lJSARTL-TR- 79-22B

AIRCRAFT CRASH SURVIVAL DESIGN GUIDE

VOLUME I1 AIRCRAFT CRASH ENVIRONMENT

AND HUMAN TOLERANCE

IMUADAO 82512 $1MULA IC

2223 SOUTH 48TH STREET

TEMPE, ARIZONA 85282

4-/ 6iN DTIC

ELECTE

JANUARY 1980 AR118

FINAL iii REPORT AA

-45 i~l.II• , Time.. sec

4 0 in in •, • -.

PREPARED FOR

APPLIED TECHNOLOGY LABORATORY

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

FORT EUSTIS, VIRGINIA 23604

4-4

A~pproved for public release; distribution unlimited.I

SO 3 26 052

................... S ; .97 .2 2. 079 2..1 ..... 2.22
APPLIED TECHNOLOGY LABORATORY POSITION STATEMENT C>

This revised edition of the Crash Survival Design Guide was

prepared for the Applied Technology Laboratory by Simula, Tnc.,

under the terms of Contract DAAJ02-77-C-0021. The original

Crash Survival Design Guide was published in 1967 as USAAVLABS

Technical Report 67-22 and subsequent revisions were published

as USAAVLABS Technical Report 70-22 and USAAMRDL Technical Report 71-22. This current edition consists of a consolidation

of design criteria, concepts, and anlytical techniques developed through research programs sponsored by thlb Laboratory

over the past 20 years into one report suitable for use as a

designer's guide by aircraft design engineers and other interested personnel.

This document has been coordinated with USAAVRADCOM, the U. S.

Army Safety Center, the U. S. Army Aeromedical Research Laboratory, and several other Government agencies active in aircraft

crashworthiness research and development.

The technical monitors for this program were Messrs. G. T.

Singley III, R. E. Bywaters, W. J. Nolan, and H. W. Holland of

the Safety and Survivability Technical Area, Aeronautical Systems Division, Applied Technology Laboratory.

Comments or suggestions pertaining to this Design Guide will C

be welcomed by this Laboratory.

DISCLAIMERS

The findings in this report ere not to be construed as on officiel Department of the Army Position unieo so

designated by other authorized documents.

When Government drawings, specifications, or other data are used for any purpose other than in connection

with a definitely related Government procurement operation, the United States Government thereby Incurs no

responsibility nor any obligation whatsoever; and the fact that the Government may have formulated, furnished, or in any way supplied the seid drawings, specifications, or other date is not to be regarded by Implication or otherwise as in any manner licensing the holder or any other person or corporation, or conveying any rights or permission, to manufacture, uN, or sell any patented Invention that mov in any way be related thereto.

Trade names cited in this report do not constitute an official endorsement or approval of the usa of such

commercial hardware or software.

DISPOSITION INSTRUCTIONS

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

Af NOTE: This is a revised edition of the Crash Survival Design Guide (formerly USAAMRDL Technical Report 71.-22). All

previous editions are obsolete and should be destroyed.

i-ii ii
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20. (Continued)

").Volume I Design Criteria ad Checklists'

Volume i),,YAircraft Crash Environment ak Human Tolerance)

Volume III rAircraft Structural Crashworthiness ' Volume Iv;Aircraft Seats, Restraints, and Litters

Volume Aircraft Postcrash Survival, 2

"•his volume (Volume II) contains information on the aircraft crash

environment, human tolerance to impact, occupant motion during a

crash, human anthropometry, and crash test dummies, all of which

serves as background for the design information presented in the

other volumes.

.I /

V.

L

UNCLASSIFIED SECURITY CLASSIFICATION tf THIS PAOEM.n Data Entered)
PREFACE

This report was prepared for the Safety and Survivability Technical Area of the Applied Technology Laboratory, U. S. Army Research and Technology Laboratories (AVRADCOM), Fort Eustis,

Virginia, by Simula Inc. under Contract DAAJ02-77-C-0021, initiated in September 1977. The Department of the Army Project

Number is IL162209AH76. This guide is a revision of USAAMRDL Te.vhnicAl Report 71-22, Crash Survival Design Guide, published

• • -6ctober 6tobe........... 1971. .........

'•!U A .S. major Army-sponsored portion of the research data contained in aircraft herein crashworthiness was taken fromconducted from 1960 to 1979. Acknowledgment is extended to the

U. S. Air Force, Federal Aviation Administration, NASA, and

U. S. Navy for their research in crash survival. Appreciation

is extended to the following organizations for providing accident case histories leading to the establishment of the impact

conditions in aircraft accidents:

U U. S. Army Safety Center, Fort Rucker, Alabama.

'9 Civil Aeronautics Board, Washington, D. C.

e U. S. Naval Safety Center, Norfolk, Virginia.

* U. S. Air Force Inspection and Safety Center, Norton t.• / Air Force Base, California.

Additional credit is due the many authors, individual companies, and organizations listed in the bibliographies for their

contributions to the field. The contributions of the following authors to previous editions of the Crash Survival Design

Guide are most noteworthy:

D. F. Carroll, R. L. Cook, S. P. Desjardins, J. K. Drummond, J. H. Haley, Jr., A. D. Harper, H. G. C. Henneberger,

N. B. Johnson, G. Kourouklis, W. H. Reed, S. H. Robertson,

L. M. Shaw, J. W. Turnbow, and L. W. T. Weinberg.

- - This volume has been prepared by Dr. D. H. Laananen. Data from

the investigations of recent Army aircraft a66id7e*5Ci-were provided by the U. S. Army Safety Center. R. F. Chandler of the

FAA Civil Aeromedical Institute furnished assistance in locating recent information on human tolerance, anthropometry, and

crash test dummies.

.3
TABLE OF CONTENTS

Section Page

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

LIST OF ILLUSTRATIONS ... ................ 7

LIST OF TABLES. . . . . . . . . . . . . . . . . . . . 10

INTRODUCTION . . . . . . . .1 .

CHAPTER 1. BACKGROUND DISCUSSION. ........... . 14

CHAPTER 2. DEFINITIONS . .. . . . ... .. .. 17

2.1 GENERAL TERMS. .............- 17

2.2 AIRCRAFT PARAMETERS. .......... 18

2.3 HUMAN BODY PARAMETERS. . . . . . . . .. 21

CHAPTER 3. AIRCRAFT CRASH ENVIRONMENT . . . . . . . 23

3.1 INTRODUCTION ........ 23

3.2 VELOCITY CHANGES IN MAJOR IMPACTS. . . . 24

3.2.1 Vertical Velocity Changes. .v. * e a . . 24

3.2.2 Longitudinal Velocity-Changes. . . . . . 25

3.2.3 Lateral Velocity Changes . . . .... . 27

3.2.4 Combined Velocity Changes .... .... 27

3.3 IMPACT ACCELERATIONS.... 29

3.3.1 Vertical Impact Accelerations... .... 30

3.3.2 Longitudinal Impact AcceleratioA&. . . . 30

3.3.3 Lateral Impact Accelerations . . . . . . 30

3.4 IMPACT ATTITUDES . ........... 31

3.5 IMPACTED TERRAIN . . . . . . . . . . . . 32

3.6 IMPACT INJURY FREQUENCIES. . . . o . . . 33

3.7 ACCIDENT INFORMATION RETRIEVAL SYSTEM. . 35

CHAPTER 4. HUMAN TOLERANCE TO IMPACT. . . . . . . . 39

4.1 INTRODUCTION 39

4.2 FACTORS AFFECTING HUMAN TOLERANCE . 40

4.3 WHOLE-BODY ACCELERATION TOLERANCE ... 43

4.3.1 Spineward (-G ) Acceleration ...... 43

4.3.2 Sternumward (IG ) Acceleration . 45

4.3.3 Headward (+Gz) z cceleration ... 45

4.3.4 Tailward (-G ) Acceleration. ...... 48

4.3.5 Lateral (G ) Acceleration .......... ... 48

4.4 HEAD IMPAC* TOLERANCE. . . . . ..... 50

4.4.1 Weighted Impulse Criterion ....... 52

'Al 5

Omw m PA= &Aw ... .. ' '

-... .... ..... i |....... . .. -i i,•:•" . ... i: ..... ; "i.,•' ... i i '*'i ... I• ~ '
TABLE OF CONTENTS (CONTD)

Section Page

4.4.2 J-Tolerance . . . ................... . 54

4.4.3 Effective Displacement Index . . . . . . 57

4.4.4 Strain Energy Considerations . . . . . . 57

4.4.5 Comparison of Head Injury Predictors . . 58

4.5 NECK IMPACT TOLERANCE . . . . . . . . 59

4.6 CHEST IMPACT TOLERANCE . . . . o . . . . 60

4.7 ABDOMINAL IMPACT TOLERANCE . . . . . . . 62

4.8 SPINAL INJURY TOLERANCE. . . . . . . . . 63

4.8.1 Dynamic Response Index . . . 63

4.8.2 Wayne State University Two-Dimensional SModel . . . : . . .. . . . . . . . . .. 671

4.8.3 Air Force Head-Spine Model . . . . . 68

4.8.4 Vertebral Properties . . . . . . . . . . 71

4.9 LEG INJURY TOLERANCE . . . . . . . . . . 72

4.10 ABBREVIATED INJURY SCALE . . . . . . . . 73

CHAPTER 5. OCCUPANT MOTION ENVELOPES. . . . . . . . 75

5.1 INTRODUCTION . . . . . . . . . . . . . . 75

5.2 FULL RESTRAINT . . . . . . . . . . .. . 76

5.23 LAP-BELT-ONLY RESTRAINT . . . . . . . . . 76 ¶1CHAPTER 6. HUMAN BODY DIMENSIONS AND MASS

DISTRIBUTION ............ . 83

6.1 INTRODUCTION . . . . . . .. ..... 83

6.2 ANTHROPOMETRY . . . . . . 83

6.2.1 Conventional Anthropometric

Measurements ..•.• . . . . . . . . . 83

6.2.2 Body Joints and Ranges of motion . . . . 85

6.3 INERTIAL PROPERTIES. . . . . . . . . . . 90

6.4 SCALING OF MEASUREMENTS. . . . . . . . . 93

CHAPTER 7. CRASH TEST DUMMIES . . . . . . . . . . . 95

7.1 INTRODUCTION . . . . . . . . . . . . . 95

7.2 DUMMY TECHNOLOGY . . . . . . . . . . . . 95

7.2.1 History of Dummy Development . . . . . . 95

7.2.2 Part 572 Dummy . . . . . ... . 96

7.2.3 Other Recent Dummy Designs, . . . . .. 97

7.3 COMPARISON OF DUMMY AND HUMAN RESPONSE . 102

7.4 SUITABILITY OF DUMMIES FOR AIRCRAFT

SYSTEM EVALUATION. . . . . . . . . . . . 103

A 6

"•~
1

TABLE OF CONTENTS (CONTD)

"Section Page

REFERENCES .............................. 107

BIBLIOGRAPHY. .............. . . . . . . . 115

LIST OF ILLUSTRATIONS

Figure

* 1 Aircraft coordinate and attitude

directions . . . ... ... . . . 18

2 Typical aircraft floor deceleration

pulse. . . . . . . . . . . 19

3 Terminology for directions of forces on

the body . f. . f. .f . . .f . . .f. . . . . 21

4 Vertical velocity changes for survivable

rotary- and light fixed-wing aircraft

accidents . . . . .. .f. .. . 25

5 Distribution of longitudinal velocity

changes for survivable rotary- and light

fixed-wing aircraft accidents . . . . . . 26

6 Design velocity changes - off-axis

requirements .f.t . .f.f.t.f. . . . . . . 28

7 Frequency of occurrence of aircraft

pitch angle at impact for survivable

attack and cargo helicopter accidents,

1971-76 . . . . . . . . . . . . . . . . . 31

8 Frequency of occurrence of aircraft

roll angle at impact for survivable

attack and cargo helicopter accidents,

1971-76 . . . f . . . . . f . a f 32

9 Frequency of injuries to body parts

in U. S. Army aircraft accidents,

1971-76 . . . . .f. . .f . .f . . . 37

7

.............................................. ................ ..............

• , n T "'" ...... T 'F 'i .. .i f i° ==4
LIST OF ILLUSTRATIONS (CONTn

Figure Page

10 Pelvic rotation and submarining caused

by high longitudinal forces combined

with moderate vertical forces. . . . . . 42

11 Duration and magnitude of spineward

acceleration endured by various

subjects . . . . . . . . . . . . . . . . 44

12 Initial rate of change of spineward

acceleration endured by various

subjects . . . . . . . . . . . . . . . 46

13 Duration and magnitude of headward

acceleration endured by various

subjects . . . . . . . . . . . . . . . . 47

14 Initial rate of change of headward

acceleration endured by various

subjects . . . . . . . . . . . . . . . . 49

15 Wayne State Tolerance Curve for the

human brain in forehead impacts

against plane, unyielding surfaces . . . 53

16 Sample calculation of a Severity Index . . . . . . . . . . . . . . . . . . 55

17 ' Damped, spring-mass system used in

computing J-tolerance ......... 56

18 Comparison of SI, EDI, and kinematics

of six frontal impacts producing linear

fracture . . . . . . . . . . . . . 0 58

19 AIS injury rating versus normalized

chest deflection ....... . . . . 61

20 Spinal-injury model. . . . . . . . . . . 64

21 Probability of spinal injury estimated

from laboratory data compared to

operational experience . . . . . . . . . 66

22 Comparison of model output and experi- mental data for 10 G runs with the

spine in the (a) erect and (b) hyperextended modes . . . . . . . . . 69

..... ............
Ii

LIST OF ILLUSTRATIONS (CONTD)

Figure Page

23 Three-dimensional head-spine model . . . 71

24 Femur injury criterion .... . . .. 74

25 Full-restraint extremity strike

envelope - side view . . ..... ... 77

26 Full-restraint extremity strike

envelope - top view.... ....... 78

27 Full-restraint extremity strike

envelope - front view. . . . . . . . . . 79

28 Lap-belt-only extremity strike

envelope - side view.. 80

S29 Lap-belt-only extremity strike

envelope top view. 81

30 Lap-belt-only extremity strike

envelope - front view .... ..... 82

31 Conventional seated anthropometric

dimensions . . . . . . . . . . . . . . . 84

32 Normal distribution curve. . . . . . . . 86

33 Path of instantaneous center of rotation

during shoulder abduction. . . . . . . . 87

34 Sitting skeletal joint locations based

on a 50th-percentile Army aviator. . . . 88

35 Joint ranges of motion... . . . . . . 89

36 Mass distribution of seated torso

referenced to the skeletal structure

for a 50th-percex-tile Army aviator . . . 92

37 Dummy external dimensions. . . . . . . . 98

38 Comparison of mean head resultant

acceleration responses for three different dummy designs . . . . . . .... 104

(9

. ,.V,
LIST OF ILLUSTRATIONS (CONTD)

Figure Page

39 Comparison of mean head resultant

acceleration responses for HSRI dummy

conducted at two laboratories ...... 105

LIST OF TABLES

Table

1 Distribution of terrain impacted by

Army aircraft during the period 1971

through 1976 . . . . . . . . . . . . . . 33

2 Frequency of major and fatal injuries

to each body part as percentages of

total major and fatal injuries . . . . . 34

3 Frequency of fatal injuries to each

body pert as percentages of all fatal

injuries . . . . . . . . . . . . . . . . 36 14 Summary of anthropometric data for U. S.

Army aviators. .......... . . . 85

5 Summary of anthropometric data for

soldiers . . . . . . . . . . . 86

6 Range of joint rotation ....... . 89

7 Center-of-mass distribution of seated

torso - 50th-percentile Army aviator . . 91

8 Segment moments of inertia about the

center of mass. . .. ........ . 93

9 Dummy external dimensions (Part 572) . . 99

10 Dummy component weights (Part 572) . . . 100

11 Center-of-gravity locations

(Part 572) . . . . . . . . . . . . . . . 100

12 Hybrid II mass moments of inertia. . . . 101

10

.........
INTRODUCTION

For many years, emphasis in aircraft accident investigation

was placed on finding the cause of the accident. Very little

effort was expended in the crash survival aspects of aviation

safety. However, it became apparent through detailed studies

of accident investigation reports that large improvements in

crash survival could be made if consideration were given in

the initial aircraft design to the following general survivability factorst

1. Crashworthiness of Aircraft Structure - The ability of the aircraft structure to maintain living space

for occupants throughout a crash.

2. Tiedown Chain Strength - The strength of the linkage preventing occupant, cargo, or equipment from becoming missiles during a crash sequence.

i 3. Occupant Acceleration Environment - The intensity

and duration of accelerations experienced by occupants (with tiedown assumed intact) during a crash.

4. Occupant Environment Hazards - Barriers, projec- tions, and loose equipment in the immediate vicinity of the occupant that may cause contact injuries.

5. Postcrash Hazards - The threat to occupant survival

posed by fire, drowning, exposure, etc., following the impact sequence.

Early in 1960, the U. S. Army Transportation Research Command*

initiated a long-range program to stvidy all aspects of aircraft safety and survivability. Through a series of contracts

with the Aviation Safety Engineering and Research Division

(AvSER) of the Flight Safety Foundation, the problems associated with occupant survival in aircraft crashes were studied

to determine specific relationships between crash forcer,

structural failures, crash fires, and injuries. A series of

reports covering this effort was prepared and distributed by

the U S Army, beginning in 1959. In October 1965, a special proje t initiated by the U. S. Army consolidated the design

criteria presented in these reports into one technical document suitable for use as a designer's guide by aircraft design

, *Now the Applied Technology Laboratory, Research and Technology Laboratories of the U. S. Army Aviation Research and De-

"velopment Command (AVRADCOM).

11

.. 4 
Ir

engineers and other interested personnel. The document was to

be a summary of the current state of the art in crash survival

design, using not only data generated under Army contracts,

but also information collected from other agencies and organizations. The Crash Survival Design Guide, first published

in 1967, realized this goal.

Since its initial publication, the Design Guide has been revised several times to incorporate the results of continuing

research in crashworthiness technology. The last revision

of TR 71-22 was the basis for the criteria contained in the

Army's crashworthiness military standard, MIL-STD-1290(AV),

"Light Fixed- and Rotary-Wing Aircraft Crashwort1 iiness" (Reference 1). This current revision, the fourth, ontains the

most comprehensive treatment of all aspects of .rcraft crash

survival now documented. It can be used as a 9gneral text to

establish a basic understanding of the crash environment and

the techniques that can be employed to improve chances for

survival. It also contains design criterie and checklists

on many aspects of crash survival and thus can be used as a

source of design requirements.

The current edition of the Aircraft Crash Survival Design

Guide is published in five volumes. Volume titlee and general

subjects included in each volume are as follows:

Volume I - Design Criteria and Checklists

Pertinent criteria extracted from Volumes II through V,

presented in the same order in which they appear in those

volumes.

Volume II - Aircraft Crash Environment and Human 1olerance

Crash environment, human tolerance to impact, military

anthropometric data, occupant environment, test dummies,

accident information retrieval.

Volume III - Aircraft Structural Crashworthiness

Crash load estimation, structuk.al response, fuselage and

landing gear requirements, rotor requirements, ancillary equipment, cargo restraints, structural modeling.

1. Military Standard, MIL-STD-1290(AV), LIGHT FIXED- AND

ROTARY-WING AIRCRAFT CRASHWORTHINESS, Department of Defense, W!ashington, D. C., 25 January 1974.

12

4. . .
Volume IV - Aircraft Seats, Restraints, and Litters

Operational and crash environment, energy attenuation,

seat design, litter requirements, restraint system design, occupant/restraint system/seat modeling.

Volume V - Aircraft Postcrash Survival

Postcrash fire, ditching, emergency escape, crash locator

beacons.

This volume (Volume II) contains information on the aircraft

crash environment and the response of the human body to such

an environment. Following a general discussion of aircraft

crashworthiness in Chapter 1, a number of words commonly referred to in discussing the crash environment are defined in

Chapter 2. Chapter 3 describes the crash environment itself,

in terms of impact conditions, terrain, and the nature and

frequency of different types of injuries. Chapter 4 discusses

the tolerance of the human body and various body parts to impact loading. Chapter 5 presents data on occupant motion during a c~ash. Chapter 6 provides data on human anthropometry

that may be useful directly, as in cockpit design, or indirectly, as in preparing input for computer simulation models

such as those discussed in Volume IV. Chapter 7 describes the

crash test dummies used in evaluating protective systems such

as seats and restraints.

13

• ,','.. ,.•t .
1. BACKGROUND DISCUSSION

This volume deals with the variables involved in the Army aircraft crash environment and the effects of forces in that environment on the human body. An understanding of the environment and of the ability of the human body to survive it is

necessary for the effective design of more crashworthy aircraft. The following background discussion presents general

considerations that are of importance in understanding and applying the included information.

The overall objective of designing for crashworthiness is to

eliminate unnecessary injuries and fatalities in relatively

mild impacts. Results from analyses and research during the

past several years ha'e shown that the relatively small cost

4 in dollars and weight of including crashworthy features is an

extremely wise investment. Consequently, new generation aircraft are being procured to rather stringent crashworthy requirements.

To maximize aircraft crashworthiness, or, in the sense being

discussed, to provide as much occupant protection as possible,

all aspects of the complete system must be considered. In

other words, every available subsystem must be employed to the

fullest extreme in order to maximize the protection afforded

to vehicle occupants. When an aircraft impacts the ground,

deformation of the ground absorbs some energy. This is an uncontrolled variable since the quality of the impacted surface

usually cannot be selected by the pilot. If the aircraft

lands in the proper attitude, the landing gear can be used

to absorb a significant amount of the impact energy. After

stroking of the gear, crushing of the fuselage provides the

next level of energy absorption. Of course, one of the functions of the fuselage is to provide a protective shell around

the occupant while energy-absorbing stroke is occurring outside the shell. The functions of the seat and restraint system are to restrain the occupant within the protective shell

during the crash sequence and to provide additional energyabsorbing stroke to further reduce the loads. The structure

and components immediately surrot)-,ding the occupant also must

be considered. Structures such as cyclic controls, glare

shields, instrument panels, and sidewalls must be delethalized in some manner if they lie within the strike envelope of

the occupant.

The original edition of this design guide dealt primarily with

modifications that could be made to existing aircraft to increase their crashworthiness; however, now, two approaches to

improving aircraft crashworthiness are open. The first approach is to influence the design of new aircraft, and the

1: 14
second is to improve the crashworthiness of existing aircraft.

Obviously, much higher levels of crashworthiness can be

achieved in the design and development of new aircraft if

crashworthiness is considered from the beginning. This is

being accomplished at the present time through the use of procurement packages that include pertinent specifications that

require certain levels of crashworthiness of various subsystems as well as for the entire aircraft. However, some of the

available potential is still being lost due to the historical

approach used in designing aircraft. The basic aircraft is

designed leaving space and providing attachment provisions for

subsystems. Later the subsystems are designed and then are

limited by the previously established, somewhat arbitrary boundary conditions. The boundary conditions may unnecessarily limit the performance of the subsystems. The better approach

is the systems approach in which all systems and subsystems

are, at least preliminarily, designed at the same time. This

approach enables subsystem considerations to affect thn larger

systems and will produce a more nearly optimum vehicle.

The same principles for improving crashworthiness 6an be ap- plied to the retrofit of existing aircraft; however, the "castin-concrete" status of existing production structure is a more

costly and difficult obstacle to overcome. When crashworthiness features must be included through retrofit, the level that

can be achieved is usually reduced. Even in retrofit situations, however, the overall objective can be met; i.e., occupant protection can be maximized to eliminate unnecessary injuries.

As mentioned above, the entire system should be considered in

any analysis resulting in apportionment of the crash energy to

be absorbed by the various components. However, any valid systems approach will consider probable alternate crash environments wherein all subsystems cannot perform their desired functions; for example, an impact situation in which the landing

gear cannot absorb its share of the impact crash energy because

of angle of impact or loss of gear. To achieve the overall

goal, therefore, minimum levels of crash protection have been

required of the various individual subsystems, such as the seat.

In earlier editions of the Design Guide, the requirements to

provide occupant protection in crashes up to and including the

severity of the 95th-percentile survivable crash pulse was expressed. With the deployment of aircraft designed for crash

safety, the link to the 95th-percentile survivable crash pulse

has been dropped, and the recommended design environment is

simply presented as the design pulse. Obviously, the severity of a 95th-percentile survivable crash pulse will be much

l15
greater for the new aircraft than for aircraft having no crashworthy requirements placed upon them during their development.

The extent of the crash protection provided to the occupant

cannot indefinitely continue to be linked to the survivability

of the crash as improved crashworthiness increases the severity of the survivable crash producing a never-ending increase

in the level of crashworthiness at the expense of aircraft performance. The crashworthiness levels recommended herein are

felt to be a near optimum mix of requirements including consideration of cost, weight, and performance. The crash environments selected for design purposes in this volume are identical to the historical 95th-percentile survivable crash pulses.

Also in earlier editions of the Design Guide, information was

provided on design of fixed-wing transport aircraft. Considering the volume of new information on crashworthy design, in

an effort to ensure that the size of this document remains

within reasonable limits, only the primary aircraft in the

Army inventory are considered. Therefore, information given

herein is intended to apply to rotary-wing aircraft and light

fixed-wing aircraft, defined by a mission gross weight of

12,500 lb or less.

16

..... ......
2. DEFINITIONS

4k. . 2.1 GENERAL TERMS

The following text defines words commonly used in discussions

of the aircraft crash environment:

* The Term "G"

The ratio of a particular acceleration (a negative

acceleration may be referred to as a deceleration)

to the acceleration due to gravitational attraction *1

at sea level (32.2 ft/sec3 ). With respect to the

crash environment, unless otherwise specified, all

acceleration values (G) are those at a point approx- imately at the center of the fuselage floor. In accordance with common practice, this report will refer to accelerations measured in G. To illustrate,

it is customarily understood that 5 G represents an

acceleration of 5 x 32.2, or 161 ft/seca. As a result, crash forces can be thought of in terms of

multiples of the weight of objects being accelerated. Therefore, also in keeping with common practice, the term G is used in this document to define

accelerations or forces.

"• Survivable Accident

An accident in which the forces transmitted to the

occupant through his seat and restraint system do

not exceed the limits of human tolerance to abrupt

accelerations and in which the structure in the oc- cupant's immediate environment remains substantially

intact to the extent that a livable volume is provided for the occupants throughout the crash sequence.

"" survival Envelope

The range of impact conditions--including magnitude

and direction of pulses and the duration of forces

occurring in an aircraft accident--wherein the occupiable area of the aircraft remains substantially

intact, both during and following the impact, and

the forces transmitted to the occupants do not ex- •' ceed the limits of human tolerance when current

state-of-the-art restraint systems are used.

It should be noted that, where the occupiable volume

is altered appreciably through elastic deformation

17
during the impact phase, survivable conditions may

not have existed in an accident that, from postcrash

inspection, outwardly appeared to be survivable.

2.2 AIRCRAFT PARAMETERS

. Aircraft Coordinates

Positive directions for velocity, acceleration, and

force components and for pitch, roll, and yaw are

illustrated in Figure 1.

z

Figure 1. Aircraft coordinate and attitude directions.

2.8 )
* Velocity Change in Major Impact (AV)

The decrease in velocity of the airframe during the

major impact, expressed in feet per second. The ma- Sor impact is the one in which highest forces are

incurred, not necessarily the initial impact. For

the acceleration pulse shown in Figure 2, the major

impact should be considered ended at time T2 . Elastic recovery in the structure will tend to reverse

the direction of aircraft velocity prior to T2 .

A

Peak

Average

Q)

U

Uo

T T -TTT 1 2

i•~ i T-T1

Figure 2. Typical aircraft floor deceleration pulse.

19 -- r -- --- - -
Should the velocity actually reverse, its direction

must be considered in computing the velocity,change.

For example, an aircraft impacting downward with a

vertical velocity component of 30 ft/sec and rebounding with an upward component of 5 ft/sec should

be considered to experience a velocity change

AV = 30 - (-5) = 35 ft/sec

during the major impact. After the aircraft rebounds upward, gravity will accelerate it downward

again, as illus-rated by the negative acceleration

between T2 and T3 in Figure 2.

* Longitudinal Velocity Change

The decrease in velocity during the major impact

measured along the longitudinal (roll) axis of the

aircraft. The velocity may or may not reach zero

during, the major impact. For example, an aircraft

impacting the ground at a forward velocity of 100

ft/sec and slowing to 35 ft/sec before rebounding

into the air would experience a longitudinal velocity change of 65 ft/sec during this impact.

* Vertical Velocity Change

The decrease in velocity during the major impact

measured along the vertical (yaw) axis of the aircraft. The vertical velocity generally reaches zero

J Iduring the major impact.

e Abrupt Accelerations

Accelerations of short duration primarily associated

with crash impacts, ejection seat shocks, capsule

impacts, etc. One second is generally accepted as

the dividing point between abrupt and prolonged accelerations. Within the extremely short duration

range of abrupt accelerations (0.2 sec and below),

the effects on the human body are limited to me-

.1 chanical overloading (skeletal and soft tissue

stresses), there being insufficient time for functional disturbances due to fluid shifts.

20

".....................................................
& 2.3 HUMAN BODY PARAMETERS

* Human Body Coordinates

In order to minimize the confusion sometimes created

by the terminology used to describe the directions

of forces applied to the body, a group of NATO scientists compiled the accelerative terminology table of

equivalents shown in Figure 3 (Reference 2). Terminology used throughout this guide is compatible with

the NATO terms as illustrated.

Headward o

(+Gz) Direction of .1 accelerative force

Vertical

Back to chest Headward - Eyeballs-down

(sternumward) Lateral right (+G(+Cy) Transverse

Lateral right - Eyeballsleft

Lateral left - Eyeballsright

Back to chest - Eyeballs- S~in

Lateral left Chest Chest to back - Eyeballs-

(-G to back out

y (spineward) Note:

Tailward G The accelerative force on

G)the body acts in the same direction as the arrows.

Figure 3. Terminology for directions of forces on the body.

* Human Tolerance to the Crash Environment

Obviously, the tolerance of the human body to crash

environments is a function of many variables, including the unique characteristics of each person as

2. Gell, C. F., TABLE OF EQUIVALENTS FOR ACCELERATION TERMINOLOGY, Aerospace Medicine, Vol. 32, No. 12, December

1961, pp. 1109-1111.

S4 21

4, 44
well as the loading variables. The loads applied to

the body include decelerative loads imposed by seats

and restraint systems as well as localized forces

due to impact with surrounding structure. Tolerable

levels of the decelerative loads depend on the direction of the load, the orientation of the body,

and the means of applying the load. For example,

the critical nature of the loads parallel to the occupant spine manifests itself in any of a number of

spinal fractures. Forces perpendicular to the occupant spine can produce spinal fracture through flexure that results from jackknife bending over a lapbelt-only restraint. The lap belt might inflict

injuries to the internal organs if it is not retained on the pelvic girdle but is allowed to exert

its force above the iliac crests in the soft stomach

region. Excessive rotational or translational acceleration of the head can produce concussion. Further, skull fracture can result from localized impact with

surrounding structure. Therefore, tolerance is a

function of the method of occupant restraint as well

as the variable of specific occupant makeup.

For the purposes of this document, human tolerance

is defined as a selected array of parimeters that

describe a condition of decelerative *.oading for

which it is believed there is a reasona-Ae probability for survival without major injury. As used in

this volume, designing for the limits of human tolerance refers to providing design features that will

maintain these conditions at or below their tolerable levels to enable the occupant to survive the

given crash environment.

Submarining

A rotation of the hips under and about the lap belt

as a result of a forward inertial load exerted by

deceleration of the thighs and lower legs accompanied by lap belt slippage up and over the iliac

crests. Lap belt slippage up and over the iliac

crests can be a direct result of the upward loading

of the shoulder harness straps at the center of the

lap belt.

e Dynamic Overshoot

The amplification of decelerative force on cargo or

personnel above the floor input deculerative force

(ratio of output to input). This amplification is a

result of the dynamic response of the system.

22

, ' * 

J .i
3. AIRCRAFT CRASH ENVIRONMENT

3.1 INTRODUCTION

This chapter presents the results of statistical studies conducted to determine impact conditions that occurred in accidents of various types of U. S. Army aircraft. Data were accumulated through a study of U. S. Army accidents for the

period 1 July 1960 through 30 June 1965 and 1 January 1971

through 31 December 1976. Additional data considered pertinent were obtained from Civil Aeronautics Board, U. S. Navy,

and U. S. Air Force accident reports.

Aircraft included here are rotary-wing and light fixed-wing of

mission gross weight no greater than 12,500 lb. The accident

cases selected were limited to those in which one or more of

the following factors applied: (1) Substantial structural damage, (2) postcrash fire, (3) personnel injuries, and (4) at

least one person surviving the crash. Mid-air collisions and

other accidents resulting in catastrophic uncontrolled free

falls from altitudes of a hundred feet or more were not considered. Such accidents almost invariably result in random, unpredictable crash kinematics and nonsurvivable impact

forces, and are of little value in establishing realistic

crash survival envelopes that would be useful to the aircraft

designer. Although accidents involving postcrash fire were

considered where possible, analysis of impact forces in many

of the accidents involving fire was impossible due to extensive burn damage to the aircraft, and these cases were not

used in the study. Still other accident reports simply provided insufficient or inadequate data to permit a detailed

analysis of the case.

The impact accelerations computed or estimated for each accident by the original accident investigation boards were analyzed and recalculated or estimated by the survey team. This

action was taken to minimize the variation in findings caused

by the diversity in training and experience of the various members of the boards who made the initial calculations. Those

accidents involving high impact forces that appeared to be near

the upper limits of survivability were analyzed in depth. In

the Icss severe accidents, which made up the bulk of the accidents surveyed, the impact forces given by original investigation boards or teams were analyzed primarily to eliminate any

major errors or misjudgments by the aircraft accident board.

Minor individual variations in findings in these less severe

accidents were acceptel with the assumption that the sum of

the variations would tend to be zero and would not affect the

accuracy of the final data.

23
Altogether, 563 rotary-wing accidents and 92 fixed-wing accidents were reviewed in the preparation of earlier editions of

the Design Guide. However, only 373 total cases are used in

establishing the impact conditions outlined here. Impact attitude data from an additional 108 attack and 10 cargo helicopters collected during 1971-1976 also are presented.

Analysis of the data from Army accident records showed a similarity in impact conditions between rotary- and light fixedwing STOL aircraft (0-1, U-6, U-1). The typical severe STOL

crash occurred with a large vertical velocity component in

addition to a longitudinal component. The small amount of

underfloor structure means that high vertical accelerations

result at impact. The small amount of structure below the

floor is a primary reason for high vertical acceleration in

rotary-wing aircraft also. The data showed that, except for

the lateral direction, the similarities between rotary-wing

and lignt STOL aircraft impact conditions were sufficient to

allow treating them as being the same.

3.2 VELOCITY CHANGES IN MAJOR IMPACTS

The velocity change of the airframe during major impacts (as

defined in 2.2) was difficult to obtain in the majority of the

accidents analyzed because Qf an inability to fix the impact

velocity. The impact velocities in 40 accidents were known,

and the velocity changes in the major impact pulses were estimated with errors probably not exceeding plus or minus 20

percent.

3.2.1 Vertical Velocity Changes

The vertical velocity change is defined as the decrease in velocity during the major impact measured along the vertical

(yaw) axis of the aircralt. The vertical velocity generally reaches zero during the major impact.

The curve in Figure 4 shows the distribution of vertical velocity changes experienced by rotary- and light fixed-wing

aircraft in survivable accidents involving substantial structural damage or occupant injury.

The median of the distribution in Figure 4 is a vertical velocity change of approximately 24 ft/sec, with one-half of the

accidents involving vertical velocity changes greater than

this figure and one-half having vertical velocity changes

lower than this figure. For reference, 24 ft/sec is the velocity reached by a body during a free fall of 8 ft 11 in.

Ninety-five percent of the accidents represented by the curve

in Figure 4 incurred vertical velocity changes of less than

24
"50

U 40 - -...-

44L4

30 -

U

"Median 24 ft/sec

0

-

>

i > "1

0

0 10 20 30 40 50 60 70 80 90 100

Cumulative frequency of occurrence, percent

Figure 4. Vertical velocity changes for survivable rotaryand light fixed-wing aircraft accidents.

42 ft/seu. A free fall of 27 ft 5 in. is required for a body

to reach a velocity of 42 ft/sec.

3.2.2 Longitudinal Velocity Changes

The curve in Figure 5 depicts the distribution of longitudinal

velocity changes incurred by rotary- and light fixed-wing aircraft in survivable accidents involving substantial structural

25

S. ' ii, .
-I O -- r_....._ 50 -_

C.)

40

Median = 28 ft/sec

" 20 H

10

0 10 20 30 40 50 60 70 80 90 100 3

Cumulative frequency of occurrence, percent

Figure 5. Distribution of longitudinal velocity changes

fcr survivable rotary- and-light fixed-wing aircraft accidents.

26

--. "

- ,'.-
1

damage or occupant injury. The curve gives a relative frequency of occurrence of the stated longitudinal velocity

changes.

The median of the distribution curve in Figure 5 is a longitudinal velocity change of approximately 28 ft/sec, with one-half

of the accidents involving longitudinal changes greater than

this figure and one-half experiencing longitudinal changes lower than this figure.

1Ninety-five percent of the accidents represented by the curve

in Figure 5 incurred longitudinal velocity changes of less than

50 ft/sec.

3.2.3 Lateral Velocity Changes

The accident survey produced insufficient data to plot an accurate distribution of lateral velocity changes. However, the

engineering personnel partic.'pating in the study were able to

infer from the circumstances of the accident cases studied

that velocity changes in the lateral direction generally do

not exceed 25 ft/sec for light fixed-wing aircraft and cargo

and attack helicopters. Recent studies have indicated that

30 ft/sec is a more realistic number for other rotary-wing

aircraft.

3.2.4 Combined Velocity Changes

The resultant velocity change for combined longitudinal, vertical, and lateral components of a 95th-percentile survivable

accident of rotary- and light fixed-wing aircraft does not ap- ' ~pear to exceed 50 ft/sec. The vertical or lateral components

do not exceed their individual 95th-percentile values, i.e., 42 ft/sec vertically, and 30 and 25 ft/sec laterally for

rotary- and light fixed-wing aircraft, respectively.

Figure 6 shows plots of combined longitudinal, lateral, and

vertical velocity changes for helicopters, to be used in determining intermediate velocity change components. For light

fixed-wing aircraft and for cargo and attack helicopters, Figure 6(b) will still be correct, but (c) and (d) must be altered

for a lateral velocity change of 25 ft/sec instead of 30 ft/sec.

Since those accidents with the greatest longitudinal velocity

change are not necessarily the same accidents with the greatest vertical velocity change, vertical and longitudinal percentiles shown in Figures 4 and 5 cannot be vector summed to give

the combined velocity changes.

27
to~ 41

-44 14

4 44444

0 ON

4. 1

C H 31 

Ix.4

0

ILL9

N Ln

28U
In general, the components illustrated in Figure 6 are related

by the equation

V x2 +Vy2 +V z2 ,V R2

+ + = (1)

where Vx = longitudinal velocity change, ft/sec

V = lateral velocity change, ft/sec

Vz a vertical velocity change, ft/sec

VR - resultant velocity change, ft/sec

and the axes are those illustrated in Figure 1. The curves

have been terminated at 15 degrees, based on a study of accident reports discussed in Section 3.4. It was found that 95

percent of attack and cargo helicopter accidents occur at

pitch angles between +25 degrees (nose up) and -15 degrees

(nose down). Therefore, 15 degrees above the longitudinal axis appeared to be a reasonable point to which to extend the

longitudinal-vertical conditions of Figure 6(b). For consistency, the 15-degree limit was extended to the other directions, where insufficient data existed.

As mentioned in Section 3.3, impact decelerations for design

of the overhead structure have been determined. However, these

conditions are intended for aircraft inverting after impact,

not those impacting in an inverted attitude. Therefore# no

design velocity changes are presented here for impact on the

upper parts of the aircraft.

3.3 IMPACT ACCELERATIONS

Average aircraft floor decelerations were estimated by the

following procedures:

a Calculations where deceleration distance (s) and velocity change (V) were knowni i.e., the equation

G 

V2

avg 2gs (2)

was used in those cases for which the velocity at

the end of the pulse was zero.

29

" }A
"* Comparisons of accident configurations and damage

with crash test data.

"* Observations of failure or nonfailure of lap belts,

shoulder harnesses, and seats of known strength.

* Comparisons of injuries with generally established

human tolerance limits.

3.3.1 Vertical Impact Accelerations

An examination of the frequency of occurrence of vertical accelerations for impacts of rotary- and light fixed-wing aircraft showed that 95 percent of survivable accidents involved

average accelerations of less than 24 G. Peak accelerations

can be expected to reach approximately twice the average values

if triangular approximations to the crash pulses are assumed.

3.3.2 Longitudinal Impact Accelerations

Examination of the frequency of occurrence of longitudinal accelerations for impacts of rotary- and light fixed-wing aircraft showed that 95 percent of survivable accidents involved

average accelerations of less than 15 G. Peak accelerations

would be approximately twice the average values if triangular

approximations to the crash pulses are assumed.

3.3.3 Lateral Impact Accelerations

Significant lateral accelerations were found to be present in

the crashes studied, particularly in accidents where a rotarywing aircraft autcrotated into trees or where rotor blades

struck trees or other obstacles during normal operation (Reference 3). Impact with trees often causes the fuselage to

rotate and finally impact the ground on its side. Lateral ac- celerations also result in the types of secondary impacts associated with low-angle, high-velocity crashes and with nearly

all severe accidents. The data on lateral accelerations obtained in the survey were not sufficient to allow construction

of a distribution plot; however, it is believed that these accelerations rarely exceed 16 G for light fixed-wing aircraft

and cargo and attack helicopters, although they may reach

18 G for other helicopters.

3. Gupta, B. P., HELICOPTER OBSTACLE STRIKE TOLERANCE CONCEPTS ANALYSIS, Bell Helicopter Textron; Technical Report

78-46, Applied Technology Laboratory, U. S. Army Research

and Technology Laboratories (AVRADCOM), Fort Eustis, Virginia, April 1979, AD A069877.

30

0i 3WI

m'A

AJ
3.4 IMPACT ATTITUDES

Analysis of data for survivable accidents involving attack and

cargo helicopters from 1971 through 1976 resulted in the distributions of pitch and roll angles presented in Figures 7 and

8. The distribution of yaw angles appears much like the roll

angle histogram of Figure 8, with 72 percent of the impacts

occurring at zero or negligible yaw. Accidents reported as

nonsurvivable were not included in the analysis, but they often occurred at extremes of attitude. In a number of nonsurvivable accidents, the attitude at impact was not reported.

However, the number of these accidents with unknown attitude

does not appear to constitute more than 10 percent of the total number of accidents. Therefore, the data indicate that

existing attack and cargo helicopters are most likely to crash

at a small roll angle. I

40

S20-

' 2 7 1 I I Ii 

S-30 -20 -10 0 10 20 30

Pitch, degrees

Figure 7. Frequency of occurrence of

aircraft pitch angle at impact for survivable attack

and cargo helicopter accidents, 1971-76.

31

.. .
80 +

a,. 60_j

S40

0 20

1 7 7

-30 -20 -10 0 10 20 30

Roll, degrees

Figure 8. Frequency of occurrence of aircraft roll angle at impact for survivable attack

and cargo helicopter accidents, 1971-76.

3.5 IMPACTED TERRAIN '

The probability of impacting a given type of terrain may influence an aircraft design in several ways. For example, the

types of landing gear and the escape systems installed may be

based on the terrain most likely to be encountered in operation. A search of accident records at the U. S. Army Safety

Center resulted in the distribution of impacted terrain shown

in Table 1.

As discussed in Volume III, the lower fuselage stucture should

"be designed to minimize the effects of plowing or earth scooping during a longitudinal impact. For a primarily vertical

impact, loose soil would prove beneficial through additional

energy being absorbed by soil compaction. On the other hand,

in a primarily horizontal (longitudinal and/or lateral) impact, loose soil may increase the deceleration level if the

gear do not bend or fail or if the belly skin ruptures to

cause an earth-scooping effect.

32(

• .44
TABLE 1. DISTRIBUTION OF TERRAIN IMPACTED

BY ARMY AIRCRAFT DURING THE PERIOD 1971 THROUGH 1976

Terrain Percentage of impacts

Sod 43

Trees 30

Rocks 10

Prepared surface 7

Bog 6

Water 2

Snow 1

Ice 1

3.6 IMPACT INJURY FREQUENCIES

The data contained in this section were obtained from a search

of accident records at the U. S. Army Safety Center, and are

based on reports of accidents that involved a total of 4550

occupants during the period 1971 through 1976. The intention

here is to inform the designer of aircraft systems as to the

location of problem areas and the significance of the information contained in subsequent chapters, which deal with human

tolerance to impact and the occupant environment.

Table 2 displays the frequencies of major and fatal injuries

to various body parts as percentages of all major and fatal

injuries reported. The injuries are broken down according to

aircraft type in order to point out the effect of this variable where it is significant. It can be noted that the frequency of serious vertebral injuries is lower for light fixedwing aircraft and cargo helicopters than for the other helicopter types. A reasonable explanation for this observation is

that the vertical component of acceleration experienced by the

occupants of light fixed-wing aircraft and cargo helicopters

during impact is lower than for the other aircraft. As pointed

out in Chapter 4, the vertical component of acceleration is

likely to cause vertebral damage. Fixed-wing aircraft, often

33
00 1 Jo a 0 0 m~ N cm

CA1- .-4 -4 H H

04 0'

P4- r4 4)k-i N- r-4 0ý LA N- O L

1-4 C4.4

4us

u 9: ~ 1 IV 0 P LA P-i 0 LM HP- 4 H a 0 0 * S 5 4 *.- r- (n N N N V14.

(d H

43 V) 0 H

H 4)

H to H w 41O C h P- H- LA co P~H -4 J

a) -W- N H- PH N H

S0 0 r-4 CD~H 0 ýo 0%

.O N~- r4 9 0 0 0; N4

H LI44 4r-HH

a 0 4)

O 41441

9 JHr- LA N 0 0) N Ch w. N %0O

0 04

-ri 0 w. H LA t. w co fn P-4

H- tE-4 4J. Or0 N 0 ti 4i L

4JHHHH H

4.4 1) 9 S "4 ) )

(d1. 04- M *a)

.D 0) - E-44-4M

34
subject to a stall-spin type of crash, may have a lower frequency of flat impacts than helicopters. Cargo helicopters,

being larger, have a greater crujh distance beneath the floor

that can attenuate the vertical impact forces.

The seriousness of head injuries is demonstrated by the data

of Table 2 where they appear as the leading cause of major and

fatal injuries. Their significance is further amplified by

the data of Table 3, where head injuries are shown to account

for approximately 31 percent of all fatal injuries, regardless

of aircraft type.

The information presented in Tables 2 and 3 is summarized in

Figure 9.

The influence of rollover conditions on the incidence of fatalities was studied in 198 major Army aircraft accidents during 1970-71. The aircraft studied were utility, attack, observation, and cargo helicopters. The accident data indicated

that the probability of a major accident resulting in fatal

injuries increases if a rollover occurs. Of 104 rollover accidents, 51 accidents, or 49 percent, resulted in fatalitiesg

of 94 no-roll accidents, 30 accidents, or 32 percent, resulted

in fatalities (Reference 4).

3.7 ACCIDENT INFORMATION RETRIEVAL SYSTEM

Crash survival design criteria are based on the realities of

actual aircraft accidents. It is necessary to know such facts

as aircraft attitude and velocity at impact, and the magnitude,

duration, and direction of the impact forces. These data are

currently derived by accident investigation teams who must estimate the impact conditions from their study of the wreckage

and terrain, witness accounts, medical reports, etc. These

same teams also must try to reconstruct preimpact events and

determine possible primary and contributing cause factors so

that preventive measures can be taken to decrease the incidence of aircraft accidents. Such accident reconstruction,

even when performed by the most skilled investigators, can

produce no more than estimated numbers and probable causes.

Accurate aircraft precrash and crash data are needed to establish a sound basis for more precise design criteria and operational techniques to reduce accidents and improve crash survivability.

4. Haley, J. L., and Hicks, J. E., CrASHWORTHINESS VERSUS

COST: A STUDY OF ARMY ROTARY-WING AIRCRAFT ACCIDENTS IN

PERIOD JAN 70 THROUGH DEC 71, paper presented at Aircraft

Crashworthiness Symposium, University of Cincinnati,

6 - 8 October 1975.

35

----------------------------
.4-)

4.4

() ~ ~ C; ~ c C4 

k -H -

0

cn cn -I 

41 ~ r% 0 r- -

'4.E-4

ON ~ W - N) 4

Ci N ~0 OrLA 0F-04A c ~ cnrn

".4

~~E-4~L t)0( LA 0 Io.D F 0 L

to to Nc c - o o .

4-4i o c P-4 c Po -I

q ~ U .- 0r 0- 0-~0 r- Q (A-

'-4 IV 0A

En 0 fn 0 N .

oI H In 1r' H 0 r

wN ON (n MHo f

Q.J N r4

E4 "-t n ) 

4 40 > ~4- WN

36 'el
4.-

4-4

0

4-

0-

~ 44 r-4

0

44.

rnr

dFa

37~'
To answer this need, the Applied Technology Laboratory of the

U. S. Army Research and Technology Laboratories (AVRADCOM) developed a program and awarded a contract for the conceptual analysis and preliminary design of an Accident Information Retrieval System (AIRS) suitable for helicopter installation.

As part of this effort, present state-of-the-art electromechanical crash recorders were examined and found to be both too

large and too heavy for practical helicopter installation.

Results of the study, documented in Reference 5, indicate that

a solid-state system of approximately 10 lb in weight and 200

in. 3 volume is feasible for installation in new Army aircraft,

such as UTTAS or AAH. The further development and eventual

installation of the AIRS would measurably reduce the cost and

time of accident investigation while enhancing the validity of

the accident data obtained. The data that such a system would

provide should be of great value in further refining crashworthy design criteria.

A

5. Burrows, L. T., AN ECONOMICAL APPROACH TO AN ACCIDENT IN- . ,IFORMATION RETRIEVAL SYSTEM (AIRS), Applied Technology

Laboratory, U. S. Army Research and Technology Laboratories (AVRADCOM); paper presented at the 15th Annual Survival and Flight Equipment Symposium, Las Vegas, Nevada,

December 1977.

38

J/J

o -- "'• • " • •'• ' "•' "•J ~ "• i " ,,,..,,... '- ..-.. ' . .-- - ... ._.. 
4. HUMAN TOLERANCE TO IMPACT

4.1 INTRODUCTION

The objective of this chapter is to provide the designer of

aircraft systems with a summary of available information on

tolerance of the human body to forces of the type experienced

in a crash environment. Knowledge of human tolerance is vital

to understanding and effectively applying the principles of

crashworthy design that are defined elsewhere in this guide.

A great deal of research has been conducted in the field of

biodynamics, and general guidelines and approximate end points

have been determined. However, there still, remain many areas

of uncertainty and disagreement, and much more research is

needed to provide accurate, proven figures. An obvious difficulty is that the usual test method of stressing a specimen

beyond the point of failure in order to establish tolerance

limits is not possible when one is dealing with injurious and

fatal ranges of forces acting on the human body. Experiments

using human volunteers have necessarily been conducted, for

the most part, at subcritical levels. The few instances that

iihave have inadvertently provided valuable, approached but often theuniverifiable, critical range data. of forces Test

animals such as chimpanzees, monkeys, bears, pigs, and mice

have been used in attempts to establish a better definition of

the injurious and fatal ranges of forces. Human cadavers also

have been utilized as test specimens. While these approaches

have provided valuable data in many areas of investigation, a

means of reliably extrapolating these results to the tolerance

of a live human is not yet available.

As discussed in Volume IV, mathematical models of the human

body have been used successfully in studying the overall kinematic response of the body to crash forces and in evaluating

crashworthiness of vehicle interiors. Mathematical models directly related to injury prediction are discussed in this

chapter. Anthropomorphic dummies, as mechanical models, have

been refined to remarkable levels of physical resemblance to

the human body. However, interpreting the results from mathematical and mechanical simulators remains a problem, as tolerable levels of the predicted variables are neither well defined nor widely agreed apon.

S39
The following sections discuss the factors that affect human

tolerance to impact and summarize the existing data on tolerable levels for various body parts. An in-depth review of research in human tolerance before 1970 can be found in Reference 6.

4.2 FACTORS AFFECTING HUMAN TOLERANCE

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

number of variables, including characteristics of the individi ual such as age, sex, and general state of health. Military

systems can be expected to be used by personnel who are generally 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 applic.ation 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 conditions.

The overall probability of survival in a crash depends to a

large extent on the manner of restraint. It would be extremely

difficult to prevent the arms and legs from contacting the cabin interior during a severe impact, but the use of upper and

lower torso restraints to prevent such critical body parts as

the head and chest from striking surrounding structure can

significantly reduce the probability of serious or fatal injury under given crash conditions.

The method of body restraint, of all the factors affecting human tolerance, offers the designer the greatest opportunity

for effective application of crashworthy design. The effectiveness of the restraint system is dependent upon the area

over which the total force is distributed, the location on the

body at which the restraint is applied, and the degree to

which it limits residual freedom of movement (Reference 7). j

6. Snyder, R. G., HUMAN IMPACT TOLERANCE, Paper 700398,

International Automobile Safety Compendium, Society of

Automotive Engineers, New York, 1970, pp. 712-782.

7. Rothe, V. E., et al., CREW SEAT DESIGN CRITERIA FOR ARMY

AIRCRAFT, TRECOM Technical Report 63-4, U. S. Army Transportation Research Command, Fort Eustis, Virginia, February 1963.

40.

m~
The greater the contact area between the body and the restraint system, the greater the human tolerance. The re- straint system should be located on the body at those points

that are best able to withstand the loads exerted by the decelerative force and that are best able to furthcr distribute

the force to the remainder of the body. These points are primarily the pelvic girdle and the shoulder structure. An additional restraint around the rib cage has been shown to increase

tolerance to spineward, eyeballs out (-G ) accelerations. Restraint systems located over soft tissudX tend to be much less

effective, often resulting in crushing of the viscera between

the restraint system and bony structures. Residual freedom of

movement should be limited to an absolute minimum consistent

with the necessary comfort and movements required by the

duties of the occupant.

When restrained only by a lap belt, the occupant's tolerance

to abrupt acceleration is relatively low. In forward-facing

seats, a longitudinal impact will cause A rotation of the upper torso over the belt, a whipping action'of the head, and

often, impact of the upper torso on the legs, resulting in

chest, head, and neck injuries. Head injuries due to impacts with the surrounding environment are very common for occupants

restrained only with lap belts. When longitudinal forces are

combined with a vertical component, there is a tendency for Sthe occupant to slip under the belt to some degree. This MO- tion, often referred to as submarining, can shift the belt up

over the abdomen. The longitudinal component of the pulse

then causes the upper torso to flex over the belt, with the

restraining force concentrated at some point on the spine and

not on the pelvic girdle. In this configuration, tolerance is

extremely low.

rThe addition of a shoulder harness greatly reduces injuries

from head impacts and helps to maintain proper spinal alignment

for strictly vertical impact forces. However, this standard

configuration may be unsatisfactory for impacts with both vertical and longitudinal components. Pressure by the upper torso

against the shoulder straps causes these straps to pull the lap

belt up into the abdomen and against the lower margin of the

rib cage. This movement of the lap belt allows the pelvis to

move forward under the lap belt, causing severe flexing of the

S.... spinal column, as shown in Figure 10. In this flexed position,

the vertebrae are very susceptible to anterior compression

fractures and, if the lap belt slips off the top of the pelvic

bone structure (over the top of the iliac crests), severe injury can occur as a result of viscera crushing. A lap belt

tiedown strap prevents raising of the lap belt by the shoulder

harness and may nearly double the tolerance to impact forces.

41

fti

!
- .•.• Shoulder harness pull K I

Pelvic joint

low

Expected torso displacement

Figure 10. Pelvic rotation and submarining caused by

high longitudinal forces combined with

moderate vertical forces.

The restraint system used in new Army aircraft crewseats, as

defined by MIL-.S-58095(AV), Reference 8, consists of a lap

belt, lap belt tiedown strap, and two shoulder straps connected by a single-point release buckle.

The amount of slack in the restraint system can affect tolerance to a given acceleration pulse. In general, the lower the

*! elongation properties of the link between the occupant and the

8. Military Specification, MIL-S-58095(AV), SEAT SYSTEM:

CRASHWORTHY, NON-EJECTION, AIRCREW, GENERAL SPECIFICATION

FOR, Department of Defense, Washington, D. C., 27 August

1971.

42

Jm' '
I

seat, the greater the occupant's tolerance to an abrupt acceleration. A loose restraint system also will result in the occupant's receiving a significantly greater magnification of

the accelerative force applied to the seat than would occur

with a snug system. The inertia of the occupant will cause

him to maintain a near constant velocity, independent of the

decreasing velocity of the seat, until the slack in the restraint system is taken up. As this point is reached, the velocity of the occupant is abruptly reduced to that of the seat,

resulting in relatively high G levels, even exceeding those of

the seat. This is often referred to as dynamic overshoot. Dynamic overshoot is a complex phenomenon involving the elasticity, geometry, mass distribution, and, thus, the natural frequency of the occupant, and the restraint and seat systems. An

example is discussed in detail in Volume IV.

4.3 WHOLE-BODY ACCELERATION TOLERANCE

The paragraphs in this section describe experimental results

applicable to acceleration of well-restrained (including fulltorso restraint) seated occupants. The tolerance levels pre- sented here define for the designer the limits of the environment that must be provided to enable survival of aircraft

occupants in a crash.

4.3.1 Spineward (-G ) Acceleration

The magnitude and duration of the applied accelerative force

have definite effects on human tolerance, as shown in Figure

11 (Reference 9). As indicated by this curve, a spineward

chest-to-back accelerative force of 45 G has been tolerated

voluntarily by some 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 11 shows that the tolerable

limits on acceleration loading are a function of duration.

The whole-body tolerance data displayed in Figure 11 were collected for a variety of full-torso restraint 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 significantly reduced.

9. 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, D. C., June 1959.

t ,43

. ,.. ,.,.,,•I 7, 
4.)

0 4.)

41 Q)

W.14

w H 

Gon

0 l

ll

.11

0 H 00

__o 0 04 > E-4

44

0 0.

- *V.,
With respect to whole-body deceleration, the rate of onset of

the applied force also has a definite, although not yet well

understood, effect on human tolerance. Under some impact conditions, the rate of onset appears to be a determining factor,

as indicated by the diagram in Figure 12 (Reference 9). Lower

rates of onset were more tolerable than higher rates under the

test conditions present. Under other impact conditions, such

as extremely short durations that occur in impacts from free

falls, rates of onset as high as 28,000 G/sec were survivable

and appeared to have little effect on human tolerance (Reference 10). It appears that, in certain ranges, the effects of

the rate of onset are related to the natural frequencies of

the body and of the various body organs (Reference 11).

4.3.2 Sternumward (+Gx) Acceleration

The human tolerance limit for sternumward, eyeballs-in (+Gd

acceleration has not been accurately r stablished. Due to- t

high degree of restraint provided by a full-length seat back

in this configuration, it can be safely assumed that tolerance

is greater than for spineward acceleration. A maximum of 83 G

measured on the chest with a base duration of 0.04 sec was experienced on one run in a backward-facing seat. However, the

subject was extremely debilitated, went into shock following

the test, and required on-the-scene medical treatment (Reference 12). Human tolerance to sternumward acceleration, therefore, probably falls somewhere between this figure of 83 G for

0.04 sec and 45 G for 0.1 sec, which is the accepted end point

for the -G (eyeballs-out) case. As in the case of +G accelerations, fhe human subject experiments employed head festraint

so that under aircraft operational conditions tolerable levels would be lower.

4.3.3 Headward (+Gz) Acceleration

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

the force is applied perpendicular to the long axis of the

10. Snyder, R. G., HUMAN TOLERANCE TO EXTREME IMPACTS IN FREE

FALL, Aerospace Medicine, Vol. 34, No. 8, August 1963,

pp. 695-709.

11. Stapp, J. P., JOLT EFFECTS OF IMPACT ON MAN, Brooks Air

Force Base, San Antonio, Texas, November 1961.

12. Beeding, E. L., Jr., and Mosely, J. D., HUMAN DECELERATION TESTS, Air Force Missile Development Center; AFMDC

Technical Note 60-2, Holloman Air Force Base, New Mexico,

January 1960, AD 23148.

p 45

. . . . -. -' . ,..--. . . ... ... .
SUBJECT SUPPORT

SHuman Lap, shoulder, thigh,

and chest straps

0 Human Lap, shoulder, thigh, and chest straps

Human Lap, shoulder, thigh,

and chest straps

17 Chimpanzee Military lap and shoulder

straps A Chimpanzee 3-in, aotton webbing, 5 Acceleration

horizontal, 2 vertical

straps

100-, - ,

C 0

1. L-No xiotk No shock

40 Cardiovascula hok hook (onjunotivaland hemorrhage) retinal

20 1.A..'

U0.

.0G

S-Onset rate max

0 't t0

a0

.001 .002 .004 .006 .01 .02 .04.,06 .1 .2 '.4 .L ' 1

Figure 12. Initial rate of change of spineward accel-

*eration endured by various subjects. (From

Reference 9)

body parallel in a to forward the long or bac~kward axis direction This is (G s~own ) than by when a comparison applieC T(G)n !

of the curves in Figures 1 nd13. A primary reason for the

significantly lower tolerance to headward (+G ) loading is the

susceptibility of the lumbar vertebrae, whic~h must 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 (Ga).

eration endured 46by v. (FZ.' 

"'A R e f e r e n c e 9 )

bodyin forardor bckwrd irecion(G )tha whe apli.

paale to th ogaxs( hi sA nb omaio
430

EA

-_ _ _ r4

___ 4J

4J)

cu

0

4)

- -. - - - 4 1

4) - _ _ _ *rd ( -1 1- . - - -~ 

IL -,4

- ___ r~u

___ 00

41i

(U m U-)

0 0oI

(0r ~ 4 Mr 0)0 i

0) 0 H -W 

~~i UJ~ ft~'q -H IfVBV~t 

)A 47- -0ý41

4-47.
These vertical loads, therefore, result in greater stress per unit area than do sternumward or spineward loads. Finally, k

along the direction of the long axis, the body configuration

allows for greater displacement of the viscera within the body

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 viAcera than do forces applied sternumward or spineward (G ), thereby increasing the susceptibility

of the viscera to injuries.

As in the case of the longitudinal direction (Figure 12), rate

of onset also affects tolerance to vertical accelerative loads;

however, insufficient data were available to establish the limits. (Figure 14 presents one set of available data.)

4.3.4 Tailward (-Gz) Acceleration

The human tolerance limit for tailward, eyeballs-up (-G ) acceleration is approximately 15 G for a duration of 0.1 sec.

The shoulder harness/lap belt restraint has been used in all

human testing with tailward accelerations. Most experiments

also have included a lap belt tiedown strap, and the 15 G tolerance limit is based on this latter configuration.

4.3.5 Lateral (G ) Acceleration

Very little research has been conducted on human tolerance to

lateral (G ) accelerations. Two studies, one involving restraint byya lap belt alone (Reference 13) and another involving restraint by the lap belt/shoulder harness configuration

(Reference 14), provide the principal available data. In both

cases, a side panel provided additional restraint. With restraint by the lap belt alone, volunteers were able to withstand a pulse with an average peak of approximately 9 G for a

duration of approximately 0.1 sec. At this level, the tests

were discontinued due to increasing concern about lateral spi- nal flexion. In the experiments with restraint by lap belt

and shoulder harness, volunteers were able to withstand a

pulst with an average acceleration of approximately 11.5 G for

13. zaborowski, A. V., HUMAN TOLERANCE TO LATERAL IMPACT WITH

LAP BELT ONLY, Proceedings, Eighth Stapp Car Crash and

Field Demonstration Conference, Society of Automotive

Engineers, New York, 1964.

14. Zaborowski, A. V., LATERAL IMPACT STUDIES - LAP BELT

SHOULDER HARNESS INVESTIGATIONS, Proceedings, Ninth Stapp

Car Crash Conference, Society of Automotive Engineers, New York, 1965.

48

.' .'k .• . .. .' ... l " •• t ,% ". 
41

c: 10

Cj 0 0 )

.,~~ - 0

C a)

Q-C) -r

to _ z~U-0J. 

0. ~ ZC.~ >

C cc ' *~J -- H

cCca -zl)TD wnwTcCcc -~jj~ JO 

0.- CL L~C0.

-~ 0.49 cd 1-. 
a duration of approximately 0.1 sec with no permanent physiological changes. Tests were discontinued at this level due to

possible cardiovascular involvement experienced by one of the

two subjects tested. No end points for human tolerance to lateral impacts were proposed in the reports of these experiments. The only reasonable conclusions from these data at this

time are that a pulse of 11.5 G with a duration of 0.1 sec is

readily sustained by subjects restrained by a lap belt and

shoulder harness and that the human survival limit is at some

point beyond this level, probably at least 20 G for 0.1 sec.

The above values are supported by a series of human volunteer

experiments conducted to measure the inertial response of the

head and neck to +G whole-body acceleration (Reference 15).

Acceleration inputs Yanged from long-duration pulses with magnitudes of 2 to 7.5 G to short-duration pulses of 5 to 11 G.

4.4 HEAD IMPACT TOLERANCE

As indicated by the accident data discussed in Chapter 3, over

30 percent of the Army aircraft crash fatalities result from

head injuries. The injuries may result either from impact of

the head on some aircraft structure or equipment, or from head

acceleration without impact. In the case of mechanical impact,

tolerance conditions often are based on the presence or aboence of skull fracture. However, concussion can result from

nonimpact motion of the head, whether in flexion or hyperextension (References 16 and 17). Concussion that may be nonfatal in itself can temporarily immobilize an individual and

reduce his chances of survival by subjecting him to postcrash

hazards such as fire or drowning.

Fatal head injury has been shown to result from such severe

brain damage as laceration of brain tissue or shear of the

15. Ewing, C. L., et al., DYNAMIC RESPONSE OF THE HUMAN HEAD

AND NECK TO +G IMPACT ACCELERATION, Proceedings, TwentyFirst Stapp Cay Crash Conference, Society o Automotive

Engineers, New York, 1977, pp. 547-586.

16. Hollister, N. R., et al., BIOPHYSICS OF CONCUSSION, Wright

Air Development Center; WADC Technical Report 58-192,

Wright-Patterson Air Force Base, Ohio, 1958.

17. Ommaya, A. L., anid Hirsch, A. E., TOLERANCES FOR CEREBRAL

CONCUSSION FROM HEAD IMPACT AND WHIPLASH IN PRIMATES,

Journal of Biomechanics, Vol. 4, 1971, pp. 13-21.

5U
brain stem (Reference 18). Head impact studies with anesthetized monkeys and dogs have been conducted to relate the severity and duration of concussion to intracranial pressure

change (Reference 19). Moderate to severe concussion effects

were observed in the range of 30 to 90 lb/in. 3 intracranial

pressure change concurrent with head impact.

According to an hypothesis developed by Holburn, shear stresses

induced by head rotation also can produce concussion (Reference 20). Kornhauser, in Reference 21, indicated a relationship between damaging velocity and damaging accelerations as

follows:

0* *(3)

where 0 - damaging rotational velocity, rad/sec

0 a damaging rotational acceleration, rad/seca

w natural frequency of rotation of brain, rad/sec

Ommaya, et al., developed scaling factors needed to predict concussion thresholds for man from data taken on subhuman primates

(Reference 22). This study showed that 1 can be represented by

18. Goldsmith, W., BIOMECHANICS OF HEAD INJURY, In Biomechanics-Its Foundation and Objectives, ed. by Fung, Y. C.,

Perrone, N., and Anliker, M., Prentice-Hall, Inc., Englewood Cliffs, New Jersey, 1972, pp. 585,-634.

19. Gurdjian, E. S., Roberts, V. L., and Thomas, L. MI.,

TOLERANCE CURVES OF ACCELERATION AND INTRACRANIAL PRESSURE AND PROTECTIVE INDEX IN EXPERIMENTAL HEAD INJURY,

Journal of Trauma, Vol. 6, 1966, pp. 600-604.

20. Holburn, A. H. S., NECHANICS OF BRAIN INJURIES, British

Medical Bulletin, Vol. 3(6) 1945, pp. 147-149.

21. Kornhauser, M., PREDICTION AND EVALUATION OF SENSITIVITY

TO TRANSIENT ACCELERATION, Journal of Applied Mechanics, Vol. 21, 1945, p. 371.

22. Ommaya, A. K., et al., SCALING OF EXPERIMENTAL DATA ON

CEREBRAL CONCUSSION IN SUB-HUMAN PRIMATES TO CONCUSSION

THRESHOLD FOR MAN, Proceedings, Eleventh Stapp Car Crash

Conference, Society of Automotive Engineers, New York,

October 10-11, 1967.

51

., ' "." , ,,,-
the equation

c

m"r" (4)

where m - mass of the brain, gm

c - an experimentally derived constant, gMt/ 3 rad/sec2

The investigators found c a 21,600 gm2/3 rad/sec2, and further

showed that the relationship of Equation (3) produced reasonable agreement between predictions and empirical data. Limiting values thus predicted to produce a 50-percent probability

of concussion in a man having a brain mass of 1300 gm are as

follows (Reference 23):

0 a 1800 rad/secA

O - 50 rad/sec

In general, assessing the probability of injury by observation

of oversimplified parameters, such as peak acceleration of the

imposed acceleration-time environment, usually is not constructive. The problem has been to define some form of parameter

that is indicative of the degree of severity of a particular

input excitation. Various indicators have been developed,

based on experiments, and several of these for the head are

presented and discussed in Sections 4.4.1 through 4.4.5.

4.4.1 Weighted Impulse Criterion (Severity Index)

It can be seen from human tolerance data presented previously

that high forces or accelerations can be tolerated for only

very short periods of time, while lower values of these quanti- ties can be tolerated for longer periods of time. This same

relationship for head injury in forehead impacts, which was

established on the basis of impact tests performed at Wayne

State University on animals and human cadavers, is illustrated

23. Ommaya, A. K., at al., COMPARATIVE TOLERANCE FOR CEREBRAL

CONCUSSION BY HEAD IMPACT AND WHIPLASH INJURY IN PRIMATES,

1970 International Automobile Safety Conference Compendium,

Society of Automotive Engineere, New York, 1970.

52
in Figure 15 (Reference 24). In these tests, longitudinal impacts of the subject's forehead against unyielding flat surfaces were conducted, and the acceleration-time history of the

specimen head was measured at a point on the skull diametrically opposite the point of impact. The curve shown in Figure 15 was based on the observation of linear skull fracture.

oI

600 !'

500 i.. . .'

4J 400

Dangerous to life

S300

U

. 200 --

2.100-

Not, dangerous to life -

0 2 4 6 8 10 12 30 100

Time duration of effective

acceleration, msec

Figure 15. Wayne State Tolerance Curve for the

human brain in forehead impacts

against plane, unyielding surfaces.

(From Reference 24)

4 Based on such data as that collected at Wayne State, Gadd suggested a weighted impulse criterion as an evaluator of injury potential (Reference 2t). This severity index i& defined as

24. Patrick, L. M., Lissner, ii. R., and Gurdjian, E. S.,

SURVIVAL BY DESIGN - HEAD PROTECTION, Proceedigs, Seventh

Stapp Car Crash Conference, Society of Automotive Engi- Sneers, New York, 1963.-

25. Gadd, C. W., USE OF A WEICHTED-IMPULSE CRITERION FOR ESTIMATING INJURY HAZARD, Proceedings, Tenth Stapp Car Crash

Conference, Society of Automotive Engineers, New York,

53

i3.
t

SI= andt

to (5)

where SI = severity index

a = acceleration as function of time

n = weighting factor greater than I

t = time

Based on data in References 9 and 24, the exponent n has been

determined to be 2.5 for head and facial impacts. For the

load, Reference 26 suggests tolerable SI levels of 1U00 for

distributed loading and 400 for localized loading. Severity

index values exceeding 600 produced concussion in load impacts

sustained by U. S. Army aircrewmen in aircraft accidents (Reference 24). A lower value of n also has been suggested for regions of softer tissue, which behave viscoelastically.

The severity index can be calculated by dividing the time base

of the acceleration time curve into sufficient segments to define the acceleration curve. The G value then read from the

curve for the center of the increment is raised to the 2.5

power, and the result is multiplied by the time increment.

The sum of all the values obtained gives the severity index.

A severity index sample calculation is shown in Figure 16,

which is taken from Reference 26.

4.4.2 J-Tolerance

Slattenschek, after noting different head deceleration curves

when different types of windshield glass were used, developed

a method of assessing multiple impact tolerance by a "Jtolerance" value (Reference 27). A second-order vibrational

26. SAE Information Report, HUMAN TOLERANCE TO IMPACT CONDITIONS AS RELATED TO MOTOR VEHICLE DESIGN - SAE J885a,

SAE Handbook, Part 2, Society of Automotive Engineers,

Warrendale, Pennsylvania, 1979, pp. 34.114-34.117.

27. Slattenschek, A., and Tauffkirchen, W., CRITICAL EVALUATION OF ASSESSMENT METHODS FOR HEAD IMPACT APPLIED IN APPRAISAL OF BRAIN INJURY HAZARD, IN PARTICULAR IN HEAD IMPACT ON WINDSHIELDS, Paper 700426, 1970 International

Automobile Safety Conference Compendium, Society of Automotive Engineers, New York, 1970, pp. 1084, 1112.

54

A 1.Oa
80

60

U•

0

(• 4 0I

00 2 4 6 8 10 12 14 16 18 20 22 24 26 28 30 32 34

Time, msec:

S~~~~Calculations Inreta

Time Incremental Time Sl •

Increment increment Midpolt G (Time x25

number (sec) G value

7 130 0,13

1 0.001 1.40

2 0.001 38 1,400

3 0.001 23 2,500 2.50

4 0.002 27 3,800 7.60

5 0.001 33 6,300 6.30

6 0.001 40 10,000 10.00

7 0.001 38 8,800 8.80

8 1).001 47 15,000 15.00

9 0.001 75 48,000 48.00

1i0 0.001 30 57,000 57.00

11 0.001 73 46,000 46.00

12 0.001 56 23,000 23.00

13 0.001 43 12,000 12.00

14 0.002 37 8,300 16.60

"15 0.002 33 6,200 12 . 40

16 0.001 27 3,800 3. 80 17 0.001 24 2,800 2.80

18 0.007 20 1,R00 12.60

19 0.001 17 1,200 1.20

20 0.002 10 330 0.66

2S0everity Index 287-97

Figure 16. Sample calculation of a Severity Index.

(From Reference 26•

455
model, based on the Wayne State Tolerance Curve, is used to

determine the tolerable amplitude of brain motion. The response of the simple, damped, spring-mass system shown in Figure 17 is given by

x + 2Dwk + w = -b(t) (6)

where x = relative displacement of mass

= relative velocity of mass

x = relative acceleration of mass

b = acceleration of system at point N (driving acceleration)

D = damping coefficient

S= angular frequency.

- P b

c

N

D

Figure 17. Damped, spring-mass system used

in computing J-tolerance. (From

Reference 27)

The angular frequency w is calculated from the Wayne State Tolerance curve:

W J\k/mi 635 rad/sec

where m = mass

k = spring constant

56 )
The right side of the differential equation, -b(t), corresponds

to its interference function, that is to say, to the excitation

caused by the head impact; the solution of the equation is the

amplitude

X f a [b(t)]

Based on the tolerable amplitude xtolr, the tolerance value J

is given by

J x Xma---x

x tolr (7)

where J 1 is the tolerable limit.

4.4.3 Effective Displacement Index

The Effective Displacement Index (EDI), reported by Brinn and

Staffeld, is derived from a mathematical spring-mass model

based on the work described in Section 4.4.2 (Reference 28).

The peak deflection of the model, in inches, is taken as the

index of damage. Using a natural frequency of 77 Hz and a

damping value of 70.7 percent, a tolerable EDI of 0.15 in. was

obtained by fitting the Wayne State Tolerance Curve.

It should be noted here that, as in the case of the Gadd Severity Index, a tolerable EDI is based on anterior-posterior impact only, due to the unavailability of data for other directions.

4.4.4 Strain Energy Considerations

Melvin and Evans considered the basic types of skull fracture

and investigated the effects of impactor size and shape, skull

geometry, and soft tissue (Reference 29).

28. Brinn, J and Staffeld, S. E., THE EFFECTIVE DISPLACEMENT

INDEX - AN ANALYSIS TECHNIQUE FOR CRASH IMPACTS OF ANTHROPOMETRIC DUMMIES, Proceedings, Fifteenth Stapp Car Conference, Society of Automotive Engineers, New York, 1971,

pp. 817-824.

29. Melvin, J. W., and Evans, F. G., A STRAIN ENERGY APPROACH

TO THE MECHANICS OF SKULL FRACTURE, Proceedings, Fifteenth

Stapp Car Crash Conference, Society of Automotive Engineers, New York, 1971, pp. 666-685.

57

". .. .
4.4.5 Comparison of Head Injury Predictors

Hodgson and Thomas investigated skull fractures in 40 cadavers that were dropped with their heads striking rigid, flat,

hemispherical, and cylindrical surfaces on the front, side,

and rear (Reference 30). The Severity Index and the Effective

Displacement Index were compared at fracture level for all frontal impacts, anu their average values at fracture were found

to agree closely with the critical values predicted by the authors of the methods. Results for the frontal flat-plate impacts are shown in Figure 18, where A,,irefers to calculation

of the indices from resultant acceler ion at the head center

of gravity. A - refers to calculation from. anterior-posterior

acceleration m•A3ured at the point on the skull most distant

from the impact site, the condition used in the derivation of

the Wayne State Tolerance Curve.

S250 SI EDI-in.

Test no. ACG AA-P ACG AA-P 0

'H 200 - 2 749 400 .14 ... S3 2020 1800 .18 .12

4 1280 792 .17 .1.3

150 6 565 390 .13 .11

3 29 1020 724 .16 .13 '

100o

u 50 6

0)

0 .005 .010 .015 .02 .02'5 .03 .035 .04 .045

Time, sec

Figure 18. Comparison of SI, EDI, and kinematics of

six frontal impacts producing linear

fracture. (From Reference 30)

I 30. Hodgson, V. R., and Thomas, L. M., COMPARISON OF HEAD ACCELERATION INJURY INDICES IN CADAVER SKULL FRACTURE, Proceedings, Fifteenth StappCar Crash Conference, Society

iof Automotive Engineers, New York, 1971, pp. 190-206.

"58
For the drop height range that produced linear fracture in the

cadaver's skull due to frontal impact against a rigid, flat plate, the SI for the Alderson 50th-percentile dummy head ac- celeration response was significantly higher than for the cadaver. The higher acceleration measured for the dummy head

would, of course, be expected in such an impact because of the

greater rigidity of its metal skull. The EDI, on the other

hand, was essentially the same for both cadaver and dummy.

Fan analyzed predictions of the Gadd Severity Index and the

Vienna Institute of Technology Brain Model (J-tolerance) and

concluded that improvements in both techniques could be made

(Reference 31). He concluded that the SI puts too much weight

on the acceleration but ignores the time factors of an impulse. Fan's revised approach to SI calculation irivolves a successive

approximation method where variable weighting factors are ap- plied to both acceleration and time. A revised brain model

also is presented, based on the Vienna model, with additional

information included on dynamic properties of the human skullbrain system. The revised brain model utilized an equivalent

viscous damping of 40 percent of critical damping and, on comparison with the Wayne State Tolerance Curve, yielded tolerable values of brain deformation, velocity, and angular velocity of 1.25 in., 135.5 in./sec, and 175 rad/sec, respectively.

The maximum deviation from the Wayne State curve was reported

to be within 5 percent.

4.5 NECK IMPACT TOLERANCE

Tolerance of the human neck to rotation, as experienced in

whiplash, and to localized impact loading has been investigated. Mertz and Patrick determined the moment about the occipital condyles (considered to be the center for rotation of

the head with respect to the neck) at the threshold of pain for volunteer subjects. On the basis of their investigations,

tolerable levels for neck flexion (forward rotation) and neck

extension (backward rotation) of a 50th-percentile adult male

are proposed in Reference 32.

31. Fan, W. R. S., INTERNAL HEAD INJURY ASSESSMENT, Proceedings, Fifteenth Stapp Car Crash Conference, Society of

Automotive Engineers, New York, 1971, pp. 645-665.

32. Mertz, H. J., and Patrick, L. M., STRENGTH AND RESPONSE OF

THE HUMAN NECK, Proceedings, Fifteenth Stapp Car Crash

Conference, Society of Automotive Engineers, New York,

1971, pp. 207-255.

59 -l • ' II
Gadd, Culver, and Nahum, using unembalmed, elderly cadavers,

investigated the relationship between rotation and resisting

moment in hyperextension and lateral flexion (Reference 33).

4.6 CHEST IMPACT TOLERANCE

An extensive research program on impact response of the human

thorax has been reported by Kroell, Schneider, and Nahum (Reference 34). A 6-in.-diameter rigid impactor of varying mass

and moving at a range of speeds was used to strike unembalmed,

seated cadavers. Deflection and force were measured as functions of time. Figure 19 (from Reference 34) shows the Abbreviated Injury Scale (AIS) (Reference 35) plotted against

normalized chest deflection (deflection divided by chest

anterior-posterior diameter). The least-squares fit shown has

an associated correlation coefficient of 0.772. However, as

seen in Figure 19, the restrained back data appears to follow

a different trend. Scaling of the relationships between force

and penetration has been reported in Reference 36.

The Effective Displacement Index (EDI), which was discussed in

Section 4.4.3 in reference to head injury, also has been applied to chest injury by Brinn and Staffeld (Reference 28).

In agreement with the discussion in the preceding paragraphs,

they point out that there is evidence that rib cage deflection

may be the governing criterion for chest injury and that the

measurement of this deflection might be the basis for a chest

33. Gadd, C. W., Culver, C. C., and Nahum, A. M., A STUDY OF

RESPONSES AND TOLERANCES OF THE NECK, Proceedings, Fifteenth Stapp Car Crash Conference, Society of Automotive

Engineers, New York, 1971, pp. 256-268.

34. Kroell, C. K., Schneider, D. C., and Nahum, A. N., IMPACT

TOLERANCE AND RESPONSE OF THE HUMAN THORAX II, Proceedings,

Eighteenth Stapp Car Crash Conference, Society of Automotive Engineers, New York, 1974, p. 383-457.

35. THE ABBREVIATED INJURY SCALE (AIS), Joint Committee of

the American Medical Association, American Association

for Automotive Medicine, and the Society of Automotive

Engineers; American Association for Automotive Medicine,

Morton Grove, Illinois, 1976 Revision.

36. Neathery, R. F., ANALYSIS OF CHEST IMPACT RESPONSE DATA

AND SCALED PERFORMANCE RECOMMENDATIONS, Proceedings,

Eighteenth Stapp Car Crash Conference, Society of Automotive Engineers, New Yori, 1974, pp. 459-493.

60

.... , - N, 'V,.,'m 4
Ul

0 o 4-$

r4 4)

00

0H0

00 4J

t0 w

0~0

4.)) +~.

14-I4

"41a

0) 4.)

41))

4J~f > 4 u~ "

Ho -H H Id *)),rt 

~~00

SAjtnýUT SIW

61
survivability index. Howe-ex, the acute distress experienced

by human volunteers subjected 'o whole-body deceleration has

been considered a justification for an independent hazard index based on acceleration. Using a natural frequency of 15 Hz

and a damping value of 25 percent, a maximum EDI of 2.2 in.

was obtained for voluntary human exposure. Based on experience with energy-absorbing steering columns, an EDI of 2.8 in.

was suggested as a test limit for current-design anthropomorphic dummies,

On the basis of 16 dives performed by an instrumented professional high diver onto a mattress, combined with the results

of ea:lier studies, a long-duration acceleration tolerancce

level of 60 G with a pulsu duration of 100 msec has been recommended for the thorax in the anterior-posterior direction

(Reference 37).

4.7 ABDOMINAL IMPACT TOLERANCE

Relative to other body regions, little information is available

concerning abdominal tolerance to blunt impact trauma. Snyder

has reported that one of the major reasons is a "marked disagreement between medical investigators both asto the frequency

with which various abdominal organs are involved, as well as

the significance of the trauma" (Reference 6). Although a number of experimental animals, mainly hogs, have been utilized

in abdominal impact studies, the large number of variables precludes either generalization or, certainly, extrapolation of

the results to humans. The large number of organs, their complexity, and their lack of symmetry make the exact location

and direction of the impact critical. Also, there are additional considerations, such as whether or not the bladder is

full, or whether the stomach is full or empty. Summaries of

research in abdominal impact may be found in References 38

and 39.

37. Mertz, H. J., and Gadd, C. W., THORACIC TOLERANCE TO

WHOLE-.BODY DECELERATION, Proceedings, Fifteenth Stapp

Car Crash Conference, Society of Automotive Engineers,

New York, 1971, pp. 135-157.

38. Frey, C. G., INJURIES TO THE THORAX AND ABDOMEN, Paper

700195, Human Anatomy, Impact Injuries, and Human Tolerances, Society of Automotive Engineers, New York, 1970,

pp.69-76.

39. Gogler, E., et al., BIOMECHANICAL EXPERIMENTS WITH ANIMALS ON ABDOMINAL TOLERANCE LEVELS, Proceedings, TwentyFirst Stapp Car Crash Conference, Society of Automotive

Engineers, New York, 1977, pp. 713-751.

62

€ . . ~...

S. . , . • . , I I i I I -- . ... i" '•
4.8 S:'INAL INJURY TOLERANCE

Damage to the vertebral column, particularly the upper lumbar

and lower thoracic regions, occurs frequently in +Gz impact, where the force is directed parallel to the spine. A ÷G component can be expected in the use of aircraft ejection leats

and in aircraft accidents, particularly those involving heli- 2

copters, where a vertical component of impact force is usually

present. A recent summary of research on ÷G2 impact exposure

limits is contained in Reference 40.

Various mathematical models have been developed for prediction

of spinal response to +G loading. An obvious injury mechanism is the inertial loadihg sustained by the vertebrae, resulting in compression fractures. Therefore, the earliest models

J• have been one-dimensional spring-mass systems that assume all

the load to be borne by the vertebral body. One such model

that has been used extensively in ejection seat evaluation is

discussed in Section 4.8.1. However, this simplified approach

cannot predict all types of spinal injury and cannot assess

the significance of spinal curvature. More comprehensive ap- proaches have included the flexural beam model of Soechting

(Reference 41) and the discrete parameter model of Orne and

Liu (Reference 42) that accounts for the effects of eccentric

loading as well as spinal curvature. Sections 4.8.2 and 4.8.3

describe two more recently developed models that appear prom- ising for use in assessment of spinal injury potential in aircraft crashes.

S4.8.1 Dynamic Response Index

The human response to short-duration accelerations applied in

the upward vertical direction parallel to the spine (+G ) has been modeled by a single lumped-mass, damped-spring sysiem as

40. Von Gierke, H. E., and Brinkley, J. W., IMPACT ACCELERATIONS, In Foundation of Space Biology and Medicine, Volume II, Book One, Scientific and Technical Information

?Office, Nationa Aeronautics and Space Administration,

Washington, D. C., 1975, pp. 214-246.

41. Soechting, J. F., RESPONSE OF THE HUMAN SPINAL COLUMN

TO LATERAL DECELERATION, Journal of Applied Mechanics,

Vol. 40, 1973, pp. 643-649.

42. Orne, David, and Liu, Y. K., A MATHEMATICAL MODEL OF SPINAL RESPONSE TO IMPACT, Journal of Biomechanics, Vol. 4,

1971, pp. 49-71.

.4, 63

•A",

',. S. ,• ... •;•..• . i
shown in Figure 20 (Reference 43). In this model, it has been

assumed that the total body mass that acts upon the vertebrae

to cause deformation can be represented by the single mass.

In use, the relationship

2 d6 n2•

dt + 2r d + W z (8)

dt 2 ,ii1 +n 6 u

is solved through the use of a qp,1r. The third term, which

includes the deformation of the spine, 6, divided by the gravitational acceleration, g, is referred to as the Dynamic Response Index (DRI). The model is used to predict the maximum

deformation of the spine and associated force within the vertebral column for various short-duration acceleration inputs.

The spring stiffness for the model was determined from tests

of human cadaver vertebral segments, damping ratios were determined from measurements of mechanical impedance of human subjects during vibration and impact.

tl m - nmass (lb-sec 2 /in.)

6 - iefkt,-ction (in.)

Sw damping ratio

m k - stiffness (lb/in.)

z - acceleration input

(in./sec2)

i *Dynmic Rspons Inde

43 igure2:.: in-nry n n;ajral R;fereqency of43)

MAN BODY, Frost Engineering Development Corp.s AMRL TechSnical Report 66-157, Aerospace Medical Research Lab,

Wright-Patterson Air Force Base, Ohio, November 1969,

AD 701383.

rl •64

* i ra -- i~ ~- --
An analytical effort was conducted to determine the degree of

correlation between the spinal injury (DRI) model and injuries

experienced in operational aircraft ejection seats (Reference

44). Figure 21 shows the relationship between operational acceleration environments and actual spinal injury rates. The

response of the model is expressed in DRI values. It can be

seen that the injury probability does vary with the DRI but

that the cadaver data show a higher probability of injury than

do the operational data. It would be expected that the intact,

living vextebral column imbedded in the torso would be stronger

than cadaver segments; consequently, this result might be predicted.

To establish acceptable ejection seat acceleration environments, the Air Force has adopted a system using a combination

of acceleration components and the DRI.

In Specification MIL-S-9479 (Refercnce 45), the acceleration

levels to be imposed on the seat occupant are controlled by

acceleration, time, and DRI as shown in the following relationship:

kR 2 2 + ( - ÷ <. 1. 0 (9)

DRLGXL GyL) 1

Here G, and G are measured acceleration magnitudes in the x

and y •irectigns and G L and G are the limit acceleration

parameters as read froA accoleEktion versus time curves included in the specification. DRI is the DRI value computed

from Equation (7) for the positive z direction. DRI is the

limit value of the DRI. The value of DRI is 18 unriss the

resultant acceleration vector is more thar& 5 conical degrees

off the z axis and aft of the plane of the seat back, in which

case, the value of DRI is 16. The computed value for the

left-hand term of Equatikon (9) may not exceed one.

The DRI is calculated from Equation (8) with model coefficients

for the positive spinal case (eyeballs down) defined for the

Tl4.Brinkley, J. W., and Shaffer, J. T., DYNAMIC SINULATION

TECHNIQUES FOR THE DESIGN OF ESCAPE SYSTEMS: CURRENT APPLICATIONS AND FUTURE AIR FORCE REQUIREMENTS, Aerospace

"Medical Research Lab; AMRL Technical Report 71-29-2,

Wright-Patterson Air Force Base, Ohio, December 1971,

AD 740439.

45. Military Specification, WIL-S-9479, SEAT SYSTEM, UPWARD

EJECTION, AIRCRAFT, GENERAL SPECIFICATION FOR, Department

of Defense, Washington, D. C., March 1971.

.I 65
50. Operational data

40 - Cadaver data /

30 / u /

U

20 /

10-

5- /

2- D/C

1r. 11

1r%. - B /

0.5- /

,0.2- /

.,-I

04

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

F 48

*Denotes rocket catapult

Figure 21. Probability of spinal injury estimated from

laboratory data compared to operational

experience. (From Peference 44)

mean age of the Air Force flying population (age 27.9 years).

The model coefficients are as follows:

W n w 52.9 rad/sec

= 0.224

"66

MW:"7;•• •w.•• : ,,... . ... ii
* The DRI has been shown to be effective in predicting spinal injury potential for +G acceleration environments in ejection

seats. However, it sh.ld be remembered that it is a simple model of a complex dynamic system and that the correlations

made are for ejection seat acceleration-time pulses that can

vary widely from crash pulses. In particular, the rate of onset can be an order of magnitude greater than for ejection seat

pulses. Also, the position of the spine at the time of impact

can have a significant influence on the susceptibility to vertebral damage. Therefore, a helicopter pilot leaning forward

in his seat might be expected to respond differently from an upright, well-restrained ejection seat occupant, and, thus,

have lower tolerance to impact.

4.8.2 Wayne State University Two-Dimensional Model

A two-dimensional spinal model that considers the details of

load transmission among individual vertebrae has been developed by King and Prasad (Reference 46). The model considers

the natural spinal curvatures and the effects of flexion and

eccentric inertial loading on the spine. Head and neck motions are simulated, and their effects on the forces and monments in the thoracic and lumbar spine can be studied for offaxis impacts in the midsagittal plane. The input acceleration

pulse can be an arbitrary function of time. The restraint and

support systems have been included to properly simulate a

seated vehicle occupant. The experimental data for validation

of the model were obtained from cadaveric runs with the spine in the erect and hyperexteyido(I mode so that the model incorporates the ability to simiilate both spinal configurations.

The following assumptions were made in the mathematical developnents

e The 24 vertebral bodies, the head, and the pelvis

are rigid bodies constrained to move in the midsagittal plane.

* Each rigid body has three degrees of freedom in the

midsagittal plane--two translational and one rotational.

* The intervertebral discs are massless, and deformation of the spine takes place at the discs.

46. King, A. I., and Prasad, P., AN EXPERIMENTALLY VALIDATED

MODEL OF THE SPINE, Journal of Applied Mechanics, Vol. 41,

No. 3, September 1974, pp. 546-550.

67
* The discs are replaced by a system of springs and

dampers--one spring and damper for axial forces and

another spring and damper arrangement for restoring

torques due to relative angular motion between adjacent vertebral bodies.

* The facets and laminae are springs connected to the

vertebral body by a massless rigid i.od.

e Each rigid body is assumed to carry a portion of the

torso woight that is eccentric with respect to the

center line of the spine.

e The rigid bodies are arranged to simulate the spinal

curvatures as closely as possible.

Equations of motion are derived for each vertebra, resulting in

a set of 78 second-order differential equations that are solved

numerically on a digital computer.

Experiments involving the use of human cadavers were carried

out for model validation. An acceleration input was applied

at the pelvis while the top of the head was allowed to remain

stress free. The parameter used for validation was the force

between two adjacent vertebrae, and an intervertebral load

cell was developed to provide the magnitude and line of action

of the force. Comparisons of model predictions and experimental data are shown in Figure 22 for two 10 G runs with the

spine in different positions, where the loads were measured

between the second and third lumbar vertebrae. Shown are both

the loadr between vertebral bodies (IVL) and those in the facets, w' h limit relative rotation of the vertebrae. The significance of the initial curvature of the spine videnced

by the difference in response betwqee' tha erect nyperextended (backward rotation of the torso) modes.

This model appears to be potentially useful in spinal injury

prediction, provided that dynamic fracture loads for vertebrae

are known, ae discussed in Reference 47.

4.8.3 Air Force Head-Spine Model

Under the sponsorship of the U. S. Air Force Aeromedical Research Laboratory, a three-dimensional, discrete model of the

human spine, torso, and head was developed for the purpose of

47. Hakim, N. S., and King, A. I., STATIC AND DYNAMIC ARTICULAR FACET LOADS, Proceedings, TwenLieth Stapp Car Crash

Conference, Society of Automotive Engineers, New York,

1976, pp. 607-639.

68
Time, msec

50 100 150 200 250

IVL IVL (experimental)

-500 (model)V

S-1000 ... 10 IVL = Intervertebral load

0 500 -

25-50 100.. 1laD-' 00 250 -250-0 -• " Facet load

Facet load- (experimental)

(model)

(a) Erect

Time, msec

50 100 150 200 250

o V' (mod_ l)X- _ IVL (experimental)

-500 __ . ...

S-1000

IVL = Intervertebral load

0 500l

250 50 100 150 200 250

-500. Fae od (experimental)

Facet load (model)

(b) Hyperextended

Figure 22. Comparison of model output and experimental

data for i0 G runs with the spine in the

1a) erect and (b) hyperextended modes.

(From Reference 46)

evaluating mechanical response in pilot ejection. It was developed in sufficient generality to be applicable to other body

response problems, such as occupant response in ai'rcraft crash

and head-spinal system response to arbitrary loads. There are

69
no restrictions on the distribution of direction of applied

loads, therefore, a wide variety of situations can be treated.

The model has been described in Reference 48.

The anatomy is modeled by a collection of rigid bodies that

represent skeletal segments, such as the vertebrae, pelvis,

head, and ribs, interconnected by deformable elements that

represent, ligaments, cartilaginous joints, viscera, and connective tissues. Techniques for representing other aspects

of the ejection environment, such as harnesses and the seat

geometry, are included also. The model is valid for large

displacements of the spine and treats material nonlinearities.

The elements of the model are illustrated in Figure 23.

The basic model is modular in format so that various components can be omitted or replaced by simplified representations.

Thus, while the complete model is rather complex and involves

eubstantial computational effort, various simplified models,

which aru quite effective in duplicating the response of the

complete model within a range of conditions, are available.

Three methods of solution are available for the analysis: direct integration in time either by an explicit, central difference method or by an implicit, trapezoidal method, and a frequency analysis method.

A variety of coniditions have been simulated, including different rates of onset, ejection at angles, effects of lumbar curvature, and eccentric head loadings. It has been shown that

large initial curvatures and perfectly vertical acceleration

loadings, which cause large bending moments, result in substantial flexural response of the spine. It has been further shown

that the combination of the spine's low flexural stiffness,

initial curvature, and mass eccentricity are such that stability cannot be maintained in a 10 G ejection without restraints

or spine-torso-musculature interaction..

The complete models were used mainly to study the effects of

the rib cage and viscera on spinal response. The flexural

stiffness of the torso is increased substantially by a visceral

model, even though it has no inherent flexural stiffness. In

addition, the viscera provide significant reductions in the

axial loads.

48. Belytschko, T., Schwer, L., and Schultz, A., A MODEL FOR

ANALYTIC INVESTIGATION OF THREE-DIMENSIONAL HEAD-SPINE

DYNAMICS - FINAL REPORT, University of Illinois at Chicago

Circle; AMRL Technical Report 76-10, Aerospace Medical

Research Lab, Wright-Patterson Air Force Base,

Ohio, April 1976, AD A025911.

'70

-A& A
Figure 23. Three-dimensional head-spine model. (From

Reference 48)

4.8.4 Vertebral Properties

Accurate strenoth and deflection properties of the seated human

torso when exj,,aed to inertial loads are urgently needed.

Reference 49 includes a consolidation of the available

49. MODELS AND ANALOGUES FOR -,HE EVALUATION OF HUMAN BIODYNAMIC RESPONSE, PERFORMANCE, AND PROTECTION, AGARD Conference Proceedings No. 253, NATO Advisory Group ?or Aerospace Research and Development, Neuilly sur Seine, France,

1968.

71

,. ....,• . ... '..' " ... I!.
data from King, Kazarian, and Hodgson (References 46, 50, and

51) for the head and spinal column exposed to +G loading.

They present a set of recommended properties that Eay find use

in a more sophisticated model for spinal injury than the DRI.

4.9 LEG INJURY TOLERANCE

Femoral fracture due to longitudinal impact on the knee has

been studied extensively, probably because of the frequency of

this type of injury in automobile accidents.

Based on cadaver data obtained by Patrick, et al. (Reference

52), King, et al., recommended a peak fracture load of 1700 lb

as a realistib criterion (Reference 53). This value has since

been adopted for use in Federal Motor Vehicle Safety Standard

208. Recent experiments reported by Powell, et al., point

to this value as being conservative for impacts of less than

20 msec (Reference 54).

Viano has presented a criterion that assesses the dependence of

the permissible human knee load on the duration of the primary

50. Kazarian, L., and Graves, G. A., COMPRESSIVE STRENGTH

CHARACTERISTICS OF THE HUMAN VERTEBRAL CENTRUM, AMRL Technical Report 77-14, Aerospace Medical Research Laboratory,

Wright-Patterson Air Force Base, Ohio, 1977.

51. Hodgson, V. R., and Thomas, L. M., HEAD IMPACT RESPONSE -

PROPOSAL NUMBER VRI 7.2, Society of Automotive Engineers,

Inc., New York, 1975.

52. Patrick, L. M., Kroell, C. K., and Mertz, H. G., FORCES O0

THE HUMAN BODY IN SIMULATED CRASHES, Proceedings, Ninth Stapp Car Crash Conference, Society of Automotive Engi-

- I neers, New York, 1965, pp. 237-259.

* 53. King, U. J., Fan, W. R. S., and Vargovick, R. J., FEMUR

... .LOAD INJURY CRITERIA - A REALISTIC APPROACH, Proceedings,

Seventeenth Stapp Car Crash Conference, Society of Automotive Engineers, New York, 1973, pp. 509-524.

54. Powell, W. R., at al., INVESTIGATION OF FEMUR RESPONSE TO

LONGITUDINAL IMPACT, Proceedings, Eighteenth Stapp Car

Crash Conference, Society of Automotive Engineers, New

York, 1974, pp. 539-556.

72

e-

, Ft 
force exposure (Reference 55). Based on the knee impact data

from previous experiments with both fresh and embalmed cadavers, Viano suggests the following femur injury criterion (FIC)

to define a permissible peak knee load:

F(kN) = 23.14 - 0.71 T(msec), T < 20 msec

F(kN) * 8.90, T > 20 msec (10)

or, in English units,

F(lb) - 5200 -160 T(msec), T < 20 msec

F(Ib\._ 20',U, T > 20 msec (10a)

The relationship of Equati6t.o44A is illustrated in Figure 24.

4.10 ABBR4VIATED INJURY SCALt(

The Abbreviated Injury Scale (0IS), first reported in 1971,

was developed as a comprehensive system for rating tissue damage that would be acceptable to physicians, engineerg, and researchers working in automotive crash investigation. An engineer concerned with interpretation or use of crash injury data

can find information on the AIS and it,- use in References 35

and 56.

55. Viano, C. C., CONSIDERATIONS FOR A FEMUR INJURY CRITERION, PROCEEDINGS, Twenty-First Stapp Car Crash Conference,

Society of Automotive Engineers, New York, 1977, pp. 445-

473.

56. States, M. M., et al., FIELD APPLICATION AND DEVELOPMENT

OF THE ABBREVIATED INJURY SCALE, Proceedings, Fifteenth

Stapp Car Crash Conference, Society of Automotive Engineers, New York, 1971, pp. 710-738.

7 3
24.0

22.0- Axial knee impact: peak force'

20.0- femur tolerance as dependent

upon Eulse duration

18.0-F=23.14- 0.71 T, T<20 msec

S1.0F-8. 90 T > 20 insec

16.0

014.0

.)12.0

-W 10. 0

ý4 8.0 -Knee force

24.0 :eu~ +:±: 2 1 5

740 -

2- .........
5. OCCUPANT MOTION ENVELOPES

5.1 INTRODUCTION

The purpose of this chapter is to acquaint the reader with the

extent of an aircraft occupant's mot:Lon in a crash environment.

This knowledge is vital in designing for occupant protection

from injury due to impact with the aircraft interior, a topic

discussed in Volume IV.

The kinematics of body action associated with aircraft crash

impacts are quite violent, even in accidents of moderate severity. The flailing of body parts is much more pronounced when

the aircraft occupant ic restrained in a seat with a lap belt

only. lHowever, even with a lap belt and a shoulder harness that are drawn up tightly, multidirectional flailing of the head, arms, and legs, and to a lesser extent, the lateral displacement of the upper torso within its restraint harnessing,

is extensive. If it were possible to provide adequate space

within the occupant's immediate environment, this flailing action of a fully restrained ocnupant would not be a particular

problem. Since space for occupants is usually at a premium in

aircraft, especially in cockpit areas, it is not feasible to

remove structural parts of the aircraft sufficiently to keep

the occupant from striking them. The only alternative is to

design the occupant's immediate environment so that, when the

body parts do flail and contact rigid and semirigid structures,

injury potential is minimized.

An occupant who is even momentarily debilitated by having his

head strike a sharp, unyielding structural object or by a leg

injury can easily be prevented from rapidly evacuating the aircraft and may not survive a postcrash fire or a water landing.

The imortance of occupant environment designed for injury prevention, therefore, should be emphasized if crash protection

is to be ensured.

Several approaches are available to alleviate potential secondary impact problems. The most direct, approach, which should

be taken if practical, is to relocate the hazardous structure

or object out of the occupant's reach. Such action is normally

subject to tradeoffs between safety and operational or human

engineering considerations. If relocation is not a viable alternative, the hazard might be reduced by mounting the offending structure on frangible or energy-absorbing supports and

applying a padding material to distribute the contact force

over a larger area. Application of protective padding for

both energy absorption and load distribution is discussed in

Volume IV.

75 1'

. ,.I ... 
5.2 FULL RESTRAINT

Body extremity strike envelopes are presented in Figures 25

through 27 for a 95th-percentile Army aviator wearing a restraint system that meets the requirements of MIL-S-58095(AV)

(Reference 8). The restraint system consists of a lap belt,

lap belt tiedown strap, and two shoulder straps,. The forward

motion shown in Figures 25 and 26 was obtained from a test

utilizing a 95th-percentile anthropomorphic dummy subjected to

a spineward (-G ) acceleration of 30 G. The lateral motion is

based on expect'd restraint system deflections in a 30 G lateral environment.

5.3 LAP-BELT-ONLY RESTRAINT

Although upper torso restraint is required in new Army aircraft, strike envelopes for a 95th-percentile aviator wearing

lap-belt-only restraint are presented for possible use in Figures 28 through 30. They are based on 4 G accelerations and

4 in. of torso movement away from the seat laterally and forward.

76
10 in.

_________ ... Heel rest line for cockpit.

I Aircrait, foor liefor

troop compartment.

Figure 25. Full-restraint extremity strike envelopeside view.

I 77
Ii,

Figure 26. Full-restraint extremiity strike envelope

top view.
7- . -

10 in/

____ ____ he rest line for cocýEit.

A cra1ft foor l1-hW=6r

troop compartment.

* Figure 27. Full-restraint extremity strike envelopefront view.

79

p * *'7

A; t-
______ Heel rest line for cockpit.I

Aircralt foo line for

troop compartment.

Figure 28. Lap-belt-only extremity strike envelope-I

80
000

000 .00

ii f-4

81

::-77 7
04.

-4-

04

0)

.................................... .................
6. HUMAN BODY DIMENSIONS AND MASS DISTRIBUTION

6.1 INTRODUCTION

This chapter presents information on the dimensions and properties of the human body. Anthropometric measurements are external dimensions of the human body that can be used to define

aircraft requirements such as seat height and width, eye height,

or cabin height. A specialized type of anthropometric measure.

ment is the "link length," or distance between joint centers,

which can be used in locating control positions and is essential for the design of mathematical or physical simulators of

the human body. Finally, the inertial properties of the body

and parts of the body also are required in the design of human

simulators.

6.2 ANTHROPOMETRY

Anthropometry is a specialized area of physical anthropology

that is concerned with the measurement of the human body and

its parts. Two types of anthropometric measurements have been

recorded, and the use of both types in vehicle design has been

summarized in Reference 57. Conventional dimensions of the

body obtained with subjects in rigid, standardized positions

are easily obtained. Extensive collections of such data are

used in clothing design and may determine certain vehicle design parameters including seat height and eye height. A second class of data, which may be referred to as workspace dimensions, is more difficult to obtain and can be applied only to

the specific workspace studied. However, these workspace dimensions are essential in designing aircraft interiors for maximum occupant protection.

6.2.1 Conventional Anthropometric Measurements

Conventional anthropometric measurements of greatest interest

in aircraft interior design include those dimensions illustrated in Figure 31, as well as standing height and body weight.

The most recent anthropometric survey of U. S. Army aviators is

contained in Reference 58, and the dimensions of greatest potential usefulness aLe presented in Table 4. Corresponding

57. Roe, R. W., and Kyropoulos, P., THE APPLICATION OF ANTHROPOMETRY TO AUTOMOTIVE DESIGN, Paper 700553, Society of

Automotive Engineers, New York, 1970.

58. Churchill, E., et al., ANTHROPOMETRY OF U. S. ARMY AVIATORS - 1970, Anthropology Research Project, Yellow

Springs, Ohio; Technical Report 72-52-CE, U. S. Army Natick Laboratories, Natick, Massachusetts, December 1971,

AD 743528.

83

~|
Sitting height SEye hih

Functional

Sreach Shoulder breadth

Shoulder-elbow length

Elbow-fingertip'lnt lengt

Bui-l-nee enth ht

Popliteal height Hip breadth

Buttock-popliteal

Figure 31. Conventional seated anthropometric dimensions.

dimensions for nonaviators, taken from Reference 59, are listed

in Table 5. These dimensions are nude measurementsj the dimensions of bulky clothing and helmets must be considered for specific applications.

Because anthropometric surveys involve a large population sample, they generally follow a normal (bell-shaped) distribution.

A normal distribution, as shown in Figure 32, can be described

in terms of its mean or average value and its dispersion about

Sthe mean, often expressed as standard deviation.

The percentile value, which corresponds to a rank order, is a

useful statistic for designers. If a group of subjects were

ordered from least to greatest for any given measurement, such

as standing height, tLie first percentile would be that part 4 'exceeded by 99 percent of the group; the 5th percentile would

be that exceeded by 95 percent of the group. The 50th percentile, or median, would be that half of the group exceeded by the other half. For a normally distributed sample, the median

value is the same as the mean, or average.

59. THE BODY SIZE OF SOLDIERS - U. S. ARMY ANTHROPOMETRY -

1966, USANL Technical Report 72-51-CE, U. S. Army Natick

Laboratories, Natick, Massachusetts, December 1971, AD 743465.

84

• •, ? ,
TABLE 4. SUMMARY OF ANTHROPOMETRIC DATA FOR

U. S. ARMY AVIATORS (REFERENCE 58)

Percentiles (in.)

Measurement 5th 50th 95th

Weight (ib) 133.0 171.0 212.0

Stature 64.6 68.7 72.8

Seated height 33.7 35.8 37.9

Shoulder breadth 17.0 18.7 20.3

Functional reach 28.8 31.1 34.2

Hip breadth, sitting 13.2 14.8 16.7

Eye height, sitting 29.0 31.0 33.1

Knee height, sitting 19.3 20.8 22.6

Popliteal height 15.1 16.6 18.3

Shoulder-elbow length 13.3 14.4 15.6

Elbow-fingertip length 17.6 19.0 20.3

Buttock-popliteal length 17.7 19.3 21.0

Buttock-knee length 22.0 23.7 25.4

For a normal distribution, as shown in Figure 32, 68 percent

of the sample is included within plus or minus one standard

deviation of the mean, and 95 percent within plus or minus two

standard deviations.

An example of the use of the statistics of anthropometric data

in setting design limits for a vehicle dimension (seat height

adjustment range) was presented in Reference 57.

6.2.2 Body Joints and Ranges of Motion

Few body joints involve rotation about fixed axes or pivot

points. Rather, the instantaneous center of rotation may depend on position, as illustrated in Figure 33 for the shoulder

85

S.

R '
TABLE 5. SUMMARY OF ANTHROPOMETRIC DATA FOR

SOLDIERS (REFERENCE 59)

Percentiles (in.)

Measurement 5th 50th 95th

Weight (ib) 126.0 156.0 202.0

Stature 64.5 68.7 73.1

Seated height 33.3 35.7 38.1

Shoulder breadth 16.3 17.8 19.6

Hip breadth, sitting 11.9 13.0 14.5

Eye height, sitting 28.6 31.0 33.3

Knee height, sitting 19.6 21.3 23.1

Popliteal height 16.0 17.5 19.2

Shoulder-elbow length 13.3 14.5 15.7

Elbow-fingertip length 17.4 18.8 .20.4

Buttock-popliteal length 18.0 19.6 21.3

Buttock-knee length 21.6 23.4 25.3

a=Standard 1 .. deviation

44

0

C, 0

2.5 16 50 84 97.5

Percentile

Figure 32. Normal distribution curve.

86
Figure 33. Path of instantaneous center of rotation during shoulder abduction. (From

Reference 60)

joint. Dempster reported on an extensive study of workspace

requirements for seated operators, in which he determined

"link lengths" between effective joint centers for major body

parts (References 60 and 61). These link lengths have a number of crashworthiness-related applications: first, in developing or expanding the strike envelopes shown in Chapter 5;

second, in designing crash test dummies; and third, in providing numbers for mathematical simula-.ors. From the data of several investigations, the skeletal joint locations for a 50thpercentile Army aviator, illustrated in Figure 34, were developed (Reference 49).

Joint ranges of motion are required in the same areas of application listed above the link lengths. These movements, illustrated in Figure 35, are measured from a standard anatumical

position defined as an erect standing posture with the palm

surfaces of the hands positioned anteriorly. Various studies

have determined the ranges of motion that may be attained

voluntarily and under external force; Cable 6 lists angles

60. Dempster, W. T., SPACE REQUIREMENTS FOR THE SEATED OPERATOR, Wright Air Development Center; WADC Technical Report

55-159, Wright-Patterson Air Force Base, Ohio, 1955,

AO 087892.

61. Dempster, W. T., and Gaughran, G. R. L., PROPERTIES OF

BODY SEGMENTS BAAD ON SIZE AND WEIGHT, American Journal

of Anatomy, Vol. 120, 1967, pp. 33-54. ,g

87

' •i '. %" •, " ' ... 44 •
F!

IJ

Top of head (crown)I

90 RL 174.6*

80 Skull, base and

C1 intersection RT 160.3

C2-C3 hinge RL 157.1

70

C7 T hinge RLRL 148.7

.,LI • 60 '• RL ~141.6 RL R 142.4 4. '

-W 50RL 135.1

540

RL 118.7

30RL 112.8 RL 111.8

i 20 • RL 1 1 04. S___________-R 101. 6RRL 10410.9" 0.

RL 96 1

10 RL 91.2

10 RL• - 83.. 6, . . . ...

0 10 20 30 40 5 0"a,0

Horizontal scale, cm

*.*-Horizontal reference line

in centimeters

RL 50.2-

RL 42. *

Figure 34. Sitting skeletal joint locations based on a 50thpercentile Army aviator. (From Reference 49)

88
C C- 4 '," IG

,E

)P

S N

Figure 35. Joint ranges of motion.

TABLE 6. RANGE OF JOINT ROTATION (REFERENCE 62)

d Measured

Body rotation (deg) component,,

motion Symbol Motion description Voluntary Forced

Head - with A Dorsiflexion 61 77

respect to B Ventriflexion 60 76

torso C Lateral flexion 41 63

D Rotation 78 83

Upper arm - E Abduction (coronal 130 137

at shoulder plane)

F Flexion 180 185

G Hyperextension 58 69

Forearm - at H Flexion 141 146

elbow

Thigh - at I Flexion 102 112

hip J Hyperextens. on 45 54

K Medial rotation - -

L Lateral rotation - -

M Adduction - -

N Abduction 71 79

Lower leg P Flexion 125 138

at knee

89

I [ I ] I" I I-• I I l ' I I /
obtained by Glanville and Kreezer (Reference 62) for the movements defined in Figure 35.

6.3 INERTIAL PROPERTIES

Inertial properties of the human body have been used in design

of escape systems, and moments of inertia of live subjects in

a seated position were determined by Santischi, DuBois, and

Omoto iReference 63). However, anthropomorphic dummies and

mathematical simulations require inertial properties of body

segments, specifically moments of inertia, mass, and centerof-mass locations. Dempster reported on a study in which center-of-mass locations and moments of inertia with respect to

transverse (y) axes were measured on segmented cadavers (Referencres 60 and 61). Clauser, McConville, and Young determined

center-of-mass locations and developed regression equations for cadaver segments (Reference 64). rrom data of various

sources, the study presented in Reference 49 determined the segment masses and center-of-mass locations presented in Table 7

and Figure 36 for a 50th-percentile Army aviator. Chandler,

et al., measured moments of inertia with respect to six axes

for fourteen segments of six cadavers and, from them, calculated the principal moments of inertia for the segments (Reference 65). Their results are presented in Table 8.

62. Glanville, A. D., and Kreezer, G., THE MAXIMUM AMPLITUDE

AND VELOCITY OF JOINT MOVEMENTS IN NORMAL MALE HUMAN

ADULTS, Human Diology, Vol. 9, 1937, pp. 197-211.

63. Santischi, W. R., DuBois, J., and Omoto, C., MOMENTS OF

INERTIA AND CENTERS OF GRAVITY OF THE LIVING HUMAN BODY,

AMRL Technical Data Report 63-36, Aerospace Medical Research Lab, Wright-Patterson Air Force Base, Ohio, 1963.

64. Clauser, C. E., McConville, J. T., and Young, J. W.,

WEIGHT, VOLUME, AND CENTER OF MASS OF SEGMENTS OF THE

HUMAN BODY, Antioch College; AMRL Technical Report 69-70,

Aerospace Medical Research Lab, Wright-Patterscn Air Force

Base, Ohio, August 1969, AD 710622.

65. Chandler, R. F., et al., INVESTIGATION OF INERTIAL PROPERTIES OF THE HUMAN BODY, Report No. DOT-HS-801-430, U. S.

Department of Transportation, Washington, D. C., March

1975.

90
TABLE 7. CENTER-OF-MASS DISTRIBUTION OF SEATED TORSO - 50TH-PERCENTILE ARMY AVIATOR

(REFERENCE 49)

Segment Z-axis X-axis

mass location location

Body sement identity. (kg) (cm)* (cm)**

Head 4.74 77.6 10.1

Neck (CI-C7) 1.63 71.3 9.7

Upper thoracic (Tl-T3) 4.07 62.4 10.0

Upper mid thoracic (T4-T6) 4.07 55.6 9.7

Lower mid thoracic (T7-T9) 4.66 48.1 11.2

Upper arm 4.44 43.5 12.2

Lower thoracic (TIO-T12) 5.29 40.2 13.0

Lumbar (LI and L2) 4.48 31.8 13.2

Lumbar (L3 and L4) 4.87 24.5 13.1

Forearm 2.62 24.6 23.7

Lumbar (L5) 2.52 19.0 12.2

Hand 0.92 15.3 45.1

Pelvis 8.89 13.0 11.2

Thigh (hip) 15.83 7.6 27.2

Lower leg 6.38 -8.6 55.0

Foot 1.99 -37.1 59.0

TOTAL 77.40

*Location is based on floor level of zero with 50th-percentile male head crown equal to reference line of 174.6 cm.

**Location is based on seat back with reference line equal to

zero. Seat back is perpendicular to seat bottom, and torso

touchies seat barl; at head, shoulders, and buttocks.

91

r . , . ,. 4 "I'•" .":r 4
90

(4.74 kg)

I-kad

80 C

(1.63 kg)

70 Neck

CG

60 (4.07 kg)

(4.07 kg)

50-

50 (4.66 kg)

-Upr. arm CG (4.44 kg),

attached to TI-T6

4J (5.29 kg)

30 (4.48 kg)

SForearm CG (2.62 kg) (4.87 kg)- -

20, " "

2.52 kg'ý

Approx. hip cut and CG (0.92 kg)

Pel~vis CG- -lip

through pelvis • thro ugh l

(0.89 kg) CG•

0 0 20 30 40 0'6

Horizontal scale, cm -Calf ~(6.38 CGkg)

,.. I .•-•" ,-Foot [•, ~Approx. out through _ lower calf •

Figure 36. Mass distribution of seated torso referenced to

the skeletal structure for a 50th-percentile

Army aviator. (From Reference 49)

' 92 (9
TABLE 8. SEGMENT MOMENTS OF INERTIA ABOUT

THE CENTER OF MASS (REFERENCE 65)

Moment of inertia (103 gmrcm2 )* Body

segment xx zz

Head 174.00 164.40 202.90

Torso 16,194.00 10,876.00 3,785.00

Upper arm 143.60 135.20 22.00

Forearm 65.80 63.80 8.70

Hand 7.19 5.86 1.97

Thigh 1,144.00 1,190.00 218.70

Calf 393.10 391.20 28.90

Foot 32.62 30.76 7.29

*Mean values of stature and weight reported to be

172.2 cm and 69.6 kg, respectively, for sample of

six cadavers.

6.4 SCALING OF MEASUREMENTS

References 58 and 59 contain a significant volume of anthropometric data, whose statistics have been completely analyzed.

In other words, the mean, standard deviation, and percentiles

are listed for all measurements. However, the link lengths

presented in Section 6.2.2 and the inertial properties presented in Section 6.3 are based on rather small samples. If

the user of this guide wishes to scale the dimensions of Figures 34 or 36 to an occupant size other than the 50th-percentile Army aviator, it is recommended that the scaling be based on

the most similar anthropometric dimension. For example, the

lower leg length shown in Figure 34 is 41.0 cm. In order to

convert this dimension to a 95th-percentile value, it would be

multiplied by the ratio of popliteal heights from Table 4.

The 5th-percentile lower leg length is then calculated as

V.

q3
L95 L 50, PH95

- 45.2 cm

The center-of-mass location then can be determined for the new

length using the same percentage of total segment length as

shown in Figure 34. For example, according to Figure 34, the

center of mass of the lower leg is approximately 38.8 percent

of the distance down from the knee joint to the ankle joint.

Retaining this percentage for the 95th-percentile occupant,

the distance down from the knee joint to the lower leg center

of mass is 0.388 (45.2) a 17.5 cm. Scaling segment masses presents a greater problem as there is no complete set of segment

data for all size occupants. The simplest scaling approach

would be to multiply the masses of Figure 36 by the ratio of

total body masses (or weights) from References 58 or 59. For

a more rigorous approach, the regression equations of Reference 64 might be used.

94
7. CRASH TEST DUMMIES

7.1 INTRODUCTION

The technology of crash test dummies has advanced significantly

since 1968, when the first standard dummy was defined. Several

designs currently are available, and many one-of-a-kind systems have been developed by various laboratories for their own

use. However, for use in aircraft system evaluation, special consideration must be given to the effects of the vertical component of impact force, which make the aircraft crash environment quite different from that of an automobile, for which most

dummies have been developed.

This chapter briefly outlines the evolution of current dummy technology, indicates the design features that are desirable

for aircraft system testing, and summarizes research and comparative performance of dummies and humans.

7.2 DUMMY TECHNOLOGY

7.2.1 History of Dummy Development

An early dummy design by one of today's principal manufacturers

was a rugged ejection seat dummy built by Sierra Engineering

Company for the Air Force in 1949. According to Reference 66,

this dummy had limited articulation and poor biomechanical fidelity, but it filled an important need, not only for aircraft

system manufacturers but for the automobile industry as well.

A significant step toward the present anthropomorphic dummies

was made by Swearingen, who, in 1949, needed a dummy better

than the rigid test articles then available in order to evaluate explosive decompression for an aircraft cabin as a result

of window failure. He designed a 120-lb dummy with articulated

principal joints, realistic distributiov of body weight, and

centers of gravity approximating the human body. More than 4 500 blast tests were made to determine the hazard of explosive

decompression and to provide means for preventing or reducing

passenger injury. Placed in hundreds of positions, the dummy

was blown against, into, and through simulated aircraft cabin

windows. Most, of these would have been fatal to a human and

each test resulted in rather gross disintegration of the dummy.

6-. LeFevre, R. L., and Silver, J. N., DUMMIES - THEIR FEATURES AND USE, Proceedings, Automotive Safety Engineering

Seminar, Society of Automotive Engineers, New York, June

20-21, 1973.

95

ke I 
In 1951, Swearingen completed an improved dummy, capable of

withstanding 35 to 50 G, which was used in evaluating a new

safety harness for general aviation (Reference 67).

In 1954, Alderson Research Laboratories, Inc. created the first

mass production dummy, unique for its modular design. The design permitted new parts to be added as needs changed and as

knowledge grew over the subsequent decade. In 1967, both of

the major dummy manufacturers marketed new devices that featured increased articulation in the vertebral column and shoulders, as well as increased chest compliance. These changes

effected some improvement in biomechanical response but still

fell far short of what is available today.

In 1968, SAE Recommended Practice J963 was published as a partial definition of a standard 50th-percentile male anthropo- metric test device (Reference 68). J963 recommends weights,

center-of-gravity locations, dimensions for body Segments, and

the ranges of motion for body joints. Although moments of inertia and many design details were left unspecified, this was

a first step toward a standard test device. Alderson upgraded

its design to meet J963 in 1968 and 1971, while the Sierra

counterpart appeared in 1970.

7.2.2 Part 572 Dummy

The role of the anthropomorphic dummy in automobile safety

testing was formally changed in 1971 by the National Highway

Traffic Safety Administration. Prior to that time, dummies

had been used for determining relative performance of similar

safety systems. The new law carried the implications that dummies must determine the absolute potential for injury to human

occupants in an automobile crash and that different testing organizations should obtain the same results. The transition

from relative to absolute measuring instrument forced the requirement for the dummy to be a standardized test instrument,

as well as a reasonable simulation of a human being, since the

legal performance limits are based on human tolerance data.

67. Swearingen, J. J., DESIGN AND CONSTRUCTION OF A CRASH TEST

DUMMY FOR TESTING SHOULDER HARNESS AND SAFETY BELTS, Civil

Aeronautics Administration, Civil Aeronautics Medical Research Laboratory, (now FAA Civil Aeromedical Institute),

Oklahoma City, Oklahoma, April 1951.

68. SAE Recommended Practice, ANTHROPOMORPHIC TEST DEVICE FOR

USE IN DYNAMIC TESTING OF MOTOR VEHICLES, SAE J963, SAE

Handbook, Society of Automotive Engineers, Warrendale,

Pennsylvania, 1978, pp. 34.107-34.110.

96
In 1972, General Motors Corporation produced the Hybrid II

dummy, a 50th-percentile male anthropomorphic test device.

This dummy utilizes torso and limbs from the Alderson VIP-50A

dummy with modifications made to the chest to allow increased

deflection and damping. The head assembly was adapted from

the Siýrra 292-1050 design with several anatomical modifications. Both the neck and lumbar spine consist of a butyl rubber cylinder, the latter being reinforced by an internal steel

cable.

Along with a number of other modifications, the design of the

Hybrid II formed the basis for the Code of Federal Regulations,

Title 49 (49CFR) Part 572 specification for dummies (Reference

69). Its specified dimensions and inertial properties are displayed in Figure 37 and Tables 9 through 11. Segment moments

of inertia reported in Reference 70 for a Hybrid II dummy are

listed in Table 12.

7.2.3 Other Recent Dummy Designs

Various organizations involved in safety research and engineering have produced their own dummy designs through modification

of production devices, as was done in the Hybrid II development

discussed in Section 7.2.2. These dummies have generally been

designed to yield improved simulation of human response, but

none have been officially standardized for manufacturing as in

the case of the Part 572 system. They are briefly summarized

here only to provide background information regarding the state

of the art.

"Repeatable Pete" was developed by the Highway Safety Research

Institute to accurately match the dynamic response of une•-

balmed human cadavers, particularly the head acceleration, for

r9 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.

70. Massing, D. E., Naab, K. N., and Yates, P. E., PERFORMANCE

EVALUATION OF NEW GENERATION OF 50TH-PERCENTILE ANTHROPOMORPHIC TEST DEVICES: VOLUME I - TECHNICAL REPORT, Calspan Corporation; DOT-HS Technical Report 801-431, U. S.

Department of Transportation, National Highway Traffic

Safety Administration, Washington, D. C., March 1975,

PB 240-920.

97

! 7

*.. .: • - J .- -•.
P

, 4.10 440 0
TABLE 9. DUMMY EXTERNAL DIMENSIONS (PART 572)

Part 572

Figure 37 specification Designation code (in.)

Seated height A 35.7 t 0.1

Shoulder pivot height B 22.1 ± 0.3

Hip pivot height C 3.9

Hip pivot from back line D 4.8

Knee pivot from back line E 20.4 ± 0.3

Knee pivot from floor F 19.6 ± 0.3

Head back from back line G 1.7

i I Chest depth U 9.3 t 0.2

( Shoulder width I 18.1 ± 0.3

Chest circumference over

nipples K 37.4 t 0.6

Waist circumference at

minimum girth L 32.0 ± 0.6

Hip width M 14.7 ± 0.7

Popliteal height N* (17.3 ± 0.2)

Shoulder-elbow length Q* (14.1 ± 0.3)

Elbow rest height R* (9.5 1 0.5)

head width S* (6.1 ± 0.2)

Head length T* (7.7 ± 0.2)

Head segment line AA 9.3

Shoulder-thorax segment line BB 25.1

*Added to Part 572 data, SAE specification value in parentheses.

99
TABLE 10. DUMMY COMPONENT WEIGHTS (PART 572)

Part 572

specification

Segment -(b)

Head 11.2 ± 0.1

Upper torso (including lumbar spine) 41.5 ± 1.6

Lower torso (including visceral sac

and upper thighs) 37.5 ± 1.5

Upper arm 4.8 ± 0.2

Lower arm 3.4 ± 0.1

Hand 1.4 ± 0.1

Upper leg 17..6 ± 0.7

Lower leg 6.9 ± 0.3

Foot 2.8 ± 0.1

Total dummy (including instrumentation in head, torso, and femurs) 164.0 ± 3.0

TABLE 11. CENTER-OF-GRAVITY LOCATIONS (PART 572)

Part 572 specificationx z

Segment x and z reference origin (in.) (in.)

Head Back and top of head +4.0 ± 0.2 -4.7 ± 0.1

Upper torso Backline and top of head +4.1 ± 0.3 -17.2 ± 0.3

Lower torso Backline and top of head +4.9 ± 0.5 -31.0 ± 0.5

and upper

thigh

Upper arm Shoulder pivot 0.0 ± 0.3 -5.0 ± 0.3

Lower arm Elbow pivot +4.2 ± 0.3 0.0 ± 0.3

Hand Wrist pivot +2.2 ± 0.3 0.0 ± 0.3

Upper leg Knee pivot to upper leg -6.7 ± 0.3 0.0 ± 0.3

rotation center

Lower leg Knee pivot to ankle pivot 0.0 ± 0.3 -8.0 ± 0.3

Foot Ankle pivot +2.2 ± 0.3 -1.7 ± 0.3

NOTE: Axis system is shown in Figure 37 (drawing of external

dummy dimensions), using +x forward and +z up.

100

77-
TABLE 12. HYBRID II MASS MOMENTS OF (. INERTIA (REFERENCE 70)

Moment of inertia

(in.-lb-sec2)

Body segment Ix Iy z

Head 0.226 0.275

Head/neck 0.310 0.367 0.233

Upper torso (includes

lumbar spine) 2.18 1.79

Lower abdomen, pelvis,

and visceral sac 2.32* 1.73* -

Right upper arm 0.134 0.132 0.022

Right forearm (no hand) 0.012 0.068 0.071

Right upper leg 0.127 0.873 0.890

Right lower leg (no foot) 0.599 0.575 0.359

*Included lumbar spine section.

NOTES: 1. Instrumentation was installed in the head,

chest, and femurs during the measurements.

2. Estimated accuracy of measurements: ±3

percent.

more viable use in determination of head injury criteria (Reference 71). General Motors used detailed human anthropometric

data in designing the GM ATD 502 dummy (Reference 72).

71. McElhaney, J. H., Mate, P. I., and Roberts, V. L., A NEW

CRASH TEST DEVICE - "REPEATABLE PETE," Proceedings, Seventeenth Stapp Car Crash Conference, Society of Automotive

Engineers, New York, 1973, p. 467.

72. Hubbard, R. P., ANTHROPOMETRIC BASIS OF THE GM ATD 502

CRASH TEST DUMMY, Paper 750429, Society of Automotive Engineers, Warrendale, Pennsylvania, 1975.

101
S~ of A di-tinguishing the vertebral feature column for of the an average ATD 502 adult design male is the in an location erect

seated posture. In addition, General Motors has reported development of the Hybrid III dummy (Reference 73). Based on

the ATD 502, this design includes transducers for measurement

of neck loads and chest deflections.

The Transport and Road Research Laboratory in England has developed a specialized side impact dummy for use in side impact

testing of automobiles (Reference 74). The segment joints are

designed for lateral motion, and seven force traiisducers are

used to measure the load distribution between the occupant and

the vehicle interior.

7.3 COMPARISON OF DUMMY AND HUMAN RESPONSE

TLh.re are two basic .questions regarding the use of a mechanical system such as a dummy to evaluate the degree of protection

a vehicle system would afford its human occupants. First, how

closely does the dummy response simulate human response? Secondly, how does performance vary from one dummy to another and

from one test laboratory to another?

Finding an answer to the first question presents a problem.

The response of live human subjects can, of course, be determined only at safe acceleration levels, substantially below

crashworthiness design conditions. The response of human cadavers at higher acceleration levels has been used in dummy

design, but questions do exist in the quality of simulation

provided by cadavers. Recently, Walsh and Romeo reported on a

series of sled tests and full-scale car crash tests wherein

fresh, unembalmed cadavers and dummies were exposed to identical environments (Reference 75). Both belt restraints and air

73. Foster, J. K., Kortge, J. 0., and Wallanin, M. J., HYBRID

III-A BIOMECHANICALLY-BASED CRASH TEST DUMMY, ProceedingsL

Twenty-First Stapp Car Crash Conference, Society of Automotive Engineers, New York, 1977, pp. 973-1014.

74. Harris, J., THE DESIGN AND USE OF THE TRRL SIDE IMPACT

DUMMY, Proceedings, Twentieth Stapp Car Crash Conference,

Society of Automotive Engineers, Warrendale, Pennsylvania,

1976, pp. 75-106.

75. Walsh, Ml. M., and Romeo, D. M., _ESULTS OF CADAVER AND

ANTHROPOMORPHIC DUMMY TESTS IN IDENTICAL CRASH SITUATIONS,

Proceedings, Twentieth Stapp Car Crash Conference, Society

of Automotive Engineers, Warrendale, Pennsylvania, 1976,

pp. 108-131.

102

i'

., ~ •;.C.),: - .4 / - . . .*
Sbags were used. Although the overall kinematic response between cadaver and dummy agreed fairly well, some injuries that

could not have been detected with the dummy were observed in

the cadaver.

Several test programs have been conducted to compare the dynamic response of different dummy designs. Chandler and

Christian demonstrated that, as dummies become more complex,

the number of test variables may exceed those that the experimenter can control (Reference 76). The requirement for standard test practices also was noted. Massing, Naab, and Yates

compared several dummies in tests with either belt restraints,

air bags, or energy-absorbing steering columns (Reference 70).

Some of the sled tests were repeated at two different laboratories. As an example of the results, mean head resultant accelerztions for ten repeated tests on each of five dummies

with belt restraints are shown in Figure 38. Dummies D5 and

D6 are both GM-50X designs (Reference 72), and Al and A2 are

Highway Safety Research Institute (HSRI) devices (Reference 71).

Figure 39 shows the mean head accelerations for ten repeated

tests with the same HSRI dummy conducted at two different facilities, the FAA C.-vil Aeromedical Institute (CAMI) and Calspan Corporation. Because only one dummy was tested at the two

different facilities, data are insufficient to permit generalization with respect to comparative performance. However, Figure 39 indicates that differences in performance did exist.

With respect to dummies tested at the same facility, differences in repeatability were noted. The HSRI dummies were found

to be less repeatable than the GM 5OX dummies, and both of

these were found to be less repeatable than the Hybrid II for

the air bag restraint. As shown in Figure 38, differences do

exist among the performance characteristics of all the dummies.

Detailed analyses of the data are presented in Reference 70.

7.4 SUITABILITY OF DUMMIES FOR AIRCRAFT SYSTEM EVALUATION

All of the recently developed dummies described in Section 7.2

were designed for automotive testing. All are based on the

anthropometry of a 50th-percentile U. S. civilian male. In

dynamic testing of an energy-absorbing seat, design for aircraft occupant weight can play a critical role. It would be

desirable, although generally not practical, to evaluate a seat

for a range of occupant sizes. A 95th-percentile dummy would

verify the strength of the seat structure and restraint system

76. Chandler, R. F., and Christian, R. A., COMPARATIVE EVALUATION OF DUMMY PERFORMANCE UNDER -G IMPACT, Proceedings, "Thirteenth Stapp Car Crash Confererhe, Society of Automo- .IAI tive Engineers, New York, 1969, pp. 61-75.

103

.. '
0

x ~0 a 0

LM -H (D---

U0

m 4JW

N4-1

w0

44)

r- tn N ,L

1040
W- 4-4

005

Ln

00 u 

U0

P-4

0 )

0n

M CD .. LA C40

t. Ln fl. L

1050
as well as the adequacy of the energy-absorbing stroke. Testing with a 50th-percentile dummy would demonstrate the performance of the system for an occupant of average height and

weight. A 5th-percentile dVmmy would probably experience acceleraiions of higher magnitude and would establish the severity of a given set of impact conditions for the smaller occupant. However, both the expense of dummy purchase and the cost

of conducting dynamic tests may make such a test program impractical. An alternative procedure might be to establish the

occupant protection capability of a seat design by analysis

and to conduct a dynamic test with a 95th-percentile dummy to

verify system strength.

There are two additional factors that should be considered in

dummy selection for aircraft seat testing. First, some designs

are more suitable than others for testing with a headward (+G

acceleration component. None of the dummies have been designed

for accurate response to vertical impact. The spinal column,

which is a critical region of human tolerance to aircraft crash

loading, has been designed to simulate response to -Gx loading

rather than the more critical +G direction. However, the re572 dummy permits more consistent positioning than the steel

ball-and-socket configuration used in some other dummies. Instability in the latter type could affect 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 close to that for a human.

Secondly, if the results of tests conducted at different facilities are to be compared, standardization of dummies and test procedures is mandatory.

At present, it seems that use of the Part 572 dummies, modified

to improve their simulation accuracy to impact loading in the

+G direction and sized to 5th-, 50th-, and 95th-percentile

veisions of the U. S. Army aviator, provides the recommended

approach.

*

4e' 106

Sj

.......................................................................... j.. AA.

...........................................................................................
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Sill

• l l I I *• I.r * P
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"I .. .. = ...... .. ="' -- - - '•"I 'T ]•" ' '"" • ' - '-I :':- "-...-........'.-..... .;

' '' 112 I I I I i - i •7••F
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, I
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114
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117
INDEX

Abbreviated Injury Scale (AIS) 60, 73

Abdominal injury tolerance 62

Acceleration

abrupt 20

design values 30

G level defined 17

headward, tolerance to 45

lateral impact 30

pulse shape for design 19, 30

spineward, tolerance to 43

vertical impact 30

Accident

Information Retrieval System 35

investigations 23, 35

records 24, 33

survivable 17

Aircraft

coordinates 18

types 16, 23, 24

Air Force, U. S.

accident reports 23

Animals

experimental 39, 62

Anthropometry 83

Army, U. S.

accident reports 23

anthropometry of aviators 83

Attitude at impact

design velocity change and 29

pitch, distribution 31

roll, distribution 31

yaw, distribution 31

Cadavers

chest impact 60

dummies, comparison with 58

human tolerance experiments 39

inertial properties of 90

leg injury 72

skull fracture 53, 58

vertebral fracture 64, 68 •' ' Center of mass, human body 90, 93

Chest impact tolerance 60

•i' Civil Aeronautics Board 23

Cockpit hazards 75

Concussion

escape prevented by 50

thresholds 54

Crash Survival Design Guide 12

S11
* INDEX (Continued)

Crashworthiness

defined 11

features of aircraft 14

system 15

Crush distance 35

Damage in estimating deceleration 29

Deceleration. See Acceleration

definition of G level 17

DRI. See Dynamic Response Index

Dummies

cadavers in design of 39, 97, 102

history 95

human response compared 5b, 102

inertial properties 97

link lengths 97

moments of inertia 97

Part 572 96

repeatability 103

SAE J963 96

seat testing 103

segment weights 97

side impact 102

Dynamic overshoot 22, 43

Dynamic Response Index 63

Earth scooping 32

Effective Displacement Index

chest injury 60

head injury 57

Ejection seats

DRI for evaluation of 63, 65

Energy absorption

cockpit area 75

ground 14

seat 14, 15

Escape

system design 63

Extremities

cockpit hazards 75

Fixed-wing aircraft 16, 23

accident investigation 23

impdct conditions 33

vertical velocity change 24

Floor deceleration 30

G, defined 17

Head injury

Army aircraft accidents 35

concussion 50

Effective Displacement Index 57

119

-. .V
I

INDEX (Continued)

fatal injuries 35

Severity Index 53

skull fracture 50, 58

Wayne State Tolerance curve 52, 56

Helicopters

attack 31

cargo 31

Human tolerance. See Tolerance

Impact attitude 31

Impact force estimates 29

Injury frequencies 33

Joint rotations 87

J-tolerance 54

Kinematics

human body 75

Landing gear 14

Leg injury 72

Link lengths 83, 93

Load factors 17

Major impact, defined 19

Mass of body segments 90

Mathematical simulators 39, 63, 83

MIL-STD-1290(AV) 12

Moments of inertia 90, 93

Navy, U. S.

accident reports 23

Neck injury 59

Nonsurvivable accidents 23, 31

Onset rate 45, 48

Padding, protective 75

Percentiles

normal distribution 84

seat testing 106

Pitch angle 31

Protective shell 14

Repeatability of dummies 103

Restraint system

failures in estimating deceleration 30

slack 42

submarining 22, 42

"tolerance, effect on 40, 43, 75

vertebral injury affected by 42, 70

Roll angle 31

Rollover accidents 35

Rotor blade impacts 30

Scaling of human body measurements 93

j ' 120

___________4N~~s} ''..
INDEX (Continued)

Seat testing

dummies 103

spinal injury criterion 63

Severity Index 53

Simulators, mathematical 39, 63, 83

Skull fracture 50, 58

Sod impacted most frequently 32

Spinal injury. See Vertebral injury Strike envelopes 75

Submarining 22, 42

Survivable accident, defined 17

Terrain impact 30, 32

Tolerance 39

abdominal impact 62

chest impact 60

deceleration estimated 30

defined 21

factors affecting 40

head injury 50

leg injury 72

neck injury 59

spinal injury 63

whole-body 43

Tree impacts 30

Velocity change

combined 27

lateral 27

longitudinal 20, 25

major impact 19, 24

vertical 20, 24

Vertebral injury 63

aircraft type 33

cadaver data 64

+G impact 46

miialignment reducing tolerance 67

properties 71

restraint affects 41

seating position affects 67

Wayne State Tolerance Curve 52, 56, 58

Weight of Army aviators 85

Yaw angle 31

121-

12, 74 .,- -T 

,
I DEPARTMENT OF THE ARMY

APPLIED TECHNOLOGY LABORATORY

US ARMY RESEARCH AND TECHNOLOGY LABORATORIES (AVRADCOM)

FORT EUSTIS. VIRCINIA 23604

READER'S SERVICE LETTER

(For use in submitting comments, recommendations, corrections, and

revisions for Aircraft Crash Survival Design Guide)

FROM:

TO: Director, Applied Technology Laboratory, US Army Research and Technology

Laboratories (AVRADCOM), ATTN: DAVDL-ATL-ASV, Fort Eustis, Virginia

23604

REMARKS:

i i

SIGNATURE: DATE:

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DEPARTMENT OF THE ARMY

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FORT EUSIIS. VIRGINIA 230o4

READER'S SERVICE LETTER

(For use in submitting comments, recommendations, corrections, and

revisions for Aircraft Crash Survival Design Guide)

FROM:

TO: Director, Applied Technology Laboratory, US Army Research and Technology

Laboratories (AVRADCOM), ATTN: DAVDL-ATL-ASV, Fort Eustis, Virginia t

23604

REMARKS:

1S

I.!

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SIGNATURE: _____ _____DATE: ____________
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APPLIED TECHNOLOGY LABORATORY US ARMY RESEARCH AND TECHNOLOGY LABORATORIES (AVRADCOM)

rORT EUSTIS. VIRGINIA 23604

READER'S SERVICE LETTER

(For use in submitting comments, recommendations, corrections, and

revisions for Aircraft Crash Survival Design Guide)

FROM: II

TO: Director, Applied Technology Laltfratory, US Army Research and Technology

Laboratories (AVRADCOM), ATTN: DAVDL-ATL-ASV, Fort Eustis, Virginia

23604

REMARKS:

S D

I iSIGNATURE: __________DATE: ________

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ri
SUPPLEMENTARY

INFORMATION..... .. .
DEPARTMENT OF THE ARMY

APPLIED TECHNOLOGY LABORATORY

US ARMY RESEARCH AND TECHNOLOGY LABORATORIES (AVRAOCOM)

FORT EUSTIS. VIRGINIA 23604

4 ERRATUM

USARTL-TR-79-22B

TITLE. Aircraft Crash Survival Design Guide

Volume II - Aircraft Crash Environment and Human Tolerance

Insert revised page 28.

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DEPARTMENT OF THE ARMY

APPLIED TECHNOLOGY LABORATORY

US ARMY RESEARCH AND TECHNOLOGY LABORATORIES (AVRADCOM)

17ORT EUSTIS. VIRGINIA 23604

f) ERRATA

USARTL TR 79-22B

TITLE: Aircraft Crash Survival Design Guide

Volume II - Aircraft Crash Environment and Human Tolerance

Make the following changes:

Page 54, first paragraph --

Change (Reference 24),to (Reference 77)

Page 114 -- Add:

77. Slobodnik, B. A., SPH-4 HELMET DAMAGE AND HEAD INJURY

CORRELATION, USAARL Report No. 80-7, U. S. Army Aeromedical

Research Laboratory, Fort' Rucker, Alabama, September 1980.

I.4,

/•