 0.0100

0.005 0

CAMI NADC Simula WSU

Figure 77. Test facility irpact condition series, duration

of seat pan z-axis acceleration at 23 G.

25

2O0 0

20 0

0o

• 15

10

o Calculated from measured acceleration

"A 0 S O M -L A p r e d i c t i o n

5

CAMT NADC Simula WSU

Figure 78. Test famility impact condition series, DRI.

"123

" ., . . ... .. . .... . . . ., .. .

• •,,'','". .*' _, .'-,, s.,-" .,,,.!...Ž t ,.'L. - .. '..'. . - , .. . .. .. ... ... -.",•.. . .• .. . . -. ' ' "..
3000

2250

0 • °

1500

'0

750

0 SOM-LA prediction

,~00

CAMI NADC Simula WSU

Figure 79. Test facility impact condition series, maximum

lumbar spine axial compressive load.

3000

":., 2250 0

P4 0

00

-,F

.4 0 1500

* 0°

"0 SOM-LA prediction

0

CAMI NADC SS.mul& WSU

Figure 80. Test facility impact condition series, maximum

lumbar spine y-axis moment (flexion).

124

" , ", " ' ' "*,""" "'"- ."' , "*" ".';." *."". " *". ". *""""'T. '" ' ""-'"' "" "" ' '" .7.: . . - • . •' •. ?, ',.., -,• ,,, .'•-,' . ,, s.,'",. ', ,'* ,- .' .... . ",.' .,,.*.* • , • .. . -•. -... -,". . . . . . . ," .. .. .. " "-'
.* 6.1.6 Dummy Type Series

Most of the testing for response comparisons of various dummy

models was accomplished using a rigid seat. As described in Section 4.1.1, the seat back was horizontal for all tests, thereby

aligned with the impact velocity vector on the CAMI sled facility.

No cushions were used, and the seat pan was mounted on a six-axis

load cell. In order to remove the dynamic effects of the stroking seat and the interaction between the occupant and seat, a

14.5-G trapezoidal-shaped input pulse was used. The actual pulse

from Test A80-026, which used the baseline Hybrid II dummy, is

shown in Figure 81; others were nearly identical.

20. 1 1-

15.

• 10.

r4

A 5.

U

*o.

.0. -5. -0

-10I I I

0 0.05 0.10 0.15 0.20 0.25

Time (sec)

Figure 81. Sled deceleration, rigid seat test A80-026.

As listed in Table 18, three 50th-percentile dummies were compared, and three tests were conducted with each dummy. The peak

dummy accelerations and seat pan force average for the three tests

are presented in Table 18. For a given dummy, measured acceleratit.s varied by a maximum of ±1 G from the mean. The acceleration

valv.es presented in Table 18 indicate that for x-axis response in

a vc-1.al test, the Hybrid II and VIP-50 are similar. However,

the e..zptomeric lumbar spine in the Hybrid II appears to significantly .iter the z-axis response. The peak z-axis accelerations

measured for the Hybrid II exhibit variations for the pelvis,

125

r4 . W
TABLE 18. AVERAGE PEAK ACCELERATIONS AND SEAT PAN LOADS FOR

RIGID SEAT COMPARISON TESTS OF 50TH-PERCENTILE

DUMMIES (AVERAGE OF THREE TESTS)

Dummy Peak Acceleration (G) Seat Pan

Descrip- Pelvis Pelvis Chest Chest Head Head z Force

tion x z x z x z (ib)

Hybrid II 9.6 42.9 6.6 26.2 12.0 33.4 2710

Alders~n VIP-50- 10.0 35.3 5.3 37.4 14.7 39.0 2960

"AlderIon

CG-50d 16.9 39.9 14.3 41.3 56.6 46.8 5500

Notes: 1. Test- A80-0251 A80-026, and A80-027.

2. Tests A80-022, A80-023, and A80-024.

3. Tests A80-032, A80-033, and A80-034.

chest, and head, while the essentially incompressible column of

the VIP-50 maintains a fairly uniform acceleration level throughout the body. Response of the older design CG-50 is significantly different from the other two dummy types, particularly in the

head acceleration and the load transmitted to the seat pan (as

shown in Figure 82).

A second series of tests for comparison of dummy types used three

95th-percentile dummies commonly used for dynamic qualfication tests of energy-absorbing seats: the VIP-95 (with elastomeric

spine), the CG-95, and the Sierra 292-895. These tests were conducted with the Black Hawk seat in the baseline test configura- tion. Results of these tests, presented in Table 19, are inconclusive due to the lack of complete data sets and small sample size. However, the seat pan response appears to be consistent

among the three dummies. Results from the two VIP-95 dummies cannot be directly compared since dummy 4a (Test A81-123) had been

reworked to include an Ensolite pad in the buttocks area, not

present in dummy 4b. Also, dummy 4a had been modified to include

load cells in the lumbar and neck regions.

6.1.7 Dummy Percentile Series

Evaluation of seat and occupant performance as functions of dummy percentile was complicated by the lack of similar anthropomorphic

test dummies. Two sets of baseline configuration tests were run

with a Hybrid II 50th-percentile dummy and a VIP-95 dummy with an

elastomeric spine similar to the Hybrid II configuration. Results

of the test series are presented in Table 20. In the final test

126
6000

I ,CG-50

Part 572

5000"

o-o-o-o-o VIP-50

14000--

44I

"- 3000-

> 0

S 2000- 0

S0000

00

0 0.05 0.01 0.15 0.20

Time (sec)

Figure 82. Comparison of seat pan force measured in rigid

seat tests with three 50th-percentile dummies.

TABLE 19. COMPARISON OF RESPONSE FOR 95T11-PERCENTILE

DUMMIES IN TIHE BASELINE TEST CONFIGURATION

Duration of

.umm Scat Peak Accolration AG) Stor Pan z Du seat J Acc keleration

Deone p- Stroke Seat Seat Pelvis Pelvis Chest Cheat Iiead Head at 23 G ton Test No. (in Pa x Pan z X . K 2 K I (scc) DRI

La tal Pan z2 I9.95a.632 (sac) 9.

Alderson AB1-123 1 '

3 13.2 29.6 28.1 24.8 35.0 41.5 26.8 N/M4 N/N 0.012 17.2 VIP-95 AO0-073°* 12.0 29.3 25.1 21.8 41*2 15..4 29.8 S3.6 32.2 0.005 19.7

Alderson

CO-95 A81-125 13.1 20.4 2)9.6 N/H N/N 57.0 34.2 N/N N/N 0.012 18.3

, 292-895 ASI-126 11.0 29.8 31.9 22.8 63.3 N;/H N/N N/N N/M 0.010 20.0

..Notes , 1. Dummy 4a (Onsolite pd under polvis, as noted in Table 4) with lumbar and neck

load cells.

2. Dummy 4b.

3. with alastomeric spine. 4. Not measured,

127

W ... .. ,....

W- C ' + . , - . . . - •
TABLE 20. SUMMARY OF THE EFFECT OF ANTHROPOMORPHIC DUMMY

PERCENTILE ON PEAK ACCELERATION VALUES

Duration

of Seat

Pan z

CAMI Seat Peak Acceleration (G) Acceler- CAMI eat -ation

Dummy Test Stroke Seat Seat Pelvis Pelvis Chest Chest Head Head at 23 G

Percentile No. (in.) Pan x Pan z x z x z x z (sec) DRI

50th A80-053 10.6 25.1 26.4 30.5 30.5 18.1 30.4 31.5 36.5 0.007 21.0

Percentile

A80-054 10.5 26.8 29.4 25.1 30.8 24.1 25.8 38.5 36.5 0.016 19.7

A81-1241 11.4 32.3 26.0 22.2 28.5 35.5 28.1 32.5 28.5 0.006 19.4

95th A80-073 12.0 29.3 25.1 21.8 41.2 15.4 29.8 53.6 32.2 0.005 18.0

Percentile 

A81-123 3 3 13.2 29.6 28.1 24.8 34.5 41.5 26.8 N/A N/A 0.012 17.2

Notes: 1. Load cell in lumbar region.

2. Dummy 4b.

3. Dummy 4a, Ensolite pad in buttocks, load cells in lumbar spine and neck.

of each series (test numbers A81-124 and A81-123) the anthropo- morphic test dummies were modified by the addition of a six-axis

load cell at the base of the lumbar spine which was used to obtain load and moment data. Measured load and moment data for

these two tests are presented in Table 21.

TABLE 21. SUMMARY OF THE EFFECT OF ANTHROPOMORPHIC DUMMY

PERCENTILE ON SPINAL LOADS AND MOMENTS

Peak Lumbar

Dummy Test Peak Lumbar Load (lb) Moment (in.-lb)

Percentile No. x _ z Resultant x _ z Resultant

50th

Percentile A81-124 360 60 1220 1220 360 1250 93 1260

95th

Percentile A81-123 593 153 1230 1260 653 1500 107 1510

When a 95th-percentile dummy is used in place of the 50thpercentile device, the increase in seat stroke of approximately 2 in. is predictable by the increase in vertical effective weight. However, one might also assume that there would be a general decrease in acceleration for the 95th-percentile dummy, although

4 

this 

Table is 

20. 

not 

The 

readily apparent from the peak values presented in explanation is that the peak values, due to dy- namic overshoot, are not strongly affected by the occupant size,

128

-ri- r - r~r ~ - .e.r.'r '
while the average acceleration during the crash sequence is lower

for a heavier occupant (provided that the energy-absorbing limit

load is fixed). This is evident from the seat pan DRI, which

shows a definite trend of lower values for increasing occupant size.

Loads and moments presented in Table 21 are intended for use in

validation of computer models and correlation with injury cxiteria, and are not intended for use as a quantitative measure of

spinal damage based on known vertebral segment strength distributions. The data presented indicate that axial spinal force

(z-axis) is not a function of occupant size, while the moment

associated with forward rotation of the body (y-axis) does show

"an increasing trend with body size. However, these data should

be used with caution because the chest x-axis acceleration measured in tests with the modified dummies shows a substantial increase in peak value. This is possibly an indication that the

installation of the load cell may have altered the rotational

stiffness of the torso. Also, as noted in Section 5.6, a second

load path carries part of the axial force and bending moment in

the torso, perhaps a greater percentage in the dummy than in a

human, where the abdominal region is not as stiff.

6.1.8 Cadavers Versus Anthropomorphic Dummies

A series of tests is being conducted to establish decelerative

spinal fracture loads for occupants of energy-absorbing seats,

as discussed in Cha.ter 2. That program has not yet been com-

"pleted, and the test data have not yet been analyzed in detail.

However, enough tests have been completed to determine characteristic responses of cadavers in the Black Hawk seat and to justify

the use of the Part 572 dummy as a human surrogate for seat testing. Although not specifically part of the sensitivity analysis being discussed in this chapter, the cadaver test program has

significant implications on the development of more rigorous seat

criteria. A comparison of cadaver and dummy test results is

presented here.

The cadaver test series is being conducted at Wayne State University under the baseline test conditions. Similar tests were conducted at WSU with the Part 572 dummy that was instrumented to

measure spinal forces and moments. Results of comparable cadaver

and dummy tests are shown in Table 22. Resultant peak body accelerations are presented for comparison because the accelerometer

orientation in the cadaver did not necessarily correspond to the

standard dummy coordinate system.

Required seat stroke values presented in Tables 22 and 23 indicate

that, in general, the Part 572 dummy requires slightly less stroke

distance than does the cadaver. The seat pan vertical accelerations, presented in Figures 83 and 84 for the vertical and combined tests, respectively, show that the interaction between the

129

W~4*" @ ~ -V too * * 9-f I. * '.

-. ~- 7 c--
TABLE 22. COMPARISON OF SEAT AND OCCUPANT RESPONSE

FOR DUMMIES AND CADAVERS

Duration of

Seat Pan z

Seat Peak Acceleration (G) Acceleration

Stroke Seat Seat Pelvis Chest Head at 23 G

(in.) Pan x Pan z Resultant Resultant Resultant (sec) DRI

Dummy,

Vertical 6.3 15.3 28.9 40.3 43.4 41.7 0.011 18.4

Cadaver,

Vertical 7.6 17.5 26.5 33.9 N/A 49.7 0.004 21.8

Dummy, Combined 4.5 27.8 23.9 35.2 31.6 46.9 0.004 17.8

Cadaver, 4.5 37.3 25.4 22.8 N/A 59.6 0.004 22.2

Combined 5.5 40.5 21.0 44.4* N/A 97.3* 0. 19.3

b*Impact between mouth-mount accelerometer and thigh.

".47 '4'

TABLE 23. COMPARISON OF SEAT STROKE FOR CADAVERS

AND DUMMIES IN THE UH-60A CREWSEAT

Dummy Tests

Cadaver Tests 5Oth% Hybrid II

Occupant Seat Occupant Seat

Weight Stroke Weight Stroke

Test Description (lb) (in.) (ib) (in.)

14.5-G E/A, Vertical Orientation, 166 7.6 164 6.3

42-45 G Peak Input Acceleration 160 7.4 164 7.0

140 7.1

148 7.1

14.5-G E/A, Combined Orientation, 166 5.5 164 4.5

42-45 G Peak Input Acceleration 140 4.5

11.5-G E/A, Combined Orientation, 218 9.4 164 6.5

42-45 G Peak Input Acceleration 141 7.0

160 9.0

8.5-G E/A, Combined Orientation, 200 13.1

42-45 i Peak Input Acceleration 143 9.7

165 8.9

I, 130

•7•• '-U- - • •'-% ' , i. '--"r"...-. '- 4 •% • •,• " • . -" " ". . ,. " "."," " . .. . . . 4

', ".•,,,.',".-, ,' ,,,*.",,..."• .4,<' 4 - '4 '4.,, .... , , ," .'"". 4 ,...,..'' " . ."-"... . . ....... .. . " ." . ."
32 ' I

28 Cadaver (AFO2O) S.. .. Cadaver (AF028)

S2424 Part 572 Dummy

o

"4 20

*112

r 2\

S o.,. .,I\ -

0 K8 • 1

-4

0 0.02 0.04 0.06 0.08 0.10 0.12 0.14

Time (sec)

Figure 83. Seat pan vertical acceleration for a Part 572 dummy

and two cadavers measured in vertical mode tests with

14.5-G energy absorbers.

32 ' , I .4"

28 --- -Cadaver (AF021)

-+---Cadaver (AF029)

"24 /,. - Part 572 dummy

'0"

'Z 20

16 , II

)+ +, 12 // ',,, \

.1 ii

+ ,8 . . . . . 1.. 

U

I N 00 \. /--9

4.2 00 .6 00 .0 0 1 0'.1

Figur etpnvrtclaclrtonfraPr 84 7

'i.M

I..

01.0 . .

* 1+ . . . ,) ,+ + 0.0 - % " 0.0 • 0 . 06) . . ', ". . . .

S. .. • • -. ,- ,- .,-.• . ,,., ... i, : ,. .) + ,( ,, ,~t iIm,, . e:: (. : _ .- •. .,.,,),.,, .j• ,. p• ,) ,,T , +,,o , i))
Part 572 dummy and seat pan is very similar to the response measured with human cadavers. Also, a comparison of the required

seat stroke for cadavers and dummies, shown in Figure 85, indicates that the occupant type has little effect on this parameter.

The limit load for the cadavers in Figure 85 is normalized by the

cadaver weight; that is, the effective limit load is equal to the

measured energy absorber load divided by the effective vertical

weight of the seat and occupant. The comparison between body

accelerations for the dummy and cadaver does not show a good correlation.

30

Cadavers Dummies

25 A

*'0'

20

E Vertical mode cadaver tests

0. 0

Comb ned '50., ,

modocadverteat 8 0 . ---. •

0 I , I , I, I I

78 9 10 11 12 13 14 15 16

I E~•tectivo onerg7 absaorber lim~t load ccli

Figure 85. Comparison of normalized seat stroke for

cadavers and dummies tested with the

Wayne State University input pulse shape.

Results of this limited comparison between occupant types indicate that seat performance criteria based on seat pan accelera-

•tion may not be as sensitive to occupant type as a criterion

"based on body segment acceleration. However, it may also indicate that injury mechanisms within the body, e.g., spinal deformation, cannot be reliably predicted from seat pan acceleration,

as internal body response can vary significantly for various oc-

* cupant types with similar inputs from the seats.

44r

*.'1 132

S':5
6.1.9 Seat Orientation Series

In order to investigate the effect of aircraft (i.e., seat) orientation on seat and occupant response, tests were conducted

with both rigid and energy-absorbing seats. The baseline rigid seat tests at CAMI were conducted with the seat vertical axis

(and seat back) parallel to the impact velocity vector. Tests

were also conducted with the seat vertical axis pitched forward

30 degrees and rolled 10 degrees. Results are compared in

Table 24 for both the Hybrid II and VIP-95 dummies. The input

conditions, in terms of sled velocity and deceleration, were

held fairly constant as the orientation was changed. Both the

seat pan force and resultant pelvis acceleration were essentially unaffected by the change in orientation. However, the resultant

accelerations of the chest and/or head increased considerably in the pitched and rolled configuration, tending to indicate the

increased severity of the environment.

TABLE 24. COMPARISON TEST RESULTS FOR SEAT ORIENTATION SERIES, RIGED SEAT

Seat Peak

Pitch/ Sled Initial Seat Pan

Roll Test Dummy Accel Velocity Pelvis Pelvis Pelvis Chest Chest Chest Head Head Head z-force

(deg) No. Percentile (G) (ft/sec) X Z Resultant x z Resultant x z Resultant (ib)

0/0 ABO-026 50 14.4 42.1 8.22 42.1 42.9 6.04 25.7 26.4 12.1 32.2 34.4 2,650

0/0 A80-027 50 14.4 41.9 6.11 43.3 43.7 7.05 26.7 27.6 11.8 34.0 36.0 2,550

30/10 A80-038 50 15.6 42.6 15.8 38.4 41.5 29.2 24.8 38.3 18.6 31.4 36.5 2,580

"" 0/0 A80-031 95 14.1 42.1 3.70 30.9 31.1 7.4 25.1 26.2 10.4 26.4 28.4 2,880

30/10 ABO-037 95 14.7 42.3 17.5 25.2 30.7 20.2 22.5 30.2 23.1 25.8 34.6 2,850

Orientation of the energy-absorbing seat was also varied; two

orientations were used at CAMI and three at NADC. Results for

this series are summarized in Table 25. At CAMI, Test A80-078

maintained the same resultant sled acceleration and velocity as

the baseline case, while Test A8O-081 increased the resultant to

maintain the same vertical components. In both cases, x- and z- accelerations were increased above the levels measured in the

baseline configuration. At NADC, the drop height and, thus, the

impact velocity and acceleration were held constant while the

seat orientation was changed. There, as at CAMI, measured accelerations increased with increasing pitch angle, indicating an

environment increasingly severe to the occupant. The same trend

was not observed for the DRI or duration of seat pan acceleration

above 23 G.

As shown in Table 25 and in Figure 86, the effect of orientation

on seat stroke is evident. In Figure 86, results of those tests

with similar resultant sled impact conditions are plotted. The

reduced seat stroke with increasing aircraft pitch is expected

as the vertical component of input acceleration is reduced, and

probably accounts for increased seat and dummy accelerations.

133

. . ' .... " .... " . . " " " 0 "V " "4 W ' •PA
TABLE 25. COMPARISON OF TEST RESULTS FOR SEAT ORIENTATION SERIES WITH ENERGY-ABSORBING SEAT

Duration of

"Seat Peak Total Peak Deceleration (G) Seat Pan a

Pitch/ Input Velocity Seat Acceleration

Roll Test Decel Change Stroke Seat Seat Pelvis Pelvis Chest Chest Head Head at 23 G

(deg) No. (G) (ft/sec) (in.) Pan x Pan z x z x z x z (sac) DRI

17/0 A80-053 40.7 44.8 10.6 25.3 26.4 17.1 30.5 18.1 30.4 31.5 36.5 0.007 21.0

17/0 A80-054 40.2 45.2 10.5 22.1 29.4 10.1 30.8 24.1 25.8 38.5 28.2 0.016 19.7

34/10 A80-078 39.9 44.0 6.7 30.2 34.1 30.5 49.9 37.2 40.2 29.8 50.4 0.010 18.7

34/10 A80-081 45.2 53.0 9.0 29.4 41.8 36.9 50.9 59.6 36.2 100. 46.8 0.015 20.6

0/0 N-195 42.7 41.0 12.1 18.1 34.5 17.7 34.4 4.67 33.7 15.4 37.7 0.006 21.1

"13/0 N-189 44.9 40.4 10.8 16.6 36.3 15.0 38.8 10.1 36.1 24.7 40.8 0.006 20.8

13/0 NJ-190 45.0 40.5 11.6 16.8 34.6 18.7 40.1 14.7 37.4 22.7 42.8 0.007 20.7

30/10 N-196 46.7 41.3 7.3 28.8 35.5 18.7 43.1 26.8 40.8 33.1 44.8 0.014 20.3

a? Peak &A Peak Lap Peak Shoulder Peak Footrest

.:.Pitch/Rol 

ih lt I Test .

load (ib) ... Belt Load (lb) Belt Load (1b) GSap G Strap Load (ib)

gAngle No. Right Left Right Left Right Left Load (1b) z

N 17/0 A80-053 1380. 1590. 373. 213. 200. 147. 593. 1760. 2200.

17/0 A80-054 1449. 1550. 300. 233. 87, 250. 560. 1900. 2120.

34/10 ABO-078 1330. 1400. 200. 887. 620. 413. 613. 1930. 2120.

34/10 ABO-081 1370. 1430. 707. 1550. 400. 540. 773. 2020. 2470.

0/0 N-195 1410. 1420. 348. 462. 26.8 40.2 610. 1740. 2330,

13/0 N-189 1449. 1457. 342. 335. 188. 221. 476. 1670. 2310.

13/0 N-190 1370. 1415. 361. 508. 181. 355. 515. 1890 2410.

30/10 N-196 1480. 1530. 100. 616. N/M* 107. 449. 1650. 2250.

'Not tulsuured.

The effect of orientation on spinal loads is shown in Table 26,

where results of two CAMI tests with the modified 50th-percentile

dummy are presented. Note that in Table 26 "flexion* refers to

forward bending of the torso which causes the chest to approach

*= the knees. *Extension," on the other hand is a backward bending

so that the distance between the chest and the knees is increased. 4,,• Although the vertical component of input velocity and acceleration

"remained approximately the same, the axial compressive load in the

lumbar spine increased by 43 percent. With the greater horizontal

input, the maximum lumbar moment was increased by 84 percent.

6.1.10 Ramped Energy Absorber Series

Tests were conducted at both CAMI and NADC using energy absorbers

of *ramped* design, whose loads increased linearly with stroke.

134

% . .... '. . . - q m. m .m. •- 1 mm.. • m . . .i. . . . . .
15 1

Av- 42 ft/sec

12 AL

0 . °

9

U 0

0

4 0

U, 4j 6

3 0 CAMI test Least-Squares Approximation

A NADC test S = -0.175 0p + 13.0

0 SOM-LA prediction R a -0.95

0_. I , I

0 5 10 15 20 25 30 35

Seat pitch anglo, 8 (DEG)

p

%I

Figure 86. Seat orientation series, maximum seat stroke.

TABLE 26. IWMT OF SEAT ORIENtATITON O LUMIAR FORK•S AND MOMEWS

FInput (duset R tAel-ration (0t

17/0 ABI-124 45.0 11,2 -43.0 41.3 -12.1 39.5 1220 360 1255 843

S 34/10 AS-127 $2,4 20.9 -42 .8 4.0 -26.4 3942 1740 577 2307 1395

As illustrated in Figure 18, the load for the Type 1 device increased from a load factor of 9.5 G to 19.5 G through approximately 12 in. of stroke. The steeper slope of the Type 2 device

increased from 7.25 G to 21.7 G in 12 in. of stroke. The energy absorber force-time histories measured in the CAMI tests are

compared in Figure 87 with that for the baseline vertical test.

135

. . 4 .'. . 4 -.-. . 4, .. . . ..
Baseline

A80-054

J~1390 lb

Type I RaMp

A80-065

As0-066

'.4"

* .4•.

',• Figure 87. Comparison of energy absorber force-time

histories for ramped El& series*

Although the load-deflection characteristics. of the two ramped

-,devices were different* the trends in test results were similar,

shown by the data in Table S~As 27, these special devices performed

less "efficiently than the standard square-wave typedei.Th

ramped devices caused the Part 572 dummy to utilize more than

•: 1.5 in. additional stroket while the measured accelerations and

:.•. calculated DR1s were actually higher.

::.• 6.1.11 Movab~le' seat Weight. Series

:- The seat veight variable was included in the parametric test

•: series because of the effect it has in performarice of several pro-

,4 duction seats. Crewseat qualification tests for the Army UH-60A

• ; and AN-64A helicopters and Navy SH-608 helicopter indicated that

.• the maximum •seat pan z-axis acceleration was a function of seat

136

1. -

4.•.,. . •,. ' F .:.t •:,•I .. ',.• ... " 1 .... ' o .. • , . . , . . e ... , ,. ..
TABLE 27. COMPARISON OF RESULTS FOR RAMPED ENERGY ABSORBER LOAD SERIES

Duration of

Energy Peak Total Seat Pan z

"Absorber Input Velocity Seat Peak Acceleration (G) Acceleration

Plastic Test Accel Change Stroke Seat Seat Pelvis Pelvis Chest Chest Head Head at 23 G

slope No. (G) (ft/sec) (in.) Pan x Pan z x z X z x z (sec) DRI

Type 2* A80-066 41.9 43.7 12.3 27.9 34.8 19.4 48.8 19.7 34.8 26.8 24.8 0.018 27.4

Type 1* A80-065 40.9 43.4 12.0 24.2 32.3 18.7 43.5 31.4 26.8 36.1 20.1 0.009 23.7

N-194 46.6 41.4 11.8 13.4 28.5 16.4 30.8 13.0 28.4 24.7 20.1 0.027 27.0

Baseline* A80-053 40.7 44.8 10.6 25.3 26.4 17.1 30.5 18.1 30.4 31.5 36.5 0.007 21.0

A80-054 40.2 44.2 10.5 22.1 29.4 10.1 30.8 24.1 24.8 38.5 28.2 0.016 19.7

Peak E/A Peak Lap Peak Shoulder Peak Peak Footrest

Load (ib) Belt Load (lb) Belt Load (lb) Goad (ib) Test Test -_~d(b -~G Strap

Parameter No. Right Left Right Left Right Lett Toad (ib) x z

Type 1 A80-066 1960. 2260. 415. 449. 261. 147. 744. 1940. 2360.

Type 2 A80-065 1730. 1740. 321. 401. 254. 114. 562. 1900. 2340.

N-194 1770. 1910. 509. 563. 188. 18C. 610. 1390. 2220.

Baseline A80-053 1380. 1590. 373. 213. 200. 147. 593. 1760. 2200.

A80-054 1420. 1550. 300. 233. 87. 250. 560. 1900. 2120.

*See accompanying load/deflection curves.

bucket inertial and stiffness properties. These properties play

a major roie in the interaction between the occupant and seat pan

which influences the body dynamics.

2 The ideal parametric test series for this variable would have included a seat bucket in which the weight and stiffness could be

varied independently, and energy absorber loads could be adjusted

to maintain a constant energy absorber limit load factor as seat

weight was changed. Unfortunately, constraints placed on this

research program made it prohibitive to adjust seat stiffness or

to design energy absorbers with the correct inversion loads to

maintain a constant limit load factor of 14.5 G. Therefore, in

this test series, the effect of seat bucket stiffness on performance was not investigated. Movable seat weight was modified by adding steel plates to the Kevlar portion of the UH-60A production

seat from which the armor had been removed. Seat weights for this

series were 35.0, 60.6, and 120.0 lb, with 35.0 lb being the unmodified weight of the Kevlar bucket and 60.6 lb corresponding to

the weight of the production armored UR-60A seat. The energy absorbers for this series all had the same nominal load correspond- ing to 14.5 G for the 60.6-lb seat. The effective lizait-load factor varied in this series from 16.6 G for the 35-lb bucket to

11.2 G for the 120-lb bucket. Therefore, the measured test re-

".1 sults for this series included the effects of two variables: the

movable seat weight and the energy absorber limit-load factor.

%I i The sensitivity analysis for this series was conducted by plot-

"ting trend lines for the test data and also showing the expected

137

?]19.Mr ~ ' I '4' 5 15 5
sensitivity due to the change in E/A limit-load factor which was

deduced from the curves presented in Section 6.1.4 for the energy

absorber limit-load series. It was assumed that any differences

in the trend lines must be attributable to the effect that the

movable mass of the seat bucket has on seat and occupant interaction.

"The test values for this series are compared in Table 28, and

sensitivity curves are shown in Figures 88 through 92. The seat

stroke curves shown in Figure 88 indicate that the trend is entirely due to the effect of energy absorber limit load and not

"seat weight. Therefore, if the energy absorber loads had been

modified to maintain a constant limit-load factor, then similar

stroke distances would have been achieved. This may appear to be

an obvious conclusion; however, there was uncertainty at the beginning of this effort concerning the effect that body dynamics

might have on stroke.

TABLE 28. COMPARISON OF TEST RESULTS FOR SEAT MOVABLE WEIGHT SERIES

EffectDuration of

'-a Peak Total k Acceleration Seat Pan z

Movable Input Velocity L1, -G Acceleration

Weight Test Acc.l Change Factor Seat Seat Pelvis Pelvis Chest Chest Head Head at 23 G

(lb) No. (G) (ft/sec) (G) Pan x Pan z x z x z x z (spc) DRI

35.0 A80-060 41.5 44.5 16.6 35.1 29.6 24.0 42.8 2..7 26.8 37.8 28.9 0.012 21.9

60.6 ASO-053 40.7 44.4 14.5 25.3 2A.4 17.1 30.5 18.1 30.4 31.5 36.5 0.007 21.0

"60.6 A80-054 40.7 44.2 14.5 22.1 29.4 10.1 30.8 24.1 25.8 38.5 28.2 0.016 19.7

120.0 A80-061 40.7 43.8 11.2 23.4 22.9 17.1 32.8 38.5 21.1 43.5 22.4 0. 16.4

oaba Peak B/A Peak Lap Peak Shoulder Peak Footrest N- ovable Seat ft k

ight Test $t Load (Ilb) Belt Load (1b) Belt Load (1b) tlb) load 

(1b) No*. (in.) Right left -Right Left Right Left 'Load 11b) X z

35.0 ABO-060 8.6 1400. 1450. 255. 389. 181. 147. 482. 1930. 2520.

60.6 A9O-053 10.6 1380. 1590. 373. 213. 200. 147. 593, 1760. 2200.

60.6 AB0-054 10.5 1420. 1550. 300. 233. 87. 250. 560. 1900. 2120.

120.0 A80-061 14.0 1370. 1430. 52). 389. 255. 255, 791. 1720. 2450.

The next two parameters, maximum seat pan z-axis acceleration and

Dynamic Response Index-IDRI) (shown in Figures 89 and 90), are

both based on the performance parameter, seat pan acceleration,

used in the current criterion. The trend seen in recent qualification testing that additional weight in the seat bucket acts to

*• reduce the oscillations of the seat pan is present in the parametric test series. This is exhibited by the fact that the two

sensitivity lines diverge as weight increases, indicating that

the reduction in vertical seat pan acceleration cannot be entirely ,., explained by the reduction in energy absorber limit-load factor.

The seat weight does appear to play a significant role in reducing

accelerations measured on the seat. However, the final two graphs

138

. . l "i . . -, 9. 9 , - p ' w. , -.

""""" ".."'.

7. -7 7
20

0 Measured value

0 SOM-LA prediction 0

16

1

41

W Sensitivity due to limit

load modification

4 S~S Least-Squares = 0.062 W + Approximation 6.64

R = 0.997

I I .. . 4

0 20 4) 60 80 100 120 140

Movable seat weight, W (lb)

Figure 88. Movable seat weight series, maximum seat stroke.

0 Measured value

0 SON-LA prediction

40

2 0 * C° .4

Sensitivity due to limit • • load modification

10 -Last-Squarta 6pproximation

G -- 0.080 W 32.6 OP Rt - -0.919

o..

0 20 40 60 s0 100 120 140

Movable seat Veight, W (lb)

Figure 89. Movable seat weight series, maximum

seat pan z-axis acceleration.

139

"W1

A. %
25

0 Measured value

'44 •-.. 0 0 SOM-LA prediction

--'.

N20

15

10

"_____Sensitivity due to limit

load modification

5. Least-Squares Approximation

DRI = -0.065 W + 24.2

R = -0.975

0 I I

%0 20 40 60 80 100 120 140

Movable seat weight, W (lb)

Figure 90. Movable seat weight series,

Dynamic Response Index.

"2500

-0 0 SON-LA prediction

% j ,2000

1500

1000

S " Sensitivity due to limit

"load modification

-- -. Least-Squares Approximation

0L0e, SSO0 -7.66 W 4 2610.

R -0.962

0

0 20 40 60 so 100 120 140

Movable Seat veight. (1t) 'b)

"Figure 91. Movable seat weight series, maximum lumbar

spine axial load (compression).

140

'3⁄4%
-2500

0 SOM-LA prediction 0

2000

1500

4)

c

0 1000

14

Sensitivity due to limit ra load modification

500 Least-Squares Approximation

M = 7.81 W + 1480.

R = 0.809

tIII I I

0 20 40 60 80 100 120 140

Movable seat weight, W (lb)

Figure 92. Movable seat weight series, maximum

lumbar spine moment (flexion).

"presented in this section for spinal loads and moments indicate

that the apparent reduction in seat pan acceleration is not passed

on in the form of reduced spinal loading. The trends seen in

these curves can be almost entirely attributed to the change in

energy absorber limit load.

There are two conclusions to be drawn from this series. First, the results of this series support the theory that the maximum

seat pan accelerations occur because the occupant is not fully coupled to the seat pan, and these accelerations actually have

little effect on the maximum loads to which the body is subjected.

The second conclusion is that the practice of using the measured

seat pan accelerations for determining hazard potential is inaccurate. This indicator is sensitive to localized conditions which

are a function of bucket mass, and possibly also stiffness. Loads

and moments measured in the spine are not sensitive to these localized, design-dependent conditions.

6.1.12 Seat-Frame Spring-Rate Series

For one test at CAMI, the top of the Black Hawk seat frame was

secured rigidly to the sled with steel brackets, as illustrated

in Figure 93. The intention of this test was to isolate the effect of longitudinal seat elasticity on seat and dummy response by comparison of results with those for the baseline configura- tion.

141

.,**•.. . . *; • * + *. .+4,+ *'. *.. , . . . . , + .+ . .. l, + , •. .,. .4 . .• . . ,+, . - . , , . + lli+ 'qlI +- + l I il • # +l . • * 0 P l l % I iI + I • + • .• . + , . •... ". #qt*`A* "I1+ .I .6%,*,: i,+'t * "- + ++ " ' ::":: ., ;:,.+ : . . ,+ , , ,... .. , - + •, ++'+ ,u''i ,,+ J :,= •""+*+ " ¶. * • +" , t+.+' ._'• . .*" 5 *-'• ,* .*.* '+'''+++" .. 4 / , ''.4 +,+ L". . . - ."' •/ L,, ':• P
OCCUPANT ACCELEROMETERS -4001

HEAD, TRIAXIAL I 3-AXIS LOAD • iFOOTREST CELL

CHEST, TRIAXIAL

PELVIS, TRIAXIAL • BELT

'*"-; ', • j•TRIAXIAL

'.,•"•ACCELEROMETER

DISPLAEMENTPRIMARY AND

,DISPLACEMENT REDUNDANT SLED

TRANSDUCERS ACCELEROMETERS

We

•.•AND 0

,•.* \•-SEAT FRAME

PLATFORM TRIAXIAL ACCELEROMETER SLED TRAVEL

.- ;-:;~ ~/17'*8E TAE

Figure 93. Test configuration (for stiffened

seat frame test).

Results are summarized in Table 29, where it appears that seat

.V stroke and DRI were essentially not affected by the frame modification. However, x-components of dummy pelvis and chest acceleration were reduced, as shown in Figure 94, while z-components

were increased.

The results of this test are not conclusive because, although the

frame was stiffened, seat bucket elasticity remained a factor.

6.1.13 Seat Cushion Stiffness Series

Two tests were conducted to determine the effect of seat cushion

stiffness. For one CAMI test (A80-068) a block of rigid plastic foam was contoured to cushion shape; for another, layers of

soft upholstery foam were used. In designing the cushions, an

effort was made to achieve a thickness that would seat the dummy in the same position as in the baseline case. Force-deflection

characteristics for each of the cushions were determined by com-

.1 pression under the VIP-95 dummy lower torso segment, as shown in

Figure 95. load-deflection characteristics for the two cushions

are compared in Figure 96 with the baseline configuration, the

production UH-60A Black Hawk cushion.

'1 142

,•. .... S ,.......- S S.................................-.

. - , . , * • 5 * ,: * • *-' • • " • S* ; .•'* "" . .' . -. "j > .) 5* =** " . .9*" "
TABLE 29. COMPARISON OF TEST RESULTS FOR STIFFENED SEAT FRAME SERIES

"Duration of

Peak Total Peak Deceleration (G) Seat Pan z

Seat Input Velocity Seat Acceleration

Frame Test Decel Change Stroke Seat Seat Pelvis Pelvis Chest Chest Head Head at 23 G

- Stiffness No. (G) (ft/sec) (in.) Pan x Pan z x z x z x z (sec) DRI

Baseline A80-053 40.7 44.8 10.6 25.3 26.4 17.1 30.5 18.1 30.4 31.5 36.5 0.007 21.0

I 80-054 40.2 44.2 10.5 22.1 29.4 10.1 30.8 24.1 25.8 38.5 28.2 0.016 19.7

Stiffened A80-059 40.4 43.7 10.4 28.1 32.2 21.1 37.5 11.4 34.2 20.1 36.8 0.013 19.4

Seat Peak E/A Peak Lap Peak Shoulder Peak Footrest Forward

tload (lb) Belt load (ib) Belt Load (lb) "leak Load (lb) Displ. at

Frame Test -G Strap top of

. Stiffness No. Right Left Right Left Right Left Load (lb) x z Guide Tube

Baseline A80-053 1380. 1590. 373. 213. 200. 147. 593. '760. 2200. 1.40

A80-054 1420. 1550. 300. 233. 87. 250. 560. 1900. 2120. 1.51

-- •.4

Stiffened A80-059 1390. 1470. 570. 429. 141. 248. 717. 1850. 2230. 0.00

---- Baseline

"-•Stiff frame

.', ¶ Pelvis x acceleration

LI

"IA"

01 %p

z m1'

Seat pan a acce•eration

Figure 94. Comparison of response with stiffened seat frame.

* *143

". .... 1 - .... ...... 
- ft

-- .•.-

Figure 95. Measurement of seat cushion forcedeflection characteristics.

Results for the cushion series are compared in Table 30, where ,, no variation in seat stroke or DRI from the baseline case is apparent. The rigid cushion increased the natural frequency of

the system, so that acceleration waveforms exhibited both higher .*' frequency and greater magnitudes, as shown in Figure 97 for the

seat pan and Figure 98 for the pelvis. Also, the soft cushion

permitted the dummy to reach a significantly higher velocity relative to the seat before bottoming out on the seat pan. The - result was an increased pelvic acceleration, as shown in Figure 98.

Although only one actual seat cushion was compared with two extremes of stiffness, comparison of pelvic acceleration data indi-

"cates that the baseline cushion performs significantly better than .. one that is either very rigid or very soft.

•, 6.2 ADDITIONAL PERFORMANCE VARIABLES

section SThis presents an examination of six variables that, although not originally conceived as part of the test program, would

•144

-ft."-U 1W. U V V*WW

ft. f "t •.••;•1-""'".'•.. •,'•......'1•........•............•" ,•' ,'....... ,. - ,•- -,,..... -

-I f".'t-' .'" .,,,'*'v'.',:.'f. t.tt~tt .- '.vf v ." ".. " .. ' .t . ' t . f.v t.'. . v ' . . . v ' . .. v . . vt
"3000 I 1

Baseline

- ----- Rigid cushion

-.--.- Soft cushion

0-0-0 VIP-95 buttocks

2500

:1 /

-2000 01

II

00 0

1500 o

S101

1100020 2.

0%

4'..! //

500 ."b

0 0.5 1.0 1.5 2.0 2.5 3.0

Deflection (in.)

I.

Figure 96. Load-deflection curves for cushions

used in cushion spring rate series.

STABWL 30. COMPARISON OF TEST RESU;TS FOR CUSHION SPRING RATE SERIES

Duration of,

Peak Total Seat Pan z

Input Velocity Seat -Aceleration

Cushion Test Decal Change Stroke Seat Seat Pelvis Pelvis Chest Cheat Head Head at 23 G

Stiffness No. (0) (ft/sac) i iLn Pan x Pan ax x a a (sac) DRI

Soft A80-067 41.2 44.0 10.5 28.8 26.1 13.7 43.6 25.1 32.5 30.2 46.9 0.008 19.5 .

Bahsline A80-053 40.7 44.8 10.6 25.3 26.4 17.1 30.5 18.1 30.4 31.5 36.5 0.007 21.0

A80-054 40.2 44.2 10.5 22.1 29.4 10.1 30.8 24.1 25.8 38.5 28.2 0.016 19.7

Rigid , 0o-068 41.2 44.1 10.6 29.3 29.0 25.4 49.2 26.8 31.8 20.5 28.2 0.008 20.4

* Peak E/A Peak Lap Peak Shoulder Peak Footrest

toad (Ib) Belt j Lj ( b Halt Load (lb) - e tr•a [jd - b)

Stiffness No. Rijht Loft Right Loaf t Right 1Af t j2LLoa (bL

Soft A80-067 IS20. 157U. 374. 174. 261. 187. 555. 1810. 2320.

"B. faseine A80-053 1380. 1590. 373. 213. 200. 147. 591. 1760. 2200.

A8O-054 1420. 1550. 300. 233. 87. 250. S60. 1900. 2120.

Rigid 480-068 1380. 1460. 509. 389. 322. 355. ?S0. 1970. 2260.

5, 145

,.`..

*. . **.*.,~..*.* . .. ,, ... .... ... ...

* . b..S Ais ~*I~ ~ t%
4%:

0.

0

rn 0 _ I H 0to

_•-jo .fu• co .,•

•%':> S.... I

'444

04Q

0*

4-C)

t54 (1)w U

.. o 44

0~.9-I

0 1 4J

u N W)

U co

Uu c

* .*00 5. ,,a'o m2 "

o o 0 

.,

;:• 146

7N oo

.. , , , 0. B -" °;,• i." .' ..... ' ... ., -"'44.. ... .. , . . '- d. l....- . ... .. .. . ...- - •, .,- .• ,,.. oo., , , , ,•,.,, . .. . , , ., % ,,..,. , ,.,. ; • .• 00 .. ,, . - . . -".. ,.. ,.-. .

*;•-.•¶:,• • .,, ,','..'., .'. . ,'% .-.% ,, ••% ,,, ,• •, •,• ," " ." ," C ." •,,. ,• ." • ," ,- , .4-')., 0',
be significant in developing a new seat performance specification.

The degradation of dummy performance after repeated testing was

determined by examining comparable tests at various intervals in

the test program. Program SOM-LA was used to provide a simulated

parametric test matrix for four of the parameters. In the final

section, a comparison of calibration techniques for accelerometers

is presented with an example showing the magnitude of performance

differences that can occur.

6.2.1 Dummy Performance Degradation

The performance repeatability of the Hybrid II dummy, S/N 870,

was examined by conducting rigid seat baseline tests with this

dummy at intervals in the test program. The tests were designed

to ensure that multiple tests would not degrade the performance

to such an extent that the results of the parametric test matrix

would be suspect. Figures 99 and 100 show a comparison of the

force applied to the seat pan by the dummy buttocks and the resultant pelvic acceleration for tests conducted at three periods

in the program. Test A80-025 was the first test in the research

effort with the Hybrid II dummy. Twenty-two tests were subsequently conducted at CAMI with this dummy, followed by three tests

(beginning with A80-083) repeating the conditions of A80-025 to

compare performauce. The final tests in this series, beginning

with A81-004, were conducted following the 12 NADC and Simula

energy-absorbing seat tests.

4500

A80-025

- --- A80-083

--. A81-004

S3 0 0 0 -

0

*1 1500 .

0. .050 .100 .150

Time (sec)

SFigure 99. Comparison of seat pan z-axis force for

"dummy performance degradation tests.

147

•"÷ .•'".,•B ,," 4 *, *B ..... • .... •4* 0 ...... ,O...................•...... O . ..,e ... l , .. . g, .. - *4,.• • ,'.,4'. a. * * * *-. , - ' .- . k *• • • - - •'- .-- ' -. *-. . .I. ' . •. .- 1⁄2 T•. .•
'60

60 - - " A80-025

A80-083

-.... A81-004

40

$4 ii .1

S20

LI 5

0

0 .050 .100 .150

Time (sec)

Figure 100. Comparison of pelvic resultant

acceleration for dummy performance degradation tests.

Physical degradation in the form of a decrease in buttock cushioning properties did occur by Test A81-004. It was theorized that

this reduction in cushioning might result in a large dynamic overshoot in the rigid seat test; therefore, the restraint system was

given more preload in Test A81-004 than in the previous rigid seat

tests. The result of this modified test procedure was a decrease

in the measured seat pan force for Test A81-004. However, the

accelerations recorded in the dummy pelvis do not vary notably in

any of the three tests illustrated in Figure 100.

It was concluded from this series that performance of this dummy

was not significantly altered by repeated testing. A standardized,

instrumented dummy based on the Part 572 specification should be

capable of providing repeatable results for many qualification

tests. However, it is suggested that a periodic examination of

the dummy be conducted followed by replacement of degraded components.

6.2.2 Footrest Location and Angle __

It was noted during the cadaver test program that the lo'diiveasured on the footrest varied considerably depending on o,.cupant

anthropometry. This indicated that the load path through thQ legs

reacting vertical and longitudinal inertial forces could be altered by the relative angles of the leg segments. A series of

148

554 % -Xw ' S '-. •.'.- ...... I " •• .. '1 " "*.. . . ' 5I' .5 .S..... 5 .s" •.V.. .... • ss ... .. 5a. ... ..

* A .',','...t 4, -'.. , s,I,•,, s,.' ,, ,, . ,5 " . 5 .• *•• ... ,, -..
six SOM-LA simulations was used to examine the effect of footrest

location and angle on seat and occupant response. Conditions for

these simulations are shown in Figure 101, and a comparison of

results is presented in Table 31.

Test Conditions

42 G Sled test

0 +130 A/C 42 ft/sec initial velocity

I 14.5-G E/As

C.) Part 572 dummy

4J

0 •Footrest friction

42 ft/sec coefficient =0.35

101.

Time

Figure 101. Test conditions for SOM-LA simulations

to determine the effect of footrest

angle and location.

TABLE 31. COMPARISON OF SOM-LA SIMULATIONS TO DETERMINE THE

EFFECT OF FOOTREST ANGLE AND LOCATION ON SEAT AND

OCCUPANT RESPONSE

Loc,• tion Peak Peak Peak Peak Peak

* Footrest Forward Seat Seat Pan Pelvic Chest Lumbar Lumbar

Angle of SRP Stroke z-Accel z-Accel z-Accel Load Moment

(Deg)in.) (in.)., .... (G) (in.-lb),

40 33 10.8 30.2 26.5 36.0 2350. 1474. ý_X

20 33 10.0 30.3 24.3 3647 2396* 1264.

0 33 11.5 30.3 26.6 33.8 2297. 1633.

40 26 9.2 30.6 26.6 35.4 2327. 2498.

0 26 9.8 30.4 25.3 35.8 2349. 2131.

0 19 7.6 30.1 26.1 29.9 2014. 2660.

149
The SOM-LA simulations indicate that the footrest plays an important role in determining how the loads are transmitted to the

body. For example, the predicted lumbar spine moment increased

by 110 percent when the feet were positioned 14 in. farther away

from the seat on a foctrest set at an angle of 20 degrees above

the plane of the floor. The maximum seat stroke varied by as

much as 3.9 in., which implies that the legs are bypassing the

energy-absorption mechanism and transmitting crash loads directly "to the body, thus reducing the required seat stroke.

"The objective of this series was not to infer that the feet can

be positioned to minimize load, although the great variation in

results indicates that this may be possible. Rather, the goal was

to show that there can be a significant effect on response due to

foot positioning. Future qualification testing specifications

must contain provisions for attaining repeatable foot constraints.

It is recommended that footrest angle and position correspond to

expected foot position at impact. Also, in the parametric test

series the dummy's feet were taped securely to the footrest, probably affecting system response. As mentioned in Section 7.1.4,

future qualification test specifications should address the question of whether to tape the feet.

6.2.3 Landing Gear Acceleration Ramp

Recent qualification test specifications of energy-absorbing seats - have not been consistent in reducing the energy-absorption capability of the aircraft landing gear in the input acceleration

pulse. The qualification test for the AH-64A Apache, which has

energy-absorbing landing gear that absorb a significant amount of

crash energy, did include a 5.35-G acceleration plateau for approximately 100 msec prior to the main impact pulse. However,

the test pulse for the U11-60A crewseat qualification test was the

nominal, triangular pulse described in NIL-S-58095(AV). Hearon,

et al., (Reference 15) have shown that a very short duration, low

level acceleration just prior to the main impact can have a very .significant protective effect by minimizing dynamic overshoot.

The simulations presented here are designed to show what effect

the landing gear acceleration ramp has on seat and occupant response.

Three input acceleration pulses are shown in Figure 102: the

AlI-64A crewseat qualification test pulse, the main segment of

this pulse without tho landing gear acceleration ramp and with

a correspondingly reduced velocity change, and a modified principal pulse that includes the entire 50 ft/sec velocity change of

the 95th-percentile accident. Simulations for the firsc two input pulses presented in Table 32 compare the condition~s when the

seat and occupant encounter the main impact pulse (corresponding

, to fuselage impact) at the same velocities with and without the

benefit of preload. In the second simulation, without the benefit

of preload, the seat requires an additional 0.8 in. of stroke dis-

"tance and the vertical accelerations and spinal load increase by

"at least 30 percent. Only the lumbar spine moment appears to be

150

4. 4 • '..• .' 4. , . . ... - . ' , .. . .. " • .**... . 10 AP. . _W W 41 . ,

'--4. -7
50 48.8 G

0

( 25 Av 50 ft/sec

U

U

5.35 G

0. 0.025 0.050 0.075 0.100 0.125 0.150

"50 48.8 G

50

m 25 33.8 ft/sec

(D

0 

1

' o 0 I I

0. 0.025 0.050 0.075 0.100

50 48.8 0 flTM /50

0

65v 50.0 ft/sec 25

U

4. 0

0. 0.025 0.050 0.075 0.100

•* •Time (sec)

Figure 102. Input acceleration pulses for simulations examining the effect of landing gear acceleration on seat and occupant response.

"less hazardous without preloading, possibly due to the duration

of the applied acceleration. The third simulation shows the ef- fect on seat and occupant response if the fuselage were required "_..4, to absorb the entire impact pulse. The results show an increase

",1 in seat stroke of 12 in. when the simulation is porformed without the effect of the landing gear. Body segment accelerations and spinal load are ver'y similar for the second and third cases, indicating that they are a result of the initial dynamic overshoot

151

• • , •• --,.~ .- , '.-, ,,,- -,.-=-------. - 4-- ..-'. •-...,:- -, , ., T 9. :.. . . .

,,.,,,. .\ T,, . -,, *-, •, , ,, - ,,, , ,. . ........ , .,.,.. .. . , .. [ .. .... '. .. . ., . . . .. ' -. .
TABLE 32. COMPARISON OF PREDICTED SEAT AND OCCUPANT RESPONSE

_ _FOR INPUT DECELERATION PULSES WITH AND WITHOUT

LANDING GEAR RAMP

Peak Peak Peak Peak Peak

Input Seat Seat Pan Pelvic Chest Lumbar Lumbar

Pulse Av Stroke z-Accel z-Accel z-Accel Load Moment

Shape (ft/sec) (in.) (G) (G) (G) (Ib) (in.-lb)

50.0 9.2 24.9 21.6 25.3 1742. 1801.

. 33.8 10.0 34.8 28.9 36.6 2441. 732.

50.0 21.2 34.8 28.6 33.2 2263. 2404.

due to lack of preload and not necessarily a function of crash

energy. However, the spinal moment does show a large increase

between the two cases.

Results presented in this section indicate that the preload due

to landing gear acceleration can have a significant effect on reducing injury potential. Future specifications must account for

this effect in a consistent manner for qualification testing.

6.2.4 Test Orientation to Gravity Vector

The facilities test variable discussed in Section 6.1.5 actually

"included the influence of two variables: the input pulse shape

and the orientation of the decelerator mechanism to the gravity

vector. It was known prior to conducting the parametric test

series that gravity can influence occupant response. Por a horizontal decelerator conducting a "vertical" impact test, gravity

acts to resist flexion of the upper torso. In this orientation,

"the gravity vector is equal to about 10 percent of the component

of the impact acceleration that it opposes. Making the same comparison S• for a vertical decelerator, gravity opposes the main inpkit

152

* Vow -- * V 9.9. . r--9'. .- ----- . .. . -, - .-- OI.. # 

1... *',°:[- t, *4(" " ",' -. . . * *, -" ",; ,'-'.' '" *"." " . " ..*""- "'o""""." ". ." ' . " " . . . , ,• ,; A, ,. '-. . .-. ,. . .9 . .i ., , . ,,q9 9. . . 9.. ... . .. . . . . . .. . .. . .. .

. .• . ', . ,' .1'. . '. , . , -. . • , . . . . - .9. .. 9 •. , .
acceleration pulse for which it is equivalent to only about 2.5

percent. An attempt was made to adjust the test conditions to

counteract the effect of gravity by adding additional seat pitch for the horizontal decelerator test conditions. The two simulations presented in this section were conducted to determine

whether the vertical and modified horizontal decelerator test

conditions were equivalent.

Test conditions for each of the simulations, shown in Figure 103,

were identical with the exception of seat orientation to gravity.

For the vertical test condition, seat pitch was +13 degrees, which

oriented the seat back plane at zero degrees to the gravity vector. The horizontal condition had a seat pitch of +17 degrees

resulting in a 73-degree angle between the seat z-axis and the

gravity vector. Results of the simulations for these two test

conditions are shown in Table 33.

42 G

0• Test Conditions

42 ft/sec initial velocity

14.5-G E/As

U

, mPart 572 dummy

'U,

42 ft/sec\

Time

Figure 103. Conditions for aircraft orientation simulation.

''•1

TABLE 33. COMPARISON OF SEAT AND OCCUPANT RESONSE FOR VERTICAL

AND HORIZONTAL TEST CONDITIuN SIMULATIONS

Peak

Seat Peak Peak Peak Peak

Decelerator Seat Seat Pan Pelvic Chest Lumbar Lumbar

Test Pitch Stroke z-Accel z-Accel z-Accel Load Moment

Condition (deg) (in.) (G) (G) . (GJ (ib) (in.-lb)

Horizontal 17 9.8 30.2 26.5 36.0 2350. 1474.

Vertical 13 11.3 29.7 26.8 34.6 2332. 1937.

153

S *... S 9 

.. •.•.* . .r -.. .. .. '..',,., * 4.. •-. . . : ** "....... ," . .
"There does not appear to be a significant difference in the z-axis

accelerations or axial spinal load for the two test conditions.

However, SOM-LA predicts a difference in seat stroke of 1.5 in.

and a 31 percent increase in lumbar moment for tha vertical decelerator test condition. It appears that the additional seat

pitch added to the horizontal test condition was not sufficient

to produce similar flexion in the upper torso as measured by the

lumbar moment. If additional seat pitch was used to overcome

*' gravity, then the disparity in seat stroke would be even greater.

Obviously, the vertical and horizontal test conditions used in

the simulations did not result in predictions of identical response. Additional work needs to be conducted with simulations

"to identify comparable test conditions, followed by dynamic tests

to verify equivalency. Also, because SOM-LA applies a 1-G force

to initially seat the occupant in the cushions, it does not truly

"simulate the free-fall conditions of a vertical drop test. This

factor would have to be taken into account in future simulations,

perhaps modifying the computer program for simulating a drop test.

6.2.5 Tolerances on Input Deceleration Pulse

Allowed by MIL-S-58095(AV)

Dynamic test requirements for Test No. 1 of MIL-S-58095(AV), shown

in Figure 104, allow the rise time of the impact acceleration

pulse to vary between 0.4 and 0.8 of the total pulse duration.

Program SOM-LA was used to conduct a parametric matrix of simula-

. tions for the two limits on the rate of onset with 50th- and 95thpercentile occupants and 11.5- and 14.5-G energy absorbers (based

on a 50th-percentile occupant). The nominal and actual input

pulses used for the simulations are shown in Figures 105 and 106,

respectively. The eight cases comprising the parametric series

are described in Table 34, with results listed in Tables 35 and

36 for the 14.5- and 11.5-G series, respectively.

In each of the four series comparing the minimum and maximum rate

of onset cases, the minimum rate of onset case was always less

severe, There is a minimum of 1-in. difference in seat stroke,

and seat pan and chest z-axis accelerations are greatly reduced.

Maximum lumbar spine load is also decreased significantly for the

minimum rate of onset cases, while the lumbar moment is not affected to a large degree.

7". .' Results of the rate of onset test series, described in Section

6.1.3, did not demonstrate this significant effect of rate of onset on system response, and when those tests were simulated with

SOM-LA, the accelerations did not exhibit any significant variation, thus agreeing with the test data. Seat stroke, on the other

*• hand, was 1.5 in. greater for the simulation of the higher-rateof-onset test (A80-058) than for the baseline tests (A80-053 and

A80-054). The analysis of pulse tolerances described here are

not intended to be used quantitatively but only to determine possible significance of variables that should be investigated in

greater detail by further testing.

"154
LOAD z -

•1 !- DUMMYLOAINERTIA/

3040 100

"bX Y

CHANGEIN

VELOCITY= 50 FPS

Nominal test pulse

"-.1065 SEC tThe rise time for the K G

nominal pulse may vary Maximum rate of

between the two onset pulse

values illustrated 0.4 T

(= time)

PEAK G Minimum rate of

onset pulse

Figure 104. Dynamic test requirements for Test No. 1

of N4L-S-58095(AV).

6.2.6 Method of Transducer Calibration

At all test facilities the input (sled or drop cage) acceleration

was measured using two accelerometers in order to ensure acquisition, of that important parameter. The same NADC-owned strain-gagetype accelerometer was used as the primary transducer in all tests

•,•I 155

•'e:.•'lle~•,• w"'••p' •' •"• •.* I5 " " •"IIS.4.* ,...... - "q'.... ...... "I5..... ' •' ... ... "'

- ,' ,.-- "...- . ,-5,,.. . ,....,...'.........,• . . '''i~ i-i
60

50 •:•~~ •'" 8 G

40

'~30 I LV = 0 ft/sec

U) ~20I I

10

{ 0.0325 sec

0

0 .025 .050 .075 .100

Time (sec)

Figure 105. Nominal acceleration pulse for

Test No. 1., MIL-S-58095(AV).

60-

50 - 0.026 sec 0.052 sec

-0- 48 G

40C)

?.4

10

0;0 S0 *

0 .025 .050 .075 .100

4*

Time (sec)

Figure 106. Minimum and maximum rate of onset conditions for

simulations of Test No. 1, MIL-S-58095(AV).

156

-4,0
A TABLE 34. MATRIX OF SOM-LA SIMULATIONS FOR

EVALUATION OF SENSITIVITY TO TEST

PULSE TOLERANCES AT 11.5- AND

14.5-G ENERGY ABSORBER LIMIT LOAD

SETTINGS

Minimum Maximum

Rate of Onset Occupant Rate of Onset

920 C/sec Percentile 1850 C/sec

11.5 G 14.5 G 50 11.5 G 14.5 G

11.5 G 14.5 G 95 11.5 G 14.5. G

TABLE 35. COMPARISON OF PREDICTED SEAT AND OCCUPANT RESPONSE TO

MIL-S-58095(AV) DYNAMIC TEST 1 WITH 14.5-G E/A LIMIT

LOAD

Peak Peak Peak Peak Peak

Occupant Seat Seat Pan Pelvic Chest Lumbar Lumbar

Input Percen- Stroke z-Accel z-Accel z-Accel Load Moment ., Pulse tile (in.) (G) (G) (G) (ib) (in.-lb)

Max RO 50 12.5 33.2 27.4 39.0 2770. 4153.

Min RO 50 11.4 22.0 27.2 30.9 2350. 4424.

Max RO 95 14.9 32.3 26.2 31.5 3010. 6100.

Min RO 95 13.8 20.8 25.0 24.5 2460. 6160.

TABLE 36. COMPARISON OF PREDICTED SEAT AND OCCUPANT RESPONSE TO - MIL-S-58095(AV) DYNAMIC TEST 1 WITH 11.5-G E/A LIMIT 2 •LOAD

Peak Peak Peak Peak Peak

Occupant Seat Seat Pan Pelvic Chest Lumbar Lumbar

Input Percen- Stroke z-Accel z-Accel z-Accel Load Moment

Pulse tile _(in..) (G (G) (G) (lb) (in.-lb)

Max RO 50 18.6 28.3 23.4 30.6 2363. 4170.

Min RO 50 17.0 21.8 24.3 25.0 1970. 4130.

Max RO 95 20.7 27.1 22.6 25.4 2560. 5800.

Min RO 95 19.3 20.9 22.4 19.3 2230. 5910.

,.4. 157

.: .1 *4

*. *. % : *:, ."41 . . .... : . *I . .... .. .. * . ,q ..... qP . ' " .. - , ... .• ... O . " , .• . -, . ... -
at all facilities, and the redundant transducer was furnished by

the individual facility performing the test. At Simula, where a

different method of calibration was used for accelerometers, a

significant difference was observed between primary and redundant

accelerometer measurements. For example, in Simula Test SEAC-1,

the primary and redundant accelerometers (both Bell and Howell

4-202-0001) produced peak values of 35.4 and 40.1 G, respectively.

The pulses were of identical shape, but the difference in magnitudes produced a difference in computed velocity change, 38.8

versus 42.5 ft/sec. (The actual impact velocity was 31.1 ft/sec, -- and the remainder of the velocity change resulted from the rebound,

which was significant).

The primary accelerometer was calibrated statically at NADC, using

a centrifuge, whereas the redundant Simula transducer was dynamically calibrated in a frequency regime comparable to that expected in the testing: 10-2000 Hz. Although CAMI and NADC both

employ centrifuge calibrations, they use different techniques to

compensate for lead resistance. In an attempt to understand the

differences in results, prior to their dummy verificiation tests

(A81-004 through -006), CAMI recalibrated the NADC and Simula accelerometers using their centrifuge. Using the calibration resistance (R c) value from NADC, their accelerometer produced an output

that was 3.75 G low at 100 G. Simula's accelerometer, using their

"R value, was 5.0 G high at 100 G. When these two transducers

were mounted side by side with a CAMI accelerometer for tests

A81-004 through -006, using the CAMI calibration, all three produced identical results.

It should also be noted that each facility used signal conditioning equipment with different characteristics. This means that the

control accelerometer was placed into a resistive balancing network which varied at each test facility. Full bridge-type transducers are susceptible to balance circuit errors because shunttype balancing is essential. This follows because the bridge circuit is usually enclosed within the transducer. Errors from this

requirement arise when a transducer is calibrated in a system which uses a given value of balancing resistor, but is then operated in the test with a signal conditioner having completely

different balancing circuit characteristics. This error is normally eliminated by the system (end-to-end) calibration method

practiced by the test facility.

The static/centrifuge method can be performed more readily in a

number of laboratories, while the dynamic/shaker method requires

more complex apparatus. However, the dynamic method considers

the response characteristics of the transducer in an environment

which may more closely resemble that experienced in the test.

The results here demonstrate the need for specifying calibration

methods in revised test specifications.

*4

158

Vitt
7.0 DISCUSSION OF RESULTS

This chapter presents conclusions drawn from results of the para-

"metric test series and sensitivity analysis as applied to the

*u overall goal of developing more efficient energy-absorbing seating systems. The U.S. Army's four-part research effort will ultimately lead to an improved, comprehensive specification for these

systems that will ensure procurement of optimal designs.

7.1 INPUT ACCELERATION PULSE

The shape of the acceleration pulse used as inpit for crash testing has a significant influence on seat and occupant response.

Each facility used for testing will have a characteristic pulse

shape dependent on the type of decelerator mechanism used, as illustrated by the results of the test facility series described in

Section 6.1.5. Any of the test pulses may resemble the acceleration environment that might be experienced in an aircraft accident. However, some facilities are able to provide a much better

* approximation of the nominal triangular pulse shape.

It was found that seat and occupant response to a specific input

acceleration pulse was very sensitive to the peak acceleration

and velocity change. The rate of onset of the acceleration pulse

"appears to have a secondary influence on response compared with

the above two variables. However, computer simulations presented

in Section 6.2.5 show that the tolerances allowed for rate of onset of the input pulse for Dynamic Test No. 1 of MIL-S-58095(AV)

are large enough to allow significant variations in response.

Orientation of the test facility decelerator mechanism to the

gravity vector was also shown to have a marked influence on stroke

for the seat orientations used in this program.

Simulations with Program SOM-LA indicated that the low-level acceleration provided by stroking landing gear can create a protective effect by causing a preload in the seat and body that minimizes dynamic overshoot.

7.2 PERFORMANCE CRITERION

The current criterion in MIL-S-58095(AV) for evaluation of energyabsorbing seat performance is based on acceleration measured in

the vertical direction on the seat pan. Knowledge concerning the

dynamics of impact and the interaction between the seat and occupant, and the typical injury-causing mechanisms has been greatly

expanded since this specification was written. For example, it

is now known that maximum seat pan acceleration does not occur

concurrently with maximum vertical loads on the body; and that

the typical lower spine injuries, anterior wedge fractures, are

as much a function of torso flexion as axial loading along the

spine. Vertical seat pan acceleration is unrelated to bending

that occurs in the spine, and it is also very sensitive to localized seat pan properties that do not influence injury potential

as was shown in the movable seat weight series. '4

159

.4'. 'W. -9.9 1W %W W -A W-
Another parameter investigated for correlation with injury poten- tial, the Dynamic Response Index, does have the advantage of accounting for the cumulative, time-dependent effects of accelera- tion. The DRI, therefore, can provide a measure of test severity

that considers the effects of multiple acceleration peaks on occupant response. However, it suffers from some of the same disadvantages as the direct use of the maximum seat pan acceleration because it must be calculated using measured seat pan acceleration

as input to a simple dynamic model; and it is also unrelated to

moment generated in the spine due to bending during the signifi- cant torso rotations likely to occur in helicopter accidents.

The only parameters investigated that provided predictable trends

in the injury-causing mechanisms were actual measurements of lum- bar spine load and moment. The feasibility of this technique was

demonstrated in this study by modifying two Alderson anthropomor- ' _. phic dummies, a 50th-percentile Hybrid II and a 95th-percentile

VIP-95, and repeating several key tests from the parametric test matrix. Although results from the modified dummy test series indicated some differences in performance with the load cell in-

*. stallation, it was apparent that the concept was sound. The occupant model in Program SOM-LA was correlated to the modified

dummy response characteristics and was able to show important trends in the spinal injury-causing mechanism for all of the

series in the parametric test matrix.

7.3 SEAT DESIGN PARAMETERS

The energy absorber limit load plays a very important role in determining the efficiency of a seat design. For any given stroke

distance and maximum force, a constant-force (square-wave) type device can absorb the greatest amount of energy. The design prob- lem becomes one of selecting a limit load as high as possible while not exerting injurious forces on the occupant. The actual

limit load setting that provides optimal protection from spinal injury was not examined in the parametric study; however, results

of the related cadaver and human volunteer test series combined

with accident injury data should supply this information when

completed. The energy absorber limit-load series conducted in

the parametric program determined the sensitivity of seat per- formance to limit-load setting. It was determined that the relationship between seat stroke and limit load is not linear;

* i.e., required stroke distance increases more rapidly than the

reduction in limit load would indicate.

In terms of occupant response, the spinal load decreases propor- tionally with a reduction in energy absorber limit load. Unex- - . pectedly, however, predicted spinal y-axis moment increases with

decreasing limit load. It appears that the moment is being influenced by the magnitude of seat stroke, which, due to the kine-

":.A matics of the constrained legs, causes rotation of the lower torso.

This trend may not be realized if the feet were not tied to the

footrest as the test procedures in this program specified.

S.4 160

-. "--p .''•- -3" - T

• * ',* • " -- . . . .. ' , - , " - 'A. + ' " " , - ' " , ' • . - ... ' J '- " > .+ ." ", . -' -,' •,- -'", -. . " '
Although it was once believed that ramped limit-load energy absorbers would provide a less injurious environment compared with

constant load devices, results showed that they produced a more

severe occupant loading. Also, the ramped E/Als that were tested

are less efficient, thereby requiring a greater stroke distance

and increasing the hazard of "bottoming out."

It was determined that movable seat weight was not a factor influencing the potential for injury as measured by spinal loads

and moments. However, vertical seat pan acceleration was a function of seat weight, possibly due to dynamic overshoot of the

structure that occurs when the occupant motion is not strongly

coupled to the seat pan. This series reinforced the conclusion

that seat pan acceleration is not a good indicator of test severity. The effect of seat bucket stiffness on occupant response

was not examined and may be a significant influence on seat and

occupant interaction.

7.4 TESTING PROCEDURES

"-- Procedures were developed in the parametric test program to prepare the seat and occupant for testing. These procedures were

very important in obtaining repeatable test conditions, particularly pretest measurement of body segment targets and the small,

repeatable initial tension given to the restraint system. It is

recommended that these or similar procedures be adopted as a

standard for qualification testing. However, one procedure that

will require further analysis is placement of the feet. In the

"parametric test series it appeared that taping the feet to the

footrest was an unrealistic constraint and possibly altered the

kinematics of the body during the crash sequence. Simulations

with Program SOM-LA indicated that footrest location and angle

can influence seat and occupant response to the impact by altering the load path between the seat and the floor.

Response of various types of anthropomorphic dummies currently used for qualification testing was found to be dissimilar enough

to prohibit interchangeable use ir testing. The one standardized

test device# the Hybrid II, conforming to the Part 572 specification, was found to provide repeatable test results without sustaining significant degradation after use in 36 dynamic tests.

This model dummy was also found to be capable of being modified

to allow measurement of spinal loads and moments with the addition

of a six-axis load cell at the base of the spine. The elastomeric

spine in the Hybrid II was designed to approximate human response Oil• in flexion. It would be desirable to develop a spine optimized

to replicate combined axial and flexural response with an integral

load and moment measuring capability.

The discussion presented on instrumentation calibration in Section 6.2.6 is critically important in terms of qualification testing. When the instrumentation is responsible for providing data

used to assess compliance with the procurement specification, it

is important to ensure its repeatability and its calibration for

161

Y . . " .- " * . ,
the regime in which it will be required to operate. A further

examination of calibration methods may be required. Data filtering is also an important area requiring additional analysis. Current Army specifications do not require a standardized filtering

"procedure, such as SAE J211 (Reference 25), although this recommended procedure is followed by many laboratories. However, it

is not 'lear whether the tolerances allowed in this specification

are too 'reat; also, it has been suggested by participants in this

program that SAE J211 does not make use of the performance abilities of digital filters.

162&

lip.

.4..'-7

* 4

*. 54

2. 5. Society of Auitomotive Engineers, Instrumentation for Impact

'". TeSts, SAE Recommended Practice SAE J211"JUN80, Transactions,

•-• SAE Handbook, Society of Automotive Engineers, Inc., Warreni date, Pennsylvania, 1981.

" ~162

; - - - V- - '.. .- .---

-' 4". -"
8.0 CONCLUSIONS

"Based on the results of this program, the following conclusions

can be drawn:

1. The shape of the input acceleration pulse has a significant effect on seat and occupant response.

2. Different test facilities will produce different results

in terms of seat and occupant response. Differences can

be attributed to test orientation, i.e., horizontal or

vertical, as well as to the characteristic pulse shape.

"3. Measurement of spinal force and moment provides the

most reliable means of relating test performance to

spinal injury. Seat pan acceleration, on the other

hand, is not a good indicator of test severity or injury potential.

4. Stroke distance and energy-absorber limit load are not

•' linearly related; i.e., required stroke distance increases more rapidly than a reduction in limit load

would indicate.

5. Increasing ramp energy absorbers are less efficient

than constant load (square-wave) devices and potentially

more hazardous to the occupant.

6. Simulations of crash tests indicate that placement of

the dummy feet can significantly influence seat and occupant response in a dynamic test.

7. The Part 572 dummy provides repeatable results, and

extensive testing produced no measurable degradation

-0'1 of performance. However, it is not certain that the

elastomeric spine of this dummy provides a response

similar to that of a human spine.

8. Simulations indicate that low-level acceleration provided by the landing gear portion of the crash pulse

can reduce the magnitude of dynamic overshoot, hence

lowering tho peak accelerations and loads in the body.

C.. 163

V~l. *

. 4 C, A S * - -. . . ' Y .. . 4 , ., 5 *•' .. . ., V # *' * .. .• . .. - ei 4 . -1 * . .

",, S- .,__ • . -% . , . • . -
9.0 RECOMMENDATIONS

;" 9.1 FOLLOW-ON RESEARCH

The results of the parametric test program and sensitivity analy- -

sis presented in this report constitute a tremendous step in attaining the goals of the Army's research efforts in energyabsorbing seating. However, there are several important, unanswered questions that must be examined. Those areas requiring

additional work are discussed below.

9.1.1 Energy Absorber Limit-Load Threshold

"The threshold, or upper limit, of energy absorber load that pre-

"", vents a significant portion of spinal fracture injuries is being

investigated in the cadaver and human volunteer testing programs

"discussed in the Introduction. However, it is equally important

to maintain an on-going effort to monitor injury data in accidents

* involving current, production energy-absorbing seats. The accident conditions and aviator population involved can never be duplicated in laboratory conditions. It may be desirable to use a

statistical approach to minimize the total number or the severity

of injuries by combining the statistical distributions of spinal

strengths and accident severities with the sensitivity curves presented here to set an optimal energy absorber load level.

"9.1.2 Standardized Test Dummy

A standard test dummy for energy-absorbing seat testing needs to

be developed. Results of the study to determine anthropometry

for the military aviator population can be incorporated with

methods developed in the instrumented dummy research effort to

produce a standardized device. If funding constraints should prohibit development of a new dummy, then an economical alternative

* : would be to modify the spinal properties of the Part 572 design

to meet vertical testing requirements. The spine must be instrumented to provide a measure of forces and moments. The optimum

location for force and moment measurement is not known, and should

be determined if a new lumbar spine is designed. In the existing

Part 572 system, the base of the spine is the only practical loca-

"tion for a transducer. Also, it may be necessary to scale the

design of the 50th-percentile dummy to develop 95th-and 5th-

*i percentile versions for testing the occupant size range for variable load energy absorber systems now under development.

The ultimate goal of the standardized dummy development will be

"2 to provide a means for quantitative assessment of injury potential rather than just a comparison of relative test severity.

Achieving this end will require conduclIng a matrix of dynamic

tests with the standardized dummy, replicating cadaver and human

volunteer experiments, as well as performing a number of the sensitivity tests conducted in this pl.ogram. The purpose would be

164

-, •

4. - -.- --.. ,., -. • -,.•- • * -•/ -' , .... . 1 .. . .
to develop actual spinal load and moment measurements that represent the spinal injury threshold for the aircraft crash environment. An adjunct to this effort would be to correlate Program

SOM-LA to the new dummy response characteristics to minimize the

number of dynamic tests required for future parametric studies.

* 9.1.3 Refinement of Qualification Test Methods .4

An analysis of the tolerances a'lowable on the input acceleration

"pulse needs to be conducted. This would require the development

'4• of a set of reasonable seat/occupant response variations and then

the use of the sensitivity analysis presented in this report to

define the allowable deviations from the desired test pulse for

parameters such as peak acceleration, velocity change, and rate

of onset. It would then be desirable to conduct dynamic tests

with the new standardized test durmmy to verify predicted performN ance variations and explore the range of spinal loads and moments

that these variations produce.

An effort should be undertaken to determine if an equivalent test

condition can be achieved for vertical iad horizontal decelerator

test facilities. The best approach for accomplishing this would -- '• be the use of SOM-LA simulations, followed by a set of dynamic

tests to verify the results. If equivalent test conditions cannot

be produced, then it must be decided whether the resulting performance variations exceed the tolerable limits thus requiring one

decelerator mechanism type or orientation.

The final area requiring additional research in order to develop

an improved specification is calibration of transducers, along

with data processing. A comparison needs to be made of transducers to be used for determination of performance compliance and

their available calibration techniques. The type of transducers

to be examined would include accelerometers for input acceleration

and load cells for measuring spinal load and moments. Procedures

for signal conditioning and filtering must Le explored to develop

a set consistent with the magnitude of allowable errors.

9.2 SUGGESTED MODIFICATIONS TO CURRENT ENERGY-ABSORBING

SEAT SPECIFICATIONS

The following areas have been identified as requiring modification

in MIL-S-58095(AV), MIL-STD-1290(AV), and possibly the Aircraft

Crash Survival Design Guide, TR-79-22, to achieve the Army's goals . f improved energy-absorbing seating systems:

* Define closer tolerances on the input deceleration pulse

and provide definitions for each prescribed variable.

I Specify equivalent test conditions for horizontal

and vertical test facilities and types of decelerator

mechanisms.

165

A.1"' W- W

4.%
, Require the use of the standardized test dummy with

applicable instrumentation.

• Possibly modify the required test matrix to include

"tests at the limits of occupant size range, i.e., 95th

and 5th percentile, possibly including female occupants.

* Specify required energy absorber limit-load factor.

" Add performance assessment criteria based on maximum

allowable spinal load and moment values for the standardized test dummy.

. Provide a standard procedure for test dummy positioning

and seat preparation prior to qualification testing.

* In addition to testing of the seat itself, require an

evaluation of equipment and structural interfaces applicable to specific aircraft that will affect seat performance. Floor structure and attachment should be as

close as possible to actual aircraft structure for the

seat qualification test, and procedures for providing

floor deformation should be specifically defined.

Then such equipment as helmet, instrumentation consoles,

and sights should be included in an additional test to

evaluate crew station safety performance.

* Footrest location for testing shnuld be required in

system specification, but default position can be

included in more general seat specifications.

. Standardize calibration procedures for transducers used

for assessing compliance with specification requirements.

9 Specify data filtering procedures and filter character-

*. istics.

-7 .. -'•

* ..

'.".' . •166

•-•• S....• ." 'V "* ....- 1* t .... * " ~ 'U " • " " "V "#' "'W "W'
"I..' 10.0 REFERENCES

"1. Singley, G. T., III, and Desjardins, S. P., "Crashworthy

Helicopter Seats and Occupant Restraint Systems," in Operationel Helicopter Aviation Medicine, AGARD Conference Proceedings No. 255, North Atlantic Treaty Organization, Advisory Group for Aerospace Research and Development, Neuilly

-, sur Seine, France, May 1978.

2. Crash Survival Design Guide, Dynamic Science, A Division of

-. Marshall Industries, USAAMRDL Technical Report 71-22, Eustis

"Directorate, U.S. Army Air Mobility Research and Development

Laboratory, Fort Eustis, Virginia, 1971, AD 733358.

3. Military Standards, MIL-STD-1290(AV), Light Fixed - Rotarywing Aircraft Crashworthiness, Department of Defense, Washington, D.C., 25 January 1974.

4. Military Specification, MIL-S-58095(AV), Seat System: Crashworthy, Non-Ejection, Aircrew, General Specification For,

Department of Defense, Washington, D.C., 27 August 1971.

41 5. Desjardins, S. P., and Harrison, H., The Design, Fabrication,

and Testing of an Integrally Armored Crashworthy Crewseat,

Dynamic Science, Division of Marshall Industries; USAAMRDL

"Technical Report 71-54, Eustis Directorate, U.S. Army Air

Mobility Research and Development Laboratory, Fort Eustis,

Virginia, January 1972, AD 7421733.

"6. Desjardins, S. P., and Laananen, D. H., Aircraft Crash Sur- ..,vival Design Guide, Volume IV - Aircraft Seats, Restraints, -Lters and Padding, Simula Inc., USARTL-TR-79-22D, Applied Technology Laboratory, U.S. Army Research and Technology

Laboratories (AVRADCOM), Fort Eustis, Virginia, June 1980,

AD A088441.

7. Laananen, D. H., Aircraft Crash Survival Design Guide, Volume II - Aircraft sh Environment and Human Tolerance, Nimla S'i •Inc., usARTL-TR-79-'22B,, AppilTed T;ecý;oiogy"La o-ratory,

"U.S. Army Research and Technology Laboratories (AVRADCOM),

-Port Eustis, Virginia, January 1980, AD A082512.

8v Eiband, A. M., Human Tolerance to Rapidly Applied Aceelerations: A Summary of the Literature, NASA Memorandum 5-19-9E

' NatiMonal Aeronautics and Space Administration, Washington,

D.C., June 1959.

9. Brinkley, J. W., Weis, E. B., Clarke, N. P., and Temple,

W.E., A Study of the Effect of Five Orientations of the

Acceleration Vector on Human Response, AMRL Memo M-28,

Aerospace Medical Research Laboratory, Wright-Patterson Air Force Base, Ohio, 1963.

167

44,

• • . _ , o .* .* . . . •%••."T-T÷- •-"." . . ... T"":'- '"" " ' .. .. ' • -". .... " ••
10. Reader, D. C., The Restraint Afforded by the USAF and Proposed RAF IAM Seat Harnesses for the F-Ill Under High Forward

and Lateral Decelerations, IAM Report 421, Institute of Aviation Medicine, Farnborough, Hampshire, United Kingdom, 1967.

11. Crash Survival Design Guide, Dynamic Science, A Division of

Marshall Industries, USAAMRDL Technical Report 70-22, Eustis

Directorate, U.S. Army Air Mobility Research and Development

Laboratory, Fort Eustis, Virginia, 1970, AD 695648.

12. Domzalski, L., and Singley, G. T., III, Joint Army/Navy Test

Program for Black Hawk Seat Systems, Report NADC-79229-60 * *~ (Draft), Naval Air Development Center, Warminster, Pennsylvania, 1979.

13. Dummer, R. J., Qualification Test Report, 613-1787 COOL

Qualification Testing of Armored Crashworthy Aircrew Seat,

RA-305252-1, for Sikorsky Aircraft Contract 576344, Norton

Company, Worcester, Massachusetts, October 1978.

14. Chandler, R. F., Dynamic Test of Joint Army/Navy Crashworthy Armored Crewseat, Protection and Survival Laboratory, Memoran-

* U dum Report AAC-119-80-2, Civil Aeromedical Institute, Mike

Monroney Aeronautical Center, Federal Aviation Administration, Oklahoma City, Oklahoma, 25 April 1980.

15. Hearon, B. F., Raddin, J. H., and Brinkley, J. W., *Guidance

"for the Utilization of Dynamic Preload in Impact Injury Prevention," in Impact Injury: Mechanisms, Prevention and Cost,

North Atlantic Treaty Organization, Advisory Group of Aerospace Research and Development, Neuilly sur Seine, France,

* April 1982.

16. Applied Technology Laboratory, U.S. Army Research and

Technology Laboratory (AVRADCOM), Contract DAAK51-C-0016,

"Provide Hardware, Test Support and Engineering Services*

Contracted with Simula Inc., Tempe, AZ, May 1979.

17. Laananen, D. H., and Coltman, J. W., Measurement of Spinal Loads in Two Modified Anthropomorphic Dummies, Simula Inc.

TA-82465, final Rport, Contract DAMD17-81-C-1175, U.S. Army Aeromedical Research Laboratory, Fort Rucker, Alabama,

May 1982.

"18. Desjardins, S. P., et al., Crashworthy Armored Crewseat for

the UH-60A Black Hawk, paper presented at 35th Annual National Forum of the American Helicopter Socity, Washington, D.C., May 1979.

19. Code of Federal Regulations, "Anthrompomorphic Test Dummy,, Title 49, Chapter 5, Part 572, Federal Register, Vol. 38,

No. 62, April 2, 1973, pp. 8455-8458.

S. +

168

"*4,'. + a' . *• * *+ *I .

I -iI * *• ++ - . .. *.iI + + + + l , + m++ * + 4+b + *++ *,. ' + ,- ' ++ + " + - , + " +
20. Military Specification, MIL-S-9479, Seat System Upward

"Ejection, Aircraft, General Specification for, Department

"of Defense, Washington, D.C.

21. Stech, E. L., and Payne, P. R., Dynamic Models of the Human

Body, Frost Engineering Development Corp. AMRL Technical

Report 66-157, Aerospace Medical Research Laboratory, Wright- Patterson Air Force Base, Ohio, November 1969, AD 701383.

22. Brinkley, J. W., and Shaffer, J. T., Dynamic Simulation Techniques for the Design of Escape Systems: Current Applications and Future Air Force Requirements, AMRL Technical

Report 71-29-2, Aerospace Medical Research Laboratory;

Wright-Patterson Air Force Base, Ohio, December 1971,

AD 740439.

"23. Laananen, D. H., Bolukbasi, A. 0., and Coltman, J. W.,

Computer Simulation of an Aircraft Seat and Occupant in a

Crash Environment, Volume I - Technical Report, TR-82401,

Simula Inc., DOT/FAA/CT-82/33-I, Federal Aviation Administration Technical Center, Atlantic City Airport, New Jersey,

September 1982.

24. Laananen, D. H., Coltman, J. W., and Bolukbasi, A. 0., Computer Simulation of an Aircraft Seat and Occupant in a Crash Environment , Volume II - Program SOM-LA User Manual, TR-81415,

Si Simula Inc., DOT/FAA/CT-82/33-I, Federal Aviation Administration Technical Center, Atlantic City Airport, New Jersey,

September 1982.

25. Society of Automotive Engineers, Instrumentation for Impact

Tests, SAE Recommended Practice SAE J211 JUN80, Transactions,

SAE Handbook, Society of Automotive Engineers, Inc., Warrendale, Pennsylvania, 1981.

169

0 1. A- .0 _W 'W 
APPENDIX A

CAMI TEST A81-121

Rigid Seat

50th-Percentile Dummy * .(modified for spinal load measurement)

AV = 42.4 ft/sec

G peak = 14.1

44n

W"°

4.4

. 6w4 

0

A >

0 00

.170

iv U wl

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:, %••,•., -, ,",, •i•,_'*•'• ,', J ,' , , -, , ,, . . * .. .. . A . ..-. • A .

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u .z

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94

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APPENDIX B

CAMI TEST A81-122

Rigid Seat

95th-Percentile Dummy (modified for spinal load measurement)

Av = 42.5 ft/sec

G = 12.

Gpeak 13.

nw

S.,9 JIA

4. * -

M U

*44J

-*1.1

.qA

-f, ,n

.4 4

°p.-.,

179

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APPENDIX C

CAMI BASELINE TEST A80-053

Energy-Absorbing Seat

50th-Percentile Dummy

•v = 44.8 ft/sec

G peak 40.7

I-U)

I/ w a) I

tn

zM

tn U 0 a -0 ,

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+ • ,% < , + + • + , , *,- , , . ,, + . V + ', . . . . . . . . I•.. n , , - + + + + + +' ' + : + • ' IA " " " " + .. a 
APPENDIX D

NADC BASELINE TEST N-189

Energy-Absorbing Seat

50th-Percentile Dummy

AV= 40.4 ft/sec

Gea 44.9

pe

wi i

430

1.4

z 0.040

.44

.. 4.4ILn

9 d* 9 a t

in 49 P) MJ ~ - 1 t

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d w w mt

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SIMULA BASELINE TEST SEAC-1

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WAYNE STATE UNIVERSITY TEST 159

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